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Experimental analysis of assessing of the tripping effectiveness of miniature circuit breakers in an electrical installation fed from a synchronous generator set
Electric Power Systems Research 142 (2017) 341–350
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Experimental analysis of assessing of the tripping effectiveness of miniature circuit breakers in an electrical installation fed from a synchronous generator set Krzysztof Ludwinek a , Jerzy Szczepanik b,∗ , Maciej Sułowicz b a b
˛ Faculty of Electrical Engineering, Automatic Control and Computer Science, Kielce University of Technology, Al. Tysiaclecia P. P. 7, 25-314 Kielce, Poland Faculty of Electrical Engineering and Computer Science, Cracow University of Technology, ul. Warszawska 24, 31-155 Krakow, Poland
a r t i c l e
i n f o
Article history: Received 20 January 2016 Received in revised form 13 August 2016 Accepted 23 September 2016 Keywords: Synchronous generator Generating set Short circuit Overcurrent protection Miniature circuit breaker
a b s t r a c t Due to the supply reliability concerns or to supply remote loads in some distribution networks, the synchronous generators are used as auxiliary power sources. In Poland the generators are usually connected to the network of TN (earthed neutral point) structures. Both TN-C and TN-S or their mixed arrangements TN-C-S are used. Overcurrent protection in these circuits, according to Polish and international standards and the recommendation of SEP (Association of Polish Electrical Engineers), should also act as anti-shock protection. The protection is considered to be effective (for 400 V electrical installations) if the overcurrent protection trip takes less than 0.4 s (maximum allowable time of touching voltage occurrence). The investigation performed by authors, showed that the producers of the generators usually use B or C characteristic circuit breakers and for large generator sets (greater than 100 kVA) they recommend load split to the value of one fifth of the generator maximum power. Additionally, the most frequently excitation forcing equipment is mounted to keep short circuit current three times greater than nominal current of the generator. To assess the effectiveness of the anti-shock protection measures in circuits fed from auxiliary generator and for current lower than 32 A, the set up including a self-exiting 5.5 kVA and 400 V generator, overcurrent protection and load was built in laboratory. The set up uses miniature circuit breakers and enables short circuits in different instants, compared to the initial phase angle of the generator voltage. Results: of the performed short circuits showed, that for the self-excited generators (majority of small generator sets with stator current lower than 32 A) the overcurrent protection cannot act as anti-shock protection, when breakers allowing full utilisation of the generator power. The alternative is to limit generator load or to introduce excitation forcing increasing device price. The proposed solution for these generators given by authors is to modify anti-shock protection system in such way, that the tripping occurs always in the time below 0.4 s. Modifications include not only introduction of the measurements of currents, but also the measurements of the generator voltage. The short circuit is indicated not only by current increase, but also by voltage drop during short circuit. Experimental: device utilizing both measurements was build and tested for different initial short-circuit instants, with respect to the synchronous generator phase voltages. In every case, the switch-off time of the short circuit was not longer than 0.4 s, what means that the protection against the risk of electric shock was effective. The settings of the device are fully adjustable, what means that levels of the measured quantities indicating trip as well as the reaction time are adjustable. © 2016 Elsevier B.V. All rights reserved.
1. Introduction
∗ Corresponding author. E-mail addresses: [email protected] (K. Ludwinek), jerzy [email protected] (J. Szczepanik), [email protected] (M. Sułowicz). http://dx.doi.org/10.1016/j.epsr.2016.09.028 0378-7796/© 2016 Elsevier B.V. All rights reserved.
Protection measures against electric shock in low-voltage TNconnected supply systems are described in many publications [1–12]. Installations employing generating sets for backup power must meet the requirements regarding electric shock prevention in the same way as grid-powered installations [2,3,5,7,9–12]. Protection against electric shock via automatic power supply switch-off imposes a permissible time (for switching off) in the case of insula-
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Fig. 1. Time–current curves of miniature circuit breakers a) type B, b) type C, c) type D.
tion fault. In the case of short-circuit in TN systems at the 230/400 V level, for overcurrent protection and working currents ≤32 A and for residual current devices (RCDs), the maximum switch-off time is set to 0.4 s [5]. In TN systems with a separate earthing conductor (PE), protection against electric shock is most often realized through residual current devices (RCDs) and fuses or miniature circuit breakers (MCBs) [8–18]; when a separate PE conductor is not available, the protection is based only on overcurrent devices. TN systems ensure a negligibly small risk of electric shock if the shortcircuit impedance of the PE conductor ZPE or PEN conductor ZPEN (PEN is a conductor that combines the functions of both a PE conductor and an N conductor) is small enough and the touch voltage Ut (as the drop voltage on the impedance of the ZPE or ZPEN conductors) does not exceed the permissible value Ut = 50 V [1,5]: ZPE = ZPEN ≤
Ut Ik
(1)
where Ik is the current flowing via the PE or PEN conductor. In the case of fuse-based overcurrent protection, Joule’s integral (which is the sum of a pre-arc thermal integral calculated from t = 0 to t = tp and the arc thermal integral calculated from t = tp to t = t0 ) is the crucial factor determining the tripping of the fuse under fault conditions [13]:
2
tP
ik2 dt
I t= 0
t0
ik2 dt
+
(2)
tP
where tO is the operating time, tp is the pre-arcing time (the thermal pre-arcing integral usually acting on the narrowing of a fuse link), and (tO − tp = ta ) is the arcing time. From expression (2), the thermal equivalent current Ith is frequently calculated [13,14] as follows:
Ith =
I2t tw
(3)
where tw is the duration of the thermal equivalent current. For a fuse, the Joule integral (2) is given in the catalogues by individual producers [15]. The response time of a fuse is determined by its time–current characteristics [15]. This time is dependent on equivalent short-circuit current Ith (3). The tripping condition of an MCB in the shorted circuit assumes that the RMS value of the short-circuit current does not reach values smaller than that of the assumed equivalent tripping current of the short-circuit release. It is very difficult to assess the tripping effectiveness of an MCB in electrical installations fed from low-power synchronous generator sets, up to several kVA, while considering the breaker’s time–current characteristics owing to relatively low values of short-circuit currents and high contents of transients in shortcircuit currents. In these electrical installations, the only reliable method to assess the tripping effectiveness of MCBs is the exper-
Fig. 2. An example of a power protection variant in TN-S system and synchronous generating set working as a backup power supply.
imental one. The practical choice of the nominal current for the circuit breaker is most commonly derived from the single-phase IN1ph or three-phase IN3ph rated current of the synchronous generator and the character of the load (most commonly R, RL or motor). Depending on the load characteristics, circuit breakers are chosen based on B, C or D time–current characteristics (see Fig. 1). Table 1 presents the tripping times of miniature circuit breakers of B, C or D time–current characteristics in relation to the k-ratio of the short-circuit current Ik = kIN (where IN is the rated current) [16–18]. Tripping times of miniature circuit breakers in relation to B, C or D characteristics are current dependent, thus to achieve fast enough switching off times the value of short circuit current has to have large values for a certain time. Minimal values of the currents required to achieve a breaker with a known characteristic to trip in time below 0.1 s are stated in the IEC standards [17,18]. For a tripping time 0.4 s, the thermal equivalent current can be estimated from characteristics (Fig. 1) or determined during the experiment. In a low-voltage electrical installation, the stability of the power source (of the electrical power system) is, by definition, the essence
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Table 1 Main technical details of elements and devices in control system of the 5.5 kVA synchronous generator supporting protection against electric shock. No. in Fig. 6
Element or device in the control system Fig. 6
Main technical details
1 2 3 4 5
Salient pole synchronous generator Single phase bridge rectifier Push-pull converter Signal conditioning board Programmable Logic Controller
6
Voltage transducers
SN = 5.5 kVA, UN = 400 V, In 3ph = 7.9 A, In 1ph = 13.8 A, nN = 3000 rpm KBPC610: 6 A, 1000 V ±15 V (1 A), 24 V (2 A) ULN 2804: 24 V, 8 × 500 mA CP1H −XA (Omron), digital (24 V): 24 inputs and 16 outputs, analog: 4 inputs and 2 outputs LV 25-NP (LEM): UN = 250 V - nominal measured voltage, IPN = 10 mA - primary nominal current RMS, tr = 40 s - response time to 90% IPN ±0.9% accuracy of UNph (TA = 25 ◦ C), linearity error <0.2% LA 100-P (LEM) IPN = 100 A - primary nominal current RMS, tr = <1 s - response time to 90% IPN ±0.45% accuracy of IPN (TA = 25 ◦ C), linearity error <0.15% di/dt accurately followed >200 A/s MCB time–current characteristics: B6, B8, B10, B13 C6, C8, C10, C13 SSR-480D125 Solid State Relays 480 V, 125 A, 0.02 ms turn “on” time
Current transducers
7
Overcurrent protection
8
Relay output board
Solid state relays
of measuring fault loop impedance (owing to the fact that electrical power system is usually modelled as a Thévenin’s equivalent circuit) [9]. The measurements of fault loop impedance are based on measurements of voltage drop on the internal impedance of the supply system. The measurements are usually taken by applying a large load (controlled short circuit) between each phase conductor and protective earth (PE) conductor or protective earth/neutral (PEN) conductor. However, as shown in this paper, in the case of electrical installations, after the switch from grid power to backup power utilizing synchronous generators, this backup voltage source may no longer be modelled as a Thévenin’s equivalent circuit, which utilizes a stationary voltage source and constant impedance. Thus, the process of the selection of the protection devices in such installations may not be undertaken in the same way as those supplied from electrical power systems. Under short-circuit conditions in electrical installations powered from a synchronous generating set, the shape of the instantaneous current is significantly influenced by transient phenomena occurring in the generator and AVR system powering the excitation circuit of the synchronous generator (working in conjunction with a closed feedback loop) [3,19–27]. In many reference guides and practical recommendations, the use of excitation controllers and systems for adjusting the voltage and power of synchronous generators cooperating with the power grid are mentioned [19–25]. The proposed control systems are used for: regulation and stabilization of the voltage and frequency of a generator working in a power system, limitation of interferences and failures, specification of the value of excitation current, specification of the power angle, the provision of limitations with long-term current overload protection, etc. In systems used at present to control synchronous generators working in combustion-engine-driven generation sets (as a backup power supply), there are no regulators controlling voltage or power in the case of a single-phase short-circuit to N or PE conductor, which would enable power switch-off in a time shorter than that required by law (0.4 s). This article presents the case of a single-phase short-circuit and the failure of the application of the overcurrent protection as an anti-shock measure. There are many manufacturers of generator sets on the market right now and some of them offer several types of generators. Every generator set is characterised by individual parameters (electrical and mechanical) and its own system of excitation control and forcing.
Usually, the manufacturers of the generator sets are not sharing any data regarding work and the characteristics of the applied AVR unit. The simulations performed using Matlab Simulink software on a system consisting of 22 kVA generator sets, using data obtained from the manufacturer and standard Simulink models (only data regarding the generator, but not the AVR system and speed control were available) showed, that there is a large difference between real short circuit current (obtained during laboratory experiments) and waveforms obtained from simulations. The characteristics of MCB (from different manufacturers) also show certain inaccuracies and should be modelled as band characteristics, what introduces some uncertainty into the determination of the tripping times of the breakers. These facts clearly show that adequate data regarding MCB switching off times during work with a small generator set can be obtained only from the experimental investigations. 2. Development of the control system supporting protection against electric shock In low and high power combustion-engine-driven generation sets, self-exciting synchronous generators with electromagnetic excitation are used more often than those with AVR systems, which are used for higher-power applications [3,19,22]. The generators with self-exciting arrangements can power a single load or groups of loads and can work in an IT, TT, TN-S or TN-C-S system [19]. To support the operation of a synchronous generator working as a backup power system and control stator currents and voltages (of a self-exciting synchronous generator), an AVR or electromagnetic excitation control system is usually used [3,22–27]. Generally, the tasks of AVR or electromagnetic control systems include the following: • maintaining short-circuit current at the level of approx. 2.5IN –6IN in the case single-phase short circuits (as in a conventional automatic voltage regulation or electromagnetic control systems used in synchronous generator sets)—according to [3], a short-circuit current should be only 3IN ; • maintaining an effective value of phase stator voltages, most often at the level of UNf in a no-load condition or a voltage with permissible deviation ıUst [3] in a load condition; • switching off the circuit in the case of overcurrent and risk of insulation damage [3].
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Fig. 3. Spectral reactance in d-axis vs. frequency for a 5.5 kVA synchronous generator.
In a short circuit, the correct operation of the anti-shock protection depends on a residual-current-operated circuit breaker, fuses or miniature circuit breakers. The speed of activation of residual circuit breakers is not questionable. However, owing to their frequent failures, they cannot (in terms of anti-shock protection) be the only devices to determine power supply switch-off, especially, for example, in firefighting vehicles [11,12]. That is why the standards [11,12,17,18], apart from residual circuit breakers, also necessitate the usage of overcurrent protection (fuses or any overcurrent circuit breakers) as an anti-shock protection supplement. In the case of a synchronous generating set (working in an autonomous system) the short-circuit condition results in the reaction of current, voltage and frequency controllers [19,21–23]. This has a significant impact on the waveform of the short-circuit current, which depends on changes in synchronous generator reactance (during the short circuit), the initial voltage phase, the distance of the short-circuit from the generator terminals and, among other things, the type of control system used for excitation voltage and current [19]. As shown in [9], given the examined 5.5 kVA synchronous generating sets—either in the case of external circuit insulation damage with additional resistance (modelling line conductor and PE or PEN conductor) or in the case of the “complete” insulation damage of a load (metallic short circuit at the terminal) − the reaction time of fuses is not always shorter or equal to that specified in the literature [5]. As shown in [19–21], the changes of examined synchronous generator reactances during the short-circuit process have a significant impact on the waveforms of short-circuit currents. Thus, when the transient phenomena become relatively long, the changes of reactance have a substantial influence on the value of the Joule integral (2) and thermal equivalent current (3). An example of a characteristic of spectral reactance in d-axis vs. frequency for a 5.5 kVA (400 V) salient pole synchronous generator is shown in Fig. 3. At the first moment during the short circuit, the value of reactance Xd is determined with f → ∞. The reactance changes shown in Fig. 3 are determined by a standstill frequency response (SSFR) test [28,29]. Usually, in electrical installations, the values of the short circuit currents, where circuit is supplied from mains, are calculated from the measurements of a fault loop impedance using Thévenin’s equivalent circuit. Such measurements are not credible for generator sets, where impedance changes (Fig. 3) accordingly to the generator state (subtransient, transient and steady state). Modern measuring equipment uses relatively small currents to determine circuit impedance, thus only steady state generator impedance is measured. During short circuits, the impedance of the synchronous generator changes, what causes the fault loop impedance change. This fact raises questions regarding the effectiveness of protection against electric shock carried out by a fast automatic power switch-off [9]. Assessing the tripping effectiveness of miniature circuit breakers in relation to permissible time is carried out during experimental investigations using the measurement set. Fig. 4 shows a
Fig. 4. Simplified diagram of the measurement set for investigation of the examined 5.5 kVA salient pole synchronous generator during protection against electric shock.
Fig. 5. Simplified diagram of the electronic phase detector 0–90◦ .
simplified block diagram of the measurement set for investigating the overcurrent protection of the examined 5.5 kVA salient pole synchronous generator during protection against electric shock. To ensure the same phase angle of the phase voltage supply (0◦ or 90◦ ), the short circuit is applied with the use of a special Electronic Phase Detector (Fig. 5) [9]. The short circuit conditions (Fig. 5) can be obtained (at the generator terminals or at the end of the line and N conductors) using an antiparallel connection of two thyristors (E2) and a phase controlled switch drive system (E5). The initial phase angles of the switching “on” of the generator phase voltage using thyristors (E2) is determined by the electronic pulse generator (E3). The microcontroller (E6) is used to control the phase trigger system (E5) and to calculate the first magnitude of short circuit current Im , duration of the thermal equivalent current tw and Joule’s integral I2 t according to expression (3) and RMS values of the thermal equivalent current Ith on the basis of the instantaneous current measurement using Hall-effect transducer (E1) and the special formed of the analog signal in system (E4). These data are displayed on the console screen (E7). The application of the microprocessor enables changing of the initial phase angles of the switching “on” of the phase voltage from ␣ = 0◦ to ␣ = 90◦ by the user through the console switches (E7) [9]. During the experimental investigations, the overcurrent protections of 8A type B and C (TCC) (for three-phase rated load current In 3ph = 7.9 A) and 13 A type B (for single-phase rated load current In 1ph = 13.8 A) were used. The presented measurement set up enables the influence of additional phase and PEN conductor resistance (Rd = RL + RPEN ) on the value of touch voltage UT to be determined for the equivalent thermal current Ith (Rd ). The examined 5.5 kVA synchronous generator is driven by a DC motor. The rotational speed of 3000 rpm is set by the frequency measurement
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Fig. 6. Control system of a synchronous generator supporting protection against electric shock.
(50 Hz) of the output phase stator voltage in the no-load state. The field current is adjusted through the electronic excitation controller system (manufacturer built-in AVR). The laboratory test bed with the proposed new control system (see Fig. 6) was built from the following blocks: three-phase synchronous generator set (1), power supply (2, 3), PLC-based controller (5), measuring and signal conditioning block (6), overcurrent protection (7), electromagnetic compatibility and relay output board (8). The output board consists of three capacitors 3C (facilitating the process of self-exciting and ensuring suitable electromagnetic compatibility), relays K1–K4, terminals L1, L2, L3 and N. The system of measurement and signal conditioning board (6) are connected to the generator terminals and consist of four transducers for measuring the phase-line voltages PU1, PU2, PU3 and excitation voltage PU4 and four transducers for measuring stator currents PI1, PI2, PI3 and field current PI4 [30–32]. The signal conditioning is performed using 2nd-order low-pass Bessel filters. The measured signals from block (6) are transferred to the PLC controller (5) inputs. The PLC is programmed to control output contactors S1–S4 (output board 8) via the signal conditioning board (4). The outputs of board 4 are two state outputs suitable to control output relays. Table 1 shows the most important technical details of elements and devices included in the control system of the 5.5 kVA salient pole synchronous generator supporting protection against electric shock (Fig. 6). The measured signals were acquired using digital four-channel oscilloscope type MSO 3014 (Tektronix) offering 100 MHz bandwidth and 1 mega-point record length on all channels. The new control system developed by the authors (Fig. 6) enables full power receipt switch-off in short-circuit conditions (if the RMS currents Ia , Ib or Ic measured with transducers PI1–PI3 are greater than set up level or/and voltage at any supply phase is lowered). The conditions, during which voltage is asymmetrical and lowered occur during asymmetrical short circuits (single phase to ground, two phases to ground). At relatively high-resistance faults (for example, during functional insulation damage) the asymme-
try of the RMS voltage measured with transducers PU1–PU3 can be used as the supply circuit trip signal. The examples shown in the results section were using the 20% asymmetry level (asymmetry of the generator’s phase voltages measured with the voltage transducers PU1, PU2 and PU3 in relation to the PE conductor) as an indicator to supply circuit switch off. Of course, in the proposed device, the asymmetry level is programmable and can be set between 10% to 80%. However, to assure the proper work of the system within the standards (10% voltage drops as well as asymmetry are allowed), during the experiments maximum sensitivity level of 20% was established. Thus, the short-circuit condition is the condition in which at least one of the PI1, PI2 or PI3 phase current transducers measures a current value greater than the set value (set within the range of 2IN –6IN ), with a simultaneous decrease of voltage in short-circuited phases and an increase in other phase voltages (asymmetry level increased above set value). The voltages in non-short circuited stator windings are increased, what is caused by the field current forcing by AVR system. The delay of the switch-off, after fault detection, can be adjusted by user from 0.2 s to 5 s using interface E7; thus, the time of the presence of dangerous voltage may be limited, depending on the effective value of the phase-line voltage and power supply conditions [5]. In the case activation failure of the standard overcurrent protection of a residual circuit-breaker or the continuation of a short circuit after the permissible time (0.4 s for 230/400 V), the short-circuited wiring is switched off.
3. Assessing of the tripping effectiveness of miniature circuit breakers in relation to permissible trip time without the new auxiliary control system To assess the effectiveness of the anti-shock protection, several experiments were conducted, without and with the proposed support circuitry. At the beginning of the experiments evaluating tripping times of the overcurrent protection devices (MCBs) with tripping characteristics B and C and with tripping current values 8 A and 13 A were performed, The values of the breakers currents
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Fig. 8. Registered waveforms for 8 A MCB with C (TCC) a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0 , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90 , ub − 90◦ and uc − 90◦ (for ia − 90◦ ). ◦
Fig. 7. Registered waveforms for 8 A MCB with tripping characteristic curve type B a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0 , ub − 0◦ and uc − 0 (for ia − 0◦ ); c) induced stator phase voltages ua − 90 , ub − 90◦ and uc-90◦ (for ia − 90◦ ).
◦
◦
◦
◦
were near the nominal values of current for the researched generator (In3ph = 7.9 A and In1ph = 13.8 A). In Figs. 7–10 the fault inception point (FTP) for the voltage ua phase angle 0◦ and 90◦ (FTP-0deg and FTP-90deg) and MCB opening point (MCB-OP0◦ and MCB-OP 90◦ ) are presented. Figs. 7 and 8 present the registered single-phase short circuit ia − 0◦ (for the phase angle 0◦ of voltage ua ), ia − 90◦ (for the phase angle 90◦ of voltage ua ), induced stator phase voltages ub and uc under no-load conditions and during the single-phase short circuit of the investigated salient pole synchronous generator with overcurrent protection of 8 A with TCC types B (Fig. 8) and C (Fig. 9). As shown in Figs. 7 and 8 for the examined 5.5 kVA synchronous generator, in the case of a short circuit in any initial phase of voltage and overcurrent protection of 8 A, for tripping characteristic curve B, the protection is activated in less than one period of 0.02 s; for tripping characteristic curve C (and in the same way for type D—Figs. 1 and 2) the interrupt reaction time will not cause the activation of protection in the required time (less than 0.4 s). Figs. 9 and 10 present the registered single-phase short-circuit ia − 0◦ and ia − 90◦ and induced stator phase voltages ua , ub and uc under the no-load condition and during the single-phase short circuit of the investigated salient pole synchronous generator with
the rotor skew for overcurrent protection of 13 A with tripping characteristic curve B (see Fig. 9) and 6 A with C (see Fig. 10). As shown in Figs. 9 and 10, for the examined salient pole synchronous generator with Sn = 5.5 kVA, Un = 230/400 V in the case of a short circuit in any initial phase of voltage, the overcurrent protection • of 13 A for type B and 0◦ initial phase of voltage causes the activation of the protection in an operating time of less than 0.02 s and for a 90◦ initial phase of voltage causes the activation of the protection in an operating time of less than 0.16 s. • of 6 A for type C causes the activation of the protection is less than 0.02 s. From the analysis of Figs. 7–10 (short circuits without the control system shown in Fig. 6 and without the additional external resistance Rd = RL + RPEN − Fig. 4), it can be concluded that the greatest value of overcurrent protection type C should not exceed 6 A, if the required tripping time is less than 0.4 s. The above result raises questions about the utilisation of the full power of the proposed generator because the maximum tripping current of the C type characteristic fuse ensuring proper anti-shock protection is 6 A. The conclusion from this investigation is that when using the MCBs, either generator load and/or load type must be limited, or safety rules must be revised.
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Table 2 Comparison of characteristic values for MCB during short circuits in system powered from generator. Miniature Circuit Breaker TCC
B C
Phase (deg)
0 90 0 90
6A
8A 2
2
10 A 2
2
13 A 2
2
Im (A)
I t (A s)
Ith (A)
tw (ms)
Im (A)
I t (A s)
Ith (A)
tw (ms)
Im (A)
95 89 95 83
33 34 35 33
57.1 53.2 56.4 44.6
10 12 11 17
95 86 97 93
35 33 508 503
56.4 44.6 31.8 31.1
11 17 503 520
99 136 38.0 94 98 89 130 37.0 95 89 operating time is always greater than 0.4 s operating time is always greater than 0.4 s
Fig. 9. Registered waveforms for 6 A MCB with TCC a) single-phase short-circuit ia − 0◦ and ia − 0◦ ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc -90◦ (for ia − 90◦ ).
A detailed comparison of maximum values of current Im , Joule’s integral I2 t, thermal equivalent current Ith and duration of thermal equivalent current tw obtained from experimental investigations for single-phase short circuits and miniature circuit breakers is presented in Table 2. The range of fuses includes ones with time–current characteristics B (6 A, 8 A, 10 A and 13 A) and C (6 A and 8 A). The values presented in Table 2, obtained from the experiment, show, that for the investigated 5.5 kVA salient pole synchronous generator, the requirements of electric shock protection (tripping time less than 0.4 s) are satisfied only for the following MCB’s:
I t (A s)
Ith (A)
tw (ms)
Im (A)
I2 t (A2 s)
Ith (A)
tw (ms)
42 206
59.0 41.8
12 118
Fig. 10. Registered waveforms for 13 A with TCC B a) single-phase short-circuit ia − 0 and ia − 90 ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0 ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc − 90◦ (for ia − 90 ). ◦
◦
◦
◦
• B in the 6–13 A range of rated current and in tripping times from 10 ms to 118 ms, • C only in the 6 A range of rated current in tripping time 11–17 ms The miniature circuit breakers with time-current characteristic C (8 A, 10 A and 13 A) do not guarantee compliance with the requirements of electric shock protection through automatic disconnection of the supply under short-circuit conditions because their operating time is greater than the required 0.4 s. For these breakers (C8 A, C10 A and C13 A), several measurements of the
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Fig. 11. Registered waveforms with the use of the control system (Fig. 7) for 8 A with TCC C overcurrent protection a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc − 90◦ (for ia − 90◦ ).
short-circuit waveforms showed relatively small values of the RMS thermal equivalent current Ith (3) for the time tw = 0.4 s.
4. Experimental investigations with the use of the new auxiliary control system supporting the operation of a synchronous generator The conclusions from the investigation presented in the previous chapter clearly state that overcurrent protection support devices are necessary to ensure full utilisation of the generator power. This was the reason for the development of the circuitry shown in Fig. 6. The presented auxiliary control system of the examined 5.5 kVA synchronous generator ensures that the time of persistence of dangerous prospective touch voltage during a short circuit directly at the generator terminal, or in the case of electric insulation failure with additional resistance, is no longer than allowed. Fig. 11 presents the registered single-phase short-circuit ia − 0◦ , ia − 90◦ and induced stator phase voltages ua , ub and uc under the no-load condition and during the single-phase short circuit of the investigated salient pole synchronous generator for overcurrent protection of 8 A with tripping characteristic curve C, with the use of the control system supporting the operation of a synchronous generator. In this case, the S1 relay (see Fig. 6) is switched off when the set time equal to 0.4 s elapsed.
Fig. 12. Registered waveforms using the control system for 8 A with C (overcurrent protection trip characteristic curve) and during the high resistance fault (Rd = 3.43 ) of: a) single-phase circuit ia − 0◦ and ia − 90◦ b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc -90◦ (for ia − 90◦ ).
Previous investigations showed that the use of type C overcurrent protection with a rated current greater than 6 A caused the persistence of dangerous prospective touch voltage for more than 0.4 s. From Fig. 11, it is visible that the supporting system switched off the short-circuited branch within the set time (0.4 s), since fault occurs at 0.1 s and switching “off” the circuit takes place at the 0.5 s time mark. Fault generation was synchronized witch oscilloscope and on the following figures occurs always at 0.1 s for “0◦ ” switching angle and is delayed of 0.005 s for “90◦ ” switching. Fig. 12 presents the recorded single-phase short circuit ia − 0◦ , ia − 90◦ and induced stator phase voltages ua , ub and uc under the noload condition, and during the high resistance fault (Rd = 3.43 ) of the investigated salient pole synchronous generator. The tests were carried out for 8 A overcurrent protection with tripping characteristic curve C (the same for 8 A with tripping characteristic curve B), using the control system supporting the operation (see Fig. 2) of the examined synchronous generator. In this case, relay K1 is switched off. The results revealed in Figs. 11 and 12, show that for the examined 5.5 kVA synchronous generator in the case of the utilisation of the proposed protection support circuitry, the switch-off during the fault always occurs within 0.4 s. This happens not only in the case of fault close to the generator when the short-circuit currents are quite high but also when a large impedance fault (insulation defect)
K. Ludwinek et al. / Electric Power Systems Research 142 (2017) 341–350
is simulated by applying additional resistance. The necessary conditions for the proposed circuitry to switch off during a single-phase fault include two basic requirements: the current in one of the phase conductors must increase above a certain adjustable level, and asymmetry of the generator voltages must exceed 20%.
5. Conclusions On the basis of experimental investigations without the control system presented in this paper, it can be concluded that the overcurrent protections with tripping characteristics B and C selected for rated currents of the examined synchronous generator (one phase rated load and three phase rated load), both for single-phase and three-phase loads, can be ineffective as anti-shock protection. Performed tests aimed at determining the effectiveness of a fast automatic power switching-off of the electrical installation powered by a synchronous generator showed that the application of smaller-than-rated or fast switching breakers is required to satisfy the stated anti-shock protection criteria, which can limit the power or type (electric machines) of the loads. The problem is associated with the limitation of the short-circuit current in circuits where only overcurrent protection is used, what can cause long persistence of an effective touch fault voltage greater than 50 V between a given fault point and reference earth, which is not allowed by given standards. The other option to switch off the short-circuited wiring during the required time may be to individually select output section wiring with the use of a separate overcurrent protection of trip characteristic curves B or C selected on each section. This increases the cost of electrical system, since in order to decrease system impedance the diameters of wires have to be increased. The overcurrent protections of trip characteristic curves B or C in the sections are selected by measuring the Joule integral of the short circuit current. This solution in the output sections will indeed enable power switch-off in the required permissible time, but still does not eliminate the causes of output power limitation (in the sections). The research shown in this article concentrated on a 5.5 kVA low-power generator used for laboratory set up, however, the problem of the anti-shock protection is also present in large power generator units. This fact was noticed by the authors during professional expertises performed on such objects as hospitals or airports. For example, for 200 kVA unit (airport radar system), the manufacturer of the generator uses excitation forcing to assure the proper value of short circuit current (above 3IN ) and suggests that the current of the single output should not exceed 20% of generator nominal current. Such arrangement allows the limitation of the value of circuit breaker current to 0.2IN , thus creates the condition when short-circuit current is 15 times greater than nominal current of the breaker, however, the load of single circuit powered by the generator is limited to 40 kVA. This approach works well for large generator sets (400 kVA and larger units) mounted, for example in hospitals, where no single loads above 20% of generator power are expected. The auxiliary protection system presented in this paper is one of the means to solve that problem. The proposed circuitry uses both voltage and current measurements to detect single-phase or two phase to ground short circuits (short circuits of phases to neutral or earthing conductor) what is a novel approach at low voltage (below 1 kV) level. This arrangement allows also switch-off of the load when the short-circuit current is limited by additional impedance introduced by the nonzero impedance fault (insulation burn down). Moreover, if the device is mounted in each section of output circuitry there is possible (by applying proper settings) to maintain the selectivity of the section protection and overcurrent protection of the generator.
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The drawback of the presented control system is that it requires quite costly elements as voltage and current measuring sensors and output switches (for each output section) that are able to turn off short-circuit currents. However, it is possible to build an integrated measurement and control system with a single voltage sensor for a single switchgear and separate current sensors for each output, what will lower overall costs of the device. The cost of electronic control board is almost negligible when compared to costs of the switches at the output. As mentioned above, these switches have to be able to turn off short circuit currents for switchgears supplied from the generator (auxiliary supply) or from the mains during normal supply conditions. References [1] M. Barrett, K. O’Connell, A.C.M. Sung, Analysis of transfer touch voltages in low-voltage electrical installations, Build. Serv. Eng. Res. Technol. 31 (1) (2010) 27–38. [2] F.R. Blánquez, C.A. Platero, E. Rebollo, F. Blázquez, On-line stator ground-fault location method for synchronous generators based on 100% stator low-frequency injection protection, Electr. Power Syst. Res. 125 (2015) 34–44. [3] LV Generator Protection, Low Voltage Expert Guide No. 8, Schneider Electric, 2001. [4] HD 60364-5-551-2010, Electrical installations of buildings - Part 5–55: Selection and erection of electrical equipment - Other equipment - Clause 551: Low-voltage generating sets, 2010. [5] HD 60364-4-41-2005, Low-voltage electrical installations −part 4–41: Protection for safety: Protection against electric shock, 2005. [6] HD 60364-5-54-2007, Electrical installations of buildings −part 5–54: selection and erection of electrical equipment, Earthing arrangements, Protective conductors and protective bonding conductors, 2007. [7] HD 60364-5-551-2010, Electrical installations of buildings - Part 5–55: Selection and erection of electrical equipment - Other equipment - Clause 551: Low-voltage generating sets, 2010. [8] S. Czapp, Protection against electric shock using residual current devices in circuits with electronic equipment, Elektronika ir Elektrotechnika 21 (4) (2007) 51–54. [9] P. Makarski, System for measuring tripping time in safety fuses and circuit breakers, Polish Patent No. PL392550-A1, PL218499-B1. [10] ISO 8528-5:2013, Reciprocating internal combustion engine driven alternating current generating sets - Part 5: Generating sets, 2013. [11] DIN 14687-2007-02, Firefighting Equipment - Permanently installed generators (generating sets) less than 12 kVA for the use in firefighting vehicles, 2007 (In German). [12] DIN 14686/A1:2015-03. Firefighting equipment - Switch cabinets for fixed generators (generating sets) > = 12 kVA in fire-brigade vehicles, Amendment A1, 2015 (In German). [13] J. Nahman, D. Stojanovi, Calculation of thermal equivalent short − time current, Facta Univ. Ser. Electron. Energetics 1 (1) (1998) 103–114. [14] M. Gianpasquale, Thermal equivalent short-time currents in three-phase balanced short circuits, IEEE Trans. Power Deliv. 9 (2) (2004) 897–898. [15] Special purpose fuses, ETI. Catalog (2015). [16] ASTI: Miniature circuit breakers and Residual current devices, ETI Catalogue (accessed 11.08.16). [17] EN 60898-1:2003/A13:2012 E. Electrical accessories - Circuit breakers for overcurrent protection for household and similar installations - Part 1: Circuit-breakers for a.c. operation, 2012. [18] EN 60898-2:2006. Electrical accessories - Circuit-breakers for overcurrent protection for household and similar installations Part 2: Circuit-breakers for a.c. and d.c. operation, 2006. [19] G. Andersson, Modelling and Analysis of Electric Power Systems. Power Flow Analysis. Fault Analysis. Power Systems Dynamics and Stability, ITET ETH, Zurich, 2004. [20] J. Pyrönen, T. Jokinen, V. Hrabcová, Design of rotating electrical machines, John Wiley, 2012. [21] R. Wamkeue, N.E.E. Elkadri, I. Kamwa, M. Chacha, Unbalanced transient-based finite-element modeling of large generators, Electr. Power Syst. Res. 56 (3) (2000) 205–210. [22] C.S. Hoong, S. Taib, K. Mieee, S. Rao, I. Daut, Development of automatic voltage regulator for synchronous generator, in: National Power & Energy Conference (PECon) 2004, Proceedings, Kuala Lumpur, Malaysia, Nov. 29–30, 2004, pp. 180–184. [23] IEEE Std 421.5-2005. IEEE Recommended practice for excitation system models for power system stability studies, 2005. [24] IEEE Std 421.4-2004. IEEE Guide for the preparation of excitation system specifications, 2004. [25] P421.2/D20, Nov 2013. IEEE Draft guide for identification, testing, and evaluation of the dynamic performance of excitation control systems, 2013. [27] M. Zielichowski, T. Szlezak, Improvements in ground-fault protection of unit-connected generator stator winding using laboratory test environment, Electr. Power Syst. Res. 78 (9) (2008) 1635–1639.
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[28] K. Ludwinek, Staszak J. Possibility of graphical environment applications for ˛ evaluating of equivalent circuit parameters and time constants, Przeglad Elektrotechniczny 12 (2011) 195–200. [29] A.M.A. Oteafy, J.N. Chiasson, S.A. Zaid, Development and application of a standstill parameter identification technique for the synchronous generator, Int. J. Electr. Power Energy Syst. 81 (2016) 222–231.
[30] K. Ludwinek, Measurement of momentary currents by hall linear sensor, ˛ Elektrotechniczny 10 (2009) 182–187. Przeglad [31] Voltage Transducer LV 25-P, Ver. 19, Datasheet 2014. [32] Current Transducer LA 100-P, Ver. 12, Datasheet 2016.
Contents lists available at ScienceDirect
Electric Power Systems Research journal homepage: www.elsevier.com/locate/epsr
Experimental analysis of assessing of the tripping effectiveness of miniature circuit breakers in an electrical installation fed from a synchronous generator set Krzysztof Ludwinek a , Jerzy Szczepanik b,∗ , Maciej Sułowicz b a b
˛ Faculty of Electrical Engineering, Automatic Control and Computer Science, Kielce University of Technology, Al. Tysiaclecia P. P. 7, 25-314 Kielce, Poland Faculty of Electrical Engineering and Computer Science, Cracow University of Technology, ul. Warszawska 24, 31-155 Krakow, Poland
a r t i c l e
i n f o
Article history: Received 20 January 2016 Received in revised form 13 August 2016 Accepted 23 September 2016 Keywords: Synchronous generator Generating set Short circuit Overcurrent protection Miniature circuit breaker
a b s t r a c t Due to the supply reliability concerns or to supply remote loads in some distribution networks, the synchronous generators are used as auxiliary power sources. In Poland the generators are usually connected to the network of TN (earthed neutral point) structures. Both TN-C and TN-S or their mixed arrangements TN-C-S are used. Overcurrent protection in these circuits, according to Polish and international standards and the recommendation of SEP (Association of Polish Electrical Engineers), should also act as anti-shock protection. The protection is considered to be effective (for 400 V electrical installations) if the overcurrent protection trip takes less than 0.4 s (maximum allowable time of touching voltage occurrence). The investigation performed by authors, showed that the producers of the generators usually use B or C characteristic circuit breakers and for large generator sets (greater than 100 kVA) they recommend load split to the value of one fifth of the generator maximum power. Additionally, the most frequently excitation forcing equipment is mounted to keep short circuit current three times greater than nominal current of the generator. To assess the effectiveness of the anti-shock protection measures in circuits fed from auxiliary generator and for current lower than 32 A, the set up including a self-exiting 5.5 kVA and 400 V generator, overcurrent protection and load was built in laboratory. The set up uses miniature circuit breakers and enables short circuits in different instants, compared to the initial phase angle of the generator voltage. Results: of the performed short circuits showed, that for the self-excited generators (majority of small generator sets with stator current lower than 32 A) the overcurrent protection cannot act as anti-shock protection, when breakers allowing full utilisation of the generator power. The alternative is to limit generator load or to introduce excitation forcing increasing device price. The proposed solution for these generators given by authors is to modify anti-shock protection system in such way, that the tripping occurs always in the time below 0.4 s. Modifications include not only introduction of the measurements of currents, but also the measurements of the generator voltage. The short circuit is indicated not only by current increase, but also by voltage drop during short circuit. Experimental: device utilizing both measurements was build and tested for different initial short-circuit instants, with respect to the synchronous generator phase voltages. In every case, the switch-off time of the short circuit was not longer than 0.4 s, what means that the protection against the risk of electric shock was effective. The settings of the device are fully adjustable, what means that levels of the measured quantities indicating trip as well as the reaction time are adjustable. © 2016 Elsevier B.V. All rights reserved.
1. Introduction
∗ Corresponding author. E-mail addresses: [email protected] (K. Ludwinek), jerzy [email protected] (J. Szczepanik), [email protected] (M. Sułowicz). http://dx.doi.org/10.1016/j.epsr.2016.09.028 0378-7796/© 2016 Elsevier B.V. All rights reserved.
Protection measures against electric shock in low-voltage TNconnected supply systems are described in many publications [1–12]. Installations employing generating sets for backup power must meet the requirements regarding electric shock prevention in the same way as grid-powered installations [2,3,5,7,9–12]. Protection against electric shock via automatic power supply switch-off imposes a permissible time (for switching off) in the case of insula-
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Fig. 1. Time–current curves of miniature circuit breakers a) type B, b) type C, c) type D.
tion fault. In the case of short-circuit in TN systems at the 230/400 V level, for overcurrent protection and working currents ≤32 A and for residual current devices (RCDs), the maximum switch-off time is set to 0.4 s [5]. In TN systems with a separate earthing conductor (PE), protection against electric shock is most often realized through residual current devices (RCDs) and fuses or miniature circuit breakers (MCBs) [8–18]; when a separate PE conductor is not available, the protection is based only on overcurrent devices. TN systems ensure a negligibly small risk of electric shock if the shortcircuit impedance of the PE conductor ZPE or PEN conductor ZPEN (PEN is a conductor that combines the functions of both a PE conductor and an N conductor) is small enough and the touch voltage Ut (as the drop voltage on the impedance of the ZPE or ZPEN conductors) does not exceed the permissible value Ut = 50 V [1,5]: ZPE = ZPEN ≤
Ut Ik
(1)
where Ik is the current flowing via the PE or PEN conductor. In the case of fuse-based overcurrent protection, Joule’s integral (which is the sum of a pre-arc thermal integral calculated from t = 0 to t = tp and the arc thermal integral calculated from t = tp to t = t0 ) is the crucial factor determining the tripping of the fuse under fault conditions [13]:
2
tP
ik2 dt
I t= 0
t0
ik2 dt
+
(2)
tP
where tO is the operating time, tp is the pre-arcing time (the thermal pre-arcing integral usually acting on the narrowing of a fuse link), and (tO − tp = ta ) is the arcing time. From expression (2), the thermal equivalent current Ith is frequently calculated [13,14] as follows:
Ith =
I2t tw
(3)
where tw is the duration of the thermal equivalent current. For a fuse, the Joule integral (2) is given in the catalogues by individual producers [15]. The response time of a fuse is determined by its time–current characteristics [15]. This time is dependent on equivalent short-circuit current Ith (3). The tripping condition of an MCB in the shorted circuit assumes that the RMS value of the short-circuit current does not reach values smaller than that of the assumed equivalent tripping current of the short-circuit release. It is very difficult to assess the tripping effectiveness of an MCB in electrical installations fed from low-power synchronous generator sets, up to several kVA, while considering the breaker’s time–current characteristics owing to relatively low values of short-circuit currents and high contents of transients in shortcircuit currents. In these electrical installations, the only reliable method to assess the tripping effectiveness of MCBs is the exper-
Fig. 2. An example of a power protection variant in TN-S system and synchronous generating set working as a backup power supply.
imental one. The practical choice of the nominal current for the circuit breaker is most commonly derived from the single-phase IN1ph or three-phase IN3ph rated current of the synchronous generator and the character of the load (most commonly R, RL or motor). Depending on the load characteristics, circuit breakers are chosen based on B, C or D time–current characteristics (see Fig. 1). Table 1 presents the tripping times of miniature circuit breakers of B, C or D time–current characteristics in relation to the k-ratio of the short-circuit current Ik = kIN (where IN is the rated current) [16–18]. Tripping times of miniature circuit breakers in relation to B, C or D characteristics are current dependent, thus to achieve fast enough switching off times the value of short circuit current has to have large values for a certain time. Minimal values of the currents required to achieve a breaker with a known characteristic to trip in time below 0.1 s are stated in the IEC standards [17,18]. For a tripping time 0.4 s, the thermal equivalent current can be estimated from characteristics (Fig. 1) or determined during the experiment. In a low-voltage electrical installation, the stability of the power source (of the electrical power system) is, by definition, the essence
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Table 1 Main technical details of elements and devices in control system of the 5.5 kVA synchronous generator supporting protection against electric shock. No. in Fig. 6
Element or device in the control system Fig. 6
Main technical details
1 2 3 4 5
Salient pole synchronous generator Single phase bridge rectifier Push-pull converter Signal conditioning board Programmable Logic Controller
6
Voltage transducers
SN = 5.5 kVA, UN = 400 V, In 3ph = 7.9 A, In 1ph = 13.8 A, nN = 3000 rpm KBPC610: 6 A, 1000 V ±15 V (1 A), 24 V (2 A) ULN 2804: 24 V, 8 × 500 mA CP1H −XA (Omron), digital (24 V): 24 inputs and 16 outputs, analog: 4 inputs and 2 outputs LV 25-NP (LEM): UN = 250 V - nominal measured voltage, IPN = 10 mA - primary nominal current RMS, tr = 40 s - response time to 90% IPN ±0.9% accuracy of UNph (TA = 25 ◦ C), linearity error <0.2% LA 100-P (LEM) IPN = 100 A - primary nominal current RMS, tr = <1 s - response time to 90% IPN ±0.45% accuracy of IPN (TA = 25 ◦ C), linearity error <0.15% di/dt accurately followed >200 A/s MCB time–current characteristics: B6, B8, B10, B13 C6, C8, C10, C13 SSR-480D125 Solid State Relays 480 V, 125 A, 0.02 ms turn “on” time
Current transducers
7
Overcurrent protection
8
Relay output board
Solid state relays
of measuring fault loop impedance (owing to the fact that electrical power system is usually modelled as a Thévenin’s equivalent circuit) [9]. The measurements of fault loop impedance are based on measurements of voltage drop on the internal impedance of the supply system. The measurements are usually taken by applying a large load (controlled short circuit) between each phase conductor and protective earth (PE) conductor or protective earth/neutral (PEN) conductor. However, as shown in this paper, in the case of electrical installations, after the switch from grid power to backup power utilizing synchronous generators, this backup voltage source may no longer be modelled as a Thévenin’s equivalent circuit, which utilizes a stationary voltage source and constant impedance. Thus, the process of the selection of the protection devices in such installations may not be undertaken in the same way as those supplied from electrical power systems. Under short-circuit conditions in electrical installations powered from a synchronous generating set, the shape of the instantaneous current is significantly influenced by transient phenomena occurring in the generator and AVR system powering the excitation circuit of the synchronous generator (working in conjunction with a closed feedback loop) [3,19–27]. In many reference guides and practical recommendations, the use of excitation controllers and systems for adjusting the voltage and power of synchronous generators cooperating with the power grid are mentioned [19–25]. The proposed control systems are used for: regulation and stabilization of the voltage and frequency of a generator working in a power system, limitation of interferences and failures, specification of the value of excitation current, specification of the power angle, the provision of limitations with long-term current overload protection, etc. In systems used at present to control synchronous generators working in combustion-engine-driven generation sets (as a backup power supply), there are no regulators controlling voltage or power in the case of a single-phase short-circuit to N or PE conductor, which would enable power switch-off in a time shorter than that required by law (0.4 s). This article presents the case of a single-phase short-circuit and the failure of the application of the overcurrent protection as an anti-shock measure. There are many manufacturers of generator sets on the market right now and some of them offer several types of generators. Every generator set is characterised by individual parameters (electrical and mechanical) and its own system of excitation control and forcing.
Usually, the manufacturers of the generator sets are not sharing any data regarding work and the characteristics of the applied AVR unit. The simulations performed using Matlab Simulink software on a system consisting of 22 kVA generator sets, using data obtained from the manufacturer and standard Simulink models (only data regarding the generator, but not the AVR system and speed control were available) showed, that there is a large difference between real short circuit current (obtained during laboratory experiments) and waveforms obtained from simulations. The characteristics of MCB (from different manufacturers) also show certain inaccuracies and should be modelled as band characteristics, what introduces some uncertainty into the determination of the tripping times of the breakers. These facts clearly show that adequate data regarding MCB switching off times during work with a small generator set can be obtained only from the experimental investigations. 2. Development of the control system supporting protection against electric shock In low and high power combustion-engine-driven generation sets, self-exciting synchronous generators with electromagnetic excitation are used more often than those with AVR systems, which are used for higher-power applications [3,19,22]. The generators with self-exciting arrangements can power a single load or groups of loads and can work in an IT, TT, TN-S or TN-C-S system [19]. To support the operation of a synchronous generator working as a backup power system and control stator currents and voltages (of a self-exciting synchronous generator), an AVR or electromagnetic excitation control system is usually used [3,22–27]. Generally, the tasks of AVR or electromagnetic control systems include the following: • maintaining short-circuit current at the level of approx. 2.5IN –6IN in the case single-phase short circuits (as in a conventional automatic voltage regulation or electromagnetic control systems used in synchronous generator sets)—according to [3], a short-circuit current should be only 3IN ; • maintaining an effective value of phase stator voltages, most often at the level of UNf in a no-load condition or a voltage with permissible deviation ıUst [3] in a load condition; • switching off the circuit in the case of overcurrent and risk of insulation damage [3].
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Fig. 3. Spectral reactance in d-axis vs. frequency for a 5.5 kVA synchronous generator.
In a short circuit, the correct operation of the anti-shock protection depends on a residual-current-operated circuit breaker, fuses or miniature circuit breakers. The speed of activation of residual circuit breakers is not questionable. However, owing to their frequent failures, they cannot (in terms of anti-shock protection) be the only devices to determine power supply switch-off, especially, for example, in firefighting vehicles [11,12]. That is why the standards [11,12,17,18], apart from residual circuit breakers, also necessitate the usage of overcurrent protection (fuses or any overcurrent circuit breakers) as an anti-shock protection supplement. In the case of a synchronous generating set (working in an autonomous system) the short-circuit condition results in the reaction of current, voltage and frequency controllers [19,21–23]. This has a significant impact on the waveform of the short-circuit current, which depends on changes in synchronous generator reactance (during the short circuit), the initial voltage phase, the distance of the short-circuit from the generator terminals and, among other things, the type of control system used for excitation voltage and current [19]. As shown in [9], given the examined 5.5 kVA synchronous generating sets—either in the case of external circuit insulation damage with additional resistance (modelling line conductor and PE or PEN conductor) or in the case of the “complete” insulation damage of a load (metallic short circuit at the terminal) − the reaction time of fuses is not always shorter or equal to that specified in the literature [5]. As shown in [19–21], the changes of examined synchronous generator reactances during the short-circuit process have a significant impact on the waveforms of short-circuit currents. Thus, when the transient phenomena become relatively long, the changes of reactance have a substantial influence on the value of the Joule integral (2) and thermal equivalent current (3). An example of a characteristic of spectral reactance in d-axis vs. frequency for a 5.5 kVA (400 V) salient pole synchronous generator is shown in Fig. 3. At the first moment during the short circuit, the value of reactance Xd is determined with f → ∞. The reactance changes shown in Fig. 3 are determined by a standstill frequency response (SSFR) test [28,29]. Usually, in electrical installations, the values of the short circuit currents, where circuit is supplied from mains, are calculated from the measurements of a fault loop impedance using Thévenin’s equivalent circuit. Such measurements are not credible for generator sets, where impedance changes (Fig. 3) accordingly to the generator state (subtransient, transient and steady state). Modern measuring equipment uses relatively small currents to determine circuit impedance, thus only steady state generator impedance is measured. During short circuits, the impedance of the synchronous generator changes, what causes the fault loop impedance change. This fact raises questions regarding the effectiveness of protection against electric shock carried out by a fast automatic power switch-off [9]. Assessing the tripping effectiveness of miniature circuit breakers in relation to permissible time is carried out during experimental investigations using the measurement set. Fig. 4 shows a
Fig. 4. Simplified diagram of the measurement set for investigation of the examined 5.5 kVA salient pole synchronous generator during protection against electric shock.
Fig. 5. Simplified diagram of the electronic phase detector 0–90◦ .
simplified block diagram of the measurement set for investigating the overcurrent protection of the examined 5.5 kVA salient pole synchronous generator during protection against electric shock. To ensure the same phase angle of the phase voltage supply (0◦ or 90◦ ), the short circuit is applied with the use of a special Electronic Phase Detector (Fig. 5) [9]. The short circuit conditions (Fig. 5) can be obtained (at the generator terminals or at the end of the line and N conductors) using an antiparallel connection of two thyristors (E2) and a phase controlled switch drive system (E5). The initial phase angles of the switching “on” of the generator phase voltage using thyristors (E2) is determined by the electronic pulse generator (E3). The microcontroller (E6) is used to control the phase trigger system (E5) and to calculate the first magnitude of short circuit current Im , duration of the thermal equivalent current tw and Joule’s integral I2 t according to expression (3) and RMS values of the thermal equivalent current Ith on the basis of the instantaneous current measurement using Hall-effect transducer (E1) and the special formed of the analog signal in system (E4). These data are displayed on the console screen (E7). The application of the microprocessor enables changing of the initial phase angles of the switching “on” of the phase voltage from ␣ = 0◦ to ␣ = 90◦ by the user through the console switches (E7) [9]. During the experimental investigations, the overcurrent protections of 8A type B and C (TCC) (for three-phase rated load current In 3ph = 7.9 A) and 13 A type B (for single-phase rated load current In 1ph = 13.8 A) were used. The presented measurement set up enables the influence of additional phase and PEN conductor resistance (Rd = RL + RPEN ) on the value of touch voltage UT to be determined for the equivalent thermal current Ith (Rd ). The examined 5.5 kVA synchronous generator is driven by a DC motor. The rotational speed of 3000 rpm is set by the frequency measurement
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Fig. 6. Control system of a synchronous generator supporting protection against electric shock.
(50 Hz) of the output phase stator voltage in the no-load state. The field current is adjusted through the electronic excitation controller system (manufacturer built-in AVR). The laboratory test bed with the proposed new control system (see Fig. 6) was built from the following blocks: three-phase synchronous generator set (1), power supply (2, 3), PLC-based controller (5), measuring and signal conditioning block (6), overcurrent protection (7), electromagnetic compatibility and relay output board (8). The output board consists of three capacitors 3C (facilitating the process of self-exciting and ensuring suitable electromagnetic compatibility), relays K1–K4, terminals L1, L2, L3 and N. The system of measurement and signal conditioning board (6) are connected to the generator terminals and consist of four transducers for measuring the phase-line voltages PU1, PU2, PU3 and excitation voltage PU4 and four transducers for measuring stator currents PI1, PI2, PI3 and field current PI4 [30–32]. The signal conditioning is performed using 2nd-order low-pass Bessel filters. The measured signals from block (6) are transferred to the PLC controller (5) inputs. The PLC is programmed to control output contactors S1–S4 (output board 8) via the signal conditioning board (4). The outputs of board 4 are two state outputs suitable to control output relays. Table 1 shows the most important technical details of elements and devices included in the control system of the 5.5 kVA salient pole synchronous generator supporting protection against electric shock (Fig. 6). The measured signals were acquired using digital four-channel oscilloscope type MSO 3014 (Tektronix) offering 100 MHz bandwidth and 1 mega-point record length on all channels. The new control system developed by the authors (Fig. 6) enables full power receipt switch-off in short-circuit conditions (if the RMS currents Ia , Ib or Ic measured with transducers PI1–PI3 are greater than set up level or/and voltage at any supply phase is lowered). The conditions, during which voltage is asymmetrical and lowered occur during asymmetrical short circuits (single phase to ground, two phases to ground). At relatively high-resistance faults (for example, during functional insulation damage) the asymme-
try of the RMS voltage measured with transducers PU1–PU3 can be used as the supply circuit trip signal. The examples shown in the results section were using the 20% asymmetry level (asymmetry of the generator’s phase voltages measured with the voltage transducers PU1, PU2 and PU3 in relation to the PE conductor) as an indicator to supply circuit switch off. Of course, in the proposed device, the asymmetry level is programmable and can be set between 10% to 80%. However, to assure the proper work of the system within the standards (10% voltage drops as well as asymmetry are allowed), during the experiments maximum sensitivity level of 20% was established. Thus, the short-circuit condition is the condition in which at least one of the PI1, PI2 or PI3 phase current transducers measures a current value greater than the set value (set within the range of 2IN –6IN ), with a simultaneous decrease of voltage in short-circuited phases and an increase in other phase voltages (asymmetry level increased above set value). The voltages in non-short circuited stator windings are increased, what is caused by the field current forcing by AVR system. The delay of the switch-off, after fault detection, can be adjusted by user from 0.2 s to 5 s using interface E7; thus, the time of the presence of dangerous voltage may be limited, depending on the effective value of the phase-line voltage and power supply conditions [5]. In the case activation failure of the standard overcurrent protection of a residual circuit-breaker or the continuation of a short circuit after the permissible time (0.4 s for 230/400 V), the short-circuited wiring is switched off.
3. Assessing of the tripping effectiveness of miniature circuit breakers in relation to permissible trip time without the new auxiliary control system To assess the effectiveness of the anti-shock protection, several experiments were conducted, without and with the proposed support circuitry. At the beginning of the experiments evaluating tripping times of the overcurrent protection devices (MCBs) with tripping characteristics B and C and with tripping current values 8 A and 13 A were performed, The values of the breakers currents
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Fig. 8. Registered waveforms for 8 A MCB with C (TCC) a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0 , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90 , ub − 90◦ and uc − 90◦ (for ia − 90◦ ). ◦
Fig. 7. Registered waveforms for 8 A MCB with tripping characteristic curve type B a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0 , ub − 0◦ and uc − 0 (for ia − 0◦ ); c) induced stator phase voltages ua − 90 , ub − 90◦ and uc-90◦ (for ia − 90◦ ).
◦
◦
◦
◦
were near the nominal values of current for the researched generator (In3ph = 7.9 A and In1ph = 13.8 A). In Figs. 7–10 the fault inception point (FTP) for the voltage ua phase angle 0◦ and 90◦ (FTP-0deg and FTP-90deg) and MCB opening point (MCB-OP0◦ and MCB-OP 90◦ ) are presented. Figs. 7 and 8 present the registered single-phase short circuit ia − 0◦ (for the phase angle 0◦ of voltage ua ), ia − 90◦ (for the phase angle 90◦ of voltage ua ), induced stator phase voltages ub and uc under no-load conditions and during the single-phase short circuit of the investigated salient pole synchronous generator with overcurrent protection of 8 A with TCC types B (Fig. 8) and C (Fig. 9). As shown in Figs. 7 and 8 for the examined 5.5 kVA synchronous generator, in the case of a short circuit in any initial phase of voltage and overcurrent protection of 8 A, for tripping characteristic curve B, the protection is activated in less than one period of 0.02 s; for tripping characteristic curve C (and in the same way for type D—Figs. 1 and 2) the interrupt reaction time will not cause the activation of protection in the required time (less than 0.4 s). Figs. 9 and 10 present the registered single-phase short-circuit ia − 0◦ and ia − 90◦ and induced stator phase voltages ua , ub and uc under the no-load condition and during the single-phase short circuit of the investigated salient pole synchronous generator with
the rotor skew for overcurrent protection of 13 A with tripping characteristic curve B (see Fig. 9) and 6 A with C (see Fig. 10). As shown in Figs. 9 and 10, for the examined salient pole synchronous generator with Sn = 5.5 kVA, Un = 230/400 V in the case of a short circuit in any initial phase of voltage, the overcurrent protection • of 13 A for type B and 0◦ initial phase of voltage causes the activation of the protection in an operating time of less than 0.02 s and for a 90◦ initial phase of voltage causes the activation of the protection in an operating time of less than 0.16 s. • of 6 A for type C causes the activation of the protection is less than 0.02 s. From the analysis of Figs. 7–10 (short circuits without the control system shown in Fig. 6 and without the additional external resistance Rd = RL + RPEN − Fig. 4), it can be concluded that the greatest value of overcurrent protection type C should not exceed 6 A, if the required tripping time is less than 0.4 s. The above result raises questions about the utilisation of the full power of the proposed generator because the maximum tripping current of the C type characteristic fuse ensuring proper anti-shock protection is 6 A. The conclusion from this investigation is that when using the MCBs, either generator load and/or load type must be limited, or safety rules must be revised.
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Table 2 Comparison of characteristic values for MCB during short circuits in system powered from generator. Miniature Circuit Breaker TCC
B C
Phase (deg)
0 90 0 90
6A
8A 2
2
10 A 2
2
13 A 2
2
Im (A)
I t (A s)
Ith (A)
tw (ms)
Im (A)
I t (A s)
Ith (A)
tw (ms)
Im (A)
95 89 95 83
33 34 35 33
57.1 53.2 56.4 44.6
10 12 11 17
95 86 97 93
35 33 508 503
56.4 44.6 31.8 31.1
11 17 503 520
99 136 38.0 94 98 89 130 37.0 95 89 operating time is always greater than 0.4 s operating time is always greater than 0.4 s
Fig. 9. Registered waveforms for 6 A MCB with TCC a) single-phase short-circuit ia − 0◦ and ia − 0◦ ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc -90◦ (for ia − 90◦ ).
A detailed comparison of maximum values of current Im , Joule’s integral I2 t, thermal equivalent current Ith and duration of thermal equivalent current tw obtained from experimental investigations for single-phase short circuits and miniature circuit breakers is presented in Table 2. The range of fuses includes ones with time–current characteristics B (6 A, 8 A, 10 A and 13 A) and C (6 A and 8 A). The values presented in Table 2, obtained from the experiment, show, that for the investigated 5.5 kVA salient pole synchronous generator, the requirements of electric shock protection (tripping time less than 0.4 s) are satisfied only for the following MCB’s:
I t (A s)
Ith (A)
tw (ms)
Im (A)
I2 t (A2 s)
Ith (A)
tw (ms)
42 206
59.0 41.8
12 118
Fig. 10. Registered waveforms for 13 A with TCC B a) single-phase short-circuit ia − 0 and ia − 90 ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0 ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc − 90◦ (for ia − 90 ). ◦
◦
◦
◦
• B in the 6–13 A range of rated current and in tripping times from 10 ms to 118 ms, • C only in the 6 A range of rated current in tripping time 11–17 ms The miniature circuit breakers with time-current characteristic C (8 A, 10 A and 13 A) do not guarantee compliance with the requirements of electric shock protection through automatic disconnection of the supply under short-circuit conditions because their operating time is greater than the required 0.4 s. For these breakers (C8 A, C10 A and C13 A), several measurements of the
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Fig. 11. Registered waveforms with the use of the control system (Fig. 7) for 8 A with TCC C overcurrent protection a) single-phase short-circuit ia − 0◦ and ia − 90◦ ; b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc − 90◦ (for ia − 90◦ ).
short-circuit waveforms showed relatively small values of the RMS thermal equivalent current Ith (3) for the time tw = 0.4 s.
4. Experimental investigations with the use of the new auxiliary control system supporting the operation of a synchronous generator The conclusions from the investigation presented in the previous chapter clearly state that overcurrent protection support devices are necessary to ensure full utilisation of the generator power. This was the reason for the development of the circuitry shown in Fig. 6. The presented auxiliary control system of the examined 5.5 kVA synchronous generator ensures that the time of persistence of dangerous prospective touch voltage during a short circuit directly at the generator terminal, or in the case of electric insulation failure with additional resistance, is no longer than allowed. Fig. 11 presents the registered single-phase short-circuit ia − 0◦ , ia − 90◦ and induced stator phase voltages ua , ub and uc under the no-load condition and during the single-phase short circuit of the investigated salient pole synchronous generator for overcurrent protection of 8 A with tripping characteristic curve C, with the use of the control system supporting the operation of a synchronous generator. In this case, the S1 relay (see Fig. 6) is switched off when the set time equal to 0.4 s elapsed.
Fig. 12. Registered waveforms using the control system for 8 A with C (overcurrent protection trip characteristic curve) and during the high resistance fault (Rd = 3.43 ) of: a) single-phase circuit ia − 0◦ and ia − 90◦ b) induced stator phase voltages ua − 0◦ , ub − 0◦ and uc − 0◦ (for ia − 0◦ ); c) induced stator phase voltages ua − 90◦ , ub − 90◦ and uc -90◦ (for ia − 90◦ ).
Previous investigations showed that the use of type C overcurrent protection with a rated current greater than 6 A caused the persistence of dangerous prospective touch voltage for more than 0.4 s. From Fig. 11, it is visible that the supporting system switched off the short-circuited branch within the set time (0.4 s), since fault occurs at 0.1 s and switching “off” the circuit takes place at the 0.5 s time mark. Fault generation was synchronized witch oscilloscope and on the following figures occurs always at 0.1 s for “0◦ ” switching angle and is delayed of 0.005 s for “90◦ ” switching. Fig. 12 presents the recorded single-phase short circuit ia − 0◦ , ia − 90◦ and induced stator phase voltages ua , ub and uc under the noload condition, and during the high resistance fault (Rd = 3.43 ) of the investigated salient pole synchronous generator. The tests were carried out for 8 A overcurrent protection with tripping characteristic curve C (the same for 8 A with tripping characteristic curve B), using the control system supporting the operation (see Fig. 2) of the examined synchronous generator. In this case, relay K1 is switched off. The results revealed in Figs. 11 and 12, show that for the examined 5.5 kVA synchronous generator in the case of the utilisation of the proposed protection support circuitry, the switch-off during the fault always occurs within 0.4 s. This happens not only in the case of fault close to the generator when the short-circuit currents are quite high but also when a large impedance fault (insulation defect)
K. Ludwinek et al. / Electric Power Systems Research 142 (2017) 341–350
is simulated by applying additional resistance. The necessary conditions for the proposed circuitry to switch off during a single-phase fault include two basic requirements: the current in one of the phase conductors must increase above a certain adjustable level, and asymmetry of the generator voltages must exceed 20%.
5. Conclusions On the basis of experimental investigations without the control system presented in this paper, it can be concluded that the overcurrent protections with tripping characteristics B and C selected for rated currents of the examined synchronous generator (one phase rated load and three phase rated load), both for single-phase and three-phase loads, can be ineffective as anti-shock protection. Performed tests aimed at determining the effectiveness of a fast automatic power switching-off of the electrical installation powered by a synchronous generator showed that the application of smaller-than-rated or fast switching breakers is required to satisfy the stated anti-shock protection criteria, which can limit the power or type (electric machines) of the loads. The problem is associated with the limitation of the short-circuit current in circuits where only overcurrent protection is used, what can cause long persistence of an effective touch fault voltage greater than 50 V between a given fault point and reference earth, which is not allowed by given standards. The other option to switch off the short-circuited wiring during the required time may be to individually select output section wiring with the use of a separate overcurrent protection of trip characteristic curves B or C selected on each section. This increases the cost of electrical system, since in order to decrease system impedance the diameters of wires have to be increased. The overcurrent protections of trip characteristic curves B or C in the sections are selected by measuring the Joule integral of the short circuit current. This solution in the output sections will indeed enable power switch-off in the required permissible time, but still does not eliminate the causes of output power limitation (in the sections). The research shown in this article concentrated on a 5.5 kVA low-power generator used for laboratory set up, however, the problem of the anti-shock protection is also present in large power generator units. This fact was noticed by the authors during professional expertises performed on such objects as hospitals or airports. For example, for 200 kVA unit (airport radar system), the manufacturer of the generator uses excitation forcing to assure the proper value of short circuit current (above 3IN ) and suggests that the current of the single output should not exceed 20% of generator nominal current. Such arrangement allows the limitation of the value of circuit breaker current to 0.2IN , thus creates the condition when short-circuit current is 15 times greater than nominal current of the breaker, however, the load of single circuit powered by the generator is limited to 40 kVA. This approach works well for large generator sets (400 kVA and larger units) mounted, for example in hospitals, where no single loads above 20% of generator power are expected. The auxiliary protection system presented in this paper is one of the means to solve that problem. The proposed circuitry uses both voltage and current measurements to detect single-phase or two phase to ground short circuits (short circuits of phases to neutral or earthing conductor) what is a novel approach at low voltage (below 1 kV) level. This arrangement allows also switch-off of the load when the short-circuit current is limited by additional impedance introduced by the nonzero impedance fault (insulation burn down). Moreover, if the device is mounted in each section of output circuitry there is possible (by applying proper settings) to maintain the selectivity of the section protection and overcurrent protection of the generator.
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The drawback of the presented control system is that it requires quite costly elements as voltage and current measuring sensors and output switches (for each output section) that are able to turn off short-circuit currents. However, it is possible to build an integrated measurement and control system with a single voltage sensor for a single switchgear and separate current sensors for each output, what will lower overall costs of the device. The cost of electronic control board is almost negligible when compared to costs of the switches at the output. As mentioned above, these switches have to be able to turn off short circuit currents for switchgears supplied from the generator (auxiliary supply) or from the mains during normal supply conditions. References [1] M. Barrett, K. O’Connell, A.C.M. Sung, Analysis of transfer touch voltages in low-voltage electrical installations, Build. Serv. Eng. Res. Technol. 31 (1) (2010) 27–38. [2] F.R. Blánquez, C.A. Platero, E. Rebollo, F. Blázquez, On-line stator ground-fault location method for synchronous generators based on 100% stator low-frequency injection protection, Electr. Power Syst. Res. 125 (2015) 34–44. [3] LV Generator Protection, Low Voltage Expert Guide No. 8, Schneider Electric, 2001. [4] HD 60364-5-551-2010, Electrical installations of buildings - Part 5–55: Selection and erection of electrical equipment - Other equipment - Clause 551: Low-voltage generating sets, 2010. [5] HD 60364-4-41-2005, Low-voltage electrical installations −part 4–41: Protection for safety: Protection against electric shock, 2005. [6] HD 60364-5-54-2007, Electrical installations of buildings −part 5–54: selection and erection of electrical equipment, Earthing arrangements, Protective conductors and protective bonding conductors, 2007. [7] HD 60364-5-551-2010, Electrical installations of buildings - Part 5–55: Selection and erection of electrical equipment - Other equipment - Clause 551: Low-voltage generating sets, 2010. [8] S. Czapp, Protection against electric shock using residual current devices in circuits with electronic equipment, Elektronika ir Elektrotechnika 21 (4) (2007) 51–54. [9] P. Makarski, System for measuring tripping time in safety fuses and circuit breakers, Polish Patent No. PL392550-A1, PL218499-B1. [10] ISO 8528-5:2013, Reciprocating internal combustion engine driven alternating current generating sets - Part 5: Generating sets, 2013. [11] DIN 14687-2007-02, Firefighting Equipment - Permanently installed generators (generating sets) less than 12 kVA for the use in firefighting vehicles, 2007 (In German). [12] DIN 14686/A1:2015-03. Firefighting equipment - Switch cabinets for fixed generators (generating sets) > = 12 kVA in fire-brigade vehicles, Amendment A1, 2015 (In German). [13] J. Nahman, D. Stojanovi, Calculation of thermal equivalent short − time current, Facta Univ. Ser. Electron. Energetics 1 (1) (1998) 103–114. [14] M. Gianpasquale, Thermal equivalent short-time currents in three-phase balanced short circuits, IEEE Trans. Power Deliv. 9 (2) (2004) 897–898. [15] Special purpose fuses, ETI. Catalog (2015). [16] ASTI: Miniature circuit breakers and Residual current devices, ETI Catalogue (accessed 11.08.16). [17] EN 60898-1:2003/A13:2012 E. Electrical accessories - Circuit breakers for overcurrent protection for household and similar installations - Part 1: Circuit-breakers for a.c. operation, 2012. [18] EN 60898-2:2006. Electrical accessories - Circuit-breakers for overcurrent protection for household and similar installations Part 2: Circuit-breakers for a.c. and d.c. operation, 2006. [19] G. Andersson, Modelling and Analysis of Electric Power Systems. Power Flow Analysis. Fault Analysis. Power Systems Dynamics and Stability, ITET ETH, Zurich, 2004. [20] J. Pyrönen, T. Jokinen, V. Hrabcová, Design of rotating electrical machines, John Wiley, 2012. [21] R. Wamkeue, N.E.E. Elkadri, I. Kamwa, M. Chacha, Unbalanced transient-based finite-element modeling of large generators, Electr. Power Syst. Res. 56 (3) (2000) 205–210. [22] C.S. Hoong, S. Taib, K. Mieee, S. Rao, I. Daut, Development of automatic voltage regulator for synchronous generator, in: National Power & Energy Conference (PECon) 2004, Proceedings, Kuala Lumpur, Malaysia, Nov. 29–30, 2004, pp. 180–184. [23] IEEE Std 421.5-2005. IEEE Recommended practice for excitation system models for power system stability studies, 2005. [24] IEEE Std 421.4-2004. IEEE Guide for the preparation of excitation system specifications, 2004. [25] P421.2/D20, Nov 2013. IEEE Draft guide for identification, testing, and evaluation of the dynamic performance of excitation control systems, 2013. [27] M. Zielichowski, T. Szlezak, Improvements in ground-fault protection of unit-connected generator stator winding using laboratory test environment, Electr. Power Syst. Res. 78 (9) (2008) 1635–1639.
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