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15 pages, 8380 KiB  
Article
Design and Analysis of a Low Torque Ripple Permanent Magnet Synchronous Machine for Flywheel Energy Storage Systems
by Yubo Sun, Zhenghui Zhao and Qian Zhang
Energies 2024, 17(24), 6337; https://doi.org/10.3390/en17246337 - 16 Dec 2024
Viewed by 381
Abstract
Flywheel energy storage systems (FESS) are technologies that use a rotating flywheel to store and release energy. Permanent magnet synchronous machines (PMSMs) are commonly used in FESS due to their high torque and power densities. One of the critical requirements for PMSMs in [...] Read more.
Flywheel energy storage systems (FESS) are technologies that use a rotating flywheel to store and release energy. Permanent magnet synchronous machines (PMSMs) are commonly used in FESS due to their high torque and power densities. One of the critical requirements for PMSMs in FESS is low torque ripple. Therefore, a PMSM with eccentric permanent magnets is proposed and analyzed in this article to reduce torque ripple. Cogging torque, a significant contributor to torque ripple, is investigated by a combination of finite element analysis and the analytical method. An integer-slot distribution winding structure is adopted to reduce vibration and noise. Moreover, the effects of eccentric permanent magnets and harmonic injection on the cogging torque are analyzed and compared. In addition, the electromagnetic performance is analyzed, and the torque ripple is found to be 3.1%. Finally, a prototype is built and tested, yielding a torque ripple of 3.9%, to verify the theoretical analysis. Full article
(This article belongs to the Special Issue Energy, Electrical and Power Engineering: 3rd Edition)
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<p>Topology of proposed PMSM.</p>
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<p>Winding connection.</p>
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<p>Analysis model of surface-mounted PMSM.</p>
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<p>Bread-type eccentric permanent magnet.</p>
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<p>Influence of eccentricity on torque performance.</p>
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<p>Cogging torque of PMSM with different permanent magnets.</p>
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<p>Air gap magnetic densities of PMSM with different permanent magnets. (<b>a</b>) Radial air gap magnetic densities. (<b>b</b>) Tangential air gap magnetic densities. (<b>c</b>) Harmonic order.</p>
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<p>Cogging torque contribution of different harmonics. (<b>a</b>) PMSM with original permanent magnets. (<b>b</b>) PMSM with eccentric permanent magnets.</p>
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<p>Permanent magnet with third harmonic injection.</p>
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<p>Cogging torque of PMSM with harmonic injection. (<b>a</b>) Effect of harmonic injection. (<b>b</b>) Contribution of harmonics.</p>
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<p>Load electromagnetic performance. (<b>a</b>) Magnetic field line. (<b>b</b>) Magnetic flux density.</p>
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<p>Back electromotive force of PMSM.</p>
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<p>Torque performance of PMSM. (<b>a</b>) Cogging torque. (<b>b</b>) Torque.</p>
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<p>Vibration acceleration of PMSM with different permanent magnets.</p>
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<p>Prototype.</p>
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<p>Experimental platform.</p>
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<p>Vibration and noise test platform.</p>
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<p>Comparison of experimental and simulated results. (<b>a</b>) Back electromotive force of prototype. (<b>b</b>) Comparison of back electromotive force coefficient.</p>
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<p>Experimental results. (<b>a</b>) Torque. (<b>b</b>) Vibration acceleration.</p>
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14 pages, 51619 KiB  
Article
Current Harmonics Suppression of Six-Phase Permanent-Magnet Synchronous Motor Drives Using Back-Electromotive Force Harmonics Compensation
by Po-Sheng Huang, Cheng-Ting Tsai, Jonq-Chin Hwang, Cheng-Tsung Lin and Yu-Ting Lin
Energies 2024, 17(24), 6280; https://doi.org/10.3390/en17246280 - 12 Dec 2024
Viewed by 471
Abstract
This paper investigates a back-electromotive force (EMF) harmonic compensation strategy for six-phase permanent-magnet synchronous motors (PMSMs) to reduce current harmonics and improve system performance. Ideally, the back-EMF waveform should be perfectly sinusoidal. However, manufacturing imperfections such as suboptimal magnetic circuit design, uneven winding [...] Read more.
This paper investigates a back-electromotive force (EMF) harmonic compensation strategy for six-phase permanent-magnet synchronous motors (PMSMs) to reduce current harmonics and improve system performance. Ideally, the back-EMF waveform should be perfectly sinusoidal. However, manufacturing imperfections such as suboptimal magnetic circuit design, uneven winding distribution, and mechanical eccentricity introduce low-order spatial harmonics, particularly the 5th, 7th, 11th, and 13th orders, which distort the back-EMF, increase current harmonics, complicate control, and reduce efficiency. To address these issues, this study proposes a compensation strategy utilizing common-mode and differential-mode current control. By injecting the 6th and 12th harmonics into the decoupled voltage commands along the d-axis and q-axis, the strategy significantly reduces current harmonic distortion. Experimental validation was conducted using a TMS320F28386D microcontroller, which controlled dual inverters via PWM signals and processed real-time current feedback. Rotor position feedback was provided by a resolver to ensure precise and responsive motor control. At a rotational speed of 900 rpm, with a peak phase current Im of 200 A and an IGBT switching frequency of 10 kHz, the phase-a current total harmonic distortion (THD) was reduced from 11.86% (without compensation) to 6.83% (with compensation). This study focused on mitigating harmonics below the 14th order. The experimental results demonstrate that the proposed back-EMF harmonic compensation strategy effectively minimizes current THD, highlighting its potential for improving the performance and efficiency of multi-phase motor systems. Full article
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<p>Six-phase PMSM back-EMF measurement system block diagram.</p>
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<p>Phase-<span class="html-italic">a</span> and phase-<span class="html-italic">x</span> back-EMF measurement.</p>
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<p>The waveform of the measured and reconstructed phase-<span class="html-italic">a</span> and phase-<span class="html-italic">x</span> back-EMF: (<b>a</b>) phase-<span class="html-italic">a</span> and phase-<span class="html-italic">x</span> back-EMF waveform; (<b>b</b>) zoomed-in waveform.</p>
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<p>Six-phase PMSM common-mode and differential-mode current closed-loop control block.</p>
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<p>Simulated phase-<span class="html-italic">a</span> current waveform and harmonic histogram: (<b>a</b>) Without compensation. (<b>b</b>) With compensation.</p>
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<p>Histogram of the THD reduction in simulation result.</p>
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<p>Photos of the six-phase PMSM testbench: (<b>a</b>) Setup with the dynamometer driving the six-phase PMSM. (<b>b</b>) Six-phase PMSM drive system.</p>
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<p>Actual phase-<span class="html-italic">a</span> current waveform and harmonic histogram: (<b>a</b>) Without compensation. (<b>b</b>) With compensation.</p>
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<p>Histogram of the THD reduction in actual result.</p>
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15 pages, 6360 KiB  
Article
Experimental Determination of Influences of Static Eccentricities on the Structural Dynamic Behavior of a Permanent Magnet Synchronous Machine
by Julius Müller, Marius Franck, Kevin Jansen, Gregor Höpfner, Jörg Berroth, Georg Jacobs and Kay Hameyer
Machines 2024, 12(9), 649; https://doi.org/10.3390/machines12090649 - 16 Sep 2024
Viewed by 720
Abstract
In electrified vehicles, the masking noise behavior of internal combustion engines is absent, making the tonal excitation of the electric machine particularly noticeable in vehicle acoustics, which is perceived as disturbing by consumers. Due to manufacturing tolerances, the tonal NVH characteristics of the [...] Read more.
In electrified vehicles, the masking noise behavior of internal combustion engines is absent, making the tonal excitation of the electric machine particularly noticeable in vehicle acoustics, which is perceived as disturbing by consumers. Due to manufacturing tolerances, the tonal NVH characteristics of the electric machine are significantly influenced at wide frequency ranges. This paper presents a systematic exploration of the influence of static eccentricity as one manufacturing tolerance on the NVH behavior of Permanent Magnet Synchronous Machines (PMSMs). The study utilizes a novel test bench setup enabling isolated variations in static eccentricity of up to 0.2 mm in one PMSM. Comparative analysis of acceleration signals reveals significant variations in the dominance of excitation orders with different eccentricity states, impacting critical operating points and dominant frequency rages of the electric machine. Despite experimentation, no linear correlation is observed between increased eccentricity and changes in acceleration behavior. Manufacturing eccentricity and deviations in rotor magnetization are discussed as potential contributors to the observed effects. The findings emphasize static eccentricity as a critical parameter in NVH optimization, particularly in electrified powertrains. However, the results indicate that further investigations are needed to explore the influence of eccentricities and magnetization deviations on NVH behavior comprehensively. Full article
(This article belongs to the Special Issue Noise and Vibrations of Electrical Machines)
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<p>PMSM test bench setup.</p>
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<p>Sensor setup at test bench.</p>
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<p>Schematic setup of the test bench environment.</p>
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<p>Geometric determination of the rotor position.</p>
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<p>Laser signal from sensor 1 at 100 rpm.</p>
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<p>Measured static eccentricities for different eccentricity states.</p>
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<p>Rotor speed and position during the measured run-up scenarios.</p>
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<p>Total acceleration level of sensor number 5 in normal direction for different static eccentricity states.</p>
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<p>Acceleration signal of the first static eccentricity state in the normal direction at sensor position number 5.</p>
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<p>Acceleration signal of the sixth static eccentricity state in normal direction at the sensor position number 5.</p>
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<p>Order level for different eccentricity states in the normal direction, at sensor position number 5.</p>
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<p>Order level for different eccentricity states in the normal direction, at sensor position number 5.</p>
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10 pages, 2920 KiB  
Communication
Optimization Design of Cogging Torque for Electric Power Steering Motors
by Guoguang Zhang and Peng Hou
Machines 2024, 12(8), 517; https://doi.org/10.3390/machines12080517 - 30 Jul 2024
Cited by 1 | Viewed by 804
Abstract
Excessive cogging torque can cause torque fluctuations, noise, and vibration in electric power steering (EPS) motors, which is a key factor in the high-precision and high-performance optimization design of EPS motors for electric vehicles. This article takes a 12-slot 10-pole electric power steering [...] Read more.
Excessive cogging torque can cause torque fluctuations, noise, and vibration in electric power steering (EPS) motors, which is a key factor in the high-precision and high-performance optimization design of EPS motors for electric vehicles. This article takes a 12-slot 10-pole electric power steering motor for a certain car as an example. By establishing the corresponding electromagnetic field model and theoretical analysis of the motor, the influence of the pole arc coefficient and eccentricity parameters of the permanent magnet on the cogging torque of the electric power steering motor is explored. A comprehensive optimization scheme for reducing the cogging torque of the motor structure is proposed. The effectiveness of the designed scheme was verified through finite-element simulation and experimental testing of motor electromagnetism. Compared with the original design, the optimized structure of the EPS motor resulted in an 86.62% reduction in cogging torque during experimental testing. Full article
(This article belongs to the Special Issue Optimal Design and Drive of Permanent Magnet Synchronous Motors)
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<p>Finite-element simulation calculation model of the EPS motor.</p>
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<p>Cogging torque of the EPS motor with different pole arc coefficients.</p>
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<p>Permanent magnet with an unequal thickness.</p>
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<p>Cogging torque with different eccentricities.</p>
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<p>The variation curve of cogging torque with the polar arc coefficient and permanent eccentricity.</p>
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<p>Optimized EPS motor magnetic density cloud diagram and magnetic field line diagram.</p>
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<p>Cogging torque of the EPS motor before and after optimization.</p>
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<p>Schematic diagram of the motor cogging torque testing bench.</p>
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<p>EPS motor cogging torque on-site testing platform.</p>
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20 pages, 7543 KiB  
Article
Analytical Modeling of Eddy Current Losses and Thermal Analysis of Non-Uniform-Air-Gap Combined-Pole Permanent Magnet Motors for Electric Vehicles
by Shilun Ma, Jianwei Ma, Keqi Chen and Changwei Li
Machines 2024, 12(6), 377; https://doi.org/10.3390/machines12060377 - 31 May 2024
Viewed by 794
Abstract
In order to solve the problem of large eddy current losses and high temperature rises caused by a large number of permanent magnets, a new type of combined-magnetic-pole permanent magnet motor is proposed in this paper. The sinusoidally distributed subdomain model of a [...] Read more.
In order to solve the problem of large eddy current losses and high temperature rises caused by a large number of permanent magnets, a new type of combined-magnetic-pole permanent magnet motor is proposed in this paper. The sinusoidally distributed subdomain model of a non-uniform-air-gap rotor was established using the Laplace equation, and the analytical expression of eddy current losses in the rotor in a uniform air gap and non-uniform air gap was derived. The effect of the rotor’s eccentricity on eddy current losses was obtained. According to the characteristics of the distributed winding of the non-uniform-air-gap combined-pole permanent magnet motor, an equivalent treatment was performed to obtain the equivalent thermal conductivity value; to establish an equivalent thermal network model of the motor; determine the temperature of each component of the motor; and verify the correctness of the thermal network model through magnetothermal bidirectional coupling. Finally, an experimental platform was set up to carry out temperature rise experiments on the two prototypes. The experimental results show that a non-uniform-air-gap rotor structure can effectively reduce a rotor’s eddy current losses and motor temperature rise, as well as verify the accuracy of the analytical model’s calculation results. Full article
(This article belongs to the Section Vehicle Engineering)
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Figure 1
<p>Structure diagram of an ITRPMM. 1. Stator core; 2. armature winding; 3. semicircular radial magnetic field permanent magnet (SRPM); 4. radial magnetic field rectangular permanent magnet (RRPM); 5. rotor core; 6. tangential magnetic field rectangular permanent magnet (TRPM).</p>
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<p>Rotor structure with a non-uniform air gap.</p>
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<p>Distribution of the air gap when the magnetic field is sinusoidal.</p>
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<p>Eddy current loss model of a silicon steel sheet.</p>
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<p>Influence of rotor eccentricity on eddy current losses. (<b>a</b>) Relationship between eddy current losses and rotor eccentricity. (<b>b</b>) Finite element simulation results with and without rotor eccentricity.</p>
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<p>Equivalent thermal network model of the motor.</p>
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<p>Equivalent model of armature winding. (<b>a</b>) Initial model of stator winding. (<b>b</b>) Equivalent model of stator winding.</p>
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<p>Flowchart of magnetothermal bidirectional coupling.</p>
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<p>Temperature rise of each part of the motor with a uniform air gap structure at a rated load for 120 min of operation. (<b>a</b>) Temperature rise of the rotor core. (<b>b</b>) Temperature rise of the permanent magnet. (<b>c</b>) Temperature rise of the stator core. (<b>d</b>) Temperature rise of the armature winding. (<b>e</b>) Temperature rise of the motor housing. (<b>f</b>) Temperature rise of the shaft.</p>
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<p>Temperature rise of each part of the motor with a non-uniform air gap structure at a rated load for 120 min of operation. (<b>a</b>) Temperature rise of the rotor core. (<b>b</b>) Temperature rise of the permanent magnet. (<b>c</b>) Temperature rise of the stator core. (<b>d</b>) Temperature rise of the armature winding. (<b>e</b>) Temperature rise of the motor housing. (<b>f</b>) Temperature rise of the shaft.</p>
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<p>Prototype and temperature rise experiment platform. (<b>a</b>) Uniform air gap rotor structure. (<b>b</b>) Non-uniform air gap rotor gap rotor structure. (<b>c</b>) Stator and armature winding. (<b>d</b>) Diagram of the test platform.</p>
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<p>Mechanical characteristic curves of the ITRPMM.</p>
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<p>Thermal imaging camera.</p>
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<p>Steady-state temperature cloud.</p>
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<p>Maximum temperature rise contrast curves of the armature winding.</p>
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<p>Experimental platform of the no-load back EMF.</p>
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<p>Measured waveform of the no-load back EMF of the prototype at rated speed. (<b>a</b>) No-load back EMF waveform with a uniform air gap. (<b>b</b>) No-load back EMF waveform with a non-uniform air gap.</p>
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<p>Harmonic amplitude of the no-load induced electromotive force.</p>
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22 pages, 7584 KiB  
Article
Diagnostics of Interior PM Machine Rotor Faults Based on EMF Harmonics
by Natalia Radwan-Pragłowska and Tomasz Wegiel
Energies 2024, 17(9), 2198; https://doi.org/10.3390/en17092198 - 3 May 2024
Cited by 1 | Viewed by 732
Abstract
This article presents a detailed study on the diagnosis of rotor faults in an Interior Permanent Magnet Machine based on a mathematical model. The authors provided a wide literature review, mentioning the fault diagnosis methods used for Permanent Magnet Machines. The research emphasizes [...] Read more.
This article presents a detailed study on the diagnosis of rotor faults in an Interior Permanent Magnet Machine based on a mathematical model. The authors provided a wide literature review, mentioning the fault diagnosis methods used for Permanent Magnet Machines. The research emphasizes the necessity of precise assumptions regarding winding construction to accurately analyze the additional harmonics appearing in rotor faults caused by electromotive force (EMF), i.e., rotor eccentricity and magnet damage. The article also discusses specific features appearing in the spectrum of air gap permeance functions and the impact of rotor eccentricity and magnet damage on PM flux density distribution and as a consequence on EMF stator windings. The novelty of the presented content is the analysis of induced EMFs for cases of the simultaneous occurrence of rotor eccentricity and PM damage. The findings of this study provide valuable insights for the diagnosis and understanding of internal asymmetries in Interior PM Machines. Full article
(This article belongs to the Special Issue New Solutions in Electric Machines and Motor Drives: 2nd Edition)
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<p>A simplified cross-section of an interior PM machine.</p>
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<p>Explanation of the calculation of magnetic line lengths.</p>
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<p>PM flux density distribution in the air gap for symmetrical and slot-less machines.</p>
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<p>Modeling damage to one magnetic pole.</p>
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<p>Modeling damage to one pair of magnetic poles.</p>
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<p>A simplified model illustrating the location of the windings of a PM machine.</p>
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<p>Stator winding factors of an example PM machine according to “<span class="html-italic">p</span>” harmonic index.</p>
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<p>Permeance function: (<b>a</b>) full air-gap symmetry; (<b>b</b>) static eccentricity <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>; (<b>c</b>) dynamic eccentricity <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>; and (<b>d</b>) mixed eccentricity <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>.</p>
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<p>Flux density: (<b>a</b>) symmetry of air-gap; (<b>b</b>) symmetry of air-gap and PM demagnetization.</p>
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<p>Flux density: (<b>a</b>) static eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>); (<b>b</b>) static eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>) and PM demagnetization.</p>
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<p>Flux density: (<b>a</b>) dynamic eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>); (<b>b</b>) dynamic eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>) and PM demagnetization.</p>
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<p>Flux density: (<b>a</b>) mixed eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math>); (<b>b</b>) mixed eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math>) and PM demagnetization.</p>
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<p>Example FEM calculations of flux density distribution.</p>
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<p>Stator phase back EMF: (<b>a</b>) air-gap symmetry; (<b>b</b>) air-gap symmetry and PM demagnetization (analytical calculations—dashed line ending with the marker “o”; FEM analysis—solid line).</p>
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<p>Stator phase back EMF: (<b>a</b>) static eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>); (<b>b</b>) static eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>) and PM demagnetization (analytical calculations—dashed line ending with the marker “o”; FEM analysis—solid line).</p>
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<p>Stator phase back EMF: (<b>a</b>) dynamic eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>); (<b>b</b>) dynamic eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>) and PM demagnetization (analytical calculations—dashed line ending with the marker “o”; FEM analysis—solid line).</p>
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<p>Stator phase back EMF: (<b>a</b>) mixed eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math>); (<b>b</b>) mixed eccentricity (<math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">s</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">d</mi> </msub> <mo>=</mo> <mn>0.1</mn> </mrow> </semantics></math>) and PM demagnetization (analytical calculations—dashed line ending with the marker “o”; FEM analysis—solid line).</p>
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<p>Simplified cross-section of the machine with rotor eccentricity.</p>
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19 pages, 11485 KiB  
Article
Research on Cogging Torque Reduction of Direct-Drive Type Dual-Rotor Radial Flux Permanent Magnet Synchronous Motor for Electric Propulsion Aircraft
by Haomin Li, Minglu Feng, Li Lei, Yingsan Geng and Jianhua Wang
Energies 2024, 17(7), 1583; https://doi.org/10.3390/en17071583 - 26 Mar 2024
Cited by 1 | Viewed by 994
Abstract
The direct-drive type high-torque-density motor is one of the most promising solutions of electric propulsion for aircraft. The cogging torque of the direct-drive motor causes torque ripple, vibration, and noise, which seriously affect the stability and reliability of the electric propulsion system for [...] Read more.
The direct-drive type high-torque-density motor is one of the most promising solutions of electric propulsion for aircraft. The cogging torque of the direct-drive motor causes torque ripple, vibration, and noise, which seriously affect the stability and reliability of the electric propulsion system for aircraft. In this paper, a novel direct-drive type high-torque-density dual-rotor radial flux permanent magnet synchronous motor (DRPMSM) is proposed, and its cogging torque is weakened by permanent magnet shape design and pole-arc coefficient combination. An eccentric and chamfered permanent magnet shape is proposed, and the influence of eccentric distance and chamfer angle on cogging torque is clarified. Through the pole-arc coefficient combination design of the inner and outer rotors of the DRPMSM, the cogging torques of the inner and outer rotor are phase reversed, thus further reducing the total cogging torque of the DRPMSM. By employing the response surface method, a mathematical model is established for the cogging torque of the DRPMSM in relation to the pole-arc coefficients, chamfer angle, and eccentric distance of the permanent magnets. The parameters that minimize the cogging torque of the DRPMSM are obtained using a genetic algorithm. A prototype is manufactured according to the optimized parameters, and experimental results validate the correctness of the theoretical analysis and the effectiveness of the optimization design method. Full article
(This article belongs to the Section F: Electrical Engineering)
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<p>Structure of the DRPMSM.</p>
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<p>2D finite element model of the DRPMSM.</p>
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<p>Two types of magnetic circuit of the DRPMSM: (<b>a</b>) N-S type; (<b>b</b>) N-N type.</p>
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<p>Flux line distribution of the N-S type DRPMSM (<b>a</b>) with cooling holes and (<b>b</b>) without cooling holes.</p>
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<p>Flux density distribution of rated load condition.</p>
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<p>Permanent magnet shape: (<b>a</b>) inner permanent magnet shape; (<b>b</b>) outer permanent magnet shape.</p>
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<p>The influence of eccentric distance and chamfer angle on the fundamental Fourier coefficient <span class="html-italic">B<sub>r</sub></span><sub>6</sub> of the outer cogging torque: (<b>a</b>) eccentric distance; (<b>b</b>) chamfer angle.</p>
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<p>The influence of eccentric distance and chamfer angle on the amplitude of the outer cogging torque: (<b>a</b>) eccentric distance; (<b>b</b>) chamfer angle.</p>
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<p>The influence of eccentricity and chamfer angle on the outer rotor’s torque performance: (<b>a</b>) cogging torque; (<b>b</b>) output torque; (<b>c</b>) torque ripple.</p>
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<p>Pole-arc combination of inner and outer permanent magnets.</p>
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<p>Pole-arc combinations that make <span class="html-italic">B<sub>r</sub></span><sub>6</sub> equal to zero.</p>
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<p>Comparison of cogging torque before and after pole-arc combination.</p>
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<p>The variation of <span class="html-italic">B<sub>r</sub></span><sub>6_out</sub> with <span class="html-italic">α<sub>N</sub></span><sub>2</sub>.</p>
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<p>The variation of the outer cogging torque with <span class="html-italic">α<sub>N</sub></span><sub>2</sub>: (<b>a</b>) outer cogging torque waveforms; (<b>b</b>) peak value of the outer cogging torque at first half-cycle.</p>
Full article ">Figure 14 Cont.
<p>The variation of the outer cogging torque with <span class="html-italic">α<sub>N</sub></span><sub>2</sub>: (<b>a</b>) outer cogging torque waveforms; (<b>b</b>) peak value of the outer cogging torque at first half-cycle.</p>
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<p>The variation of total cogging torque with <span class="html-italic">α<sub>N</sub></span><sub>2</sub>: (<b>a</b>) total cogging torque waveforms; (<b>b</b>) peak value of the total cogging torque at first half-cycle.</p>
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<p>The variation of outer rotor’s torque performance with the pole-arc combination: (<b>a</b>) torque ripple; (<b>b</b>) average output torque; (<b>c</b>) cogging torque.</p>
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<p>The cogging torque optimization flowchart of the DRPMSM.</p>
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<p>The optimized cogging torque of DRPMSM.</p>
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<p>Prototype of the designed DRPMSM.</p>
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<p>Peak cogging torque value and measured value of the DRPMSM before and after optimization.</p>
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18 pages, 12466 KiB  
Article
Electromagnetic Vibration Analysis of Transverse Flux Permanent Magnet Linear Submersible Motor for Oil Production
by Mei Zhao, Yihao Li, Sicheng Zuo, Pingpeng Tang, Tong Yao, Huaqiang Zhang and Shunjie Wu
Energies 2023, 16(23), 7911; https://doi.org/10.3390/en16237911 - 4 Dec 2023
Viewed by 1046
Abstract
A transverse flux linear motor is a special type of linear motor with a high thrust force density, and it has broad application prospects in the field of linear direct-drive systems. In the process of oil production, the vibration of the linear motor [...] Read more.
A transverse flux linear motor is a special type of linear motor with a high thrust force density, and it has broad application prospects in the field of linear direct-drive systems. In the process of oil production, the vibration of the linear motor poses a significant amount of harm to the system due to its special slender structure. This paper focuses on the electromagnetic vibration of a transverse flux permanent magnet linear submersible motor (TFPMLSM). Firstly, the no-load air gap flux density is calculated based on the field modulation principle. Secondly, the radial electromagnetic force (REF) of the TFPMLSM is calculated, and the finite element method (FEM) is used to analyze the time-space and spectral characteristics of the REF. Then, the influence of secondary eccentricity on the frequency spectrum of the REF is further concluded. Finally, the natural frequencies of each vibration mode are calculated using the modal superposition method and the influence of the REF on the motor vibration is obtained through magnetic-structural coupling analysis. The research results found that the motor does not cause resonance at low speeds, and the fundamental frequency of REF has the greatest impact on electromagnetic vibration. Full article
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Figure 1
<p>Primary structure of the transverse flux permanent magnet linear submersible motor (TFPMLSM): (<b>a</b>) primary unit; (<b>b</b>) lamination A of primary iron core; (<b>c</b>) lamination B.</p>
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<p>Secondary structure of TFPMLSM: (<b>a</b>) mover; (<b>b</b>) lamination C of secondary iron core; (<b>c</b>) lamination D.</p>
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<p>The arrangement of the mover iron poles.</p>
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<p>3-D finite element model of the TFPMLSM.</p>
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<p>Size parameters of motor. (<b>a</b>) Front view of motor. (<b>b</b>) Side view of motor.</p>
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<p>Axial flux path of TFPMLSM.</p>
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<p>Flux path of one sixth single-phase TFPMLSM model. (<b>a</b>) Maximum flux path in a positive direction. (<b>b</b>) Maximum flux path in a negative direction.</p>
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<p>MMF distribution of PMs and permeance distribution of iron poles. (<b>a</b>) Modal of stator PMs and mover iron poles; (<b>b</b>) modal of mover PMs and stator iron poles; (<b>c</b>) permeance of mover iron poles; (<b>d</b>) permeance of stator iron poles; (<b>e</b>) MMF of stator PMs; (<b>f</b>) MMF of mover PMs.</p>
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<p>Electromagnetic force in three directions.</p>
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<p>Air gap flux density distribution under tooth II.</p>
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<p>Axial segmentation of the TFPMLSM.</p>
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<p>REF distribution and its FFT under no-load state. (<b>a</b>) The spatiotemporal distribution of REF (section I); (<b>b</b>) harmonic distribution (section I); (<b>c</b>) the spatiotemporal distribution of REF (section II); (<b>d</b>) harmonic distribution (section II). Note that different colors in (<b>b</b>) and (<b>d</b>) represent different magnitudes of REF.</p>
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<p>REF distribution and its FFT under load state. (<b>a</b>) The spatiotemporal distribution of REF; (<b>b</b>) harmonic distribution. Note that different colors in (<b>b</b>) represent different magnitudes of REF.</p>
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<p>Static eccentricity of mover.</p>
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<p>Spatial distribution and harmonics of REF versus eccentricity. (<b>a</b>) Spatial distribution of REF; (<b>b</b>) harmonic distribution.</p>
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<p>Time distribution and harmonic of axial electromagnetic force versus eccentricity. (<b>a</b>) Time distribution of REF; (<b>b</b>) harmonic distribution.</p>
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<p>Time distribution of REF and its DC component versus eccentricity. (<b>a</b>) Time distribution of REF; (<b>b</b>) influence of eccentricity on the DC component of REF.</p>
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<p>Time distribution of AEF and its DC component versus eccentricity. (<b>a</b>) Time distribution of AEF; (<b>b</b>) harmonic distribution.</p>
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<p>Stator structure of modal analysis.</p>
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<p>The first 20 modes and their natural frequencies under unconstraint and fixed constraint.</p>
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<p>Location of vibration measurement points.</p>
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<p>Electromagnetic vibration response of stator tooth surface and outer surface. (<b>a</b>) Stator tooth surface vibration response; (<b>b</b>) vibration response of stator outer surface.</p>
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<p>Electromagnetic vibration response of stator tooth surface and outer surface with 40% mover eccentricity. (<b>a</b>) Stator tooth surface vibration response; (<b>b</b>) vibration response of stator outer surface.</p>
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17 pages, 4369 KiB  
Article
Suspension Flux Internal Model Control of Single-Winding Bearingless Flux-Switching Permanent Magnet Motor
by Yao Chen, Wanneng Yu, Rongfeng Yang and Bowen Cui
Actuators 2023, 12(11), 404; https://doi.org/10.3390/act12110404 - 28 Oct 2023
Viewed by 1510
Abstract
A suspension flux internal model control method is proposed to address the problem of the strong coupling of a single-winding bearingless flux-switching permanent magnet motor leading to a significant ripple of the rotor radial displacement. Firstly, based on air-gap magnetic field modulation theory, [...] Read more.
A suspension flux internal model control method is proposed to address the problem of the strong coupling of a single-winding bearingless flux-switching permanent magnet motor leading to a significant ripple of the rotor radial displacement. Firstly, based on air-gap magnetic field modulation theory, the stator flux equation considering rotor dynamic eccentricity is established to reveal the relationship between the eccentric rotor and the magnetic field. Secondly, according to the dynamic characteristics of the motor and the variation law of the air-gap magnetic field, the suspension-plane flux is substituted into the rotor dynamic model, and the suspension flux-dynamics internal model and corresponding output are constructed, respectively. Finally, a complete control strategy is established, and the rotor is stably suspended by PWM control. The simulation and experimental results show that the proposed method has better steady-state and dynamic performance than traditional PID control, and the maximum radial displacement ripples of the rotor are reduced by 53% and 50% in steady-state and dynamic operation. Full article
(This article belongs to the Special Issue Vibration Control Using Electromagnetic Actuators)
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<p>Topology of BFSPMM.</p>
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<p>Rotor-motion coordinate system of BFSPMM.</p>
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<p>Cascade model of air-gap magnetic field modulation theory.</p>
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<p>Modulation function diagram of BFSPMM.</p>
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<p>Control strategy.</p>
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<p>Experimental tests: (<b>a</b>) BFSPMM unit; (<b>b</b>) Controller and driver.</p>
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<p>Simulation results: (<b>a</b>) Rotor displacement under PID control; (<b>b</b>) Suspension current under PID control; (<b>c</b>) Rotor displacement under proposed control; (<b>d</b>) Suspension current under proposed control.</p>
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<p>Simulation results: (<b>a</b>) Rotor displacement under PID control; (<b>b</b>) Suspension current under PID control; (<b>c</b>) Rotor displacement under proposed control; (<b>d</b>) Suspension current under proposed control.</p>
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<p>Comparison of the steady-state experiment: (<b>a</b>) Traditional PID method; (<b>b</b>) Proposed method.</p>
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<p>Comparison of the dynamic-state experiment: (<b>a</b>) Rotor displacement under PID control; (<b>b</b>) Suspension current under PID control; (<b>c</b>) Rotor displacement under proposed control; (<b>d</b>) Suspension current under proposed control.</p>
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<p>The connection between the proposed method and LADRC.</p>
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17 pages, 23093 KiB  
Article
Analysis of Direct Torque Control Response to Stator and Rotor Faults in Permanent Magnet Synchronous Machines
by Ibrahim M. Allafi and Shanelle N. Foster
Energies 2023, 16(19), 6940; https://doi.org/10.3390/en16196940 - 3 Oct 2023
Viewed by 1559
Abstract
Direct-torque-control-driven permanent magnet synchronous machines eliminate the need for a position sensor while providing improved torque dynamics. However, the structure, regulation principle and nature of compensation of hysteresis-based controllers used in direct torque control impacts performance under faulty operating conditions. This paper analyzes [...] Read more.
Direct-torque-control-driven permanent magnet synchronous machines eliminate the need for a position sensor while providing improved torque dynamics. However, the structure, regulation principle and nature of compensation of hysteresis-based controllers used in direct torque control impacts performance under faulty operating conditions. This paper analyzes the reaction of direct torque control to the presence of various faults that occur in permanent magnet synchronous machines. The analysis presented reveals that the direct torque control injects a negative sequence voltage and manipulates the torque angle to meet the control objectives when a fault occurs. The co-simulation of finite element analysis and a multi-physic circuit simulator is used to validate the response of the hysteresis-based controller to the machine health. The results indicate that the hysteresis comparators have the ability to mask the impact of the faults in the direct-torque-control-driven permanent magnet synchronous machines. Full article
(This article belongs to the Special Issue Condition Monitoring and Failure Prevention of Electric Machines)
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<p>Electromagnetic torque and stator flux linkage control loops in DTC-driven PMSM.</p>
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<p>Torque and stator flux linkage variations due to hysteresis comparators. (<b>a</b>) Actual torque variation within torque hysteresis bands. (<b>b</b>) Actual flux variation within flux hysteresis bands.</p>
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<p>Frequency spectrum of the IPMSM torque magnitude under both FOC (<b>left</b>) and DTC (<b>right</b>) drives. The 6th and 12th harmonics are not observed in the DTC drive.</p>
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<p>Frequency spectrum of the torque angle in DTC-driven PMSM.</p>
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<p>Sector partitions and the available set of voltage vectors in the DTC drive.</p>
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<p>Illustration of the error between the actual and estimated stator flux linkages after the occurrence of a fault.</p>
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<p>Path of the estimated stator flux linkage based on the voltage vector selection table shown in <a href="#energies-16-06940-t001" class="html-table">Table 1</a>.</p>
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<p>Optimal path of the actual stator flux linkage based on the voltage vector selection table shown in <a href="#energies-16-06940-t001" class="html-table">Table 1</a>.</p>
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<p>Path of the actual stator flux linkage based on the voltage vector selection table shown in <a href="#energies-16-06940-t001" class="html-table">Table 1</a>.</p>
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<p>Machine model in MAXWELL coupled with DTC circuit in SIMPLORER including the TTSC fault circuit on phase A winding.</p>
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<p>Sequence components of the commanded voltages in DTC-driven PMSM under healthy operating conditions.</p>
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<p>Sequence components of the commanded voltages in DTC-driven PMSM under TTSC fault with different severity levels. The healthy sequence components, from <a href="#energies-16-06940-f011" class="html-fig">Figure 11</a>, are shown in the faded color. (<b>a</b>) Case 1: (15, 0.5 <math display="inline"><semantics> <mo>Ω</mo> </semantics></math>). (<b>b</b>) Case 2: (30, 0.25 <math display="inline"><semantics> <mo>Ω</mo> </semantics></math>).</p>
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<p>Sequence components of the commanded voltages in DTC-driven PMSM under HRC fault with different severity levels. The healthy sequence components, from <a href="#energies-16-06940-f011" class="html-fig">Figure 11</a>, are shown in the faded color. (<b>a</b>) Case 1: 100%. (<b>b</b>) Case 2: 150%.</p>
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<p>Sequence components of the commanded voltages in DTC-driven PMSM under demagnetization fault with different severity levels. The healthy sequence components, from <a href="#energies-16-06940-f011" class="html-fig">Figure 11</a>, are shown in the faded color. (<b>a</b>) Case 1: 1 Magnet. (<b>b</b>) Case 2: 3 Magnets.</p>
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<p>Sequence components of the commanded voltages in DTC-driven PMSM under eccentricity fault with different severity levels. The healthy sequence components, from <a href="#energies-16-06940-f011" class="html-fig">Figure 11</a>, are shown in the faded color. (<b>a</b>) Case 1: 40%. (<b>b</b>) Case 2: 60%.</p>
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<p>The magnitude of the magnet and reluctance torque multipliers, cos (m) and cos (d), in (<a href="#FD18-energies-16-06940" class="html-disp-formula">18</a>) and (<a href="#FD19-energies-16-06940" class="html-disp-formula">19</a>).</p>
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<p>The magnitude of the offset error in (<a href="#FD18-energies-16-06940" class="html-disp-formula">18</a>) and (<a href="#FD19-energies-16-06940" class="html-disp-formula">19</a>).</p>
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20 pages, 12070 KiB  
Article
Performance Optimization of Ultralow-Frequency Electromagnetic Energy Harvester Driven by Eccentric mass
by Jintao Liang, Chao Zhang and Kangqi Fan
Machines 2023, 11(7), 743; https://doi.org/10.3390/machines11070743 - 15 Jul 2023
Cited by 2 | Viewed by 1426
Abstract
Driven by an eccentric mass through a two-layered cantilevered plectrum, the electromagnetic energy harvester (EEH) can convert low-frequency mechanical vibrations into continuous uni-directional rotation. To optimize the performance of the EEH, electromagnetic analysis of the EEH was conducted. Three-phase winding permanent magnet (PM) [...] Read more.
Driven by an eccentric mass through a two-layered cantilevered plectrum, the electromagnetic energy harvester (EEH) can convert low-frequency mechanical vibrations into continuous uni-directional rotation. To optimize the performance of the EEH, electromagnetic analysis of the EEH was conducted. Three-phase winding permanent magnet (PM) topology was employed, and combinations of different coils and magnet pole numbers were designed. Then, the finite element method (FEM) was applied to analyze the influence of the combinations of the coils and pole numbers as well as the PM dimensions on the three-phase induced voltage. Prototypes with different configurations were fabricated and the analysis effectiveness was confirmed. Furthermore, different types of stator yokes were designed to enhance the magnetic field. Compared to the original prototype, the output voltage of the optimal prototype increased by 0.5 V with the same rotation speed, and the harmonic components were sufficiently low. Then, experiments with excitation by linear reciprocating motions and swing motions were conducted. Under different exciting conditions, the optimal prototype can also induce the highest voltage amplitude. With an increase in the weight of the eccentric mass, a long duration can be reached that lasts up to 12 s. In summary, the proposed optimization can achieve a high-efficiency and high-power density EEH. Full article
(This article belongs to the Special Issue Mechatronic Systems: Developments and Applications)
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<p>Structure of the proposed EEH: (<b>a</b>) cross-section view; (<b>b</b>) Exploded view.</p>
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<p>Principle of the proposed EEH.</p>
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<p>Coil and phase vectors with different configurations: (<b>a</b>) 12C/10P; (<b>b</b>) 12C/14P; (<b>c</b>) 18C/16P; (<b>d</b>) 18C/20P.</p>
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<p>Initial prototype of the proposed EEH.</p>
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<p>Electromagnetic FEM model of the EEH: (<b>a</b>) geometric model; (<b>b</b>) mesh generation.</p>
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<p>Three-phase voltage waveforms with initial model at 600 rpm: (<b>a</b>) FEM result; (<b>b</b>) Experimental result: (<b>c</b>) Result comparison in one period of Phase A.</p>
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<p>Experiment device for voltage measurement.</p>
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<p>FEM results of initial 12C-10P model with different <span class="html-italic">b<sub>pm</sub></span>: (<b>a</b>) Voltage waveforms; (<b>b</b>) Harmonic magnitude of voltage.</p>
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<p>FEM results of initial 12C-10P model with different <span class="html-italic">b<sub>pm</sub></span>: (<b>a</b>) Voltage waveforms; (<b>b</b>) Harmonic magnitude of voltage.</p>
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<p>Non-teeth stator core structure.</p>
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<p>Voltage amplitude of 12C-10P FEM model with different material.</p>
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<p>FEM results of steel-yoke 12C-14P model with different <span class="html-italic">b<sub>pm</sub></span>: (<b>a</b>) Voltage waveforms; (<b>b</b>) Harmonic magnitude of voltage.</p>
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<p>Comparison of voltage performance with different configurations: (<b>a</b>) Voltage amplitude; (<b>b</b>) Voltage amplitude versus PM mass.</p>
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<p>Prototypes with different configurations: (<b>a</b>) 12C-14P with <span class="html-italic">b<sub>pm</sub></span> = 6 mm; (<b>b</b>) 18C-20P with <span class="html-italic">b<sub>pm</sub></span> = 4 mm.</p>
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<p>Experiment results of three-phase voltage waveforms at 600 rpm: (<b>a</b>) 12C-10P with <span class="html-italic">b<sub>pm</sub></span> = 5 mm (initial); (<b>b</b>) 12C-14P with <span class="html-italic">b<sub>pm</sub></span> = 6 mm (optimal); (<b>c</b>) 18C-20P with <span class="html-italic">b<sub>pm</sub></span> = 4 mm; (<b>d</b>) Result comparison in one period of Phase A.</p>
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<p>Experiment results of three-phase voltage waveforms at 600 rpm: (<b>a</b>) 12C-10P with <span class="html-italic">b<sub>pm</sub></span> = 5 mm (initial); (<b>b</b>) 12C-14P with <span class="html-italic">b<sub>pm</sub></span> = 6 mm (optimal); (<b>c</b>) 18C-20P with <span class="html-italic">b<sub>pm</sub></span> = 4 mm; (<b>d</b>) Result comparison in one period of Phase A.</p>
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<p>Experiment device for swing excitation.</p>
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<p>Output voltage of the three prototypes under swings with 2.5 Hz, 60°: (<b>a</b>) 12C-10P with <span class="html-italic">b<sub>pm</sub></span> = 5 mm (initial); (<b>b</b>) 12C-14P with <span class="html-italic">b<sub>pm</sub></span> = 6 mm (optimal); (<b>c</b>) 18C-20P with <span class="html-italic">b<sub>pm</sub></span> = 4 mm.</p>
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<p>Output voltage of the three prototypes under swings with 2.5 Hz, 60°: (<b>a</b>) 12C-10P with <span class="html-italic">b<sub>pm</sub></span> = 5 mm (initial); (<b>b</b>) 12C-14P with <span class="html-italic">b<sub>pm</sub></span> = 6 mm (optimal); (<b>c</b>) 18C-20P with <span class="html-italic">b<sub>pm</sub></span> = 4 mm.</p>
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<p>Frequency response of output voltage under different swing angles: (<b>a</b>) Swing 40°; (<b>b</b>) Swing 60°.</p>
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<p>Frequency response of output voltage under different swing angles: (<b>a</b>) Swing 40°; (<b>b</b>) Swing 60°.</p>
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<p>Output voltage of the 12C-14P prototype with two eccentric mass structures in ultra-low swing frequency (1Hz): (<b>a</b>) With initial structure under swing; (<b>b</b>) With whole steel block structure under swing; (<b>c</b>) With whole steel block structure after swing remove.</p>
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<p>Output voltage of the 12C-14P prototype with two eccentric mass structures in ultra-low swing frequency (1Hz): (<b>a</b>) With initial structure under swing; (<b>b</b>) With whole steel block structure under swing; (<b>c</b>) With whole steel block structure after swing remove.</p>
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<p>Experiment device for linear reciprocation.</p>
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<p>Frequency response of output voltage under 50 mm linear motion.</p>
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<p>Output voltage of the original and optimal prototypes under linear reciprocation with 2.5 Hz (<b>a</b>) With original prototype; (<b>b</b>) With optimal prototype.</p>
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<p>Output voltage of the original and optimal prototypes under linear reciprocation with 2.5 Hz (<b>a</b>) With original prototype; (<b>b</b>) With optimal prototype.</p>
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14 pages, 6895 KiB  
Article
Characteristic Analysis of a New Structure Eccentric Harmonic Magnetic Gear
by Libing Jing, Youzhong Wang, Dawei Li and Ronghai Qu
Actuators 2023, 12(6), 248; https://doi.org/10.3390/act12060248 - 14 Jun 2023
Viewed by 1539
Abstract
An eccentric harmonic magnetic gear (EHMG) is better suited for situations requiring larger transmission ratios than magnetic-field-modulated magnetic gears. In the meantime, to increase the torque density even further, a new structure for EHMGs is presented in this paper. The stator’s permanent magnets [...] Read more.
An eccentric harmonic magnetic gear (EHMG) is better suited for situations requiring larger transmission ratios than magnetic-field-modulated magnetic gears. In the meantime, to increase the torque density even further, a new structure for EHMGs is presented in this paper. The stator’s permanent magnets (PMs) are irregularly distributed, while the rotor’s PMs are applied to a fan-shaped structure. Moreover, a Halbach array is adopted in both the rotor and the stator. A two-dimensional finite element (FE) model of the proposed EHMG is developed, and the flux density distribution and torque of the EHMG are calculated and verified via FE analysis. When compared to a conventional EHMG, the presented model’s torque increases from 38.04 Nm to 50.41 Nm. In addition, for the sake of avoiding the oscillation and noise caused by resonance, a modal analysis of the proposed model is conducted and the consequences show that it has better antivibration properties. Finally, a prototype is made, a test bench is established, and the correctness and effectiveness of the proposed model are verified. Full article
(This article belongs to the Special Issue Power Electronics and Actuators)
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<p>Structure of EHMG. (<b>a</b>) Conventional model; (<b>b</b>) proposed model.</p>
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<p>Relative placement of stator and rotor at different angles. (<b>a</b>) 0°; (<b>b</b>) 135°; (<b>c</b>) 225°; (<b>d</b>) 360°.</p>
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<p>The unequal blocks of PMs.</p>
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<p>The fan-shaped structure of PMs.</p>
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<p>Magnetic flux distribution. (<b>a</b>) Conventional; (<b>b</b>) proposed.</p>
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<p>Flux density. (<b>a</b>) Radial flux density; (<b>b</b>) tangential flux density.</p>
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<p>Harmonic pole pair distribution spectrum of flux density. (<b>a</b>) Radial component; (<b>b</b>) tangential component.</p>
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<p>Static torque.</p>
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<p>Steady-state torque.</p>
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<p>Space distribution of the radial electromagnetic force density.</p>
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<p>Modal shapes of rotor. (<b>a</b>) Second order; (<b>b</b>) third order; (<b>c</b>) fourth order; (<b>d</b>) fifth order; (<b>e</b>) sixth order.</p>
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<p>Prototype parts and test bench. (<b>a</b>) High-speed rotor; (<b>b</b>) low-speed rotor; (<b>c</b>) stator; (<b>d</b>) test bench.</p>
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<p>Output torque.</p>
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<p>Measured efficiency.</p>
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19 pages, 8229 KiB  
Article
The Neutral Voltage Difference Signal as a Means of Investigating Eccentricity and Demagnetization Faults in an AFPM Synchronous Generator
by Alexandra C. Barmpatza
Machines 2023, 11(6), 647; https://doi.org/10.3390/machines11060647 - 14 Jun 2023
Cited by 4 | Viewed by 1461
Abstract
This article investigates the neutral voltage difference signal, VNO signal, for fault diagnosis. The aforementioned signal is the signal of the voltage between the common star point of the stator and the common star point of the load. The under-study faults are [...] Read more.
This article investigates the neutral voltage difference signal, VNO signal, for fault diagnosis. The aforementioned signal is the signal of the voltage between the common star point of the stator and the common star point of the load. The under-study faults are demagnetization and static eccentricity faults, while the machine in which the faults are investigated is an axial flux permanent magnet (AFPM) synchronous generator, suitable for wind power applications. This study was conducted using a 3D finite element method (3D-FEM), and the machine’s FEM model was validated through experiments. This method is one of the most accurate methods for electrical machine computation, allowing for a detailed study of electromagnetic behavior. The components that constitute the VNO signal were determined using a 3D-FEM software program (Opera 18R2). Subsequently, further analysis was performed using MATLAB R2022b software, and a fast Fourier transform (FFT) was applied to this signal. In all the investigated faulty cases, new harmonics appeared, and the healthy amplitudes of most of the already existing harmonics increased. These findings can be used for fault identification. The analysis revealed that the harmonic frequency of 1.5fs was the most dominant in the case of demagnetization, while in the case of static eccentricity, the most dominant harmonic was a frequency equal to the machine’s operating frequency, fs. The novelty of this study is that this signal has not previously been used for fault identification, especially in AFPM synchronous machines. This signal depends on EMF voltage and stator phase currents but is less sinusoidal. Consequently, it can detect faults in cases where the aforementioned signals cannot be used for detection. Full article
(This article belongs to the Section Electrical Machines and Drives)
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<p>The FEM model of the machine in which the faults are studied.</p>
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<p>The experimental setup in the laboratory.</p>
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<p>The stator current of phase A in the healthy generator when the speed was 600 rpm and the machine fed a 30 Ohm resistive load: blue line—waveform generated using the FEM model; red line—waveform generated during the experimental procedure.</p>
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<p>The electrical circuit of the stator of the machine.</p>
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<p>The V<sub>NO</sub> signal waveform for the healthy and faulty cases with (<b>a</b>) 20% partial demagnetization and (<b>b</b>) 50% partial demagnetization.</p>
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<p>The V<sub>NO</sub> signal waveform for the healthy and faulty cases with (<b>a</b>) 20% partial demagnetization and (<b>b</b>) 50% partial demagnetization.</p>
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<p>The V<sub>NO</sub> signal spectra for the healthy and faulty cases with (<b>a</b>) 20% partial demagnetization and (<b>b</b>) 50% partial demagnetization.</p>
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<p>The V<sub>NO</sub> signal spectra for the healthy and faulty cases with (<b>a</b>) 20% partial demagnetization and (<b>b</b>) 50% partial demagnetization.</p>
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<p>The stator current spectra for the healthy case and the case where one magnet was demagnetized generated during the experimental procedure.</p>
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<p>Schematic diagram of static eccentricity: (<b>a</b>) angular; (<b>b</b>) axial.</p>
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<p>The V<sub>NO</sub> signal waveform for the healthy and faulty cases with (<b>a</b>) 30% static angular eccentricity and (<b>b</b>) 40% static angular eccentricity.</p>
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<p>The V<sub>NO</sub> signal spectra for the healthy and faulty cases with (<b>a</b>) 30% static angular eccentricity and (<b>b</b>) 40% static angular eccentricity.</p>
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<p>The V<sub>NO</sub> signal waveform for the healthy and faulty cases with (<b>a</b>) 1 mm static axial eccentricity and (<b>b</b>) 2 mm static axial eccentricity.</p>
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<p>The V<sub>NO</sub> signal waveform for the healthy and faulty cases with (<b>a</b>) 1 mm static axial eccentricity and (<b>b</b>) 2 mm static axial eccentricity.</p>
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<p>The V<sub>NO</sub> signal spectra for the healthy and faulty cases with (<b>a</b>) 1 mm static axial eccentricity and (<b>b</b>) 2 mm static axial eccentricity.</p>
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<p>The V<sub>NO</sub> signal spectra for the healthy and faulty cases with (<b>a</b>) 1 mm static axial eccentricity and (<b>b</b>) 2 mm static axial eccentricity.</p>
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19 pages, 13312 KiB  
Article
Optimal Design of Marine Motors for Joint Efficiency and Economic Optimization
by Yingshui Sun, Yuxuan Fang, Qiao Zhang and Qing Liu
Energies 2023, 16(12), 4588; https://doi.org/10.3390/en16124588 - 8 Jun 2023
Cited by 3 | Viewed by 1199
Abstract
The permanent magnet synchronous motor (PMSM) has been widely used in the field of ship electric propulsion due to its advantages of a small size, light weight, low loss, and high efficiency. In this paper, a 100 kW ship-side thruster motor was taken [...] Read more.
The permanent magnet synchronous motor (PMSM) has been widely used in the field of ship electric propulsion due to its advantages of a small size, light weight, low loss, and high efficiency. In this paper, a 100 kW ship-side thruster motor was taken as the research object, and the problem of the high harmonic content of the air gap magnetic flux density in the motor was addressed by designing a rotor eccentricity. On this basis, the hybrid Taguchi method of genetic algorithm was used to optimize the rotor structural parameters with increasing the efficiency and reducting the cost of the motor as the optimization objectives. The results show that the performance and economy of the motor have been greatly improved after optimization. Finally, the motor weight reduction hole was designed, and a prototype was manufactured and tested. The test data are within the allowable range compared with the simulation data, verifying the effectiveness of the multiobjective optimization algorithm proposed in this paper. Full article
(This article belongs to the Section F: Electrical Engineering)
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<p>Finite Element Model of Electric Motor.</p>
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<p>Magnetization Curve of 50WW470.</p>
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<p>Electric Motor Flux Map.</p>
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<p>Surface Shape Optimization of Rotor Before and After.</p>
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<p>Comparison of Air Gap Flux Density before and after Optimization.</p>
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<p>Comparison of Back EMF before and after Optimization.</p>
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<p>Genetic Algorithm Flowchart.</p>
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<p>Rotor structural parameters.</p>
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<p>Taguchi Optimization Process Diagram.</p>
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<p>Torque waveform of motor before and after optimization.</p>
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<p>Comparison of magnetic flux density distribution before and after weight reduction design.</p>
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<p>Stress distribution in rotor core.</p>
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<p>Transient temperature field simulation of motor.</p>
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<p>Physical image of prototype.</p>
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<p>Motor testing platform.</p>
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<p>Motor torque-efficiency curve.</p>
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<p>Back electromotive force waveform.</p>
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<p>The temperature variation curve of the test motor winding.</p>
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19 pages, 2749 KiB  
Article
Analytical Calculation of Air Gap Magnetic Field of SPMSM with Eccentrically Cut Poles Based on Magnetic Pole Division
by Jiahe Zhang, Jiapei Hu, Guobiao Gu and Fangmian Du
Energies 2023, 16(11), 4450; https://doi.org/10.3390/en16114450 - 31 May 2023
Viewed by 1413
Abstract
In the design process of surface-mounted permanent magnet motor (SPMSM) for industrial robots and computer numerical control (CNC) machine tools, to pursue the sinusoidal nature of the back electromotive force, the magnetic poles in the form of eccentric pole cutting structure are often [...] Read more.
In the design process of surface-mounted permanent magnet motor (SPMSM) for industrial robots and computer numerical control (CNC) machine tools, to pursue the sinusoidal nature of the back electromotive force, the magnetic poles in the form of eccentric pole cutting structure are often used. To analyze the no-load air gap magnetic field of the SPMSM with eccentrically cut poles simply and accurately, a subdomain model magnetic field analytical calculation method based on equal-area integral block processing of permanent magnets is proposed. The problem that the traditional subdomain analysis model cannot be directly applied to the SPMSM with eccentrically cut poles of unequal thickness is solved. The proposed method considers the influence of stator slotting and the actual permeability of permanent magnets, and directly obtains the fundamental wave and harmonic components of the no-load air gap flux density by solving the subdomain model. The finite element method (FEM) is used to directly calculate the air gap magnetic field for verification. The results of the analytical method and the no-load air gap magnetic density calculated by the FEM are consistent, which verifies the accuracy of the proposed analytical method and can quickly guide the design of the SPMSM with eccentrically cut poles. Full article
(This article belongs to the Section F: Electrical Engineering)
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<p>(<b>a</b>) Schematic diagram of the block method of eccentrically cut poles. (<b>b</b>) Cross-sectional view of SPMSM with eccentrically cut poles.</p>
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<p>(<b>a</b>) Schematic diagram of the block method of eccentrically cut poles. (<b>b</b>) Cross-sectional view of SPMSM with eccentrically cut poles.</p>
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<p>Flow chart of the subdomain analysis method.</p>
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<p>(<b>a</b>) Diagram of the magnetic density produced by each small pole block and the radial component of the resultant magnetic density. (<b>b</b>) Diagram of the magnetic density produced by each small pole block and the tangential component of the resultant magnetic density.</p>
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<p>Magnetic flux density diagram of SPMSM with eccentrically cut poles.</p>
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<p>(<b>a</b>) The comparison between the calculation results of the no-load air gap flux density radial component analytical method and the FEM calculation results. (<b>b</b>) The comparison between the calculation results of the no-load air gap flux density tangential component analytical method and the FEM calculation results.</p>
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<p>(<b>a</b>) The comparison between the calculation results of the no-load air gap flux density radial component analytical method and the FEM calculation results. (<b>b</b>) The comparison between the calculation results of the no-load air gap flux density tangential component analytical method and the FEM calculation results.</p>
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<p>Induced electromotive force.</p>
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