A New Parallel Double Excitation Synchronous Machine - IEEE Xplore

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This paper presents 3-D finite-element analysis of a new double excitation synchronous ... and design aspects of this new machine are presented in the paper.
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A New Parallel Double Excitation Synchronous Machine B. Nedjar1 , S. Hlioui1 , Y. Amara2 , L. Vido1 , M. Gabsi1 , and M. Lécrivain1 SATIE, ENS Cachan, CNRS, F-94230 Cachan, France GREAH, EA 3220, Université du Havre, 76063 Le Havre, France This paper presents 3-D finite-element analysis of a new double excitation synchronous machine. It is shown that the machine has true field regulation capability. The principle of operation and design aspects of this new machine are presented in the paper. Comparison of 3-D FEA with an experimental study done on a prototype having a different rotor structure is also investigated. Index Terms—Electromagnetic analysis, field weakening, hybrid excitation, permanent-magnet machines, synchronous machines, 3-D finite-element method.

I. INTRODUCTION

D

OUBLE excitation (or hybrid excitation) consists of combining wound field and permanent-magnet excitations in the same synchronous machine [1]–[17]. The goal of double excitation principle is to combine advantages of permanent-magnet machines, high power density and efficiency, with these of wound field synchronous machines, good field weakening capability. Double excitation machines can be divided into two categories: series double excitation and parallel double excitation machines [8], [11]. Machine presented in this paper belongs to the second category (parallel double excitation machines). First, a nonexhaustive state of the art of double excitation structures is presented. The principle of operation and design aspects of this new machine are then discussed. Finally, comparison with a parallel double excitation prototype having a different rotor structure is also investigated. II. DOUBLE EXCITATION SYNCHRONOUS MACHINES

The double excitation principle allows a wide variety of structures to be realized. Many criteria can then be chosen for the classification of double excitation machines. Classical criteria used for classification of other types of electric machines can be used; as an example: 1) radial field and axial field machines; 2) 2-D and 3-D structure machines. Fig. 1(a) shows an example of an axial flux double excitation machine [16] and Figs. 1(b) and (c) show examples of double excitation machines with 2-D and 3-D structures, respectively [17], [11]. 3-D structures are in general more difficult to analyze and manufacture. Regarding the particular structure of double excitation machines, the presence of two excitation flux sources, two criteria seem more specific for classification of these machines: 1) By analogy with electric circuits, the first criterion concerns the way the two excitation flux sources are combined: series and parallel double excitation machines [8], [11]. Fig. 1. Different double excitation structures. Manuscript received October 06, 2010; revised March 04, 2011; accepted March 15, 2011. Date of publication April 05, 2011; date of current version August 24, 2011. Corresponding author: S. Hlioui (e-mail: sami.hlioui@satie. ens-cachan.fr). 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.2011.2134864

2) The second criterion concerns the localization of the excitation flux sources in the machine: both sources in the stator, both sources in the rotor and mixed localization. By mixed localization it is meant that one source (excitation coils or permanent magnets) is located in the rotor or the

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Fig. 2. 3-D cut view of the double excitation machine.

stator and the other source in the stator or the rotor, respectively. Having excitation coils in the stator is favored to avoid sliding contacts. The first criterion is more linked to flux control capability of double excitation machines. Since excitation flux created by excitation coils should pass through the permanent magnets, which , in series double excihave a low relative permeability tation machines, the flux control capability of the series double excitation structures should be less efficient than that of the parallel double excitation structures. The second criterion is more linked to ease of manufacture and operation of double excitation machines. Indeed, having both excitation sources (excitation coils and permanent magnets) located in the stator presents some advantages from manufacturing and operating point of views because they are fixed: it is far more easier to evacuate losses from fixed parts than moving ones; the presence of excitation flux sources in the stator implies a completely passive rotor which means no need of a containment system and an improved high speed operating capability. The studied double excitation machine has excitation coils located in the stator. Open circuit flux control principle of studied machine is described in following section. III. PRINCIPLE OF OPERATION AND DESIGN ASPECTS A. Principle of Operation Fig. 2 shows a cut view of studied machine. It combines a permanent-magnet excitation with a wound field excitation. Excitation coils are located in the stator, on top of armature end windings, thereby avoiding sliding contacts. Radially magnetized rare earth permanent magnets are located in the rotor. This machine has 12 magnetic poles . The stator is composed of a laminated core, solid iron yoke and end-shields, conventional ac three-phase windings and two excitation annular coils. Solid iron components (external yoke and end-shields) provide a low reluctance path for wound field excitation flux. The rotor is, amongst other things, composed of two solid iron collectors and 12 rare earth permanent magnets. Parts on top and below magnets can either be laminated or massive. Effect of laminating these parts or not on machines performance is investigated. Rotor back iron is magnetically insulated from other ferromagnetic parts of the rotor.

Fig. 3. PM excitation flux trajectories: (a) active flux lines, (b) nonactive flux lines.

Fig. 3(a) shows principal flux trajectory of PM excitation flux. This flux circulates from one pole to another as for classical surface PM machines. These flux lines participate to power conversion contrary to these shown in Fig. 3(b). Flux lines shown in Fig. 3(b) are designated as nonactive because they do not pass through armature windings and therefore do not participate to power conversion. Design procedure should increase reluctance path for these lines. Fig. 4 shows wound field excitation flux trajectories. The machine has two annular excitation coils. Each coil is acting in one kind of magnetic poles. The flux created by an excitation coil passes one time through active part’s air gap (homoplar path). Depending on DC excitation current direction, excitation coils can either be used to enhance or decrease excitation flux passing through armature windings. Finite-element calculations, shown in the following section, will assess the effectiveness of excitation flux control in this machine. Advantage related to the structure of this machine, no sliding contacts and good excitation flux control, should be, however, counterbalanced by a significant increase in material requirement. Stator and rotor core require tangential and axial flux conduction capacity; reasons which make the machine heavier than classical PM machines. B. Design Aspects The first element which has been look at is the spacing between claws on top of permanent magnets (Fig. 5). Fig. 5(a) shows leakage flux lines, which increase if spacing between claws is reduced. This leakage flux path reduces air gap flux and then machines performance. Fig. 5(b) shows the same rotor but with an increased spacing between claws. Increasing this spacing will impact on wound field excitation flux cross-section.

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Fig. 7. Stator/rotor air gaps. Fig. 4. Wound field excitation flux trajectories.

Fig. 5. Leakage flux between rotoric claws.

Fig. 8. 3-D finite-element mesh.

Fig. 6. Rotor with empty spaces filled with permanent magnets. (a) 3-D view. (b) Cut view.

These components of the machine are perfect cylinders made of solid iron and as a consequence the air gap between them can be reduced as regards to air gap thickness in active part of the machine (Fig. 7). Finite-element calculations with an air gap thickness of half that of active part have been done. Results of this analysis are reported in the next section. Section IV presents finite-element analysis of the studied machine. 3-D FEA is used to determine flux distribution and the flux control capability of this machine. It is also used to study effect of some design parameters on machine performance. Main machine data are identical to that of a previous prototype. A comparison study between studied machine and the prototype is presented in this paper.

Effect of this parameter (spacing between claws) is investigated in the next section. Something else which can be done to reduce this leakage flux and in the same time increase air gap flux is to fill empty spaces by permanent magnets (Fig. 6). Fig. 6 shows a 3-D view of the rotor with all spared spaces filled with permanent magnets. For some configurations the area available for magnets is wide and therefore inexpensive ferrite magnets can be utilized without the performance penalty usually associated with their low residual flux density. At least ferrite permanent magnets can be used to fill empty spaces. Since wound field excitation flux has an important axial component (Fig. 4), air gap flux control can be improved by using solid iron for claws instead of laminated sheets. Using solid iron implies, however, higher iron loss and it also offers a low reluctance path for nonactive flux lines [Fig. 3(b)]. For ease of manufacturing rotor back iron can also be realized using solid iron. Another parameter which has been studied is the air gap thickness between rotoric flux collectors and statoric end-shields.

IV. FINITE-ELEMENT ANALYSIS The structure of studied machine requires the use of 3-D finite-element analysis. Fig. 8 shows the 3-D finite-element mesh of studied machine. The nonlinearity of B-H curves of machine’s different parts is considered in this finite-element analysis. Fig. 9 shows two cross sections of the initial machine design. Fig. 9(a) shows an axial cross section (perpendicular to rotation axis) in the active part of the machine and Fig. 9(b) shows a cross section parallel to the axis of the machine [AA plane, Fig. 9(a)]. These figures also shows the main dimensions of the machine. Table I gives the machine’s main data. The initial design parameters of this machine have been derived from a simple analytical model based on a simple reluctance network. For the double excitation circuit’s design, the principle of equalization of flux cross-sections has been used , Fig. 9(b)]. This machine will be referred next as [ “Machine 1”.

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TABLE I DOUBLE EXCITATION SYNCHRONOUS MACHINE DATA

Fig. 9. Initial design main dimensions (Machine 1).

The stator structure and dimensions of this machine is identical to a previously built prototype [11]. The smallest wound field excitation flux cross-section in the stator is the section of the cylinders located at the inner radius of end shields [ , Fig. 9(b)]. The design of the rotor claws was based on the equalization of contact sections between claws and rotoric flux collector with the section [Fig. 9(b)]. Fig. 8 shows the 3-D mesh of 1/12 of the machine (only one pole pair is considered (1/6 of the machine) and half of the machine’s axial length). Finite-element calculations are done on 1/6 of the machine. The machine’s model should consider the entire axial length of the machine. The mesh of only 1/12 of the machine is shown to highlight the smoothness of the mesh. The finite-element calculations are done considering two air volumes at axial ends of the machine (Fig. 10). Fig. 10 shows the different boundary conditions imposed to nodes belonging to outer bounding surfaces of the finite-element model. Dirichlet

boundary condition is applied to bounding surface in both axial limits of the finite-element model and for radii equal to 0 m and (machine’s external radius), respectively. The 3-D mesh of the structure is obtained by extruding a 2-D mesh. In the 2-D geometry, it is necessary, from the very start, to envisage all surfaces which make it possible by extrusion to create volumes of the structure in 3-D. The developed model takes into account the rotor movement. The air-gap is divided into two parts (Fig. 11); a part is linked to the rotor and the other part to the stator. When creating the 2-D geometry, used for the extrusion, at the border of the two parts of air-gap, two lines are confused. One line belongs to the half air-gap linked to the stator and the other one to that linked to the rotor. The same number of elements is imposed on these lines. The number of elements imposed on these lines is equal to 120. These elements are uniformly distributed. As mentioned earlier, due to symmetry considerations, finite-element calculations are carried out on a sixth of the machine, that is to say 60 , which means that there is a node every 0.5 on the lines at the border between the two air-gap parts. The simulation of the movement of the rotor, in this case, can be realized only between two angles which are multiple of 0.5 ; the intermediate positions must be also multiple of 0.5 . For the position where, the lines at the border are completely confused, the nodes on these lines, having same location, are merged. For other positions, where the lines at the border are not any more completely confused (Fig. 12), the steps to be followed are: 1) on the air-gap, the nodes in parts where these lines are intersected, it is sufficient to merge nodes having same location; 2) remaining nodes should be coupled as shown in Fig. 12(b). Fig. 13 shows different electromotive force (EMF) waveforms per turn, at 1000 rpm, for different excitation current values. It can be seen that the double excitation is very effective. The EMF maximum value has been nearly doubled when enhancing air gap excitation flux compared to the case where no field current is applied. It can be noticed that air gap excitation

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Fig. 12. Movement consideration in the finite-element model.

Fig. 10. 3-D finite-element model. (a) 3-D view. (b) Side view. (c) Front view.

Fig. 11. Air-gap division in two parts.

flux can be completely cancelled by applying an adequate field current value.

Fig. 13. EMF waveforms for different excitation currents.

Fig. 14 (Machine 1) shows variation of maximum air gap flux versus field magnetomotive force (MMF). It can be seen that a wide range of flux control can be achieved. Fig. 14 compares flux control capability of different machines. All these machines have same basic structure. Machine 2 has the same structure as machine 1 but uses solid iron claws. It can be seen that flux control is greatly improved by using solid iron claws. However, an increase of rotor’s iron loss can be feared. Machine 3 has different claws shapes [Fig. 5(b)] compared to machine 1 [Fig. 5(a)]. The leakage flux between claws has been reduced by nearly 10% compared to machine 1, but it hasn’t affected the open circuit air gap flux when no field current is applied. Furthermore,

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Fig. 14. Flux control capability for different machines.

Fig. 15. Sensitivity analysis for the spacing between claws. (a) Angle definition. (b) Flux control capability.

effectiveness of flux control, as easily expected, has been reduced compared to machine 1. Fig. 6 shows rotor of machine 4. All empty spaces have been filled with ferrite permanent magnets ( T). Compared to other machines the open circuit air gap flux has been increased, but the field weakening capability has been reduced. For a field MMF of ATs the flux is reduced in same proportions as machine 1, but since flux value has been increased, when no field current is applied, cancellation of open circuit air gap is no more possible.

MMF = +1500

MMF = 0

MMF =

Fig. 16. Air-gap flux density distribution for different field ATs (Machine 1). (a) Field AT. (b) Field AT. (c) Field AT.

01500

A sensitivity analysis concerning the spacing between claws was carried out using the 3-D finite-element model. Fig. 15 illustrates results of this study. Claws shape corresponds to the one shown in Fig. 5(a) for an angle equal to 3 (Machine 1) and to the one shown in Fig. 5(b) for an angle equal to 16 (Machine 3), respectively. This study confirms previous conclusions concerning the comparison between machines 1 and 3. Calculations done with an air gap thickness of 0.25 mm between rotoric flux collectors and statoric end-shields has shown practically no difference with previous case ( mm). Double excitation flux passes through laminated claws which

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Fig. 18. Lamination effect modeling.

Fig. 19. Stator of double excitation prototype: Details of (a) armature and excitation windings, and (b) end shields.

axial penetration of excitation flux created by field coil is limited by laminated nature of claws [Figs. 16(a) and (c)]. Fig. 17 shows the same distribution for machine 2. Since claws are made of solid iron, excitation flux created by field coil will act over the entire length of active part and air gap flux control is more effective. Flux density is more uniformly distributed, over axial length, compared to machine 1. Machines laminated parts have been modeled with anisotropic magnetic properties. The equivalent value of laminated parts permeability in axial direction has been derived based on experience acquired with this kind of machines [9]. Fig. 18 illustrates how the value of relative permeability in z direction is estimated. Laminated parts are considered as succession of ferromagnetic and nonmagnetic (lamination insulation and parasitic air gaps) materials. A packing factor , defined as the total length of ferromagnetic steel parts divided by total laminated pack length (active length), is set to 97%. Equation (1) gives then the value of the equivalent relative permeability in axial direction

MMF = +1500

MMF = 0

MMF =

Fig. 17. Air-gap flux density distribution for different field ATs (Machine 2). AT. (b) Field AT. (c) Field (a) Field AT.

01500

offer a high reluctance path. Equivalent reluctance of this path should be higher than that offered by air gap between rotoric flux collectors and statoric end-shields and as consequence reduces the effect of this last one. Fig. 16 shows radial component distribution, of air gap flux density, over a magnetic pole (North Pole) in the active part of machine 1, for three different values of excitation current. For a null excitation current air gap flux density distribution is only due to permanent magnets [Fig. 16(b)]. It can be noticed that

(1) where is the relative permeability of ferromagnetic parts. With a value of (corresponding to linear part of the B(H) curve) and , a value of is obtained. The next section presents a comparison of flux control capability of studied machine with a previously built double excitation prototype. V. COMPARISON WITH ANOTHER PROTOTYPE Fig. 19 shows the stator of a 3 kW machine prototype to which studied machine has been compared. Both machines

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Fig. 20. Rotor of double excitation prototype. (a) Lamination sheet. (b) Final assembly. TABLE II DOUBLE EXCITATION SYNCHRONOUS PROTOTYPE DATA

Fig. 21. Excitation flux control characteristic for prototype machine (3-D FE analysis and experimental results).

(studied machine and prototype) have same stator (same stator and overall dimensions) but different rotor structures. Fig. 20 shows details of prototype’s rotor. Rotors of both machines have same overall dimensions but different structures. Figs. 20(a) and (b) show respectively lamination sheet used to build prototype’s rotor and assembled rotor. The prototype has 12 poles as studied machine. This prototype uses ferrite permanent magnets. The principle of flux concentration helps to reach high values of air gap flux density. Table II gives some complementary data concerning this prototype. Before comparing flux control capability of both machines, a finite-element analysis of prototype machine (Figs. 19 and 20) has been conducted. This study helps to estimate to which extent the FEA is effective and accurate. Fig. 21 shows flux variations versus field ampere turns for prototype machine. This figure also compares results from 3-D FEA and experimental results. It can be noticed that experimental and FEA results agree well. The air-gap flux changes with a variation of %, when air gap flux is enhanced, and

MMF = +750

Fig. 22. EMF waveforms for different field MMF for prototype machine (3-D AT. (b) Field FE analysis and experimental results). (a) Field AT. AT. (c) Field

MMF = 0

MMF = 0750

%, when it is weakened, with respect to the no-field excitation flux. 3-D FEA has also been used to calculate EMF waveforms for different excitation currents ( AT [Fig. 22(a)], 0 AT [Fig. 22(b)] and AT [Fig. 22(c)]. Figs. 22(a), (b), and (c) compare EMF waveforms obtained with the 3-D FEA and experimental measurements for different field MMF. The 3-D FEA is used to compute flux variations over an electrical period; the EMF waveforms are then obtained by means of numerical derivation of flux waveforms. EMF measurements

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alytical models are more convenient to use in design optimization process, especially for machines with 3-D structure (magnetic flux flowing in the three dimensions). REFERENCES

Fig. 23. Flux control capability comparison.

are made at a rotational speed of 170 rpm. As seen an excellent agreement between FE results and measurements is obtained. This study confirms the fact that the finite-element analysis is effective and accurate. Fig. 23 compares flux variations versus field ampere turns for studied machine (3-D finite-element analysis) and prototype machine (experimental measurements). The goal here is to compare performance of both machines in terms of flux control capability. Although, performance of studied machine seems to be lower than that of prototype machine no definitive conclusions can be drawn. It should be noticed that both machines have not been optimized and that more investigations are needed. However, the fact that air gap excitation flux can be completely canceled for the new double excitation machine constitutes an interesting characteristic. A characteristic which can be advantageous in case of a fault accruing during operation, as a phase short-circuit. Comparing different machines is not an easy task. Machines should be first optimized for different applications implying different power levels and volume constraints before establishing advantages and drawbacks of each structure. VI. CONCLUSION This paper presents a new double excitation machine with good field weakening capability. This structure belongs to parallel double excitation machines. Excitation coils located in the stator allow an effective air gap flux control while avoiding permanent-magnet demagnetization risk. Comparison with a double excitation prototype having same stator structure and overall dimensions, but different rotor structure, is also presented. This machine can be either used as a motor or generator. Some design parameters of this structure have been investigated using 3-D FEA. In order to assess the effectiveness and accuracy of the finite-element method, a 3-D finite-element analysis has also been applied to the double excitation prototype (having the same stator, but a different rotor structure). Comparison of FE and experimental results tends to confirm this fact. This study will be used to help establish an analytical model, based on reluctance network, of the new structure [18], [19]. An-

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