Influence of Magnetic Materials on Claw Pole Machines ... - IEEE Xplore

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Valeo Electrical Systems, F94046 Créteil Cedex, France. This paper presents a study on magnetic materials for claw pole machines which are widely used ...
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IEEE TRANSACTIONS ON MAGNETICS, VOL. 46, NO. 2, FEBRUARY 2010

Influence of Magnetic Materials on Claw Pole Machines Behavior Li Li1 , Afef Kedous-Lebouc1 , Albert Foggia1 , and Jean-Claude Mipo2 G2Elab, Grenoble Electrical Engineering Laboratory, Grenoble INP-UJF-CNRS UMR 5269, ENSE3-BP 46-F38402, Saint-Martin-d’Hères Cedex, France Valeo Electrical Systems, F94046 Créteil Cedex, France This paper presents a study on magnetic materials for claw pole machines which are widely used nowadays in the automobile industry for their simple manufacturing and cost reasons. The finite element analysis with the loss surface model is used to evaluate the iron losses for three categories of material: M800-50A, M330-35A, and AFK502. Comparisons are made in terms of output current, efficiency, as well as iron loss as main characteristics of the machine. On the other hand, the experimental method of loss separation provides a clear idea of the loss distribution in the machines, from which the iron loss can be deduced. A qualitative comparison is made between the two methods, which proves to be coherent. Index Terms—Claw pole machine, electric machines, electrical engineering, finite element analysis, iron loss estimation, loss surface model, magnetic materials.

I. INTRODUCTION HE claw pole machine (Fig. 1) has a large application in the automobile sector. The competitive automobile market makes this domain under constant development. Despite the numerous advantages, its efficiency, around 70%, is always a major concern. Meanwhile, the over-heating introduces additional losses, failures, or full stops. Therefore, many efforts have been carried out in order to analyze its behavior and improve its performance [1]–[5]. The objective of this study is to have a comparison of different materials, to research how to improve the performances of this machine by using high-quality magnetic materials.

T

Fig. 1. Claw pole machine. (a) Stator. (b) Rotor.

TABLE I MAIN CHARACTERISTICS

II. DESCRIPTION OF CLAW POLE MACHINES The particularity of the machine is its claw pole rotor which consists of an excitation winding enclosed by the claws and probably the inter-polar magnets. The machine, of which the main characteristics are presented in Table I, delivers a dc current, through a bridge rectifier, to a battery of 13.5 V. A double three-phase system (Fig. 2) is used to reduce the current ripple, the harmonic rate, and to some extent the iron losses. Furthermore, rare-earth permanent magnets are inserted in the interpolar regions which not only provide additional flux but also prevent flux leakage between two successive poles. III. LOSS SURFACE MODEL FOR IRON LOSS PREDICTION The LS model developed in G2elab [6] is a macroscopic dynamic and scalar hysteresis model which links the applied field H to the average flux density B considered in the cross section of the sheet. It includes all the dynamic effects and assumes that

Fig. 2. Double three-phase system.

Manuscript received June 20, 2009; revised August 31, 2009; accepted September 01, 2009. Current version published January 20, 2010. Corresponding author: L. Li (e-mail: [email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TMAG.2009.2032520

the material behavior is completely defined knowing the instanwhich is a major taneous value of B and its time variation factor for loss estimation [7], [8]. This model is identified by a unique characteristic surface (Fig. 3) obtained experimentally under a trian-

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LI et al.: INFLUENCE OF MAGNETIC MATERIALS ON CLAW POLE MACHINES BEHAVIOR

Fig. 3. Characteristic surface

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H (B; dB=dt) obtained at 1.75 T.

gular and frequency variable B. For an arbitrary B(t) signal, the field H(t) calculated contains two terms:

(1) is determined by a simple static hysteresis model and is extracted from the dynamic part of ( ) surface. This LS model is applied to estimate iron loss in electrical machines [9]. The model is coupled to a finite element simulation in order to obtain the local B(t) variations and to calculate then the local and global magnetic losses. It has been validated for several machines at variable operating points. In most cases, its precision is better than 20%.

Fig. 4. Axial component of the flux density in direction z. (a) Observation support. (b) Axial component of the flux density.

IV. LAMINATION EFFECT IN THE SIMULATIONS Due to the trapezoidal form of the claw poles, previous studies have shown that the magnetic flux density doesn’t propagate only in the lamination plane of the stator. There exists an axial component of the magnetic flux density whose frequency is twice the synchronous frequency because of the reluctance variation of the three-dimensional rotor. This axial component is an important source of iron loss because of its higher frequency knowing that the loss by eddy currents is proportional to the square of the frequency. Therefore, it is important to take into account this factor for the iron loss estimation. In the reality, the magnetic circuit of the stator is anisotropic because of the lamination. The anisotropic equivalent model, implemented in the software FLUX, is used for our simulations. For this anisotropic model, the relative permeability in the axial direction is defined by:

(2) where is the local permeability of the ferromagnetic material, which depends on the working point on the magnetic characteristic; and is the iron filling factor, ratio of iron thickness/total thickness. The simulation results obtained when lamination is taken into account are compared to those obtained with bulk materials. The

Fig. 5. Iso-flux density of stator at 3000 rpm with the same scale from 0 to 2 T. (a) Isotropic stator. (b) Laminated stator.

results, at 3000 rpm, show that the axial component of the flux density is largely reduced (Fig. 4) due to the lamination effect. The laminated stator is more saturated in the sheet plane (Fig. 5) and the machine output current (Fig. 6) is slightly lower, which is probably because of the reduction of the effective iron length. The iron loss is estimated to 43 W and 56 W, respectively, for the bulk and the laminated stator. The difference can be explained by the fact that since the axial component of the flux density is less important in a laminated stator, then the flux density in the laminations is greater than in the case of massive stator. Furthermore, the eddy currents generated by the axial component in the lamination planes are not presently taken into account in the iron loss estimation.

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IEEE TRANSACTIONS ON MAGNETICS, VOL. 46, NO. 2, FEBRUARY 2010

Fig. 6. Current at 3000 rpm with massive and sheet iron. Fig. 7. Energy balance of a claw pole machine. TABLE II IRON LOSS ESTIMATED AT 6000 RPM

TABLE III IRON LOSS ESTIMATED AT 10 000 RPM

V. COMPARISON OF DIFFERENT MATERIALS A. Iron Loss Estimation for M800-50A and M330-35A The simulations are performed for three categories of materials: M800-50A, M330-35A, and AFK502. In automobile applications, the current value is specified according to the speed of the vehicle in order to guarantee the expected operation of every electrical equipment. Under these conditions, the claw pole machines have to be optimized in terms of efficiency, which means loss minimization. We compared the two materials for three load values: 20 A, 110 A, 200 A, at two operating speeds: 6000 rpm (Table II) and 10 000 rpm (Table III) by using the method described in the previous section. The stator iron loss is not so sensitive to the frequency as we thought, which is probably because the machine is practically short-circuited at high speeds, and thus the machine is less saturated. The rotor iron loss estimation is more delicate. The loss generated by the eddy currents is very important because of the bulk material. The analytical method is not so accurate and the finite element analysis requires an excessively fine meshing due to the skin effect. The meshing becomes unrealizable for high frequencies. The surface impedance method should be efficient for the iron loss estimation in the rotor, but it needs adaptation in our case. B. Experimental Loss Separation The principle of loss separation is to establish a method for loss determination which is based on the current and voltage measurements. The energy balance of a claw pole machine is

Fig. 8. Different losses in a claw pole machine with maximal charge 200A.

TABLE IV IRON LOSS MEASURED AND PREDICTED AT 6000 RPM FOR A M800-50A STATOR

shown in Fig. 7. The iron loss can be obtained by this loss separation method using the following equation: (3) is the absorded mechanical power; is the output; where is mechanical and ventilation losses; is the rotor Joule is the stator Joule loss; and is the diode Joule loss. loss; Fig. 8 shows the evolution of the different losses at different speeds. It proves that iron loss increases with the speed, which is one of the most influential factors. The comparison between experimental results and simulations is listed in Table IV. This is a qualitative comparison in which experimental results include both the rotor iron loss and the stator iron loss. In the machine design process, the thermal aspect becomes a critical point because higher output current is demanded in a reduced size. The machine should be designed according to the loss evacuation capacity of the ventilation system. The loss surface model provides a loss assessment at the beginning of this

LI et al.: INFLUENCE OF MAGNETIC MATERIALS ON CLAW POLE MACHINES BEHAVIOR

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Fig. 9. B(H) curves characterized in G2elab and used for prototype realization. Fig. 10. Iso-flux density at 1800 rpm with the same scale from 0 to 2 T. TABLE V CURRENT COMPARISON FOR DIFFERENT MATERIALS

process, and can be integrated as a constraint in the optimization cycle. C. Noble Material: AFK502 In this section, the traditional materials, M800-50A and SAE1006, of the claw pole machines are replaced by the noble iron-cobalt alloy, Fe49%Co2%V (AFK502); the B(H) curves are shown in Fig. 9. The comparisons are made, in Table V, essentially for two operating speeds: the base speed 1800 rpm and at 6000 rpm. The machine with an AFK502 rotor generates a current 20% higher than its counterpart at 1800 rpm. This result is very interesting. On one hand, a higher power machine can be achieved. On the other hand, for the same output, efficiency can be increased thanks to a reduction of the armature turn number and thus the Joule loss. The AFK502 rotor is less saturated than that of SAE1006 as is shown in Fig. 10. The fact of using AFK502 does not improve the current at high speeds is because the machine is practically under short circuit and is far from being saturated. VI. CONCLUSION The finite element analysis with the loss surface model is adopted to estimate the iron losses in a series of claw pole machines. The results confirm the importance of iron loss modeling during the machine design. Besides, the experimental loss separation, a method based on the electrical measurements gives an

evaluation of the loss distribution. Especially, it provides a qualitative comparison with the loss surface model, which proves to be totally reasonable. The simulations using AFK502 confirm that the characteristics of the machine can be improved by replacing the traditional materials by special magnetic materials which are more effective in terms of magnetic saturation polarization, permeability, magnetic loss, etc. Therefore, it is possible to reduce the machine dimensions and thus improve the power/weight ratio. But a simple replacement does not guarantee a significant improvement of the performances, because the topology of the machine is not necessarily optimized. A loss surface model for the material AFK502 is under construction for further study and a prototype with a rotor of AFK502 is on the way. REFERENCES [1] M. Hecquet and P. Brochet, “Modeling of claw-pole alternator using permeance network coupled with electric circuits,” IEEE Trans. Magn., vol. 31, no. 3, pp. 2131–2134, May 1995. [2] S. Kuppers and G. Hennenberger, “Numerical procedures for the calculation and design of automotive alternators,” IEEE Trans. Magn., vol. 33, no. 2, pp. 2022–2025, Mar. 1997. [3] L. Albert et al., “Sizing of automotive claw-pole alternator based on analytical modeling,” in ICEM, Cracow, Poland, Sep. 2004, p. 400. [4] Y. Huang et al., “Design and analysis of a high-speed claw pole motor with soft magnetic composite core,” IEEE Trans. Magn., vol. 43, no. 6, pp. 2492–2494, Jun. 2007. [5] C. Kaehler and G. Henneberger, “Transient 3-D FEM computation of eddy-current losses in the rotor of a claw-pole alternator,” IEEE Trans. Magn., vol. 40, no. 2, pp. 1362–1365, Mar. 2004. [6] T. Chevalier et al., “A new dynamic hysteresis model for electrical steel sheet,” Phys. B: Phys. Condens. Matter, vol. 275/1–3, pp. 197–201, 2000. [7] G. Bertotti, “General properties of power losses in soft ferromagnetic materials,” IEEE Trans. Magn., vol. 24, no. 1, Jan. 1988. [8] M. A. Raulet et al., “The magnetic field diffusion equation including dynamic hysteresis: A linear formulation of the problem,” IEEE Trans. Magn., vol. 40, no. 2, Mar. 2004. [9] T. Chevalier et al., “Estimation of magnetic loss in an induction motor fed with sinusoidal supply using a finite element software and a new approach to dynamic hysteresis,” IEEE Trans. Magn., vol. 35, no. 5, pp. 3400–3402, Sep. 1999.