Optimization Design of Novel Consequent Pole Motor for Electric Power Steering System

Abstract

This paper proposes a novel consequent pole machine (CPM) to improve cogging torque, torque ripple, and torque per magnet volume compared to a conventional CPM. The proposed structure consists of a consequent pole rotor that reduces permanent magnet (PM) material. Different from conventional surface permanent magnet machines (SPMs) and conventional CPMs, the proposed CPM has an additional iron pole (AIP) extension under the magnet. This structure is suggested in order to increase the magnet pole angle while maintaining the magnet volume by partially substituting the magnet with iron. The design of the proposed machine was optimized to achieve higher torque and lower torque ripple. Comparisons of the proposed CPM with a conventional SPM and CPM were conducted to show the superiority of the proposed CPM utilizing AIP. The results showed that the proposed CPM exhibits 68% lower cogging torque, 75% lower torque ripple, and 16% higher torque per magnet volume with the same magnet volume compared with the conventional CPM.

1. Introduction

Permanent magnet synchronous motors (PMSMs) play a crucial role in various industrial applications due to their high efficiency and power density, particularly in electric vehicles and industrial machinery. In high-performance applications like electric vehicles, PMSMs are essential because they provide excellent torque density and efficiency. However, with the rising cost and limited availability of rare-earth permanent magnets (PMs), there has been an increasing demand for motor designs that maintain performance while reducing PM usage. Researchers have proposed various solutions to address this challenge, focusing on reducing reliance on rare-earth magnets while optimizing motor performance.

One promising solution is the consequent pole motor (CPM), which alternates between PM and iron poles. This design offers the potential to significantly reduce the amount of magnet material used. However, existing CPM designs often suffer from increased cogging torque, torque ripple, and reduced torque per magnet volume, leading to overall performance degradation compared to traditional PMSMs [1,2,3]. These issues highlight the need for design optimizations to improve the performance of CPMs while reducing PM usage.

Efforts to reduce PM usage have focused primarily on optimizing rotor designs to enhance flux utilization. Flux-focusing structures have been widely adopted to concentrate magnetic flux between magnets and iron poles, thus improving torque density without increasing magnet volume [4,5,6]. For instance, modular and spoke-type rotor designs have been shown to significantly boost air-gap flux density, enhancing performance without increasing magnet size and reducing reliance on rare-earth materials. Hybrid rotor designs, which combine consequent poles with spoke-type structures, have further improved flux-weakening capabilities and torque density, making them suitable for high-performance applications [7,8]. However, these methods often involve complex geometries that are challenging to manufacture and may not adequately address torque ripple issues.

Challenges related to torque ripple and cogging torque remain unresolved in CPM designs. To address these issues, researchers have introduced asymmetric magnetic pole designs, which smooth the air-gap flux density and reduce electromagnetic imbalances, thereby minimizing torque ripple [9,10,11]. Moreover, techniques such as multi-arc poles and segmented pole designs have been employed to further reduce torque ripple and improve motor performance, particularly at low speeds [12,13]. These optimizations also address unipolar leakage flux, a common issue in traditional CPMs, resulting in improved motor efficiency and reliability [14,15]. However, the potential for increased manufacturing complexity and cost in these designs cannot be overlooked.

Maximizing torque output per unit of magnet material is another critical area of research. Studies have shown that PM arc optimization can significantly improve torque density in CPMs while minimizing magnet material usage [16,17]. Furthermore, designs that incorporate flux-assisted magnets and tangential magnetization have demonstrated improvements in torque performance and efficiency without increasing magnet volume [18,19]. These approaches align with the broader goal of reducing rare-earth magnet dependency while optimizing magnet utilization and maintaining high motor performance [20,21].

Building on these advancements, this paper proposes a novel CPM design that incorporates additional iron pole (AIP) extensions beneath the magnets. This design not only increases the magnet pole angle without increasing the magnet volume but also optimizes the magnetic flux path, thereby improving torque performance. By partially replacing the magnets with iron, the proposed design enhances the magnetic circuit, leading to higher torque output while maintaining the same magnet volume. Unlike previous designs that require complex manufacturing techniques [7,12], the AIP design offers an effective and practical solution with straightforward implementation.

In this paper, a six-pole, nine-slot, novel CPM was proposed to address the issues of high torque ripple and cogging torque in conventional designs. The proposed CPM features an additional rotor iron pole extension beneath the magnet pole and a larger magnet pole angle, which maintains the magnet volume of the conventional CPM while significantly enhancing torque performance. Moreover, the design of the proposed model was optimized using the Kriging model, ensuring a balance between torque stability and material efficiency. To assess the performance of the proposed model, comprehensive comparisons with a conventional SPM and a conventional CPM were conducted using two-dimensional (2-D) finite element analysis (FEA).

The main contributions of this paper are as follows:

(1)

Introduction of a novel CPM design with an additional iron pole (AIP) extension that optimizes the magnetic flux path and enhances torque performance while maintaining the same magnet volume.

(2)

Significant reduction in cogging torque (68%) and torque ripple (75%) compared to conventional CPMs, achieving improved torque stability and smooth motor operation.

(3)

Increased torque per magnet volume by 16% compared to conventional CPMs, highlighting more efficient use of magnetic material.

(4)

Presentation of a design approach that reduces reliance on rare-earth permanent magnets, contributing to cost-effective and sustainable motor development for industrial and automotive applications.

The remainder of this article is organized as follows. Section 2 provides an overview of the design parameters and configuration of the conventional SPM and CPM models. Section 3 details the development and optimization process of the proposed CPM, including key design variables and constraints. Section 4 presents the simulation methodology and performance analysis used to evaluate the models. Finally, Section 5 discusses the comparative results, highlighting the improvements achieved by the proposed and also concludes this study, summarizing the main findings and suggesting directions for future work.

2. SPM Design for Electric Power Steering Motor in Tilt and Telescopic Steering Columns

The design parameters of the surface permanent magnet machine (SPM) for the electric power steering (EPS) system, specifically for the electric tilt and telescopic steering column, are listed in

Table 1. Main parameters and specifications of the SPM.

Item	Unit	Value
		SPM
Air gap	mm	0.5
Stack length	mm	30
Rotor diameter	mm	21.9
Stator diameter	mm	38.4
Magnet thickness	mm	4
Number of turns	-	31
Slot fill factor	%	30
Rated current/phase	Arms	1.4
Rated power	W	12
Rated speed	RPM	3500
PM material	T	Br = 0.4 (Ferrite)
Core material	-	50H1300 (by NSSMC)

, and the motor’s configuration is shown in Figure 1. The motor features a nine-slot structure with concentrated windings for six poles.

To effectively reduce the magnet usage compared to the designed SPM, a CPM type was applied to propose a design that minimizes magnet consumption. For this, the flux characteristics of the SPM, conventional CPM, and proposed CPM utilizing AIP were compared using the magnetic equivalent circuit.

2.1. Design of Conventional CPM and Proposed CPM

Based on the results of the conventional SPM design, the conventional CPM and the proposed CPM were designed following the process shown in Figure 2.

Figure 3 shows a comparison of the shapes of the CPMs. In Figure 3a, the conventional CPM is shown, which used half the number of magnets compared to the conventional SPM. In this CPM, the permanent magnets were placed on alternating poles of the rotor. The structure of the proposed CPM, utilizing AIP extensions beneath the magnet poles for improved performance, is shown in Figure 3b. Despite having different magnet pole angles, the rotor structure of the proposed CPM maintained the same magnet volume as the conventional CPM. The key mechanical specifications of the three models discussed in this paper are summarized in

Table 2. Comparison of main parameters and specifications of each model.

Item	Unit	Value
		Conventional SPM	Conventional CPM	Proposed CPM
Air gap	mm	0.5	0.5	0.5
Stack length	mm	30	30	30
Rotor diameter	mm	21.9	21.9	21.9
Stator diameter	mm	38.4	38.4	38.4
Magnet thickness	mm	4	4	3
Number of turns	-	31	31	31
Slot fill factor	%	30	30	30
Rated current	Arms	1.4	1.4	1.4
Rated power	W	12	12	12
Rated speed	RPM	3500	3500	3500
PM material	T	Br = 0.4 (Ferrite)
Core material	-	50H1300 (by NSSMC)

.

2.2. Analytical Comparison of Magnetic Flux

To analyze the characteristics of the proposed CPM utilizing AIP, a comparison was made between the magnetic circuit structures of the conventional SPM and CPM. The flux values obtained from the equivalent circuit analysis were compared to highlight the advantages of the proposed CPM with AIP.

Figure 4 shows the simplified magnetic equivalent circuits for the conventional SPM, conventional CPM, and proposed CPM. To simplify the mathematical modeling, the saturation of the iron core, fringing effects, and flux leakage were neglected. Additionally, the stator and rotor back iron in all models were assumed to be identical, so their reluctances were disregarded. However, in the proposed CPM, an AIP extension was included beneath the magnet. Therefore, its reluctance was added to the equivalent magnetic circuit calculation to ensure accurate results.

The magnetic reluctance in the SPM model consisted of three main components: the air gap, the magnet, and the core. The reluctance for each section was defined as follows. The reluctance of the air gap was expressed as

$$R_g={\displaystyle \frac{l_g}{{\mu }_0{\cdot A}_g}}$$

(1)

$$R_m={\displaystyle \frac{l_m}{{\mu }_m{\cdot A}_m}}$$

(2)

where $l_g$ is the air gap length, ${\mu }_0$ is the permeability of free space, $A_g$ is the cross-sectional area of the air gap, $l_m$ is the magnet length, ${\mu }_m$ is the permeability of the magnet, and $A_m$ is the cross-sectional area of the magnet.

The core reluctance followed a similar structure but was generally much lower due to the high permeability of the core material. Thus, the total reluctance for the SPM model was given by the sum of all three components:

$$R_{total, SPM}=R_g+R_m+R_{core}$$

(3)

Consequently, the flux in the SPM model, considering the magnetomotive force (MMF) $F_{m,SPM}$, was expressed as

$${\varphi }_{SPM}\left(\theta \right)={\displaystyle \frac{F_{m,SPM}}{R_{total,SPM}}}·\mathrm{c}\mathrm{o}\mathrm{s}\left(\theta \right)$$

(4)

where $\theta$ represents the rotor position. The flux varied sinusoidally with the rotor position due to the periodic nature of the rotor’s magnetic field.

In the case of the conventional CPM model, the flux path encountered an additional component due to the alternating arrangement of the magnets. As a result, the flux had to traverse two air gaps, effectively doubling the reluctance in that section. The total reluctance for the CPM model was represented as

$$R_{total,\mathrm{C}PM}={2R}_g+R_m+R_{core}$$

(5)

However, in the CPM model, there was also an additional reluctance from the iron pole, $R_{ip}$, which was introduced into the magnetic circuit due to the presence of iron poles in the rotor structure. The reluctance of the iron pole was given by

$$R_{ip}={\displaystyle \frac{l_{ip}}{{\mu }_c·A_{ip}}}$$

(6)

$$R_{total,\mathrm{C}PM}={2R}_g+R_m+R_{ip}+R_{core}$$

(7)

where $l_{ip}$ is the length of the iron pole, ${\mu }_c$ is the permeability of the core, and $A_{ip}$ is the cross-sectional area of the iron pole. The total reluctance included the reluctance of the air gap, magnet, iron pole, and core.

Given that ${\mu }_c$≫ ${\mu }_m$, the reluctance of the magnet $R_m$ was significantly greater than that of the iron pole $R_{ip}$, and, thus, $R_{ip}$ could be neglected in most cases. Hence, the simplified flux equation for the CPM model was

$${\varphi }_{CPM}\left(\theta \right)={\displaystyle \frac{F_{m,CPM}}{R_{total,CPM}}}·\mathrm{c}\mathrm{o}\mathrm{s}\left(\theta \right)$$

(8)

In the proposed CPM model, an additional iron pole extension was introduced to reduce the magnetic air gap and optimize the magnetic flux path. This additional element increased the overall reluctance slightly but reduced the air gap, thus improving the flux linkage. The total reluctance for the proposed CPM and the reluctance of the additional iron pole extension were expressed as

$$R_{total,PCPM}={2R}_{g2}+R_{m2}+R_{ip}+R_{aip}+R_{core}$$

(9)

$$R_{aip}={\displaystyle \frac{l_{aip}}{{\mu }_c·A_{aip}}}$$

(10)

where $R_{g2}$ is the reluctance of the reduced air gap, $R_{m2}$ is the reluctance of the optimized magnet, $R_{aip}$ is the reluctance of the additional iron pole extension, $l_{aip}$ is the length of the additional iron pole extension, and $A_{aip}$ is its cross-sectional area.

Since ${\mu }_c$≫ ${\mu }_m$, both $R_{ip}$ and $R_{aip}$ could be neglected in most cases. Thus, the simplified flux equation for the proposed CPM became

$${\varphi }_{PCPM}\left(\theta \right)={\displaystyle \frac{F_{m,PCPM}}{R_{total,PCPM}}}·\mathrm{c}\mathrm{o}\mathrm{s}\left(\theta \right)$$

(11)

In the proposed CPM model, the magnet reluctance $R_{m2}$ was smaller than $R_m$ due to the optimized magnet length $l_{m2}$ and larger cross-sectional area $A_{m2}$. As a result, the total reluctance was reduced, leading to an increased flux compared to the conventional CPM.

Through these equations, the flux density of the SPM, CPM, and proposed CPM could be compared. It can be observed that the flux path was optimized in the proposed CPM due to the inclusion of the AIP, resulting in improved flux density compared to the conventional CPM design. By adjusting the pole arc angle while considering the structural constraints of the rotor, the proposed design achieved a flux value comparable to that of the SPM, while significantly improving the flux compared to the conventional CPM. The flux variations of the SPM, CPM, and proposed CPM, as a function of rotor position, are visually compared in Figure 5.

As shown, the proposed CPM exhibited a higher flux density across most of the rotor positions compared to the conventional CPM, confirming the effectiveness of the AIP in optimizing the magnetic flux path. Although the SPM model showed the highest peak flux, the proposed CPM achieved flux values comparable to the SPM in certain regions, despite using less magnet material. The conventional CPM, on the other hand, displayed the lowest flux density, further emphasizing the performance improvements gained through the proposed design.

It was confirmed that the proposed CPM improved the flux value compared to the conventional CPM by applying the AIP. Additionally, an optimal design was carried out to maximize the performance of the proposed CPM.

3. Optimal Design of Proposed CPM

The magnet pole structure in the proposed CPM was the most crucial factor that affected the torque characteristics of the machine. Therefore, the PM arc shape, the magnet pole angle, and the consequent pole angle were considered the key design parameters to improve cogging torque, output torque, and torque ripple. Figure 6 shows the rotor structure with the design variables.

In Figure 6, θ₁ represents the angle of the magnet pole arc and θ₃ indicates the pole arc angle of the core where no magnet is attached. x refers to the length between the two ends of the pole arc, and this length, along with the center point of the pole arc, was used to draw the arc [22]. The radius of the arc was determined by moving along the y-axis until the two ends of the arc coincided. θ₂ was determined by θ₁ and θ₃.

To verify that the correct variables were chosen for optimization, a sensitivity analysis was performed. The effect of the design variables on the optimization objectives is illustrated in Figure 7. Since the stator windings were identical across all three motor models, it was assumed that efficiency, output power, and maximum power would not vary significantly. Therefore, the optimization primarily focused on torque characteristics, such as cogging torque, output torque, and torque ripple.

For the proposed CPM, the design variables of the PM arc $\mathrm{x}$, magnet pole angle ${\mathsf{\theta }}_1$, and CP angle θ₃ showed high sensitivity to the cogging torque, the output torque, and the torque ripple. Based on the sensitivity analysis results, the variables and their ranges were defined.

To optimize the design, the following conditions were applied to the variables.

(1)

${\mathit{V}}_m$ = ${\mathit{V}}_{m\_\mathit{CPM}}$

(2)

θ₁ + 2 × θ₂ + θ₃ = 120°

(3)

45° ≤ θ₁ ≤ 65°, 30°≤ θ₃ ≤ 45°

(4)

x + y + z = 4 mm

(5)

0 ≤ x ≤ 0.8 mm, with z depending on θ₁

where ${\mathit{V}}_m$ is the volume of the magnet and ${\mathit{V}}_{m\_CPM}$ is the volume of the magnet in the conventional CPM. θ₁, θ₂, and θ₃ are the mechanical angles of the rotor, and θ₂ depended on θ₁ and θ₃ based on condition 2. To maintain the same magnet volume as the conventional CPM, when the magnet pole angle θ₁ was increased, the AIP extension $\mathit{z}$ was added. During the optimization process, the magnet volume was fixed to match that of the conventional CPM. The objective function was defined to maximize the output torque while minimizing cogging torque and torque ripple.

An outline of the optimized design process is shown as a flow chart in Figure 8.

First, conditions such as 1–5 were assigned to the objective functions and the design variables. Then, the Latin hypercube sampling (LHS) technique was applied to create the sample models. To calculate the sample model, the FEA method was used. Based on the analysis results of the samples, an approximate model was generated by the Kriging method [23,24,25]. Optimized results were obtained using GA on the Kriging model. As a result, it was validated by 2-D FEA.

4. Comparison of Conventional and Proposed Machines

This section examines the electromagnetic performance of the proposed CPM and compares it to that of the conventional SPM machine and the conventional CPM. The flux density and flux line plot of the proposed CPM are shown in Figure 9.

Figure 10, Figure 11 and Figure 12 show the back EMF, back EMF Fourier transform, cogging torque, and torque waveform, respectively.

Figure 10a,b show the even-order harmonics generated by the unbalanced structure of CP in the conventional CPM. These harmonics were significantly reduced by the compensation of the extended magnet pole angle and the additional iron pole extension under the magnet pole in the proposed CPM. Hence, the back EMF total harmonic distortion (THD) was reduced from 5.2% to 3.35% in the proposed machine.

In Figure 11, the cogging torque of the proposed CPM demonstrates a significantly lower peak-to-peak value compared to both the conventional SPM and CPM designs. This improvement was directly attributed to the additional iron pole extension, which effectively increased the number of iron poles. By increasing the number of iron poles, the frequency of the cogging torque was raised, which, in turn, reduced the peak-to-peak amplitude. The additional iron pole extension also ensured that the magnetic flux distribution was more uniform, reducing the negative effects of magnetic interaction between the rotor and stator and leading to smoother operation.

In Figure 12, the proposed machine shows a slight reduction in average torque, 6.1% lower than that of the conventional SPM. However, the proposed CPM exhibited a 6.8% reduction in torque ripple compared to the SPM, indicating improved torque stability. Furthermore, when compared to the conventional CPM, the proposed CPM demonstrated a 17.9% increase in average torque and a 74.9% reduction in torque ripple. These results suggest that the proposed design effectively optimizes both torque performance and ripple reduction, making it highly suitable for applications that demand smooth and reliable operation.

The overall performance is summarized in

Table 3. Comparison of three model topologies.

Item	Unit	Value
		Conventional SPM	Conventional CPM	Proposed CPM
Back EMF	V	4.04	3.2	3.81
THD	%	5.2	22.9	3.35
Cogging torque	mNm	2.13	4.88	1.57
Torque	mNm	32.9	26.2	30.9
Torque ripple	%	13.2	49	12.3
Magnet volume	mm3	5397	2698.5	2698.5
Torque per magnet volume	${\displaystyle \frac{\mathrm{m}\mathrm{N}\mathrm{m}}{\mathrm{m}{\mathrm{m}}^3}}$	${6.1\times 10}^{-3}$	${9.7\times 10}^{-3}$	${11.5\times 10}^{-3}$

. The performance analysis in Table 3 further demonstrates the efficiency of the proposed CPM in terms of torque per magnet volume. The proposed CPM achieved 45% and 16% higher torque per magnet volume than the conventional SPM and CPM, respectively. This improvement highlights the capability of the proposed design to deliver higher torque output without increasing the consumption of permanent magnets (PMs). The optimization of the rotor design, especially with the inclusion of the iron pole extension, results in more efficient use of the magnetic material and overall enhanced performance.

These findings, although based on 2D FEM simulations, are consistent with trends observed in similar rotor design optimizations. The reduction in cogging torque and torque ripple, combined with the increased torque density, highlights the effectiveness of the proposed CPM in applications requiring smooth torque and high-performance density.

The speed–torque–current characteristics were calculated based on motor performance parameters, and efficiency was determined as the relationship between input and output. The efficiency equation was given by

$$\eta ={\displaystyle \frac{{\mathit{P}}_{\mathit{o}}}{{\mathit{P}}_{\mathit{i}}}}={\displaystyle \frac{{\mathit{\omega }}_{\mathit{m}}{\mathit{T}}_{\mathit{e}}}{{\mathit{V}}_{\mathit{t}}\mathit{I}}}$$

(12)

where ${\mathit{P}}_{\mathit{o}}$ represents the output power and ${\mathit{P}}_{\mathit{i}}$ denotes the input power, expressed as ${\mathit{\omega }}_{\mathit{m}}{\mathit{T}}_{\mathit{e}}$ and ${\mathit{V}}_{\mathit{t}}\mathit{I}$, respectively.

Additionally, the motor’s performance, including its T-N characteristics, was analyzed using the torque constant results from the FEM analysis and the motor parameters provided in Table 2. This analysis offered valuable insights into the motor’s performance under various load conditions.

Table 4. Motor T-N characteristics analysis using motor parameters.

Conventional SPM				Conventional CPM				Proposed CPM
${\mathit{T}}_{\mathit{e}}$ [mNm]	${\mathit{\omega }}_{\mathit{m}}$ [rpm]	$\mathit{I}$ [A]	$\mathit{\eta }$ [%]	${\mathit{T}}_{\mathit{e}}$ [mNm]	${\mathit{\omega }}_{\mathit{m}}$ [rpm]	$\mathit{I}$ [A]	$\mathit{\eta }$ [%]	${\mathit{T}}_{\mathit{e}}$ [mNm]	${\mathit{\omega }}_{\mathit{m}}$ [rpm]	$\mathit{I}$ [A]	$\mathit{\eta }$ [%]
0	3657	0.14	0.0	0	4592	0.18	0.0	0	3894	0.15	0.0
10	3547	0.56	73.5	10	4418	0.70	73.1	10	3768	0.60	73.4
20	3436	0.98	81.4	20	4243	1.23	80.2	20	3643	1.05	81.1
30	3325	1.40	82.7	30	4069	1.76	80.8	30	3518	1.49	82.2
40	3215	1.83	82.0	40	3894	2.29	79.3	40	3392	1.94	81.3
50	3104	2.25	80.4	50	3720	2.81	76.9	50	3267	2.39	79.5
60	2993	2.67	78.3	60	3545	3.34	74.1	60	3141	2.84	77.3
70	2883	3.09	76.0	70	3371	3.87	71.0	70	3016	3.29	74.7
80	2772	3.51	73.5	80	3196	4.40	67.7	80	2890	3.74	72.0
90	2661	3.93	70.9	90	3022	4.92	64.3	90	2765	4.18	69.2
100	2551	4.35	68.2	100	2847	5.45	60.8	100	2639	4.63	66.3
110	2440	4.77	65.4	110	2673	5.98	57.2	110	2514	5.08	63.3
120	2329	5.20	62.6	120	2498	6.51	53.6	120	2388	5.53	60.3
130	2219	5.62	59.7	130	2324	7.04	50.0	130	2263	5.98	57.3
140	2108	6.04	56.9	140	2149	7.56	46.3	140	2138	6.42	54.2
150	1997	6.46	54.0	150	1975	8.09	42.6	150	2012	6.87	51.1
160	1887	6.88	51.0	160	1800	8.62	38.9	160	1887	7.32	48.0
170	1776	7.30	48.1	170	1626	9.15	35.2	170	1761	7.77	44.8
180	1665	7.72	45.2	180	1451	9.67	31.4	180	1636	8.22	41.7
190	1555	8.14	42.2	190	1277	10.20	27.7	190	1510	8.67	38.5
200	1444	8.57	39.2	200	1102	10.73	23.9	200	1385	9.11	35.4

and Figure 13 present the performance characteristics of three motor types—conventional SPM, conventional CPM, and proposed CPM—comparing speed, current, power, and efficiency over a torque range from 0 to 200 mNm.

When considering the harness resistance losses between the motor terminal and the power supply, as well as the correlation with measured values, the actual efficiency may have been lower than the calculated values. However, since the three models were expected to have identical stator specifications and similar mechanical losses, the overall efficiency trends of the three models were anticipated to closely match the simulation results.

5. Conclusions

This paper proposed a novel six-pole, nine-slot, concentrated winding CPM featuring an AIP extension beneath the magnet pole. The design increased the magnet pole angle while maintaining the same magnet volume as the conventional CPM, leading to substantial performance improvements. Specifically, the proposed machine achieved a 16% increase in torque per magnet volume compared to the conventional CPM and reduced cogging torque and torque ripple by 68% and 75%, respectively. The AIP extension, which introduced additional iron poles that improved magnetic flux distribution and reduced electromagnetic imbalances, gave enhanced performance results. In comparison to a study by Akihisa et al. [13], which also proposed a consequent pole PM motor with the same magnet volume as a conventional CPM, our work stands out by emphasizing the significant reduction in torque ripple. While Akihisa et al. focused primarily on torque and speed performance without addressing torque ripple mitigation, our design highlights the importance of reducing torque ripple for smoother motor operation.

Our findings indicate that the proposed CPM structure, with its balanced approach to torque output and torque ripple reduction, is particularly well-suited for applications where smooth and efficient performance is critical, such as EPS systems. The proposed design not only maintains efficient PM utilization but also provides clear advantages in terms of performance and cost-effectiveness. Future work will involve experimental validation and 3D modeling to account for additional factors like end-leakage and three-dimensional flux distributions. These efforts will further establish the practical viability of the design and its applicability in real-life automotive and industrial systems. The results indicate that the proposed CPM structure holds significant potential to reduce dependence on rare-earth materials while delivering high-performance metrics, making it a compelling choice for contemporary engineering applications.