SiC-Based High E ﬃ ciency High Isolation Dual Active Bridge Converter for a Power Electronic Transformer

: This paper discusses the beneﬁts of using silicon carbide (SiC) devices in a three-stage modular power electronic transformer. According to the requirements to be fulﬁlled by each stage, the second one (the DC / DC isolation converter) presents the most estimable improvements to be gained from the use of SiC devices. Therefore, this paper is focused on this second stage, implemented with a SiC-based dual active bridge. Selection of the SiC devices is detailed tackling the e ﬃ ciency improvement which can be obtained when they are co-packed with SiC antiparallel Schottky diodes in addition to their intrinsic body diode. This e ﬃ ciency improvement is dependent on the dual active bridge operation point. Hence, a simple device loss model is presented to assess the e ﬃ ciency improvement and understand the reasons for this dependence. Experimental results from a 5-kW Dual Active Bridge prototype have been obtained to validate the model. The dual active bridge converter is also tested as part of the full PET module operating at rated power.

As is clear in Figure 1, the structure is fully modular as it is formed by several identically-stacked cells. Thanks to this modular structure, the number of cells of the CHB is chosen in such a way that the cell voltage (V cell ) is equal to the required V dcLV . In this way the design of this DAB is, to some extent, simplified using a unity transformation ratio (i.e., no step-up/down). This PET topology requires high galvanic isolation between HV and LV sides. The isolation is provided by the DAB high frequency transformer (HFT). Table 1 shows the characteristics of the developed PET.

Use of WBG for PET
The superior material properties of WBG semiconductors allow power devices to operate at higher temperatures, voltages and switching frequency in comparison to Si counterparts [4,21]. Among WBG materials, SiC presents the most mature technology for high voltage devices [4]. Both 1.2 kV and 1.7 kV SiC MOSFETs are already available on the market with a wide range of current ratings [5,22,23]. The use of SiC MOSFETs not only introduces a relevant improvement to the efficiency of fast switching power converters, but also enables going to higher switching frequencies at high blocking voltage which cannot be achieved using available Si IGBTs. However, at limited switching frequency requirements (<10 kHz), especially for high power (>100 A), Si IGBTs are still the preferred choice due to cost-effectivity and reliability in addition to SiC MOSFET higher dv/dt, di/dt, and EMI issues [24].
Accordingly, to analyze the merits of integrating SiC devices in the PET, the device requirements for each of the three stages of the PET are identified below.

Device Requirements per PET Stage
In the addressed PET topology, since no step-up/down is needed, the power devices employed on both transformer sides (i.e., HV and LV sides) have the same voltage rating. These include devices of the CHB full bridge (FB), the two DAB full bridges (FB1, FB2), and the DC/AC converter devices (see Figure 1). While this is true, the specifications in terms of current ratings and commutation requirements differ significantly.
CHB: its power devices do not need to commutate fast due to the multilevel nature of the topology [25,26]. Moreover, C cell size is not determined by the cell switching frequency. Therefore, SiC switching devices are not of special merit here, even when going to higher V cell (i.e., to reduce the number of stacked modules), still Si IGBT would be the selected option [14]. On the other hand, SiC free-wheeling antiparallel diodes are quite interesting in this case since the CHB full bridge is required to handle positive and negative currents (i.e., not only during the dead time). Therefore, SiC diodes can effectively improve the CHB efficiency due to their reduced conduction forward voltage drop as well as lower reverse recovery time compared to Si diodes [27].
DAB: high switching frequencies can provide significant reduction of the size and weight of the converter magnetics and the input and output capacitors. Additionally, going to higher PET module voltage is not achievable using Si devices, unless the switching frequency is reduced to few kHz. Regarding the antiparallel diode, the advantages of using SiC are still a controversial issue [14,28] and, therefore, will be analyzed in this work. 3P4L DC/AC: commutation requirements are not high (in the range of few kHz [29]), and since it interfaces the LVAC grid, high blocking voltages are not required. Therefore, Si devices are a good candidate for this stage.
Considering the above discussion, for this PET, 1.7 kV Si IGBT devices with SiC freewheeling diodes are used for the CHB FB. As for the DAB, 1.2 kV SiC MOSFETs are used and possible enhancements introduced by a SiC antiparallel diode will be discussed in detail in Section 4. Finally, 1.7 kV Si IGBTs are used for the 3P4L converter.
In the next section, the possible achievable enhancements gained by employing SiC in the DAB are discussed highlighting the practical limitations intrinsic to this PET topology.

Benefits and Practical Limitations of Using SiC MOSFETs for the DAB Converter
The use of SiC MOSFETs for the DAB converter in this PET structure has two main benefits: 1.
It enables increasing the switching frequency of the DAB and, therefore, decreasing the size of the HFT [30,31].

2.
The high blocking voltages enable increasing V cell and, therefore, for a given V acHV , the number of stacked PET modules can be reduced (see Figure 1) [7,17]. This is also advantageous for the HFT. Since each module handles more power, this leads to a higher transformer power density. In order to see clearly the effect of this, Figure 2 shows the HFT power density as a function of the cell voltage where the PET total power and V acHV are fixed (i.e., the number of PET modules vary). A comparative analysis regarding this relation is previously presented in [17] showing an improvement in the HFT power density as V cell increases.

124
In the next section, the possible achievable enhancements gained by employing SiC in the DAB 125 are discussed highlighting the practical limitations intrinsic to this PET topology.

127
The use of SiC MOSFETs for the DAB converter in this PET structure has two main benefits: 2. The high blocking voltages enable increasing V cell and, therefore, for a given V acHV , the number 131 of stacked PET modules can be reduced (see Figure 1) [7,17]. This is also advantageous for the 132 HFT. Since each module handles more power, this leads to a higher transformer power density.

133
In order to see clearly the effect of this, Figure 2 shows the HFT power density as a function of 134 the cell voltage where the PET total power and V acHV are fixed (i.e., the number of PET modules 135 vary). A comparative analysis regarding this relation is previously presented in [17] showing an 136 improvement in the HFT power density as V cell increases.

139
However, these SiC potential benefits may be compromised by certain practical implementation 140 constraints.

141
On one hand, increasing V cell has several adverse effects. High DC link voltages create practical 142 problems for feeding the control circuitry in each cell. As commercial auxiliary power supplies (APS)

143
do not provide the required isolation [29], each module circuitry has to be supplied from its DC link, where the HV and LV sides have separate APSs [32]. Commercial APSs can be used for voltages 145 under 1 kV, otherwise, custom solutions must be implemented, such as the modular ISOP topology 146 proposed in [29]. Consequently, various aspects must be considered for the selection of V cell .

147
On the other hand, the size reduction of the HFT, in this particular case, is constrained by the 148 high isolation required by the PET. This isolation imposes minimum clearance distances between 149 windings, which compromises the window utilization factor resulting in a physical limit on further 150 size reduction.

152
The DAB (see Figure 3)   However, these SiC potential benefits may be compromised by certain practical implementation constraints.
On one hand, increasing V cell has several adverse effects. High DC link voltages create practical problems for feeding the control circuitry in each cell. As commercial auxiliary power supplies (APS) do not provide the required isolation [29], each module circuitry has to be supplied from its DC link, where the HV and LV sides have separate APSs [32]. Commercial APSs can be used for voltages under 1 kV, otherwise, custom solutions must be implemented, such as the modular ISOP topology proposed in [29]. Consequently, various aspects must be considered for the selection of V cell .
On the other hand, the size reduction of the HFT, in this particular case, is constrained by the high isolation required by the PET. This isolation imposes minimum clearance distances between windings, which compromises the window utilization factor resulting in a physical limit on further size reduction.

SiC-Based DAB Converter
The DAB (see Figure 3) is selected for the intermediate stage of the PET, as it provides galvanic isolation as well as bidirectional power flow [33,34]. The DAB is based on two active bridges interfaced through an HFT, which provides the required galvanic isolation. This converter provides bidirectional operation by controlling the phase shift between the AC voltages generated by both bridges (V 1 and V 2 ). Also, this converter can have relatively high efficiency due to the soft-switching operation of all the devices at nominal conditions (zero-voltage switching, ZVS) [35,36].

169
The boost converter was chosen as a preliminary test bench due to its simplicity and rapid 170 prototyping, but foremost, due to its similar operation to a DAB converter as it has two switching 171 devices in a leg with an inductance connected to the middle point. However, the differences in 172 operation between a CCM boost and a DAB are well understood and, therefore, perspective devices 173 resulting from the first selection stage are then tested in a DAB converter prototype.

174
The test bench used commercial driver boards from CREE and a commercial FPGA-based 175 controller platform (BASYS2). Tests in the boost converter were done at 2 kW 400/800 V for switching 176 frequencies of 30, 50, and 100 kHz and a dead time of 500 ns. Table 2 summarizes the results. The 177 efficiency is calculated using the input and output DC voltages and currents of the converter 178 measured using digital multimeters. It is observed that all the devices show a high efficiency barely 179 affected by the increase in switching frequency which was increased by a factor of more than three.

SiC Device Selection
Regarding the selection of the SiC devices for the DAB, 6.5, 10, and 15 kV SiC MOSFETs have been reported for laboratory prototypes [37][38][39][40][41][42], but are far from being a viable commercial alternative yet. Current ratings offered in commercially available 1.7 kV SiC MOSFETs are still limited [23], and their commutation characteristic must be improved. The 1.2 kV SiC devices remain as the most mature technology available in the SiC device market. Consequently, these devices have been selected for the DAB. The DC link voltage is, therefore, set to 800 V. Seven 1.2 kV SiC commercial MOSFETs were selected; two power modules and five discrete N-channel SiC MOSFETs. A comparative analysis of these devices was carried out in a boost converter operating in continuous conduction mode (CCM) (see Figure 4).

169
The boost converter was chosen as a preliminary test bench due to its simplicity and rapid 170 prototyping, but foremost, due to its similar operation to a DAB converter as it has two switching 171 devices in a leg with an inductance connected to the middle point. However, the differences in 172 operation between a CCM boost and a DAB are well understood and, therefore, perspective devices 173 resulting from the first selection stage are then tested in a DAB converter prototype.

174
The test bench used commercial driver boards from CREE and a commercial FPGA-based 175 controller platform (BASYS2). Tests in the boost converter were done at 2 kW 400/800 V for switching 176 frequencies of 30, 50, and 100 kHz and a dead time of 500 ns. Table 2 summarizes the results. The 177 efficiency is calculated using the input and output DC voltages and currents of the converter 178 measured using digital multimeters. It is observed that all the devices show a high efficiency barely 179 affected by the increase in switching frequency which was increased by a factor of more than three.  The boost converter was chosen as a preliminary test bench due to its simplicity and rapid prototyping, but foremost, due to its similar operation to a DAB converter as it has two switching devices in a leg with an inductance connected to the middle point. However, the differences in operation between a CCM boost and a DAB are well understood and, therefore, perspective devices resulting from the first selection stage are then tested in a DAB converter prototype.

180
The test bench used commercial driver boards from CREE and a commercial FPGA-based controller platform (BASYS2). Tests in the boost converter were done at 2 kW 400/800 V for switching frequencies of 30, 50, and 100 kHz and a dead time of 500 ns. Table 2 summarizes the results. The efficiency is calculated using the input and output DC voltages and currents of the converter measured using digital multimeters. It is observed that all the devices show a high efficiency barely affected by the increase in switching frequency which was increased by a factor of more than three. Since the performance of all seven MOSFETs is comparable, the selection of the adequate option was mainly based on the size and the price. The specifications for the DAB converter are shown in Table 1. Accordingly, the peak current handled by the devices can be calculated from Equation (1) [35], where T is half the switching period, d is the phase shift, L lk is the leakage inductance, v o is the output voltage, v i is the input voltage and n is the HFT turns ratio: The current peak is calculated for the maximum phase-shift, this is selected to be 0.35 according to [35]. Based on Equation (1), this current is approximately equal to 11 A. The device current rating is selected to be twice the magnitude of the peak current handled by the devices to keep a safety margin. Therefore, the minimum required current rating is 22 A. This eliminates the two modules, and the CREE 60 A discrete, as the size and price are not justified in this case. The remaining four discrete devices are almost equally favored except for the ROHM SCH2080KE, as it includes a SiC Schottky barrier diode (SBD) co-packaged with the MOSFET.

Antiparallel SBD for a DAB Converter
Observing the efficiency of the boost converter using ROHM SCH2080KE versus for example ROHM SCT2080KE (without an additional SBD), it is consistently higher without SiC SBD. For the boost converter, this is logical as an additional antiparallel diode increases the output capacitance (see Table 2) and since hard-switching occurs, this increases the switching losses. Although, this makes sense and is simple to understand for the boost, for the DAB, it is more complicated as ZVS is implemented. That being the case, it is not valid to make a selection between both devices unless the additional diode behavior is studied for the DAB operation to identify if it improves or worsens its efficiency. This issue is addressed as follows, including: (1) understanding the potential effects introduced by a SiC antiparallel diode in a DAB converter, (2) developing a simple analytical loss model to estimate the possible efficiency improvement introduced by the SiC diode at a certain DAB operating point, and (3) validation of the proposed model using experimental results in a DAB prototype.
The two devices used in the analysis are the ROHM devices (i.e., SCH2080KE and SCT2080KE) as it is the same die but one packed with a SiC antiparallel SBD [22]. Characteristics of both devices are provided in Table 3. The diode forward voltage, V F-diode , is obtained from the datasheet at the value of I p_lk (see Equation (1)) where the employed phase shift is that corresponding to 5 kW. In order to provide a qualitative preliminary understanding of the possible effects of the SiC SBD on the DAB efficiency, the waveforms of the DAB are shown in Figure 5.

215
The potential advantages of using SCH2080KE +SBD in comparison to SCT2080KE are analyzed 216 as follows:

217
• Since VF-Diode of the additional SiC SBD is more than three times lower than that of the body diode 218 (see Table 3), the power losses due to diode conduction during the dead times are lower using 219 SCH2080KE +SBD.

220
• The diode can enter into conduction while the MOSFET is ON. This can be seen from Figure 5, The potential advantages of using SCH2080KE +SBD in comparison to SCT2080KE are analyzed as follows: • Since V F-Diode of the additional SiC SBD is more than three times lower than that of the body diode (see Table 3), the power losses due to diode conduction during the dead times are lower using SCH2080KE +SBD.

•
The diode can enter into conduction while the MOSFET is ON. This can be seen from Figure 5, when the current through the MOSFET is negative (i.e., source to drain), a forward voltage drop, V F-MOSFET , is applied to the diode. If V F-MOSFET is higher than the diode knee voltage (V knee ), then the diode conducts. If this case is true, current is shared between the MOSFET and the diode and, therefore, conduction losses are reduced.

•
Unlike boost converter operation, soft switching is adopted in DAB devices at turn ON and, therefore, the additional output capacitance introduced by the additional SiC diode, illustrated in Figure 4, does not significantly penalize the switching losses (see Table 2).
However, these advantages can be compromised by several situations resulting in almost no advantages of having a SiC SBD, these cases are summarized as follows: • From Figure 6, diodes do not operate during the whole period of the dead time. During (A), the output parasitic capacitances of the MOSFETs (see Figure 4) in one leg are exchanging the voltage (i.e., one is discharging while the other is charging) and during (B), diodes conduct due to circulating currents. In some cases, (B) can tend to zero depending on the value of the current charging/discharging the parasitic capacitors and on the dead time.

•
If the current through the MOSFETs is not high enough to produce a V F-MOSFET > V knee , then the diode will never conduct during MOSFET ON time.

•
The turn OFF switching losses should be analyzed as the enhancement introduced by the SiC SBD in the conduction losses can, in some cases, be compromised by the increase in turn OFF switching losses (due to the extra capacitance).
However, these advantages can be compromised by several situations resulting in almost no 230 advantages of having a SiC SBD, these cases are summarized as follows:

231
• From Figure 6, diodes do not operate during the whole period of the dead time. During (A), the 232 output parasitic capacitances of the MOSFETs (see Figure 4) in one leg are exchanging the 233 voltage (i.e., one is discharging while the other is charging) and during (B), diodes conduct due 234 to circulating currents. In some cases, (B) can tend to zero depending on the value of the current 235 charging/discharging the parasitic capacitors and on the dead time.

236
• If the current through the MOSFETs is not high enough to produce a V F-MOSFET > V knee , then the 237 diode will never conduct during MOSFET ON time.

238
• The turn OFF switching losses should be analyzed as the enhancement introduced by the SiC 239 SBD in the conduction losses can, in some cases, be compromised by the increase in turn OFF 240 switching losses (due to the extra capacitance).

241
As a conclusion to the previous discussion, it is important to develop a simple model to

249
Accordingly, i lk is considered flat (I 1 = I 2 in Figure 5) and, therefore, as an approximation, the peak

254
Since t (B) = dead time -t (A) , therefore diode operates for 570 ns for the case of the SiC SBD and 255 588 ns for the body diode. This time represents around 95% of the total dead time (600 ns) and, 256 therefore, the reduction in conduction losses during diode operation is relevant to consider.

257
Regarding the diode conduction interval during MOSFET ON time, first, the diode 258 characteristics are identified from the MOSFET datasheet (see Table 3). Figure 7 shows the forward  As a conclusion to the previous discussion, it is important to develop a simple model to determine, for a certain DAB operation point, if an extra antiparallel SiC diode co-packed with the SiC MOSFET is worthy or avoiding it is better. This is introduced in the next section.

Diode Conduction Intervals
Regarding diode conduction interval during the dead time, in order to estimate the time duration (B) (see Figure 6), the capacitor charging time during (A) is estimated using Equation (2), where C o(er) is the MOSFET effective output capacitance given by the data sheet and V 1 = V i = V o . Accordingly, i lk is considered flat (I 1 = I 2 in Figure 5) and, therefore, as an approximation, the peak current (I p_lk ) is considered to be equal during all the switching transitions.
Since t (B) = dead time − t (A) , therefore diode operates for 570 ns for the case of the SiC SBD and 588 ns for the body diode. This time represents around 95% of the total dead time (600 ns) and, therefore, the reduction in conduction losses during diode operation is relevant to consider.
Regarding the diode conduction interval during MOSFET ON time, first, the diode characteristics are identified from the MOSFET datasheet (see Table 3). Figure 7 shows the forward voltage drop for both diodes (V F-Diode ) as a function of the current through the diode. Additionally, the voltage drop on the MOSFET due to its ON resistance (V F-MOSFET ) as a function of the current it is conducting is illustrated.
Based on the data in Table 3, I min SiC is 6.8 A while I min body is 11.

276
First, the periods where the diode can conduct (when it is forward biased) are identified as 277 shown in Figure 8. It can be seen that, the diodes of all the ON MOSFETs can conduct during (1) and

278
(4) as the current through all the devices is negative (see Figure 5), while during (3) and (6)

282
Accordingly, to simplify the power loss estimation the following assumptions are made:

283
• As I min SiC is around 75% of I pk-lk , therefore, it is assumed that the diode operates only during It is possible to see that each diode conducts when the current through the MOSFET is above a certain minimum value. It is clearly lower in the case of the SiC SBD. This can be estimated from Equation (3), where R ON-MOSFET stands for ON-state drain-to-source MOSFET resistance: Based on the data in Table 3, I min SiC is 6.8 A while I min body is 11.2 A. Since the theoretical calculated I p_lk = 9 A, therefore it is not possible that the body diode enters into conduction while the MOSFET is ON. On the other hand, the SiC diode would conduct when the current exceeds 6.8 A.

Estimation of the Power Losses
Power losses are estimated for both cases of the DAB using MOSFETs with a SiC antiparallel diode (SCH2080KE) and with only the body diode (SCT2080KE). Theoretical power loss estimation will be compared to the experimental efficiency results presented in the next section.
First, the periods where the diode can conduct (when it is forward biased) are identified as shown in Figure 8. It can be seen that, the diodes of all the ON MOSFETs can conduct during (1) and (4) as the current through all the devices is negative (see Figure 5), while during (3) and (6) only the diodes of the secondary bridge can conduct.
Accordingly, to simplify the power loss estimation the following assumptions are made: • As I min SiC is around 75% of I pk-lk , therefore, it is assumed that the diode operates only during period (3) and (6) when the peak current is passing through the devices (periods (1) and (4) are neglected).

•
The power loss estimation is performed only during the periods with different operation for the case of the SiC SBD and the body diode. In other words, only the difference in losses between both cases is considered. These intervals are: (3), (6) and the dead times.
Power losses are estimated for both cases of the DAB using MOSFETs with a SiC antiparallel 272 diode (SCH2080KE) and with only the body diode (SCT2080KE). Theoretical power loss estimation 273 will be compared to the experimental efficiency results presented in the next section.

274
First, the periods where the diode can conduct (when it is forward biased) are identified as 275 shown in Figure 8. It can be seen that, the diodes of all the ON MOSFETs can conduct during (1) and

276
(4) as the current through all the devices is negative (see Figure 5), while during (3) and (6)  280 Accordingly, to simplify the power loss estimation the following assumptions are made:

281
• As I min SiC is around 75% of I pk-lk , therefore, it is assumed that the diode operates only during 282 period (3) and (6) when the peak current is passing through the devices (periods (1) and (4)   During intervals (3) and (6), the conduction and switching losses of all the devices are estimated, while during the dead times, only the diodes conduction losses are estimated.

(a) Losses during dead time:
During one switching period (T sw ), four intervals of dead time take place. During each interval two diodes conduct. Accordingly, the diodes total conduction losses are estimated using Equation (4), where V F-Diode is the diode forward voltage at I p_lk obtained from the diode characteristic curve given by the datasheet (linearized in Figure 7). (3) and (6) (Figure 8): Losses during these two intervals are divided into: (1) turn OFF switching losses (P sw ) and (2) MOSFET and diode conduction losses (P MOSFET and P Diode ). The switching losses are straightforward, estimated from MOSFET E OFF using Equation (5).
On the other hand, conduction losses are not straight forward. As noticed from Figure 8, in both (3) and (6) intervals, the primary bridge devices are different from the secondary ones. For example, during (3), for the primary, only the MOSFETs conduct (i.e., all the current flows through the MOSFETs, S1 & S4). In this case, four MOSFETs of the primary bridge conduct during (3) and (6) (two in each interval). Therefore, the total primary conduction losses can be easily estimated using Equation (6).
However, for the secondary, the current is shared between the diode and the MOSFET (S5, D5 and S8, D8). Accordingly, the current conducted by each element has to be estimated. Since the voltage drop on the MOSFET is equal to the diode V F-Diode , then: where I M and I D are the currents through the MOSFET and the diode respectively and R D is the diode dynamic resistance.
Since the current is shared by both elements, then, I M and I D can be obtained from Equations (7) and (8): Finally, the total conduction losses of the secondary bridge during T sw can be calculated from Equations (9) and (10): The values of the power loss components estimated in this section are summarized in Table 4 where the enhancement to the DAB efficiency is estimated to be approximately 0.22% given the rated power is 5 kW.

Experimental Validation of the Proposed Loss Model
A DAB test bench was constructed and tests were performed at the nominal operation defined in Table 1 and the results are summarized in Table 5. It was observed that the measured efficiency of the DAB converter composed of MOSFETs with an additional SiC antiparallel SBD is higher than that with only the body diode. The enhancement observed from the experimental results is around 0.3%, which validates the calculation introduced previously resulting in 0.22%. These results validate the proposed hypothesis and approach.

Experimental Results of the PET Module
The developed full-scale PET module is shown in Figure 9. Its structure is that schematically shown in Figure 1b. Table 6 summarize the main components of the PET module.

333
One of the key aspects in the design of this HFT is the high galvanic isolation required between 334 its primary and secondary sides, being 24 kV in this PET. This presents significant challenges 335 compared to the ones considered in literature [43,44]. Moreover, the DAB power transfer inductance 336 (L lk ) is magnetically integrated in the HFT by making use of its series leakage inductance [43]. This 337 reduces size and cost but, on the other hand, imposes additional constrains on the HFT design due 338 to the required accuracy of L lk .

339
The HFT design is, therefore, a tradeoff between four variables: transferred power capability, 340 temperature rise (i.e., losses), size and cost. Several design iterations are performed to achieve the 341 required isolation with an acceptable tradeoff between these variables. The HFT design was 342 previously presented in [45]. The main experimental validation tests are provided in this work for 343 completeness.
344 Figure 10a shows laboratory developed HFT for test purposes and Figure 10b shows the final 345 factory encapsulated HFT. A UU core structure is used typically selected in literature for separate 346 winding [46,47]. The core ferrite material is selected to be Ferroxcube® 3C90 based on the DAB 347 switching frequency [43]. An epoxy resin providing 15 kV/mm and exhibiting acceptable thermal 348 conductance of 3 W/mK was used for encapsulation.

Developed HFT
One of the key aspects in the design of this HFT is the high galvanic isolation required between its primary and secondary sides, being 24 kV in this PET. This presents significant challenges compared to the ones considered in literature [43,44]. Moreover, the DAB power transfer inductance (L lk ) is magnetically integrated in the HFT by making use of its series leakage inductance [43]. This reduces size and cost but, on the other hand, imposes additional constrains on the HFT design due to the required accuracy of L lk .
The HFT design is, therefore, a tradeoff between four variables: transferred power capability, temperature rise (i.e., losses), size and cost. Several design iterations are performed to achieve the required isolation with an acceptable tradeoff between these variables. The HFT design was previously presented in [45]. The main experimental validation tests are provided in this work for completeness. Figure 10a shows laboratory developed HFT for test purposes and Figure 10b shows the final factory encapsulated HFT. A UU core structure is used typically selected in literature for separate winding [46,47]. The core ferrite material is selected to be Ferroxcube ® 3C90 based on the DAB switching frequency [43]. An epoxy resin providing 15 kV/mm and exhibiting acceptable thermal conductance of 3 W/mK was used for encapsulation. Figure 11 shows a schematic representation of the HFT design from ANSYS PEmag ® software (Canonsburg, PA, USA). previously presented in [45]. The main experimental validation tests are provided in this work for 343 completeness. Figure 10a shows laboratory developed HFT for test purposes and Figure 10b shows the final 345 factory encapsulated HFT. A UU core structure is used typically selected in literature for separate 346 winding [46,47]. The core ferrite material is selected to be Ferroxcube® 3C90 based on the DAB 347 switching frequency [43]. An epoxy resin providing 15 kV/mm and exhibiting acceptable thermal

355
To verify the achieved isolation, a high potential test is done for both porotypes and results are 356 compared. A Hipot tester is used to apply a voltage potential of up to 24 kV between both windings.

357
The leakage current flowing from the winding with the higher potential, through air/resin, is 358 recorded and shown in Figure 12. A significant diversion between both prototypes is clear at higher 359 voltage potentials. Having lower leakage currents means that successive partial discharge due to high 360 dv/dt is avoided which would lead to eventual insulation breakdown [48].

363
To locate the hottest spot and verify its temperature rise, five NTC sensors were mounted inside 364 the HFT (see Figure 11). The location of this spot differs in encapsulated (NTC 1) and non-365 encapsulated prototypes (NTC 4). The temperature profile of the encapsulated HFT hottest spot 366 during a four-hour DAB test at rated operation is shown in Figure 13. The steady state temperature, 367 under natural convection, was recorded to be 60 °C. To verify the achieved isolation, a high potential test is done for both porotypes and results are compared. A Hipot tester is used to apply a voltage potential of up to 24 kV between both windings. The leakage current flowing from the winding with the higher potential, through air/resin, is recorded and shown in Figure 12. A significant diversion between both prototypes is clear at higher voltage potentials. Having lower leakage currents means that successive partial discharge due to high dv/dt is avoided which would lead to eventual insulation breakdown [48].

357
To verify the achieved isolation, a high potential test is done for both porotypes and results are 358 compared. A Hipot tester is used to apply a voltage potential of up to 24 kV between both windings.

359
The leakage current flowing from the winding with the higher potential, through air/resin, is 360 recorded and shown in Figure 12. A significant diversion between both prototypes is clear at higher 361 voltage potentials. Having lower leakage currents means that successive partial discharge due to high 362 363

365
To locate the hottest spot and verify its temperature rise, five NTC sensors were mounted inside 366 the HFT (see Figure 11). The location of this spot differs in encapsulated (NTC 1) and non-367 encapsulated prototypes (NTC 4). The temperature profile of the encapsulated HFT hottest spot 368 during a four-hour DAB test at rated operation is shown in Figure 13.  To locate the hottest spot and verify its temperature rise, five NTC sensors were mounted inside the HFT (see Figure 11). The location of this spot differs in encapsulated (NTC 1) and non-encapsulated prototypes (NTC 4). The temperature profile of the encapsulated HFT hottest spot during a four-hour DAB test at rated operation is shown in Figure 13. The steady state temperature, under natural convection, was recorded to be 60 • C.  Temp.
(°C) Figure 13. Experimental results. Hottest spot temperature profile for the encapsulated High Frequency Transformer (HFT) using 3C90 ferrite core measured using a Negative Temperature Coefficient (NTC) sensor.
A summary of the final HFT design is shown in Table 7.  Figure 14 shows the performed test connection diagram. A DC power supply is connected to the LVDC side to provide V dcLV and the power is transferred from FB2 to FB1. The DAB current control regulates the transferred power using single phase shift (SPS) modulation while the CHB full bridge regulates the cell capacitor voltage (V cell ).

371
A summary of the final HFT design is shown in Table 7.  380 Figure 15 shows the experimental results for the PET module nominal operation (described in Table 1). Figure 15a shows the DAB waveforms, where FB2 gates are leading FB1 gates and the phase  Figure 15 shows the experimental results for the PET module nominal operation (described in Table 1). Figure 15a shows the DAB waveforms, where FB2 gates are leading FB1 gates and the phase shift between both controls the magnitude of the transferred power [35]. 382 Figure 15 shows the experimental results for the PET module nominal operation (described in Table 1). Figure 15a shows the DAB waveforms, where FB2 gates are leading FB1 gates and the phase 384 385  Figure 15b shows the CHB waveforms, where V CHB and i CHB are the CHB full-bridge output voltage and current respectively. V cell is regulated at 800 V by the CHB voltage control.

Conclusions
This paper analyzes the use of SiC devices in three-stage modular PETs. It has been shown that SiC MOSFETs can be especially advantageous in the isolation stage, as they combine high blocking voltage and high switching frequencies, leading to higher efficiency, size reduction of the isolation transformer, and higher power density. On the other hand, limited merit is achieved using SiC in the front end AC/DC and the LV side DC/AC converters. However, this partially depends on the application of the PET and the selected topology.
The paper details the selection of SiC MOSFETs for the DAB DC/DC isolation stage experimentally comparing different commercial devices. The benefits of a SiC antiparallel SBD is investigated. It is concluded that the SiC diode improves the efficiency of the DAB as it reduces the total conduction losses of the device. However, this improvement strongly depends on two critical design aspects: the employed dead time interval and the time interval during which the current is driven by both the MOSFET and the diode at the same time (due to the voltage drop in the MOSFET on resistance directly biasing the diode). Generally, the efficiency improvement can be more relevant for a flatter inductor current shape and when the DAB is working around its nominal operation point. A loss model was presented to assess the introduced efficiency improvement and was experimentally validated.
Experimental results showing the operation of the full-scale DAB converter at rated conditions as part of the PET module are provided.   MOSFET effective output capacitance