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dc choppers with circuit breaker functionality for HVDC transmission lines
Electric Power Systems Research 116 (2014) 106–116
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Electric Power Systems Research journal homepage: www.elsevier.com/locate/epsr
Novel dc/dc choppers with circuit breaker functionality for HVDC transmission lines Liangyi Tang a,∗ , Bin Wu b , Venkata Yaramasu b , Weirong Chen a , Hussain S. Athab b a b
School of Electrical Engineering, Southwest Jiaotong University, 610031 Chengdu, China Department of Electrical and Computer Engineering, Ryerson University, Toronto, ON M5B 2K3, Canada
a r t i c l e
i n f o
Article history: Received 10 December 2013 Received in revised form 8 May 2014 Accepted 29 May 2014 Keywords: HVDC circuit breakers dc/dc chopper HVDC transmission Multi-terminal HVDC Closed-loop control
a b s t r a c t This paper proposes two novel dc/dc choppers with dc circuit breaker functionality for HVDC transmission lines. These converters feature step-up and/or step-down functions with instant interruption of shortcircuit faults. Compared to other state-of-the-art converter topologies described in the literature, the proposed topologies have better control functions, lower manufacturing costs and reduced energy losses. A double closed-loop controller is designed to regulate the inductor current and output voltage of the converters. An auxiliary controller is also proposed to ensure proper shut-down of the circuit breaker in the event of dc short-circuit faults. The simulation and experimental results are presented to validate the effectiveness of the proposed converter configurations and control scheme. © 2014 Elsevier B.V. All rights reserved.
1. Introduction 1.1. Research objective As prices of fossil fuels rise and as concern about the dangers of global warming grows, electricity production from renewable energy sources such as wind and photovoltaic (PV) energy conversion systems is a subject of great interest nowadays. Many megawatt-level PV plants and offshore wind farms have been installed around the globe [1,2]. The high voltage direct current (HVDC) transmission is a promising solution to integrate the power generated from these renewable energy sources into the existing power grid. The technological advancements in power electronics have enabled the lower cost and higher performance operation of HVDC transmission lines [3,4]. Earlier HVDC transmission lines were mainly configured as two terminal systems with limited applications, and recently there has been significant interest for multi-terminal HVDC and HVDC grids [5–7]. They allow integration of many energy sources and the tapping of industrial and consumer loads [8,9]. A typical multiterminal HVDC system with such a flexible power integration
∗ Corresponding author at: School of Electrical Engineering, Southwest Jiaotong University, No. 111, Er Huan Lu, Bei Yi Duan, Chengdu, Sichuan 610031, China. Tel.: +86 136 9941 8986. E-mail addresses: [email protected], [email protected] (L. Tang). http://dx.doi.org/10.1016/j.epsr.2014.05.021 0378-7796/© 2014 Elsevier B.V. All rights reserved.
and allocation is shown in Fig. 1. The important technical and operational requirements for multi-terminal HVDC system are summarized below: • Depending on the different power requirements, the tapped loads demand step-up or step-down of the transmission-level dc voltages. When the tapped loads need more power, the output voltage should be forced to step-up, and when some of the tapped loads shut down, the output voltage should be forced to stepdown. For example, a 200 kV dc voltage should be stepped-up to 300 kV or stepped-down to 100 kV. • In the event of short-circuit fault at the low voltage terminal of the HVDC transmission line, the main HVDC transmission lines should be isolated from the fault instantly to ensure the whole grid safety [10]. To fulfill the above two important requirements, dc/dc choppers as depicted in Fig. 1 are used. They perform stepup and step-down functions in addition to the dc circuit breaker (DCCB) functionality. 1.2. Research background Various dc circuit breakers are researched for HVDC systems such as: mechanical DCCB [11], solid-state DCCB [12], hybrid DCCB [13] and power electronic converter based DCCB [14–19]. A comprehensive analysis with fault clearance time, control complexity, costs and losses indicates that the power electronic converter based DCCB is a good choice because it has shorter fault clearance time and also can change the transmission voltage level which is fit for
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Fig. 1. Multi-terminal HVDC system with a dc circuit breaker (dc/dc chopper).
Fig. 2. The four-switch dc/dc choppers employed in HVDC systems [18].
flexible power supply. Recently some scholarly works have been carried-out to integrate the circuit breaker function into the dc/dc choppers [14–19]. The circuit breaker illustrated in Fig. 1 can perform fault clearance in addition to the step-up and/or step-down functions. A multi-functional four-switch dc/dc chopper with short-circuit protection functionality has been analyzed for high power application in [18] and the topology is shown in Fig. 2. Comparing with the mechanical DCCB and the hybrid DCCB, this configuration has more functions [19]. However, this configuration is costly due to the greater number of active switches which are realized by MV-IGBT’s or IGCT’s. Many active switches need to be connected in series and in parallel to handle the high voltage level and high peak fault current. For example, the test system showed in Fig. 1 needs a string with 134 IGBT’s (4.5 kV, 650 A) in series and 3 strings in parallel. Moreover, compared with diodes, the price of IGBT’s or IGCT’s is much higher. The control complexity is also greater as it needs to generate more gating signals. A modified converter configuration is described in [19] with one active switch group replaced by a diode group and the topology is shown in Fig. 3. These two choppers can offer bidirectional power flow. But, most of the ac or dc tapped loads do not need to provide power back to the main HVDC lines [8,9]. In addition, the cost of the DCCB is very high because of the huge number of the active switches and control devices. In order to reduce cost and complexity of DCCB, two new converter configurations are proposed. These configurations deal with the unidirectional power flow and make them more suitable for the HVDC systems. The proposed configurations use less number of active switches compared to the existing ones, and thus they decrease the overall cost, control complexity and power losses. These choppers can step-up and/or step-down the transmissionlevel dc voltages in addition to the circuit breaker functionality.
The comparison between proposed and existing DCCB’s is given in Table 1. The one-switch topology saves 2412 IGBT’s compared the four-switch topology and the two-switch topology saves 1608 IGBT’s. The cost of one IGBT module includes the cost of itself and the control platform which produces the switching signals. The losses of one IGBT module include conduction losses and switching losses which are decided according to the curves in manufacturer’s sheets and the circuit parameters. The losses of the active switches are the main losses in the IGBT-based converters [19]. In addition, the control scheme development is as important as of the converter configuration. In [19], double-loop control algorithm has been proposed for the operation of DCCB. However, this control scheme cannot interrupt and isolate the short-circuit faults completely. To solve this issue, a novel auxiliary controller is also proposed to properly shut-down the chopper in the event of short-circuit faults. The proposed converter configurations and the control scheme are verified through MATLAB simulations and dSPACE DS1103 real-time controller based experiments.
Fig. 3. Three-switch topology [19].
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Table 1 Summary of comparison between proposed and existing DCCB’s. Type of topology
Voltage steppinga
Bi-directional or not
Control complexity
Cost for DCCB
No. of stringsb
Total no. of IGBT’sc
Total no. of diodesd
Four-switch topology Three-switch topology Two-switch topology One-switch topology
Step down and step up Step down Step down and step up Step down
Yes Yes No No
Highest High Medium Low
Highest High Medium Low
8 6 4 2
3216 2412 1608 804
0 804 804 804
a
At the same power flow direction. For both positive and negative poles. For fulfilling the voltage and current level, 134 IGBT’s (5SNA0650J450300, 4.5 kV, 650 A) are connected in series and 3 strings are connected in parallel. d For fulfilling the voltage level and giving a path for the fault peak current, 134 diodes (5SDF07F4501, 4.5 kV, 650 A) are connected in series and 3 strings are connected in parallel. b c
Fig. 4. Proposed dc/dc choppers (circuit breakers) with minimal number of active switches.
1.3. Article structure
2.2. Operating principle of proposed topologies
This paper is organized as follows: the proposed power converter topologies for circuit breaker and their operating principles are discussed in Section 2. The control strategy along with the auxiliary controller is presented in Section 3. The mathematical relation between circuit parameters and total decay time is presented in Section 4. The simulation and experimental results are presented in Sections 5 and 6, respectively. Finally in Section 7, our conclusions are given.
The operating principle of proposed two-switch topology is shown in Fig. 5. During turn-on period, the active switches S1 and S4 are conducting and the energy is stored in dc inductor Ldc . The energy stored in Ldc is discharged to the load through D3 during turn-off period. The mode of operation considering continuous inductor current is shown in Fig. 5c. During one switching period Ts , the average dc voltage across the inductor is zero assuming a loss-less system: Vi DTs + (−Vo )(1 − D)Ts = 0
(1)
2. Proposed converter topologies
From the above expression, the relation between the input and output voltage can be obtained as:
2.1. Configuration of proposed topologies
Vo =
A bipolar dc/dc chopper shown in Fig. 2a was described in [18], where a four-switch chopper as shown in Fig. 2b is connected in cascade to achieve bipolar configuration. In this paper, the proposed converter topologies for circuit breaker function are shown in Fig. 4. Two-switch topology is shown in Fig. 4a and one-switch topology is shown in Fig. 4b.
where D is duty cycle, given by D=
Vi D 1−D
ton Ts
(0 ≤ D ≤ 1)
(2)
(3)
From (2) it can be noticed that the operation of the proposed topology is similar to the buck-boost converter, where the output
Fig. 5. Operating principle of proposed two-switch topology (circuit breaker).
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Fig. 6. Operating principle of proposed one-switch topology (circuit breaker).
Fig. 7. Classical double closed-loop control scheme.
voltage (Vo ) is higher than the input voltage (Vi ) when D > 0.5, and Vo is lower than Vi when D < 0.5. The output voltage equals input voltage when D = 0.5. The operating principle of the one-active-switch topology is shown in Fig. 6. It is a classical buck converter employed as dc circuit breaker. So the following analysis can be carried-out: (Vi − Vo )DTs + (−Vo )(1 − D)Ts = 0
(4)
The input and output voltages can be related as Vo = DVi
(0 ≤ D ≤ 1)
(5)
It can be noticed from (5) that the output voltage of this dc circuit breaker is a fraction of the input voltage.
sent to the controller through low-pass filters (LPF). The lower and upper saturation limits for the iL∗ are set to 0 and rated current of the converter, respectively. The duty cycle saturation limits are set to 0 and 1. The PI controllers are used in both the control loops to regulate the vo and iL . It can be noticed from Fig. 7 that when the converter operates at the rated conditions, the control loops produce the rated iL∗ and D. However, this control scheme cannot distinguish the normal operation and recovery period from the faulty operation of the transmission lines and thus produces gating signals even during the short-circuit faults. In other words, this controller cannot interrupt the short-circuit faults completely. This is an undesirable characteristic of the classical double closed-loop controller.
3. Proposed control scheme
3.2. Proposed control scheme
3.1. Conventional control scheme
To overcome this problem, a novel auxiliary controller is proposed as shown in Fig. 8 to be used in conjunction with the classical controller. The auxiliary controller performs zero crossing detection on the duty cycle, and when the duty cycle value falls below zero, it sends a trigger signal to a two-port switch. With the trigger signal being applied, the two-port switch produces the zero duty cycle value throughout the operation of converter. When the fault is completely isolated/cleared, the trigger signal can be
The conventional double closed-loop controller is shown in Fig. 7, where the outer-loop controls the output voltage (vo ) of the converter, while the inner-loop controls the inductor current (iL ). The outer control loop produces the inductor current reference iL∗ to be used in the inner control loop. The inner loop produces the duty cycle D for the converter. The feedback signals vo and iL are
Fig. 8. Proposed double closed-loop control scheme with additional auxiliary controller.
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Table 2 Parameters used in the simulation and experimental tests. Variable
Description
Simulation
Experimental
Vi Vo Ldc Ci Co RT LT Ro Po fsw Kpv Kiv Kpi Kii
Input dc voltage (V) Output dc voltage (V) dc inductor (mH) Input dc-link capacitance (F) Output dc-link capacitance (F) Transmission resistance () Transmission inductance (mH) Load resistance () (step-down/step-up) Output power (W) (step-down/step-up) Switch frequency (Hz) Vo PI controller proportional gain Vo PI controller integral gain iL PI controller proportional gain iL PI controller integral gain
200 k 100 k/300 k 800 22 47 3 20 600 16.7 M/150 M 1500 0.05 50 0.2 600
100 50/150 10 1000 1500 0.5 1 20 125/1125 1500 0.01 30 0.3 300
disabled to restore normal operation. This algorithm is applicable for the proposed converters and also for the conventional converters in [18,19].
4. Equivalent circuit and total decay time during faults It is important to establish the mathematical relation between the circuit parameters and fault related parameters. In this work, two fault related parameters are defined: fault clearance time tf and total decay time td . The fault clearance (isolation) time tf corresponds to the time between fault occurrence and shutdown of converter (by adjusting duty cycle D to zero with the help of proposed control scheme). Even the converter is isolated from the fault by the proposed control scheme, the inductor current iL becomes zero after a specific time depending upon the converter and transmission line parameters. The time between turn-off of dc circuit breaker (D = 0) and zero inductor current value is defined as the total decay time td . The fault clearance time tf is related with the control parameters and circuit parameters. For the given values of circuit parameters, the fault clearance time is decided by the PI control parameters: Kpv , Kiv , Kpi , and Kii . The Kpv and Kiv are PI controller proportional and integral gain of the output voltage vo adjustment. The Kpi and Kii are PI controller proportional and integral gain of the inductor current iL adjustment. The circuit parameters Ldc and Co , and transmission line parameters RT and LT influence the total decay time td . To derive a mathematical relation between decay time and converter/transmission line parameters, an equivalent circuit is derived during faulty condition and presented in Fig. 9. The equivalent circuit corresponds to the turn-off period of dc circuit breaker (refer to Figs. 5b and 6b), and short circuited (across the dc circuit breaker) transmission line. The energy in the output filter capacitor Co and dc inductor Ldc is dissipated by the transmission line
Fig. 9. Equivalent circuit during a faulty condition.
impedance. According to the Kirchhoff’s current theorem, the node current equation is: iL = iT + iC
(6)
The output filter capacitor current is given by iC =
Co dvo dt
(7)
According to the Kirchhoff’s voltage theorem, the circuit voltage equations are:
vo =
LT diT + RT iT dt
vL + vo =
Ldc diL + vo = 0 dt
(8)
(9)
By substituting (6) and (8) in (9), Ldc diL LT d(iL − iC ) + + RT (iL − iC ) = 0 dt dt
Fig. 10. Simulation results for conventional four-switch topology without the auxiliary controller.
(10)
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Fig. 11. Simulation results for conventional four-switch topology with the auxiliary controller.
Fig. 12. Simulation results for proposed topologies with auxiliary controller during step-down at different transmission line impedance.
By substituting (9) in (7), the capacitor current can be expressed as, iC =
Co dvo L Co d2 iL = − dc 2 dt dt
(11)
The differential equation of iL can be obtained by substituting (11) in (10) as demonstrated below: LT Ldc Co d3 iL L Co RT d2 iL (LT + Ldc )diL + dc + =0 3 dt + RT iL dt dt 2
(12)
According to the characteristic roots of Eq. (12), the time constant of the circuit can be calculated as [20]: =
(−Ldc Co RT ) −
1 3 3
3LT Ldc Co [Ldc Co 3 RT 3 − LT Ldc 2 Co 2 RT (LT + Ldc )] 3
(13)
For the circuit shown in Fig. 13, the total decay time td can be assumed to be 3–4 times the circuit time constant , i.e., td =3 =
(−Ldc Co RT ) −
1 3 3
9LT Ldc Co
[Ldc Co 3 RT 3 − LT Ldc 2 Co 2 RT (LT + Ldc )] 3
(14) 5. Simulation verification and results 5.1. Simulation set-up In order to verify the proposed control scheme and converters, simulations were carried out on a high-voltage system using MATLAB/Simulink software. The parameters of the system are shown in Table 2. The performances of the proposed converters are also compared with that of the four-switch converter introduced in literature [18]. The value of the dc inductor Ldc is selected such that the converter operates in continuous conduction mode with 5% ripple in the inductor current. The value of the output filter capacitor
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Fig. 13. Simulation results for conventional four-switch topology with auxiliary controller during step-down at different circuit parameters Ldc and Co .
Co is chosen such that the output voltage contains 10% ripple under the rated operating conditions of the converter. The transmission line model parameters line resistance RT and line inductance LT are also considered in the simulation tests. A 200 km transmission line is chosen with a line inductance value of 0.11155 mH/km and line resistance value of 0.014 /km [21]. The values of RT and LT used in the simulation tests are 3 and 20 mH, respectively. In the dc system, the line resistance is the main factor of line losses when it operates in normal condition. When the fault occurs, the line inductance prevents the rate of change of fault current. In all the tests, the converter is assumed to be operating in steady-state before the occurrence of short-circuit fault. 5.2. Comparison between conventional and proposed control scheme To validate the auxiliary controller and to demonstrate its significance, a test with conventional double closed-loop controller
(without the auxiliary controller) has been carried out with the conventional four-switch converter, the results are presented in Fig. 10. The step-down operation from 200 kV to 100 kV is considered with a duty cycle value of less than 0.5. A short-circuit fault is applied at the output of the converter at 0.6 s. When a short-circuit fault occurs, the output voltage sharply decreases to zero. The double closed-loop controller forces the converter to turn-off rapidly by bringing the duty cycle D to zero. But as presented in the Fig. 10, the duty cycle increases again after about 10 ms because of the double closed-loop adjustment. This scenario makes the switches on and the current increases again. This leads to an oscillation in the circuit. Because the short-circuit fault is still in circuit, the output voltage vo cannot be recovered back to 100 kV, and the inductor current is maintained at a value higher than its rated value. This is an undesirable operation in HVDC systems as it might cause subsequent faults and destroy the whole system. The results with the auxiliary controller are shown in Fig. 11. The peak value of the fault current is also reduced from 600 A to 400 A
Fig. 14. Simulation results for proposed topologies with auxiliary controller during step-down.
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Fig. 15. Simulation results for existing and proposed topology with auxiliary controller during step-up.
approximately compared to the previous case (Fig. 10) because of instant interruption of short-circuit fault. In addition, the auxiliary controller ensures that the converter is completely shut-down after the duty cycle becomes zero. The fault clearance time tf of the converter is measured as the time between the occurrence of fault and the instant the duty cycle becomes zero. In this test, a fast clearance time of 6.5 ms was achieved. This time is very low compared to the mechanical/hybrid breakers and thus the fault current is limited to a safe value without employing any other current limiters. After the fault is cleared, the system can be recovered back to normal operation by resetting the auxiliary controller. The proposed auxiliary controller can be applied to any dc circuit breakers based on converters which are already described in literatures or the proposed topologies. The analysis presented here also applies for step-up operation of the circuit breakers. 5.3. Analysis of fault clearance time In order to explore the relationship between the fault clearance time tf and transmission line impedance, simulation studies have been conducted and presented in Fig. 12. The transmission line impedance is assumed to be 0.5 pu, 1 pu, and 1.5 pu of the rated impedance. The detailed fault clearance time values are shown in Table 3. The results indicate that as the transmission line impedance increases, the fault clearance time also increases linearly. Furthermore, the circuit parameters Ldc and Co can also influence the fault clearance time value as demonstrated in Fig. 13. The detailed fault clearance time values are included in Table 4. From this analysis, it is observed that as the Ldc and Co values have also linear
Fig. 17. Experimental set-up: (A) dSPACE DS1103 R&D controller; (B) CP1103 connector; (C) interface board; (D) gate drivers; (E) IGBT module; (F) dc inductor; (G) input filter capacitor; (H) output filter capacitor; (I) dc power supply; and (J) voltage and current transducers. Table 3 Relationship between the fault clearance time and the transmission line impedance. Case
1
2
3
4
5
6
ZT (pu) RT () LT (mH) tf (ms)
0.5 1.5 10 5.5
0.75 2.25 15 6.0
1 3 20 6.4
1.25 3.75 25 6.8
1.5 4.5 30 7.2
1.75 5.25 35 7.6
relationship with the fault clearance time. A very low values of Ldc and Co increases the inductor current and output voltage ripple, while high values of Ldc and Co increases the cost of the system. In this work, the Ldc and Co values are chosen with a compromise between the ripple of waveform and cost of the system. 5.4. Proposed topologies with new control scheme 5.4.1. Step-down operation The performance of the proposed topologies during step-down operation is shown in Fig. 14. Similar to the four-switch topology presented in Fig. 11, these converters operate at a duty cycle value Table 4 Relationship between the fault clearance time and the circuit parameters Ldc and Co .
Fig. 16. Block diagram of experimental set-up.
Case
1
2
3
4
5
6
Ldc and Co (pu) Ldc (mH) Co (F) tf (ms)
0.5 400 23.5 4.2
0.75 600 35.25 5.2
1 800 47 6.2
1.25 1000 58.75 7.3
1.5 1200 70.5 8.2
1.75 1400 82.25 9.1
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Fig. 18. Performance of conventional four-switch topology without the auxiliary controller.
Fig. 19. Performance of conventional four-switch topology with the auxiliary controller.
of less than 0.5. The response time of the converters to the shortcircuit faults is almost same as that of four-switch topology. By substituting the converter and transmission line parameters given in Table 2 into (14), the decay time td = 0.294 s can be obtained. This value matches with the simulation results shown in Fig. 14. These results validate that the proposed topologies can perform similar to four-switch topology, however with lower number of number of active switches. 5.4.2. Step-up operation The comparison between the four-switch converter topology and proposed two-switch topology is presented in Fig. 15 for the step-up operation. The output voltage reference is set at 300 kV and thus both the converters operate at a duty cycle higher than 0.5. When the short-circuit fault occurs at 0.6 s, both the converters are completely shut-off with almost identical response times. These results prove that with the auxiliary controller, the proposed converter configurations can perform better dc circuit breaker performance. Compared with the four-switch topology, the proposed converter configurations employ less number of active switches.
6. Experimental verification and results 6.1. Experimental set-up To validate the simulation results presented earlier, a lowpower experimental set-up is developed as shown in Fig. 16. The photograph and main elements of the prototype are shown in Fig. 17. The active switches are implemented using Semikron SKM75GB123D dual-pack IGBT modules. In order to compare and verify the performance of the four-switch topology with conventional and new controllers, two dual-pack IGBT modules are selected. The double closed-loop method and auxiliary controller are implemented using a host PC running with MATLAB-Simulink software through real-time interface (RTI). The output voltage vo and inductor current iL measurements are made using LEM LV25-P and LA55-P transducers, respectively, and sent to dSPACE DS1103 controller through the CP1103 I/O connector. The IGBT gate drivers are based on SKHI22B dual cores which are powered with 0/15 V supply. An interface board with MC14504BCP and TLP521-4 is used between the dSPACE and SKHI22B gate drivers for TTL to CMOS logic
Fig. 20. Experimental results for proposed topologies with auxiliary controller during step-down.
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Fig. 21. Experimental results for the conventional and proposed topology with auxiliary controller during step-up.
conversion and signal isolation, respectively. The DC power supply (Vi ) has been obtained by Xantrex XDC-600-20. The experimental set-up has a limitation on the maximum current of 40 A and the transmission line equivalent parameters RT and LT are chosen such that the fault current is less than the maximum value allowed by the prototype converter. The parameters of the prototype converter are given in Table 2. The four IGBT’s along with their free-wheeling diodes are used to implement different converter topologies such as four-switch converter [18], and proposed two and one active switch converters. For four-switch converter, all the four IGBT’s are utilized by sending four gating signals from the DS1103 controller. For two-switch converter, the gating signals for S2 and S3 are disabled and freewheeling diode D3 is utilized. Similarly for one-switch topology, the gating signals for S2 , S3 and S4 are disabled.
6.2. Comparison between conventional and proposed control scheme The experimental results for a four-switch converter during step-down operation with the conventional double closed-loop controller (without the auxiliary controller) are shown in Fig. 18a, where the response of the system is different from the high-power simulations presented in the previous section. But it carries same meaning that the classical controller initiates the gating signals when the fault is still in the circuit. The low-power simulation results are presented in Fig. 18b to better compare to the experimental results. The experimental results and corresponding low-power simulation results are presented in Fig. 19 by considering the auxiliary controller. During the short-circuit fault, the converter is completely shut-down. These results prove that the auxiliary controller is essential to ensure safe operation of the HVDC system.
6.3. Proposed topologies with new control scheme 6.3.1. Step-down operation The experimental results with the proposed topologies during the step-down operation are presented in Fig. 20. The results with two- and one-switch topologies are presented in Fig. 20a and b, respectively. These converters present similar performance in stepdown operation and in disabling the gating signals during shortcircuit faults. 6.3.2. Step-up operation The experimental results with four-switch topology and the proposed two-switch topology are presented in Fig. 21a and b, respectively. These converters present similar levels of performance, but the proposed topology uses less number of active switches compared to the conventional four-switch topology. The experimental results presented here validate the proposed converter configurations and control scheme. The experimental results are very similar to those of the low power simulation results. 6.3.3. Transition between step-down and step-up operation The proposed two-switch converter can change its operation mode from step-down to step-up and vice versa by simple controller at the same power flow direction. The experimental results are presented in Fig. 22, where a step change in the operation of the converter is applied. As shown in Fig. 22a, the reference output voltage is changed from 50 V to 150 V. The output voltage and inductor current increase so as to force the duty cycle to change from less than 0.5 to greater than 0.5. Similar results are presented for stepup to step-down operation in Fig. 22b. The transition from one operation to another operation has been accomplished smoothly, without any overshoots or oscillations in the system response which is significative for flexible power supply applications.
Fig. 22. Experimental results for proposed two-switch topology during transition between the step-up and step-down operation.
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7. Conclusions This paper proposed an auxiliary controller to ensure the shortcircuit faults shut-down completely for all the relative circuit breaker topologies in the HVDC system. The fault clearance time of the proposed controller is shorter around 6.5 ms compared those of mechanical and hybrid DCCB. Based on dc/dc choppers, this paper proposed two power converter topologies as dc circuit breakers for the HVDC transmission applications. The proposed converters perform multiple functions such as dc circuit breaker and step-up and/or step-down operation. They exhibit similar performance, but with less number of active switches and simpler control scheme, which offer great reduction in the cost, complexity, switching losses and more suitable for HVDC transmission system. Besides, a kind of circuit breaker with simpler controller was developed which can switch between step-up and step-down modes at the same power flow direction, which is advantageous to the flexible power supply and is promising for implementation in the practical HVDC systems. Acknowledgements The authors are thankful for the support from the Laboratory for Electric Drive Applications and Research (LEDAR) and Centre for Urban Energy (CUE) of Ryerson University in Canada. References [1] Renewable Energy Policy Network for the 21st Century (REN21), Renewables Global Status Report, 2013, Available at: http://www.ren21.net (accessed 08.13). [2] Global Wind Energy Council (GWEC), Wind Report, 2013, Available at: http://www.gwec.net (accessed 08.13). [3] O. Gomis-Bellmunt, J. Liang, J. Ekanayake, N. Jenkins, Voltage–current characteristics of multiterminal HVDC-VSC for offshore wind farms, Electr. Power Syst. Res. 81 (2011) 440–450. [4] N. Barberis Negra, J. Todorovic, T. Ackermann, Loss evaluation of HVAC and HVDC transmission solutions for large offshore wind farms, Electr. Power Syst. Res. 76 (2006) 916–927.
[5] D. Jovcic, Offshore wind farm with a series multiterminal CSI HVDC, Electr. Power Syst. Res. 78 (2008) 747–755. [6] O. Gomis-Bellmunt, J. Liang, J. Ekanayake, R. King, N. Jenkins, Topologies of multiterminal HVDC-VSC transmission for large offshore wind farms, Electr. Power Syst. Res. 81 (2011) 271–281. ˜ [7] M. Aragüés-Penalba, A. Egea-Àlvarez, O. Gomis-Bellmunt, S. Andreas, Optimum voltage control for loss minimization in HVDC multi-terminal transmission systems for large offshore wind farms, Electr. Power Syst. Res. 89 (2012) 54–63. [8] A. Ekstrom, P. Lamell, International Conference on HVDC Tapping Station: Power Tapping from a DC Transmission Line to a Local AC Network, AC and DC Power Transmission, 17–20 September 1991, 1991, pp. 126–131. [9] D. Pietersen, 7th AFRICON Conference on Tapping Small Amounts of Power from HVDC Transmission Lines Using Parallel Cascaded Converters, Africa, 15–17 September, 2004, pp. 691–696. [10] C.M. Franck, HVDC circuit breakers: a review identifying future research needs, IEEE Trans. Power Delivery 26 (2011) 998–1007. [11] B. Bachmann, G. Mauthe, E. Ruoss, H.P. Lips, J. Porter, J. Vithayathil, Development of a 500 kV Air blast HVDC circuit breaker, IEEE Trans. Power Appar. Syst. 104 (1985) 2460–2466. [12] J. Magnusson, R. Saers, L. Liljestrand, G. Engdahl, Separation of the energy absorption and over-voltage protection in solid-state breakers by the use of parallel varistors, IEEE Trans. Power Electron. (2014). [13] L. Novello, F. Baldo, A. Ferro, A. Maistrello, E. Gaio, Development and testing of a 10-kA hybrid mechanical – static DC circuit breaker, IEEE Trans. Appl. Supercond. 21 (2011) 3621–3627. [14] Jovcic D., Wu B, Fast fault current interruption on high-power DC networks, in: 2010 IEEE Power and Energy Society General Meeting, 25–29 July, 2010, pp. 1–6. [15] D. Jovcic, B.T. Ooi, Tapping on HVDC lines using DC transformers, Electr. Power Syst. Res. 81 (2011) 561–569. [16] D. Jovcic, B.T. Ooi, Theoretical aspects of fault isolation on high-power direct current lines using resonant direct current/direct current converters, IET Gener. Transm. Distrib. 5 (2011) 153–160. [17] K.A. Corzine, R.W. Ashton, A new Z-source DC circuit breaker, IEEE Trans. Power Electron. 27 (2012) 2796–2804. [18] C. Zhan, C. Smith, A. Crane, A. Bullock, D. Grieve, DC transmission and distribution system for a large offshore wind farm, in: 9th IET International Conference on AC and DC Power Transmission, 19–21 October, 2010, pp. 1–5. [19] M. Hajian, D. Jovcic, B. Wu, Evaluation of semiconductor based methods for fault isolation on high voltage DC grids, IEEE Trans. Smart Grid 4 (2013) 1171–1179. [20] K. Erwin, Advanced Engineering Mathematics, 10th ed., Willey-IEEE Press, John Wiley & Sons, Inc., Jefferson City, USA, 2010. [21] M. Shukr, D.W.P. Thomas, P. Zanchetta, VSC-HVDC transmission line faults location using active line impedance estimation, in: 2012 IEEE International Energy Conference and Exhibition (ENERGYCON), 9–12 September, 2012, pp. 244–248.
Contents lists available at ScienceDirect
Electric Power Systems Research journal homepage: www.elsevier.com/locate/epsr
Novel dc/dc choppers with circuit breaker functionality for HVDC transmission lines Liangyi Tang a,∗ , Bin Wu b , Venkata Yaramasu b , Weirong Chen a , Hussain S. Athab b a b
School of Electrical Engineering, Southwest Jiaotong University, 610031 Chengdu, China Department of Electrical and Computer Engineering, Ryerson University, Toronto, ON M5B 2K3, Canada
a r t i c l e
i n f o
Article history: Received 10 December 2013 Received in revised form 8 May 2014 Accepted 29 May 2014 Keywords: HVDC circuit breakers dc/dc chopper HVDC transmission Multi-terminal HVDC Closed-loop control
a b s t r a c t This paper proposes two novel dc/dc choppers with dc circuit breaker functionality for HVDC transmission lines. These converters feature step-up and/or step-down functions with instant interruption of shortcircuit faults. Compared to other state-of-the-art converter topologies described in the literature, the proposed topologies have better control functions, lower manufacturing costs and reduced energy losses. A double closed-loop controller is designed to regulate the inductor current and output voltage of the converters. An auxiliary controller is also proposed to ensure proper shut-down of the circuit breaker in the event of dc short-circuit faults. The simulation and experimental results are presented to validate the effectiveness of the proposed converter configurations and control scheme. © 2014 Elsevier B.V. All rights reserved.
1. Introduction 1.1. Research objective As prices of fossil fuels rise and as concern about the dangers of global warming grows, electricity production from renewable energy sources such as wind and photovoltaic (PV) energy conversion systems is a subject of great interest nowadays. Many megawatt-level PV plants and offshore wind farms have been installed around the globe [1,2]. The high voltage direct current (HVDC) transmission is a promising solution to integrate the power generated from these renewable energy sources into the existing power grid. The technological advancements in power electronics have enabled the lower cost and higher performance operation of HVDC transmission lines [3,4]. Earlier HVDC transmission lines were mainly configured as two terminal systems with limited applications, and recently there has been significant interest for multi-terminal HVDC and HVDC grids [5–7]. They allow integration of many energy sources and the tapping of industrial and consumer loads [8,9]. A typical multiterminal HVDC system with such a flexible power integration
∗ Corresponding author at: School of Electrical Engineering, Southwest Jiaotong University, No. 111, Er Huan Lu, Bei Yi Duan, Chengdu, Sichuan 610031, China. Tel.: +86 136 9941 8986. E-mail addresses: [email protected], [email protected] (L. Tang). http://dx.doi.org/10.1016/j.epsr.2014.05.021 0378-7796/© 2014 Elsevier B.V. All rights reserved.
and allocation is shown in Fig. 1. The important technical and operational requirements for multi-terminal HVDC system are summarized below: • Depending on the different power requirements, the tapped loads demand step-up or step-down of the transmission-level dc voltages. When the tapped loads need more power, the output voltage should be forced to step-up, and when some of the tapped loads shut down, the output voltage should be forced to stepdown. For example, a 200 kV dc voltage should be stepped-up to 300 kV or stepped-down to 100 kV. • In the event of short-circuit fault at the low voltage terminal of the HVDC transmission line, the main HVDC transmission lines should be isolated from the fault instantly to ensure the whole grid safety [10]. To fulfill the above two important requirements, dc/dc choppers as depicted in Fig. 1 are used. They perform stepup and step-down functions in addition to the dc circuit breaker (DCCB) functionality. 1.2. Research background Various dc circuit breakers are researched for HVDC systems such as: mechanical DCCB [11], solid-state DCCB [12], hybrid DCCB [13] and power electronic converter based DCCB [14–19]. A comprehensive analysis with fault clearance time, control complexity, costs and losses indicates that the power electronic converter based DCCB is a good choice because it has shorter fault clearance time and also can change the transmission voltage level which is fit for
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Fig. 1. Multi-terminal HVDC system with a dc circuit breaker (dc/dc chopper).
Fig. 2. The four-switch dc/dc choppers employed in HVDC systems [18].
flexible power supply. Recently some scholarly works have been carried-out to integrate the circuit breaker function into the dc/dc choppers [14–19]. The circuit breaker illustrated in Fig. 1 can perform fault clearance in addition to the step-up and/or step-down functions. A multi-functional four-switch dc/dc chopper with short-circuit protection functionality has been analyzed for high power application in [18] and the topology is shown in Fig. 2. Comparing with the mechanical DCCB and the hybrid DCCB, this configuration has more functions [19]. However, this configuration is costly due to the greater number of active switches which are realized by MV-IGBT’s or IGCT’s. Many active switches need to be connected in series and in parallel to handle the high voltage level and high peak fault current. For example, the test system showed in Fig. 1 needs a string with 134 IGBT’s (4.5 kV, 650 A) in series and 3 strings in parallel. Moreover, compared with diodes, the price of IGBT’s or IGCT’s is much higher. The control complexity is also greater as it needs to generate more gating signals. A modified converter configuration is described in [19] with one active switch group replaced by a diode group and the topology is shown in Fig. 3. These two choppers can offer bidirectional power flow. But, most of the ac or dc tapped loads do not need to provide power back to the main HVDC lines [8,9]. In addition, the cost of the DCCB is very high because of the huge number of the active switches and control devices. In order to reduce cost and complexity of DCCB, two new converter configurations are proposed. These configurations deal with the unidirectional power flow and make them more suitable for the HVDC systems. The proposed configurations use less number of active switches compared to the existing ones, and thus they decrease the overall cost, control complexity and power losses. These choppers can step-up and/or step-down the transmissionlevel dc voltages in addition to the circuit breaker functionality.
The comparison between proposed and existing DCCB’s is given in Table 1. The one-switch topology saves 2412 IGBT’s compared the four-switch topology and the two-switch topology saves 1608 IGBT’s. The cost of one IGBT module includes the cost of itself and the control platform which produces the switching signals. The losses of one IGBT module include conduction losses and switching losses which are decided according to the curves in manufacturer’s sheets and the circuit parameters. The losses of the active switches are the main losses in the IGBT-based converters [19]. In addition, the control scheme development is as important as of the converter configuration. In [19], double-loop control algorithm has been proposed for the operation of DCCB. However, this control scheme cannot interrupt and isolate the short-circuit faults completely. To solve this issue, a novel auxiliary controller is also proposed to properly shut-down the chopper in the event of short-circuit faults. The proposed converter configurations and the control scheme are verified through MATLAB simulations and dSPACE DS1103 real-time controller based experiments.
Fig. 3. Three-switch topology [19].
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Table 1 Summary of comparison between proposed and existing DCCB’s. Type of topology
Voltage steppinga
Bi-directional or not
Control complexity
Cost for DCCB
No. of stringsb
Total no. of IGBT’sc
Total no. of diodesd
Four-switch topology Three-switch topology Two-switch topology One-switch topology
Step down and step up Step down Step down and step up Step down
Yes Yes No No
Highest High Medium Low
Highest High Medium Low
8 6 4 2
3216 2412 1608 804
0 804 804 804
a
At the same power flow direction. For both positive and negative poles. For fulfilling the voltage and current level, 134 IGBT’s (5SNA0650J450300, 4.5 kV, 650 A) are connected in series and 3 strings are connected in parallel. d For fulfilling the voltage level and giving a path for the fault peak current, 134 diodes (5SDF07F4501, 4.5 kV, 650 A) are connected in series and 3 strings are connected in parallel. b c
Fig. 4. Proposed dc/dc choppers (circuit breakers) with minimal number of active switches.
1.3. Article structure
2.2. Operating principle of proposed topologies
This paper is organized as follows: the proposed power converter topologies for circuit breaker and their operating principles are discussed in Section 2. The control strategy along with the auxiliary controller is presented in Section 3. The mathematical relation between circuit parameters and total decay time is presented in Section 4. The simulation and experimental results are presented in Sections 5 and 6, respectively. Finally in Section 7, our conclusions are given.
The operating principle of proposed two-switch topology is shown in Fig. 5. During turn-on period, the active switches S1 and S4 are conducting and the energy is stored in dc inductor Ldc . The energy stored in Ldc is discharged to the load through D3 during turn-off period. The mode of operation considering continuous inductor current is shown in Fig. 5c. During one switching period Ts , the average dc voltage across the inductor is zero assuming a loss-less system: Vi DTs + (−Vo )(1 − D)Ts = 0
(1)
2. Proposed converter topologies
From the above expression, the relation between the input and output voltage can be obtained as:
2.1. Configuration of proposed topologies
Vo =
A bipolar dc/dc chopper shown in Fig. 2a was described in [18], where a four-switch chopper as shown in Fig. 2b is connected in cascade to achieve bipolar configuration. In this paper, the proposed converter topologies for circuit breaker function are shown in Fig. 4. Two-switch topology is shown in Fig. 4a and one-switch topology is shown in Fig. 4b.
where D is duty cycle, given by D=
Vi D 1−D
ton Ts
(0 ≤ D ≤ 1)
(2)
(3)
From (2) it can be noticed that the operation of the proposed topology is similar to the buck-boost converter, where the output
Fig. 5. Operating principle of proposed two-switch topology (circuit breaker).
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Fig. 6. Operating principle of proposed one-switch topology (circuit breaker).
Fig. 7. Classical double closed-loop control scheme.
voltage (Vo ) is higher than the input voltage (Vi ) when D > 0.5, and Vo is lower than Vi when D < 0.5. The output voltage equals input voltage when D = 0.5. The operating principle of the one-active-switch topology is shown in Fig. 6. It is a classical buck converter employed as dc circuit breaker. So the following analysis can be carried-out: (Vi − Vo )DTs + (−Vo )(1 − D)Ts = 0
(4)
The input and output voltages can be related as Vo = DVi
(0 ≤ D ≤ 1)
(5)
It can be noticed from (5) that the output voltage of this dc circuit breaker is a fraction of the input voltage.
sent to the controller through low-pass filters (LPF). The lower and upper saturation limits for the iL∗ are set to 0 and rated current of the converter, respectively. The duty cycle saturation limits are set to 0 and 1. The PI controllers are used in both the control loops to regulate the vo and iL . It can be noticed from Fig. 7 that when the converter operates at the rated conditions, the control loops produce the rated iL∗ and D. However, this control scheme cannot distinguish the normal operation and recovery period from the faulty operation of the transmission lines and thus produces gating signals even during the short-circuit faults. In other words, this controller cannot interrupt the short-circuit faults completely. This is an undesirable characteristic of the classical double closed-loop controller.
3. Proposed control scheme
3.2. Proposed control scheme
3.1. Conventional control scheme
To overcome this problem, a novel auxiliary controller is proposed as shown in Fig. 8 to be used in conjunction with the classical controller. The auxiliary controller performs zero crossing detection on the duty cycle, and when the duty cycle value falls below zero, it sends a trigger signal to a two-port switch. With the trigger signal being applied, the two-port switch produces the zero duty cycle value throughout the operation of converter. When the fault is completely isolated/cleared, the trigger signal can be
The conventional double closed-loop controller is shown in Fig. 7, where the outer-loop controls the output voltage (vo ) of the converter, while the inner-loop controls the inductor current (iL ). The outer control loop produces the inductor current reference iL∗ to be used in the inner control loop. The inner loop produces the duty cycle D for the converter. The feedback signals vo and iL are
Fig. 8. Proposed double closed-loop control scheme with additional auxiliary controller.
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Table 2 Parameters used in the simulation and experimental tests. Variable
Description
Simulation
Experimental
Vi Vo Ldc Ci Co RT LT Ro Po fsw Kpv Kiv Kpi Kii
Input dc voltage (V) Output dc voltage (V) dc inductor (mH) Input dc-link capacitance (F) Output dc-link capacitance (F) Transmission resistance () Transmission inductance (mH) Load resistance () (step-down/step-up) Output power (W) (step-down/step-up) Switch frequency (Hz) Vo PI controller proportional gain Vo PI controller integral gain iL PI controller proportional gain iL PI controller integral gain
200 k 100 k/300 k 800 22 47 3 20 600 16.7 M/150 M 1500 0.05 50 0.2 600
100 50/150 10 1000 1500 0.5 1 20 125/1125 1500 0.01 30 0.3 300
disabled to restore normal operation. This algorithm is applicable for the proposed converters and also for the conventional converters in [18,19].
4. Equivalent circuit and total decay time during faults It is important to establish the mathematical relation between the circuit parameters and fault related parameters. In this work, two fault related parameters are defined: fault clearance time tf and total decay time td . The fault clearance (isolation) time tf corresponds to the time between fault occurrence and shutdown of converter (by adjusting duty cycle D to zero with the help of proposed control scheme). Even the converter is isolated from the fault by the proposed control scheme, the inductor current iL becomes zero after a specific time depending upon the converter and transmission line parameters. The time between turn-off of dc circuit breaker (D = 0) and zero inductor current value is defined as the total decay time td . The fault clearance time tf is related with the control parameters and circuit parameters. For the given values of circuit parameters, the fault clearance time is decided by the PI control parameters: Kpv , Kiv , Kpi , and Kii . The Kpv and Kiv are PI controller proportional and integral gain of the output voltage vo adjustment. The Kpi and Kii are PI controller proportional and integral gain of the inductor current iL adjustment. The circuit parameters Ldc and Co , and transmission line parameters RT and LT influence the total decay time td . To derive a mathematical relation between decay time and converter/transmission line parameters, an equivalent circuit is derived during faulty condition and presented in Fig. 9. The equivalent circuit corresponds to the turn-off period of dc circuit breaker (refer to Figs. 5b and 6b), and short circuited (across the dc circuit breaker) transmission line. The energy in the output filter capacitor Co and dc inductor Ldc is dissipated by the transmission line
Fig. 9. Equivalent circuit during a faulty condition.
impedance. According to the Kirchhoff’s current theorem, the node current equation is: iL = iT + iC
(6)
The output filter capacitor current is given by iC =
Co dvo dt
(7)
According to the Kirchhoff’s voltage theorem, the circuit voltage equations are:
vo =
LT diT + RT iT dt
vL + vo =
Ldc diL + vo = 0 dt
(8)
(9)
By substituting (6) and (8) in (9), Ldc diL LT d(iL − iC ) + + RT (iL − iC ) = 0 dt dt
Fig. 10. Simulation results for conventional four-switch topology without the auxiliary controller.
(10)
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Fig. 11. Simulation results for conventional four-switch topology with the auxiliary controller.
Fig. 12. Simulation results for proposed topologies with auxiliary controller during step-down at different transmission line impedance.
By substituting (9) in (7), the capacitor current can be expressed as, iC =
Co dvo L Co d2 iL = − dc 2 dt dt
(11)
The differential equation of iL can be obtained by substituting (11) in (10) as demonstrated below: LT Ldc Co d3 iL L Co RT d2 iL (LT + Ldc )diL + dc + =0 3 dt + RT iL dt dt 2
(12)
According to the characteristic roots of Eq. (12), the time constant of the circuit can be calculated as [20]: =
(−Ldc Co RT ) −
1 3 3
3LT Ldc Co [Ldc Co 3 RT 3 − LT Ldc 2 Co 2 RT (LT + Ldc )] 3
(13)
For the circuit shown in Fig. 13, the total decay time td can be assumed to be 3–4 times the circuit time constant , i.e., td =3 =
(−Ldc Co RT ) −
1 3 3
9LT Ldc Co
[Ldc Co 3 RT 3 − LT Ldc 2 Co 2 RT (LT + Ldc )] 3
(14) 5. Simulation verification and results 5.1. Simulation set-up In order to verify the proposed control scheme and converters, simulations were carried out on a high-voltage system using MATLAB/Simulink software. The parameters of the system are shown in Table 2. The performances of the proposed converters are also compared with that of the four-switch converter introduced in literature [18]. The value of the dc inductor Ldc is selected such that the converter operates in continuous conduction mode with 5% ripple in the inductor current. The value of the output filter capacitor
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Fig. 13. Simulation results for conventional four-switch topology with auxiliary controller during step-down at different circuit parameters Ldc and Co .
Co is chosen such that the output voltage contains 10% ripple under the rated operating conditions of the converter. The transmission line model parameters line resistance RT and line inductance LT are also considered in the simulation tests. A 200 km transmission line is chosen with a line inductance value of 0.11155 mH/km and line resistance value of 0.014 /km [21]. The values of RT and LT used in the simulation tests are 3 and 20 mH, respectively. In the dc system, the line resistance is the main factor of line losses when it operates in normal condition. When the fault occurs, the line inductance prevents the rate of change of fault current. In all the tests, the converter is assumed to be operating in steady-state before the occurrence of short-circuit fault. 5.2. Comparison between conventional and proposed control scheme To validate the auxiliary controller and to demonstrate its significance, a test with conventional double closed-loop controller
(without the auxiliary controller) has been carried out with the conventional four-switch converter, the results are presented in Fig. 10. The step-down operation from 200 kV to 100 kV is considered with a duty cycle value of less than 0.5. A short-circuit fault is applied at the output of the converter at 0.6 s. When a short-circuit fault occurs, the output voltage sharply decreases to zero. The double closed-loop controller forces the converter to turn-off rapidly by bringing the duty cycle D to zero. But as presented in the Fig. 10, the duty cycle increases again after about 10 ms because of the double closed-loop adjustment. This scenario makes the switches on and the current increases again. This leads to an oscillation in the circuit. Because the short-circuit fault is still in circuit, the output voltage vo cannot be recovered back to 100 kV, and the inductor current is maintained at a value higher than its rated value. This is an undesirable operation in HVDC systems as it might cause subsequent faults and destroy the whole system. The results with the auxiliary controller are shown in Fig. 11. The peak value of the fault current is also reduced from 600 A to 400 A
Fig. 14. Simulation results for proposed topologies with auxiliary controller during step-down.
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Fig. 15. Simulation results for existing and proposed topology with auxiliary controller during step-up.
approximately compared to the previous case (Fig. 10) because of instant interruption of short-circuit fault. In addition, the auxiliary controller ensures that the converter is completely shut-down after the duty cycle becomes zero. The fault clearance time tf of the converter is measured as the time between the occurrence of fault and the instant the duty cycle becomes zero. In this test, a fast clearance time of 6.5 ms was achieved. This time is very low compared to the mechanical/hybrid breakers and thus the fault current is limited to a safe value without employing any other current limiters. After the fault is cleared, the system can be recovered back to normal operation by resetting the auxiliary controller. The proposed auxiliary controller can be applied to any dc circuit breakers based on converters which are already described in literatures or the proposed topologies. The analysis presented here also applies for step-up operation of the circuit breakers. 5.3. Analysis of fault clearance time In order to explore the relationship between the fault clearance time tf and transmission line impedance, simulation studies have been conducted and presented in Fig. 12. The transmission line impedance is assumed to be 0.5 pu, 1 pu, and 1.5 pu of the rated impedance. The detailed fault clearance time values are shown in Table 3. The results indicate that as the transmission line impedance increases, the fault clearance time also increases linearly. Furthermore, the circuit parameters Ldc and Co can also influence the fault clearance time value as demonstrated in Fig. 13. The detailed fault clearance time values are included in Table 4. From this analysis, it is observed that as the Ldc and Co values have also linear
Fig. 17. Experimental set-up: (A) dSPACE DS1103 R&D controller; (B) CP1103 connector; (C) interface board; (D) gate drivers; (E) IGBT module; (F) dc inductor; (G) input filter capacitor; (H) output filter capacitor; (I) dc power supply; and (J) voltage and current transducers. Table 3 Relationship between the fault clearance time and the transmission line impedance. Case
1
2
3
4
5
6
ZT (pu) RT () LT (mH) tf (ms)
0.5 1.5 10 5.5
0.75 2.25 15 6.0
1 3 20 6.4
1.25 3.75 25 6.8
1.5 4.5 30 7.2
1.75 5.25 35 7.6
relationship with the fault clearance time. A very low values of Ldc and Co increases the inductor current and output voltage ripple, while high values of Ldc and Co increases the cost of the system. In this work, the Ldc and Co values are chosen with a compromise between the ripple of waveform and cost of the system. 5.4. Proposed topologies with new control scheme 5.4.1. Step-down operation The performance of the proposed topologies during step-down operation is shown in Fig. 14. Similar to the four-switch topology presented in Fig. 11, these converters operate at a duty cycle value Table 4 Relationship between the fault clearance time and the circuit parameters Ldc and Co .
Fig. 16. Block diagram of experimental set-up.
Case
1
2
3
4
5
6
Ldc and Co (pu) Ldc (mH) Co (F) tf (ms)
0.5 400 23.5 4.2
0.75 600 35.25 5.2
1 800 47 6.2
1.25 1000 58.75 7.3
1.5 1200 70.5 8.2
1.75 1400 82.25 9.1
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Fig. 18. Performance of conventional four-switch topology without the auxiliary controller.
Fig. 19. Performance of conventional four-switch topology with the auxiliary controller.
of less than 0.5. The response time of the converters to the shortcircuit faults is almost same as that of four-switch topology. By substituting the converter and transmission line parameters given in Table 2 into (14), the decay time td = 0.294 s can be obtained. This value matches with the simulation results shown in Fig. 14. These results validate that the proposed topologies can perform similar to four-switch topology, however with lower number of number of active switches. 5.4.2. Step-up operation The comparison between the four-switch converter topology and proposed two-switch topology is presented in Fig. 15 for the step-up operation. The output voltage reference is set at 300 kV and thus both the converters operate at a duty cycle higher than 0.5. When the short-circuit fault occurs at 0.6 s, both the converters are completely shut-off with almost identical response times. These results prove that with the auxiliary controller, the proposed converter configurations can perform better dc circuit breaker performance. Compared with the four-switch topology, the proposed converter configurations employ less number of active switches.
6. Experimental verification and results 6.1. Experimental set-up To validate the simulation results presented earlier, a lowpower experimental set-up is developed as shown in Fig. 16. The photograph and main elements of the prototype are shown in Fig. 17. The active switches are implemented using Semikron SKM75GB123D dual-pack IGBT modules. In order to compare and verify the performance of the four-switch topology with conventional and new controllers, two dual-pack IGBT modules are selected. The double closed-loop method and auxiliary controller are implemented using a host PC running with MATLAB-Simulink software through real-time interface (RTI). The output voltage vo and inductor current iL measurements are made using LEM LV25-P and LA55-P transducers, respectively, and sent to dSPACE DS1103 controller through the CP1103 I/O connector. The IGBT gate drivers are based on SKHI22B dual cores which are powered with 0/15 V supply. An interface board with MC14504BCP and TLP521-4 is used between the dSPACE and SKHI22B gate drivers for TTL to CMOS logic
Fig. 20. Experimental results for proposed topologies with auxiliary controller during step-down.
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Fig. 21. Experimental results for the conventional and proposed topology with auxiliary controller during step-up.
conversion and signal isolation, respectively. The DC power supply (Vi ) has been obtained by Xantrex XDC-600-20. The experimental set-up has a limitation on the maximum current of 40 A and the transmission line equivalent parameters RT and LT are chosen such that the fault current is less than the maximum value allowed by the prototype converter. The parameters of the prototype converter are given in Table 2. The four IGBT’s along with their free-wheeling diodes are used to implement different converter topologies such as four-switch converter [18], and proposed two and one active switch converters. For four-switch converter, all the four IGBT’s are utilized by sending four gating signals from the DS1103 controller. For two-switch converter, the gating signals for S2 and S3 are disabled and freewheeling diode D3 is utilized. Similarly for one-switch topology, the gating signals for S2 , S3 and S4 are disabled.
6.2. Comparison between conventional and proposed control scheme The experimental results for a four-switch converter during step-down operation with the conventional double closed-loop controller (without the auxiliary controller) are shown in Fig. 18a, where the response of the system is different from the high-power simulations presented in the previous section. But it carries same meaning that the classical controller initiates the gating signals when the fault is still in the circuit. The low-power simulation results are presented in Fig. 18b to better compare to the experimental results. The experimental results and corresponding low-power simulation results are presented in Fig. 19 by considering the auxiliary controller. During the short-circuit fault, the converter is completely shut-down. These results prove that the auxiliary controller is essential to ensure safe operation of the HVDC system.
6.3. Proposed topologies with new control scheme 6.3.1. Step-down operation The experimental results with the proposed topologies during the step-down operation are presented in Fig. 20. The results with two- and one-switch topologies are presented in Fig. 20a and b, respectively. These converters present similar performance in stepdown operation and in disabling the gating signals during shortcircuit faults. 6.3.2. Step-up operation The experimental results with four-switch topology and the proposed two-switch topology are presented in Fig. 21a and b, respectively. These converters present similar levels of performance, but the proposed topology uses less number of active switches compared to the conventional four-switch topology. The experimental results presented here validate the proposed converter configurations and control scheme. The experimental results are very similar to those of the low power simulation results. 6.3.3. Transition between step-down and step-up operation The proposed two-switch converter can change its operation mode from step-down to step-up and vice versa by simple controller at the same power flow direction. The experimental results are presented in Fig. 22, where a step change in the operation of the converter is applied. As shown in Fig. 22a, the reference output voltage is changed from 50 V to 150 V. The output voltage and inductor current increase so as to force the duty cycle to change from less than 0.5 to greater than 0.5. Similar results are presented for stepup to step-down operation in Fig. 22b. The transition from one operation to another operation has been accomplished smoothly, without any overshoots or oscillations in the system response which is significative for flexible power supply applications.
Fig. 22. Experimental results for proposed two-switch topology during transition between the step-up and step-down operation.
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7. Conclusions This paper proposed an auxiliary controller to ensure the shortcircuit faults shut-down completely for all the relative circuit breaker topologies in the HVDC system. The fault clearance time of the proposed controller is shorter around 6.5 ms compared those of mechanical and hybrid DCCB. Based on dc/dc choppers, this paper proposed two power converter topologies as dc circuit breakers for the HVDC transmission applications. The proposed converters perform multiple functions such as dc circuit breaker and step-up and/or step-down operation. They exhibit similar performance, but with less number of active switches and simpler control scheme, which offer great reduction in the cost, complexity, switching losses and more suitable for HVDC transmission system. Besides, a kind of circuit breaker with simpler controller was developed which can switch between step-up and step-down modes at the same power flow direction, which is advantageous to the flexible power supply and is promising for implementation in the practical HVDC systems. Acknowledgements The authors are thankful for the support from the Laboratory for Electric Drive Applications and Research (LEDAR) and Centre for Urban Energy (CUE) of Ryerson University in Canada. References [1] Renewable Energy Policy Network for the 21st Century (REN21), Renewables Global Status Report, 2013, Available at: http://www.ren21.net (accessed 08.13). [2] Global Wind Energy Council (GWEC), Wind Report, 2013, Available at: http://www.gwec.net (accessed 08.13). [3] O. Gomis-Bellmunt, J. Liang, J. Ekanayake, N. Jenkins, Voltage–current characteristics of multiterminal HVDC-VSC for offshore wind farms, Electr. Power Syst. Res. 81 (2011) 440–450. [4] N. Barberis Negra, J. Todorovic, T. Ackermann, Loss evaluation of HVAC and HVDC transmission solutions for large offshore wind farms, Electr. Power Syst. Res. 76 (2006) 916–927.
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