# Improving thermal management of SiC MOSFET through cold plates and/or power modules optimization

W Cherief, A Casado-Ramoneda, J Giolat, R Khazaka, Y Avenas, M Debeze

► To cite this version:

W Cherief, A Casado-Ramoneda, J Giolat, R Khazaka, Y Avenas, et al.. Improving thermal management of SiC MOSFET through cold plates and/or power modules optimization. The 26th European Conference on Power Electronics and Applications, GDR SEEDS France & EPE Association, Mar 2025, Paris, France. 10.34746/epe2025-0121. hal-05097167

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Submitted on 4 Jun 2025

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# Improving thermal management of SiC MOSFET through cold plates and/or power modules optimization

W. Cherief<fnref n="1" />, A. Casado-Ramoneda<fnref n="2" />,<fnref n="3" />, J. Giolat<fnref n="1" />, R. Khazaka<fnref n="2" />, Y. Avenas<fnref n="3" />, M. Debeze<fnref n="1" />

## Keywords

«Brazing cold plate», «Silicon carbide», «Heat spreader», «optimized cold plate», «Power module».

## Abstract

An experimental and numerical investigation on the impact of internal heat spreader and optimized cold plate for cooling SiC power module is presented. Silver-diamond (AgD) composites are used as heat spreaders under the SiC chips, while the cold plates are fitted with specific geometries to enhance convective heat transfers. This study is conducted for a flow regime varying from 500 < Reynolds < 8000. Thermal results show that the use of heat spreaders induces a 20% reduction of Rth(j-a). Similarly, an optimized cold plate design would reduce the cold plate thermal resistance by more than 55%. The study demonstrates that the heat spreader efficiency is better in the case of the use of an optimized design cold plate.

## Introduction

The fast development of wide band gap devices, such as SiC MOSFETs, induces an increase of thermal constraints in power modules. Compared to Si IGBTs, SiC MOSFETs have a lower area for the same current and voltage ratings [1], hence, if the same packaging technology is maintained, the power density of the module is improved but the heat flux densities are increased.

Fig. 1 presents the typical temperature drops in a classical power module that includes chips, a metallised ceramic substrate, a baseplate, a thermal interface material (TIM) and attachment materials. In the case of CAS325M12HM2 SiC power module from CREE, McPherson et al. [2] study the proportional contribution of the

different layers of the power module regarding its total junction to heatsink thermal resistance. Fig. 1b displays that, apart from chips, each element of the assembly has a non-negligible contribution on the total temperature drop in the assembly.

<figure><img src="image_2.png" /><figcaption>Fig. 1: Classical power module – (a) typical assembly – (b) inner temperature drops – in brackets, the percentage of the total temperature drop from [2]</figcaption></figure>

To improve the thermal management of power modules, it is therefore important to use materials with high thermal conductivity. The goal is to reduce conductive thermal resistance. For instance, solders under the chips can be replaced by silver or copper sintered joints [3]. Also, ceramic materials with high thermal conductivity like aluminium nitride (AlN) are of interest. Other authors propose to integrate solid high conductivity materials [4] or two-phase devices

<footnote n="1">¹ MERSEN France La Mure SAS, 38350 La Mure, France</footnote>

<footnote n="2">² Safran Tech, Electrical & Electronic Systems Research Group, 78772 Châteaufort, France</footnote>

<footnote n="3">³ Univ. Grenoble Alpes, CNRS, Grenoble INP, G2Elab, 38000 Grenoble, France</footnote>

<footnote>Email: yvan.avenas@g2elab.grenoble-inp.fr</footnote>

<footnote>wahid.cherief@mersen.com</footnote>
[5] to optimize the thermal conduction inside the base plate.

However, for high temperature or high reliability applications, it is not always possible to choose materials with the highest thermal conductivity. For instance, active metal brazing (AMB) substrates show better mechanical behaviour than direct bonded copper (DBC), but the thermal conductivity of silicon nitride (Si₃N₄) is lower than the one of AlN [6]. Moreover, to reduce thermo-mechanical stresses due to large amplitude thermal cycling, copper baseplates are often replaced by materials with lower thermal expansion coefficient (CTE) like AlSiC [7] or aluminium-graphite (AlG) [8]. However, these materials often have reduced thermal conductivities compared to copper. Therefore, for these applications, it is necessary to find other ways to reduce the junction temperature.

In this paper, it is proposed to evaluate the integration of high thermal conductivity silver-diamond (AgD) composites as heat spreaders directly under the chips as presented in Fig. 2. With this technique, the temperature drops inside the lower thermal conductivity parts of the power module, solders, Si₃N₄ ceramic, baseplate and TIM could be reduced effectively. Moreover, the low CTE of these composites is of great interest for high temperature and high reliability applications.

<figure><img src="image_3.png" /><figcaption>Fig. 2: Power module with AgD thermal spreader under the chips</figcaption></figure>

Finally, to cancel the temperature drop due to the thermal interface material (TIM), it is also possible to use direct liquid cooling. In [2], the authors show that the total thermal resistance of SiC power modules can be largely reduced using this technique. Another important way to reduce the junction temperature is to use double side cooling as presented in [9]. However, these two solutions induce an important modification of the global mechanical assembly. If the designer wants to keep the same type of assembly and the same type of power module, these solutions cannot be used. Therefore, one must increase the thermal efficiency of the cold plate to reduce the thermal resistance due to the heat exchange and

consequently lower the junction temperature of the SiC chip. This resistance can be optimized by increasing the exchange surface and/or the heat transfer coefficient. These two parameters are intimately linked. The heat transfer coefficient *h* will depend on the geometry of the surface exchange, the flow rate and the flow regime, and the thermo-physical properties of the fluid. In the same way, the modification of the surface exchange is associated by an increase in the coefficient *h* but also by an increase in the friction factor.

Mersen cooling division uses several geometric configurations to enhance heat transfers in cold plates. From literature point of view, offset strip fins (OSF) and wavy technologies show better thermal performance. Gong et al [10] indicate an increase of thermal performance by 55% compared to straight channels. Aneesh et al [11] study different wavy geometries. They find that trapezoidal channels highlight very high thermal performances, but with higher pressure drops. In addition, manufacturing this kind of geometry could be very difficult from industrial point of view. Other geometries for intensification of the heat exchange like pin fin or louvered fin could be used to decrease the convective thermal resistance and consequently reduce chip temperature. However, the integration of these technologies in a cold plate often results in a compromise between thermal performances and the generated pressure losses. This is why each technology is optimized for a specific operating case and especially for a given power module packaging.

Finally, it follows that for a good reduction of the thermal stress of a power module chip, an optimization on the entire module, TIM and cooler chain is necessary.

Thus, the aim of this paper is to discuss and evaluate some solutions to improve the thermal management of high reliability power modules by keeping the classical mechanical arrangement using indirect liquid cooling: power modules with baseplates fixed on cold plates. In the following section, the global methodology will be presented. It includes the description of samples with and without inner heat spreaders, the fabrication of the tested cold plates, the description of the experimental setup and the presentation and discussion of simulation and experimental results.
# Implementation methodologies

## Power module description

Power modules are implemented at Safran-Tech with SiC MOSFETs (surface 5 mm x 5 mm). Fig. 3 shows two studied power modules: the first one (Fig. 3a) has a classical configuration and the other one (Fig. 3b) includes an internal heat spreader under the chip. This spreader is made with silver diamond (AgD) which has a higher thermal conductivity and a lower CTE compared to copper (Table I). The die-attach and the internal spreader attach are made with 80Au20Sn solder alloy given its better thermo-mechanical performance compared to SAC305 for small surfaces [12]. The Active Metal Brazing (AMB) substrate is made of copper and Silicon Nitride (Si₃N₄). As mentioned earlier, it is chosen given its high lifetime under large amplitude thermal cycles [13]. Despite its lower thermal conductivity compared to copper, the baseplate is made with Aluminum Graphite (ALG1808) since it has a lower density and a closer CTE to the AMB substrate one. The baseplate is attached to the substrate with SAC305 InForm solder [14]. This type of solder is dedicated for large surfaces and contains a copper mesh that maintains the solder thickness and reduces air voids generated in the solder solidification process. Thus, it allows to improve both the reliability and the effective thermal resistance.

<figure><img src="image_4.png" /><figcaption>Fig. 3: Power modules for the tests – (a) typical assembly – (b) assembly with AgD thermal spreader</figcaption></figure>

<table><caption>Table I: Thermal properties of the assembly materials at 25 °C.</caption><thead><tr><td>Material</td><td>Thermal conductivity λ<sub>∞</sub> W·m<sup>-1</sup>·K<sup>-1</sup></td><td>Density, ρ<sub>s</sub> kg·m<sup>-3</sup></td><td>CTE, ppm·K<sup>-1</sup></td></tr></thead><tbody><tr><td>SiC</td><td>300</td><td>2330</td><td>2.6</td></tr><tr><td>Si<sub>3</sub>N<sub>4</sub> ceramic</td><td>65</td><td>3110</td><td>3.3</td></tr><tr><td>ALG1808</td><td>180</td><td>2200</td><td>7.9</td></tr><tr><td>Copper, Cu</td><td>400</td><td>8950</td><td>17</td></tr><tr><td>Silver Diamond AgD</td><td>920</td><td>5600</td><td>6.1</td></tr><tr><td>80Au20Sn</td><td>57</td><td>14700</td><td>16</td></tr><tr><td>SAC305</td><td>60</td><td>7330</td><td>40</td></tr></tbody></table>

## Cold plates

The aluminium cold plates used in this study are manufactured at Mersen and assembled by vacuum brazing. This well-controlled process makes it possible to obtain high-end plates in terms of reliability and durability (over 20 years of lifetime). This method of assembling cold plates by brazing allows to easily integrate intensification geometries to enhance heat exchanges and consequently obtain efficient cold plates. In this sense, other technical assemblies of cold plates such as Friction Stir Welding (FSW) could show some limitations and can be used for simple plates (without intensification geometry like offset strip fins). Furthermore, Mersen cold plates with indirect cooling obtained by vacuum brazing, show an excellent resistance to thermal cycling issue thanks to the cover plate which allows a better heat spreading. Likewise, the leaks risk is non-existent thanks to the monolithic composition of the cold plate and the grain-to-grain connection. Finally, the risks linked to corrosion are considerably reduced and the maintenance is almost non-existent.

Thus, four aluminium cold plates were manufactured and assembled by vacuum brazing to run this study:

- CP_V1: smooth straight channel (hydraulic diameter = 3 mm)
  CP_V2; CP_V4 and CP_V5: these cold plates are fitted with a special geometry in order to increase the heat exchange. The channels of CP_V2 channel are presented in [15]. While the CP_V4 is fitted with offset strip fins. Finally, the CP_V5 is specific geometry represented by periodic repetition of a half moon
The overall dimension of each cold plate is 100 mm x 65 mm x 19 mm. G641 thermal grease from Novagard company is used to ensure the thermal connection between the power module and the cold plate. In addition, 3 K-type thermocouples are installed on the surface of the plate. Fig. 4 shows the power module/heat sink assembly and the thermocouples positions.

<figure><img src="image_5.png" /><figcaption>Fig. 4: Power module and cold plate assembly.</figcaption></figure>

## Characterization tools

The closed loop system used to characterize the assembly of Fig. 4 is described in [15]. The junction temperature measurement is based on the use of a thermosensitive electrical parameter: the voltage across the body diode of the MOSFET [16] with $V_{GS}=-10$V and $I_D=-100$ mA.

The fluid used in this study consists of a mixture of water ethylene glycol (EG50%). This type of fluid is widely used in cooling of power electronics. Table II shows the thermo-physical properties of the coolant for two different temperatures.

<table><caption>**Table II: Thermo-physical properties of EG50%**</caption><thead><tr><th>EG50%</th><th>ρ<br>kg.m<sup>-3</sup></th><th>Cp<br>J.kg<sup>-1</sup>.K<sup>-1</sup></th><th>λ<br>W.m<sup>-1</sup>.K<sup>-1</sup></th><th>μ<br>Pa.s</th></tr></thead><tbody><tr><td>T = 50°C</td><td>1042</td><td>3456</td><td>0.418</td><td>0.00174</td></tr><tr><td>T = 40°C</td><td>1048</td><td>3409</td><td>0.418</td><td>0.00214</td></tr></tbody></table>

## Study parameters

The unique chip of the power module under study dissipates 100 W. Given the surface area of the chip, the power density to be evacuated is 400

W.cm⁻². The Reynolds number based on the hydraulic diameter varies over a range from 500 to 8000 corresponding to the volume flow rate from 0.5 to 6.5 L.min⁻¹.

The thermal resistance of the cold plate is defined by:

$$R_{th} = \frac{T_w - T_a}{P} \qquad (1)$$

where $T_w$ and $T_a$ are respectively the surface wall temperature of the cold plate and the inlet bulk fluid temperature. $P$ (W) is the power losses. This equation is also used to calculate the global junction to ambient thermal resistance $R_{th(j-a)}$:

$$R_{th(j-a)} = \frac{T_j - T_a}{P} \qquad (2)$$

with $T_j$ is the junction temperature of the semi-conductors.

Otherwise, the pumping power is calculated by:

$$P_p = Q_v \Delta P \qquad (3)$$

where $P_p$ (W) is the pumping power, $Q_v$ (m³.s⁻¹) the volume flow rate and $\Delta P$ (Pa) the pressure drop.

# Results and discussion

## Pressure drops

The pressure drops measurement for the four cold plates described above with different volume flowrates at isothermal condition ($T_a = 50°C$) are given in Fig. 5. This figure shows that the pressure drops of CP_V1, with a smooth straight channel, are important compared to the optimized cold plates. The reduction of the hydraulic diameter of this plate induces a significant increase in the pressure loss. However, the design of this cold plate was carried out with the aim of having a reasonable thermal performance compared to the optimized plates. In the case of a larger hydraulic diameter (standard commercial plate), the thermal performances would certainly be very low, which would risk to reduce the quality of our comparison in this study.

Furthermore, the CP_V4 plate shows the lowest pressure losses, this is remarkable with the increase in flow rate. For $Q_v = 6$ L.min⁻¹, the V5 plate shows an increase in $\Delta P$ of 105% compared to the V4 plate. Similarly, plate V2 shows a 54%
increase in $\Delta P$ compared to plate V4. The CP_V4 cold plate has a geometry that allows for a locally reduced flow speed. In fact, the fluid is distributed over several mini channels. This therefore reduces the pressure loss.

Finally, Fig. 5. highlights that the cold plates V2, V4 and V5 cause turbulence in the fluid with small volume flow rate (Reynolds number). This is an interesting result regarding thermal stresses. Even with a weak flow rate given the existing turbulence, the heat is well evacuated from the SiC chips. This highlights the possibility of working with low flow rates to evacuate high power densities.

<figure><img src="image_6.png" /><figcaption>Fig. 5: Pressure drops comparison between four different geometries of cold plates at $T_a = 50^\circ C$ (fluid: EG50%).</figcaption></figure>

## Cold plates thermal resistance

Fig. 6 shows the evolution of the thermal resistance of the cold plates as a function of the pumping power. These curves are obtained at isothermal conditions: dissipation power 100 W, $T_a = 34^\circ C$ and the fluid used is EG50%. The power module used does not have an AgD spreader under the chip. The CP_V5 plate shows the lowest thermal resistance, very close to the one of CP_V2 cold plate. For a pumping power of 1 W, the CP_V5 cold plate shows a lower $R_{th}$ by 12% compared to the V2 cold plate, by 29% compared to the V4 and by 57% compared to the V1. The CP_V1 cold plate shows a higher $R_{th}$ for an identical pumping power. This therefore highlights the interest of using optimized plates with specific geometries.

In the following sections and concerning the study of the power module with and without integrated spreader, we will conduct experimental and numerical comparisons between the CP_V1 and CP_V2 plates. This approach is motivated by

the facility of numerical studies of the two cold plates. The Cp_V1 plate will be used as a reference.

<figure><img src="image_7.png" /><figcaption>Fig. 6: Comparison of thermal resistance for the studied cold plates at $T_a = 34^\circ C$ (fluid: EG50%).</figcaption></figure>

## Impact of heat spreader on the thermal resistance of the single chip power module

In this section, we experimentally study the impact of the spreader on the $R_{th-ja}$ of the single-chip power module. Fig. 7 shows the evolution of $R_{th-ja}$ and the junction temperature as a function of the two cold plates CP_V1 and CP_V2 for $Q_v = 3$ L.min$^{-1}$. It is clear that AgD spreader allows a significant reduction in the junction temperature. The spreader eliminates the hot spot effect in the module and extends the heat exchange surface. Thus, the spreader which has a thermal conductivity of 900 W.m$^{-1}$.K$^{-1}$ allows 17% reduction of the global $R_{th}$. This reduction is well marked with the optimized cold plate CP_V2. The junction temperature of the module with spreader and the CP_V2 plate is 98$^\circ C. A gain of 6.5$^\circ C$ is obtained with the optimized plate compared to the CP_V1 cold plate. Finally, note that this single-chip power module configuration does not explicitly show the impact of an optimized cold plate. The gains are lower for the case of the CP_V1 cold plate. This explains that for single chip modules the spreader has a better performance in high heat exchange environments. In addition, the impact of the TIM (thermal interface material) in this configuration is not negligible, which may partially mask the optimized cold plate performances. In real operation of power modules in converters, this case study (single chip) is not feasible except for discrete components such as TO247. Current and future developments in power modules show an increase in chip density in reduced surfaces. Thus,
the thermal behaviors are certainly modified and are not suitable for single-chip components.

<figure><img src="image_8.png" /><figcaption>Fig. 7: Impact of AgD spreader and optimized cold plate on thermal resistance junction to ambient for CP_V1 et CP_V2. ($T_a$ = 34°C, fluid: EG50% and $Q_v$ = 3 L.min⁻¹).</figcaption></figure>

In order to go further in our study and to confirm the hypotheses made above, we carried out a comparative study by numerical simulations on a power module which contains 4 chips (with and without spreaders). The objective of this step is to work across a real case of a power module. In this approach we have therefore simulated the plates CP_V1 and CP_V2. In [15] and [17], we demonstrate the compliance between measurement and simulation of CP_V2 cold plate technology. The CP_V1 is made with simple smooth parallel channel. The simulation of this type of geometry is quite simple and generally is consistent with experimental reality. Finally, the power losses in the power module $P_{d-tot}$ = 280 W (each chip dissipates 70 W). The results of this approach are given in the following section.

## Effect of heat spreader on the thermal resistance of a multi-chip power module

Fig. 8 shows the power module studied in this section. It is composed of 4 identical SiC MOSFET chips (size 5 mm x 5 mm). The different layers of power modules remain identical to the description given in Fig. 3.

The three-dimensional numerical simulations are solved using Ansys software based on the finite volume method. For the momentum equation for the coolant, the inlet volume flow rate is specified, while the outlet pressure is set to 0 Pa. We assume that there is no slip at the wall. For energy equation, the inlet temperature is given ($T_a$ = 34°C). The k-epsilon model is used to simulate flow characteristics for turbulent flow conditions.

<figure><img src="image_9.png" /><figcaption>Fig. 8: Multi-chip power module fitted with AgD spreaders.</figcaption></figure>

The heat transfer between plates and ambient, viscous dissipation, gravity and body forces are not taken into consideration.

Fig. 9a and Fig. 9b show the mapping temperature ($Q_v$ = 5 L.min⁻¹) for both cases with and without spreader for the CP_V2 cold plate. Fig. 10 shows a synthesis of the simulation scenarios for $Q_v$ = 5 L.min⁻¹. From these two figures, we can conduct the following quantitative analysis:

- *Impact of the optimized cold plate CP_V2:* In the case of a power module without spreader, the cold plate CP_V2 allows a $R_{th}$ reduction of 19% compared to the plate CP_V1. For the case with spreader, the CP_V2 has a 22% $R_{th}$ reduction.

- *Impact of the spreader:* the impact of the spreader on $R_{th}$ with the CP_V1 plate allows a reduction of 17%. In the case of the CP_V2 plate, a reduction of $R_{th}$ of 21% is observed.

Furthermore, the interest of using an optimized plate is well demonstrated by the comparison between the two plates CP_V1 and CP_V2. The junction temperature with spreader for CP_V1 is almost similar to the junction temperature of CP_V2 without spreader. In other words, the efficiency of the spreader integrated in a power module is improved with the use of an optimized plate. It then results that the hybridization between a power module with a spreader and an optimized cold plate allows for better thermal management, an increase in the reliability of the power module and consequently an increase of the conducted power of the module.

Finally, the observations obtained in this section are in the same order of magnitude with those obtained in the case of double side cooling power module [18]. The main interest in this study
comparing to the double side cooling is to save
the classical heat sink/power module assembly.

<figure><img src="image_10.png" /><figcaption>Fig. 9: Power module and cold plate mapping
temperature for Q_v = 5 L/min: a) with spreader,
b) without spreader.</figcaption></figure>

<figure><img src="image_11.png" /><figcaption>Fig. 10: Performance of multi chips power
module with and without spreader and
optimized cold plate. (T_a = 34°C, fluid: EG50%
and Q_v = 5 L.min⁻¹).</figcaption></figure>

Conclusion

The main objective of this work is to study the
impact of an inner spreader (integrated under the
chip) and an optimized cold plate on the thermal

performances of a power module. The assembly
of the latter was done with respect to mechanical
and thermal constraints. Different typologies of
cold plates and power modules were studied. The
main results are summarized as follow:

- Cold plates fitted with a heat exchange intensification geometry allow better cooling of the power modules. A reduction of more than 55% in the thermal resistance is obtained compared to a simple cold plate.

- The heat spreader shows better efficiency when it is mounted with an optimized cold plate. A maximum reduction of 36% on the R<sub>th-ja</sub> is obtained.

The perspectives of this work are based on the
experimental validation of the results obtained by
simulation. A comparison approach between
different spreaders must also be undertaken. A
densification of the chips within the module is
also necessary to be close to the commercialised
power modules.

Finally, an economic study must be conducted in
order to properly quantify the cost of the spreader
on the overall price of the power module and thus
its interest for the thermal management of power
modules.

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