High performance cooling system for automotive inverters - IEEE Xplore

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and the Innovative electronics Manufacturing Research Centre (IeMRC) for their support of this work. Keywords. « Automotive application », « Cooling » ...

High performance cooling system for automotive inverters Cyril Buttay1 , Jeremy Rashid2 , C. Mark Johnson1 , Peter Ireland3 , Florin Udrea2 , Gehan Amaratunga2 , Rajesh K. Malhan4 1 School

of Electrical and 2 Univ. of Cambridge Electronic Engineering Engineering Department Univ. of Nottingham (Electrical Eng. Division) University Park 9 JJ Thompson Ave. Nottingham NG7 2RD Cambridge CB3 0FA UK UK

3 Department

of Engineering Science Oxford University Park Road Oxford OX1 3PJ, UK

4 DENSO CORPORATION Research Laboratories Nisshin-shi Aichi 470-0111 Japan

[email protected]

Acknowledgments The authors would like to thank the UK Engineering and Physical Sciences Research Council (EPSRC) and the Innovative electronics Manufacturing Research Centre (IeMRC) for their support of this work.

Keywords « Automotive application », « Cooling », « Power converters for HEV », « Packaging », « thermal design ».

Abstract A novel double-side cooled power module is presented which delivers superior cooling performance with the potential for improved robustness to thermal cycling. The semiconductor dies are sandwiched between conventional DBC substrates, the substrates being directly cooled rather than through a conventional heat spreader heat sink assembly. A theoretical analysis is presented illustrating that direct cooling can offer a lower total thermal resistance provided the heat transfer coefficient at the cooled surface is sufficiently high. Experimental results demonstrate the effectiveness of the selected impingement cooling technique when applied in both single- and double-side cooled formats. Measurements on the double-side cooled structure show a total thermal resistance (junction to ambient) that is less than 40% of the junction to case resistance of a conventional module. Similar improvements are observed in the transient thermal impedance (step response) curve indicating that thermal cycling ranges will be reduced under all operational conditions

1

Introduction

Hybrid Electric Vehicles (HEVs) include both an Internal Combustion Engine (ICE) and one or several electrical motors. Due to the presence of the ICE, the ambient temperature in the engine compartment is usually pretty high, and can be more than 100°C in some locations [1], with some strong thermal cycling [2]. In such a high temperature ambient, the power electronic converters that drive the electrical motors require efficient thermal management since the electronic components have a limited maximum temperature, typically around 150 to 175°C for silicon-based semiconductors and 150°C for high-temperature ceramic capacitors. Therefore, the thermal resistance between the active part of a component and the environment has to be kept to a minimum. The structure of a classical power module is described in figure 1(a). As can be seen, heat is extracted from the silicon device through a stack of many layers of different materials. This structure shows poor

Figure 1: Structure of a a classical power module and b the proposed sandwich structure

reliability to thermal cycling, as each layer expands differently with temperature. It has been shown that the interfaces, especially the copper/baseplate and the die/wirebond are very sensitive to this phenomenon [3], [4] A so-called “sandwich” structure is presented in figure 1(b). In this structure, the device is connected to two Direct Bonded Copper (DBC) substrates, one on top of the device, and one on bottom. Therefore, the heat generated by the device can be extracted from both sides. Furthermore, this structure is cooled by spraying water directly to the back of the DBC substrates[6], which simplifies the thermal stack, reducing the number of layers, hence, the thermal resistance of the whole. Finally, compared to other sandwich structures [5], a specific mechanical arrangement of the substrates and dies has been chosen such that a uni-axial compressive force is applied to the dies during operation. This improves resistance to fatigue of the interconnects through thermal cycling. In this paper, we present in detail this sandwich structure and the impingement cooling system.

2

Principle

2.1 Direct Cooling

(a)

(b)

Figure 2: In the standard water-cooled structure (a), the substrate is attached to a baseplate, itself secured to a waterplate. The water flows through paths machined in the plate, and carries the heat away to an external heatsink. In the direct-cooled structure (b) we spray the coolant directly onto the back of the substrate.

The current state-of-the-art cooling of power semiconductors is depicted in figure 2(a). In this structure, the baseplate of the power module is bolted to a water plate. An intermediate layer of thermal paste is used to compensate for surface roughness and make good thermal contact. As the heat is carried through the material of the water-plate, this material has to offer a low thermal resistance. This is why metals (especially copper or aluminium) are usually employed. This results in a heavy structure (the baseplate is 2 to 5-mm thick, and the water plate is more than 1-cm thick), with most of the material used only to carry the heat. In the direct-cooled structure (figure 2(b)), the water is in direct contact with the back of the power module substrate. Using this method, the heat-exchanger can be made out of any material (even plastics), as it is only used to guide the water, not for heat transport. Also, as there is no need to compensate for surface roughness, the substrate can be mounted onto the heat-exchanger with little clamping force (just

enough to seal properly). In this case, DBC substrates can have enough mechanical strength to be used without a baseplate. In addition to giving mechanical strength to the package, the baseplate is also used as a heat-spreader, to increase the active area in contact with the heatsink. On the other hand, if no heat-spreading is needed (i.e if the heat-exchanger is efficient enough), using a baseplate merely adds two layers (solder and the baseplate itself), and increases rather than reduces the thermal resistance.

Figure 3: Simulated structure and parameters. The ambient temperature is 20°C.

To analyse the effect of the baseplate on the thermal resistance, the structure in figure 3 has been simulated using FreeFEM ++[7]. Please note that 3-dimensional effects (e.g corners of the die), as well as non-linearity (variation of thermal conductivity with temperature, phase change of the coolant) are not taken into account in this simple 2-D model. Simulations have been carried out on two structures (with and without baseplate), with two widths (10 mm and 50 mm, see figure 3), and various values of heat transfer coefficient on the back, to represent the effectiveness of the heat removal.

(a)

(c)

(b)

(d)

IsoValue 20 30 40 50 60 70 80 90 100 110 120 130 140 150

Figure 4: Temperature plot obtained from finite-element simulations of the structure presented in figure 3, with a cell pitch of 20 mm (corresponding to the results in figure 5(b)). The DBC substrates are attached to a baseplate in (a) and (c), while they are in direct contact with the coolant in (b) and (d). An exchange coefficient of 8 kW.m−2 .K −1 is considered in figures (a) and (b). In figures (c) and (d), it is 40 kW.m−2 .K −1 .

Some of the results on the 10 mm-wide structure can be seen in figure 4. In figures 4(a) and 4(b), an exchange coefficient of 8 kW.m−2 .K −1 is applied on the bottom of the structure. This value corresponds to the lower end of the direct cooling performance (10 to 40 kW.m−2 .K −1 )[6]. In these conditions, the maximum temperatures on top of the die are 124°C (a) and 134°C (b), both with a coolant temperature of 20°C. In figures 4(c) and (d), a better heat-exchange coefficient is considered (40 kW.m−2 .K −1 ). From this, it is obvious that the overall temperatures are lower when the heat removal is more efficient. Furthermore, the hottest points in figures 4(c) and 4(d) are identical (73°C), which indicates that the spreading effect of the baseplate is no longer needed.

3

without baseplate with 2mm-thick copper baseplate Thermal resistance [K/W]

Thermal resistance [K/W]

3 2.5 2 1.5 1 0.5

without baseplate with baseplate

2.5 2 1.5 1 0.5

0

10

20

30

40

50

60 2

Exchange coefficient[kW/m /K]

(a)

70

80

0

10

20

30

40

50

60

70

80

2

Exchange coefficient[kW/m /K]

(b)

Figure 5: Comparison of direct cooling performance with a base plate and without (2-D FEM simulations) as a function of the exchange coefficient. In (a), we consider a large area for heat spreading (100 mm pitch between chips), while in (b), the pitch is 20 mm, closer to the actual configuration.

The evolution of the junction-to-ambient thermal resistance with the exchange coefficient value is given in figures 5(a) (50 mm-width for the structure in figure 3) and (b)(10 mm). In the case of a 50 mm-wide structure (which can be considered as a semi-infinite condition as the halfdie is only 2.5 mm-wide), there is plenty of room for the heat to spread, hence the better performance of the baseplate module at the lower exchange coefficient values. However, when this exchange coefficient increases, most of the heat is drained vertically, with little spreading effect. Thus, the baseplate becomes first unnecessary (around 40 kW.m−2 .K −1 ), and then a burden, as the added layers are responsible for an increase in RT H compared to the baseplateless design. In the case of a 10 mm-wide structure (close to the layout in the module that is presented below), the advantage of using a baseplate becomes marginal (figure 5(b)), even at the lowest exchange coefficient values: there is indeed little room to spread the heat over. From these simulations, it can be concluded that, providing the solid-to liquid heat transfer is efficient enough, it is possible to get rid of the baseplate without any loss in thermal performance. This results in a reduction of weight (the baseplate is by far the heaviest part of the structure). This should also translate in an increase in reliability, as there are less layers in the thermal stack.

2.2 Double-Side Cooling It can be seen from figure 5 that for the highest values (above 50 kW.m−2 .K −1 ), the exchange coefficient has only a small influence on the thermal resistance. This means that in this area, the substrate itself is the limiting factor. Further decrease of RT H requires to change either the materials (especially replacing the alumina layer with, for example, aluminium nitride), or the structure. An obvious solution to improve cooling is to use both sides of the die (see figure 1(b)). This requires some changes in the electrical connections of the devices, as it is then no longer possible to use wirebonds, as in figure 1(a)), but it doubles the effective area of the heat dissipation path, theoretically allowing for a twofold reduction in thermal resistance. A so-called “sandwich” structure, in which the power dies are soldered between two DBC substrates has been built and is shown in figure 6. The design of this structure is explained in details in . The DBC substrates use a thin layer of alumina (0.3 mm) as the dielectric, and thick (0.6 mm) layers of copper with a two-level etching to form mechanical features. Three sandwich structures such as those presented in figure 6 sit between two heat-exchangers (see figure 7(a)) to form a three-phase inverter. In this structure, the heat-exchangers also provide direct cooling to the driver and capacitor boards. The hole matrix that can be seen in the heat-exchangers in figure 7(a) is employed to spray the cooling water onto the power modules, increasing the exchange coefficient.

Figure 6: A prototype sandwich structure. This module is a half-bridge, with 2 diodes and 2 IGBTs. Note the ridges, obtained by a two-level etching of the top copper layer, that improve the sealing with the heat-exchangers.

(a)

(b)

Figure 7: (a) an exploded view of an integrated inverter, using the proposed cooling system. The actual inverter is shown in (b)

In the proposed system, the power modules are not bonded directly to the heat-exchanger. They are loosely clamped, and a silicon rubber bead (about 0.5mm thick) provides the sealing required. This way, virtually no mechanical stresses are transmitted between the modules and the heat-exchangers, even if material with very different coefficients of thermal expansion are used. In the case of the inverter of figure 7(b), heat exchangers have been made in various materials, including brass, aluminium, carbonloaded polymer and glass-loaded nylon.

3

Results

Two heat-exchangers, specifically designed for cooling performance measurement have been made (figure 8). Compared to those presented in figure 7(b), they are designed to cool one power module only, without the capacitor board and the drivers, and have one inlet and outlet per plate, instead of one set for both.

3.1 Pressure drop The pressure drop versus flow-rate plot of the single-module test heat-exchanger is shown in figure 9. It can be seen that the heat-exchangers offer a moderate pressure drop, even at relatively high flowrates. Furthermore, these results are obtained for a double-side cooling where both heat-exchangers are connected in series. It is theoretically possible to divide the pressure drop by a factor of 4 by connecting the heat-exchangers in parallel. At the moment, the heat-exchangers are subdivided into two series-connected cells. Again, it is possible to reduce the pressure drop by connecting these two cells in parallel. Finally, a complete inverter uses

Figure 8: Test system, designed for cooling performance measurements. This hosts a single power module, and sensors allows for measurement of flow-rate, pressure drop, and temperature of the cooling fluid. 1.6 P = 0.15 F2

1.4 Pressure drop [bar]

1.2 1 0.8 0.6 0.4 0.2 0 -0.2 0

0.5

1

1.5 2 Flowrate [l/min]

2.5

3

3.5

Figure 9: Pressure drop vs flow rate for a one-module, double side heat-exchanger.

three power modules, whose heat-exchangers can be connected either in series or parallel. Therefore, this cooling technique offers a lot of flexibility to accommodate the requirements of the cooling circuit.

3.2 Cooling performance 3.2.1

Thermal Resistance

A specific test setup is used to measure the cooling performance of the sandwich structure. It works as follows: one of the devices of the power module is connected to a high current source (here 40 A through one of the SiC SBDs) to make it self-heat. Then this high-current source is disconnected and the device is forward biased by a low-current source (here 50 mA). The forward voltage across the device is then monitored, as this voltage is an image of the temperature in the semiconductor die. Previous calibration measurements have shown a dependency of -1.8 mV/°C on the forward voltage of the diodes used at a biasing current of 50 mA. The current and voltage across the device are also monitored during the power pulse, so it is possible to calculate the power dissipated and then the thermal resistance of the package: RT H ja =

∆T P

(1)

The results are listed in table I, for three test conditions: double and single side cooling with impingement, and double side cooling without impingement (the spray plate that creates the jet effect is removed). As expected, the double-side cooling with impingement is the most efficient, but not as far as twice as good than the single-side cooling. This can be explained by two reasons:

• the structure is not completely symmetric: the back contact of the dies has a larger area than the top contact (5×5 mm versus 3×3) • when measuring the single-side cooling performance, some of the heat can escape through the top surface of the sandwich package.

Table I: Thermal resistance, junction to ambient [K/W]

1 l/min 2 l/min 3 l/min

0.8 0.7 Thermal impedance [K/W]

single-side cooling 0.66 0.6 0.58

Transient Thermal Impedance

0.6 0.5

0.6

0.5 l/min 1 l/min 1.5 l/min 2 l/min 2.5 l/min 3 l/min

Thermal impedance [K/W]

3.2.2

double-side cooling 0.45 0.39 0.36

0.4 0.3 0.2 0.1 0 -0.1

0.5 0.4

0.5 l/min 1 l/min 1.5 l/min 2 l/min 2.5 l/min 3 l/min

0.3 0.2 0.1 0 -0.1

0.0001

0.001

0.01 Time

(a)

0.1

1

10

0.0001

0.001

0.01 Time

0.1

1

10

(b)

Figure 10: Transient thermal impedance (step response), junction to ambient, for a diode, as a function of the flow-rate of coolant for single-side (a), and double-side (b) direct cooling.

Using the test setup, it is also possible to monitor the cooling-down transient of the die, and then compute the step response, as shown in figure 10. Both single and double-side cooling exhibit the same behaviour up to 0.1s, because at the shortest time scales the thermal impedance is dictated by the package itself (mainly by the thermal properties of the dies and the DBC substrates). In the slower time scales, as we approach steady state, the thermal impedance becomes more dependant on the efficiency of the impingement mechanism. This is where the difference between single and double-side cooling appears clearly, although not with a 1:2 ratio, because of the reasons exposed above. It is also visible from figure 10 that the presented structure has a very fast response, as it takes about 1 s for the diode to reach the ambient temperature (in a classical power module, attached to a heatsink, this time would be more than 10 s). This faster response in itself may not be desirable under all circumstances as the “thermal buffering” effect of the base-plate in classical power modules is often used to absorb short duration transients, for example during over-current conditions. However, as is evident from figure 11, the sandwich module thermal step response is lower than that of the classical package for all pulse durations. Thus it has an improved ability to deal with transients as well as lower steady state thermal resistance. From the manufacturer¢s datasheet [9], the thermal resistance, junction-to-case, of an equivalent IGBT in a classical package is 1 K/W, whereas the thermal resistance, junction-to-ambient, of the presented package can be lower than 0.4 K/W. Note that the junction to ambient thermal resistance of the classical package must also incorporate the resistances of the heatsink and of the thermal grease, so it is possible to conclude that the presented package offers an overall thermal resistance which is at most a third of that of the classical package

Sandwich Classical package

Thermal resistance [K/W]

1 0.8 0.6 0.4 0.2 0 0.0001

0.001

0.01

0.1

1

10

1

Time

Figure 11: Comparison of the transient impedance, junction to ambient, of the sandwich package with a 5×5 mm SiC diode (measured, with a flow rate of 3 l/min) and the transient impedance, junction to case, of a Si-IGBT of the same size, in a classical package with a baseplate (values from the datasheet [9])

3.3 Temperature Increase of the Cooling Fluid

Dissipated Power [W]

70 60

1.6

Power (calorimetric measurement) Power (electric measurement) ∆T

1.4 1.2

50

1

40

0.8

30

0.6

20

0.4

10

0.2

0

∆T [K]

80

0

2

1.

1

0.

in

l/m

l/m

5

in

l/m

l/m

5

in

in

Figure 12: Measurement of the power dissipated by one of the diodes, using calorimetric and electric methods. The difference is the power dissipated in the ambient air surrounding the heat-exchanger. ∆T is the difference between the temperature of the coolant between the inlet and the outlet.

As shown in section 3.1, the presented cooling technique works with a fairly low pressure-drop and high flow-rate. One of the main advantages is that the temperature difference between the inlet and the outlet of the heat-exchanger is kept to a minimum. This can be seen on figure 12, where a ∆T as low as 0.4 K is achieved under 2 l/min and a power dissipation of almost 60 W. A low ∆T is important to ensure there is no hotspot in the system, which might cause reliability issues (due to mechanical stresses), or electrical unbalance (when different devices of the systems work at different temperatures). The power dissipated by the device (measured from the voltage drop and the current flowing through the diode) is also plotted in figure 12. It is compared to the power being actually removed by the cooling fluid (obtained from ∆T and the flow-rate). As can be seen, above 1 l/min, almost all heat is carried by the fluid. This means that there is negligible convection losses from the heat-exchanger walls, and thus that the heat-exchanger can be made of any material. At the lowest flow-rate, however, a non negligible part of the dissipated power is taken away through convective effects from the heat-exchanger itself (the test heat-exchanger is made out of brass, which is a pretty good thermal conductor).

Conclusion A novel double-side cooled power module has been presented which delivers superior cooling performance with the potential for improved robustness to thermal cycling. Analysis of conventional and novel package structures has identified that direct cooling (i.e. without a heat-spreader or base-plate) can offer a lower total thermal resistance provided the heat transfer coefficient at the cooled surface is sufficiently high. The impingement cooling method chosen for the presented module provides direct cooling of the substrate tiles, resulting in a lightweight and compact inverter that is compatible with the automotive cooling circuit in term of flow-rate and pressure drop. Experimental results, showing the effectiveness of the cooling technique when applied in both singleand double-side cooled formats, have been presented. Measurements on the double-side cooled structure show a total thermal resistance (junction to ambient) that is less than 40% of the junction to case resistance of a conventional module. Similar improvements are observed in the transient thermal impedance

References [1] J. G. Kassakian and D. J. Perreault, “The future of electronics in automobiles,” in Proceedings of the ISPSD’2001 Conference, Osaka, Japon, May 2001, pp. 15–19. [2] The changing automotive environment: high-temperature electronics, Johnson, R.W.; Evans, J.L.; Jacobsen, P.; Thompson, J.R.; Christopher, M, IEEE Transactions on Electronics Packaging Manufacturing, Volume 27, Issue 3, July 2004 Page(s):164 - 176 [3] Thermal fatigue resistance evaluation of solder joints in IGBT power modules for traction applications, Thebaud, J.-M.; Woirgard, E.; Zardini, C.; Sommer, K.-H., Power Electronics Specialists Conference, 2000. PESC 00, Volume 3, 18-23 June 2000 Page(s):1285 - 1290 vol.3 [4] Three-Dimensional Packaging for Power Semiconductor Devices and Modules, Calata, J.N.; Bai, J.G.; Xingsheng Liu; Sihua Wen; Guo-Quan Lu, IEEE Transactions on Advanced Packaging, Volume 28, Issue 3, Aug. 2005 Page(s):404 - 412 [5] Double-sided cooling for high power IGBT modules using flip chip technology, Gillot, C.; Schaeffer, C.; Massit, C.; Meysenc, L., IEEE Transactions on Components and Packaging Technologies, Volume 24, Issue 4, Dec. 2001 Page(s):698 - 704 [6] Cooling jets, Chambers, Andrew; Dailey, Geoffrey; David R. H; Ireland, Peter; patent, US application number 20050074325, 22 sep 2004, available online at http://appft1.uspto.gov/netahtml/PTO/ search-bool.html [7] Hecht F., Pironneau O., Le Hyaric A., Ohtsuka K., FreeFem++ Manual, available from http://www. Freefem.org. [8] Compact Double-Side Liquid-Impingement-Cooled Integrated Power Electronic Module, C. M. Johnson, C. Buttay, S. J. Rashid, F. Udrea, G. A. J. Amaratunga, P. Ireland and R. K. Malhan, Proceedings of ISPSD’07, may 2008, Jeju, Korea. [9] GB15RF120K, IGBT PIM module, datasheet from International Rectifier, available online: http://www. irf.com/product-info/datasheets/data/gb15rf120k.pdf

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