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Progress of the ITER NBI acceleration grid power supply reference design
Fusion Engineering and Design 88 (2013) 956–959
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Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Progress of the ITER NBI acceleration grid power supply reference design Vanni Toigo a , Loris Zanotto a,∗ , Marco Bigi a , Hans Decamps b , Alberto Ferro a , Elena Gaio a , Daniel Gutiérrez c , Kazuki Tsuchida d , Kazuhiro Watanabe d a
Consorzio RFX, Associazione EURATOM-ENEA sulla Fusione, Corso Stati Uniti 4, I-35127 Padova, Italy ITER Organization, Route de Vinon sur Verdon, 13115 Saint Paul Lez Durance, France Fusion For Energy, C/Josep Pla 2, 08019 Barcelona, Spain d Japan Atomic Energy Agency, 801-1 Mukoyama, Naka, Ibaraki-ken 311-0193, Japan b c
h i g h l i g h t s This paper reports the progress in the reference design of the Acceleration Grid Power Supply (AGPS) of the ITER Neutral Beam Injector (NBI) A critical revision of the main design choices is presented in light of the definition of some key interface parameters between the two AGPS subsystems. The verification of the fulfillment of the requirements in any operational conditions is reported and discussed.
a r t i c l e
i n f o
Article history: Available online 12 February 2013 Keywords: ITER NBI Power Supply
a b s t r a c t This paper reports the progress in the reference design of the Acceleration Grid Power Supply (AGPS) of the ITER Neutral Beam Injector (NBI). The design of the AGPS is very challenging, as it shall be rated to provide about 55 MW at 1 MV dc in quasi steady-state conditions; moreover, the procurement of the system is shared between the European Domestic Agency (F4E) and the Japanese Domestic Agency (JADA), resulting in additional design complication due to the need of a common definition of the interface parameters. A critical revision of the main design choices is presented also in light of the definition of some key interface parameters between the two AGPS subsystems. Moreover, the verification of the fulfillment of the requirements in any operational conditions taking into account the tolerance of the different parameters is also reported and discussed. © 2013 Elsevier B.V. All rights reserved.
1. Introduction The Acceleration Grid Power Supply (AGPS) is the most powerful among the power supply systems [1] of the ITER Neutral Beam Injector (NBI) [2,3]. It has to feed the multistage grid system to accelerate negative ions up to −1 MeV (Deuterium) or −870 keV (Hydrogen), with a beam current of 40 A or 46 A respectively. The challenging requirements, both in terms of ratings and of specific constraints related to the particular application called for in-depth studies to outline the reference design and to advance the iterative process of cross-check verifications toward its assessment. The development of the AGPS reference scheme was carried out starting from the analysis of the existing NBIs rated for the highest voltages. The NBI AGPS scheme closest to the ITER one in terms of requirements was the JT-60 U one, based on negative ions and designed for a voltage up to 500 kV dc, half of the total acceleration voltage required for ITER [4–6].
∗ Corresponding author. Tel.: +39 049 8295898. E-mail address: [email protected] (L. Zanotto). 0920-3796/$ – see front matter © 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.fusengdes.2013.01.013
With respect to the JT-60 U scheme, some modifications and optimizations have been introduced in the AGPS reference scheme for ITER to cope with the more demanding requirements and to take advantage from the latest technologies available on the market. The ITER AGPS reference scheme, shown in Fig. 1, consists of an ac/dc stage feeding via dc links five Neutral Point Clamped (NPC) three-phase inverters [7], each connected to a step-up transformer feeding a six-pulse three phase bridge rectifier; the rectifiers are connected in series at the output side to obtain the nominal acceleration voltage of −1 MV, with availability of the intermediate potentials [6,8]. The procurement of the AGPS is shared between the European Domestic Agency (F4E), which is providing the low voltage power conversion equipment, namely AGPS Conversion System (AGPS-CS), and the Japanese Domestic Agency (JADA), providing the high voltage part, downstream of the inverter units, called AGPS Direct Current Generator (AGPS-DCG). A strong integration effort in the development of the design between ITER Organization (IO), JADA, F4E and Consorzio RFX, is necessary, as the selection of the scheme and the choice of parameters on one subsystem has a strong technical and cost impact on the other and
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
957
Table 2 AGPS dynamic and precision requirements.
Fig. 1. AGPS reference scheme.
vice versa. To study the operation of the AGPS reference scheme and to verify the fulfillment of the requirements, a first set of parameters was assumed and a detailed numerical model was developed; these studies are described in [7,9]. After recalling the main requirements of the AGPS, this paper presents the progress on the definition of some key interface parameters between the two subsystems and, in this respect, a critical revision of the main design choices, in particular the selection of the inverter and input stage topologies and of the inverter modulation technique. Moreover, the choice of the dc-link voltage, turn ratio and short-circuit impedance of the step-up transformer is justified by analyzing the performance with respect to the variation of operational conditions and parameters tolerance. 2. Main requirements The main requirements of the system have been discussed in detail in [7]. The AGPS shall provide the currents and voltages reported in Table 1, for Deuterium and Hydrogen operation. On top of the beam power, the AGPS must compensate for the losses due to the stripped electrons and to the co-extracted electrons, which generate additional current on the grids. The total rated power delivered by the AGPS shall be about 55 MW for both Deuterium and Hydrogen operation. From the requirements reported in Table 1, it follows that the step-up transformers shall be insulated for up to −1 MV dc between primary and secondary winding. Diode rectifiers and filters must be insulated for the same voltage as well. The ratings of each AGPS stage can be derived from the table increasing the maximum current up to 66 A to take into account a margin of about 5%. The AGPS shall operate within 20%–100% of the full output voltage. The regulation of the output voltage is obtained by proper control of the NPC inverter. Precision and dynamic requirements at
Parameter
Value
Voltage regulation range Voltage resolution Output voltage accuracy for 1 h operation Maximum voltage fluctuation Maximum Voltage Ripple Rise time of the output voltage Max. settling time
20%–100% 1 kV ±2% ±2.5% at flat-top ±5% 80 ms 50 ms
beam turn-on and after a breakdown of the grids are summarized in Table 2. Note that the design shall guarantee an output voltage rise time of 80 ms, a settling time of 50 ms and a voltage ripple within ±5% of the output voltage. Moreover, in case of breakdown, the AGPS must switch off within 150 s from the breakdown (50 s for the detection plus 100 s for turning-off) to protect the grids and, after a suitable time (20 ms), the voltage must be reapplied with an appropriate ramp up to the set-point, with the dynamic and precision already described. An additional important requirement for the design of the insulation of the high voltage components is that in case of loss of inverter control, i.e. if the inverters fail to regulate the output voltage, the AGPS output voltage must never exceed the nominal value of the insulation voltage to ground (1.2 MV) in every possible condition. 2.1. Interface parameters A step forward in the assessment of the interface between AGPSCS and AGPS-DCG has been done in 2011 and, in particular, a list of key parameters has been defined and agreed; Table 3 summarizes them. The choice of their value is crucial, as explained in the following sections, to address the design of both systems towards the fulfillment of the requirements, and, at the same time, to guarantee their feasibility and optimize their procurement. 3. Critical revision of the reference scheme On the basis of the definition of the interface parameters, a critical revision of the reference scheme has been performed to evaluate the adequacy and convenience of the main design choices. 3.1. Inverter scheme The inverter design is based on a NPC scheme whereas in JT60 U it was based on single phase H-bridges [10]. The selection of the NPC three-phase scheme has several advantages. The NPC scheme, based on the present availability of switching components rated for higher voltage, allows the increase of the dc-link voltage with respect to the H-bridge, with consequent reduction of the inverter rated current and consequent reduction of losses on inverter switches; moreover, the primary winding of the stepup transformers can be delta connected instead of open delta,
Table 1 Voltage and current rating of the grids. Parameter
Main supply Grid 1 Grid 2 Grid 3 Grid 4 Grounded conductora
D− , 1 MeV, 40 A
H− , 870 keV, 46 A
Table 3 AGPS main interface parameters.
Voltage
Current
Voltage
Current
Parameter
Value
−1000 kV −800 kV −600 kV −400 kV −200 kV 0 kV
59.4 A 3.0 A 5.3 A 3.1 A 2.7 A 45.3 A
−870 kV −696 kV −522 kV −348 kV −174 kV 0 kV
62.8 A 2.4 A 4.3 A 3.1 A 2.1 A 50.9 A
dc-Link voltage dc-Link voltage max. variation in stationary regime Inverter topology Inverter output frequency Inverter total output inductance Step-up transformer short-circuit voltage Step-up transformer turn-ratio Step-up transformer winding connections
6500 V ±6% NPC 150 Hz 110H 12% ± 1.2% 1/18.2 Delta/Star
a Beam current + accelerated electrons at the output of the grounded grid + grounded grid current.
958
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
to obtain the same output voltage. In both PWM and square-wave modulation the instantaneous output voltage is given by Eq. (2). Therefore, in case of loss of inverter control and in no load conditions, due to the filter capacitances which are charged on the peak of the applied voltage, a higher output voltage would be generated by the PWM. Moreover, voltage steps are applied to the windings at a much lower frequency with respect to PWM. These features result in benefits for the design of the insulation. 3.3. Regulation of the dc-link voltage
Fig. 2. Output voltage waveforms of one phase of an NPC inverter in square-wave modulation. Upper plot: phase voltage, lower plot: line voltage.
therefore requiring only three bushings instead of six and reducing the phase current on the windings. 3.2. Modulation technique Square-wave modulation at 150 Hz has been selected to regulate the output voltage of the NPC inverters; an example of the output voltage waveforms produced by a NPC inverter in square-wave modulation is shown in Fig. 2. This technique has been selected instead of Pulsed Width Modulation (PWM), because of its advantages. On top of limiting the switching frequency of the inverter active components, the square wave modulation maximizes the AGPS output voltage for a given dc-link voltage and step-up transformer turn ratio, while the higher voltage harmonic content at the inverter output plays a minor role in this case being the voltage finally rectified at the AGPS output. Considering a single AGPS stage (see Fig. 3), the no-load output voltage Ed0 (average value) at full modulation and neglecting the dc filter effect, for the sinusoidal (PWM) case and the square waveform case respectively, is given by Eqs. (1) and (2). √ 3 V 1.65 Edo = 1.35 · V2rms = 1.35 · · √dc = (1) · Vdc n n 2 2 (2) · Vdc n √ √ where 3/n is the voltage ratio of the step-up transformer, Vdc / 2 the maximum, theoretical rms value of the sinusoidal line voltage generated by a PWM modulation. Comparing (1) and (2), it can be noted that for the same dc-link voltage (Vdc ) and turn ratio (n), the output voltage (Edo ) is higher in the square wave modulation case. In other words, a lower dc link voltage (about 21%) is necessary Edo =
Fig. 3. Simplified scheme of an AGPS stage.
The drawback of the square-wave modulation is that the per unit output voltage ripple is strongly dependant on the inverter modulation index, and the compliance with the ripple specification is not met for the full range of variation of the AGPS output voltage. As a consequence, it is necessary to reduce the dc-link voltage according to a regulation scheme aimed at minimizing the ripple when the AGPS output voltage decreases. An input stage based on thyristor controlled rectifiers was selected to achieve the regulation of the dc-link voltage. To optimize the design, a common dc-link for the five stages has been chosen. The scheme has been described and investigated in details in [8], where it has been shown to be effective in maintaining the ripple within the specification. The introduction of the regulation makes the dc-link voltage independent on the voltage variations of the supplying network, expected to be within ±10%, since they are compensated by the action of the thyristor rectifiers. An immediate advantage is that the design of the insulation of the AGPS-DCG shall consider only the prescribed maximum dc-link voltage in stationary conditions, i.e. 6.5 kV +6%. 4. Verification of the interface parameters choice The choice of the key interface parameters and of the subsystems topologies was made trying to reach the best compromise between the full compliance with the requirements on one hand and feasibility and procurement simplification of both the AGPS-CS and AGPS-DCG subsystems on the other hand. Once verified that the selected values satisfy the feasibility issues, the fulfillment of the requirements has been verified by means of numerical simulations utilizing the detailed AGPS model [7,9]. Nevertheless, a simple and effective way to check that a set of parameters can satisfy the requirements in all the operative conditions and not exceed the overvoltage limit at the high voltage side of the AGPS has been worked out on the basis of the main equations governing the design of the AGPS. The parameters are the dc-link voltage, the inverter output inductance, the turn ratio and short-circuit impedance of the step-up transformers and the equations are summarized in Appendix and plotted in Fig. 4. This plot shows the output voltage of AGPS as a function of the inverter modulation index for different choices in the parameter space and for different load conditions. The plot is obtained considering a turn ratio n = 1/18.2 and shortcircuit impedance Lsc of 12%. A tolerance of ±1.2% has been assumed on this value, according to indications from industry supporting JAEA (Lsc2 = 0.2 ± 0.02 H on the step-up transformers rated to operate a −800 kV and −1 MV, Lsc2 = 0.23 ± 0.023 H on the three other step-up transformers). Three groups of graphs are represented, one for the dc-link voltage nominal value and the other two for the prescribed nominal variation of the dc-link voltage (±6%). Within each group, the graphs are obtained for the nominal short-circuit impedance and for the upper and lower limits of the tolerance. The load conditions for each group have been selected considering the worst case for the requirement under investigation. Intermediate and lower groups are obtained in Deuterium operation at full current (66 A), to test the fulfillment of the voltage requirements in nominal operation and when the dc-link is at the lower boundary.
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
Appendix A.
125 0
10 A Output Voltage Ed x 5 [kV] .
120 0
Vd-6%, LscMax Vd-6%, Lsc Vd-6%, LscMin Vd, LscMax Vd, Lsc Vd, LscMin Vd+6%, LscMax Vd+6%, Lsc Vd+6%, LscMin
115 0
110 0
66 A
105 0
66 A
100 0
950 0,88
959
0,9
0,92
0,94
0,96
0,98
1
Modulation Index [ ]
Fig. 4. Total output voltage as a function of the inverter modulation index. Id = 66 A or 10 A, n = 1/18.2, Vsc = 12%.
The upper plot has been obtained considering a load current of 10 A, corresponding to an operation near no-load condition, in order to verify the output voltage in case of loss of inverter control. The area highlighted in light brown in the plot corresponds to the specified operating region. With the given choice of parameters, the AGPS is able to meet the requirements. However, there is a risk of slightly exceeding 1.2 MV at the output in case of loss of inverter control at low load currents, as the voltage can be as high as 1.23 MV. But this excess of about 2% is considered acceptable by the AGPS-DCG designers.
Considering Fig. 3, with inverter in square-wave modulation, the average output voltage Ed of the AGPS is given by: Ed = E(ı)d0 − 6 · f · LTOT · Id
(1)
where E(ı)d0 is the no-load output voltage, f the frequency of the modulation, LTOT the total inductance seen by the inverter and reported at the secondary side of the step-up transformer, Id the current supplied at the output. The load is assumed as a current generator and the effect of output filters is neglected. Eq. (1) is commonly used in sinusoidal regime for accounting commutation voltage drops, but it can be shown that it is valid for square waveforms also. It can be shown that: E(ı)d0 =
2 · Vdc · k ı n
(2)
where Vdc is the dc-link voltage, k(ı) a function of the inverter modulation index ı and n the turn ratio of the step-up transformer. k(ı) is equal to ı for ı near 1 [7]. LTOT is given by the following expression: √ 2 3 LTOT = · LINV + LSC2 (3) n where LINV is the inverter total output inductance and LSC2 is the step-up transformer short-circuit inductance seen at the secondary side. Combining (1), (2) and (3) it is possible to plot the output voltage as a function of the interface parameters in various conditions, accounting for tolerance. References
5. Conclusions The progress in the design of the ITER NBI AGPS has been presented. The AGPS reference scheme is the result of a careful optimization and selection process from existing schemes, which considered the demanding requirements of the application. The NPC inverter scheme, combined with the square-wave modulation technique, allows achieving the required performance and optimizing the design of both AGPS-CS and AGPS-DCG. The assessment of the interface parameters, which has been deeply discussed and analyzed, has proved to be satisfactory in all operative conditions, while avoiding at the same time the possibility of exceeding the over-voltage limit at the output in case of loss of inverter control. Acknowledgement and disclaimer This work was set up in collaboration and financial support of Fusion for Energy. The views and opinions expressed herein do not necessarily reflect those of the ITER Organization.
[1] V. Toigo, et al., Design status of power supply systems for SPIDER and MITICA NBI experiments, in: 27th Symposium on Fusion Technology, Liège, September 24–28, 2012. [2] ITER technical basis, IAEA, Vienna 2002, ITER EDA, Doc. Series N.24, Plant Description Document, Sec. 2.5.1. [3] R.S. Hemsworth, A. Tanga, V. Antoni, Status of the ITER neutral beam injection system, Review of Scientific Instruments 79 (Feburary) (2008). [4] P.L. Mondino, The ITER pulsed power supply system, in: Proceeding on 17th IEEE/NPSS Symposium Fusion Engineering, 1998, pp. 491–496. [5] K. Watanabe, et al., Development of a dc 1 MV power supply technology for NB injectors, Nuclear Fusion 46 (2006) S332–S339. [6] K. Watanabe, et al., Design of a −1 MV dc UHV power supply for ITER NBI, Nuclear Fusion 49 (2009) 055022. [7] L. Zanotto, A. Ferro, V. Toigo, Assessment of performance of the Acceleration Grid Power Supply of the ITER neutral beam injector, Fusion Engineering and Design 84 (2009) 2037–2041. [8] E. Gaio, V. Toigo, et al., The alternative design concept for the ion source power supply of the ITER neutral beam injector, Fusion Engineering and Design 83 (2008) 21–29. [9] A. Ferro, L. Zanotto, V. Toigo, Control strategy for the acceleration voltage of the ITER neutral beam injector, IEEE Transactions on Plasma Science 40 (3–1) (2012) 564–569. [10] M. Kuriyama, et al., High energy negative-ion based neutral beam injection system for JT-60 U, Fusion Engineering and Design 26 (1995) 445–453.
Contents lists available at ScienceDirect
Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Progress of the ITER NBI acceleration grid power supply reference design Vanni Toigo a , Loris Zanotto a,∗ , Marco Bigi a , Hans Decamps b , Alberto Ferro a , Elena Gaio a , Daniel Gutiérrez c , Kazuki Tsuchida d , Kazuhiro Watanabe d a
Consorzio RFX, Associazione EURATOM-ENEA sulla Fusione, Corso Stati Uniti 4, I-35127 Padova, Italy ITER Organization, Route de Vinon sur Verdon, 13115 Saint Paul Lez Durance, France Fusion For Energy, C/Josep Pla 2, 08019 Barcelona, Spain d Japan Atomic Energy Agency, 801-1 Mukoyama, Naka, Ibaraki-ken 311-0193, Japan b c
h i g h l i g h t s This paper reports the progress in the reference design of the Acceleration Grid Power Supply (AGPS) of the ITER Neutral Beam Injector (NBI) A critical revision of the main design choices is presented in light of the definition of some key interface parameters between the two AGPS subsystems. The verification of the fulfillment of the requirements in any operational conditions is reported and discussed.
a r t i c l e
i n f o
Article history: Available online 12 February 2013 Keywords: ITER NBI Power Supply
a b s t r a c t This paper reports the progress in the reference design of the Acceleration Grid Power Supply (AGPS) of the ITER Neutral Beam Injector (NBI). The design of the AGPS is very challenging, as it shall be rated to provide about 55 MW at 1 MV dc in quasi steady-state conditions; moreover, the procurement of the system is shared between the European Domestic Agency (F4E) and the Japanese Domestic Agency (JADA), resulting in additional design complication due to the need of a common definition of the interface parameters. A critical revision of the main design choices is presented also in light of the definition of some key interface parameters between the two AGPS subsystems. Moreover, the verification of the fulfillment of the requirements in any operational conditions taking into account the tolerance of the different parameters is also reported and discussed. © 2013 Elsevier B.V. All rights reserved.
1. Introduction The Acceleration Grid Power Supply (AGPS) is the most powerful among the power supply systems [1] of the ITER Neutral Beam Injector (NBI) [2,3]. It has to feed the multistage grid system to accelerate negative ions up to −1 MeV (Deuterium) or −870 keV (Hydrogen), with a beam current of 40 A or 46 A respectively. The challenging requirements, both in terms of ratings and of specific constraints related to the particular application called for in-depth studies to outline the reference design and to advance the iterative process of cross-check verifications toward its assessment. The development of the AGPS reference scheme was carried out starting from the analysis of the existing NBIs rated for the highest voltages. The NBI AGPS scheme closest to the ITER one in terms of requirements was the JT-60 U one, based on negative ions and designed for a voltage up to 500 kV dc, half of the total acceleration voltage required for ITER [4–6].
∗ Corresponding author. Tel.: +39 049 8295898. E-mail address: [email protected] (L. Zanotto). 0920-3796/$ – see front matter © 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.fusengdes.2013.01.013
With respect to the JT-60 U scheme, some modifications and optimizations have been introduced in the AGPS reference scheme for ITER to cope with the more demanding requirements and to take advantage from the latest technologies available on the market. The ITER AGPS reference scheme, shown in Fig. 1, consists of an ac/dc stage feeding via dc links five Neutral Point Clamped (NPC) three-phase inverters [7], each connected to a step-up transformer feeding a six-pulse three phase bridge rectifier; the rectifiers are connected in series at the output side to obtain the nominal acceleration voltage of −1 MV, with availability of the intermediate potentials [6,8]. The procurement of the AGPS is shared between the European Domestic Agency (F4E), which is providing the low voltage power conversion equipment, namely AGPS Conversion System (AGPS-CS), and the Japanese Domestic Agency (JADA), providing the high voltage part, downstream of the inverter units, called AGPS Direct Current Generator (AGPS-DCG). A strong integration effort in the development of the design between ITER Organization (IO), JADA, F4E and Consorzio RFX, is necessary, as the selection of the scheme and the choice of parameters on one subsystem has a strong technical and cost impact on the other and
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
957
Table 2 AGPS dynamic and precision requirements.
Fig. 1. AGPS reference scheme.
vice versa. To study the operation of the AGPS reference scheme and to verify the fulfillment of the requirements, a first set of parameters was assumed and a detailed numerical model was developed; these studies are described in [7,9]. After recalling the main requirements of the AGPS, this paper presents the progress on the definition of some key interface parameters between the two subsystems and, in this respect, a critical revision of the main design choices, in particular the selection of the inverter and input stage topologies and of the inverter modulation technique. Moreover, the choice of the dc-link voltage, turn ratio and short-circuit impedance of the step-up transformer is justified by analyzing the performance with respect to the variation of operational conditions and parameters tolerance. 2. Main requirements The main requirements of the system have been discussed in detail in [7]. The AGPS shall provide the currents and voltages reported in Table 1, for Deuterium and Hydrogen operation. On top of the beam power, the AGPS must compensate for the losses due to the stripped electrons and to the co-extracted electrons, which generate additional current on the grids. The total rated power delivered by the AGPS shall be about 55 MW for both Deuterium and Hydrogen operation. From the requirements reported in Table 1, it follows that the step-up transformers shall be insulated for up to −1 MV dc between primary and secondary winding. Diode rectifiers and filters must be insulated for the same voltage as well. The ratings of each AGPS stage can be derived from the table increasing the maximum current up to 66 A to take into account a margin of about 5%. The AGPS shall operate within 20%–100% of the full output voltage. The regulation of the output voltage is obtained by proper control of the NPC inverter. Precision and dynamic requirements at
Parameter
Value
Voltage regulation range Voltage resolution Output voltage accuracy for 1 h operation Maximum voltage fluctuation Maximum Voltage Ripple Rise time of the output voltage Max. settling time
20%–100% 1 kV ±2% ±2.5% at flat-top ±5% 80 ms 50 ms
beam turn-on and after a breakdown of the grids are summarized in Table 2. Note that the design shall guarantee an output voltage rise time of 80 ms, a settling time of 50 ms and a voltage ripple within ±5% of the output voltage. Moreover, in case of breakdown, the AGPS must switch off within 150 s from the breakdown (50 s for the detection plus 100 s for turning-off) to protect the grids and, after a suitable time (20 ms), the voltage must be reapplied with an appropriate ramp up to the set-point, with the dynamic and precision already described. An additional important requirement for the design of the insulation of the high voltage components is that in case of loss of inverter control, i.e. if the inverters fail to regulate the output voltage, the AGPS output voltage must never exceed the nominal value of the insulation voltage to ground (1.2 MV) in every possible condition. 2.1. Interface parameters A step forward in the assessment of the interface between AGPSCS and AGPS-DCG has been done in 2011 and, in particular, a list of key parameters has been defined and agreed; Table 3 summarizes them. The choice of their value is crucial, as explained in the following sections, to address the design of both systems towards the fulfillment of the requirements, and, at the same time, to guarantee their feasibility and optimize their procurement. 3. Critical revision of the reference scheme On the basis of the definition of the interface parameters, a critical revision of the reference scheme has been performed to evaluate the adequacy and convenience of the main design choices. 3.1. Inverter scheme The inverter design is based on a NPC scheme whereas in JT60 U it was based on single phase H-bridges [10]. The selection of the NPC three-phase scheme has several advantages. The NPC scheme, based on the present availability of switching components rated for higher voltage, allows the increase of the dc-link voltage with respect to the H-bridge, with consequent reduction of the inverter rated current and consequent reduction of losses on inverter switches; moreover, the primary winding of the stepup transformers can be delta connected instead of open delta,
Table 1 Voltage and current rating of the grids. Parameter
Main supply Grid 1 Grid 2 Grid 3 Grid 4 Grounded conductora
D− , 1 MeV, 40 A
H− , 870 keV, 46 A
Table 3 AGPS main interface parameters.
Voltage
Current
Voltage
Current
Parameter
Value
−1000 kV −800 kV −600 kV −400 kV −200 kV 0 kV
59.4 A 3.0 A 5.3 A 3.1 A 2.7 A 45.3 A
−870 kV −696 kV −522 kV −348 kV −174 kV 0 kV
62.8 A 2.4 A 4.3 A 3.1 A 2.1 A 50.9 A
dc-Link voltage dc-Link voltage max. variation in stationary regime Inverter topology Inverter output frequency Inverter total output inductance Step-up transformer short-circuit voltage Step-up transformer turn-ratio Step-up transformer winding connections
6500 V ±6% NPC 150 Hz 110H 12% ± 1.2% 1/18.2 Delta/Star
a Beam current + accelerated electrons at the output of the grounded grid + grounded grid current.
958
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
to obtain the same output voltage. In both PWM and square-wave modulation the instantaneous output voltage is given by Eq. (2). Therefore, in case of loss of inverter control and in no load conditions, due to the filter capacitances which are charged on the peak of the applied voltage, a higher output voltage would be generated by the PWM. Moreover, voltage steps are applied to the windings at a much lower frequency with respect to PWM. These features result in benefits for the design of the insulation. 3.3. Regulation of the dc-link voltage
Fig. 2. Output voltage waveforms of one phase of an NPC inverter in square-wave modulation. Upper plot: phase voltage, lower plot: line voltage.
therefore requiring only three bushings instead of six and reducing the phase current on the windings. 3.2. Modulation technique Square-wave modulation at 150 Hz has been selected to regulate the output voltage of the NPC inverters; an example of the output voltage waveforms produced by a NPC inverter in square-wave modulation is shown in Fig. 2. This technique has been selected instead of Pulsed Width Modulation (PWM), because of its advantages. On top of limiting the switching frequency of the inverter active components, the square wave modulation maximizes the AGPS output voltage for a given dc-link voltage and step-up transformer turn ratio, while the higher voltage harmonic content at the inverter output plays a minor role in this case being the voltage finally rectified at the AGPS output. Considering a single AGPS stage (see Fig. 3), the no-load output voltage Ed0 (average value) at full modulation and neglecting the dc filter effect, for the sinusoidal (PWM) case and the square waveform case respectively, is given by Eqs. (1) and (2). √ 3 V 1.65 Edo = 1.35 · V2rms = 1.35 · · √dc = (1) · Vdc n n 2 2 (2) · Vdc n √ √ where 3/n is the voltage ratio of the step-up transformer, Vdc / 2 the maximum, theoretical rms value of the sinusoidal line voltage generated by a PWM modulation. Comparing (1) and (2), it can be noted that for the same dc-link voltage (Vdc ) and turn ratio (n), the output voltage (Edo ) is higher in the square wave modulation case. In other words, a lower dc link voltage (about 21%) is necessary Edo =
Fig. 3. Simplified scheme of an AGPS stage.
The drawback of the square-wave modulation is that the per unit output voltage ripple is strongly dependant on the inverter modulation index, and the compliance with the ripple specification is not met for the full range of variation of the AGPS output voltage. As a consequence, it is necessary to reduce the dc-link voltage according to a regulation scheme aimed at minimizing the ripple when the AGPS output voltage decreases. An input stage based on thyristor controlled rectifiers was selected to achieve the regulation of the dc-link voltage. To optimize the design, a common dc-link for the five stages has been chosen. The scheme has been described and investigated in details in [8], where it has been shown to be effective in maintaining the ripple within the specification. The introduction of the regulation makes the dc-link voltage independent on the voltage variations of the supplying network, expected to be within ±10%, since they are compensated by the action of the thyristor rectifiers. An immediate advantage is that the design of the insulation of the AGPS-DCG shall consider only the prescribed maximum dc-link voltage in stationary conditions, i.e. 6.5 kV +6%. 4. Verification of the interface parameters choice The choice of the key interface parameters and of the subsystems topologies was made trying to reach the best compromise between the full compliance with the requirements on one hand and feasibility and procurement simplification of both the AGPS-CS and AGPS-DCG subsystems on the other hand. Once verified that the selected values satisfy the feasibility issues, the fulfillment of the requirements has been verified by means of numerical simulations utilizing the detailed AGPS model [7,9]. Nevertheless, a simple and effective way to check that a set of parameters can satisfy the requirements in all the operative conditions and not exceed the overvoltage limit at the high voltage side of the AGPS has been worked out on the basis of the main equations governing the design of the AGPS. The parameters are the dc-link voltage, the inverter output inductance, the turn ratio and short-circuit impedance of the step-up transformers and the equations are summarized in Appendix and plotted in Fig. 4. This plot shows the output voltage of AGPS as a function of the inverter modulation index for different choices in the parameter space and for different load conditions. The plot is obtained considering a turn ratio n = 1/18.2 and shortcircuit impedance Lsc of 12%. A tolerance of ±1.2% has been assumed on this value, according to indications from industry supporting JAEA (Lsc2 = 0.2 ± 0.02 H on the step-up transformers rated to operate a −800 kV and −1 MV, Lsc2 = 0.23 ± 0.023 H on the three other step-up transformers). Three groups of graphs are represented, one for the dc-link voltage nominal value and the other two for the prescribed nominal variation of the dc-link voltage (±6%). Within each group, the graphs are obtained for the nominal short-circuit impedance and for the upper and lower limits of the tolerance. The load conditions for each group have been selected considering the worst case for the requirement under investigation. Intermediate and lower groups are obtained in Deuterium operation at full current (66 A), to test the fulfillment of the voltage requirements in nominal operation and when the dc-link is at the lower boundary.
V. Toigo et al. / Fusion Engineering and Design 88 (2013) 956–959
Appendix A.
125 0
10 A Output Voltage Ed x 5 [kV] .
120 0
Vd-6%, LscMax Vd-6%, Lsc Vd-6%, LscMin Vd, LscMax Vd, Lsc Vd, LscMin Vd+6%, LscMax Vd+6%, Lsc Vd+6%, LscMin
115 0
110 0
66 A
105 0
66 A
100 0
950 0,88
959
0,9
0,92
0,94
0,96
0,98
1
Modulation Index [ ]
Fig. 4. Total output voltage as a function of the inverter modulation index. Id = 66 A or 10 A, n = 1/18.2, Vsc = 12%.
The upper plot has been obtained considering a load current of 10 A, corresponding to an operation near no-load condition, in order to verify the output voltage in case of loss of inverter control. The area highlighted in light brown in the plot corresponds to the specified operating region. With the given choice of parameters, the AGPS is able to meet the requirements. However, there is a risk of slightly exceeding 1.2 MV at the output in case of loss of inverter control at low load currents, as the voltage can be as high as 1.23 MV. But this excess of about 2% is considered acceptable by the AGPS-DCG designers.
Considering Fig. 3, with inverter in square-wave modulation, the average output voltage Ed of the AGPS is given by: Ed = E(ı)d0 − 6 · f · LTOT · Id
(1)
where E(ı)d0 is the no-load output voltage, f the frequency of the modulation, LTOT the total inductance seen by the inverter and reported at the secondary side of the step-up transformer, Id the current supplied at the output. The load is assumed as a current generator and the effect of output filters is neglected. Eq. (1) is commonly used in sinusoidal regime for accounting commutation voltage drops, but it can be shown that it is valid for square waveforms also. It can be shown that: E(ı)d0 =
2 · Vdc · k ı n
(2)
where Vdc is the dc-link voltage, k(ı) a function of the inverter modulation index ı and n the turn ratio of the step-up transformer. k(ı) is equal to ı for ı near 1 [7]. LTOT is given by the following expression: √ 2 3 LTOT = · LINV + LSC2 (3) n where LINV is the inverter total output inductance and LSC2 is the step-up transformer short-circuit inductance seen at the secondary side. Combining (1), (2) and (3) it is possible to plot the output voltage as a function of the interface parameters in various conditions, accounting for tolerance. References
5. Conclusions The progress in the design of the ITER NBI AGPS has been presented. The AGPS reference scheme is the result of a careful optimization and selection process from existing schemes, which considered the demanding requirements of the application. The NPC inverter scheme, combined with the square-wave modulation technique, allows achieving the required performance and optimizing the design of both AGPS-CS and AGPS-DCG. The assessment of the interface parameters, which has been deeply discussed and analyzed, has proved to be satisfactory in all operative conditions, while avoiding at the same time the possibility of exceeding the over-voltage limit at the output in case of loss of inverter control. Acknowledgement and disclaimer This work was set up in collaboration and financial support of Fusion for Energy. The views and opinions expressed herein do not necessarily reflect those of the ITER Organization.
[1] V. Toigo, et al., Design status of power supply systems for SPIDER and MITICA NBI experiments, in: 27th Symposium on Fusion Technology, Liège, September 24–28, 2012. [2] ITER technical basis, IAEA, Vienna 2002, ITER EDA, Doc. Series N.24, Plant Description Document, Sec. 2.5.1. [3] R.S. Hemsworth, A. Tanga, V. Antoni, Status of the ITER neutral beam injection system, Review of Scientific Instruments 79 (Feburary) (2008). [4] P.L. Mondino, The ITER pulsed power supply system, in: Proceeding on 17th IEEE/NPSS Symposium Fusion Engineering, 1998, pp. 491–496. [5] K. Watanabe, et al., Development of a dc 1 MV power supply technology for NB injectors, Nuclear Fusion 46 (2006) S332–S339. [6] K. Watanabe, et al., Design of a −1 MV dc UHV power supply for ITER NBI, Nuclear Fusion 49 (2009) 055022. [7] L. Zanotto, A. Ferro, V. Toigo, Assessment of performance of the Acceleration Grid Power Supply of the ITER neutral beam injector, Fusion Engineering and Design 84 (2009) 2037–2041. [8] E. Gaio, V. Toigo, et al., The alternative design concept for the ion source power supply of the ITER neutral beam injector, Fusion Engineering and Design 83 (2008) 21–29. [9] A. Ferro, L. Zanotto, V. Toigo, Control strategy for the acceleration voltage of the ITER neutral beam injector, IEEE Transactions on Plasma Science 40 (3–1) (2012) 564–569. [10] M. Kuriyama, et al., High energy negative-ion based neutral beam injection system for JT-60 U, Fusion Engineering and Design 26 (1995) 445–453.