New High Voltage 2-Pole Concentrated Winding and Corresponding Rotor Design for Induction Machines

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IECON2015-Yokohama November 9-12, 2015 New High Voltage 2-Pole Concentrated Winding and Corresponding Rotor Design for Induction Machines Oleg Moros, Gurakuq Dajaku FEAAM GmbH D-85577 Neubiberg,

de Dieter Gerling Chair of Electrical Drives and Actuators Universitaet der Bundeswehr Muenchen D-85577 Neubiberg, Germany Dieter.Gerling@unibw.

1 IECON2015-Yokohama November 9-12, 2015 New High Voltage 2-Pole Concentrated Winding and Corresponding Rotor Design for Induction Machines Oleg Moros, Gurakuq Dajaku FEAAM GmbH D Neubiberg, Germany Christian Klusmann Universitaet der Bundeswehr Muenchen D Neubiberg, Germany Dieter Gerling Chair of Electrical Drives and Actuators Universitaet der Bundeswehr Muenchen D Neubiberg, Germany Abstract This paper presents a design of a two-pole high voltage induction machine with concentrated coils wound on twelve stator teeth. Using a combined wye-delta winding a magnetomotive force distribution with a low content of space harmonics is created. For suppression of still existing undesirable slot harmonics, a special interconnection of rotor bars is examined. In addition to the design with a finite element method, an analytical approach was made in order to verify the obtained results. Particular attention is paid to the following points: - Reduction of installation space - Reduction of production costs - Evaluation of currents and magnetic fields Keywords high voltage, tooth concentrated winding, distortionfree MMF, multi-cage rotor, low cost, low winding factor, induction machine I. INTRODUCTION Recently, the industry is more and more interested in single tooth concentrated windings. Most of them are designed as permanent magnet synchronous motors (PMSM) which operate either with higher spatial harmonics or with a multiple arrangement of the fundamental harmonic around the machine circumference. Hence they are designed with a high number of poles. Unfortunately, such a machine design cannot cover many other low-pole industrial applications like those in difficult operating environments e.g. chemical, oil and gas, water and wastewater sectors, which are currently covered by two- or four-poles high voltage (HV) induction motors (IM) with distributed windings. Synchronous machines with concentrated windings achieve very high power densities and efficiencies and would therefore be well suited for such applications, but another very important requirement which is very often imposed on such a machine is the direct-on-line (DOL) run-up capability. Synchronous machines would need either an auxiliary winding or an auxiliary start-up engine to accelerate the machine up to synchronous speed, which is not applicable as soon as a large load being present already during startup. Furthermore, in order to achieve the desired speed, a gear set would be necessary for a high-pole machine, which is disadvantageous for both the required installation space and the overall efficiency. It is general knowledge that IMs require an almost ideal sinusoidal shape of magnetic flux density in the air gap, which is only possible with a sinusoidal magnetomotive force (MMF) distribution. A distortion-free MMF is the main reason to use distributed windings for IMs. Fig. 1 illustrate, that the resulting large end windings require nearly the half of the given axial installation space and reduce the overall power density of the machine. In HV machines this disadvantage is extremely apparent. By using the new proposed machine design with concentrated coils, many benefits can be achieved: - Higher stack length in constant installation space or - Reduction of installation space (Fig. 2) - Increase of the overall power density - Savings in material and manufacturing costs - Enabling of modular assembly Fig. 1. Exemplary ratio of the stack length to the entire length of the conventional 2-pole HV machine with distributed windings. Fig. 2. Stator end windings of a HV machine with distributed winding (left) and single tooth concentrated coils in modular assembly (right) in comparison /15/$ IEEE

2 II. REQUIREMENTS AND ASSUMPTIONS The general geometrical and electrical parameters e.g. total machine length, diameter, load dependent currents and efficiencies which are used as reference in the new design can be obtained as a fixed value or a recommendation from literature [1], [2] and from official announcements of the respective manufacturers or suppliers e.g. ABB Group, Siemens AG, WEG Industries, etc. Some other parameters e.g. air gap length may vary depending on the application. For induction machines, the size of the air gap δ as a function of the pole pitch τ p with different numbers of poles is shown in Fig. 3 [3]. Especially for IMs, the air gap is a very critical part. A lot of performance parameters like magnetizing current, power factor, over-load capability, cooling, acoustic and magnetic noise, rotor radial movement, tooth pulsation etc. are affected by length of the air gap. However, the air gap has only a slight influence on machine efficiency, which varies between 95.8% for HV modular IMs and 97.1% for HV cast iron IMs [4]. Table I summarizes the main requirements imposed on the new machine design. Since for this machine design and application only air cooling is provided, especially in the rotor a sufficient cooling should be ensured by means of cooling channels, and the rotor losses should be kept at minimum. The temperature requirement of 393K results from the operating temperature of several reference models, so that a fair comparison is ensured. In a low-pole electrical machine with a large shaft diameter, particular attention should be paid to the shaft material. Improvement of magnetic properties of the shaft material reduces the magnetic flux density in the rotor yoke and therefore reduces the rotor losses. Higher shaft magnetic conductivity becomes important especially when the effective rotor yoke width is reduced due to a high number of cooling channels. Fig. 3. Air gap lengths of IM vs. pole pitch [3]. III. WINDING LAYOUT AND ROTOR DESIGN The objective of the presented winding is to provide a rotating electric machine, which is cheap in both material and production, hence with tooth concentrated coils but small amount of parasitic harmonics of magnetomotive force, which is essential for use in IMs. For the presented winding to be efficient a sinusoidal MMF is necessary. To create such, the electric current loading has to be sinusoidal as well. The proposed winding can achieve such a current loading if aligned as contemplated. The winding used in this design on twelve stator teeth consists of two winding systems and provides a combined wye-delta winding with concentrated coils. It is obvious, that each winding phase works with another current vector, so that this kind of winding is often called semi-six-phase winding [5]. For a better optical comparison Fig. 4 shows the current phasors of a usual threephase system and of the six-phase system. TABLE I. MAJOR MACHINE SPECIFICATIONS Dimension Geometrical Requirements Stack length Outer stator diameter Inner stator diameter Shaft diameter Air gap length Electrical Requirements Phase number Number of poles Rated phase voltage Rated current Rated frequency Output Requirements Rated mech. Power Rated torque Torque ripple Efficiency Value 764 mm 800 mm 430 mm 235 mm mm kv 115A 50 Hz 1 MW 3.2 knm < 3 % 95.8 % Fig. 4. Current phasor diagrams; usual three-phase winding (left) and six- or semi-six-phase winding (right). In each phase of each winding system exactly two stator teeth are wound lying exactly 180 degree geometrically and electrically apart from each other. In this way each winding system generates a fundamental wave with the winding factor ξ 1 =0.259 and the 5 th and the 7 th space harmonics with the winding factor ξ 5,7 = Fig. 5 shows the winding diagram of the proposed machine design. In this interconnection the shift angle between currents of the two winding systems is 90 degrees. The 5 th and the 7 th harmonics can be suppressed by

3 geometrical shifting of winding systems by the same angle to each other, which corresponds to exactly three stator teeth. The fundamental harmonic is not affected by this arrangement and can be used as the operating wave. For a complete suppression, the flux produced by the unwanted space harmonics of both winding systems must be equal. Because of the wye-delta circuitry, the wye current has higher amplitude by the factor 3. To adjust the field amplitude of the delta winding to the same value, the number of turns is to be extended by the same factor (1). With this choice not only a good MMF distribution is achieved (see Fig. 7), but also the current densities in both winding systems are equal, which enables a uniform heating and cooling of the machine in total and of the individual machine parts [6]. 3 n Y (1) Having a double-layer winding, the current in every slot consist of two components with a 30 degree electrical phase difference. Comparing two adjacent slots, the resulting currents from each also have a difference of 30 degree. So to acquire a 2 pole sinusoidal current loading and MMF on the stator with 12 slots, the winding provides the best prerequisites. Although the slot harmonics of ordinal number 11 and 13 still have significant proportions in the air gap, with the proposed rotor design not each harmonic contained in the air gap exerts Fig. 5. Proposed winding design; combined wye-delta winding (left) and the one-phase winding scheme (right) Fig. 6. Current loading for the proposed 12-teeth / 2-poles concentrated winding. n Δ Fig. 7. MMF space harmonics for the proposed 12-teeth / 2-poles concentrated winding. an influence on the formation of torque. According to (2) and (4) an appropriate choice of the number of bars connected in parallel with an end ring can lead to the suppression of several harmonics [8]. Fig. 8 shows that the choice of three independent cages [9-10] with eleven bars each reduces the coupling factor of the 11 th harmonic completely to zero while the 13 th harmonic is damped by about 83%. The corresponding stator and rotor circuits are shown in Fig. 9. ˆ 2πrl pπ pα s Φ = B sinc sinc N2 N2 2 with magnetic flux Φ, harmonic ordinal number, magnetic flux density B, stator inner radius r, stack length l, fundamental wave s number of pole pairs p, Number of rotor bars connected in parallel N 2 and skew angle α s. A skewing of rotor bars would not only increase the production costs, but also the risk and value of the inter-bar currents [7]. To avoid these risks, additional manufacturing efforts and cost expenditures skewing is not considered in the proposed machine design (3). ˆ 2π rl Φ = B ξc, (3) N 2 where pπ ξc, = sinc (4) N2 with ξ C, being coupling factor. In Fig. 10 two options of the proposed multi-cage rotor are shown. For the most compact structure a configuration with one conventional end ring can be chosen. For further experimental investigations, the model with three-cage interconnection on both machine sides is well suited. Since three consecutive rotor bars belong to different electrical branches, this construction allows measurements of the inter-bar voltages and inter-bar currents in practice. For the proposed machine design the number of rotor cages and bars was limited to 3 x 11. (2)

Coupling Faktor ξ C, 1 0.8 0.6 0.4 0.2 0-0.2 Fundamental harmonic 11 th harmonic 13 th harmonic -0.4 0 5 10 15 20 25 30 35 40 Number of Rotorbars Fig. 8.

4 Coupling Faktor ξ C, Fundamental harmonic 11 th harmonic 13 th harmonic Number of Rotorbars Fig. 8. Coupling factors for three significant space harmonics. With further increases in dimensions, die casting of rotor bars is no longer an economic process, even in case of aluminum die-casting. In case of copper even more thermal and mechanical difficulties occur, making die-casting very expensive or not yet possible for certain sizes and geometries. IV. SIMULATION RESULTS In this section, the merits of the proposed solution in terms of MMF harmonic effect mitigation are illustrated by means of finite element simulations. The previously mentioned new squirrel cage induction motor design has the following geometrical and electrical characteristics TABLE II. NEW WINDING DESIGN SPECIFICATIONS Dimension Winding type Coil type Number of stator slots Winding factor Air gap length Number of turns per phase Copper filling Number of rotor cages Number of rotor bars per cage Value Combined wye-delta Concentrated mm 2 x (75+130/ 3) 55 % 3 11 Fig. 9. General representation of the stator and rotor circuits. For this machine design the worst case of operating environment is assumed and hence, the air gap is increased to 200 % of the size of the recommended value shown in Fig. 3. The number of turns per phase and winding system has the major influence on the stator main inductivity and therefore on the machine behavior in general. Because of the unusually low winding factor and (5) a high number of turns is needed [7]. L 3 w ξ 4 lr μ 2 p πδ 1 1 1, m = 0 2 (5) Fig. 10. Possible multi-cage rotor configurations; including one conventional end ring (left); containing three independent squirrel cages (right). The next possible solutions would be 39 bars for the fully suppressed 13 th harmonic or 44 bars with 4 cages, which would increase the production costs unnecessarily. In case of copper rotor bars a particularly economical and reliable mass production of large machines can be achieved with prefabricated rotor bars. Compared to copper die-casting, this method offers several advantages: - Exclusion of material defect / gas pockets - Thermal savings - No material aging by thermal processing (T>1400K) - Automation of manufacturing processes with vacuum permeability μ 0, overall number of turns w 1, winding factor ξ 1 and air gap length δ. Precise observance of (1) is important for minimization of rotor losses and torque pulsation. A reasonable setting provides high efficiencies and torque ripple below 2 % of rated torque. Generally known, air gap fields differ from the analytically determined MMFs. Especially harmonics with a low ordinal number are damped stronger at high load conditions by the reluctance of stator iron. It s difficult to investigate these effects analytically, but they must be considered in FEA. As expected, simulation results with a 2D-FEA application show a symmetrical 2-pole flux distribution (see Fig. 11). To verify the optically detected results, the radial component of the air gap field shown in Fig. 11 is analyzed using FFT. Its result is shown in Fig. 12 and there are absolutely no components of the 5 th or the 7 th harmonics. Due to the proposed rotor a considerable proportion of 11 th and 13 th harmonics doesn t have any influence on torque formation which can be proved with FFT analysis of rotor currents in Fig. 14 and Fig

1600 1400 1200 Rotor bar current [A] 1000 800 600 400 200 0 0 5 10 15 20 25 30 Harmonic Order Fig. 15. Spectral analysis of rotor bar current. Fig. 11.

Spectral analysis of radial air gap flux density. As has been shown above, there is almost exclusively the fundamental frequency in rotor bars.

However, the requirements were fully met, even keeping the stack length constant, so instead of ca. 1.5 m (reference design) less than 1 m axial installation space is required.

5 Rotor bar current [A] Harmonic Order Fig. 15. Spectral analysis of rotor bar current. Fig. 11. Magnetic flux lines and flux density distribution Fig. 12. Radial air gap flux density distribution. Air gap magnutude [p.u.] Harmonic Order Fig. 13. Spectral analysis of radial air gap flux density. As has been shown above, there is almost exclusively the fundamental frequency in rotor bars. Profound investigations showed that the air gap harmonics strengthen current displacement effects in rotor bars and thus affect the machine efficiency negatively. However, the requirements were fully met, even keeping the stack length constant, so instead of ca. 1.5 m (reference design) less than 1 m axial installation space is required. Table III summarizes the main output parameters of the machine. In spite of the increased eddy currents in the rotor, the losses are relatively small, which is especially important for an air cooled application. Since the air gap length was selected to be substantially larger than usual, magnetizing current increases affecting the power factor negatively. An application specific adjustment of the air gap length can strongly increase the power factor or improve many other machine parameters. Since the 11 th and the 13 th harmonics still produce losses, in order to reduce the rotor heat generation and to improve the thermal suitability a further increase of the air gap is considerable. TABLE III. ANALYSIS RESULTS AT RATED POWER OPERATION Losses Value Percentage Rotor eddy losses Stator I 2 R losses Core losses Windage, friction, etc. (0.01 x P m) Total losses 4.5 kw 17.0 kw 10.2 kw 10,0 kw 41.7 kw 10,8 % 40.8 % 24.4 % 24.0 % 100 % Mech. power Torque Overall efficiency Power factor Torque pulsation 1 MW 3.2 knm Nm 96,0% 1.6% Fig. 14. Rotor bar current vs. time V. CONCLUSION This paper presents a new three phase concentrated winding with 12 slots, 2 poles developed as a combined wye-delta winding and two corresponding rotor designs. Despite a small number of slots, the new unconventional winding with a low winding factor, but a good MMF spectrum could be created,

6 which is essential for use in IMs. Using a multi cage rotor slot harmonics could be suppressed, and only the fundamental frequency can be detected in rotor bars. Contrary to the assertion, that highly efficient operation is only possible with high winding factor, in this paper it has been clearly shown, that not the winding factor, but rather the MMF distribution function is the most important input for the efficiency of IMs. The objective of this paper is to present a new machine design, which is cheap in both material and production. By using concentrated coils about 33% axial installation space savings could be achieved and therefore the overall power density increases. A low harmonic content in the MMF ensures a lowloss operation and allows the usage of the machine in aircooled applications. REFERENCES [1] W.J. Bonwick: "The Principles of Large AC Generators and Motors", Centre for Electrical Power Engineering, Monash University, Melbourne, Victoria, Australia, [2] E. Bolte: "Elektrische Maschinen, Grundlagen, Magnetfelder, Wicklungen, Asynchronmaschinen, Synchronmaschinen, Elektronisch kommutierte Gleichstrommaschinen", Springer Berlin Heidelberg Verlag, Berlin, Germany, (in german) [3] J. Klamt: "Berechnung und Bemessung elektrischer Maschinen", Springer Berlin Heidelberg Verlag, Germany, (in german) [4] ABB Group: "High voltage induction motors", technical catalog for IEC motors, BU Motors and Generators, Sept [5] K. Henning, W. Hofmann: "Combined stator windings in electric machines with same coils", XXth International Conference on Electrical Machines (ICEM), Marseille, France, [6] O. Moros, D. Gerling: "Geometrical and Electrical Optimization of Stator Slots in Electrical Machines with Combined Wye-Delta Winding", 2014 XXI International Conference on Electrical Machines (ICEM), Berlin, Germany, [7] A. Behdashti; M. Poloujadoff: "A New Method for the Study of Inter-Bar Currents in Polyphase Squirrel-Cage Induction Motors", IEEE Transactions on Power Apparatus and Systems, [8] D. Gerling: "Electrical Machines, Mathematical Fundamentals of Machine Topologies", Springer Verlag, Germany, [9] C.E. Linkous: "High Efficiency Induction Motor with Multi-Cage Rotor", Unated States Patent, US , Fort Wayne, Ind, [10] C.E. Linkous: "Method of Making High Efficiency Induction Motor with Multi-Cage Cirquit Rotor", Unated States Patent, US , Fort Wayne, Ind,

Key Factors for the Design of Synchronous Reluctance Machines with Concentrated Windings

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