# Single Active Bridge DC/DC Converter as Auxiliary Power Supply for Medium-Voltage Multilevel Converters

Patrick Lenzen, Martin Pfost

► To cite this version:

Patrick Lenzen, Martin Pfost. Single Active Bridge DC/DC Converter as Auxiliary Power Supply for Medium-Voltage Multilevel Converters. The 26th European Conference on Power Electronics and Applications, GDR SEEDS France & EPE Association, Mar 2025, Paris, France. 10.34746/epe2025-0242. [https://doi.org/10.5088405]

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<figure><img src="image_1.png" /></figure>

Distributed under a Creative Commons Attribution 4.0 International License
Single Active Bridge DC/DC Converter as Auxiliary
Power Supply for Medium-Voltage Multilevel Converters

Patrick Lenzen, Martin Pfost

TU DORTMUND UNIVERSITY

Martin-Schmeisser-Weg 4

44227 Dortmund, Germany

Tel.: +49 231-7557844

E-Mail: patrick.lenzen@tu-dortmund.de

URL: https://ewa.etit.tu-dortmund.de

Acknowledgment

This work was supported by the German Fed-
eral Ministry for Economic Affairs and Climate
Action under grant 03EI6110D. The authors are
responsible for the content of this publication.

Keywords

<<DC-DC converter>> , <<Single Active Bridge>> ,
<<Modular Multilevel Converters (MMC)>>

Abstract

Modular multilevel converters are used in various
applications such as HVDC, STATCOM or PV.
A robust auxiliary power supply is essential to
ensure full control of each stage at all times, even
during failure scenarios. Due to the high voltage
between the cell and the external input power,
high isolation is required. This paper presents
a Single Active Bridge (SAB) auxiliary power
supply for MMCs. The proposed design achieves
high isolation and is capable of powering more
than 100 MMC cells simultaneously. The high
stray inductance resulting from the isolation dis-
tance can be included in the converter design. The
system can handle asymmetric load changes of up
to 24% on a single stage without causing over-
voltage conditions. Furthermore, a novel solution
is presented to address highly asymmetrical load
scenarios.

**Introduction**

Modular multilevel converters (MMCs) have a
wide range of applications. They are used in
HVDC applications, motor drives, PV and STAT-

COM products [1]. The MMC topology reduces
the number of required transformers and reduces
the filter size [2]–[4].

Besides these advantages, one problem is bal-
ancing the voltage across the capacitor voltage
at each module. One solution is presented in [5]
by controlling the dead time of each bridge cell.
Moreover, another challenge is to provide auxili-
ary power to each cell of the MMC. An auxiliary
power supply is required to power the gate driver
and the controller at each level. If the converter
is to be used on the medium-voltage grid, a high
isolation capability according to the grid voltage
is required. In addition, it is necessary to have full
control access to each bridge cell during start-up
or in the event of a grid failure. Therefore, the
auxiliary power supply should be independent of
the medium-voltage grid. Currently, there are no
low-cost commercial DC/DC converters available
with this level of isolation.

Previously, various solutions for auxiliary
power supplies have been proposed, including
wireless charging converters as in [6]. In this
configuration multiple standalone CLLC-CL con-
verters are required to power multiple bridge
cells. For a multilevel converter with a large
number of cells, these solutions require a large
number of converters and primary full bridges.
Furthermore, the additional space required for the
auxiliary supply at each cell is often not avail-
able. Additionally, this topology requires multiple
inductors and capacitors, all of them must be
tuned. Alternatively, [7] employes a LLC with ad-
ditional transformers between the output windings
to address the output voltage stability concerns.
Nevertheless, the integration of these supplemen-
tary windings makes the structure too costly and
complex for an auxiliary power supply.
This paper presents a Single Active Bridge DC/DC converter auxiliary power supply for a medium-voltage modular multilevel converter. In the case of the Single Active Bridge only one primary full bridge and a straightforward secondary sides are necessary. Furthermore, this topology is designed to serve auxiliary power for a 200 kVA MMC. The input voltage of the MMC is 20 kV AC rms. The auxiliary power consumption per cell has been quantified by measurements and is 3.6 W. Nevertheless, it is necessary to consider reserve power and set the required auxiliary power to 5 W. The mean auxiliary power of each cell is typically consistent. In event of a failure, an asymmetric load is to be expected. A potential solution for asymmetric loads is presented in the last chapter. The input voltage of the auxiliary supply would be rectified line voltage, here 325 V. It should be noted that lower or higher voltages are also possible with buck/boost PFCs. The auxiliary supply can be uncontrolled with an output voltage of the auxiliary power supply between 10 V – 27 V. This is the typical input voltage of an off-the-shelf DC/DC converter (with lower isolation capabilities) which is connected to the output of the SAB to generate constant voltages for each gate drive and the microcontroller.

## Proposed System

The Single Active Bridge (SAB) consists of a full bridge on the primary side and of a full bridge rectifier on each cell on the secondary side. For each single cell the primary side and the secondary sides are connected by single transformers, as illustrated in Fig. 1. The series inductor $L_{ser}$ may be implemented as discrete component or alternatively, may be realized as stray inductance of the transformer. The output voltage $V_{o,n}$ is buffered by a capacitor. Furthermore, in order to prevent an overvoltage in the event of an unexpected load reduction, a TVS diode is employed to clamp the output voltage to 27 V. The input voltage of the converter is dependent upon the number of cells within the MMC. In our MMC, 60 cell bridges are anticipated. In order to reduce the primary current, the maximum possible input voltage without the use of a PFC is used. Here, the rectified line power of 325 V is utilized as the input for the SAB, resulting in a voltage for each cell of

$$V_{in,aux} \leq \frac{325 \text{ V}}{60} = 5.4 \text{ V}. \qquad (1)$$

<figure><img src="image_2.png" /><figcaption>Fig. 1: Single Active Bridge with SiC full bridge on the primary side and multiple output rectifier bridges is shown. The primary and secondary sides are connected by ring core transformers.</figcaption></figure>

The design of the SAB is based on a simple transformer design; consequently, this design is presented first.

## Coil Design

The coil design is based on the design proposed by [7]. This will guarantee that the configuration is also free of partial discharge. The toroidal cores of the TDK R63/38/25 model with the N87 material have been selected. A photograph of the ring core is provided in Fig. 2. The primary winding, comprising the high-voltage cable, is routed through the central axis of the core. The secondary windings are wound around the core. Here, litz wire with low isolation capability is possible because the primary winding is already isolated against the HV voltage. Furthermore, the coupling capacitance is simulated with Ansys Maxwell and amounts to 0.12 pF, which is comparable to previously reported designs [8].

<figure><img src="image_3.png" /><figcaption>Fig. 2: The practical realization of a transformer is illustrated in the figure, which depicts the primary winding formed by an HV cable through the center of the core. The cross-section of the cable is marked for reference.</figcaption></figure>
<table><caption>TABLE I: Transformer parameters for different transfer ratios $n_{\text{Tr}}$</caption><thead><tr><th>n<sub>TR</sub></th><th>2</th><th>3</th><th>4</th></tr></thead><tbody><tr><td>Primary ind. L<sub>1</sub> [µH]</td><td>5.03</td><td>5.17</td><td>5.13</td></tr><tr><td>Secondary ind. L<sub>2</sub> [µH]</td><td>20.20</td><td>46.10</td><td>81.30</td></tr><tr><td>Mutual ind. M [µH]</td><td>9.80</td><td>15.04</td><td>19.97</td></tr><tr><td>Coupling coefficient k</td><td>0.973</td><td>0.974</td><td>0.978</td></tr><tr><td>Mutual ind. (APR) L<sub>m</sub> [µH]</td><td>4.76</td><td>4.90</td><td>4.91</td></tr><tr><td>Stray ind. (APR) L<sub>σ</sub> [µH]</td><td>0.27</td><td>0.27</td><td>0.22</td></tr></tbody></table>

In order to ascertain the parameters of the entire converter, the parameters of the transformer are measured. The number of primary turns is set to one because the winding is routed only once through the core. The straightforward concept would no longer be viable if additional primary windings were employed. The number of secondary turns is not fixed and may be modified according to the desired output voltage. Fig. 3 depicts the practical measurement with three secondary windings. The resonance frequency is at 700 kHz. A notable increase in inductance is observed from 200 kHz onwards. Therefore, the switching frequency should be below 200 kHz to ensure a stable operation. Between 10 kHz-200 kHz, the inductance of the transformer is nearly constant. The measurement of the transformer parameters for a varying transfer ratio

$$n_{\text{TR}} = \sqrt{L_2/L_{L1}} \qquad (2)$$

can be found in Tab. I. The values are used to calculate an all-primary-referred (APR) equivalent circuit of a single transformer, see Fig. 4. In the circuit simulation, a single transformer is connected in series for each cell. In order, to get the parameters of the APR circuit, it is first necessary to calculate the coupling coefficient as

$$k = \frac{M}{L_1 L_2} \qquad (3)$$

This can be used to determine the stray inductance $L_{\text{ser}}$ and $L_{\text{m}}$ of the APR circuit as

$$L_{\text{ser}} = (1 - k^2)L_1 \quad \text{and} \qquad (4)$$

$$L_{\text{m}} = k^2 L_1. \qquad (5)$$

The parameters for the different transfer ratios $n_{\text{Tr}}$ are outlined in Tab. I. The parameters of the transformers will be used for the following circuit design.

<figure><img src="image_4.png" /><figcaption>Fig. 3: Transformer parameters as measured with an Keysight E4990A impedance analyzer</figcaption></figure>

<figure><img src="image_5.png" /><figcaption>Fig. 4: All-primary-referred equivalent circuit for one transformer</figcaption></figure>

## Parameter Determination and Control

The design guides of the SAB for duty cycle control is detailed presented in [9], for switching frequency control in [10] and for a combination of duty and switching frequency control in [11]. The duty cycle $d$ is defined as

$$d = \frac{dT}{T} \qquad (6)$$

with $dT$ as the phase shift and $T$ as the switching period, shown in Fig. 5. In all previously named publications the transformer is depicted in an idealized manner. Furthermore, large series inductances in comparison to small series resistances are used.

As illustrated in the section on coil design, on the primary side only a full bridge is mandatory, eliminating the need for an additional $L_{\text{ser}}$ due to the substantial stray inductance of the transformer. However, the series inductance and mutual inductance are insufficient for the ideal calculation. Consequently, simulations are employed to ascertain the parameters. Exemplary waveforms of the proposed system are shown in Fig. 5. The ripple of the current is predominant determined by the mutual inductance and the series resistance.

The subsequent parameters have been predetermined prior to the simulation process. In order to maintain the control system as simple as possible, the auxiliary converter is left uncontrolled. A
<figure><img src="image_6.png" /><figcaption>Fig. 5: The output voltage $V_{o,1}$ of stage 1, the switch-point voltage $v_{zw}$, and the primary transformer current $i_L$ are shown. The load for all stages is $P = 5$ W.</figcaption></figure>

fixed duty cycle and a fixed frequency are employed for the converter. In our case, a switching frequency of 50 kHz is used as the inductances of the transformer remains constant within the specified frequency range, see Fig. 3. It should be noted that the maximum auxiliary power differs for each MMC. Accordingly, the duty cycle $d$ is used to regulate the maximum output voltage $V_{o,n}$ in accordance with the specific system parameters, see Fig. 5. The impact of the duty cycle $d$ on the output power is shown in Fig. 6. In the simulation, a transfer ratio of $n_{Tr} = 2$ is employed. It was found that there was only a slight increase in the output voltage for a duty cycle $d$ that was above 0.3. Therefore, for simulations with the maximum power of $P = 5$ W a duty cycle of 0.3 is used.

<figure><img src="image_7.png" /><figcaption>Fig. 6: Cell output voltage $v_{o,n}$ for a variation of the duty cycle $d = dT/T$ for a system with a transfer ratio $n_{Tr} = 2$, switching frequency $1/T = 50$ kHz, and input voltage $v_{sw,n} = 5.4$ V per cell.</figcaption></figure>

In the simulations, the transfer ratio $n_{Tr}$ is varied in order to obtain the expected output voltage, $v_{o,n}$. The simulations are conducted with

<table><caption>TABLE II: Simulation results for different transfer ratios $n_{Tr}$</caption><thead><tr><td>output voltage $v_{o,n}$</td><td>max current max($i_L$)</td><td>$n_{Tr}$</td></tr></thead><tbody><tr><td>9.01 V</td><td>4.53 A</td><td>2</td></tr><tr><td>14.32 V</td><td>4.50 A</td><td>3</td></tr><tr><td>19.30 V</td><td>4.76 A</td><td>4</td></tr></tbody></table>

the maximum power of $P = 5$ W. The results are presented in Tab. II. The maximal current exhibits only minimal variation because the ripple of the current $i_L$ is predominantly defined by the value of the mutual inductance and the series resistance, as previously stated. Accordingly, a transfer ratio of $n_{Tr} = 3$ is selected because for $n_{Tr} = 2$ the resulting output voltage $v_{o,n}$ is below the previously defined minimum voltage of 10 V. Furthermore, the transfer ratio of $n_{Tr} = 4$ is not selected because the resulting output voltage is too high for solutions where a less power is required. Consequently, the output voltage should be close to the minimum voltage in order to have a margin in the event of a lower auxiliary load.

## Experimental Results

To build up a down-scaled test system, a system with five cells is chosen. For five cells the input voltage $V_{in,aux}$ should be $5 \cdot 5.4$ V = 27 V. Here, a slightly lower power supply with 24 V is used because this is a standardized input voltage. The experimental setup of the SAB DC/DC converter is shown in Fig. 7. For simplicity, the primary winding is realized with stranded wire instead of HV cable. The primary full bridge uses a 1000 V SiC full bridge to make the setup capable of supplying more than 100 MMC cells if using a boost PFC on the input. The full bridge is mounted on a motherboard and controlled by a TI C2000 F280049 microcontroller. The output load of each stage is 40 Ω, which results in 3.6 W power consumption at 12 V, corresponding to the expected auxiliary load of a MMC cell.

Additionally, it is important to minimize ringing in this system to prevent common mode current. This is achieved by zero voltage switching (ZVS) during all switching events, see $v_{sw}$ during the dead time in Fig. 8. In addition, RC snubbers across the secondary diodes reduce voltage ringing across the diodes in the off-state.
<figure><img src="image_8.png" /><figcaption>Fig. 7: Experimental setup of the SAB. For simplicity, the HV cable is replaced by a stranded wire with lower isolation capability. The motherboard with a full bridge is marked with 2. The output stages with the output load are denoted by 1.</figcaption></figure>

<figure><img src="image_9.png" /><figcaption>Fig. 8: The output voltage $V_{o,1}$, $V_{o,2}$ of stage 1 and 2, the switch-point voltage $v_{zw}$, and the primary transformer current $i_L$ are shown. The load for all stages is 40 Ω.</figcaption></figure>

## Asymmetrical Load

The output voltage of all stages is 13 V when the output load is identical in all stages. To evaluate the output voltage variation with a load change, a load reduction of 24% is performed in cell 1, see Fig. 9. The voltage in cell 1 increases from 13 V to 17 V and decreases to 11.6 V in all other cells. The voltages are acceptable because they are within the specified range of 10 V – 27 V. If the load is reduced in more than one cell, the voltage change is even smaller.

<figure><img src="image_10.png" /><figcaption>Fig. 9: The output voltage $V_{o,1}$, $V_{o,2}$ of stage 1 and 2, the switch-point voltage $v_{zw}$, and the primary transformer current $i_L$ are shown. The load for stage 1 is 100 Ω and for all other stages is 40 Ω.</figcaption></figure>

If a high asymmetrical load is required, the output voltage of the cell with low power will exceed 27 V with the previous setup. The TVS diode clamps the voltage, but this will generate losses. To overcome this efficiency loss, the secondary side of each transformer can be changed to a center-tapped transformer, see Fig. 10. Due to the change of the transformer, a different input voltage is required to reach the same output voltage. Here, the input voltage is increased from 24 V to 40 V. However, it is still possible to generate auxiliary power for 100 MMC cells with a boost PFC. The center point of the transformer is connected to ground with a switch. If the output voltage is low, the switch closes and transfer power to the output. Once the output voltage reaches a certain load, the switch will be turned off so that the output voltage does not exceed the limit. This can be easily controlled by a 2-point controller. It must be ensured that the switch is closed during the start-up phase to power the controller.

The simulation results for an auxiliary supply for a five cell MMC are shown in Fig. 11. The switch is open in cell 1 and closed in cells 2-5. The output voltage of stages 2-5 is only slightly increased and does not become unstable. The output voltage of the stage 1 is zero.

<figure><img src="image_11.png" /><figcaption>Fig. 10: Output stage with a center-tapped transformer</figcaption></figure>
<figure><img src="image_12.png" /><figcaption>Fig. 11: Simulation result for a converter with 5 output stages. The switch of output stage 1 is open and the switches of the other stages are closed.</figcaption></figure>

## Conclusions

A Single Active Bridge DC/DC converter is presented as an attractive solution to apply auxiliary power to all stages of a MMC. The simple structure has low ringing because of ZVS. Moreover, a load reduction in one cell of 24% is possible. A further reduction can be achieved by adding an additional switch between the center-tapped transformer and ground. The converter is suitable for high voltage applications with a maximum input voltage of 20 kV AC rms.

## References

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