Hybrid HVDC circuit breaker with self-powered gate drives

: The ever increasing electric power demand and the advent of renewable energy sources have revived the interest in high-voltage direct current (HVDC) multi-terminal networks. However, the absence of a suitable circuit breaker (CB) or fault tolerant voltage source converter station topologies with the required characteristics (such as operating speed) have, until recently, been an obstacle in the development of large scale multi-terminal networks for HVDC. This study presents a hybrid HVDC CB concept which is capable of meeting the requirements of HVDC networks. Simulation results are presented which are validated by experimental results taken from a 2.5 kV, 700 A rated laboratory prototype.


Introduction
The interest in high-voltage direct current (HVDC) multi-terminal systems has been revived in recent years. The ever increasing demand for electric power and the advent of renewable energy sources such as off-shore wind power [1][2][3] and solar thermal generation in deserts [4] require an electric transmission system that bridges very long distances with low losses. Most of the HVDC lines realised so far have been point-to-point links [5]. Linking more than two HVDC terminals to form a meshed multi-terminal HVDC network would have several advantages which makes the realisation of HVDC networks very attractive [6,7]. The renewed interest in HVDC networks was initiated by the advances in HVDC technology, such as the availability of the voltage source converter (VSC) HVDC. However, the acceptance of HVDC networks with respect to efficiency, reliability, and controllability will strongly depend on the availability of HVDC circuit breakers (CBs) and fault tolerant VSC topologies making them both important enabling technologies [8][9][10][11]. This paper focuses on a particular type of Hybrid DC CB with self-powered gate drives.
Compared with AC networks, DC current interruption is inherently difficult, mainly due to the absence of a natural current zero crossing in DC systems. Furthermore, DC CBs are required to clear a fault very quickly, typically within a few milliseconds, and to dissipate the large amount of energy which is stored in the system inductances [5]. Commercially existing high voltage (HV) DC switches have operating times of several tens of milliseconds, which although suitable for load current switching, is too slow for clearing faults to meet the needs of a reliable HVDC network [12]. Therefore, new technologies for fast and reliable HVDC CBs are required.
A purely solid-state CB for HVDC networks was proposed and patented in 1997 [13]. Although a very fast response time can be achieved, this arrangement cannot be considered as a practical solution, due to the high steady-state losses which are dissipated in the HV switching branch. In 2011, a hybrid CB solution was demonstrated at CIGRE, Bologna [12]. This topology can be considered as an updated version of that considered in [13], that is, a low voltage (LV) branch consisting of ultra-fast mechanical switch and an insulated gate bipolar transistor (IGBT) is introduced to carry the nominal current and thus avoid the high steady-state losses. The solution presented however, does not have self-powered gate drives for the auxiliary arm. This would be advantageous since powering such circuits in HVDC systems is a well-known challenge. Other solutions have also been proposed in this highly active research area [14][15][16][17].
This paper presents an alternative hybrid HVDC CB topology based on [16]. This topology has a fast response time, thus enabling its use with state of the art switching hardware and furthermore it utilises self-powered gate drives in the HV branch, thus eliminating the need for external gate power supplies for the CB.
Simulation results, using PLECS software, are presented and validated using experimental results taken from a 2.5 kV, 700 A rated laboratory prototype.

Principles of hybrid CBs
A CB is required to accomplish the following basic requirements: (i) large current handling capability without excessive losses; (ii) capability of very fast transition from the conducting to the blocking state in case of a fault, without damaging itself in the process; and (iii) after current interruption, when it is open, the dielectric strength must be built up in order to block any current despite the high potentials at the terminals [18][19][20][21][22][23][24]. Traditional mechanical circuit breakers (MCBs) have a very small contact resistance in the closed position, with galvanic separation in the open state. However, they have a long reaction time due to the need to extinguish the fault in an arc chute [18][19][20][21][22].
Solid-state circuit breakers (SSCBs), based on high-power semiconductors, offer many advantages when compared with conventional solutions. In SSCBs, due to the absence of moving parts, there is no arcing, contact erosion, or bounce. SSCBs are therefore faster, minimising fault duration. However, because of the semiconductor's on-state resistance resulting in heating and power loss, the SSCBs fail as far as the first basic requirement of a CB mentioned above is concerned.
Since a SSCB (MCB) can still perform admirably for one of these requirements where an MCB (SSCB) fails, a parallel combination of semiconductor(s) and an MCB can combine the advantages of both, resulting in a hybrid CB. A hybrid CB allows a combination of the MCB's current-carrying function and SSCB's high-speed arc-less

IET Power Electronics
Special Issue on Converters and Semiconductor Circuit Breakers for HVDC and DC Grids interrupting function. This concept of a hybrid DC CB is shown in Fig. 1a, where the parallel combination of a semiconductor device and mechanical breaker branch is clear.
The normal current flows through the main contact for low power dissipation. The semiconductor device connected in parallel to the main contact operates after a fault occurrence and performs the fault current interruption. It is to be noted that as the semiconductor device is turned off to finally clear the fault, the voltage across it rises due to the stored energy in the line inductance. This voltage may reach a very high value and destroy the hybrid CB if no additional components are used for reducing the voltage. As shown in Fig. 1a, an overvoltage protection device is connected in parallel to the semiconductor device to reduce the peak voltage. Therefore, by limiting the voltage magnitude during turn-off, the energy absorber protects the hybrid CB from damage and also dissipates the energy stored in the line inductance to reduce the current to zero. The hybrid CB consists of three main sections: main LV branch, auxiliary HV branch and extinguishing RCD snubber plus surge arrester branch. A brief description of each part is given in the following sections. Considering a 120 kV HVDC system, Table 1 gives a typical specification for the proposed hybrid HVDC CB [25].

Main branch
The main LV switching branch consists of a semiconductor based commutating switch, for example, an IGBT, and an ultra-fast mechanical switch. A surge arrester A M is employed in parallel with the commutating IGBT, in order to provide a bypass at the beginning of a fault current commutation. Fig. 1c shows the detailed structure of the main branch for unidirectional current flow. The arrangement of the mechanical switch may vary depending on the voltage and switching requirements of the system.
Recent publications indicate that contact opening times for mechanical switches are in the region of 2 ms and that this is likely to be improved in the future [12][13][14][15][16][17]23]. In this paper, it is assumed that the mechanical switch takes 500 µs to operate and a further 300 µs for the contacts to open and support full voltage. These values are based on extrapolation of current state of the art mechanical switch capabilities to the near future. Based on this capability, the shape of the transient recovery voltage produced by the HV surge arrester string can be profiled to be a gentle ramp rather than a steep voltage ramp as in the usual approach, in order to fully utilise the capability of the mechanical switch. Using such a gentle voltage ramp, that is, voltage rising from zero to its target level from the time range 500-800 µs, brings a significantly lower peak current seen by the HV IGBTs and less energy dissipated by the HV surge arresters.
The commutating LV IGBT switch is used to commutate the fault current from the main branch to the auxiliary HV switching branch when a fault occurs, in order to reserve a sufficient time (i.e. at least 500 µs) for an actuator to open the mechanical switch. During normal operation, the current flows only through the main branch, resulting in a relatively low on-state voltage. Therefore, the issue of high steady-state losses, normally associated with solid state devices in the main current path is successfully overcome. When a fault occurs, the LV commutating switch is turned off and the current commutates to the auxiliary HV branch (assuming it is gated at the appropriate time). When the main LV branch is no longer carrying current, the mechanical switch opens, thus protecting the LV switch from the voltage that builds up across the  auxiliary branch. The required voltage rating of the LV switch is thus significantly reduced in comparison with a component that remains in the main current path throughout the switching cycle. Its voltage rating is dictated by the clamping voltage of the parallel-connected surge arrester. Considering the nominal load current level, a suitably rated single or a number of parallel IGBTs can be used to construct the LV commutating switch. The surge arrester, A M , is used to clamp the voltage across the commutating switch at a safe level, it also generates a driving voltage for the current pulse required for the 'self-powered' gate drive circuits of the auxiliary branch.

Auxiliary branch
When a fault is detected in the HVDC grid, the fault current is transferred from the main LV branch to the auxiliary HV branch. After the fault current has been successfully transferred the fast mechanical switch is opened and the cells of the auxiliary branch are turned off to allow the HV surge arresters to extinguish the current.
The counter voltage produced by the HV surge arrester is typically 1.5 times the supply voltage. To withstand it, a large number of HV switch modules need to be connected in series. For a self-powered implementation, it is preferable to use only one series IGBT switch in each module. Hence, considering a 3.3 kV, 1.2 kA IGBT switch and additional 30% headroom, the number of auxiliary modules for each CB, n, can be calculated by Although a large reactor L is employed to limit the rise rate of the fault current, the peak current endured by the auxiliary branch is still quite high, that is, 7.5 kA. To carry this current safely, a suitably rated IGBT or parallel connection of IGBTs is required. An extensive cooling system is not required, since the current will only flow through the auxiliary branch under fault conditions.

Extinguishing branch
Each cell of the HV auxiliary branch is connected in parallel with a capacitive snubber and surge arrester arrangement as shown in Fig. 1d where L S represents a lumped stray inductance of the circuit; R S is a discharge resistor of the RCD snubber; A H , A L , C H and C L represent a HV surge arrester, LV surge arrester, HV capacitor and LV capacitor, respectively. The energy stored in C L is used to gate the devices of the associated HV auxiliary module.

HV surge arrester A H :
The HV arrester, A H, is responsible for producing the counter voltage which in combination will reduce the fault current to zero. Considering the system requirement and total auxiliary module number, the rating of each A H can be determined.

LV surge arrester A L :
The LV arrester, A L, is used to limit the voltage across C L at a safe level to drive LV level electronics. The LV limb (including C L and A L ) in each cell provides a means of powering the gate drive circuitry of the HV IGBTs. This eliminates the need to continuously power the control unit of the auxiliary HV branch.

Stray inductance L S :
When the auxiliary HV branch is switched on, the fault current starts to commutate to it. Once the current is fully transferred, the mechanical switch is turned off at zero current with LV stress. Fig. 2a shows the theoretical current waveform of the CB during the commutation event, where I MAIN and I AUX represent the main branch and auxiliary branch currents, respectively. I COM_1 and I COM_2 represent the initial and post commutation currents. The initial commutation current is given by (2), where I TH and v are the threshold level of fault detection and the rate of rise of fault current, respectively, as given in Table 1; t CHARGING is the charging time of the LV capacitors (the duration after fault detection before the auxiliary branch is gated on).  Fig. 2b shows an equivalent circuit used for the commutation analysis, where V clamp is the clamping voltage of the main branch (i.e. 2.5 kV); nL S is the lumped stray inductance of the n auxiliary modules; V F_ALL is the lumped forward voltage of all the HV IGBTs. From [26], the forward voltage of the selected HV IGBT is taken as 2.8 V. Considering a 20 µs charging time and a total of 80 auxiliary modules, the current commutation time and the post commutation current can be calculated by (3) and (4).
A large stray inductance results in a long commutation time and thus a high post commutation current. As a result, the time left for opening the mechanical switch is reduced accordingly. It may be even shorter than the minimum requirement (800 μs) for opening the mechanical switch when the fault current reaches the maximum breaking threshold (7.5 kA). To reserve sufficient time for an actuator to open the mechanical switch, a relatively short commutation time is required. This provides an upper limit for the allowable stray inductance of each auxiliary module. Due to the large number of modules required, the effect of the IGBT forward voltage drop on the commutation time should be taken into account (2.8 V × 80 = 224 V).
On the basis of these considerations, the stray inductance L S for each auxiliary module is taken to be 0.5 μH. This is a conservative design (i.e. readily achievable in practice), and the value is sufficiently large to highlight its influence in the simulation results.

RCD snubber:
Diodes D S and D P ensure that the energy stored in capacitor C L is protected and reserved for further switching (re-closing) actions. Such an event may be required if the fault has not been cleared. The state of the transmission line can be checked by re-triggering the entire auxiliary arm IGBTs. A faulted line would then be opened again; a safe transmission line would have conduction passed back to the main branch before turning the auxiliary branch off. The selection of R S is only determined by the peak discharging current endured by the HV IGBT, which occurs at the re-closing event, given by The aim of employing the LV limb (including C L and A L ) in each auxiliary module is to provide a means of powering the gate drive circuitry. This obviates the need to continuously power the control unit of the auxiliary HV switching branch with obvious implications for cost reduction. Initially, when the 2 kA threshold level is triggered, the LV IGBT switch of the main branch is switched off. Hence, the current flows through the surge arrester A M as a bypass, creating a charging voltage for each capacitor in the auxiliary branch. Considering one auxiliary module, an equivalent charging circuit can be drawn as shown in Fig. 2c, where V clamp /n is the effective charging voltage; L S represents the lumped stray module inductance; C H and C L are the HV and LV capacitors, respectively.
Once C L is charged, the energy stored in it is used to turn on the associated auxiliary HV switching module (i.e. to provide power to the gate drive and associated circuitry). As each auxiliary module may contain a number of IGBTs in parallel, C L must provide enough charge to gate all of these devices. Fig. 2d shows the voltage waveform assumed on the LV capacitor C L during the IGBTs turn on process (the charge is assumed to be taken from the capacitor at a constant current during the turn-on time). After providing the required gate charge, the voltage across C L , V LV , drops to voltage V LV_E (which must be high enough to assure proper turn-on and correct operation of the associated electronic circuitry).
The C H and C L capacitors are designed by considering the resonant circuit formed by the capacitors and the stray inductance, L S, shown in Fig. 2c. Since C L needs to be quickly charged to enable a fast response time of the auxiliary branch to the occurrence of a fault, the impedance of C H can be set much smaller than that of C L .
With the aim of achieving a compact system volume, the relationship between the two capacitors is further analysed to find the minimum value of the HV capacitor. Assuming the break down voltage of the LV arrester is relatively large, that is, the LV capacitor voltage will not be clamped during the charging event, then the prospective voltage across the two capacitors is given by (6) and (7).
Considering the gate charge requirement, Q, the voltage across the LV capacitor, V LV , drops to the end point voltage V LV_E (i.e. ≥15 V) after the IGBT turn-on process.
From (6) to (8), the relationship between the two capacitors can be derived as As both C L and C H must be positive, the minimum of C L can be found Taking the derivative of C H with respect to C L (9), the corresponding value of C L that yields the minimum value for C H is given by Substituting (11) into (9) yields, To reserve some charge for associated logic gate circuits and in this case assuming four parallel connected IGBTs, a charge of Q equal to 240 µC is considered [26]. The corresponding C H and C L are  (11) and (12), which yields approximately 15 µF for both. Table 2 summarises the design results for each auxiliary module. The charging time of the LV capacitor can also be found Simulation results  closing the fault switch. When a fault threshold is reached, the LV commutating IGBT is turned off. This causes the current to flow through the arrester, A M , as a bypass. The current pulse generated by A M is used to charge the two capacitors of each cell. The energy stored in C L is then used for turning on the HV IGBT in the auxiliary branch. Once the auxiliary branch is turned on, the current is fully commutated to it in a time determined by the value of the stray inductance. The mechanical switch is then turned off at zero current. When the mechanical switch is fully off, the HV IGBTs are turned off to commutate the current to the extinguishing branch, which reduces the current to zero.
Figs. 3b and c show the current waveforms when the CB performs a fault current breaking action, where the curves represent the current through the main branch, the current through the auxiliary HV switch branch, the current through the extinguishing branch and the load side current. A fault occurs at a simulation time of 10 ms, and the current starts to build up at the rate of 10 kA/ms. Once the threshold of 2 kA is met, the LV commutating IGBT is turned off. Therefore, the fault current flows through the surge arrester A M , and the two capacitors of each snubber circuit are charged by a current pulse. Once the auxiliary HV branch is turned on (after the LV capacitor is charged), the fault current is fully commutated to it in about 50 μs, and the mechanical switch can then be opened at zero current with LV stress. After 500 μs, the 80 auxiliary modules are turned off sequentially, in order to commutate the fault current into the extinguishing branch. As a result, the HV surge arrester string forces the fault current to reduce to zero. The total fault clearing time is about 2.75 ms.   Fig. 3d shows the corresponding voltage across the CB during this fault clearing event. The total 80 auxiliary modules have been further divided into four groups. When the mechanical switch contacts begin to open these four groups are sequentially switched off within 300 μs. Therefore, a step-like voltage ramp is applied, which brings a significantly lower peak current seen by the HV IGBTs and less energy dissipated by the HV arresters. Fig. 4 shows the detailed voltage waveform of one auxiliary module, where curves represent the voltage across the auxiliary module, the voltage across C H , the voltage across C L and the voltage across the snubber diode D. Figs. 4b and c show the zoomed-in details indicating that the voltage across C L remains constant hence we can use it for switching on the auxiliary arm cells.
As shown in Fig. 4b, when the extinguishing branch sees the current pulse, the two capacitors are charged to 31 V quickly, due to their identical capacitance. The energy stored in each C L is used to turn on associated auxiliary HV switches. A 3 μs, 40 A current pulse has been employed to represent the IGBT gate charge which  Once the auxiliary HV switches are turned on, the voltage across each module drops to 3 V (i.e. the IGBT forward voltage), and the HV capacitor C H is quickly depleted. Owing to the series diode Ds, C L is protected from being discharged, allowing for a re-triggering operation, if required. At a simulation time of 10.6 ms, as shown in Fig. 4c, the HV IGBTs are turned off (staggered as discussed above). The fault current is then commutated to the associated RCD snubber plus surge arrester branch. C L is quickly charged and the voltage across it is clamped at 55 V by A L , while the HV capacitor C H is charged to the breakdown level of A H . When the voltage produced by the CB equals the supply voltage 120 kV, the fault current reaches its peak (i.e. at time 10.8 ms).
The counter voltage produced by the surge arrester string is 1.5 times the supply voltage V DC (at the peak current) and forces the fault current to reduce to zero. When the fault current is cleared, the voltage across each module is 1.5 kV (i.e. 1.5 kV × 80 = V DC ). Each C L is protected from discharge by the associated D S and D P. Fig. 5a shows the total energy dissipation of the HV surge arrester during the current breaking action. Fig. 5b compares the energy absorption of each auxiliary module in the different groups.
When the mechanical switch is initially open, the 20 auxiliary modules of Group A are first turned off. Following that, the modules of Group B, C and D are sequentially switched off after every 100 μs. Consequently, the HV surge arresters of Group A consume the maximum energy and that of Group D the least.

Validation of the proposed hybrid HVDC CB
A scaled-down laboratory prototype of the proposed hybrid HVDC CB was used to validate the simulation results. The basic test set-up is as shown in Fig. 6a. For simplicity and ease of implementation, a series connection of five HV cells was used for the auxiliary branch, with each HV auxiliary cell using a single IGBT.
Operation of the circuit is as follows. A capacitor bank is charged to the operating voltage (V DC ) from a rectifier source. Following isolation of the source from the capacitor bank, a semiconductor switch is used to emulate a fault and current rises in the main branch (which is already gated on) limited by the inductor. When the threshold current is detected, the LV IGBT of the main branch is turned off. Hence, the fault current flows through the surge arrester, A M , as a bypass resulting in a charging voltage to charge the capacitors C L and C H of each cell. A summary of the parameters used for the experimental validation is given in Table 3.
The LV switching device is implemented using a 3.3 kV/1.2 kA IGBT from Dynex Semiconductor Ltd. A similar device is used to emulate the operation of the ultra-fast mechanical switch. An auxiliary arm comprising five HV IGBTs rated at 1.7 kV, 800 A is used to demonstrate the capabilities of the hybrid HVDC CB. A PCB was used to interface the HV IGBTs with surge arresters and RCD snubber circuit as shown in Fig. 6b.
A self-powered gate drive (SPGD) has been designed using a complete discrete device based solution. In contrast with the standard gate drive, the SPGD has no externally derived power supply but relies on the energy stored locally on the LV capacitor to turn the auxiliary branch HV IGBTs on and off. Fig. 6c shows the circuit diagram of the SPGD used in this work. As discussed before, when the main branch LV IGBT is turned off, a current pulse quickly charges the LV capacitor to about 25 V. During the charging period, the voltage on the LV capacitor is monitored and the status sent to the controller. As soon as the voltage requirement is achieved, the controller sends an acknowledgment signal to the circuit to gate the IGBT. A push-pull arrangement is used to achieve the required current gain to drive the IGBT.

Experimental results
Experimental results are presented in this section to validate the operation of the proposed hybrid HVDC CB. A capacitor bank is charged to the operating voltage (2.5 kV) from a source. Following the isolation of the source from the capacitor bank, the fault switch is engaged and current rises in the main branch (which is already gated on). When the threshold for fault current is detected (200 A in this case), the LV IGBT of the main branch is turned off, while the auxiliary branch remains off. Hence, the fault current flows through the surge arrester A M as a bypass resulting in a charging voltage for the C L and C H capacitors in the auxiliary arm. The charging time is approximately 20-30 µs representing a half cycle of the resonant period created by the series combination of the two 15 µF capacitors and the loop inductance (about 5 µH). No additional inductance was used, since the inherent stray inductance in the construction was adequate. Clearly, in a CB built at full voltage the self-inductance created by the valve layout would likely be large enough that no inductance would need to be addedcare must be taken to limit the inductance, however, since it affects the commutation time from the main branch onto the auxiliary branch.
Once C L is charged to around 20 V, the energy stored in it is used to turn on the associated HV IGBT. After turning on the IGBTs the fault current commutates from the main branch to the auxiliary branch in approximately 20 µs. The commutation time is dictated by the current at the point of commutation (about 300 A), the loop inductance (5 µH) and the clamping voltage of the LV switch (165 V) giving a theoretical value of (600 × 5/165 = 18 µs) which agrees well with the observations. Fig. 7a shows the current through the main branch, auxiliary branch and the voltage across the LV capacitor during the current build-up, resonant charging, current commutation and breaking periods. A detailed view of the current build-up, resonant charging and current commutation events is shown in Fig. 7b. It should be noted that the spikes and notches in the LV capacitor voltage waveform are a result of the high dv/dt and changes in common mode voltage which cannot be rejected by the differential probe used. The results in Figs. 7a and b are consistent with those shown in Figs. 3b and c, and therefore validate these simulation results.
The fault current through the DC reactor and voltage seen by the mechanical switch during the fault clearing event are shown in Fig. 7c. Five cells were used for the auxiliary branch. When the mechanical switch is ready, these five cells are switched off within 400 µs, that is, 100 µs delay between each other. Thus, a step-like voltage ramp is applied as shown in the lower plot of Fig. 7c. Fig. 8 shows the relevant waveforms of the complete hybrid HVDC CB. These include the currents through the main and auxiliary branches, the voltages across the mechanical switch and DC reactor and the voltage across C L and SPGD demand voltage.

Conclusion
This paper has presented a novel hybrid HVDC CB with self-powered gate drives. This CB has a very fast response in breaking fault currents. The auxiliary HV branch switches are gated by using energy stored in the LV capacitors thereby eliminating the need to continuously power the gate drive circuitry of this branch. The operation of the hybrid breaker has been shown through simulations and validated with experimental results taken from a 2.5 kV, 700 A rated scaled down laboratory prototype. Extension of the design for system voltages and currents can be readily achieved by adding more cells in series and parallel to meet the voltage and current requirements, respectively. It should be noted that the clearing time of the proposed technique is strongly affected by the switching time of the disconnector switch and the current state of the art has switching speeds of around 2 ms. The reliability of such a device is subject to further work and it is likely that a crowbar system would be used to bypass faulty cells [16].