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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.&lt;br/&gt;&lt;br/&gt;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.&lt;br/&gt;&lt;br/&gt;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)).&lt;br/&gt;&lt;br/&gt;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.&lt;br/&gt;&lt;br/&gt;Fig. 1. Vectors in permanent magnet synchronous reluctance machines.&lt;br/&gt;(a) for conventional RMRS machines, (b) for proposed ACRS machines.&lt;br/&gt;&lt;br/&gt;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.&lt;br/&gt;&lt;br/&gt;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.&lt;br/&gt;&lt;br/&gt;Fig. 2. Schematic of axially combined rotor structure (ACRS).&lt;br/&gt;&lt;br/&gt;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.&lt;br/&gt;&lt;br/&gt;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).&lt;br/&gt;[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).

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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).

technical paper

This paper overviews recent developments and novel topologies of a new class of asymmetric rotor pole interior permanent magnet (AIPM) machines for torque enhancement. The principle for torque enhancement by utilizing magnetic-field-shifting is introduced firstly. AIPM machine topologies are then categorized in terms of asymmetries of PM configurations and rotor core geometries. Subsequently, various AIPMs in each category are described and analyzed with respect to merits and demerits. Finally, electromagnetic performances of selected AIPM machines are compared. INDEX TERMS: Asymmetric interior permanent magnet (AIPM), Magnetic-field-shifting (MFS), PM torque, Reluctance torque I. Introduction Interior permanent magnet (IPM) machines having PMs buried inside the rotor cores are of growing interests due to superior torque density, high efficiency, wide constant power operating speed range, and high rotor robustness [1]. In IPMs, the reluctance torque resulted from the rotor saliency can be utilized to enhance the torque density without increasing the PM usage. However, in conventional IPM machines with symmetrical rotor structures, the maximum PM and reluctance torque components are achieved at different current angles whose difference is theoretically 45 electrical degrees. Consequently, only a portion of both torque components can be utilized at the current angle of maximum synthetic torque. To deal with this issue, a novel design concept for torque enhancement of IPMs by making the maximum PM and reluctance torque components to be reached with a smaller difference of current angles has been firstly proposed by setting the optimal displacement angle between the surface-mounted PM rotor and the reluctance rotor in a two-part rotor PM machine [2] and by employing asymmetric rotor flux barriers in a V-shape IPM [3], respectively. To utilize the concept designated as magnetic-field-shifting (MFS) effect in this paper, the PM field and reluctance axes of IPM machines need to be shifted, resulting in a new class of novel asymmetric rotor IPM machines. The torque enhancement principle of AIPM is illustrated in Fig. 1. This paper overviews various novel AIPM topologies which are categorized in terms of asymmetries of PM configurations and rotor core geometries, analyses various AIPMs in each category with respect to merits and demerits, and compares electromagnetic performances of selected AIPM machines. II. AIPM Machine Topologies In this section, novel AIPM machine topologies in literature are categorized according to whether their PM configurations and rotor core geometries are symmetrical or asymmetric, respectively. The proposed categorization method and some examples reported in references for each category are shown in Fig. 2 (a). An AIPM topology with the asymmetric hybrid-layer PM configuration and asymmetric rotor cavity has been proposed in [4], which achieves torque enhancement with simple structure due to utilizing MFS effect and also shows reduced torque ripples. A hybrid-pole AIPM machine is proposed in [5] that employs V-shape and inset PMs in adjacent poles alternatively, which has asymmetric PMs due to shifted positions of inset PMs but symmetrical rotor core geometry. This hybrid-pole AIPM has higher torque density compared with a V-shape IPM with the same PM usage. An AIPM machine with asymmetric PMs and symmetrical rotor structure is also reported in [6] that can achieve fully aligned maximum PM and reluctance torque components, although the increase of synthetic torque is not notable compared with a PM-assisted reluctance machine because of significant reduction of maximum reluctance torque. The asymmetric flux barrier in one side of symmetrical V-shape PMs in each pole is used in [3] to utilize the MFS effect by shifting the angle of reluctance axis and to clearly increase the synthetic torque. Similar to [2], a two-part rotor design with a SPM rotor and a multi-layer reluctance rotor is proposed in [7] that exhibit symmetrical PM configuration and rotor cavity design, respectively. The MFS effect is utilized by adjusting the displacement angle between both rotor parts, which can also achieve fully aligned maximum PM and reluctance torque components and shows a clear torque enhancement. Four selected AIPM machines with rotor topologies proposed in references [3]-[5] and [7] are redesigned by using the same stator, rotor diameter and PM usage to the conventional Toyota Prius 2010, and their torque performances at 750 r/min and 236A are compared in Fig. 2 (b). Comprehensive overview of more AIPM topologies will be given in the full paper.

**References:**

[1] Z. Zhu and D. Howe, Proc. IEEE, vol. 95, p.746-765 (2007). [2] P. Winzer and M. Doppelbauer, 2014 Int. Conf. Electr. Mach. (ICEM), Berlin, 2014, p. 641-647. [3] W. Zhao, F. Zhao, T. A. Lipo, and B.-I. Kwon, IEEE Trans. on Magn., vol. 50, no. 11, p. 1-4 (2014). [4] Y. Xiao, Z. Q. Zhu, J. T. Chen, D. Wu, and L. M. Gong, 2020 Int. Conf. Electr. Mach. (ICEM), Gothenburg, 2020, p. 26-32. [5] G. Xu, G. Liu, W. Zhao, Q. Chen, and X. Du, IEEE Trans. Ind. Electron., vol. 66, no. 4, p. 2580-2591 (2018). [6] F. Xing, W. Zhao, and B.-I. Kwon, IET Electr. Power Appl., vol. 13, no. 5, p. 573-580 (2018）. [7] H. Yang, Y. Li, H. Y. Lin, Z. Q. Zhu, et al., 2017 IEEE Energy Convers. Congr. Exposition (ECCE), Cincinnati, 2017, pp. 2362-2367.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/339b97d20cc7f72fd8fd7d97e5455425.jpg)
Fig.1.Principle of torque enhancement of AIPM. (a) Conventional IPM. (b) AIPM with reduced △β. Tsyn, Tpm and Trel denote maximum synthetic, PM, and reluctance torques, respectively, and △β denote the current angle difference between maximum PM and reluctance torque components.

![](https://assets.underline.io/uploads/markdown_image/1/image/4d44aa957f2775e6ab24c779072c37ef.jpg)
Fig.2. Categorization of some typical AIPM machine topologies. In (a), PMs in red and blue denote N- and S-poles.

 Novel Asymmetric Rotor Pole Interior Permanent Magnet Machines with Enhanced Torque Density: An Overview

I. Introduction With the rapid development of electric vehicles and hybrid electric vehicles, they put forward more critical and complicated requirements for driving motor drive systems. Double stator permanent magnet (DSPM) motors have become promising candidates for driving motors, due to its advantages of high torque/power density and relatively compact structure. DSPM motor can be considered as the integration of inner and outer motors. By the flexible control of inner and outer motors of double stator motors, different operation modes of motors can be realized. In [1], a double stator permanent magnet motor is proposed, and corresponding multi-mode characteristics are investigated. It is noted that current research on double stator motors mainly biased on proposing novel motor topologies, realizing different operation modes, and verifying motor electromagnetic performances. However, the reasonable power split and allocation of inner and outer motors are seldom investigated, especially for the multi-mode operation. Hence, it is very necessary and essential to purposefully determine power ratio of inner and outer motors, considering design requirements of multiple operation modes. Split ratio, which termed as the ratio of inner and outer diameter of stator, often determines motor magnetic and electrical loading. It possesses significant influences on motor performances, including torque output capability, loss characteristics, and efficiency. So, motor split ratio is considered as one of most crucial motor design parameters and often investigated during the motor initial design stage. In [2], the split ratio of surface-mounted permanent magnet motor is analyzed and the optimal motor split ratio is decided by analytical method. While for double stator motors, split ratios of inner and outer motors determine not only motor magnetic and electrical loading, and but also the power split characteristics of the whole motor. However, it is noted current publications of motor split ratio mainly focus on the split ratio of single-stator motors which mainly work on rated condition. Since double stator motor are required to work in different operation modes, traditional design methods of motor split ratio cannot be applied for double stator motors directly. Consequently, it has been a hot but challenging issue on how to determine split ratio of double stator permanent magnet motors during the initial design stage, considering the design requirements of multiple operation modes. In this paper, a double stator permanent magnet (DSPM) motor is investigated, and the relationship between motor split ratio and power split characteristics of different motor operation mode is analyzed in detail. Furthermore, by the purposeful design and optimization of motor split ratio, design requirements of multiple operation modes can be satisfied comprehensively. II. Motor topology and split ratio investigation Fig. 1 presents the configuration of the investigated DSPM motor. As can be seen, inner stator and middle rotor form inner motor, while outer motor consists of outer stator and middle rotor. In order to realize multimode operation effectively, the torque-speed curves of inner and outer motor are designed purposefully, as shown in Fig. 2. The inner motor is designed to have a relatively flat torque curve over a wide speed range, which is utilized to augment the output torque of outer motor, especially in the range of torque of outer motor is dropping off. In addition, driving requirements of whole DSPM motor are arranged purposefully as optimization objectives for the inner and outer motors of DSPM motor. So, by the purposeful and reasonable design and optimization of split ratios of inner and outer motors, design requirements of multi-mode operation can be satisfied. III. Optimization results and performances validation In order to satisfy design requirements of multiple operation modes, response surface method and multi-objective genetic algorithm are utilized for DSPM motor, the optimal motor split ratio and pole teeth ratio can be determined. Distribution of optimization objectives and design parameters are presented in Fig. 3 and Fig. 4 respectively. And, based on finite element analysis, electromagnetic performances of DSPM motor can be obtained. Fig. 5(a) depicts the flux lines and flux densities distributions of DSPM motor after optimization. Table I presents the multi-mode performance comparison of initial and optimal DSPM motor. And, torque performances of overall DSPM motor before and after optimization are shown in Fig. 5(b). As can be seen, after optimization, the average output torque increases from 71.83Nm to 82.81Nm, while the torque ripple reduces from 10.93% to 7.78%. Electromagnetic performances verify that, by the reasonable design and cooperation of inner and outer motors of DSPM motor, design requirements of different operation mode can be realized more effectively, which provides a research path for the optimization design of multi-port PM motors. More detailed theoretical analysis, optimization design results, and experimental validations will be presented in the full paper.

**References:**

[1] Y. Wang, W. Xu, X. Zhang, and W. Ma, “Harmonic analysis of air gap magnetic field in flux-modulation double-stator electrical-excitation synchronous machine,” IEEE Trans. Ind. Electron., vol. 67, no. 7, pp. 5302-5312, Jul. 2020. [2] Y. Ni, Z. Liu, B. Xiao, and Q. Wang, “Optimum split ratio in surface-mounted permanent magnet machines with pieced halbach magnet array,” IEEE Trans. Energy Convers., vol. 35, no. 4, pp. 1877-1885, Dec. 2020.

**Images:**
![](https://assets.underline.io/uploads/markdown_image/1/image/d6b43446fde4a1a6ef7b49cb73d872c0.jpg)

![](https://assets.underline.io/uploads/markdown_image/1/image/ae52277ced06753e709cdcc2710dce16.jpg)

 Split Ratio Investigation of Double Stator Permanent Magnet Motor Considering Multimode Operation

High torque-density, high efficiency and robust rotor structure are the key features of flux switching motors (FSMs) [1]. Permanent magnet (PM) FSMs are preferred choice, however, these machines consume high amount rare-earth magnet and its resources are depleting with the passage of time. The degradation of performance with increase in temperature and uncontrolled flux are the other concerns with PMFSMs. The flux of electrically excited (EE) FSM can be controlled electronically, performance is not degraded with increase in temperature and overall cost is low. The simple manufacturing process of single phase EEFSM, easy maintenance, longer lifetime and low cost are the main characteristic that prioritized it for high speed applications [2-3]. Authors [4-6] have proposed novel single phase EEFSM designs, they have successfully accomplished their targets in the form of good average torque and efficiency etc. These designs have overlapped windings and segmented or modular rotor. High copper losses are caused by overlapped windings and segmented or modular rotor is not suitable for high speed applications. High power density, torque and efficiency are key factors for motors used in high speed applications. Flux switching motor (FSM) designs with the conventional stator slot have relatively good average torque but high copper losses and low efficiency. This paper has proposed an optimized octane modular stator (OMS) 8slots-6poles (8S-6P) single phase electrically excited (EE) FSM design. The geometric representation of the OMS design is given in Fig.1 (a), Rse, Rre, βsθs, βrθr, and θ are stator outer radius, rotor outer radius, stator slot opening, rotor slot opening and rotor position respectively. The proposed OMS design has high copper slot filling factor and high efficiency than the conventional stator trapezoidal slot [3]. Moreover, the purpose of this paper is to analyze OMS EEFSM design by means of simplified analytical method. For electromagnetic performance investigation numerical based technique finite element analysis (FEA) is powerful tool. But, computational complexity, high computational time and high drive storage compel researchers to model and analyze electric machines and especially FSMs analytically. In this paper, doubly-salient air-gap permeance, magneto-motive force (MMF) and inductance calculation is performed to obtain radial magnetic flux density distribution in the air-gap and electromagnetic torque, accounts the influence of all parts of EEFSM. This simplified approach has reduced the computational time, drive storage and computational complexity. Geometric optimization (GO) is implemented in multiple steps to optimize the initial design. The analytical results are validated via JMAG v18.1 finite element analysis (FEA) simulator. Finally, prototype is fabricated and experimentally tested. Prototype stator without windings and with windings are shown in Fig.1 (b) and (c) respectively. The analytical results has good agreement with FEA and experimental results and the deviancy is from 3 to 4%. Overall GO implementation, average torque of the proposed optimized EEFSM design is enhanced 42.5% and comparison of the electromagnetic performance is plotted in Fig.2.

**References:**

[1] X. Sun, Z. Shi, G. Lei, Y. Guo, and J. Zhu, “Multi-objective design optimization of an IPMSM based on multilevel strategy”, IEEE Transactions on Industrial Electronics (2020). [2] M. Lihong, Z. Gangxu, Z. Tao, L. Qing and L. Baolian, “Multi-level optimization design for a Flux-Concentrating Permanent-Magnet Brushless Machine Considering PM demagnetization limitation”, IEEE Transactions on Magnetics (2020). [3] B. Khan, F. Khan, W. Ullah, M. Umair and S. Hussain, “Slot Filling Factor Calculation and Electromagnetic Performance of Single Phase Electrically Excited Flux Switching Motors”, Applied Computational Electromagnetics Society Journal, 35(8), pp.922-928, (2020). [4] M. F. Omar, E. Sulaiman, M. Jenal, R. Kumar and R. N. Firdaus, “Magnetic flux analysis of a new field-excitation flux switching motor using segmental rotor”, IEEE Transactions on Magnetics, 53(11), pp.1-4, (2017). [5] Y. J. Zhou and Z. Q. Zhu, “Comparison of low-cost single-phase wound-field switched-flux machines”, IEEE Transactions on Industry Applications, 50(5), pp.3335-3345, (2014). [6] J. Yuan, D. Meng, G. Lian, J. Zhang, H. Li and F. Ban, “The stator slot-type optimization of electrical excitation flux-switching motor and its maximum torque/copper loss control”, IEEE Transactions on Applied Superconductivity, 29(2), pp.1-5, (2019).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/182a3564aced1ceec14372100b2edd6c.jpg)
Fig.1: Proposed OMS single phase 8s-6p EEFSM (a) geometry (b) stator (c) stator with windings

![](https://assets.underline.io/uploads/markdown_image/1/image/43f48b1a116f8f4d1324142f6b37b415.jpg)
Fig.2: Performance of OMS 8S-6P EEFSM (a) Flux density (b) Flux linkage (c) Torque (d) Back-emf

 Analytical Modelling, Optimization and Electromagnetic Performance Analysis of Electrically Excited Flux Switching Motor

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

I.Introduction The vertical axis wind turbine has been widely concerned because of its advantages such as multiphase wind, stable force, easy installation and maintenance, and small starting wind speed. Vertical axis wind turbines mostly use permanent magnet synchronous generators or doubly fed induction wind turbines, and most of them are axial flux permanent magnet generators[1]. Compared with doubly fed induction wind turbine, permanent magnet synchronous generator (PMSG) has the advantage of high power density and is the main research object of vertical axis wind turbine. However, the direct coupling of the high power axial flux permanent magnet synchronous generator and wind turbine causes the starting of the wind turbine to become more difficult. In order to improve the starting performance of vertical axis wind turbines, maglev technology is considered to reduce the starting resistance as the development direction of vertical axis wind turbines. Therefore, an asymmetric-primary axis-flux hybrid- electricity generator is proposed in this paper, and the decoupling control of levitation force and output power is implemented using asymmetric primary and hybrid excitation. II.operation principle The proposed asymmetric-primary axis-flux hybrid-excitation generator for the vertical axis wind turbine structure is shown in the Fig.1. The generator is composed of double stator primary and single actuator. The double stator primary has an asymmetric structure design, and the rotor shaft is connected with the wind turbine shaft to realize direct drive wind power generation. The rotor adopts the spacing design of permanent magnet and modulating tooth to reduce the consumption of permanent magnet. Both the excitation winding and the armature winding are located in the primary stator, so the mixed excitation and magnetic field can be adjusted without adding any mechanism such as brush slip ring, which reduces the complexity of the structure. In general, the proposed generator has the following characteristics (1) The flat structure shortens the axial length of the generator and is suitable for direct drive of vertical axis wind power generation system. (2) The levitation force is provided by bilateral unbalance design of axial flux. The two sides adopt different primary sizes, and then the axial electromagnetic force imbalance of the bilateral air gap forms the suspension force, which avoids adding active and passive levitation mechanisms and reduces the complexity of the overall system. (3) The levitation force and output power are controlled by hybrid excitation. The primary tooth DC excitation is adopted to realize zero crosslinking of the excitation with permanent magnet and armature, and the levitation force regulation does not affect the output power. The system can start at low wind speed, and the levitation force and total output power can be adjusted according to the change of wind speed. Through free distribution of bilateral power to achieve the total power following the wind speed and the levitation force following the lift force of the wind turbine. (4) Using magnetic field modulated axial magnetic field structure, the rotor structure is simple and reliable. If the pole pair numbers of rotor and armature are pr and ps, and the number of the stator tooth is pn, there is pr = pn±ps. The design makes it have the magnetic field modulation effect, which can increase the magnetic field change rate under the low speed, thus increasing the generation power. (5) The use of concentrated armature windings and excitation windings reduces the end, thus reducing magnetic leakage and losses. III.simulation and results A three-dimensional finite element model of the asymmetric-primary axis-flux hybrid-excitation generator is established, and the no-load characteristics and load characteristics of the generator are obtained. The no-load characteristics under different excitation current are simulated, and the flux linkage of armature coils and DC excitation coils are shown in Fig.2. We can see that when different excitation currents are applied, the crossing flux linkage between DC excitation winding and secondary permanent magnet is almost 0, and DC flux has little effect on the armature flux. Therefore, When the excitation changes, the unilateral levitation force changes only in amplitude, but there is no harmonic phase shift, and positioning torque is almost unaffected. he proposed generator has good decoupling performance of levitation force, output power, and DC excitation. In addition, the optimization design of the generator is further studied such as pole arc coefficient, positioning moment and pole slot ratio. IV.conclusion In this paper, we proposed an asymmetric-primary axis-flux hybrid-excitation generator, studied the no-load characteristics, load characteristics and excitation characteristics, and optimized the structural parameters of the generator. The simulation results are obtained, and state that the generator can realize the decoupling of the levitation force of excitation and power, which is suitable for vertical axis wind turbine.

**References:**

[1]Ghulam Ahmad, Uzma Amin, Design, construction and study of small scale vertical axis wind turbine based on a magnetically levitated axial flux permanent magnet generator[J]. Renewable Energy, 101(2017): 286-292, 2017

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/d924fadd5a021735274eef34834cd215.jpg)
Fig.1. The structure of the proposed asymmetric-primary axis-flux hybrid-excitation generator

![](https://assets.underline.io/uploads/markdown_image/1/image/8d0f10e0585dcc735e6d439ba8d2b921.jpg)
Fig. 2. The performance under different excitation current

 Research on an Asymmetric-primary Axis-flux Hybrid-excitation Generator for the Vertical Axis Wind Turbine

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.

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