novel high-performance synrm design method: an easy

Copyright (c) 2013 IEEE. Personal use is permitted. For any other purposes, permission must be obtained from the IEEE by emailing [email protected].

This article has been accepted for publication in a future issue of this journal, but has not been fully edited. Content may change prior to final publication.

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Abstract—The complex rotor structure of the Synchronous

Reluctance Machine (SynRM) is analyzed in this paper. Three macroscopic design parameters are introduced: insulation ratios in the d- and q-axis and the rotor slot pitch in the d-axis controller angle. These parameters are optimally linked to the microscopic and detailed SynRM rotor geometry (barriers, insulation layer and segments, magnetic layers inside the rotor) parameters (dimensions) by introducing and combining a general rotor arrangement with an analytical explanatory theory. This theory represents the anisotropic behavior of the SynRM rotor structure according to literature. Based on these parameters, a novel, simple, fast and systematic design procedure for a SynRM rotor with specific stator structure is developed and presented. A SynRM rotor can be competitively optimized with respect to an Induction Machine (IM) by a limited number of Finite Element Method (FEM) sensitivity analysis studies of the macroscopic design parameters. The machine torque can be maximized by finding the best insulation ratios, while the torque ripple can be minimized by determining the best rotor slot pitch in the d-axis. Both these optimizations can be defined independently of stator structure. The method is validated by design (using this procedure), prototype and measurement of a specific SynRM machine with 3 barriers and stator standard frame size of 160 (IEC). A heat-run test was done for both the SynRM and its corresponding IM with the same stator and test bench.

Index Terms— Optimization, Synchronous Reluctance, Test, Torque, Torque Ripple.

NOMENCLATURE ALA Axially Laminated Anisotropy Barrier Air (insulation) layer inside the SynRM rotor Di Position of the end of barrier no. i (i = 1, ..., k) (middle) in the

air-gap or rotor slot openings in the air-gap DTC Direct Torque Control e.g. For Example fdi Average MMF over the segment no. i (i = 1, ..., k+1) due to

mmfd FEM Finite Element Method fqi Average MMF over the segment no. i (i = 1, ..., k+1) due to

mmfq GP General Purpose machines, e.g. for pump, fan, compressor IM Induction Machine IPM Interior Permanent Magnet, synchronous machine

Manuscript received Jan. 30 and accepted for publication May 25, 2013. Copyright (c) 2009 IEEE. Personal use of this material is permitted.

However, permission to use this material for any other purposes must be obtained from the IEEE by sending a request to [email protected].

This work was supported by ABB, Corporate Research, Sweden. R. R. Moghaddam is with ABB Corporate Research, Västerås, Sweden (e-

mail: [email protected]). F. Gyllensten is with ABB Motors and Generators, Västerås, Sweden (e-

mail: [email protected]).

k Number of barriers in the rotor per pole kw Insulation Ratio (total insulation over total iron) inside the

rotor of an ALA-SynRM kwd Insulation ratio (total insulation over total iron) in the d-axis

over ld kwq Insulation ratio (total insulation over total iron) in the q-axis la Total insulation materials in the q-axis la+ly Rotor outer radius minus machine's shaft radius lad Total insulation materials in the d-axis over ld ld Cross-section of the rotor in the d-axis Ldm Air-gap Magnetizing Inductance in d-axis of reference frame Lqm Air-gap Magnetizing Inductance in q-axis of reference frame ly Total iron materials in the q-axis and the d-axis MMF Magneto Motive Force due to the current in the stator winding mmfd d-axis component of the stator MMF (fundamental in pu.) mmfq q-axis component of the stator MMF (fundamental in pu.) p Machine's pole pair number PF Power Factor pi Permeance of the barrier no. i (i = 1, ..., k) = Sbi / W1i PM Permanent Magnet PMaSynRM Permanent Magnet assisted SynRM pp Peak-to-peak Sbi Length of the barrier no. i between the q-axis and the d-axis

(i = 1, ..., k) Segment Iron (magnetic) layer inside the SynRM rotor Si Size of the segment no. i (i = 1, ..., k+1) in the d- and q-axis SynRM Synchronous Reluctance Machine T Torque TLA Transversally Laminated Anisotropy UCG Uncontrolled Generation Mode in PM machines VSD Variable Speed Derive W1i Size of the barrier no. i in the q-axis (i = 1, ..., k) Wid Size of the barrier no. i in the d-axis (i = 1, ..., k) Yqi Radial position of the barrier no. i in the q-axis (i = 1, ..., k) α Mechanical angel from the d-axis in the rotor reference frame αm Rotor slot pitch angle (mechanical) β Rotor slot pitch, αm , controller angle Δfi Differential average MMF over the barrier no. i (i = 1, ..., k)

due to mmfq , which is equal to ( fqi+1 – fqi ) η Efficiency ξ Air-gap Saliency Ratio = Ldm / Lqm

I. INTRODUCTION

HE Synchronous Reluctance Machine (SynRM) design has been a major subject of research and development

since the introduction of this technology in 1923 [1]. These research has been focused under three different categories: firstly, maximizing torque capability, [1]- [3], [7]- [10], [12]- [18], [41], secondly, minimizing torque ripple [1]- [3], [19]- [28], [30]- [38], [42], [43] and thirdly, evaluation and benchmarking its performance versus traditional / new industrial solution(s), Induction Machine (IM) / Permanent Magnet (PM) and Interior PM (IPM) machines respectively [2]- [4], [17], [44]- [53]. These attempts have even supported and pushed the introduction of an industrial solution based on

Novel High Performance SynRM Design Method an Easy Approach for a Complicated Rotor Topology

Reza – Rajabi Moghaddam, Freddy Gyllensten, Member, IEEE

T



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SynRM technology in the market for General Purpose (GP) applications such as pumps, fans and compressors, [4], [54]. Recent advances in control and power electronics technologies as well as economical and technical difficulties which have become apparent regarding PM machines have contributed to interest in the SynRM. Using PM materials in a SynRM machine improves its performance. This strategy does raise some problems in the machine operation as well as production. One example is that the PM materials are sensitive to the temperature, therefore, increased temperature in PM during normal operation can reduce the machine torque capability. Demagnetization [5] is another risk in PM machines due to high temperature as well as transient conditions in short circuit conditions and UCG [55] operation of these machines. Using PM material clearly increases the initial cost as well as production cost (handling the PM material in production line) and maintenance cost during operation. The availability of PM material is a risk which can further increase the PM machine cost. In PM machines field-weakening and normal operation requires more complicated control methods in comparison to IM and SynRM for both hardware and software developments due to protection issues in faulty conditions.

Industrial introduction of the SynRM requires development of a fast, accurate, effective/efficient design method. This method should be competitive/compatible with the traditional design methods that are used in electrical machines design, e.g. the fast analytical method in IM design, [39], [40].

In this paper the major behavior and characteristics of an anisotropic structure, suitable for high performance SynRM rotor geometry design, are distinguished and discussed. This issue is based on the combination of the already existing concepts [1]- [3], [7]- [10], [12]- [18] and utilizing a previous advanced conceptual theory for anisotropic structure modeling [13] that analytically explains the SynRM rotor anisotropic structure behavior [1]- [3], [8], [12]- [14], [16]. Secondly, a novel, simple and systematic method for independent torque maximization [20], [41] and torque ripple minimization [20],

[42], [43], in SynRM rotor design for a known stator structure is developed. The carefully selected general rotor shape (Fig. 1) some optimum distribution rules from analytical anisotropy theory [13] and some new design parameters are used to develop a novel FEM aided fast rotor design optimization procedure for SynRM. The goal is to find the simplest, but not necessarily the best, SynRM optimal rotor which can compete with the corresponding IM. Thirdly, a high performance SynRM rotor has been designed according to the methods presented in this paper. This rotor has been designed for use with an IM (ABB) stator with IEC standard frame size 160. Finally, the performance of the IM and SynRM machines are compared under Variable Speed Drive (VSD) conditions.

II. SYNRM TORQUE OPTIMIZATION

A. An Introduction on Torque Maximization

The SynRM rotor has a complex structure and a lot of geometrical parameters are involved in the machine dimensioning and optimization. The previous attempts to tackle the complexity nature of the problem can be found in [2], [9], [10], [12]- [14], [18]. In the approach discussed in [13], [14], an analytical method based on a lumped equivalent magnetic circuit of the machine was used without considering the saturation effect. The method came up with an optimal distribution of insulation material inside the rotor [13], [14]. This method is developed more by implementing a detailed lumped equivalent circuit, including saturation in the simulation by saturable relative permeability, and varying the rotor geometrical parameters freely, in [10], [18]. This method is not accurate enough, because FEM is used partially to correct the optimization. Due to variation of all geometrical parameters in this method, too many machines have to be analyzed and evaluated. The air-gap flux density and SynRM performance are analyzed in [2], [15], that mainly follows Kostko’s method [1]. Similar method is used in modeling of the saturable relative permeability [10] and in effective air-gap in presence of saturation [18]. However, in [9], [12], the direct analytical approach for SynRM design is not followed. But, a FEM parameter sensitivity analysis on a SynRM to investigate the effect of geometrical parameters on machine inductances is used, where the optimization procedure is based on variation of the geometrical parameters directly similar to [10] and analyzing each machine with FEM. In [2], [8], [16] a macroscopic parameter, insulation ratio (kw), for optimizing mainly the Axially Laminated Anisotropy (ALA) rotor structure is introduced. This approach does not deal with the Transversally Laminated Anisotropy (TLA) rotor in general [2]. A similar analysis has been given somehow in [2], [3], [13] when the total amount of air (la) in the rotor is discussed.

B. Objective Facts: The Nature of The Problem

The SynRM rotor complexity naturally increases the optimization steps and time. Especially it is essential to consider the saturation effect [12]. The SynRM optimization, based on pure mathematical algorithms, is studied in e.g. [9]-

~ m

Y q1

S1

S2

d

q

ld

~ m

2~ m

21

321

11 WWlaxisqinairtotal

SSSlaxisqinirontotal

a

y

)21(

21

dd

dd

WWld

WW

ly

ladwd

k

Sb1)1(

11W

12W

ll ya S1

dW2S 2

dW1

p2

A

B

S3

S3

)2( k

Sb2

321

2111

SSS

WW

lyla

wqk

Fig. 1: Proposed rotor geometry and related microscopic and macroscopicparameters definition for SynRM with two interior barriers, [43].



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[12], [18], [56]. It is possible to overcome the nonlinearity nature of the problem by using FEM, especially saturation, but mathematical optimization can be avoided.

C. Suitable Rotor Arrangement and Parameterization in Microscopic Term

The first step is a suitable rotor geometry selection for SynRM. As Kostko has shown [1], the conventional salient pole reluctance machine is not a right choice for SynRM rotor arrangement. Instead, a SynRM with interior flux barriers has to be developed [1]- [3], [8], [9], [13], [16], [18], [41], [47]. The next crucial issue is a simple, flexible and general rotor barrier shape. The findings in [20], [47] show that the multiple flux barrier rotor can be microscopically parameterized as is demonstrated in Fig. 1. The key fact is that the rotor structure has to fully block the q-axis flux while minimally affecting the d-axis flux.

D. Problem Parameterization in Macroscopic Term

Using a rational and general rotor shape will reduce the number of geometric parameters. But they are still microscopic that are related with the barriers and segments dimensions. It is difficult to use directly these microscopic parameters. Independent studies have tried to deal with macroscopic parameters [2], [3], [8], [13], [16]. A macroscopic geometry parameter is theoretically and analytically connected to the machine inductances in [2], [3], [8], [13], [16]. Vagati [2], [3], [13] emphasizes on the total amount of insulation along the q–axis and inside the rotor (la, see Fig. 1). Matsuo/Lipo [2], [16] and Staton/Miller [2], [8] introduce the insulation ratio (kw). The machine anisotropic structure can be characterized by these two parameters (la and kw) through analysis of their effect on (Ldm – Lqm) and ξ.

E. Optimization Strategy and Methodology

The next most important issues are: How should this insulation be introduced in the rotor ? What is the best distribution of the insulation ? The derivations in [2], [3], [13] show each barrier width W1i must follow a distribution rule in order to optimally utilize la . Consequently the q–axis flux is minimized. This rule is expressed according to (1).

bj

bi

j

i

j

i

S

S

W

W

f

f

1

1

(1)

In equation (1), i and j (i ≠ j and = 1, ... , k) are the barriers number, the ∆fi parameter is the difference in the average per unit MMF, sin(pα) , over barrier i , if a q-axis MMF, mmfq is applied to the machine, ∆fi = fqi+1 – fqi , see Fig. 2 (top), for details refer to [2], [3], [13], [15]. The barrier length Sbi for barrier one and two are shown in Fig. 1.

Implementing (1) can be combined with an assumption on the barrier’s permeance pi = Sbi /W1i [2], [3], [13], [15]. The constant and equal permeance (pi = pj), the rotor with “homogenous anisotropic structure” concept, can be combined with (1) this gives (2).

2

)1(

f

f

1

11.

j

i

j

i

j

i

W

Wconst

p

passumption (2)

Similarly there must be a concept to utilize the total iron (ly, see Fig. 1). Generally increasing air reduces Lqm effectively, and increasing iron increases Ldm significantly. Thus, the segment optimal distribution rule has to be strongly interconnected to the d–axis flux maximization problem. A straight forward assumption for segment size Si is that it should be proportional to the average d–axis MMF, which that segment is facing in air-gap. This gives (3).

dj

di

j

i

S

S

f

f

(3)

In equation (3), i and j (i ≠ j and = 1, ... , k+1) are the

B

S1 S 2 S 3 Sk

p2

2 m0

23 m

25 m

2)12( k

m

2)12( k

m

...

...

m

d mmfd

,

Sk 1

A

fd1 fd

2

fd3

fd4

fdk

fdk 1

...)( pCos

D1 D2 D3

Dk

B

S1

S 2 S 3 Sk

p2

2 m0

23 m

25 m

2)12( k

m

2)12( k

m

...

...

m

d mmfq

,

Sk 1

A

fq1

fq2

fq3

fq4

fqk

fqk 1

...

q

f1

f2

f3

fk

1

fk

)( pSin

D1 D2 D3

Dk

Fig. 2: (top) Per-unit MMF distribution over segments due to the q-axis MMF, (bottom) per-unit MMF distribution due to the d-axis MMF over the segments, for the geometry in Fig. 1, [43].



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segments number, fdi parameter is the average per unit MMF, cos(pα), over segment i , if a d-axis MMF, mmfd is applied to the machine, see Fig. 2 (bottom). By (3), it is ensured that the flux densities in all segments are the same and iron utilization in the rotor will be increased.

III. SYNRM TORQUE RIPPLE OPTIMIZATION

A. An Introduction on Torque Ripple Minimization

Torque ripple and other secondary effects such as rotor iron losses, vibration and noise, are important issues in SynRM design [2], [13], [19]- [38], [42], [43], similar to the IM, [39], [40] and IPM machine [3], [5], [6], [11], [26], [31], [34]- [37], [44]. An analytical study of torque ripple and its main root-cause is addressed in [2], [3], [13], [20]- [25], [27], [28], [31], [34]- [37]. A FEM based investigation on SynRM torque ripple is presented in [19], [20], [22]- [27], [30], [34], [36]. Some methods that proposed by different authors to reduce the SynRM torque ripple can be found in [2], [3], [19]- [27], [34], [37], [38]. The constant rotor slot pitch is a general, simple, accepted rule and state-of-the-art method in ripple reduction for all conventional electrical machines including GP machines, but this is not the only way for SynRM torque ripple minimization, some examples can be found in [19]- [21], [23]- [26], [30]- [32], [37], [38]. E.g. selecting independently, different slot opening position for different barriers, this kind of optimization is considered in [23]- [26], [31], [37], [38] or adjusting the air-gap dimension for controlling the air-gap permeance [31], [33], [37], [38] or using different pole pitch sizes in different poles [32] or using mutual harmonic cancelation [24]- [26], [36]. These studies show the complexity of the problem, especially if the iron losses are considered, [20], [29], [35], [36]. E.g. [36] describes a possible trade-off between torque ripple and iron losses in SynRM design. Interaction between the stator slots, and the rotor slots and its magnetic reaction to the stator MMF, plays an important role in the torque ripple developed by the machine [2], [3], [13], [21]- [29], [27], [31], [32], [34], [36]- [38].

B. Objective Facts: The Nature of The Problem

Similar to any other electrical machine, SynRM rotor slot pitch, see αm in Fig. 1, has a big influence on torque ripple. The second most important factors are the constant rotor slot pitch [2], [3], [20], [22], [27], [36] and a rotor with a homogeneous anisotropic structure [2], [3], [13], [41]. In SynRM, the idea of αm = cte. was used in [2], [13], [21], [22], [27], [36] for torque ripple control. Even the Kostko’s design, the first SynRM, somehow considers the constant rotor slot pitch [1]. The idea, described in [22], [27], was based on the selection of a suitable number of barriers, assuming constant rotor slot pitch is used, this takes place, in the most famous cases, when e.g. 2β = αm or β = αm in Fig. 1, and the number of the stator slots is known. The main claim is that the torque ripple is minimized, if there is a relation between the rotor and the stator slot numbers. An intelligent idea that is used for

adapting the constant rotor slot pitch to the SynRM unsymmetrical rotor structure especially in the q-axis region of the rotor is the definition of an imaginary rotor slot opening in each pole of the machine. This imaginary slot opening is placed in the last segment, here S3 area, and one end of the opening is represented by point B in Fig. 1 and Fig. 2.

C. Optimization strategy and methodology

One difficulty that arises in such a condition is that the rotor optimization for minimum torque ripple is inter connected to the stator shape through optimization. The situation gets more complicated if some Permanent Magnet (PM) materials are introduced inside the flux barriers of the rotor, such a machine is known as IPM and/or PM assisted SynRM (PMaSynRM). The PM material also introduces MMF and consequently distortion in the resultant air-gap flux density, then the idea presented in [22], [27] for torque ripple reduction, will be distorted too, and faces new problem.

A possibility for generalizing this method is to keep the rotor slot pitch constant mainly in the d-axis for barriers number 1 to k, see Fig. 1 and Fig. 2, while 2β is allowed to vary (±) for a certain number of barriers (k).

Introducing this method does not change the equal rotor slot pitch rule for the major part of the rotor (d-axis), except in the area near to the q-axis. But, the interlock between torque ripple and the number of barriers is eliminated. The ripple can be minimized for any number of barriers and any stator structure by performing some limited number of FEM sensitivity analysis on angle β , for any machine type such as SynRM, IPM and PMaSynRM. The optimal slot pitch for any number of barriers can be found for torque ripple minimization by adjusting αm with β. Another parameter that can be used to reduce the ripple further is e.g. adjusting the rotor slot pitch by choosing a suitable radial position of the barrier in the q-axis Yq, see Fig. 1. E.g. see the effect of Yq on torque ripple in [20], [47], when the optimal kwq is determined.

IV. FEM DESIGN AND TYPICAL RESULTS ANALYSIS

Assume that the stator geometry is fixed, la + ly = cte. , then for each set of: pole number, barrier number, insulation ratio in the q- and d-axis and angle β values, the rotor microscopic dimensions can be calculated using (2) and (3). For simplicity saturation, stator slotting, iron MMF drop, imperfect winding and MMF effects are disregarded without any loss in the theory’s generality [1]- [3]. Instead, the resultant geometry can be analyzed by FEM.

21

2

k

pm

(4)

As a first step, assume that β is known and as a starting point the constant rotor slot pitch is suggested. Then, based on the number of barriers k and assumed β , the position of end points of barriers in the air-gap, points Di (for i = 1, ..., k),



5

will be as shown in Fig. 2. For this conditions the rotor slot pitch, αm , can be determined using Equation (4).

Consider that Δfi for each barrier and fdi for each segment are calculated, knowing αm and using Fig. 2, then it is straight forward to calculate each barrier and segment size in the q-axis using (2) and (3). First, by distributing the iron, ly

, (magnetic material) between different segments according to (5), note that (ly + la) and kwq are known:

,1

)(

,,...,2

,2

1

1

2

1

2

1

2

1

2

1

wq

ayk

ii

ii

k

lllyS

kifd

fd

S

S

fd

fd

S

S

(5)

second by distributing the insulation, la, (air) between different barriers according to the Equation (6):

.1

1

)(1

,,...,21

1

1

2

11

wq

ayk

ii

ii

k

lllaW

kif

f

W

W

(6)

Each barrier dimension in the d-axis also can be calculated according to Equation (7), if barrier dimension in the q-axis and insulation ratio in the d-axis are known.

kiW

W

W

Wi

axisq

i

axisdd

d ,...,21

1

1 1

(7)

A. Typical Effect of Insulation ratio in the q-axis

The effect of the q-axis insulation ratio on the torque of an specific SynRM machine with 2 barriers, at β =cte., is modeled (FEM) and the result is shown in Fig. 3. Here for each kwq it is assumed that kwd = ½ kwq . The optimum insulation ratio in the q-axis for maximum torque is around 0.6 ~ 0.7, [41].

B. Typical Effect of Insulation ratio in the d-axis

If kwq is set to its optimum (kwq = 0.6), then the effect of kwd on torque is shown in Fig. 4, [41]. Notice that kwd ≤ kwq at the optimal point [20], [47].

C. Typical Effect of Angle β on Torque ripple

The effect of β on a specific SynRM is investigated to demonstrate the torque ripple variation due to this design parameter change [43]. As was shown earlier, a SynRM with 4 barriers rotor structure exhibits a high torque ripple when 2β = αm [20]. The insulation ratio in the d-axis, kwd , and the q-axis, kwq , are kept constant and equal to the optimum values for maximum torque. Optimal kwd is 0.3 and kwq is 0.7 for

0

5

10

15

20

25

30

35

0,1 0,5 0,9 1,3 1,7 2,1kwq, insulation ratio in q-axis @ kwd = 0.5 x kwq

To

rqu

e (N

.m)

Torque

laly

lad

ly

lylakwq

lyladkwd

Fig. 3: Torque (FEM) for different kwq when kwd = ½ kwq and β = cte. [41].

0

5

10

15

20

25

30

35

0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9kwd, insulation ratio in d-axis @ kwq=0.6

To

rqu

e (N

.m)

Torque

laly

lad

ly

ly

lakwqly

ladkwd

Fig. 4: Torque (FEM) for different kwd when kwq = 0.6 and β = cte. [41].

0

6

12

18

24

30

36

-5 -3 -1 1 3 5 7 9 11 13 15 17

Torq

ue

[N.m

]

β [º] @ kwd = 0,3 & kwq = 0,7

Average torque

Torque Ripple pp

Fig. 5: Simulated (FEM) torque and torque ripple (peak-to-peak (pp)) curves for a SynRM with 4-pole, 4 barriers rotor structure, see Fig. 6, as a function of β , when kwd is 0.3 and kwq is 0.7, and end points Di (for i = 1, ..., k) are adjusted by Yq [43].

Fig. 6: Effect of β on the rotor geometry structure when kwd is 0.3 and kwq

is 0.7, see Fig. 5, end points are adjusted by Yq [43].



6

specific machine with 4 barriers structure, here, [20]. The effect of β on torque and torque ripple is shown in

Fig. 5. Actual machine structure variation due to changing β is also demonstrated in Fig. 6. The simulated (FEM) variation of torque versus time for some β cases is also plotted in Fig. 7. The corresponding β , when 2β = αm , is 4.5º mechanical. By increasing β torque ripple (pp) is reduced from 40 % , for β around 4.5º, to 13% , for β around 8º. Angle β has a negligible effect on the average torque see Fig. 5, because the insulation ratios are the same here for all machines [43].

The torque ripple minimization shows that, torque maximization and torque ripple minimization, for a certain number of barriers and stator structure, can be achieved independently by optimizing β with equal rotor slot pitch in the d-axis. The rotor slot pitch in the d-axis is the most effective parameter for the reduction of torque ripple without significantly affecting the average torque, but this is not true for the rotor slot pitch in the q-axis, [43].

V. EXPERIMENTAL VERIFICATION

Measurement and test, for some examples refer to [4], [17], [45]- [53], are the best methods to evaluate the effectiveness of the design procedure of the SynRM that is discussed in this paper. Therefore, a high performance SynRM rotor with three barriers is designed and fine tuned (the design and fine tuning steps description are out of the scope of this paper and it will be presented in the future). This machine has the same stator as its corresponding standard IM with frame size 160. The stator outer diameter is 254 mm, inner diameter is 165 mm and the stack length of the machine is 180 mm.. Then this SynRM rotor is prototyped and its performance is measured and compared with its IM counterpart. The rotor laminations and the prototype of this machine are shown in Fig. 8 (a) and (b) and the test-bench for the measurement is shown in Fig. 8 (c).

A. Heat–Run Test Conditions

The test conditions are similar to the test situations in [17]. However an industrial inverter is used. The inverter control has a built in Direct Torque Control (DTC) that is modified for SynRM control. No speed sensor is used in the test set-up. After thermal steady-state is reached, the most important mechanical and electrical parameters are read from the

instruments. The system performance of the whole drive is measured. Both IM and SynRM are tested with the same set up. Both drive and non-drive end shaft temperatures are measured by a mobile temperature probe immediately after test. The test results on SynRM and its counterpart IM at 1500 rpm is summarized in TABLE 1. The SynRM performance before prototyping and measurements is (analytically + FEM) evaluated and shown in this table (Calc.), as well.

Load Machine SynRM or IM under Test Torque meter

Fig. 8: (a) laminations of the prototyped SynR machine and (b) complete machine. (c) The heat-run test-bench for IM and SynRM.

TABLE 1 HEAT-RUN TEST MEASUREMENTS ON SYNRM AND IM.

Machine Type IM SynRM SynRM

Evaluation Type Measur. Measur. Calc.

Speed [rpm] 1498 1501 1500

f1 [Hz] 51 50 50

Slip [%] 2.099 0 0

V1ph [V rms] 215 205 219

I1 [A rma] 29 31 31

Losses [W] 1280 928 1028

Stator Copper Losses [W] 536 613 611

Friction Losses[W] 86 86 78

Rest Losses [W] 658 229 338

T [Nm] 88.3 87.7 87.2

T / I1 [Nm / A] 0.26 0.23 0.23

Pout [kW] 13.9 13.8 13.7

Pin [kW] 15.1 14.7 14.7

S1in [kVA] 18.6 19.2 20.6

Motor η [%] 91.5 93.7 93.0

PF1 0.82 0.77 0.71

η · PF1 0.746 0.717 0.664

1 / (η · PF1) 1.34 1.39 1.5

Winding Temp. Rise [K] 66 61 61

Housing Temp. Rise [K] 42 36 NA

Shaft on N-side Temp. Rise [K] 32 19 NA

Shaft on D-side Temp. Rise [K] NA 21 NA

Inv. Avr. Switch. Freq. [kHz] 4 4 NA

Inverter Losses [W] 442 442 NA

Inverter η [%] 97.2 97.1 NA

System η [%] 88.9 91.0 NA

15

20

25

30

35

40

45

0 0.001 0.002 0.003 0.004 0.005 0.006 0.007 0.008 0.009 0.01

torq

ue

[N

.m]

Time [s]

β = 0º β = 4.5º β = 8.5º β = 15º

Fig. 7: Simulated (FEM) torque versus time for a SynRM with 4-pole, 4 barriers rotor structure, see Fig. 5, for different β angles.



7

B. Heat–Run Test Results Analysis

Torque capability of the SynRM closely follows the IM, as is expected, see the T/I1 row of TABLE 1. It is notable that the SynRM voltage here is 5 % lower than for the IM, which makes a T/I1 comparison a little inaccurate, since the winding and/or control could be adjusted in order to reach same nominal voltage, which consequently affects current, too. The expected efficiency improvements for the SynRM in comparison to the IM according to [4], [20], [47] for output power 13.5 kW is around 2.9 %-units. The measurements show the corresponding efficiency improvements of 2.2 %-units, which are in good agreement with expectation. The SynRM has lower winding temperature rise than the IM by 5 K. The SynRM machine’s shaft temperature is also lower than the IM by 13 K. Lower temperatures in both windings and shaft can directly increase the lifetime of the SynRM in comparison to the IM. The SynRM housing is cooler than the IM roughly by 10 K. Power factor of the SynRM is lower than the IM by 5 %-units. However, the factor 1/(η·PF), which is the per-unit inverter current, for both machines are very close. Consequently, the required extra kVA for SynRM is just around 0,6 kVA. The measured inverter performance of the SynRM and IM drive, see TABLE 1, clearly shows that the machine type does not affect the inverter. The inverter efficiency is the same for both machines with the same output power and speed. Consequently, better SynRM efficiency directly increases the overall system efficiency. The improvement of the SynRM based drives system efficiency in comparison to the IM drive for output power of 13.5 kW is around 2.1 %-units.

VI. CONCLUSION

A novel, simple, fast and systematic design procedure for SynRM is presented. This procedure, combines analytical methods and FEM simulations for optimizing the rotor of SynRM for an specific stator structure. Torque and torque ripple of the SynRM can be maximized and minimized respectively and independently, by running some limited number of FEM sensitivity analysis on three design parameters. The insulation ratios in d- and q-axis of the machine for torque maximization and the rotor slot pitch in the d-axis for torque ripple minimization are introduced and used by this procedure.

The performance of a SynRM, which has been designed using this procedure, is measured and compared with the performance of the IM, with the same standard frame size of 160, for validation of this method.

ACKNOWLEDGMENT

Authors are thankful to C. Sadarangani, H. Lendenmann, I. Stroka and R. Kanchan from ABB, Corporate Research, Sweden, for their valuable help and support.

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Reza Rajabi Moghaddam received the B.Sc. degree in electrical power engineering from Sharif University of Technology, Tehran, Iran, in 1997 and the M.Sc. and Ph.D. degrees in electrical engineering from the Royal Institute of Technology (KTH), Stockholm, Sweden, in 2007 and 2011, respectively. Since 2006, he has been with ABB Corporate Research, Västerås, Sweden, as a Research Engineer, Scientist and technology development project's leader in various areas. His interests include

electrical machines and drives with electrical machine design orientation.

Freddy Gyllensten received the M.Sc. degree in electrical engineering from Chalmers University of Technology, Gothenburg, Sweden, in 1994 and the Ph.D. degree in electrical engineering from the Royal Institute of Technology, Stockholm, Sweden, in 2004. Since 1996 till 2010, he has been with ABB Corporate Research, Västerås, Sweden, as a Research Engineer and Project Manager in various development projects. Now he is with ABB Motors and Generators, Västerås, Sweden, as R&D

Manager, since 2010. His research interests include electrical machines and drives.

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