Novel cooling methods using Flux-barriers

Novel cooling methods using Flux-barriers
A. Nollau and D. Gerling
Φ
Abstract – This paper presents new cooling methods for
permanent magnet synchronous machines (PMSM) with flux
barriers. For a vehicle powertrain application the permanent
magnet synchronous machines is a very popular choice,
because of high power density, a high efficiency and the small
package. However, this machine is susceptible to suffer
insulations failures of coils and demagnetization of magnets
under severe thermal condition. Therefore, it is important to
have an efficient cooling method for the PMSM to generate
high heat dissipation. The configurations of the new cooling
methods are presented in this paper. Three different topologies
are compared with a standard water cooling jacket
configuration. A finite volume Computational Fluid Dynamic
(CFD) model is applied to have a very accurate and
comprehensive simulation.
Index Terms-- Cooling Method, CFD, FEA, permanent
magnet machine, thermal analysis
I.
the temperature distribution and related thermal quantities
in a system or component. Typical quantities of interest are
temperature distribution, thermal gradients and thermal flux.
This paper present new cooling methods for a PM
machine with flux barriers and compares the results with a
standard water cooling method.
II.
PROPOSED COOLING APPROACH
A.
Machine Design
The researched machine is 24-Teeth/28-Poles permanent
magnet (PM) – machine presented in [1]. The following
Fig. 1 shows the geometry of the designed PM machine.
The stator consists of 24 slots and the rotor consists of 28
rectangular permanent magnets inset in the rotor core.
INTRODUCTION
T
HE permanent magnet (PM) machine is the perfect
candidate for a drive train application due to a high
efficiency, compactness, fast dynamics and a wide
constant power region. Especially the PM synchronous
machine with fractional slot concentrated windings (FSCW)
is widely used. The advantage of this solution is a short and
complex end-winding, a high slot filling factor, low cogging
torque and a low cost manufacturing process [1].
Of course, there are drawbacks when a PM synchronous
machine is in use. A main problem is the temperature
sensitivity of permanent magnet material. This sensitivity
can lead easily to an irreversible demagnetization of the
permanent magnets. In the application of a hybrid electrical
vehicle (HEV), the thermal load environment is more
complicated. Based on the small package area and a
demanding drive cycle, permanent magnet demagnetization
and insulation failure are considerations in the design of
electric motor cooling [2]. Therefore, an efficient cooling
method is very important to prevent the machine from
damages. The prediction of the temperature distribution in
different parts of the machine can be performed by using a
lumped parameter thermal network [3] or a finite-element
method. The finite-element method gives a more detailed
distribution inside the electrical machine [4]-[5].
To calculate the thermal behavior of the machine
ANSYS Fluent is used. This is a highly sophisticated finite
element modeling and analyzing tool. It can be used to
analyze complex problems in fluid and thermal processes. A
fluent analysis calculates the fluid velocity, flow quantity
and pressure drop. In advance a thermal analysis calculates
Fig. 1. New 24-slots/28-poles motor design. Marked in light blue is the
stator and rotor iron, marked in blue and red are the 28 rectangular
magnets in the rotor, marked in green, blue and red are the winding
topology at the stator
Table 1 shows the main geometry data for the used PM
machine. The presented stator structure uses twelve simple
concentrated coils, twelve stator core modules and also
twelve additional stator teeth components of “T-shape”,
which are used as flux-barriers. These flux-barriers
overcome the drawback of concentrated windings, because
they reduce the air-gap flux-density sub-harmonics.
The obtained results show that the torque capability is
increased compared to a conventional design and the torque
ripples are reduced. With these flux-barriers the reduction
of sub-harmonics leads to a decrease of losses inside the
PM machine. Therefore, the new machine design has a
higher power density and higher efficiency compared to a
standard PM machine. With this new stator design the new
machine can be up to 20% shorter than a conventional
machine design. Therefore, the amount of material for rare
Alexander Nollau is with the Institute for Electrical Drives,
Universitaet
der
Bundeswehr
Muenchen,
85579,
Neubiberg,
Germany
978-1-4799-4775-1/14/$31.00 ©2014 IEEE
1322earth magnets, copper and iron core can be reduced by more
(phone: 0049-89-6004-4416; e-mail: [email protected]).
than 15%.
Dieter Gerling is Head of the Institute for Electrical Drives,
Universitaet der Bundeswehr Muenchen, 85579, Neubiberg, Germany (email: [email protected]).
TABLE I
MAIN GEOMETRY DATA
Outer rotor diameter
Outer stator diameter
Gap length
Active length
Magnet cross sectional area
Turns per coil & Parallel path
Ns – Number of stator slots
P – Number of pole pairs
240 mm
300 mm
1 mm
54 mm
96,75 mm²
8&1
24
14
Based on the challenging requirements for PM machines
in different industry applications such as the automotive
industry, this design brings a huge economic and ecological
benefit for the industry [1].
B.
Design of cooling methods using flux-barriers
The proposed cooling methods are all using the fluxbarriers in the new stator design to enhance the cooling area
and therefore, increase the amount of heat that can be
dissipated from the PM machine.
1)
Standard cooling jacket
The first design is considered to be the conventional
design using a cooling jacket around the stator. Fig. 2 shows
the design of this cooling jacket. It has one inlet and one
outlet with a diameter of 20 mm, respectively. The thickness
of the cooling jacket is 4 mm.
Fig. 3a. Axial flux-barrier cooling Fig. 3b. Flux-barrier marked in yellow
3)
Radial cooling using the flux-barriers
This cooling method is based on the conventional design
with a cooling jacket surrounding the stator. Nevertheless,
there is no sheet between the water jacket and the fluxbarriers. Therefore, the cooling fluid will flow inside the
flux-barriers and increase the cooling area significantly. It
has one inlet and one outlet with a diameter of 20 mm,
respectively. The thickness of the cooling jacket is 4 mm.
The flux-barriers are separated by a 1 mm thick iron sheet
from the air-gap. This ensures that no cooling fluid flows
into the air-gap and causes a short circuit. Fig. 4 shows the
radial cooling design.
Fig. 4. Radial cooling method marked in yellow
Fig. 2. Conventional water jacket marked in yellow.
Between the air in the flux barriers and the cooling jacket
is a thin aluminum plate to separate the two materials from
each other.
To enhance the cooling, several fluid obstacles are
installed above the flux-barriers to create a more turbulent
flow inside the cooling jacket. This will lead to a higher
heat transfer coefficient [6].
4)
Direct air-gap oil-cooling
The last presented method is an oil-cooling solution. This
method also uses the flux-barriers to guide the cooling fluid
to the air-gap. In contrast to the other methods, there is no
This method uses the flux-barriers to cool the machine in
sheet between the flux-barrier and the air-gap. The cooling
an axial direction. The flux-barriers are separated by a 1 mm
fluid can now cool directly the slot area and the rotor with
thick iron sheet from the air-gap. This ensures that no
the buried magnets. The fluid must be a non-electric
cooling fluid can go into the air-gap and causes a short
conduction material. Therefore, the direct air-gap cooling
circuit. Fig. 3a displays the cooling method and Fig. 3b
method uses transmission-oil. Fig. 5 shows the direct airshows the iron sheet that connects the stator modules. This
gap cooling. The disadvantage of this method is the higher
axial cooling has now 12 inlets and 12 outlets with an area
friction losses due to higher viscosity of oil compared to air
of 141.75 mm². Furthermore, there will be no cooling jacket
surrounding the stator which has the advantage of a smaller1323[7].
package area.
2)
Axial cooling using the flux-barriers
The power losses can be calculated analytically or using
FE methods. For each operating condition, the calculated
losses of the electric machine have to be distributed
homogeneously in the components of the PM machine1.
B.
Numerical Simulation
1)
Fig. 5. Direct air-gap oil-cooling. The fluid zone is marked in yellow
III. SIMULATION SETUP
A.
Losses & heat generation
In the finite element model different areas are used for
heat sources (losses) such as ohmic losses in stator
windings, iron losses in the yoke, iron losses in the teeth,
and power losses in the magnets. Table II shows the
material data that were applied for different parts, where k
is the thermal conductivity, cp is the specific heat capacity
and ρ is the density.
TABLE II
THERMAL DATA OF USED MATERIAL
Material
k (W/(m*K))
cp (J/(kg*K))
ρ (kg/m3)
Iron
Copper
Aluminum
NdFebmagnet
Slot
insulation
Air
Water
Engine-oil
50
387.6
210
486
381
896
7600
8978
2707
9
370
7550
0.25
2.2
2000
0.028
0.42
0.145
1013
4.12
1845
0.013
1077
889
Pcopper [W]
PIron_Stator [W]
PIron_Rotor[W]
PMagnet [W]
Peak Power
Average Power
2000
675
75
75
500
168.75
18.75
19
• The SST model incorporates a damped crossdiffusion derivative term in the ω equation.
• The definition of the turbulent viscosity is modified
to account for the transport of the turbulent shear
stress.
• The modeling constants are different.
2)
Transport Equations for the SST k-ω Model
The SST k-ω Model is represented by the following two
equations:
∂
∂
∂
∂k
(ρk ) +
( ρ kui ) =
(Γk
) + Gɶ k − Yk + Sk
∂t
∂xi
∂x j
∂x j
The power losses in different parts of the electrical
machine for different operation conditions are presented in
the Table III. The temperature distribution is studied for
two different power conditions at 2500 rpm. The first
investigation will determine the temperature distribution
under peak power conditions and thereafter, the temperature
distribution under average power conditions. The losses of
an electric machine consist of: stator iron losses, stator
copper losses, rotor iron losses, magnet losses and friction
losses.
Losses
The shear-stress transport (SST) k-ω model was
developed by Menter [8] to effectively blend the robust and
accurate formulation of the k-ε model in the near-wall
region with the free-stream independence of the k-ω model
in the far field. The k-ε model gives an isotropic turbulence,
which is constant in all directions. However, very close to
solid walls the fluctuations in the turbulence vary greatly in
magnitude and direction. Therefore, the turbulence cannot
be considered to be an isotropic one [5]. To achieve this,
the k-ε model is converted into a k-ω formulation. The
SST k-ω model is similar to the standard k-ω model, but
includes the following refinements:
These features make the SST k-ω model more accurate
and reliable for a wider class of flows than the standard k-ω
model. The reason is that the equations can be integrated
directly to the wall with no need for specialized wall
functions and no special conditions are required at solid
boundaries [9].
Units: W = watt, m = meter, K = kelvin, J = joule, kg = kilogram
TABLE II
POWER LOSSES FOR DIFFERENT OPERATING CONDITIONS
Shear-stress Transport (SST) k-ω Model
(1)
and
∂
∂
∂
∂ω
( ρω ) +
( ρωu j ) =
( Γω
) + Gɶ ω − Yω + Dω + Sω (2)
∂t
∂x j
∂x j
∂x j
In these equations, Gɶ k represents the generation of
turbulence kinetic energy due to mean velocity gradients,
and describes the modeling the turbulence production. Gω
represents the generation of ω, calculated as described for
the standard k-ω model in [10]. Γ k and Γω represent the
effective diffusivity of k and ω, respectively. Yk and Yω
represent the dissipation of k and ω due to turbulence. Dω
represents the cross-diffusion term, which blends the two
models together. Sk and Sω are user-defined source terms.
The 3D simulations presented in this paper are conducted
by
1324 solving the equations using the commercial software
1
The loss calculation for this machine is presented in a separated
paper at ICEM 2014
package Ansys FLUENT. Discretization of the transport
equation utilizes the second-order upwind scheme for all
equations. The COUPLED scheme is used as pressurebased solver. The coupled algorithm solves the momentum
and pressure-based continuity equations together. The full
implicit coupling is achieved through an implicit
discretization of pressure gradient terms in the momentum
equations, and an implicit discretization of the face mass
flux, including the Rhie-Chow pressure dissipation terms.
Convergence is achieved when the scaled residuals reach
10-4 and the mass flow rate difference is minimal, which
typically took about 450 iterations.
3)
Boundary Conditions
For a simple thermal analysis, the slot region is modeled
with a homogeneous material (copper), insulated with an
equivalent insulation layer. This insulation layer is
characterized by an equivalent thermal conductivity, keq,
which takes into account the thermal conductivity of the slot
insulation layer, the air-gap between the slot insulation and
the laminations, the insulation varnish of the windings and
the air-gaps between the conductors. It must be mentioned
here, that the simulations are made for an ideal case.
Therefore keq of the slot insulation layer is taken to be equal
to the thermal conductivity of the slot insulation material,
keq = 0.25 (W/(m*K)).
Furthermore, a heat convection rate is applied on the
inner rotor diameter to model the heat transfer to the rotor
shaft with k= 20 (W/(m*K)). All the other walls are taken to
be adiabatic and therefore, are not dissipating heat.
Simulations of the steady-state temperature distribution are
made using initial conditions. The initial temperature for all
elements of the machine is taken to be 0 °C.
A steady mass flow rate inlet condition is applied on the
inlet of the cooling system. The mass flow rate is taken to
be 10 l/min for all methods to ensure comparable accuracy
[11]. For the flow outlet the standard pressure outlet
conditions were used.
4)
temperature in the magnets is 186.08 °C. The pressure drop
of this coolant system is 2987.49 Pa.
Fig. 6. Temperature distribution inside the PM machine.
2)
Fig. 7 shows the result for the simulation under peak
power conditions. The volume-weighted average static
temperature in the slot region is 296.44 °C and with a hot
spot temperature of 297.87 °C. The maximum static
temperature in the magnets is 192.01 °C. The pressure drop
of this coolant system is 55.47 Pa.
Assumption
Fig. 7. Temperature distribution inside the PM machine.
The following assumptions are made to simplify the
numerical simulation:
• The ambient temperature is constant
• The heat losses are considered to be
homogenoulys distributed
• There is no influence of temperature rise on the
thermal property of materials
• The insulation material is evenly distributed
and the impregnation is good.
Axial cooling using the flux-barriers
3)
Radial cooling using the flux-barriers
Fig. 8 shows the result for the simulation under peak
power conditions. The volume-weighted average static
temperature in the slot region is 260.60 °C and with a hot
spot temperature of 262.60 °C. The maximum static
temperature in the magnets is 173.40 °C. The pressure drop
of this coolant system is 21197.9 Pa.
IV. SIMULATION RESULTS
A.
Peak Power Rating
1)
Standard cooling jacket
Fig. 6 shows the result for the simulation with the
previously mentioned boundary conditions. The volumeweighted average static temperature in the slot region is
1325
275.22 °C and with a hot spot temperature of 293.63 °C
under peak power conditions. The maximum static
Fig. 8. Temperature distribution inside the PM Machine.
4)
Direct air-gap oil-cooling
Fig. 9 shows the result for the simulation under peak
power conditions. The volume-weighted average static
temperature in the slot region is 274.41 °C and with a hot
spot temperature of 278.28 °C. The maximum static
temperature in the magnets is 94.58 °C. The pressure drop
of this coolant system is 52222.06 [Pa].
variation between the average temperature and the hot-spot
temperature is decreased from 18 K to 5 K for the standard
cooled machine and the radial cooled machine, respectively.
The disadvantage is the higher pressure drop that can be
explained by the integrated flow obstacles in the cooling
jacket. These flow obstacles will enhance the turbulence
generation and thus, the heat transfer coefficient of water.
Fig. 10. Average temperatures for each part of the machine and different
cooling methods
Fig. 9. Temperature distribution inside the PM Machine
B.
Average Power Rating
The results for the average power condition are displayed
in Table IV.
TABLE IV
SIMULATION RESULTS FOR AVERAGE POWER CONDITIONS
Volumeweighted average
temperature slot
region (°C)
Hot-spot
temperature
slot region
(°C)
Hot-spot
temperature
magnet (°C)
Standard
cooling
jacket
135.72
140.25
91.78
Axial
cooling
139.96
140.29
93.34
Radial
cooling
132.83
133.38
89.66
135.70
136.67
89.66
Method
Direct airgap cooling
C.
Comparison
Fig. 11. Average temperatures for each part of the machine and different
cooling methods
The axial cooling is the worst method compared to the
standard cooling method. This can be explained by the
small transfer area between the cooling fluid and the stator.
Although, the pressure drop is 50 times lower than the
standard cooling method, the average slot temperature is 21
K higher. Also for average power condition the temperature
is in every part of the machine higher by 3-5 K. Therefore,
this method can only be a solution to a small package
operation environment. This can be a case in an industrial
environment with average power ratings.
The direct air-gap oil-cooling method is the last
presented method. This method can also be an alternative
for a small package environment, because of the sufficient
cooling of the magnets and the rotor. The magnet material is
very sensitive to temperature, which lead easily to an
irreversible demagnetization of the permanent magnet. The
magnets are up to 50% cooler than with a standard cooling
jacket under peak power conditions. Nevertheless, the
drawbacks of this solution are the high pressure drop and
the increased friction loss due to the viscosity of oil that
flows in the air-gap region.
Fig. 10 and 11 show the graphic comparison between the
proposed cooling methods under peak power setting and
average power setting. The volume weighted average
temperatures are displayed for each part of the machine.
The results show clearly that the flux-barriers filled with
a cooling fluid can be used as an effective cooling method.
Especially the radial cooling with the usage of the flux
barrier is the best method. The advantage is that the cooling
area is increased and therefore, more heat can be dissipated
from the stator. Also the cooling fluid flows closer to the
V. CONCLUSION
slot region where the most heat will be generated. Thus, the1326
hot spot temperature is up to 10% cooler than with a
This paper uses a finite-element method (FEM) to model
standard cooling jacket. Another advantage is that the
the thermal distribution inside a PM machine. The studied
machine is a 24-Teeth/28-Poles PM machine. Based on the
results in the previous sections, the following conclusions
can be stated: The radial cooling with the usage of the fluxbarriers can be an effective alternative to cool a PM
machine more sufficiently. At peak power conditions the
temperature in the slot region is up to 10 % cooler than with
a standard cooling jacket. On the other side, the pressure
drop is higher to pump the cooling fluid through the fluxbarriers. Another alternative can be the direct air-gap oilcooling which cools down the magnet and rotor very
effectively. The magnets are up to 50% cooler than with a
standard cooling jacket under peak power conditions. All
cooling methods are only varying a few degrees Celsius
under average power conditions.
VI. ACKNOWLEDGEMENT
The authors gratefully acknowledge the contributions of
G. Dajaku and S.Spas.
VII.
[1]
[2]
[3]
[4]
[5]
[6]
Ping Zheng; Ranran Liu; Thelin, Peter; Nordlund, Erik;
Sadarangani, C., "Research on the Cooling System of a 4QT
Prototype Machine Used for HEV", IEEE Transactions on Energy
Conversion, vol.23, no.1, pp.61-67, March 2008
doi: 10.1109/TEC.2007.914356
[7] Qi Wenjuan, Zou Jibin, Li Jianjun, "Numerical calculation of
viscous drag loss of oil-filled BLDC motor for underwater
applications", International Conference on Electrical Machines and
Systems (ICEMS), 2010, pp.1739-1742, 10-13 Oct. 2010
[8] F. R. Menter. "Two-Equation Eddy-Viscosity Turbulence Models
for Engineering Applications", AIAA Journal, vol.32, no.8, pp.
1598–1605, August 1994
[9] E.Savory, R.J. Martinuzzi, J.Ryval, Z. Li and M. Blissitt,
“Evaluationof the Thermofluid Performance of an Automotive
Engine Cooling-Fan System Motor”, Proceedings of the Institution
of Mechanical, Part D: Journal of Automobile Engineering, January
1, 2011 225: 74-89, doi:10.1243/09544070JAUTO1416
[10] Ansys FLUENT User Guide
[11] Zhe Huang; Nategh, S.; Alakula, M.; Lassila, V.; Jinliang Yuan,
"Direct oil cooling of traction motors in hybrid drives", 2012 IEEE
International Electric Vehicle Conference (IEVC), pp.1-8, 4-8
March 2012, doi: 10.1109/IEVC.2012.6183163
VIII.
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G. Dajaku, D.Gerling, "Low Costs and High-Efficiency Electric
Machines," 2nd International Electric Drives Production
Conference 2012, EDPC-2012, Erlangen-Nürnberg, Germany, 2012.
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Analysis of Electric Machines for Hybrid Electric Vehicle
Application”, The 22nd International Battery, Hybrid and Fuel Cell
Electric Vehicle Symposium & Exposition, Yokohama, Japan, 2006.
D. Gerling, G. Dajaku: “Thermal calculation of systems with
distributed heat generation”, The Tenth Intersociety Conference on
Thermal and Thermomechanical Phenomena in Electronics
Systems, ITHERM '06, San Diego, CA, 2006.
A. Nollau, D. Gerling, "A new cooling approach for traction motors
in hybrid drives", 2013 IEEE International Electric Machines &
Drives Conference (IEMDC), pp.456-461, 12-15 May 2013.
doi: 10.1109/IEMDC.2013.6556136G.
Z. Kolondzovski, "Numerical modelling of the coolant flow in a
high-speed electrical machine", 18th International Conference on
Electrical Machines, 2008 (ICEM 2008), pp.1-5, 6-9 Sept. 2008
doi: 10.1109/ICELMACH.2008.4799884
Alexander Nollau was born in Dresden, Germany, on June 20, 1987. He
received the M.Sc. degree in electrical engineering from the Universität
der Bundeswehr München (University of Federal Defense Munich,
Germany). During his studies he had multiple stays at the Technical
University Dresden and the Lawrence Berkeley National Laboratory, USA.
He is currently with the Universität der Bundeswehr München, Chair of
Electrical Drives and Actuators and is working on his Ph. D. thesis in the
field of cooling electrical machines for traction drives.
Dieter Gerling Born in 1961, Prof. Gerling got his diploma and Ph.D.
degrees in Electrical Engineering from the Technical University of
Aachen, Germany in 1986 and 1992, respectively. From 1986 to 1999 he
was with Philips Research Laboratories in Aachen, Germany as Research
Scientist and later as Senior Scientist. In 1999 Dr. Gerling joined Robert
Bosch GmbH in Bühl, Germany as Director, being responsible for New
Electrical Drives and New Systems. Since 2001 he is Full Professor and
Head of the Institute of Electrical Drives at the University of Federal
Defense Munich, Germany.
1327
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