# Containerized AC/DC Converter Station as a Building Block for Medium-Voltage DC Grids

Thomas Brückner, Ali Sharaf Addin, Sven Marquardt, Arif Cömert, Florian Zimmert

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Thomas Brückner, Ali Sharaf Addin, Sven Marquardt, Arif Cömert, Florian Zimmert. Containerized AC/DC Converter Station as a Building Block for Medium-Voltage DC Grids. The 26th European Conference on Power Electronics and Applications, GDR SEEDS France & EPE Association, Mar 2025, Paris, France. 10.34746/epe2025-0322. hal-05074017

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# Containerized AC/DC Converter Station as a Building Block for Medium-Voltage DC Grids

Thomas Brückner, Ali Sharaf Addin, Sven Marquardt, Arif Cömert, and Florian Zimmert

UNIVERSITÄT DER BUNDESWEHR MÜNCHEN

Werner-Heisenberg-Weg 39, 85579 Neubiberg, Germany

Email: thomas.brueckner@unibw.de

URL: https://www.unibw.de

## Acknowledgements

This research work has been carried out within the project DEFINE and is funded by dtec.bw – Digitalization and Technology Research Center of the Bundeswehr, which we gratefully acknowledge. dtec.bw is funded by European Union NextGenerationEU. The authors would also like to thank Prof. Rainer Marquardt (emeritus) for inspiring discussions and advice.

## Keywords

«DC grid component», «medium-voltage grid», «MVDC», «modular multilevel converter».

## Abstract

DC grids are one solution to help balance the growing energy demand with distributed renewable generation and storage systems. To simplify the setup of medium-voltage (MV) DC grids, a simple system architecture with few standardized core components is necessary. For the coupling of new MVDC grids to existing AC grids, this paper proposes a design for a modular multilevel converter (MMC) station that could be premanufactured in factory and easily shipped to any site as a building block with up to 40-MW power rating. To achieve the required level of integration and compactness, innovative approaches for component and system-level design need to go hand in hand, as will be illustrated in the paper.

## Introduction

Global warming and limited natural resources require the replacement of fossil fuels by renewable energy sources for electricity generation as well as the transfer of many non-electric applications to become electric, such as e-mobility, heating and industrial processes. Hence, the demand of electric power is predicted to increase by a factor up to ten depending on the prognosis [1].

To integrate renewable sources like wind and solar – being highly volatile by nature – as well as the new active loads and energy storage systems, the electricity grids must not only be reinforced, but change their character from being a passive, mostly unidirectional network to become bidirectional and fully controllable. This transformation has already been started with HVDC transmission lines all over the world and needs to be extended to the medium-voltage (MV) distribution level. Therefore, MVDC grids have become subject of extensive research in recent years [2]–[5]. This includes converter hardware, control, and protection concepts [6]–[8]. Systems for MVDC point-to-point connections are commercially available, which are essentially downscaled from HVDC applications, e.g. [9].

In a first step, MVDC grids can be coupled to the existing AC grid structures where the electricity demand surpasses the AC grid capacity. The DC grids can transport additional energy, stabilize the connected AC grids in their function, and enable the direct connection of DC power sources. To do so, their key requirements are

1. dynamic controllability of power flows,

2. fault tolerance / security of energy supply,

3. improved efficiency.

A concept for a meshed, fault-tolerant DC cable grid that essentially relies on only three types of key components has been proposed in [10]. These components are:

1. Modular Multilevel Converter (MMC) with full-bridge functionality to couple the DC grid to neighboring AC grids,

2. passive DC nodes with disconnector and grounding switches for each connected line,

3. active DC nodes which include the above and an extra power-flow controller (PFC) [11].
<figure><img src="image_2.png" /><figcaption>Fig. 1: General meshed multi-terminal DC grid</figcaption></figure>

A schematic of such a grid is shown in Fig. 1. Ride-through capability in case of cable ground faults is achieved by grid meshing in conjunction with a symmetrical-monopole configuration with high-impedance grounding. Then, ground faults do not lead to short-circuit currents and load currents can be redirected by the means of PFCs, and faulty cables disconnected [10].

The fault-tolerant operation of the DC grid leads to specific requirements for the MMC stations to couple the DC and AC grids (MMC1-3 in Fig. 1). The design of such converter station is discussed in this paper. The requirements for the converter station are detailed in the next section, followed by sections on the electrical system design, the physical station layout and cooling, the design of converter key components, viz. arm inductors and MMC submodules, and, finally, efficiency and cost considerations.

## Requirements for Converter Station

An urban setting has been chosen to define the ratings and environmental conditions, since the increase in demand is expected to be especially high in urban-suburban scenarios. Moreover, in urban settings, space is extremely costly and besides electrical characteristics, aspects like integration and volume play a pivotal role.

The anticipated ratings of the converter station are given in Table I. Power and DC voltage ratings have been assumed to suit the transmission distances in urban/suburban areas and to serve a small district or industrial area, possibly with some renewable power generation. Since the converter will be connected to the AC grid via a transformer, the station could be connected to any of the typical AC voltages in urban grids (typically 10–20 kV).

<table><caption>Table I: Target ratings of converter station</caption><thead><tr><td>Parameter</td><td>Rating class 1</td><td>Rating class 2</td></tr></thead><tbody><tr><td>Rated DC voltage</td><td>20 kV</td><td>20 kV</td></tr><tr><td>Rated DC current</td><td>1 kA</td><td>2 kA</td></tr><tr><td>Rated power</td><td>20 MW</td><td>40 MW</td></tr><tr><td>Max. DC current</td><td>1.5 kA</td><td>3 kA</td></tr><tr><td>Peak power*</td><td>30 MW</td><td>60 MW</td></tr></tbody></table>

* overload for up to 5-10 minutes

## Physical and environmental requirements

The whole installation should fit into an urban setting. This imposes a number of requirements and challenges, among others:

* compact design with minimum footprint

* complete installation enclosed in some kind of building

* accepted exterior design

* high efficiency / minimized cooling

* limited noise level

* fast and easy installation

In an ideal case, the complete converter station would be premanufactured in factory and transported to site as a whole. For a system including a 20- or 40-MW transformer, it becomes intuitively clear that this is not feasible. Therefore, the target is to integrate the converter with all its auxiliaries into one building block, only. This building block, e.g. in form of a container, should be transportable by road to be installed at any site, urban, suburban or rural. Such containerized converter stations are known from solar inverters in the low-MW range, e.g. [12], or more recently from static VAR compensators (SVC) [13]. The transformer must be transported and encased separately, and is not covered in this paper.

## Electrical specification

An MMC with full-bridge functionality is required, as there are no circuit breakers foreseen on the DC side to interrupt potential fault currents. The operating area is defined as follows:

### Rated operation $P_{d,n}$

(defined for power factor $PF = 1$ at point of common coupling - PCC)

* 0.95...1.05·$V_{d,n}$ (DC voltage)

* 1...1.1·$V_{G,n}$ (AC voltage)

* $PF = 0.9...1$ (capacitive/inductive)

### Overload condition $P_{d,max} = 1.5·P_{d,n}$

* 0.95...1.05·$V_{d,n}$ / 1...1.1·$V_{G,n}$
<figure><img src="image_3.png" /><figcaption>Fig. 2: P-Q diagram for rated operation</figcaption></figure>

<figure><img src="image_4.png" /><figcaption>Fig. 3: DC power transfer specification</figcaption></figure>

Delivery and consumption of reactive power are specified in the P-Q diagram, see Fig. 2. In an overload condition, active power transfer has priority over reactive power compensation. Below 95 % DC voltage level, power must be reduced, see Fig. 3, but operation can still be useful in certain grid-fault scenarios.

## Electrical Dimensioning of MMC

An MMC in double-star configuration (Fig. 4) is applied as symmetrical monopole. To provide the required full-bridge functionality with a competitive efficiency, double-zero submodules in double connection (DZDCSM) have been chosen [14]. This submodule topology further enables a fully electronic failure management and electronic protection against explosion as well as a reduction of the capacitor size by 50 % compared to a series connection of two full bridges [15].

A submodule voltage rating of 2× 2 kV has been defined, with an implementation based on 3.3-kV SiC MOSFET devices. (More details on the submodule see below.) This provides a sufficiently

<figure><img src="image_5.png" /><figcaption>Fig. 4: Double-star MMC with DZDCSM</figcaption></figure>

fine granularity of the modulated voltages as well as a submodule rating that might also be considered for higher-voltage grids.

The ratio between the rated AC phase voltage and DC voltage has been chosen as

$$k_n = \frac{(V_{G,n} \times \sqrt{2})}{V_{d,n}/2} = 1.25 \quad (1)$$

which provides low conduction losses and a low energy pulsation in the submodule capacitors. I.e. the rated AC voltage at the converter is

$$V_{LL,n} = \frac{V_{d,n}}{2} \sqrt{3/2} \times k_n = 15.3 kV. \quad (2)$$

The number of submodules per arm has been chosen such that for any combination of AC and DC voltages within the specification range, a utilization factor of

$$r = \frac{V_d}{2} \cdot \frac{k+1}{n \cdot V_C} = 0.9 \quad (3)$$

is not exceeded to provide proper control margin. This factor describes the ratio of the maximum arm voltage required to the sum of the average capacitor voltages in the arm. Under consideration of an n+1 redundancy this leads to a total of nSM = 8 submodules. I.e., in case of a fault of one submodule (then: nSM = 7), full functionality is maintained; in case of a second fault in the same arm, by reducing the DC-link voltage, rated power can still be processed under most conditions, compare Fig. 5.

All converter components must be electrically rated for the overload condition. The SiC MOSFETs and their heatsinks must also be thermally fully dimensioned for the overload condition, while the busbars and to some extent also the arm inductors and submodule capacitors benefit from their longer thermal time constants.
<figure><img src="image_6.png" /><figcaption>Fig. 5: Utilization factor r for different number of submodules $n_{SM}$ per arm ($k_n = 1.25$, $V_{d,n} = 20$ kV, $V_d = 0.8...1.05 \cdot V_{d,n}$, $V_G = 0.9...1.1 \cdot V_{G,n}$)</figcaption></figure>

## Container Design

In order to maintain flexibility in reaching different installation locations, a standard 45-foot high-cube container [16] was chosen as the basic structure. Different structural extensions were carried out to meet the requirements of an MV converter station, see [17], for example:

* water- and dust-tight, non-flammable steel shell with back-ventilated facade and roof to reduce the internal temperature due to solar radiation,

* installation of pressure-relief flaps (safety measure in the event of an arc fault),

* installation of a scaffold on the roof to accommodate the cooling technology,

* stiffeners and mounting rails in the interior (on the ceiling and floor) specially designed for attaching heavy electrical components with reduced vibration,

* access door and maintenance door with a high security standard,

* protection against the leakage of dangerous liquids in the event of an accident (tight tray on the floor).

During the design process, static calculations were carried out to ensure the structural safety and durability of the converter station, considering static load scenarios, such as high wind loads at coastal or high snow loads at mountain regions.

Fig. 6 shows a drawing of the container prototype. The container can be prefabricated, including all electrical and cooling components of the MMC,

<figure><img src="image_7.png" /><figcaption>Fig. 6: Customized converter container</figcaption></figure>

<figure><img src="image_8.png" /><figcaption>Fig. 7: Container prototype on a truck</figcaption></figure>

which ensures a high level of workmanship and avoids error-prone work at the installation site.

The foundation of the container can be realized by six point-shaped gravel beds, which reduces foundation work at the installation site to a minimum. The facade of the prototype is modular and interchangeable, see Fig. 6. After assembly of the scaffold, the converter station attains a height of 3.88 meters. Therefore, the scaffold with the cooling components must be transported separately and mounted on site.

Road transportation of both, the container and the scaffold, is possible without major restrictions. Depending on the country, certain regulations may have to be observed. E.g., in Germany, the road transport of a container of this size requires a special permit. However, such a permit can be issued to a haulage company as a permanent permit in accordance with the German road traffic licensing regulations (StVZO §70).

## Internal MMC set-up

The container shall house the MMC with all its main and auxiliary components. This includes the submodules and (unlike in the known SVC implementations [13]) also the arm inductors. The maximum permissible payload of the container after deducting its dead weight of 8,500 kg is around 23,500 kg. A list of the expected weights
<table><caption>Table II: Component weights for line-up (B)</caption><thead><tr><td>Component</td><td>Weight [kg]</td></tr></thead><tbody><tr><td>submodules, rating class 2 (48x 200kg)</td><td>9,600</td></tr><tr><td>arm inductors (6x 350kg)</td><td>2,100</td></tr><tr><td>busbars / further electrical installation</td><td>1,200</td></tr><tr><td>cooling & air conditioning</td><td>500</td></tr><tr><td>mechanical support systems</td><td>800</td></tr><tr><td><strong>Total</strong></td><td><strong>14,200</strong></td></tr></tbody></table>

of the individual components, given in Table II, shows that component weight is not the primary design issue.

To fit the MMC, i.e., primarily the submodules and the arm inductors, both of which make up the dominant part of the volume, into the given space of the container is far more challenging. Clearance and creepage distances according to IEC 62477-2 [18] must be met.

Two possible internal setups of the MMC have been compared. Fig. 8(A) shows a conventional setup, where the three phases are aligned on top of each other, having their DC terminals at the outer ends and their AC terminals with (possibly coupled) arm inductors in the middle. Each two building blocks next to each other form one double submodule, respectively (cf. Fig. 13), for the 40-MW installation shown in the figure. The cabinets on either end of the container can house a DC disconnector switch and cable terminals, such that the DC cables could directly enter through the container bottom. AC connections would leave the container through the openings in the back wall into an adjacent container (or building) housing the AC transformer.

In the second concept, shown in Fig. 8(B), a linear unwinding of the MMC's B6 topology (cf. Fig. 9) is proposed, as it is commonly used e.g. in B6 diode rectifier stacks. It leads to an arrangement, where insulation voltages between the submodules are limited to the submodule voltage, here 4 kV, except for the connection points to the arm inductors. Each double submodule for the 40-MW installation is formed by the two building blocks on top of each other (cf. Fig. 13). Thus, this concept enables the lowest possible volume for the block of submodules. On the other hand, two DC rails are required to realize the necessary connections within the “stack” of submodules. All connections, AC and DC, are routed through the back wall into the adjacent transformer build-

<figure><img src="image_9.png" /><figcaption>Fig. 8: Comparison of concepts for MMC line-up (for 40-MW power rating)</figcaption></figure>

<figure><img src="image_10.png" /><figcaption>Fig. 9: Linear unwinding of MMC in line-up (B)</figcaption></figure>

ing, which – in this case – also needs to house the DC switchgear and DC terminals.

Table III provides a comparison of both installation concepts with regard to their feasibility and impact on the most important criteria. Based on this comparison, the choice was made to follow line-up (B). Its primary advantage lies within its scalability. Though both concepts allow installations for 20 MW or 40 MW within one container, only line-up (B) enables a design of creepage and clearance distances for 40-kV DC voltage. By using double the number of submodules per converter arm and distributing the structure to two adjacent containers of the same type, a 40-kV system could thus be built, also.

## Thermal Management

Three different cooling systems are considered: liquid cooling for the most lossy MMC components, air cooling for the remaining components, and additional air conditioning depending on the location. Ambient temperatures from –25°C to 40°C are assumed.
<table><caption>**Table III: Comparison of installation concepts**</caption><thead><tr><th rowspan="2">Criterion</th><th rowspan="2">Weight</th><th colspan="2">Line-up</th><th rowspan="2">Remarks</th></tr><tr><th>(A)</th><th>(B)</th></tr></thead><tbody><tr><td>Cooling</td><td>1</td><td>0.5</td><td>0.5</td><td>no substantial pro and cons A vs. B</td></tr><tr><td>Electro-mechanical design</td><td>2</td><td>0.5<br>(-)</td><td>1.5<br>(+)</td><td>main focus: design of submodules and insulation coordination</td></tr><tr><td>Reliability/protection against explosion</td><td>3</td><td>1.5</td><td>1.5</td><td>explosion protection through partitioning in both setups</td></tr><tr><td>Service-ability</td><td>2</td><td>1</td><td>1</td><td>service effort comparable</td></tr><tr><td>Scalability</td><td>2</td><td>0.5<br>(-)</td><td>1.5<br>(+)</td><td>Line-up A: 20 ... 40 MW / 20 kV<br>Line-up B: 20 ... 40 MW / 20 ... 40 kV</td></tr><tr><td>Integration</td><td>3</td><td>1.5</td><td>1.5</td><td>main focus on power part, switch gear in adjacent setting (transformer container)</td></tr><tr><td>Total</td><td>-</td><td>5.5</td><td>7.5</td><td></td></tr></tbody></table>

## Liquid cooling

The power semiconductor devices and the coils of the arm inductors are liquid cooled. Deionized water with a 40-% share of glycol is assumed for calculations, to provide frost protection down to -25°C. Given the requirements for maximum leakage currents between submodules and water distributor and collector pipes, and for a simple layout of cooling hoses, eight heatsinks are cooled serially, which significantly shortens the total length of cooling hoses. This accounts for eight submodules in the 20-MW system, and four submodules in the 40-MW system. The disadvantage is that the cooling conditions for the submodules are uneven, as the coolant temperature builds up progressively along the cooling path. The flow rate is set so that the gradient of the inlet temperatures across the eight heatsinks does not exceed 10 K at rated operating conditions, compare Table VI.

Based on these conditions, a table-type recooler is designed to supply the coolant at 45°C at an ambient temperature of 40°C. The recooler for the 20-MW system with a maximum cooling capacity of 170 kW comprises four 970-mm fans and is accommodated on the roof, along with the other components of the liquid-cooling system, such as the pumps, expansion tank and deionization unit. The cooling system of rating class 2 (40 MW) consists of two identical loops as of the 20-MW system. Cooling requirements for both, rated and overload conditions, can be met with a moderate fan speed maintaining low noise levels below 65 dB. For temporary overload conditions, the fan

<table><caption>**Table IV: Recooler fan parameters**</caption><thead><tr><th>Parameter</th><th>Nominal</th><th>Overload</th></tr></thead><tbody><tr><td>Rotation rate [1/min]</td><td>600</td><td>950</td></tr><tr><td>Air flow [m³/h]</td><td>14,000</td><td>20,000</td></tr><tr><td>Noise level [dB]</td><td>65</td><td>80</td></tr></tbody></table>

<figure><img src="image_11.png" /><figcaption>Fig. 10: Daily temperature cycle due to solar radiation (Abu Dhabi, mid-August condition, cloudless sky [19])</figcaption></figure>

speed can be increased to keep the cooling water inlet temperature at 45°C, see Table IV.

## Air cooling / air conditioning

The power losses from all other equipment inside the container are dissipated into the interior air. Additionally, the external energy input from solar radiation must be considered. To reduce the solar load, the outer walls of the container are clad with back-ventilated wooden facades that allow natural convection. The roof is double-walled. The effectiveness of these measures is verified by a simulation of the interior temperature of the container, whose results are shown in Fig. 10. In the simulation model, the external energy input is considered, exclusively. A subtropical location (Abu Dhabi) is assumed; the container is positioned so that its long walls face east and west. It is shown that the facades can efficiently reduce overheating due to solar radiation.

The interior air-cooling system is designed without refrigeration unit, which results in an interior temperature higher than the ambient temperature, but almost 50 % lower operating costs of the cooling system. It is coupled to the liquid-cooling system of the main components by six air-water cross-flow heat exchangers, each equipped with
<table><caption>Table V: Properties of arm inductor</caption><thead><tr><td>L</td><td>f₀</td><td>m</td><td>R<sub>DC</sub></td></tr></thead><tbody><tr><td>97.4 μH</td><td>1.58 MHz</td><td>350 kg</td><td>1.6 mΩ</td></tr></tbody></table>

<caption>Fig. 11: Design of an arm inductor for 40-MW MMC power rating [21]</caption>

high-capacity ventilators. These heat exchangers are positioned evenly along the container wall, generating a well-distributed air flow in the station. The air-cooling system is designed to ensure that the interior temperature does not exceed the ambient temperature by more than 10 K at any operating conditions. Nevertheless, the converter station can be equipped with a small refrigeration unit to dehumidify the interior air and provide additional cooling as needed.

## Component Design

The limited space in the container requires low-volume space-saving components. Since the largest volumes are taken by the energy-storage elements (submodule capacitors, arm inductors), optimizing their values is the first step for a compact design.

### Arm inductor

The required arm inductance is determined by the current dynamics and thus by the control bandwidth. Employing a fast-reacting control strategy with minimal dead times allows a drastic reduction of the inductance [20], which leads to a nominal arm inductance of $L_e = 100$ μH for the proposed system [21]. A high bandwidth (to avoid additional HF or EMI filters), low weight and low losses are further requirements.

The six arm inductors are suspended under the container ceiling. The given space above the row of submodules requires a flat inductor design, whereas length and width are not that strictly limited.

<caption>Fig. 12: Prototype of DZDCSM for 20-MW MMC class, bipolar terminal voltage of +4 kV, ±2 kV, and 0 kV, weight ≈ 80 kg [22]</caption>

A design to fulfill these requirements, proposed in [21], is shown in Fig. 11.

### Double submodule

All 48 double submodules of the MMC are lined up in a single row, comparable to books in a bookshelf. An optimization calculation confirms the intuitive result that a book-like form factor of the submodules enables the lowest overall volume, given that their large sides face their adjacent modules with low clearance, and the slim sides have highest clearance requirements towards the container walls. Most commercial submodules today also follow this form. The line-up within the length of the container requires an especially thin design in this case.

Two designs of the DZDCSM for the 20-MW and 40-MW installation have been worked out, with cylindrical and prismatic capacitors, respectively. Figure 12 shows the prototype for the 20-MW MMC class, which has been realized first [22]. It is equipped with Infineon's 3.3-kV SiC MOSFET devices (FF2000UXTR33T2M1), and two DC-capacitor banks of 2.05 mF, each. The use of cylindrical capacitors allows an especially slim design and very low inductive commutation paths. The reduced capacitor size is enabled by the submodule topology due to the fact that both capacitors are parallel connected during periods of high arm currents, for typical operating conditions [14], [15].

<figure><img src="image_12.png" /></figure>

<figure><img src="image_13.png" /></figure>
<figure><img src="image_14.png" /><figcaption>Fig. 13: Design options for DZDCSM for 40-MW MMC class, for line-ups cf. Fig. 8</figcaption></figure>

<figure><img src="image_15.png" /></figure>

<table><caption>Fig. 14: Schematic of DZDCSM and semiconductor configuration</caption><thead><tr><th colspan="3">Semiconductor configuration</th></tr><tr><th>System rating</th><th>20 MW</th><th>40 MW</th></tr></thead><tbody><tr><td>T<sub>1</sub>, T<sub>2</sub>, T<sub>3</sub>, T<sub>4</sub></td><td>2x parallel</td><td>3x parallel</td></tr><tr><td>T<sub>51</sub>, T<sub>52</sub>, T<sub>6</sub>, T<sub>7</sub>, T<sub>01</sub>, T<sub>02</sub></td><td>1x</td><td>2x parallel</td></tr></tbody></table>

The submodule for the 40-MW MMC class requires twice the capacitor value and is therefore approximately twice as large. Together with the study of the MMC line-ups, two concepts for the submodule have been compared, with two prismatic 5.15-mF capacitors next to each other or on top of each other, respectively, see Fig. 13. The design for line-up (A) imposes severe difficulties in implementing the required low-inductive buses. This is another reason, why line-up (B) had been selected. The specially-designed capacitors in welded stainless-steel cases comply with the maximum width of 19 cm and feature an ultra-low self-inductance <10 nH.

Figure 14 shows a schematic of the DZDCSM together with the number of parallel semiconductors installed per switch position, for both, 20-MW and 40-MW system rating. The dimensioning of the semiconductors has been primarily focused on high efficiency targets at nominal load.

<table><caption>Table VI: Semiconductor losses, efficiency, and coolant temperatures</caption><thead><tr><th rowspan="2"></th><th colspan="2">20-MW Rating</th><th colspan="2">40-MW Rating</th></tr><tr><th>Nominal</th><th>Overload</th><th>Nominal</th><th>Overload</th></tr></thead><tbody><tr><td>Output power [MW]</td><td>20.0</td><td>34.8</td><td>40.0</td><td>69.7</td></tr><tr><td>Semiconductor losses [kW]</td><td>36.2</td><td>134.4</td><td>84.3</td><td>320.4</td></tr><tr><td>Temperature gradient* [K]</td><td>7.2</td><td>27.0</td><td>8.4</td><td>32.0</td></tr><tr><td>Coolant return temperature [°C]</td><td>52.2</td><td>72.0</td><td>53.4</td><td>77.0</td></tr><tr><td>Semiconductor efficiency [%]</td><td>99.81</td><td>99.61</td><td>99.78</td><td>99.50</td></tr></tbody></table>

<figure><img src="image_16.png" /><figcaption>Fig. 15: Maximum junction temperatures of last submodule in the cooling chain (ambient air $T_a = 40°C$; cooling water inlet $T_{w,in} = 45°C$)</figcaption></figure>

Please note that for the 40-MW system, the switches $T_{51}$ and $T_{52}$ are only installed twice, since they are part of a half-bridge module with $T_6$ and $T_7$, respectively. By using SiC MOSFETs, the losses are highly dominated by conduction losses. The stresses on the various devices within the submodule are very different. Due to an improvement of the semiconductor characteristics within the development phase of the submodule, the MMCs of both ratings can achieve higher overload powers than the original design targets given in Table I, viz. 34.8 MW for the 20-MW system and 69.7 MW for the 40-MW system.

The design for overload operation ($P_{d,max}$) at the thermal limit (design value $T_{j,max} = 135°C$) leads to low junction temperatures at rated conditions ($P_{d,n}$), which is beneficial for a very high efficiency due to the strong temperature gradient of $R_{DS(on)}$. Table VI specifies the loss and cooling data as well as the semiconductor efficiencies for both, rated and overload operation; Figure 15 shows the corresponding maximum junction temperatures. A switching frequency of 150 Hz

<footnote>* temperature gradient of cooling water over all submodules thermally connected in series ($\Delta T_w = T_{w,out} - T_{w,in}$)</footnote>
per double submodule is assumed. Additionally, a thermal chain model is considered, where all eight double submodules within a 20-MW MMC arm or four double submodules within a 40-MW MMC arm are thermally coupled in series, respectively. The cooling water inlet temperature of the thermal chain is held constant at 45°C, also for overload conditions, as discussed earlier.

A superior semiconductor efficiency of up to 99.81 % is reached. The spread of cooling conditions for the chain-linked submodules is moderate with $\Delta T_w < 10$ K at rated power. However, the significant rise of the junction temperatures at overload is partly owed to the chain connection of the submodule heatsinks in the cooling circuit. Increased power ratings could be achieved with a higher degree of parallel cooling paths, though at the expense of a more difficult routing of the cooling hoses and a slightly reduced efficiency.

## Cost-Performance Analysis for Semiconductor Implementation

The implementation of the proposed submodule utilizing SiC devices results in significantly increased semiconductor costs. In order to justify this expense, it is necessary to achieve lower energy costs over the system's lifetime. A cost evaluation of the DZDCSM prototype for the 20-MW MMC class in comparison to conventional Full-Bridge Submodules (FBSM) equipped with Si-IGBTs or SiC-MOSFETs has already been presented in [22]. Following this study, the analysis is extended providing further results with variable output power for both, the 20-MW and 40-MW systems.

For simplicity, and due to its comparably minor impact, the temperature dependence of IGBT on-state characteristics and switching losses is disregarded and IGBT losses evaluated for $T_j = 125°C$, only. Cooling conditions are assumed as specified in the previous section, with $T_j \le 135°C$ at overload conditions for all submodules.

Furthermore, it is assumed that SiC devices are approximately three times more expensive per rated current than Si devices, which appears to be a reasonable estimate for the near term. With an assumed cost of 800 € per 3.3-kV 2-kA single-switch Si-IGBT module (FZ2000R33HE4), the cost per 3.3-kV, 1-kA SiC half-bridge module (FF2000UXTR33T2M1) is estimated as

$$3 \times 800 \text{ €} \times (1 \text{ kA} / 2 \text{ kA}) \times 2 = 2400 \text{ €}.$$

<table><caption>Table VII: Comparison of semiconductor configuration and cost</caption><thead><tr><th>Topology</th><th>20-MW System</th><th>40-MW System</th></tr></thead><tbody><tr><td>Si-2&times;FBSM</td><td>8 modules<br>2&times;T<sub>1, 2, 3, 4</sub><br>(6,400 €)</td><td>16 modules<br>2&times;2&times;T<sub>1, 2, 3, 4</sub><br>(12,800 €)</td></tr><tr><td>SiC-2&times;FBSM</td><td>8 modules<br>2&times;2&times;T<sub>1, 2, 3, 4</sub><br>(19,200 €)</td><td>12 modules<br>2&times;3&times;T<sub>1, 2, 3, 4</sub><br>(28,800 €)</td></tr><tr><td>SiC-DZDCSM</td><td>8 modules<br>2&times;T<sub>1, 2, 3, 4</sub><br>1&times;T<sub>51, 52, 6, 7</sub><br>1&times;T<sub>01, 02</sub><br>(19,200 €)</td><td>14 modules<br>3&times;T<sub>1, 2, 3, 4</sub><br>2&times;T<sub>51, 52, 6, 7</sub><br>2&times;T<sub>01, 02</sub><br>(33,600 €)</td></tr></tbody></table>

<figure><img src="image_17.png" /><figcaption>Fig. 16: Calculated semiconductor efficiencies for 20-MW and 40-MW system applying different sub-module topologies and semiconductor technologies</figcaption></figure>

Semiconductor configurations and cost for the different submodules are given in Table VII. Since one DZDCSM corresponds to two FBSM, the results are given for a pair of FBSM.

Figure 16 shows the calculated semiconductor efficiencies, where the solid curves correspond to the 20-MW system and the dotted curves to the 40-MW system. The stars denote the rated power of each system. The substitution of IGBTs by SiC devices in the FBSM cuts down losses by 50 % and more. In addition, the use of the DZDCSM offers a further reduction of up to 50 % compared to the SiC-equipped FBSM, achieving unsurpassed efficiency. The end of each curve is determined by reaching $T_j = 135°C$ of the hottest switch within the last submodule in the cooling chain.
<figure><img src="image_18.png" /><figcaption>Fig. 17: Cost per double submodule per MW output power for 20-MW and 40-MW system</figcaption></figure>

Figure 17 provides the relative cost per double submodule (one DZDCSM or two FBSMs) per MW output power, considering the semiconductor expense plus a rate of 6 €/W for loss capitalization. These results are, of course, highly dependent on the loss capitalization rate, which can vary from project to project to a large extent. In the given scenario, the replacement of IGBTs by SiC devices in the FBSM does not make much difference, whereas the application of the DZDCSM enables a significant cost reduction. Moreover, it can be observed that the lowest costs are achieved at about 30 MW for the smaller and 55 MW for the larger system. In other words, it would be beneficial to shift the rated operating points from the original targets to higher values.

The analysis in the previous chapter shows potentials for a power increase with regard to the installed SiC devices. To elaborate this further, improved cooling circuits should be investigated. Moreover, all ratings calculated in the paper are based on worst-case ambient conditions. Operation in moderate climates would also allow for higher power. However, the total capacitance of all submodule capacitors, which occupy the largest share of the container volume, is the limiting factor for an increase of continuous output power.

## Conclusion

The design of a containerized MMC station for connecting future MVDC grids to existing AC grids has been presented. The use of advanced control and submodule technologies as well as new approaches to inductor design enable the complete integration (including the arm inductors) of a highly efficient MMC, designed for 40 MW rated power, in an easily transportable 45-foot container. The use of such standardized, prefabricated building blocks can help to facilitate the installation of future DC grids and accelerate grid expansion.

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