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Design considerations of a multitasked electric machine for automotive applications
Electric Power Systems Research 131 (2016) 147–158
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Electric Power Systems Research journal homepage: www.elsevier.com/locate/epsr
Design considerations of a multitasked electric machine for automotive applications Erkan Mese a , Murat Ayaz b,∗ , M. Murat Tezcan a a b
Yildiz Technical University, Department of Electrical Engineering, Istanbul, Turkey Kocaeli University, Department of Electrical Education, Kocaeli 41050, Turkey
a r t i c l e
i n f o
Article history: Received 27 September 2014 Received in revised form 11 October 2015 Accepted 13 October 2015 Keywords: Permanent magnet synchronous machine (PMSM) Concentrated winding Magnetic decoupling Dual winding PM machine (DWPMSM) Fault tolerant operation Hybrid electrical vehicle applications
a b s t r a c t This paper introduces a new electric machine for Front-End Accessory Drive (FEAD) of Hybrid Electric Vehicles (HEVs). The novelty of the electric machine lies in its ability to perform multitasking. Besides the independent motor or independent generator operation like any other electric machine, the proposed machine is able to run as motor and generator simultaneously. A dual winding electric machine forms the core part of the proposed system. Windings of the electric machine are concentrated type so that electrical and magnetic isolation is maintained. This allows simultaneous motoring and generating operation in the single housing of electric machine. Design considerations of the proposed electric machine have been outlined in the paper. Comparison between proposed and conventional approaches has been made to highlight potential benefits of the new approach. Thermal analyses have been performed to show the suitability and limits of the multitasked electric machine in the HEV applications. Overall performance of the proposed machine is presented by experimental results. © 2015 Elsevier B.V. All rights reserved.
1. Introduction Electric machines designed with dual winding sets in their stator have recently become popular due to providing opportunities to solve many industrial problems. For instance, continuous operation of the mechanical and the electrical accessories in HEVs with stop/start functionality can be achieved by using such a dual winding electric machine which has simultaneously motor and generator operation capability [1,2]. The goal of this study is to present design considerations on a decoupled dual winding permanent magnet synchronous machine (DWPMSM) for continuous operation of accessory loads such as air conditioner compressor, power steering pump, water pump and etc. The proposed machine is able to run as motor and generator simultaneously without interaction between these operations. Concentrated winding type is used in the proposed machine to achieve decoupling between winding sets. Since each concentrated coil around any stator tooth has an independent magnetic circuit, magnetic decoupling between winding sets is possible [3–7]. Other advantages are shorter end windings, higher slot fill factor,
∗ Corresponding author. Tel.: +90 2623032307; fax: +90 2623032307. E-mail addresses: [email protected] (E. Mese), [email protected] (M. Ayaz), [email protected] (M.M. Tezcan). http://dx.doi.org/10.1016/j.epsr.2015.10.017 0378-7796/© 2015 Elsevier B.V. All rights reserved.
wider speed range of flux weakening, easier manufacturability and higher efficiency. On the other side, there are some deficiencies related to concentrated winding like noise, torque ripple, unbalanced magnetic forces due to additional harmonic contents [8–15]. Furthermore, MMF harmonic spectrum of concentrated winding machines contains both super-synchronous and sub-synchronous components. Whereas a balanced distributed winding has only super synchronous harmonic content in its MMF. Sub-synchronous space harmonics would cause higher rotor core and magnet losses because of relatively higher frequency [16–18]. Furthermore, subsynchronous space harmonics have relatively greater wavelength and this causes deeper penetration of the flux to the core and further increase in losses [17]. So, electric machines with concentrated winding would have higher core and magnet losses at high speed region. In particular, the losses become more significant in highpole count machines. This drawback of the concentrated winding has been extensively elaborated in the paper. 2. Material and methods of accessories drive systems in HEVs In hybrid electric vehicle (HEV), nonstop accessory load-driving feature is implemented with drive-by-wire concept as shown in Fig. 1(a). With this concept, each of the accessory loads is powered through its own electric motor and drive system fed by high voltage
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Fig. 1. (a) Drive by wire system, (b) proposed electrical accessory drive system.
battery. Drive by wire is an expensive method since many motors and inverters must be used. Furthermore it is an inefficient method since fractional power motor drive systems tend to be less efficient than high power motor drive systems. In addition, duty cycle of single motor is more likely filled with higher load intervals compared to the distributed multi-motor system. On the other hand, low voltage (12 V) power is provided through a DC/DC converter as shown in Fig. 1(a). Lundell alternator can also be used for 12 V power. The low efficiency of this alternator, which varies between 45% and 60%, depends on its speed and load level, is the main drawback preventing from the future vehicle applications [19]. Electrical accessory drive system with the proposed machine for driving accessory loads and charging 12 V battery in HEVs is shown in Fig. 1(b). As seen from the figure, the proposed dual winding machine runs as motor and generator simultaneously or just operates as a generator. If the engine is running the dual winding machine operates only as generator. Otherwise, simultaneous motor and generator operation takes place. With the decoupled dual winding sets, mechanical and electrical power can be generated in a single electric machine instead of two separate machines. By using electrical accessory drive system with the proposed machine several small motors and their drivers in drive by wire system are removed. Furthermore, several inherited components of conventional accessory drives of conventional ICE based vehicles (such as pumps and compressors) can be still used. These facts are considered as cost saving advantages by some automotive manufacturers. Table 1 summarizes components of the conventional and proposed accessory drive system. The high voltage DC/DC converter or Lundell alternator used for charging low voltage battery (12 V) in drive by wire system is eliminated with
the generator side of DWPMSM. On the other hand, a DC/DC converter with an uncontrolled rectifier or a controlled rectifier is to be used to regulate output voltage of the generator side of DWPMSM. However, these converter structures are more efficient and cheaper solution compared with the high voltage DC/DC converter due to lower input voltage. The output power of the DWPMSM is to be determined depending on power requirement for accessories in vehicles as shown in Fig. 2. As can be seen from the figure, DWPMSM should process around 6 kW power at 1500 rpm to drive all accessories. Power flow of the proposed accessory drive system and operation modes are detailed in Fig. 3. There are two operation modes of DWPMSM depending on mechanical power generation of internal combustion engine (ICE). Table 1 Part count comparison among different accessory drive systems. Parts
Conventional non-hybrid
Drive by wire
EADS
Crank pulley Pulley clutch Belt Water pump Alternator Air pump Air pump eMotor Power steering pump EPS pump EPS eMotor AC compressor AC comp. eMotor Auxiliary power unit (DC/DC converter) eMotor + Inverter
X
X
X X X X
X X
X X X X X X
Unique X
X
X
X Unique X Unique X X
X
X
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149
Fig. 2. Power requirements for mechanical and electrical accessories in HEVs.
In the first operation mode, ICE is off and power requirement for mechanical accessories should be provided by motor side of DWPMSM. The motor windings are driven by an inverter to generate 6 kW at 1500 rpm. Half of the generated power by the motor windings is consumed by mechanical accessories connected to shaft of DWPMSM. The remaining half of the power is used to supply electrical accessories through generator windings as illustrated in Fig. 3(a). In the second operation mode ICE is on and no motoring is required since mechanical power is generated by ICE as shown in Fig. 3(b). Therefore, the motor windings are unexcited and the generator windings generate almost 3 kW between 1500 rpm and 6000 rpm shaft speed to supply low voltage system in the vehicle.
Mechanical Accessories
3.1. Determining geometrical dimensions of DWPMSM Magnetic coupling between winding sets is to be around zero to achieve a multifunction operation in the single machine housing. Single layer concentrated winding topologies has better decoupling performance compared to double layer concentrated winding [14]. Also, surface mounted PM (SPM) structure allows to increase decoupling in the proposed machine. Besides decoupling performance of the SPMs, lower unbalanced radial forces, reduced stack
Inverter
Internal Combustion Engine
+ - 300 V
0 rpm
Comp. Pump Steering Pump
3. Calculation process of dual winding PMSM design parameters
Motor Terminals 6 kW
High Voltage Battery
DWPMSM 3 kW
3 kW
3 kW
Etc.
1500 rpm Generator Terminals
Mechanical Power Uncontrolled Rectifier
Electrical Power
DC-DC Electrical Chopper Accessories
a)
Mechanical Accessories Comp.
1000 ~ 6000 rpm
Pump Steering Pump
Inverter
Internal Combustion Engine
3 kW
3 kW
+ - 300 V Motor Terminals 0 kW 3 kW
Etc.
High Voltage Battery
DWPMSM 1000 ~ 6000 rpm Generator Terminals
Mechanical Power Uncontrolled Rectifier
Electrical Power
Electrical DC-DC Chopper Accessories
b) Fig. 3. Operating modes of DWPMSM, (a) first mode (motor/generator operation), (b) second mode (generator only operation).
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length and lower stator MMF harmonics are main advantages compared with interior PM machines. As a result of these aspects, surface mounted PM machine with single layer concentrated winding would be the first design decision in the design process. After deciding surface mounted PM machine, geometric envelope of the machine is determined. This includes roughly finding rotor diameter and axial length. Analytical method to design a PMSM is well documented in [20]. Similar method has been adopted here. Torque equation of a radial flux electric machine is given as T = kDr2 Ls
Stator Outer Diameter
Magnet Height Airgap
Tooth Widht Magnet Width
(1)
In Eq. (1) Dr is the rotor outer diameter, Ls is axial length of the stator core. Torque coefficient k which is related with the thermal loading and cooling technique of machine. In the proposed electric machine, the worst case for the copper losses occurs during motor and generator operation at 1500 rpm speed. On the other hand, the worst case for the core losses occurs during high speed operation (above 5000 rpm). In this mode, machine operates only as generator as described in Section 2. Although loss number is greater in high speed generator only mode, the duration is shorter. Hence simultaneous motor and generator operation at 1500 rpm reflects more severe thermal condition and resembles conventional electric machine where all stator coils carry equal current. By considering thermal loading cases stated above, torque coefficient is assumed to be 14,400 N m/m3 for air cooled motor. This value is adopted from [20] where it is stated that the torque coefficient can be selected between 14,000 and 42,000 N m/m3 for totally enclosed and sintered rare earth PM motor. Since the proposed machine is naturally cooled lower limit of this range is preferred. After narrowing the design space by sizing equation, in order to generate 6 kW shaft power at 1500 rpm shaft speed, detailed design steps have been taken iteratively by an analytical design tool running in the computer. Geometrical dimensions of the PMSM are shown in Fig. 4. Air gap length, stator or rotor skew angle, magnet height and slot span have significant effects on the motor performance. This section presents these effects by parametric analysis made by Ansoft–Rmxprt software. Initial design values of air gap length, stator skew angle, slot span, magnet height and width are 0.5 mm, 0◦ , 4 mm, 5 mm and 12.8 mm, respectively. Fig. 5(a) shows the results of parametric analysis made to minimize torque ripple. By analysis, torque ripple is reduced from 7.24% to 2.5%. Torque ripple minimization effort during the design stage aims to find optimum geometry for the minimum torque ripple and ignores harmonics
Yoke Height
Rotor Inner Diameter
Slot Opening
Rotor Outer Diameter Stator Inner Diamater
Fig. 4. General description of slot and tooth geometry.
in the phase current. Hence the results shown in Fig. 5(a) assume pure sinusoidal current in both winding of DWPMSM. Similar parametric analysis is performed for back EMF harmonic distortion and cogging torque. The respective results are shown in Fig. 5(b) and (c). The geometrical and electrical design parameters of DWPMSM are given in Table 2. This table summarizes the details of the DWPMSM design and they are prepared by parametric analysis. 3.2. Winding structure of DWPMSM In the literature, there are many studies about PMSM winding topologies. Beside many advantage stated before, concentrated winding topology has minimum magnetic coupling between the coils [14,21]. An important requirement in the proposed FEAD system is magnetic coupling between two winding sets must be minimum or preferably zero for proper operation. In concentrated winding, each coil has its own independent magnetic circuit as shown in Fig. 6 where short flux paths are dominating the magnetic circuit. Long flux paths also exist but they are minor. Dominance of short flux path can be interpreted as very low magnetic coupling between neighbor coils. This enables simultaneous motor and generator operation. Although magnetic independence, each coil still contributes to the total air gap flux density. Finite element analysis (FEA) is performed with Maxwell 2D to examine existence of magnetic coupling between two winding
Table 2 Electrical and physical design parameters of DWPMSM. Electrical parameters
Physical parameters Gen.
Number of turns
5
Wire diameter (mm) Rs-stator phase resistance () Lq-inductance (mH) Ld-inductance (mH) Flux linkage by magnets (V s) Stator tooth flux density (T) Airgap flux density (T) Rotor yoke flux density (T) Rated torque (N m) Output power (W) Back EMF constant (V/rad/s)
2.59 0.0015 0.043 0.043 0.0125 1.25 0.82 0.27 21.23 3000 0.285
Motor 21
2.3 0.033 0.776 0.776 0.055 1.44 0.82 0.27 38.21 6000 1.2
Stator parameters
Stator slot parameters
Stator outer Dia. (mm)
240
Hs0 (mm)
2.8
Stator inner Dia. (mm) Stator core length (mm) Stator skew angle (◦ )
140 120 11.1
Rotor Parameters Rotor outer Dia. (mm) Rotor inner Dia. (mm) Rotor core length (mm) Magnet width (mm) Magnet height (mm) Inertia (kg/m2 )
138 75 120 16.14 5.0 0.033
Hs1 (mm) Hs2 (mm) Bs0 (mm) Bs1 (mm) Bs2 (mm) Rs (mm)
0 28 5.5 8.8 14.4 4.5
Net slot area (mm2 ) Slot fill factor (%) Number of slots Number of poles
380.9 56.01 24 22
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Torque Ripple (%) 10
THD (%) 4
11.1º %4.678
0
0
Skew (º) 5
10
11.1º %0.68
2
5 0
Skew (º)
0
15
5
10
1 mm %4.227
5 0 0
2
1
10
Airgap (mm)
0 0
Slot Opening 8 (mm)
0
2
4
1
2
Airgap (mm)
1
5.5 mm %2.418
5
15
1 mm %0.68
1
10
151
6
5.5 mm %0.68
0 2
4
Slot Opening 8 (mm)
6
4
10
16.74 mm %2.418
5
0 8
10
12
14
5
16
18
0 8
Magnet Height 8 (mm)
0 4
5
10
12
14
16
18
Magnet Width 20 (mm)
1
5 mm %2.418
3
16.74 mm %0.68
2 Magnet Width 20 (mm)
6
7
5 mm %0.68
0
3
4
5
6
a)
7
Magnet Height 8 (mm)
b) Cogging Torque (mNm) 24 16 8 0 0 5 3 2 1 0
11.1º 0.55 mNm Skew (º)
10
15
1 mm 0.55 mNm 0
1
6 4 2 0
2 5.5 mm 0.55 mNm
2
4
Slot Opening 8 (mm)
6
3 2 1 0
16.74 mm 0.55 mNm 8
1
10
12
14
Airgap (mm)
16
18
Magnet Width 20 (mm)
5 mm 0.55 mNm
0
3
4
5
6
7
Magnet Height 8 (mm)
c) Fig. 5. DWPMSM, (a) torque ripple performance, (b) THD% of induced back EMF performance, (c) cogging torque performance.
sets. In Fig. 7, the effect of generator side loading on the motor side torque output is shown, where generator is connected to an uncontrolled rectifier. In conventional distributed type dual winding electric machine, even if a severe interaction between fundamental components of generator and motor magnetic fields is not observed, higher order harmonics of generator magnetic field adversely affect motor magnetic field. In concentrated winding sets, generator side loading has weaker adverse effect on motor torque as shown in Fig. 7. In particular, single layer concentrated winding
approach seems to be most effective approach in magnetic decoupling of two winding sets. 3.3. Analytical model of field distribution and decoupling between winding sets This section shows an analytical model of the magnetic decoupling between two winding sets in a single layer concentrated winding configuration. For this purpose, the magnetic field
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Fig. 6. (a) Concentrated single layer dual winding electric machine, (b) flux distribution of single layer coils.
40 30
Igen 0A
20
Igen 55 A
Igen 160 A
Igen 110 A
Distributed Winding Double Layer Con. Winding Single Layer Con. Winding
10 0
0
x = ln (r), y = , ∂z/∂s = 1/s ⇒ s = ez . In the second stage, the corner points of the Z Cartesian coordinate system are determined. The current in a coil directly affects the magnetic vector potential in the middle of the air gap and the Schwarz Christoffel transformation is used to transform the Z complex plane to a W complex plane.
Torque (Nm) 50
10
20
30
s = r cos s + jr sin s
(2)
z = log (s)
(3)
2g z = −j
Time
1 √ arctan a
Fig. 7. Motor torque variation for different generator winding load levels (n = 1500 rpm).
distribution model of a slot current is used. The magnetic field of the slot current is modeled by conformal transformation, as described in [22]. Four conformal transformations are used to perform the analytical method. Actual geometry in S complex plane, which has coordinates in the form of s = m + jn = rej , is shown in Fig. 8(a). The Z Cartesian complex coordinate system has coordinates in the form of z = x + iy shown in Fig. 8(b). Logarithmic conformal transformation is used to calculate the polar coordinates
1 + w − a/w + 1 1 + ln 2 1 − w − a/w + 1
θ2 θs/2 W=a θ1
4
n
α
Rotor y Region
3
Rr
a)
W=-1
2 Ln(Rs)
x
Rs 1
bo’ W=0
β g’
1 ln(Rr)
θs
m
+ ln (Rr) + j
g’=ln(Rs/Rr) bo’=θ2-θ1 α=π/2 β=3π/2
g
θ1 θ2
2
s 2
(4)
Eqs. (2)–(4) can be used to transform an arbitrary point in the actual geometry in the middle of air gap in the S complex plane to a complex Cartesian W complex plane having coordinates in the form of w = u + iv, where the solution exists. Fig. 9(a) shows the new coordinates of the geometry after Schwarz Christoffel transformation in the W complex plane. Solving Eqs. (3) and (4) simultaneously yields several w = u + iv values. These values are used for determining t = p + iq values in a slotless T plane, as shown in Fig. 9(b), using Eqs. (5) and (6).
5 6
w − a/w + 1 √ a
W=∞ W=-∞
b)
Fig. 8. (a) Slot opening in the S plane, (b) Half of the slot opening in Z plane.
Stator Region
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153
W=-∞ W=∞
6
θs
5 γ’ γ’=ln(Rs/Rr)
W=-1 θs/2
v
W=1
q 1 ln(Rr) Z=ln(Rr)-j∞
Z=ln(Rr)+j∞
-∞
Z=ln(Rs)+jθ1
0
Z=∞+jθ1
1
a
Z=ln(Rs)+j∞ u
Z=ln(Rs)+jθ2
p
∞
b
2 Ln(Rs)
W=0
a)
b)
Fig. 9. (a) One half of the slot opening in the W complex plane, (b) Slot opening in the T complex plane.
Bar_air(T)
Bar_air (T) 0.0002
0.0002
θ (Mech.)
θ (Mech.) π/48
π/24
π/16
π/12
-0.0002
π/8
5π/48
π/6
5π/24
-0.0002 Slot Center
Slot Center
a)
b)
Fig. 10. Analytical solution, (a) the left coil side radial flux density distribution (b) the right coil side radial flux density distribution.
The field distribution in the S complex plane is given by Eq. (7). One coil side field solution in Cartesian coordinates can be found by using Eq. (8), where the Cartesian coordinate field solution is transformed into the cylindrical coordinate system [23]. Bar is the radial component of the winding flux density distribution in the air gap.
∂t g 1 g s ⇒ t = j ln (w) + ln (Rs) + j =j w 2 ∂w Bt = −jo
∂ϕ Nc = jo I 2 ∂q
Nc 1 Bs = Bm + jBn = jo I 2 s Bar = Bm cos + Bn sin
(5) (6)
w−a w+1
magnetic field distributions yields a new flux density distribution, shown in Fig. 11. The plot shows that no magnetic field exists outside the coil region, suggesting that the neighbouring coils will not be affected by the operation of this coil [24].
Bar_air(T)
0.0002
∗
θ (Mech.)
(7) (8)
The field distributions of current in two adjacent slots are modelled by the procedure above, for a machine geometry with 24 slots, 22 poles and a single layer concentrated winding configuration. Coil turn number Nc (Nc = 10 turns) and coil current I = 10 A. The flux density distribution of the left and right sides of the coil are shown in Fig. 10(a) and (b), respectively. Superposition of the two
π/24 -0.0002
Magnetic Decoupling Zone
π/12
π/8
π/6 Magnetic Decoupling Zone
-0.0004 Slot Center
Slot Center
Fig. 11. Analytical solution of a single layer concentrated winding radial flux density distribution and magnetic decoupling zones.
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Table 3 Electrical and physical parameters comparation of different PMSMs. Electrical parameters
Machine type
Physical parameters
Machine type
Rotor data
Conven. motor
DWPMS (MG)
Conven. gen.
Thickness of mag. (mm) Type of magnet Total net weight(kg) Stator data Number of slots Outer Dia. (mm) Inner Dia. (mm) Total length (mm)* Len. of core (mm) Type of steel Slot area (mm2 ) Slot fill factor (%)
5 NdFe30 19.16
5 NdFe30 32.68
5 NdFe30 20.17
24 215 123 111.12 89.5 M1924G 357.7 52.72
24 240 140 144.96 120 M1924G 380.9 56.01
24 218 122 112.19 90 M1924G 361.39 61.02
DWPMSM No load data
Conven. motor
St. teeth flux density St. yoke flux density THD of Ind. voltage (%) Full load data Max. line voltage(V) RMS Ph. current (A) Efficiency (%) Power angle (◦ ) Torque angle (◦ ) Iron-core loss (W) Arm copper loss (W) Total loss (W) *
Motor oper.
1.27 0.46 –
1.25 0.44 –
155.75 34.83 93.54 – 25.73 116.48 143.21 413.96
156.11 34.93 92.85 – 26 190.98 116.01 461.53
Gen. oper. 1.44 0.43 0.68 36.65 137.28 90.48 40.98 – 231.19 84.527 315.72
Conven. gen. 1.42 0.42 0.16 36.53 134.25 92.34 40.01 – 147.7 101.05 248.75
Includes end winding.
3.4. Physical comparison with conventional approaches The proposed machine is compared with its alternative where two separate machines are used. To perform an accurate comparison between machines, electrical and magnetic design parameters of all machines must be equal to the proposed machine. These parameters are phase current, current density, fundamental back EMF, stator teeth and yoke flux density, magnet type and etc. as given in Table 3. Calculating the volumes and masses of the machines from their given geometrical parameters in Table 3, DWPMSM’s total volume is 6558.17 cm3 whereas single motor and single generator’s total volume is 8222.14 cm3 , respectively. Similarly DWPMSM has lower mass compared with the total weight of classical motor and generator. When total volume and weight numbers are taken into the account, dual winding structure has significant advantage. In addition all of above advantages, some volume, weight and cost savings also arise when we consider reduced number of position sensor, end bells and bearings. Reduction in these auxiliary but essential elements of the machine yields significant amount of space saving. 3.5. Thermal analysis of DWPMSM This section presents 2D numerical analysis and experimental results of DWPMSM’s thermal behavior by using finite-element analysis (FEA). Thermal analysis is done for 9 kW power ratings (6 kW for motor operation and 3 kW for generator operation). Totally enclosed and non ventilated thermal conditions are
Power Losses (W)
600 500
200
Stator Core Losses Rotor Core Losses Magnet Losses Copper Losses
Temperature (°C)
700
400 300 200 100 0 1000
2000
3000 4000 Speed (rpm)
a)
considered. That means a combination of natural convection, radiation and conduction occurs. The inside and outside fluid is considered as only air and hence named as air-cooled DWSMPM. The total loss versus speed for the DWPMSM is given in Fig. 12(a). As seen from the figure, the core losses and eddy current losses increase with increasing shaft speed as expected. Temperature versus speed variation obtained by the thermal analysis using these loss values is shown in Fig. 12(b). Fig. 13 shows the experimental and simulation thermal photographs during generator operation of the DWPMSM. Loss estimation shows that core and magnet losses have tendency to increase with speed. The reason for this trend can be explained by winding configuration. As stated in Section 1, concentrated winding causes sub-synchronous harmonics and this increases core and magnet losses. The influence of this can also be seen in the efficiency plot at high speed region. More serious influence of high losses can be seen after thermal analysis of the machine. As can be seen from Fig. 12(b), both core losses and magnet losses are causing very high temperature rise. This analysis shows that continuous operation at high speed particularly above 5000 rpm may be dangerous for the magnets. Continuous operation beyond 5000 rpm is not required for intended application of the proposed machine. Instead, intermittent and short term loading of the generator above 5000 rpm is a common loading scenario for the driving cycles under considerations. Hence with careful magnet grade selection and with due care for the operating speed of the generator, magnet temperature rise can be managed. For more demanding applications and cases, water cooling becomes mandatory.
5000
6000
150
Winding End Magnet Stator Tooth Stator Yoke
100 50 0 1000
2000
3000 4000 Speed (rpm)
b)
Fig. 12. (a) Power losses versus speed of the DWPMSM, (b) Temperature versus speed of the DWPMSM.
5000
6000
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155
Fig. 13. Thermal photographs during generator operation at 1500 rpm and fully loaded (a) experimental, (b) simulation.
4. Experimental results and discussion Performance evaluation of the proposed machine is performed by using experimental setup shown in Fig. 14. In experimental test bed, a squirrel cage induction motor is controlled by a four-quadrant adjustable speed driver that enables motoring and generating operation. Also, a rotary torque sensor between the induction motor and dual winding PMSM measures torque level dynamically. Induction machine in the experimental setup represents mechanical load in accessory drive system as well as internal combustion engine of the vehicle. In Mode 1, induction machine imitates mechanical accessory loads and runs as generator. In Mode 2, it emulates internal combustion engine and runs as motor. A resistive load bank with adjustable 3 kW power capabilities is used for emulating electrical accessory loads. An uncontrolled rectifier/DC–DC converter set is connected between generator windings and the resistive load bank in order to regulate output voltage. 4.1. Study of magnetic coupling on DWPMSM This test is performed to verify FEA results in terms of the magnetic decoupling between winding sets of the proposed electric machine. In the test, motor back EMF waveforms are checked to see the effect of generator winding load level (loaded or unloaded). The expectation is to see no difference between two back EMF waveforms. Generator windings are connected to a three phase bridge uncontrolled rectifier which has an output connected parallel to
Fig. 15. Motor Back EMF waveform with no load (yellow trace), Motor Back EMF waveform with load (white trace), Generator phase voltage waveform with load (pink trace), Generator Output current waveform (green trace) at 1500 rpm rotor shaft speed. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
a 450 VDC–2200 F capacitor. Capacitor terminals are connected to the 3 kW resistive load. Firstly, the proposed electric machine is driven at constant speed (1500 rpm) by induction motor. The uncontrolled rectifier terminals are open circuit, so the generator windings are unloaded and the motor windings’ back EMF waveform is recorded as shown in Fig. 15 (white trace). After that,
Fig. 14. Experimental setup for DWPMSM.
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Fig. 16. Operation modes of the proposed DWPMSM.
4.2. Operation modes of DWPMSM machine It has been shown in previous subsection that there is no magnetic coupling between motor and generator winding sets. Consequently, operation with only one winding set (only motor or only generator operation) or simultaneous operation of two winding sets (motor–generator operation at the same time) would be possible without deteriorating any machine performance. Operation modes of DWPMSM are shown in Fig. 16. 4.2.1. Independent motor operation for DWPMSM machine Performance tests of the DWPMSM motor operation are implemented by different mechanical load conditions. Motor only operation is performed at 1500 rpm constant speed which is almost the same speed in real accessory drive application. Motor windings are driven by a 600 V–50 A voltage source inverter. Modulation strategy of the inverter is sinusoidal PWM. Carrier frequency for the control signals is 10 kHz. The motor windings are controlled with q axis current value by the inverter because flux weakening is not needed at this speed. In Fig. 17, shaft torque waveform recorded by torque sensor and electrical terminal variables for motor operation are shown. The generator windings are open circuit in this operation mode.
Fig. 17. Motor operation mode (22.82 N m load torque).
100
Efficiency (%)
1300 W resistive load is connected to uncontrolled rectifier terminals without changing the shaft speed and motor windings back EMF waveform is shown in Fig. 15 (yellow trace). As seen from the figure, the motor windings’ back EMF waveforms are not affected by generator winding in the face of load change. As a result of the experimental test, magnetic decoupling between winding sets is assured with the proposed machine.
75 50 25
0 1000
1500 rpm
2000 3000 4000 5000 6000 Motor Output Power (W)
Fig. 18. Experimental efficiency data from independent motor operation.
The obtained machine efficiency during experimental tests is shown in Fig. 18. As seen from the figure, the motor operation efficiency of the DWPMSM varies between 88% and 93% depending on output power values. 4.2.2. Independent generator operation for DWPMSM machine For this operation, DWPMSM is driven by the squirrel cage induction motor. Motor windings are not excited and generator windings are connected to resistive load via an uncontrolled rectifier and DC/DC converter combination. DC/DC converter regulates output voltage at 13.7 V. For different loads and different shaft speeds, output data of generator winding is recorded. Output waveforms of generator operation for 2500 rpm, 100 A at 13.7 V DC are shown in Fig. 19. Independent generator efficiency is obtained experimentally. Fig. 20(a) shows the generator operation efficiency of the DWPMSM at different shaft speed and load levels. As seen from the figure, efficiency is maximized at 1500 rpm with 92% peak value.
Fig. 19. Generator operation mode (2500 rpm–13.7 V–100 A DC load).
100
100
75
75
1500 rpm 2000 rpm 2500 rpm 3000 rpm 3500 rpm 4000 rpm 4500 rpm 5000 rpm
50
25 0 200
800
1400
2000
2600
Efficiency (%)
Efficiency (%)
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50 25 0 1000
3200
157
DWPMSM LUNDELL
2000
3000 4000 Speed (rpm)
Generator Output Power (W)
a)
5000
6000
b)
40
80
30
60
40
20 10 0
Time (ms) 5
6
7
-10
8
Voltage (V)
Torque (Nm)
Fig. 20. (a) Experimental efficiency data from independent generator operation, (b) Comparison of alternators efficiency performance.
0 -20
Time (ms) 0
1
2
3
60
20
40
10 Time (ms) 1
2
3
7
8
b)
30
0
6
Experimental Simulation
-80
4
5
6
7
8
-10 -20
Experimental Simulation
-30
Current (A)
Current (A)
5
-60
a)
0
4
-40
Experimental Simulation
-20
20
20 0
Time (ms) 0
1
2
3
4
5
6
7
8
-20
-40
Experimental Simulation
-60
c)
d)
Fig. 21. Simultaneously motor–generator operation mode at 1500 rpm/13.7 V–40 A DC load (a) torque waveform, (b) generator phase voltage waveform, (c) generator phase current waveform, (d) motor winding phase current waveform.
A comparative plot is given in Fig. 20(b) showing efficiency profile for the generator operation of DWPMSM and Lundell alternator. As seen from the figure Lundell alternator efficiency varies between 45 and 60% [19]. However, generator efficiency of DWPMSM varies between 80 and 93%. This is not a surprising result because similar numbers are reported in the literature for conventional PMSM and the proposed machine’s generator performance seems to be in line with these reporting [19].
4.2.3. Experimental simultaneous motor/generator operation In this section, experimental performance of simultaneous motor/generator operation of the DWPMSM is handled. The mechanical output power required for system is only generated by the motor windings of DWPMSM driven by its own inverter. The induction machine representing the mechanical accessories in the system operates as generator with four-quadrant adjustable speed driver and consumes half of the mechanical power. Remaining
Table 4 Output performance data of simultaneous motor–generator operation at 1500 rpm. Motor–generator operation at 1500 rpm
Motor Q axis current 20
Shaft torque (N m) Motor input power (W) Mechanical out. power (W) Generator out. power (W) Total out. power (W) Machine efficiency (%)
9.54 2848.5 1497.8 1098.8 2596.6 91.1
30 4.7 2912.5 737.90 2033.2 2771.1 95.14
18.2 4368.1 2857.4 1136.6 3994.0 96.8
40 13.56 4459.4 2128.2 2046.0 4174.2 93.6
14.7 5812.8 3611.0 1964.0 5575.0 95.5
23 5832.4 2307.9 3152.4 5356.8 93.9
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mechanical power is transferred to supply electrical loads by the generator windings. Fig. 21 shows the related data of the operating point where the motor speed is set 1500 rpm and generator DC/DC converter is regulating at 13.7 V and 40 A DC. In Fig. 21(a) and (b), high frequency spikes are related PWM switching and caught by the measurement probes during tests. The experimental efficiency of DWPMSM during simultaneous motor/generator operation is given in Table 4. The efficiency of DWPMSM varies between 91% and 97% depending on sharing of mechanical and electrical power generation in the machine. The experimental efficiency values show that the proposed electric machine has similar performance compared with conventional PM machine. Furthermore, experimental result verifies simultaneous motor and generator action in single stator core. 5. Conclusions In this paper, a new electric machine is proposed for some automotive applications where multi-tasking is requested in an effort to develop more compact components. The proposed machine, referred as Dual Winding PMSM (DWPMSM), combines motor and generator functions in the same structure so that it acts as motor and generator simultaneously. After presenting application considerations of the DWPMSM, design procedure is given with emphasis on efficiency, torque density, cogging torque and torque ripple. Winding is the most important aspect of the DWPMSM development and due attention should be paid for proper operation of the machine. The study covered critical steps in the winding design such as location, mutual decoupling assurance and etc. Analytical and experimental studies have been presented in this regard. Quantitative study is presented to compare volumes and masses between conventional and proposed technology. Thermal analyses are also performed to show the limits of the proposed DWPMSM. Exhaustive experimental studies are performed to cover all operating modes of the DWPMSM in a HEV Front End Accessory Drive system. Overall results presented in this paper show that the proposed machine could be a viable solution for a HEV Front End Accessory Drive system. Further application possibilities may exist for the proposed machine in the automotive electrification. Acknowledgments This work was supported by the Scientific and Technological Research Council of Turkey under contract number 110E111. Also Joel M. Maguire from BorgWarner Corporation and Gary E. McGee from General Motors Corporation are appreciated for their inspiration and support during problem definition of accessory systems in hybrid electric vehicles.
References [1] E. Mes¸e, Y. Yas¸a, H. Akc¸a, M.G. Aydeniz, M. Garip, A new electric accessory drives system for hybrid electric vehicles, in: IEEE 2012 Energy Conversion Congress and Exposition, Raleigh, NC, 2012, pp. 1909–1916. [2] E. Mes¸e, M. Ayaz, M. Tezcan, K. Yılmaz, Design of dual winding permanent magnet synchronous machines for hybrid electric vehicle accessory drives, Int. J. Vehicle Des. 69 (1–4) (2015). [3] G. Jack, B.C. Mecrow, J.A. Haylock, A comparative study of permanent magnet and switched reluctance motors for high-performance fault–tolerant applications, IEEE Trans. Ind. Appl. 32 (1996) 889–895. [4] R.F. Schiferl, T.A. Lipo, Power capability of salient pole permanent magnet synchronous motors in variable speed drive applications, IEEE Trans. Ind. Appl. 26 (1990) 115–123. [5] A.M. EL-Refaie, T.M. Jahns, Impact of winding layer number and magnet type on synchronous surface PM machines designed for wide constant-power speed range operation, IEEE Trans. Energy Convers. 21 (2008) 53–60. [6] J. Cros, P. Viarouge, Synthesis of high performance PM machines with concentrated windings, IEEE Trans. Energy Convers. 17 (2002) 248–253. [7] A.M. El-Refaie, T.M. Jahns, D. Novotny, Analysis of surface PM machines with fractional-slot concentrated windings, IEEE Trans. Energy Convers. 21 (2006) 34–43. [8] N. Bianchi, M.D. Pré, G. Grezzani, S. Bolognani, Design considerations for fractional-slot winding configurations of synchronous machines, IEEE Trans. Ind. Appl. 42 (2006) 997–1006. [9] Z.Q. Zhu, D. Howe, Influence of design parameters on cogging torque in permanent magnet machines, IEEE Trans. Energy Convers. 15 (2000) 407–412. [10] N. Bianchi, S. Bolognani, Design techniques for reducing the cogging torque in surface-mounted PM motors, IEEE Trans. Ind. Appl. 38 (2002) 1259–1265. [11] Z.Q. Zhu, Fractional slot permanent magnet brushless machines and drives for electric and hybrid propulsion systems, in: Ecological Vehicles and Renewable Energies (EVER 09), Monaco, France, 2009, pp. 1–11. [12] S. Lin, T.X. Wu, L. Zhou, F. Moslehy, J. Kapat, L. Chow, Modeling and design of super high speed permanent magnet synchronous motor (PMSM), in: Aerospace and Electronics Conference, Dayton, OH, 2008, pp. 41–44. [13] G.H. Jang, C.I. Lee, Dual winding method of a BLDC motor for large starting torque and high speed, IEEE Trans. Magn. 41 (2005) 3922–3924. [14] M. Barcaro, N. Bianchi, F. Magnussen, Analysis and tests of a dual three-phase 12-slot 10-pole permanent-magnet motor, IEEE Trans. Ind. Appl. 46 (2010) 2355–2362. [15] M. Barcaro, N. Bianchi, F. Magnussen, Faulty operations of a PM fractional-slot machine with a dual three-phase winding, IEEE Trans. Ind. Electron. 58 (2011) 3825–3832. [16] N. Bianchi, E. Fornasiero, Impact of MMF space harmonic on rotor losses in fractional-slot permanent-magnet machines, IEEE Trans. Energy Convers. 24 (2009) 323–328. [17] N. Schofield, K. Ng, Z.Q. Zhu, D. Howe, Parasitic rotor losses in a brushless permanent magnet traction machine, in: Eighth International Conference on Electrical Machines and Drives, Cambridge, UK, 1997, pp. 200–204. [18] F. Magnussen, H. Lendenmann, Parasitic effects in PM machines with concentrated windings, IEEE Trans. Ind. Appl. 43 (2007) 1223–1232. [19] D.J. Perreault, V. Caliskan, Automotive power generation and control, IEEE Trans. Power Electron. 19 (2004) 618–630. [20] J.R. Hendershot, T.J.E. Miller, Design of Brushless Permanent Magnet Motors, Magna Physics Publishing and Clarendon Press, Oxford, 1994. [21] D. Ishak, Z.Q. Zhu, D. Howe, Comparison of PM brushless motors, with either all or alternative wound teeth, IEEE Trans. Energy Convers. 21 (2006) 95–103. [22] D. Zarko, B. Drago, T.A. Lipo, Analytical solution for electromagnetic torque in surface permanent-magnet motors using conformal mapping, IEEE Trans. Magn. 45 (2009) 2943–2954. [23] D. Zarko, B. Drago, T.A. Lipo, Analytical calculation of magnetic field distribution in the slotted air gap of a surface permanent-magnet motor using complex relative air-gap permeance, IEEE Trans. Magn. 42 (2006) 1828–1838. [24] E. Mes¸e, M. Ayaz, M. Tezcan, K. Yılmaz, E. Özdemir, A permanent magnet synchronous machine with motor and generator functionalities in single stator core, in: IEEE 2013 Compumag Conference, Budapest, Hungary, 2013, pp. 1–4.
Contents lists available at ScienceDirect
Electric Power Systems Research journal homepage: www.elsevier.com/locate/epsr
Design considerations of a multitasked electric machine for automotive applications Erkan Mese a , Murat Ayaz b,∗ , M. Murat Tezcan a a b
Yildiz Technical University, Department of Electrical Engineering, Istanbul, Turkey Kocaeli University, Department of Electrical Education, Kocaeli 41050, Turkey
a r t i c l e
i n f o
Article history: Received 27 September 2014 Received in revised form 11 October 2015 Accepted 13 October 2015 Keywords: Permanent magnet synchronous machine (PMSM) Concentrated winding Magnetic decoupling Dual winding PM machine (DWPMSM) Fault tolerant operation Hybrid electrical vehicle applications
a b s t r a c t This paper introduces a new electric machine for Front-End Accessory Drive (FEAD) of Hybrid Electric Vehicles (HEVs). The novelty of the electric machine lies in its ability to perform multitasking. Besides the independent motor or independent generator operation like any other electric machine, the proposed machine is able to run as motor and generator simultaneously. A dual winding electric machine forms the core part of the proposed system. Windings of the electric machine are concentrated type so that electrical and magnetic isolation is maintained. This allows simultaneous motoring and generating operation in the single housing of electric machine. Design considerations of the proposed electric machine have been outlined in the paper. Comparison between proposed and conventional approaches has been made to highlight potential benefits of the new approach. Thermal analyses have been performed to show the suitability and limits of the multitasked electric machine in the HEV applications. Overall performance of the proposed machine is presented by experimental results. © 2015 Elsevier B.V. All rights reserved.
1. Introduction Electric machines designed with dual winding sets in their stator have recently become popular due to providing opportunities to solve many industrial problems. For instance, continuous operation of the mechanical and the electrical accessories in HEVs with stop/start functionality can be achieved by using such a dual winding electric machine which has simultaneously motor and generator operation capability [1,2]. The goal of this study is to present design considerations on a decoupled dual winding permanent magnet synchronous machine (DWPMSM) for continuous operation of accessory loads such as air conditioner compressor, power steering pump, water pump and etc. The proposed machine is able to run as motor and generator simultaneously without interaction between these operations. Concentrated winding type is used in the proposed machine to achieve decoupling between winding sets. Since each concentrated coil around any stator tooth has an independent magnetic circuit, magnetic decoupling between winding sets is possible [3–7]. Other advantages are shorter end windings, higher slot fill factor,
∗ Corresponding author. Tel.: +90 2623032307; fax: +90 2623032307. E-mail addresses: [email protected] (E. Mese), [email protected] (M. Ayaz), [email protected] (M.M. Tezcan). http://dx.doi.org/10.1016/j.epsr.2015.10.017 0378-7796/© 2015 Elsevier B.V. All rights reserved.
wider speed range of flux weakening, easier manufacturability and higher efficiency. On the other side, there are some deficiencies related to concentrated winding like noise, torque ripple, unbalanced magnetic forces due to additional harmonic contents [8–15]. Furthermore, MMF harmonic spectrum of concentrated winding machines contains both super-synchronous and sub-synchronous components. Whereas a balanced distributed winding has only super synchronous harmonic content in its MMF. Sub-synchronous space harmonics would cause higher rotor core and magnet losses because of relatively higher frequency [16–18]. Furthermore, subsynchronous space harmonics have relatively greater wavelength and this causes deeper penetration of the flux to the core and further increase in losses [17]. So, electric machines with concentrated winding would have higher core and magnet losses at high speed region. In particular, the losses become more significant in highpole count machines. This drawback of the concentrated winding has been extensively elaborated in the paper. 2. Material and methods of accessories drive systems in HEVs In hybrid electric vehicle (HEV), nonstop accessory load-driving feature is implemented with drive-by-wire concept as shown in Fig. 1(a). With this concept, each of the accessory loads is powered through its own electric motor and drive system fed by high voltage
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Fig. 1. (a) Drive by wire system, (b) proposed electrical accessory drive system.
battery. Drive by wire is an expensive method since many motors and inverters must be used. Furthermore it is an inefficient method since fractional power motor drive systems tend to be less efficient than high power motor drive systems. In addition, duty cycle of single motor is more likely filled with higher load intervals compared to the distributed multi-motor system. On the other hand, low voltage (12 V) power is provided through a DC/DC converter as shown in Fig. 1(a). Lundell alternator can also be used for 12 V power. The low efficiency of this alternator, which varies between 45% and 60%, depends on its speed and load level, is the main drawback preventing from the future vehicle applications [19]. Electrical accessory drive system with the proposed machine for driving accessory loads and charging 12 V battery in HEVs is shown in Fig. 1(b). As seen from the figure, the proposed dual winding machine runs as motor and generator simultaneously or just operates as a generator. If the engine is running the dual winding machine operates only as generator. Otherwise, simultaneous motor and generator operation takes place. With the decoupled dual winding sets, mechanical and electrical power can be generated in a single electric machine instead of two separate machines. By using electrical accessory drive system with the proposed machine several small motors and their drivers in drive by wire system are removed. Furthermore, several inherited components of conventional accessory drives of conventional ICE based vehicles (such as pumps and compressors) can be still used. These facts are considered as cost saving advantages by some automotive manufacturers. Table 1 summarizes components of the conventional and proposed accessory drive system. The high voltage DC/DC converter or Lundell alternator used for charging low voltage battery (12 V) in drive by wire system is eliminated with
the generator side of DWPMSM. On the other hand, a DC/DC converter with an uncontrolled rectifier or a controlled rectifier is to be used to regulate output voltage of the generator side of DWPMSM. However, these converter structures are more efficient and cheaper solution compared with the high voltage DC/DC converter due to lower input voltage. The output power of the DWPMSM is to be determined depending on power requirement for accessories in vehicles as shown in Fig. 2. As can be seen from the figure, DWPMSM should process around 6 kW power at 1500 rpm to drive all accessories. Power flow of the proposed accessory drive system and operation modes are detailed in Fig. 3. There are two operation modes of DWPMSM depending on mechanical power generation of internal combustion engine (ICE). Table 1 Part count comparison among different accessory drive systems. Parts
Conventional non-hybrid
Drive by wire
EADS
Crank pulley Pulley clutch Belt Water pump Alternator Air pump Air pump eMotor Power steering pump EPS pump EPS eMotor AC compressor AC comp. eMotor Auxiliary power unit (DC/DC converter) eMotor + Inverter
X
X
X X X X
X X
X X X X X X
Unique X
X
X
X Unique X Unique X X
X
X
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149
Fig. 2. Power requirements for mechanical and electrical accessories in HEVs.
In the first operation mode, ICE is off and power requirement for mechanical accessories should be provided by motor side of DWPMSM. The motor windings are driven by an inverter to generate 6 kW at 1500 rpm. Half of the generated power by the motor windings is consumed by mechanical accessories connected to shaft of DWPMSM. The remaining half of the power is used to supply electrical accessories through generator windings as illustrated in Fig. 3(a). In the second operation mode ICE is on and no motoring is required since mechanical power is generated by ICE as shown in Fig. 3(b). Therefore, the motor windings are unexcited and the generator windings generate almost 3 kW between 1500 rpm and 6000 rpm shaft speed to supply low voltage system in the vehicle.
Mechanical Accessories
3.1. Determining geometrical dimensions of DWPMSM Magnetic coupling between winding sets is to be around zero to achieve a multifunction operation in the single machine housing. Single layer concentrated winding topologies has better decoupling performance compared to double layer concentrated winding [14]. Also, surface mounted PM (SPM) structure allows to increase decoupling in the proposed machine. Besides decoupling performance of the SPMs, lower unbalanced radial forces, reduced stack
Inverter
Internal Combustion Engine
+ - 300 V
0 rpm
Comp. Pump Steering Pump
3. Calculation process of dual winding PMSM design parameters
Motor Terminals 6 kW
High Voltage Battery
DWPMSM 3 kW
3 kW
3 kW
Etc.
1500 rpm Generator Terminals
Mechanical Power Uncontrolled Rectifier
Electrical Power
DC-DC Electrical Chopper Accessories
a)
Mechanical Accessories Comp.
1000 ~ 6000 rpm
Pump Steering Pump
Inverter
Internal Combustion Engine
3 kW
3 kW
+ - 300 V Motor Terminals 0 kW 3 kW
Etc.
High Voltage Battery
DWPMSM 1000 ~ 6000 rpm Generator Terminals
Mechanical Power Uncontrolled Rectifier
Electrical Power
Electrical DC-DC Chopper Accessories
b) Fig. 3. Operating modes of DWPMSM, (a) first mode (motor/generator operation), (b) second mode (generator only operation).
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length and lower stator MMF harmonics are main advantages compared with interior PM machines. As a result of these aspects, surface mounted PM machine with single layer concentrated winding would be the first design decision in the design process. After deciding surface mounted PM machine, geometric envelope of the machine is determined. This includes roughly finding rotor diameter and axial length. Analytical method to design a PMSM is well documented in [20]. Similar method has been adopted here. Torque equation of a radial flux electric machine is given as T = kDr2 Ls
Stator Outer Diameter
Magnet Height Airgap
Tooth Widht Magnet Width
(1)
In Eq. (1) Dr is the rotor outer diameter, Ls is axial length of the stator core. Torque coefficient k which is related with the thermal loading and cooling technique of machine. In the proposed electric machine, the worst case for the copper losses occurs during motor and generator operation at 1500 rpm speed. On the other hand, the worst case for the core losses occurs during high speed operation (above 5000 rpm). In this mode, machine operates only as generator as described in Section 2. Although loss number is greater in high speed generator only mode, the duration is shorter. Hence simultaneous motor and generator operation at 1500 rpm reflects more severe thermal condition and resembles conventional electric machine where all stator coils carry equal current. By considering thermal loading cases stated above, torque coefficient is assumed to be 14,400 N m/m3 for air cooled motor. This value is adopted from [20] where it is stated that the torque coefficient can be selected between 14,000 and 42,000 N m/m3 for totally enclosed and sintered rare earth PM motor. Since the proposed machine is naturally cooled lower limit of this range is preferred. After narrowing the design space by sizing equation, in order to generate 6 kW shaft power at 1500 rpm shaft speed, detailed design steps have been taken iteratively by an analytical design tool running in the computer. Geometrical dimensions of the PMSM are shown in Fig. 4. Air gap length, stator or rotor skew angle, magnet height and slot span have significant effects on the motor performance. This section presents these effects by parametric analysis made by Ansoft–Rmxprt software. Initial design values of air gap length, stator skew angle, slot span, magnet height and width are 0.5 mm, 0◦ , 4 mm, 5 mm and 12.8 mm, respectively. Fig. 5(a) shows the results of parametric analysis made to minimize torque ripple. By analysis, torque ripple is reduced from 7.24% to 2.5%. Torque ripple minimization effort during the design stage aims to find optimum geometry for the minimum torque ripple and ignores harmonics
Yoke Height
Rotor Inner Diameter
Slot Opening
Rotor Outer Diameter Stator Inner Diamater
Fig. 4. General description of slot and tooth geometry.
in the phase current. Hence the results shown in Fig. 5(a) assume pure sinusoidal current in both winding of DWPMSM. Similar parametric analysis is performed for back EMF harmonic distortion and cogging torque. The respective results are shown in Fig. 5(b) and (c). The geometrical and electrical design parameters of DWPMSM are given in Table 2. This table summarizes the details of the DWPMSM design and they are prepared by parametric analysis. 3.2. Winding structure of DWPMSM In the literature, there are many studies about PMSM winding topologies. Beside many advantage stated before, concentrated winding topology has minimum magnetic coupling between the coils [14,21]. An important requirement in the proposed FEAD system is magnetic coupling between two winding sets must be minimum or preferably zero for proper operation. In concentrated winding, each coil has its own independent magnetic circuit as shown in Fig. 6 where short flux paths are dominating the magnetic circuit. Long flux paths also exist but they are minor. Dominance of short flux path can be interpreted as very low magnetic coupling between neighbor coils. This enables simultaneous motor and generator operation. Although magnetic independence, each coil still contributes to the total air gap flux density. Finite element analysis (FEA) is performed with Maxwell 2D to examine existence of magnetic coupling between two winding
Table 2 Electrical and physical design parameters of DWPMSM. Electrical parameters
Physical parameters Gen.
Number of turns
5
Wire diameter (mm) Rs-stator phase resistance () Lq-inductance (mH) Ld-inductance (mH) Flux linkage by magnets (V s) Stator tooth flux density (T) Airgap flux density (T) Rotor yoke flux density (T) Rated torque (N m) Output power (W) Back EMF constant (V/rad/s)
2.59 0.0015 0.043 0.043 0.0125 1.25 0.82 0.27 21.23 3000 0.285
Motor 21
2.3 0.033 0.776 0.776 0.055 1.44 0.82 0.27 38.21 6000 1.2
Stator parameters
Stator slot parameters
Stator outer Dia. (mm)
240
Hs0 (mm)
2.8
Stator inner Dia. (mm) Stator core length (mm) Stator skew angle (◦ )
140 120 11.1
Rotor Parameters Rotor outer Dia. (mm) Rotor inner Dia. (mm) Rotor core length (mm) Magnet width (mm) Magnet height (mm) Inertia (kg/m2 )
138 75 120 16.14 5.0 0.033
Hs1 (mm) Hs2 (mm) Bs0 (mm) Bs1 (mm) Bs2 (mm) Rs (mm)
0 28 5.5 8.8 14.4 4.5
Net slot area (mm2 ) Slot fill factor (%) Number of slots Number of poles
380.9 56.01 24 22
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Torque Ripple (%) 10
THD (%) 4
11.1º %4.678
0
0
Skew (º) 5
10
11.1º %0.68
2
5 0
Skew (º)
0
15
5
10
1 mm %4.227
5 0 0
2
1
10
Airgap (mm)
0 0
Slot Opening 8 (mm)
0
2
4
1
2
Airgap (mm)
1
5.5 mm %2.418
5
15
1 mm %0.68
1
10
151
6
5.5 mm %0.68
0 2
4
Slot Opening 8 (mm)
6
4
10
16.74 mm %2.418
5
0 8
10
12
14
5
16
18
0 8
Magnet Height 8 (mm)
0 4
5
10
12
14
16
18
Magnet Width 20 (mm)
1
5 mm %2.418
3
16.74 mm %0.68
2 Magnet Width 20 (mm)
6
7
5 mm %0.68
0
3
4
5
6
a)
7
Magnet Height 8 (mm)
b) Cogging Torque (mNm) 24 16 8 0 0 5 3 2 1 0
11.1º 0.55 mNm Skew (º)
10
15
1 mm 0.55 mNm 0
1
6 4 2 0
2 5.5 mm 0.55 mNm
2
4
Slot Opening 8 (mm)
6
3 2 1 0
16.74 mm 0.55 mNm 8
1
10
12
14
Airgap (mm)
16
18
Magnet Width 20 (mm)
5 mm 0.55 mNm
0
3
4
5
6
7
Magnet Height 8 (mm)
c) Fig. 5. DWPMSM, (a) torque ripple performance, (b) THD% of induced back EMF performance, (c) cogging torque performance.
sets. In Fig. 7, the effect of generator side loading on the motor side torque output is shown, where generator is connected to an uncontrolled rectifier. In conventional distributed type dual winding electric machine, even if a severe interaction between fundamental components of generator and motor magnetic fields is not observed, higher order harmonics of generator magnetic field adversely affect motor magnetic field. In concentrated winding sets, generator side loading has weaker adverse effect on motor torque as shown in Fig. 7. In particular, single layer concentrated winding
approach seems to be most effective approach in magnetic decoupling of two winding sets. 3.3. Analytical model of field distribution and decoupling between winding sets This section shows an analytical model of the magnetic decoupling between two winding sets in a single layer concentrated winding configuration. For this purpose, the magnetic field
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Fig. 6. (a) Concentrated single layer dual winding electric machine, (b) flux distribution of single layer coils.
40 30
Igen 0A
20
Igen 55 A
Igen 160 A
Igen 110 A
Distributed Winding Double Layer Con. Winding Single Layer Con. Winding
10 0
0
x = ln (r), y = , ∂z/∂s = 1/s ⇒ s = ez . In the second stage, the corner points of the Z Cartesian coordinate system are determined. The current in a coil directly affects the magnetic vector potential in the middle of the air gap and the Schwarz Christoffel transformation is used to transform the Z complex plane to a W complex plane.
Torque (Nm) 50
10
20
30
s = r cos s + jr sin s
(2)
z = log (s)
(3)
2g z = −j
Time
1 √ arctan a
Fig. 7. Motor torque variation for different generator winding load levels (n = 1500 rpm).
distribution model of a slot current is used. The magnetic field of the slot current is modeled by conformal transformation, as described in [22]. Four conformal transformations are used to perform the analytical method. Actual geometry in S complex plane, which has coordinates in the form of s = m + jn = rej , is shown in Fig. 8(a). The Z Cartesian complex coordinate system has coordinates in the form of z = x + iy shown in Fig. 8(b). Logarithmic conformal transformation is used to calculate the polar coordinates
1 + w − a/w + 1 1 + ln 2 1 − w − a/w + 1
θ2 θs/2 W=a θ1
4
n
α
Rotor y Region
3
Rr
a)
W=-1
2 Ln(Rs)
x
Rs 1
bo’ W=0
β g’
1 ln(Rr)
θs
m
+ ln (Rr) + j
g’=ln(Rs/Rr) bo’=θ2-θ1 α=π/2 β=3π/2
g
θ1 θ2
2
s 2
(4)
Eqs. (2)–(4) can be used to transform an arbitrary point in the actual geometry in the middle of air gap in the S complex plane to a complex Cartesian W complex plane having coordinates in the form of w = u + iv, where the solution exists. Fig. 9(a) shows the new coordinates of the geometry after Schwarz Christoffel transformation in the W complex plane. Solving Eqs. (3) and (4) simultaneously yields several w = u + iv values. These values are used for determining t = p + iq values in a slotless T plane, as shown in Fig. 9(b), using Eqs. (5) and (6).
5 6
w − a/w + 1 √ a
W=∞ W=-∞
b)
Fig. 8. (a) Slot opening in the S plane, (b) Half of the slot opening in Z plane.
Stator Region
E. Mese et al. / Electric Power Systems Research 131 (2016) 147–158
153
W=-∞ W=∞
6
θs
5 γ’ γ’=ln(Rs/Rr)
W=-1 θs/2
v
W=1
q 1 ln(Rr) Z=ln(Rr)-j∞
Z=ln(Rr)+j∞
-∞
Z=ln(Rs)+jθ1
0
Z=∞+jθ1
1
a
Z=ln(Rs)+j∞ u
Z=ln(Rs)+jθ2
p
∞
b
2 Ln(Rs)
W=0
a)
b)
Fig. 9. (a) One half of the slot opening in the W complex plane, (b) Slot opening in the T complex plane.
Bar_air(T)
Bar_air (T) 0.0002
0.0002
θ (Mech.)
θ (Mech.) π/48
π/24
π/16
π/12
-0.0002
π/8
5π/48
π/6
5π/24
-0.0002 Slot Center
Slot Center
a)
b)
Fig. 10. Analytical solution, (a) the left coil side radial flux density distribution (b) the right coil side radial flux density distribution.
The field distribution in the S complex plane is given by Eq. (7). One coil side field solution in Cartesian coordinates can be found by using Eq. (8), where the Cartesian coordinate field solution is transformed into the cylindrical coordinate system [23]. Bar is the radial component of the winding flux density distribution in the air gap.
∂t g 1 g s ⇒ t = j ln (w) + ln (Rs) + j =j w 2 ∂w Bt = −jo
∂ϕ Nc = jo I 2 ∂q
Nc 1 Bs = Bm + jBn = jo I 2 s Bar = Bm cos + Bn sin
(5) (6)
w−a w+1
magnetic field distributions yields a new flux density distribution, shown in Fig. 11. The plot shows that no magnetic field exists outside the coil region, suggesting that the neighbouring coils will not be affected by the operation of this coil [24].
Bar_air(T)
0.0002
∗
θ (Mech.)
(7) (8)
The field distributions of current in two adjacent slots are modelled by the procedure above, for a machine geometry with 24 slots, 22 poles and a single layer concentrated winding configuration. Coil turn number Nc (Nc = 10 turns) and coil current I = 10 A. The flux density distribution of the left and right sides of the coil are shown in Fig. 10(a) and (b), respectively. Superposition of the two
π/24 -0.0002
Magnetic Decoupling Zone
π/12
π/8
π/6 Magnetic Decoupling Zone
-0.0004 Slot Center
Slot Center
Fig. 11. Analytical solution of a single layer concentrated winding radial flux density distribution and magnetic decoupling zones.
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Table 3 Electrical and physical parameters comparation of different PMSMs. Electrical parameters
Machine type
Physical parameters
Machine type
Rotor data
Conven. motor
DWPMS (MG)
Conven. gen.
Thickness of mag. (mm) Type of magnet Total net weight(kg) Stator data Number of slots Outer Dia. (mm) Inner Dia. (mm) Total length (mm)* Len. of core (mm) Type of steel Slot area (mm2 ) Slot fill factor (%)
5 NdFe30 19.16
5 NdFe30 32.68
5 NdFe30 20.17
24 215 123 111.12 89.5 M1924G 357.7 52.72
24 240 140 144.96 120 M1924G 380.9 56.01
24 218 122 112.19 90 M1924G 361.39 61.02
DWPMSM No load data
Conven. motor
St. teeth flux density St. yoke flux density THD of Ind. voltage (%) Full load data Max. line voltage(V) RMS Ph. current (A) Efficiency (%) Power angle (◦ ) Torque angle (◦ ) Iron-core loss (W) Arm copper loss (W) Total loss (W) *
Motor oper.
1.27 0.46 –
1.25 0.44 –
155.75 34.83 93.54 – 25.73 116.48 143.21 413.96
156.11 34.93 92.85 – 26 190.98 116.01 461.53
Gen. oper. 1.44 0.43 0.68 36.65 137.28 90.48 40.98 – 231.19 84.527 315.72
Conven. gen. 1.42 0.42 0.16 36.53 134.25 92.34 40.01 – 147.7 101.05 248.75
Includes end winding.
3.4. Physical comparison with conventional approaches The proposed machine is compared with its alternative where two separate machines are used. To perform an accurate comparison between machines, electrical and magnetic design parameters of all machines must be equal to the proposed machine. These parameters are phase current, current density, fundamental back EMF, stator teeth and yoke flux density, magnet type and etc. as given in Table 3. Calculating the volumes and masses of the machines from their given geometrical parameters in Table 3, DWPMSM’s total volume is 6558.17 cm3 whereas single motor and single generator’s total volume is 8222.14 cm3 , respectively. Similarly DWPMSM has lower mass compared with the total weight of classical motor and generator. When total volume and weight numbers are taken into the account, dual winding structure has significant advantage. In addition all of above advantages, some volume, weight and cost savings also arise when we consider reduced number of position sensor, end bells and bearings. Reduction in these auxiliary but essential elements of the machine yields significant amount of space saving. 3.5. Thermal analysis of DWPMSM This section presents 2D numerical analysis and experimental results of DWPMSM’s thermal behavior by using finite-element analysis (FEA). Thermal analysis is done for 9 kW power ratings (6 kW for motor operation and 3 kW for generator operation). Totally enclosed and non ventilated thermal conditions are
Power Losses (W)
600 500
200
Stator Core Losses Rotor Core Losses Magnet Losses Copper Losses
Temperature (°C)
700
400 300 200 100 0 1000
2000
3000 4000 Speed (rpm)
a)
considered. That means a combination of natural convection, radiation and conduction occurs. The inside and outside fluid is considered as only air and hence named as air-cooled DWSMPM. The total loss versus speed for the DWPMSM is given in Fig. 12(a). As seen from the figure, the core losses and eddy current losses increase with increasing shaft speed as expected. Temperature versus speed variation obtained by the thermal analysis using these loss values is shown in Fig. 12(b). Fig. 13 shows the experimental and simulation thermal photographs during generator operation of the DWPMSM. Loss estimation shows that core and magnet losses have tendency to increase with speed. The reason for this trend can be explained by winding configuration. As stated in Section 1, concentrated winding causes sub-synchronous harmonics and this increases core and magnet losses. The influence of this can also be seen in the efficiency plot at high speed region. More serious influence of high losses can be seen after thermal analysis of the machine. As can be seen from Fig. 12(b), both core losses and magnet losses are causing very high temperature rise. This analysis shows that continuous operation at high speed particularly above 5000 rpm may be dangerous for the magnets. Continuous operation beyond 5000 rpm is not required for intended application of the proposed machine. Instead, intermittent and short term loading of the generator above 5000 rpm is a common loading scenario for the driving cycles under considerations. Hence with careful magnet grade selection and with due care for the operating speed of the generator, magnet temperature rise can be managed. For more demanding applications and cases, water cooling becomes mandatory.
5000
6000
150
Winding End Magnet Stator Tooth Stator Yoke
100 50 0 1000
2000
3000 4000 Speed (rpm)
b)
Fig. 12. (a) Power losses versus speed of the DWPMSM, (b) Temperature versus speed of the DWPMSM.
5000
6000
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155
Fig. 13. Thermal photographs during generator operation at 1500 rpm and fully loaded (a) experimental, (b) simulation.
4. Experimental results and discussion Performance evaluation of the proposed machine is performed by using experimental setup shown in Fig. 14. In experimental test bed, a squirrel cage induction motor is controlled by a four-quadrant adjustable speed driver that enables motoring and generating operation. Also, a rotary torque sensor between the induction motor and dual winding PMSM measures torque level dynamically. Induction machine in the experimental setup represents mechanical load in accessory drive system as well as internal combustion engine of the vehicle. In Mode 1, induction machine imitates mechanical accessory loads and runs as generator. In Mode 2, it emulates internal combustion engine and runs as motor. A resistive load bank with adjustable 3 kW power capabilities is used for emulating electrical accessory loads. An uncontrolled rectifier/DC–DC converter set is connected between generator windings and the resistive load bank in order to regulate output voltage. 4.1. Study of magnetic coupling on DWPMSM This test is performed to verify FEA results in terms of the magnetic decoupling between winding sets of the proposed electric machine. In the test, motor back EMF waveforms are checked to see the effect of generator winding load level (loaded or unloaded). The expectation is to see no difference between two back EMF waveforms. Generator windings are connected to a three phase bridge uncontrolled rectifier which has an output connected parallel to
Fig. 15. Motor Back EMF waveform with no load (yellow trace), Motor Back EMF waveform with load (white trace), Generator phase voltage waveform with load (pink trace), Generator Output current waveform (green trace) at 1500 rpm rotor shaft speed. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)
a 450 VDC–2200 F capacitor. Capacitor terminals are connected to the 3 kW resistive load. Firstly, the proposed electric machine is driven at constant speed (1500 rpm) by induction motor. The uncontrolled rectifier terminals are open circuit, so the generator windings are unloaded and the motor windings’ back EMF waveform is recorded as shown in Fig. 15 (white trace). After that,
Fig. 14. Experimental setup for DWPMSM.
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Fig. 16. Operation modes of the proposed DWPMSM.
4.2. Operation modes of DWPMSM machine It has been shown in previous subsection that there is no magnetic coupling between motor and generator winding sets. Consequently, operation with only one winding set (only motor or only generator operation) or simultaneous operation of two winding sets (motor–generator operation at the same time) would be possible without deteriorating any machine performance. Operation modes of DWPMSM are shown in Fig. 16. 4.2.1. Independent motor operation for DWPMSM machine Performance tests of the DWPMSM motor operation are implemented by different mechanical load conditions. Motor only operation is performed at 1500 rpm constant speed which is almost the same speed in real accessory drive application. Motor windings are driven by a 600 V–50 A voltage source inverter. Modulation strategy of the inverter is sinusoidal PWM. Carrier frequency for the control signals is 10 kHz. The motor windings are controlled with q axis current value by the inverter because flux weakening is not needed at this speed. In Fig. 17, shaft torque waveform recorded by torque sensor and electrical terminal variables for motor operation are shown. The generator windings are open circuit in this operation mode.
Fig. 17. Motor operation mode (22.82 N m load torque).
100
Efficiency (%)
1300 W resistive load is connected to uncontrolled rectifier terminals without changing the shaft speed and motor windings back EMF waveform is shown in Fig. 15 (yellow trace). As seen from the figure, the motor windings’ back EMF waveforms are not affected by generator winding in the face of load change. As a result of the experimental test, magnetic decoupling between winding sets is assured with the proposed machine.
75 50 25
0 1000
1500 rpm
2000 3000 4000 5000 6000 Motor Output Power (W)
Fig. 18. Experimental efficiency data from independent motor operation.
The obtained machine efficiency during experimental tests is shown in Fig. 18. As seen from the figure, the motor operation efficiency of the DWPMSM varies between 88% and 93% depending on output power values. 4.2.2. Independent generator operation for DWPMSM machine For this operation, DWPMSM is driven by the squirrel cage induction motor. Motor windings are not excited and generator windings are connected to resistive load via an uncontrolled rectifier and DC/DC converter combination. DC/DC converter regulates output voltage at 13.7 V. For different loads and different shaft speeds, output data of generator winding is recorded. Output waveforms of generator operation for 2500 rpm, 100 A at 13.7 V DC are shown in Fig. 19. Independent generator efficiency is obtained experimentally. Fig. 20(a) shows the generator operation efficiency of the DWPMSM at different shaft speed and load levels. As seen from the figure, efficiency is maximized at 1500 rpm with 92% peak value.
Fig. 19. Generator operation mode (2500 rpm–13.7 V–100 A DC load).
100
100
75
75
1500 rpm 2000 rpm 2500 rpm 3000 rpm 3500 rpm 4000 rpm 4500 rpm 5000 rpm
50
25 0 200
800
1400
2000
2600
Efficiency (%)
Efficiency (%)
E. Mese et al. / Electric Power Systems Research 131 (2016) 147–158
50 25 0 1000
3200
157
DWPMSM LUNDELL
2000
3000 4000 Speed (rpm)
Generator Output Power (W)
a)
5000
6000
b)
40
80
30
60
40
20 10 0
Time (ms) 5
6
7
-10
8
Voltage (V)
Torque (Nm)
Fig. 20. (a) Experimental efficiency data from independent generator operation, (b) Comparison of alternators efficiency performance.
0 -20
Time (ms) 0
1
2
3
60
20
40
10 Time (ms) 1
2
3
7
8
b)
30
0
6
Experimental Simulation
-80
4
5
6
7
8
-10 -20
Experimental Simulation
-30
Current (A)
Current (A)
5
-60
a)
0
4
-40
Experimental Simulation
-20
20
20 0
Time (ms) 0
1
2
3
4
5
6
7
8
-20
-40
Experimental Simulation
-60
c)
d)
Fig. 21. Simultaneously motor–generator operation mode at 1500 rpm/13.7 V–40 A DC load (a) torque waveform, (b) generator phase voltage waveform, (c) generator phase current waveform, (d) motor winding phase current waveform.
A comparative plot is given in Fig. 20(b) showing efficiency profile for the generator operation of DWPMSM and Lundell alternator. As seen from the figure Lundell alternator efficiency varies between 45 and 60% [19]. However, generator efficiency of DWPMSM varies between 80 and 93%. This is not a surprising result because similar numbers are reported in the literature for conventional PMSM and the proposed machine’s generator performance seems to be in line with these reporting [19].
4.2.3. Experimental simultaneous motor/generator operation In this section, experimental performance of simultaneous motor/generator operation of the DWPMSM is handled. The mechanical output power required for system is only generated by the motor windings of DWPMSM driven by its own inverter. The induction machine representing the mechanical accessories in the system operates as generator with four-quadrant adjustable speed driver and consumes half of the mechanical power. Remaining
Table 4 Output performance data of simultaneous motor–generator operation at 1500 rpm. Motor–generator operation at 1500 rpm
Motor Q axis current 20
Shaft torque (N m) Motor input power (W) Mechanical out. power (W) Generator out. power (W) Total out. power (W) Machine efficiency (%)
9.54 2848.5 1497.8 1098.8 2596.6 91.1
30 4.7 2912.5 737.90 2033.2 2771.1 95.14
18.2 4368.1 2857.4 1136.6 3994.0 96.8
40 13.56 4459.4 2128.2 2046.0 4174.2 93.6
14.7 5812.8 3611.0 1964.0 5575.0 95.5
23 5832.4 2307.9 3152.4 5356.8 93.9
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mechanical power is transferred to supply electrical loads by the generator windings. Fig. 21 shows the related data of the operating point where the motor speed is set 1500 rpm and generator DC/DC converter is regulating at 13.7 V and 40 A DC. In Fig. 21(a) and (b), high frequency spikes are related PWM switching and caught by the measurement probes during tests. The experimental efficiency of DWPMSM during simultaneous motor/generator operation is given in Table 4. The efficiency of DWPMSM varies between 91% and 97% depending on sharing of mechanical and electrical power generation in the machine. The experimental efficiency values show that the proposed electric machine has similar performance compared with conventional PM machine. Furthermore, experimental result verifies simultaneous motor and generator action in single stator core. 5. Conclusions In this paper, a new electric machine is proposed for some automotive applications where multi-tasking is requested in an effort to develop more compact components. The proposed machine, referred as Dual Winding PMSM (DWPMSM), combines motor and generator functions in the same structure so that it acts as motor and generator simultaneously. After presenting application considerations of the DWPMSM, design procedure is given with emphasis on efficiency, torque density, cogging torque and torque ripple. Winding is the most important aspect of the DWPMSM development and due attention should be paid for proper operation of the machine. The study covered critical steps in the winding design such as location, mutual decoupling assurance and etc. Analytical and experimental studies have been presented in this regard. Quantitative study is presented to compare volumes and masses between conventional and proposed technology. Thermal analyses are also performed to show the limits of the proposed DWPMSM. Exhaustive experimental studies are performed to cover all operating modes of the DWPMSM in a HEV Front End Accessory Drive system. Overall results presented in this paper show that the proposed machine could be a viable solution for a HEV Front End Accessory Drive system. Further application possibilities may exist for the proposed machine in the automotive electrification. Acknowledgments This work was supported by the Scientific and Technological Research Council of Turkey under contract number 110E111. Also Joel M. Maguire from BorgWarner Corporation and Gary E. McGee from General Motors Corporation are appreciated for their inspiration and support during problem definition of accessory systems in hybrid electric vehicles.
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