Overvoltage Assessment of Point-to-Point VSC-Based HVDC Systems

Published by Jose A. JARDINI1, Ricardo L. VASQUEZ-ARNEZ2, Marcos T. BASSINI1, Marco A.B. HORITA1, Gerson Y. SAIKI2 and Marcos R. CAVALHEIRO3
Polytechnic School of the University of Sao Paulo (1), Foundation for the Technological Development of the Engineering Sciences (2), CTEEP Sao Paulo State Power Transmission Company (3).

Abstract. The application of voltage source converters (VSCs) into medium- and high-power transmission is currently attracting increased attention. In view of this increased attention, this article provides the simulation results of the overvoltages produced by faults occurring in the DC line of a point-to-point VSC-based HVDC system as well as in a neighbouring AC line system. The VSC converters considered here use the MMC (modular multi-level converter) technique to generate the voltage waveform. For pole-to-ground faults occurring in the DC link of a symmetrical monopole system, significantly high overvoltages may arise on the sound pole. This condition is of concern, mainly during the planning stage of the VSC-HVDC project, as it may require the installation of surge arresters with a good performance and/or also additional insulation of the line. In addition, unless the faults occurring on the DC link are quickly removed, sustained overvoltages can threaten the normal operation of the surge arresters installed on the DC side of both sound and faulty poles. Faults and other events in the AC system (near the DC link) may lead to sustained overvoltages that should also be examined regarding the response of the surge arresters.

Streszczenie. Artykuł przedstawia rezultaty symulacji przepięć występujących w sieci DC bazującej na wykorzystaniu VSC (voltage source converters). Układ VSC wykorzystuje technikę MMC (modular multi-level converter). Rozpatrzono też wpływ sąsziadującej z siecią DC sieci AC. Analiza przepięć w sieci prądu stałego HVDC wykorzystującej technikę VSC.

Keywords: Faults, MMC technique, Overvoltages, Point-to-point system, Surge arresters, VSC-HVDC.
Słowa kluczowe: przepięcia, sieci DC, zabezpieczenia przed przepięciami

Introduction

Traditionally, HVDC systems using line commutated converters (LCCs) are utilized for the transmission of bulk power over long distances. Recently, voltage sourced converters (VSCs) applied to HVDC systems also appeared as good candidates for the transmission of relatively large power (up to 1000 MW) at high voltages (up to ± 400 kV).

According to their topology, HVDC systems can be classified into various categories. However, for the purpose of this study, the configurations utilized are: a symmetrical monopole system, in which there is commonly no grounding system (Fig. 1a) and a bipolar system, with two converters in each end and where the returning path (with one pole out of service) can be accomplished through a metallic return or through the ground (Fig. 1b).

While the converters in LCC-HVDC applications use thyristors and can only control active power, the VSCHVDC configuration uses IGBTs (Insulated Gate Bipolar Transistors) as switching devices that can independently control both active and reactive power. This control is an important characteristic because in the LCC-HVDC alternative, the reactive power required by the converters, as well as the required filters, must be provided locally. Another important issue is that, due to the turn on/off characteristic of the IGBTs, the resulting voltage waveform is close to a sine wave, thereby requiring much smaller filters in the AC side.

The presence of overvoltages in DC systems using VSC-HVDC can be due to various causes, namely, faults in the system, loss of the inverter terminal, loss of load, etc. The reconnection of the rectifier, once it has been blocked, e.g., due to a fault, can also be responsible for the presence of overvoltages [1].

In underground cables, due to their physical separation and arrangement, the occurrence of pole-to-pole faults is rare. In contrast, DC overhead lines may be more prone to face such type of faults. Pole-to-pole faults are of special concern because it may cause failure of the semiconductor devices [2].

Within the VSC-HVDC technology, the Modular Multilevel Converter (MMC) has recently gained popularity due to its inherent advantages, namely: lower switching losses (i.e., lower switching frequency in each submodule), lower voltage across each switch (as only small capacitors are used in parallel with each switch), among others [3].

Recently, some MMC simulation models aimed at speeding up the computational time were proposed. For example, a detailed description of the MMC ‘average model’ and its dynamic performance under both balanced and unbalanced grid operation modes are presented in [4], [5] and [6]. A revision of the methods typically used for the protection of VSC-based HVDC systems is presented in [7]. The referred article, however, chiefly focuses on the overcurrent protection. The overvoltage protection and insulation coordination in MMC-based HVDC systems are explored in [8]. All DC lines in this reference are submarine cables, and the study conducted is focused on the insulation design of the DC line.

The effect of DC faults and their respective protection scheme in VSC-multi-terminal high voltage direct current (MTDC) transmission systems is presented in [9]. The scheme uses IGBT-based circuit breakers and its respective coordination with the converter switches (IGBTs) to block the converter during fault periods.

Fig.1. DC transmission configurations used: (a) symmetrical monopole and (b) bipolar system.

When a line-to-ground fault occurs in a bipolar system, the faulted pole rapidly discharges the capacitor(s) to ground. This discharge causes an imbalance of the DC link voltage between the positive and negative poles. As the voltage of the faulted line begins to fall, high currents flow from the capacitor(s) as well as from the AC grid. These high currents may damage the converters and the capacitors [10].

During faults in the DC link of a VSC-HVDC system, and also during contingencies in the AC system (involving line and/or transformers) nearby the converter stations, some overvoltages in the DC line can appear. Such overvoltages are of concern because they may have an influence on the design of the line and on the surge arresters located on the DC side of the converters.

In view of these design considerations, this article presents the magnitude of the overvoltages arising from faults occurring inside the DC link. In addition, the effect of line disconnections in another neighbouring AC grid configuration, due to inherent faults, is also presented.

Description of the DC System

The VSC-HVDC system used here is based on the MMC technique. Additionally, the MMC converter referred to as Model 2 in [5] was used herein. A two-pole 400-km DC cable and an 800-km overhead line were independently used in the DC link of the point-to-point DC system. The nominal DC voltage and transmitted power are equal to ±320 kV and 800 MW, respectively. Both bipolar and symmetrical monopole systems, along with the AC voltages and transformer characteristics, are shown in Figs. 2(a) and 2(b), respectively. The direction of the power flow shown in these figures was set to occur from Terminal 2 to Terminal 1; however, it could also be set to flow in the opposite direction.

Fig.2. (a) Bipolar and (b) symmetrical monopolar VSC-HVDC configurations used.

Regarding the DC link, both positive and negative cables have standard configurations, with a core (whose diameter is equal to 0.05 m), sheath and armour, as described in the Appendix section. The separation distance between both cables is 0.5 m, and the cables are buried at a depth of 1.5 m.

The overhead transmission line (OHTL) considered has a bundle of three ACSR conductors per pole (Chukar conductor) located at a height of 33.2 m at the tower and 14.1 m at midspan. The ground resistivity considered in both cases (i.e., cable and OHTL) was equal to 100 Ω–m. Likewise, a fault resistance equal to 0.01 Ω was considered in all the simulations.

Overvoltage Analysis

Both symmetrical monopole and bipolar VSC-HVDC systems were simulated using the PSCAD program [11] with the EMTDC as its numerical solver. For some other calculations the EMTP-RV program [12], was also used. The following faults were analysed for both the cable and the OHTL case:

a) Symmetrical Monopole System
1. Pole-to-pole fault
2. Pole-to-ground fault

b) Bipolar System
1. Pole-to-pole fault
2. Pole-to-ground fault

In all simulated cases, the fault (applied at t = 3.0 s and extinguished forcedly after 200 ms) was assumed to occur in the middle of the DC link (negative pole). This is because this point is one of the most critical in terms of overvoltage.

Overvoltages due to Faults in the DC Link During a fault in the DC link, two instants of overvoltage normally occur. The first one occurs at the beginning of the DC fault itself. The second one occurs at the instant when the fault is cleared. The major concern here is the overvoltages during the fault period. Note that independently of the power flow direction, both equivalent sources contribute to the fault current.

Next, the overvoltages and the behaviour of the DC link, due to the simulated faults, will be shown. The magnitude of the overvoltages at the beginning (Vdc_pos1), the middle (Vdc_pos2), and the end of the DC line (Vdc_pos3), as indicated in Fig. 2(b), will be presented. Measurements at these same points of the negative pole were also obtained. However, our main focus will be on the overvoltages created in the sound pole. A summary of all the values obtained in the study is presented in Table 1.

Fig.3. (a) DC voltage at the middle of line Vdc_pos2, Vdc_neg2, (b) voltage at Vdc_pos1, Vdc_pos3, (c) DC fault current during the pole-topole fault.
A. Use of a Cable in the DC Link

a) Symmetrical Monopole System

1. Pole-to-pole fault
The voltage at the fault point in both sound and faulted pole (Vdc_pos2 & Vdc_neg2) falls to zero (Fig. 3a), whereas the peak overvoltages at Vdc_pos1 and Vdc_pos3 (sound pole) are 369 kV and 372 kV, respectively (Fig. 3b). The peak of the DC current (20 ms after the fault was initiated) was approximately 51 kA (Fig. 3c).

2. Pole-to-ground fault
While the voltage at the negative pole (faulted pole) drops to a very low value (close to zero), the sound pole exhibits a significantly large peak overvoltage at the middle of the line (Vdc_pos2 =834 kV), as depicted in Fig. 4. The maximum overvoltages read at both the beginning and end of the sound pole (Vdc_pos1 & Vdc_pos3) were approximately 782 kV. These overvoltages are mainly due to the displacement of the (virtual) neutre in the DC link of the symmetric monopole system. Under the no fault conditions, this virtual reference is zero, with each pole operating at ±320 kV. The instant the negative pole drops nearly to zero (due to the fault) this reference is displaced; thus, resulting in higher DC voltages on the sound pole.

Fig.4. Voltage at the fault point (sound pole) for a pole-to-ground fault.

b) Bipolar System

1. Pole-to-pole fault
Close results to those presented in Section A.a.1 were obtained (i.e., the voltages at the middle point Vdc_pos2 and Vdc_neg2, fall to zero; whereas the values of Vdc_pos1 and Vdc_pos3 are equal to 302 kV and 305 kV, respectively). The peak fault current was equal to 33.7 kA.

2. Pole-to-ground fault
The peak overvoltage at the sound pole (Vdc_pos2) was equal to 386 kV (Fig. 5a), whereas at each end of the line relatively lower voltages (Vdc_pos1 = 323 kV and Vdc_pos3 = 332 kV) were obtained (Fig. 5b). These overvoltages can present no threat to the analysed system.

Fig.5. Pole-to-ground fault in the bipolar system: (a) voltage at the middle of the line (Vdc_pos2) and (b) voltage at the receiving and sending-end of the DC cable (Vdc_pos1, Vdc_pos3).
B. OHTL as DC Link

a) Symmetrical Monopole System

1. Pole-to-pole fault
The peak overvoltages read at Vdc_pos1 and Vdc_pos3 were 385 kV and 382 kV, respectively. As expected, the voltage at the middle of both poles fell to zero. The value of the peak fault current was approximately 35.7 kA.


2. Pole-to-ground fault
The greatest peak overvoltage at the sound pole is approximately 989 kV (Vdc_pos2 shown in Fig. 6b). In Figs.6(a) and 6(c), the values at each end of the DC line (Vdc_pos1 = 956 kV and Vdc_pos3 = 969 kV) are shown. The peak of the fault current was approximately 2.6 kA.

Fig.6. Pole-to-ground fault: (a) voltage at the receiving end of DC OHTL (Vdc_pos1), (b) in the middle of the DC line (Vdc_pos2), and (c) at the sending-end (Vdc_pos3).

b) Bipolar System

1. Pole-to-pole fault
The peak overvoltages at the beginning and end of the DC line were 433 kV (Vdc_pos1) and 422 kV (Vdc_pos3), as shown in Fig. 7. The peak of the fault current was approximately 26 kA.

2. Pole-to-ground fault
The peak overvoltage at the sound pole (Vdc_pos2 in Fig.8b) was approximately 634 kV. The values at both ends of the line were: Vdc_pos1 = 376 kV (Fig. 8a) and Vdc_pos3 = 394 kV (Fig. 8c). The maximum value of the fault current was equal to 18.4kA.

The overshoot inside the ellipse shown in Fig. 8(b) is also due to the contribution of the travelling waves along the DC line. This overshoot (further amplified in Fig. 9) coincides with the time taken by the travelling wave after bouncing off the converter stations at both ends (i.e., 2.8 ms after the occurrence of the fault). If in this particular conductor (OHTL) the wave propagation speed is close to the speed of light, it can be shown that the total distance travelled by this wave is equal to the length travelled on the DC line (i.e. 2×400 km in 2.8 ms). Each fast travelling wave (depicted in Fig. 10) bounces in each end of the line creating reflections which when added to the existing overvoltage it can reach values twice or more of those encountered during normal conditions.

From Table 1, it can be seen that pole-to-pole faults have the highest fault currents. Pole-to-ground faults in the case of the monopole system exhibit the highest overvoltages. Faults occurring in the DC link of the bipolar system do not create very significant overvoltages that could threaten the line’s BIL (Basic Impulse Insulation Level). The minimum BIL calculated for equipment within the substation was equal to 750 kV, whereas the BIL for the DC line resulted in 1900 kV. According to [13], the latter critical flashover voltage corresponds to a 300 kV line whose insulator chain has 18 standard disc insulators.

Fig.7. Pole-to-pole fault in the presence of the OHTL: Voltages at points Vdc_pos1 and Vdc_pos3.
Fig.8. Voltages at the fault point (pole-to-ground fault) of the sound pole.

Also, faults occurring in the dc cable resulted in relatively higher fault currents compared to the values obtained for the OHTL case (Table 1). An opposite behaviour was observed for the case of the overvoltages (i.e. the DC cable exhibited relatively lower overvoltages in relation to the OHTL).

Despite these overvoltages do not seriously threaten the line’s insulation level, it is also necessary to assess the surge arresters effectiveness for reducing such overvoltages. Such analysis is presented in the subsequent section.

Fig.9. Zoom-in of Fig. 8(b) indicating the effect of the travelling wave.
Fig.10. Travelling waves along the DC line.

Table 1. Overvoltage for the different type and fault configurations

P-P: Pole-to-pole fault.
P-G: Pole-to-ground fault.

Use of Surge Arresters Zinc Oxide (ZnO) arresters are commonly used for the protection against overvoltages in AC and DC systems. So, a ZnO arrester was placed at point Vdc_pos1 (sound pole), as depicted in Fig. 11. The rated voltage of the selected surge arrester is 240 kV (rms). Its V-I curve is presented in the Appendix. Additionally, the maximum energy that this particular arrester can absorb is 7 kJ/kV.

Fig.11. ZnO arresters placed in the sound (positive) pole.

The results obtained after the installation of a surge arrester for the symmetrical monopole configuration (pole-to-ground fault) when the DC link separately uses a cable and an OHTL, are presented in Figs. 12 and 13, respectively.

The reason for initially placing only one arrester is to better observe the effect of this overvoltage suppressor at a particular point along the line, in this case point Vdc_pos1. In practice, one arrester should be placed at each end and in each pole of the DC link. For both, cable and OHTL, the overvoltage at the protected point (Vdc_pos1) was effectively reduced to the arrester protection level (549 kV in Fig. 12a and 517 kV in Fig. 13a), but remained high in the other unprotected points.

Fig.12. (a) Overvoltage reduction due to the installation a ZnO arrester at Vdc_pos1, (b) voltage at Vdc_pos2, and (c) voltage at Vdc_pos3 (cable as DC link).
Fig.13. (a) Overvoltage reduction due to the ZnO arrester at Vdc_pos1, (b) voltage at Vdc_pos2, and (c) voltage at Vdc_pos3 (OHTL).

In Table 2, a summary of the voltages measured at Vdc_pos1, Vdc_pos2 and Vdc_pos3 is presented.

Subsequently, the surge arrester was taken to the middle of the line (positive pole). Although this option (surge arrester at middle of the line) might sound a bit uncommon, it can actually be the case of installing a “line” surge arrester on the tower itself, similar to the installation of line arresters in some ac systems.

In Table 3, the reduction of the overvoltage at point Vdc_pos2 (552 kV/320 kV=1.75 pu cable; and 532 kV/320 kV =1.66 pu OHTL) for the symmetrical monopole system, is presented. However, as expected, the voltages close to Terminals 1 and 2 (with no arresters) are still high, above 1.95 pu at both line ends. Therefore, it will be necessary to install surge arresters in more than one point.

So, surge arresters in the middle and at both ends of the DC line were placed. The three arresters effectively reduced the DC link overvoltages to the arresters protection level (Table 4). A brief analysis on the energy absorbed by the arresters is presented in Section 6.

Table 2. Overvoltage reduction at Vdc_pos1 after the installation of the ZnO surge arresters.

.

Table 3. Overvoltages at the sound pole with surge arrester installed only at middle of the (+) pole.

Energy absorbed by the arrester during the entire fault period:
DC cable case: Earr2 = 407.7 MJ
OHTL case: Earr2 = 376.4 MJ

Table 4. Voltages at the sound pole with surge arresters installed at the three points.

Energy absorbed during the entire fault period (200 ms):
DC cable case: Earr1 = 156.8 MJ, Earr2 = 162.8 MJ, Earr3 = 144.8 MJ.
OHTL case: Earr1 = 151.5 MJ, Earr2 = 153.6 MJ, Earr3 = 138.1 MJ
Energy Absorbed by the Specific Surge Arrester(s)

During prolonged (sustained) overvoltage conditions a considerably large amount of energy might need to be absorbed by the arrester. This high energy absorbed might threaten the operation of the installed surge arrester. The energy absorbed by a single arrester located at point Vdc_pos1 (Energ_p1) is shown in Fig. 14. It can be observed that the energy absorbed keeps rising beyond the arrester’s limit (1700 kJ) which imposes a high risk of damage to the arrester.

Fig.14. Energy absorbed by only one surge arrester (Energ_p1) at Vdc_pos1 during the pole-to-ground fault (OHTL).

Now, if it is to prevent the arrester from being damaged, the DC line should be disconnected in less than 5 ms after the fault has started (at t = 3.0 s), as indicated by the segmented line of Fig. 14. The installation of three arresters (at both ends and the middle of the positive pole) also reduced significantly the energy absorbed by each arrester (see Table 4). The negative pole (line) will also require similar arresters at the same locations as those indicated for the positive pole. Closer values of the absorbed energy were obtained after the installation of surge arresters in the point-to-point system using a cable as the DC link. An alternative proposal is the installation of a resistor and chopper with sufficient energy capability at both stations.

Response of the Bipolar HVDC System Towards Faults Occurring in a Nearby AC System The point-to-point bipolar system depicted in Fig. 2(a) was embedded into a digital model of an existing AC grid, whose operating voltage is equal to 500 kV (see Fig. 15). Four AC equivalent generators feed the entire AC grid (60 Hz). All of the transmission line parameters were represented through coupled pi sections to consider their values of series impedance and shunt admittance. To conduct this analysis, the set points of the transmitted power and the voltage of the point-to-point bipolar system were kept at 400 MW/pole and +/-320 kV/pole, respectively.

Whenever an event during t(-) to t(+) in the AC system occurs (near either DC link terminal), there will be a change in the voltage angles at buses 550 and 506. The t(-) time refers to the instant just before the event, whereas t(+) refers to the instant just after the event. The angle of the AC voltage generated by the VSC remains the same from the time prior to the event occurrence. From t(-) to t(+) there will occur a change in the active power of the converters at terminals 1 and 2. The resulting (instantaneous) imbalance in these powers change the voltage in the DC line; thus, the DC voltage may increase or otherwise decrease (e.g. if the input power + losses + output power within the VSC-HVDC scheme increase; then, the DC voltage will rise, boosted by the DC link capacitor, conversely, the DC voltage will decrease). After a certain time, the controls of the converters bring both powers (P1 and P2) to the pre-event condition. The presence of this type of event in the AC system shall be simulated to check the extent these DC overvoltages are and the amount of stress that they can cause upon the DC arresters.

Fig.15. Point-to-point bipolar system within a 500 kV AC grid.
Fig.16. (a) Rms voltage at terminal T2 (sending-end) and T1 (receiving-end) and (b) AC side power (receiving-end) close to Terminal 1 for fault F4.
Fig.17. Voltages at the beginning and end of the DC link (negative pole) for fault F4.

Four types of events were applied: F1: three-phase-to-ground fault (at t = 4.0s) after which the affected line (912 – 4942) trips off at t = 4.1 s (duration of fault: 200 ms); F2: sudden permanent disconnection of two parallel lines located between buses 505-506; F3: three-phase–to-ground fault (at bus 506) through a 25 Ω reactor (fault is removed after 100 ms) and, F4: three-phase fault that causes the simultaneous permanent disconnection of three parallel lines located between buses 505-506. In all cases, low oscillations of the overvoltages were observed; however, they did not affect the DC line arresters located at the same three points of each pole. As an example, the results of the disconnection of three parallel lines between buses 505-506 (Fault F4) are presented (Figs. 16 and 17).

(Vrms_T1 = 482 kV & Vrms_T2 = 513 kV) are shown in Fig. 16(a). Notice that with the disconnection of the three lines between buses 505 and 506, the rms voltage at Terminal 1 (closer to the fault point) reaches a new operative point (around 460 kV) in relation to the pre-fault value (482 kV). This is partly due to the higher losses now present in the remaining line. In order to ensure a proper operation of the converters, this voltage drop will have to be compensated by the transformer tap changer.

The rms voltages (AC side) of Terminals 1 and 2 (Vrms_T1 = 482 kV & Vrms_T2 = 513 kV) are shown in Fig. 16(a). Notice that with the disconnection of the three lines between buses 505 and 506, the rms voltage at Terminal 1 (closer to the fault point) reaches a new operative point (around 460 kV) in relation to the pre-fault value (482 kV). This is partly due to the higher losses now present in the remaining line. In order to ensure a proper operation of the converters, this voltage drop will have to be compensated by the transformer tap changer.

The oscillation and subsequent damping of the AC power (in around 3.0s) at Terminal 1 (P1) is shown in Fig. 16(b). Low oscillatory patterns of the voltages in the negative pole (peak values equal to Vdc_neg1 = 340 kV and Vdc_neg3 = 348 kV) were also observed for this fault (Fig. 17a and 17b, respectively). In conclusion, no significant threats over the DC link from faults occurring in a nearby ac system were found.

Conclusions

From the study conducted on the overvoltage condition due to faults regarding the operation of a point-to-point bipolar system, the following conclusions can be drawn:

– Pole-to-ground faults on the DC link of a symmetrical monopole system may give rise to high overvoltages on the DC side. This condition should be considered while developing a VSC-HVDC project.

– Sustained overvoltages due to prolonged faults on the DC link can threaten, and can even destroy, the normal operation of the surge arresters installed on the DC side. Therefore, it is recommended that the DC line be opened soon after (in this case in less than 5 ms) the occurrence of the fault.

– Overvoltages due to the faults occurring inside neighbouring AC grids may have an impact on the VSCHVDC system. Therefore, to evaluate the performance of the DC link, it is also important to determine the magnitude of such overvoltages. Regarding the analysed network, those overvoltages exhibited low non-jeopardizing values.

Appendix

All AC sources were represented as sources behind the source impedance.
AC source at Terminal 1 (Fig.2):
Vrms : 380 kV
Rseries : 0.15335 Ω
Rparallel : 100 MΩ
Lparallel : 15.31 mH

AC source at Terminal 2:
Vrms : 145 kV
Rseries : 0.0222 Ω
Rparallel : 100 MΩ
Lparallel : 2.23 mH

Characteristics of the surge arrester:

Table A1. V-I curve of the surge arrester used.

.

All data used in this article can be willingly sent upon request.

Cable dimensions:

Fig. A1. Composition and dimensions of the cable used at the positive and negative pole.

OHTL dimensions:

Fig. A2. Dimensions of the DC overhead line.

REFERENCES

[1] Lu W. and Ooi B., DC Overvoltage Control during Loss of Converter in Multiterminal Voltage Sourced Converter Based HVDC (M-VSC-HVDC), IEEE Transactions on Power Delivery, Vol. 18, No. 3, pp. 915–920, July 2003.
[2] García Alonso J. C., Mosallat F., Wachal R., Abdel-Hadi K., Half and Full Bridge MMC Fault Performance in VSC-HVDC Systems. In: CIGRE CE B4 Colloquium: HVDC and Power Electronics to Boost Network Performance, Brasilia, Oct. 2-3,pp. 1-7.
[3] Lesnicar A., Marquardt R., An Innovative Modular Multi-Level Converter Topology for a Wide Power Range. IEEE Power Tech Conference, Bologna, Italy, June 2003.
[4] Peralta J., Saad H., Dennetiére S., Mahseredjian J., Nguefeu S., Detailed and Averaged Models for a 401-Level MMC–HVDC System. IEEE Transactions on Power Delivery, Vol. 27, No. 3, JUL. 2012. pp. 1501-1508. DOI: 10.1109/TPWRD.2012.2188911
[5] Saad H., Dennetiere S., Mahseredjian J., Delarue P., Guillaud X., Peralta J., Nguefeu S., Modular Multilevel Converter Models for Electromagnetic Transients. IEEE Transactions on Power Delivery, Vol. PP, No. 99, 2013. DOI. 10.1109/TPWRD.2013.2285633
[6] Saeedifard M., Iravani R., Dynamic Performance of a Modular Multilevel Back-to-Back HVDC System. IEEE Transactions on Power Delivery, Vol. 25, No. 4, 2010. pp. 2903-2912.
[7] Candelaria J., Park J-D., VSC-HVDC System Protection: A Review of Current Methods. In: Proc. of the IEEE/PES Power Systems Conf. and Exp. (PSCE), Phoenix (AZ), 20-23 Mar.pp.1-7. DOI: 10.1109/PSCE.2011.5772604
[8] Gu Y., Huang X., Qiu P., Hua W., Study of Overvoltage Protection and Insulation Coordination for MMC based HVDC. In: Proc. of the 2nd Int. Conf. on Computer Science and Electronics Engineering (ICCSEE 2013), Atlantis Press, Paris, pp. 2369-2372.
[9] Tang L., Ooi B-T., Protection of VSC-Multi-Terminal HVDC against DC Faults. In: Proc. of the IEEE 33rd Annual Power Electronics Specialists Conf (PESC 02), Vol. 2, Cairns, Australia, Jun. 23-27, 2002. pp. 719-724. DOI: 10.1109/PSEC.2002.1022539.
[10] Yang J., Zheng J., Tang G., He Z., Characteristics and Recovery Performance of VSC-HVDC DC Transmission Line Fault. In: Asia Pacific Power and Energy Engineering Conference (APPEEC), Chengdu, pp. 1–4, April 2010.
[11] PSCAD/EMTDC® Program. Manitoba HVDC Research Centre, v.4.3.1.0 (x4), 2010.
[12] EMTP-RV, Electromagnetic Transients Program, v. 2.6, 2013.
[13] CIGRE B2/B4/C1.17, Impacts of HVDC Lines on the Economics of HVDC Projects. No. 388, Aug. 2009.

Authors: prof. dr. Jose Antonio Jardini, Polytechnic School of the University of Sao Paulo, E-mail: jose.jardini@gmail.com; dr. Ricardo Leon Vasquez-Arnez, Foundation for the Technological Development of the Engineering Sciences (FDTE), E-mail: ricardoleon00@yahoo.co.uk; msc. Marcos T. Bassini, Polytechnic School of the University of Sao Paulo, E-mail: mtbassini@gmail.com; msc. Marco Antonio Barbosa Horita, Polytechnic School of the University of Sao Paulo, E-mail: marcoabh@gmail.com; msc. Gerson Yukio Saiki, Foundation for the Technological Development of the Engineering Sciences (FDTE), E-mail: gsaiki@fdte.org.br; Marcos Rodolfo Cavalheiro, CTEEP Sao Paulo State Power Transmission Company, E-mail: mcavalheiro@cteep.com.br.
The correspondence address is: e-mail: ricardoleon00@yahoo.co.uk

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 91 NR 8/2015. doi:10.15199/48.2015.08.26

Circuit Overcurrent Protection

Published by Clive Kimblin, International Association of Electrical Inspectors (IAEI) Magazine, Electrical Safety – Circuit Overcurrent Protection, March 16, 2002

Overcurrent devices protect the circuit conductors and conductor insulation from overheating. They also limit the damage associated with overheating and faults in downstream equipment. Fuses performed this function during the first days of electrical distribution, but circuit breakers of ever increasing sophistication have been available since the early 1900s. This paper focuses on circuit breakers and describes the wide variety of available devices. The emphasis is on low-voltage residential, industrial and commercial applications where the circuit voltages range from 120 volts through 600 volts. This is the area that is commonly encountered by electrical inspectors. The residential area is principally served by single and two-pole molded-case circuit breakers while the industrial and commercial world is principally served by higher current single and two-pole breakers and by three-pole molded-case circuit breakers. The paper also discusses low-voltage power circuit breakers and medium-voltage vacuum circuit breakers. The paper concludes with a brief description of the protective functions available with electronic trip devices. These include trip units with adjustable settings and with integrated ground-fault detection, and circuit breakers with communication capabilities to remote monitors including via the Internet. Electronics also bring additional safety features to the industrial and commercial area such as zone-interlocking between molded-case circuit breakers and power breakers, and additional safety features to the residential area such as ground-fault circuit interrupters and arc-fault circuit interrupters combined with residential circuit breakers.

Introduction
Figure 1. An interior view of a typical single-pole miniature circuit breaker

Since the subject of this paper is overcurrent protection, the paper first addresses the meanings of the words “overload” and “overcurrent.” There is also a brief description of the method used for circuit interruption; namely, arc extinction within the circuit breaker. The paper then focuses on residential circuit breakers, sometimes referred to as miniature circuit breakers, followed by a description of industrial/commercial molded-case circuit breakers. This includes a discussion of molded-case circuit breaker maintenance. There is a subsequent description of low-voltage power circuit breakers, and of medium-voltage vacuum circuit breakers. The paper concludes with a discussion of the role of electronics in circuit protection, including reference to the latest developments in communications capability via the Internet, and mention of new safety devices such as arc-fault circuit interrupters.

Overcurrent Protection and Arc Interruption
Figure 2. The typical continuous current range is 15 A – 225 A and the typical short-circuit-current ratings are 10 k -42 k A. Representative residential circuit breakers appear

All circuit breakers have the primary function of protecting the circuit conductors by detecting and interrupting overcurrents. Such faults can involve relatively modest currents such as overloads, or the large short-circuit overcurrents associated with faults between conductors. The definitions of terms from the National Electrical Code1 are as follows:

Overcurrent. Any current in excess of the rated current of equipment or the ampacity of a conductor. It may result from overload, short circuit, or ground fault.

Overload. Operation of equipment in excess of normal, full-load rating, or of a conductor in excess of rated ampacity that, when it persists for a sufficient length of time, would cause damage or dangerous overheating. A fault, such as a short circuit or ground fault, is not an overload.

The IEC definition of overload2 provides additional clarity:

Overload. Operating conditions in an electrically undamaged circuit, which cause an overcurrent.

Figure 3. An outline of a three-phase molded-case circuit breaker

All circuit breakers interrupt current by separating current-carrying contacts. An electrical arc is established at the last point of contact and, in an alternating-current circuit, this arc conducts the circuit current until the current wave passes through zero. The arc plasma is comprised of ionized ambient material, for example, air and metal vapor, and has temperatures exceeding 5000 degrees C. The arc continuously receives power from the circuit as measured by the arc voltage multiplied by the arc current, and continually loses power due to thermal-conduction, radiation, and convection. At current zero, the power input is removed, the contact-polarity is reversed, and there is a race between the factors that tend to cool the arc plasma and the factors such as the circuit voltage that tend to cause arc reignition. It must be noted that controlled arcs, within overcurrent devices, perform an extremely useful function. If arc creation did not occur during contact separation, the circuit current would collapse to zero instantaneously causing high overvoltages in inductive elements such as motors and transformers. Circuit breakers are designed to interrupt short-circuit fault currents ranging from 10 kA in residential branch circuits through 200 kA in industrial and commercial circuits, and much of this design involves control of the associated high current arc within the circuit breaker, with arc extinction and circuit interruption at current zero.

Residential Molded-Case Circuit Breakers

Residential overcurrent protective devices such as miniature circuit breakers are designed to protect circuit conductors by opening automatically before conductor damage is caused by excessive I-squared-t ohmic heating. Protection is achieved by having the response curve of the breaker, the time-current curve, lower than the corresponding thermal damage characteristics for the conductor. In common with most circuit breakers, a charged spring causes the contacts to separate when the mechanism is tripped.

Figure 4. A range of industrial/commercial molded-case circuit breakers

Residential circuit breakers are thermal-magnetic devices. For low-current overloads, the breaker trips due to the heating of an internal bimetal. For short-circuit-fault currents, the circuit breaker must respond more rapidly, and the breaker trips “instantaneously” due to internal magnetic forces. An interior view of a typical single-pole miniature circuit breaker appears in figure 1.

Thermal trip action is achieved through the use of a bimetal heated by the load current. A bimetal consists of two strips of metal bonded together. Each strip has a different thermal rate of heat expansion. Heating of the bimetal due to overload current will cause the bimetal to bend or deflect. The metal having the greater rate of expansion will be on the outside of the bend curve. On a sustained overload, the deflecting bimetal will physically push the trip bar, causing the operating mechanism to unlatch. The time needed for the bimetal to bend, to unlatch the mechanism and to trip the circuit breaker, varies inversely with the current.

Magnetic trip action is achieved through the magnetic forces associated with high short-circuit-fault currents. An armature moves in response to these forces, unlatches the mechanism and causes the breaker to trip.

Residential circuit breakers are qualified to UL 4893 and are a general category4 that includes single and two-pole circuit breakers with continuous currents of 225 A or less, and with voltage ratings of 120 V, 127 V, 120/240 V. These breakers may also be used in industrial/commercial applications. The typical continuous current range is 15 A–225 A and the typical short-circuit-current ratings are 10 kA–42 k A. Representative residential circuit breakers appear in figure 2.

Industrial/Commercial Molded-Case Circuit Breakers

An outline of a three-phase molded-case circuit breaker is shown in figure 3. The function of the molded-case (frame) is to provide an insulated housing to mount all of the components. The operating mechanism opens and closes the three sets of contacts simultaneously (common-trip), and is driven by a spring-loaded mechanism. The springs are charged by moving the handle first to the “off” position and then to the “on” position. The motion of a trip-bar trips the mechanism and, in a thermal-magnetic breaker, this motion is again initiated either by a bimetal or by a magnetic trip. In figure 3, each pole contains an electromagnet whose winding is in series with the load current. When a short-circuit fault occurs, the current passing through the circuit conductor causes the magnetic-field-strength of the electromagnet in the breaker to rapidly increase and attract the armature. As the armature is attracted to the electromagnet, the armature rotates the trip bar causing the mechanism to unlatch and the circuit breaker to trip.

In each pole there is typically one stationary and one moveable contact. On tripping, an arc is drawn between the separating contacts, and this arc is then controlled by the arc extinguishers (arc chutes) with interruption at current zero. Industrial/commercial molded-case circuit breakers can be equipped with many features. The larger frame sizes are frequently equipped with electronic trip units to permit greater control of the time-current tripping characteristics. This permits, for example, precise coordination between overcurrent devices connected in series. Electronic trip units can also be designed to detect ground faults and earth leakage currents, and units equipped with communication capabilities can send circuit breaker status information, and additional information such as circuit energy consumption, to distant monitors or remote data acquisition systems.

All three-pole circuit breakers, and one and two-pole breakers with ampere ratings over 225 A and voltage ratings above 240 V, are usually categorized4 as industrial/commercial circuit breakers. These circuit breakers are also qualified to UL 4893. An important sub-category is the current limiting circuit breaker that is designed to cause an extremely rapid build-up of arc voltage. These breakers4, when operating within their current-limiting range, limit the let-through I-squared-t to a value less than the I-squared-t of a half-cycle wave of the symmetrical prospective short-circuit-fault current.

Typical industrial/commercial circuit breakers frame sizes range from 125 A– 3000 A, typical voltages range from 120 V – 600 V and typical short-circuit-current ratings range from 10 kA – 200 kA. Figure 4 shows a range of industrial/commercial molded-case circuit breakers.

Three major features of circuit breakers are 1) they are common trip and consequently isolate all phases of the circuit, 2) they can be provided with electronic enhancements, and 3) they can be reset multiple times without replacement. Since molded-case circuit breakers are not intended to be opened for examination and maintenance, the life and maintenance of these re-settable circuit breakers will be addressed.

All overcurrent protective devices need to be maintained. They need to be housed in an appropriate environment, and their condition needs to be checked periodically. In particular, when an overcurrent protective device operates automatically, good practice dictates that the source of the overcurrent should be located, and that the condition of the overcurrent protective device should be checked prior to circuit re-energization. The specific requirements for molded-case circuit breaker maintenance and the associated considerations for circuit breaker life are as follows.

Figure 5. Representative low-voltage power circuit breaker frames

When appropriately maintained, molded-case circuit breakers provide reliable protection for many years. The exact lifetime of the breaker, however, is determined by the circuit breaker’s operational duty and by its environment. With respect to operational duty, for most circuits there will be occasional overload conditions or low-current fault conditions. Here the operating life will be tens of years. In other circuits, there will be occasional high short-circuit-current faults. This will reduce the circuit breaker’s operating life and may necessitate circuit breaker replacement. Here it is noted that molded-case circuit breakers, when evaluated according to the standard UL 489 “Molded-case Circuit Breakers, Molded-case Switches, and Circuit Breaker Enclosures”3, are subjected to bolted fault conditions at maximum short-circuit-current rating in two separate tests. Thus circuit breakers have a finite interrupting capability, and breakers that experience multiple high short-circuit-current faults should receive a thorough inspection with replacement if necessary.

With respect to environmental effects, circuit breakers are sometimes exposed to high ambient temperatures, to high humidity and to other ambient conditions that are hostile to long performance. For example, industries may have corrosive environments or could be associated with dusty environments that could affect operating parts.

It is not intended that molded-case circuit breakers be disassembled for inspection. However, the condition of molded-case circuit breakers can be evaluated by using NEMA AB4 “Guidelines for Inspection and Preventive Maintenance of Molded-case Circuit Breakers used in Industrial and Commercial Applications.”5

This document should be referenced during periodic maintenance or during specific inspection following a high short-circuit-current fault. The document is intended to ensure that molded-case circuit breakers are well maintained, and provides guidelines for circuit breaker replacement.

NEMA AB4 is divided into separate sections dealing with:

• Inspection Procedures
• Preventive Maintenance
• Test Procedures
• Accessory Device Test Procedures

The section dealing with Inspection Procedures describes thermal and visual checks of the circuit breaker’s condition. Overheating of the circuit breaker would necessitate further investigation, and cracks in the molded case would certainly necessitate circuit breaker replacement.

The section dealing with Preventive Maintenance insures that the circuit breaker’s life is not compromised by external conditions. The objectives are that the circuit breaker operates in a clean environment, that the connections at the terminals are torqued properly and are in good condition, and that the circuit breaker is correctly wired.

Figure 6. As indicated in the cut-away view of figure 6, two contacts oppose each other within a vacuum envelope

The section dealing with Test Procedures deals with non-destructive tests that can be used to verify specific operating characteristics of molded-case circuit breakers: Mechanical Operation Test, Insulation Resistance Test, Individual Pole Resistance Test (millivolt drop test), Inverse Time Overcurrent Test, Instantaneous Overcurrent Trip Test, and Rated Hold-In Test. Non- compliance of one or more of these tests could lead to circuit breaker replacement.

In summary, following an automatic overcurrent interruption, the condition of any protective device should be checked prior to circuit re-energization. For molded-case circuit breakers, the condition of the circuit breaker is assessed without opening or disassembling the breaker. For tripping events caused by overloads and low-current faults, evaluation usually takes the form of visual inspection and mechanical operation. However, circuit breakers that have experienced multiple high short-circuit-current faults, as evidenced by conditions at the source of the faults, should receive a thorough inspection per the guidelines of NEMA AB4. This document should also be used for recommended, periodic, preventive maintenance.

Power Circuit Breakers

As a broad generalization, molded-case circuit breakers are applied downstream from low-voltage power circuit breakers, and are designed to be connected to circuits comprised of insulated wires and insulated cables rather than bare bus bars. As previously mentioned, a primary function of these molded-case circuit breakers is to protect the conductor and the conductor insulation, and the tests in the UL 489 standard consequently incorporate wire in the testing procedures. By contrast, low-voltage power circuit breakers are typically connected by buswork within switchgear. The ANSI standards6 therefore incorporate bus bar conductors into the testing procedures. Another general distinction is that the upstream low-voltage power circuit breakers typically have a “short-time-rating” current capability that permits these circuit breakers to remain closed during fault-clearance by a downstream circuit breaker. This optimizes the availability of power to the parallel downstream circuits protected by the single upstream breaker.

For power circuit breakers, typical continuous current ranges are 800 A–5000 A, typical voltage ranges are 240 V–600 V and typical short-circuit-current ranges are 40 kA–100 kA. Figure 5 shows representative low-voltage power circuit breaker frames.

The main differences between low-voltage power circuit breakers and molded-case circuit breakers are as follows7:

Low-voltage power circuit breakers are evaluated for “short-time duty cycle” tests. This test demonstrates that the low-voltage power circuit breaker can remain closed (or “hold-in”) for at least 0.5 seconds while the downstream (feeder) breaker has the opportunity to clear the fault. Further, the main circuit breaker must continue to “hold in” in the event that the downstream breaker subsequently re-closes, the fault is still present, and the downstream breaker has to re-open in order to isolate the fault.

Low-voltage power circuit breakers are also evaluated for a “short-circuit current duty cycle” test. The test demonstrates that a main low-voltage power circuit breaker can remain closed for at least 0.5 seconds while a downstream (feeder) circuit breaker has the opportunity to clear the fault, but if the fault current persists, the main circuit breaker must open and interrupt. Again, the continued closure of the main circuit breaker maintains power continuity to the unaffected downstream circuits and optimizes coordination.

Figure 7. Typical medium-voltage circuit breakers using vacuum interrupter technology

Low-voltage power circuit breakers are equipped with stored energy mechanisms. This permits sequences of contact opening, contact re-closure, and contact re-opening, which can be activated either remotely or locally.

Low-voltage power circuit breakers can be serviced and maintained. This is important for applications where circuit breaker replacement is inconvenient and extended life is important. Further, these breakers are used primarily in draw-out switchgear. Thus, low voltage power circuit breakers typically are designed with rear-mounted primary disconnect contacts to permit the circuit breaker to be connected to and disconnected from the primary circuit stabs in the switchgear.

Low-voltage power circuit breakers receive single pole tests at 87 percent of rated interruption current at line-to-line voltage. This reflects the possibility of high single-pole fault currents in the upstream circuit location. In particular, such circuit breakers are suitable for corner-grounded delta-connected transformer applications.

ULs 1066 and UL 489 both cover similar continuous current ranges. However, since low-voltage power circuit breakers are applied upstream from molded-case circuit breakers, and since they typically supply several parallel downstream circuits, low-voltage power circuit breakers are usually high-continuous-current devices.

Low-voltage power circuit breakers are all rated to carry 100 percent of their rated continuous current within the switchgear. For molded-case circuit breakers in enclosures, the maximum circuit current is 80 percent of rated current, although 100 percent rated circuit breakers are available.

Medium-Voltage Vacuum Circuit Breakers

In low-voltage circuits, the magnitude of the circuit fault current is limited by the voltage developed across the arc drawn between the separating contacts. This arc voltage of tens or possibly hundreds of volts can approach the circuit voltage with resulting current limitation. However, in medium voltage circuits of 2.3 kV through 38 kV, the arc voltage is small compared to the circuit voltage, and the circuit breaker experiences the full available fault current.

The technology of choice for medium-voltage overcurrent protection is vacuum. Here each pole of the 3-phase circuit breaker contains a vacuum interrupter of deceptively simple construction. As indicated in the cut-away view of figure 6, two contacts oppose each other within a vacuum envelope.

During overcurrent conditions, the current-carrying contacts are separated and an arc is established in the metal-vapor evaporated from local hot spots that develop on the contacts. The circuit current is conducted through the arc plasma formed from the ionized metal vapor. During current flow there is continual evaporation from the local hot spots on the contacts with continual condensation of the ionized metal vapor on the broader contact surfaces and on the vapor condensation shield. At current zero the power input to the arc is removed and the evaporation ceases. However, the loss of inter-contact ionized vapor continues and a vacuum condition is re-established. Further, the contact polarity reverses. This results in a rapid change of the inter-contact region from an electrical conductor to an insulator within microseconds of current zero.

The keys to vacuum interrupter design are the selection and creation of the contact material, the design of the contacts for arc control, and the creation of a vacuum envelope that maintains a high vacuum condition for tens of years. Figure 7 shows typical medium-voltage circuit breakers using vacuum interrupter technology.

Electronics in Circuit Protection

Thermal-magnetic trip units are economic and compact. They have been used effectively and efficiently for many years. Their function can also be performed by electronic trip units. The first use of electronics, in the 1960s, was associated with protective relays for medium voltage circuit breakers. Since the early 1970s electronic trip units have been increasingly applied to power circuit breakers and the larger frame sizes of industrial/commercial molded-case circuit breakers. The present thrust is to make electronic trip units available down to frame sizes of 250 A and below. The advantage of electronic trip units is that time-current curves can be readily adjusted; both for phase current settings and for the settings of integrated ground fault units4. This flexibility permits coordination between series connected overcurrent protective devices such that, under fault conditions, only the device immediately upstream from the fault will clear the circuit. A further advantage is that the trip characteristic is independent of ambient temperature.

The electronic circuits on circuit breakers may also incorporate communications capability. At first this was limited to applications such as zone-selective-interlocking. Here an upstream power circuit breaker is set to trip with no intentional delay, but a trip-restraint-signal from a downstream circuit breaker can cause the power breaker to remain closed for settings up to 0.5 seconds, the maximum short-time duration. When a fault occurs on the load side of a selectively coordinated downstream circuit breaker, this downstream breaker communicates that a fault has been detected and the upstream power breaker then permits the downstream breaker to interrupt the fault. However, if the fault occurs between the power circuit breaker and the downstream circuit breaker, no restraint signal is received from the downstream circuit breaker, and the power circuit breaker will clear the fault without any intentional delay.

Communications capability is now used8 to transmit data to remote monitors or data acquisition systems. The initial information was limited to open/close status. This was followed by information on “cause-of-trip” and most recently by electrical metering data and by complete power quality data. In fact, it is now possible, from a remote location, to monitor and diagnose the electrical situation in a total industrial plant based on information communicated via the Internet.

Electronic advances have also increased the safety protection of residential circuit breakers. Ground-fault circuit interrupters have been available for many years9 and these circuit breakers, in addition to protecting the branch circuit wiring against overcurrents, provide personnel protection against electrical shock in cords and equipment connected to the outlets. Arc-fault circuit interrupters10 have been introduced during the past five years. These devices recognize the specific characteristics of arcing faults, and then interrupt the circuit. When combined with residential circuit breakers and located at the origin of the branch circuit, these AFCIs mitigate the effects of electrical arcs in the branch circuit wiring and in the cords connected to the outlets. Residential circuit breakers with combined GFCI/AFCI protection are also available.

Summary

Circuit breakers protect the circuit conductors against overcurrent. This is accomplished by first detecting the overcurrent and then interrupting the overcurrent with subsequent isolation. Thermal-magnetic trip units or electronic trip units detect the overcurrent. Interruption and isolation is accomplished by drawing an arc between separating contacts, with subsequent arc extinction. Circuit breakers can be broadly classified as low-voltage residential molded-case circuit breakers, low- voltage industrial/commercial molded-case circuit breakers, low-voltage power circuit breakers and medium-voltage circuit breakers. An overall electrical distribution system can be expected to incorporate all classes of circuit breaker. Electronics has increased the sophistication of trip units, including communications capability, and has permitted additional safety features such as shock-protection through GFCIs and enhanced fire protection through AFCIs.

References

1.NFPA 70,National Electrical Code2002, Article 100, (National Fire Protection Association, Quincy, MA, 2002), p. 70-37.
2 “International Standard for Low Voltage Switchgear and Controlgear, Part 1: General Rules,” International ElectroTechnical Commission Standard IEC 60947-1, Third Edition, 1999-02.
3 “UL Standard for Safety for Molded-case Circuit Breakers, Molded-case Switches, and Circuit Breaker Enclosures,” UL-489, (Underwriters Laboratories, Ninth Edition), October 31, 1996.
4 “Molded-case Circuit Breakers and their Application,” NEMA AB3-2001, (National Electrical Manufacturers Association).
5 “Guidelines for Inspection and Preventive Maintenance of Molded-case Circuit Breakers used in Industrial and Commercial Applications,” NEMA AB4-2000, (National Electrical Manufacturers Association). Recognized as an American National Standard (ANSI).
6 “UL Standard for Safety for Low-Voltage AC and DC Power Breakers Used in Enclosures,” UL 1066, (Underwriters Laboratories, Third Edition), May 30, 1997. Recognized as an American National Standard (ANSI).
7 Kimblin, C.W. and Long, R.W., “Comparing Test Requirements for Low Voltage Circuit Breakers,” IEEE Industry Applications Magazine, January/February 2000, p. 45-52.
8 Engel, J.C., Murphy, W.D., and Oravetz, D.M., “Remote Monitoring of Circuit Breakers,” Conference Record of the 1999 IEEE Industry Applications Conference, Phoenix, Arizona, October 1999, p.2344-2347
9 “Overcurrents and Undercurrents – All about GFCIs and AFCIs,” Earl W. Roberts, (Reptec, Mystic, CT), 2000.
10 Kimblin, C.W., Engel, J.C., and Clarey, R.J., “Arc-Fault Circuit Interrupters, The New Residential Electrical-Safety Technology,” IAEI News, Volume 72, Number 4, July/August 2000, p. 26-31.

Author: Clive W. Kimblin is the manager, Applications & Standards for the Electrical Distribution Products Operations of Cutler-Hammer, He obtained a B.Sc (Physerations of Cutler-Hammer. He obtained a B.Sc (Physics) and Ph.D. (Electrical Engineering) from Liverpool University, England and an MSIE (Engineering Management) from the University of Pittsburgh. Prior to his current position, he worked at the Westinghouse Research and Development Center in Pittsburgh and at Holec/Begemann in The Netherlands. He is active within NEMA and is an IEEE Fellow.

Source URL: https://iaeimagazine.org/2002/2002march/circuit-overcurrent-protection/

Design of the Electrical Installation Protection System

Published by Electrical Installation Wiki, Chapter J. Overvoltage protection – Design of the electrical installation protection system

To protect an electrical installation in a building, simple rules apply for the choice of

• SPD(s);
• its protection system.

For a power distribution system, the main characteristics used to define the lightning protection system and select a SPD to protect an electrical installation in a building are:

• SPD
– quantity of SPD
– type
– level of exposure to define the SPD’s maximum discharge current Imax.

• Short circuit protection device
– maximum discharge current Imax;
– short-circuit current Isc at the point of installation.

The logic diagram in the Figure J20 below illustrates this design rule.

Fig. J20 – Logic diagram for selection of a protection system

The other characteristics for selection of a SPD are predefined for an electrical installation.

• number of poles in SPD;
• voltage protection level Up;
• operating voltage Uc.

This sub-section Design of the electrical installation protection system describes in greater detail the criteria for selection of the protection system according to the characteristics of the installation, the equipment to be protected and the environment.

Elements of the protection system

A SPD must always be installed at the origin of the electrical installation.

Location and type of SPD

The type of SPD to be installed at the origin of the installation depends on whether or not a lightning protection system is present. If the building is fitted with a lightning protection system (as per IEC 62305), a Type 1 SPD should be installed.

For SPD installed at the incoming end of the installation, the IEC 60364 installation standards lay down minimum values for the following 2 characteristics:

• Nominal discharge current In = 5 kA (8/20) µs;
• Voltage protection level Up (at In) < 2.5 kV.

The number of additional SPDs to be installed is determined by:

• the size of the site and the difficulty of installing bonding conductors. On large sites, it is essential to install a SPD at the incoming end of each subdistribution enclosure.

• the distance separating sensitive loads to be protected from the incoming end protection device. When the loads are located more than 10 meters away from the incoming-end protection device, it is necessary to provide for additional fine protection as close as possible to sensitive loads. The phenomena of wave reflection is increasing from 10 meters see Propagation of a lightning wave

• the risk of exposure. In the case of a very exposed site, the incoming-end SPD cannot ensure both a high flow of lightning current and a sufficiently low voltage protection level. In particular, a Type 1 SPD is generally accompanied by a Type 2 SPD.

The table in Figure J21 below shows the quantity and type of SPD to be set up on the basis of the two factors defined above.

Fig. J21 – The 4 cases of SPD implementation

Protection distributed levels

Several protection levels of SPD allows the energy to be distributed among several SPDs, as shown in Figure J22 in which the three types of SPD are provided for:

• Type 1: when the building is fitted with a lightning protection system and located at the incoming end of the installation, it absorbs a very large quantity of energy;
• Type 2: absorbs residual overvoltages;
• Type 3: provides “fine” protection if necessary for the most sensitive equipment located very close to the loads.

Fig. J22 – Fine protection architecture

Note: The Type 1 and 2 SPD can be combined in a single SPD

Common characteristics of SPDs according to the installation characteristics

Operating voltage Uc

Depending on the system earthing arrangement, the maximum continuous operating voltage Uc of SPD must be equal to or greater than the values shown in the table in Figure J23.

Fig. J23 – Stipulated minimum value of Uc for SPDs depending on the system earthing arrangement (based on Table 534.2 of the IEC 60364-5-53 standard)

N/A: not applicable
U: line-to-line voltage of the low-voltage system

a.^1,2 these values are related to worst-case fault conditions, therefore the tolerance of 10 % is not taken into account.

The most common values of Uc chosen according to the system earthing arrangement.
TT, TN: 260, 320, 340, 350 V
IT: 440, 460 V

Voltage protection level Up (at In)

The IEC 60364-4-44 standard helps with the choice of the protection level Up for the SPD in function of the loads to be protected. The table of Figure J24 indicates the impulse withstand capability of each kind of equipment.

Fig. J24 – Required rated impulse voltage of equipment Uw (table 443.2 of IEC 60364-4-44)

a.^ According to IEC 60038:2009.
b.^ This rated impulse voltage is applied between live conductors and PE.
c.^1,2 In Canada and USA, for voltages to earth higher than 300 V, the rated impulse voltage corresponding to the next highest voltage in this column applies.
d.^ For IT systems operations at 220-240 V, the 230/400 row shall be used, due to the voltage to earth at the earth fault on one line.

Fig. J25 – Overvoltage category of equipment

The “installed” Up performance should be compared with the impulse withstand capability of the loads.

SPD has a voltage protection level Up that is intrinsic, i.e. defined and tested independently of its installation. In practice, for the choice of Up performance of a SPD, a safety margin must be taken to allow for the overvoltages inherent in the installation of the SPD (see Figure J26 and Connection of Surge Protection Device).

Fig. J26 – Installed Up

The “installed” voltage protection level Up generally adopted to protect sensitive equipment in 230/400 V electrical installations is 2.5 kV (overvoltage category II, see Fig. J27).

Note: If the stipulated voltage protection level cannot be achieved by the incoming-end SPD or if sensitive equipment items are remote (see Elements of the protection system#Location and type of SPD Location and type of SPD , additional coordinated SPD must be installed to achieve the required protection level.

Number of poles

• Depending on the system earthing arrangement, it is necessary to provide for a SPD architecture ensuring protection in common mode (CM) and differential mode (DM).

Fig. J27 – Protection need according to the system earthing arrangement

a.^ The protection between phase and neutral can either be incorporated in the SPD placed at the origin of the installation, or be remoted close to the equipment to be protected
b.^ If neutral distributed

Note:

• Common-mode overvoltage
A basic form of protection is to install a SPD in common mode between phases and the PE (or PEN) conductor, whatever the type of system earthing arrangement used.

• Differential-mode overvoltage
In the TT and TN-S systems, earthing of the neutral results in an asymmetry due to earth impedances which leads to the appearance of differential-mode voltages, even though the overvoltage induced by a lightning stroke is common-mode.

2P, 3P and 4P SPDs
(see Fig. J28)

• These are adapted to the IT, TN-C, TN-C-S systems.
• They provide protection merely against common-mode overvoltages.

Fig. J28 – 1P, 2P, 3P, 4P SPDs

1P + N, 3P + N SPDs
(see Fig. J29)

• These are adapted to the TT and TN-S systems.
• They provide protection against common-mode and differential-mode overvoltages

Fig. J29 – 1P + N, 3P + N SPDs
Selection of a Type 1 SPD

Impulse current Iimp

• Where there are no national regulations or specific regulations for the type of building to be protected: the impulse current Iimp shall be at least 12.5 kA (10/350 µs wave) per branch in accordance with IEC 60364-5-534.

• Where regulations exist: standard IEC 62305-2 defines 4 levels: I, II, III and IV

The table in Figure J31 shows the different levels of Iimp in the regulatory case.

Fig. J30 – Basic example of balanced Iimp current distribution in 3 phase system
Fig. J31 – Table of Iimp values according to the building’s voltage protection level (based on IEC/EN 62305-2)

Autoextinguish follow current Ifi

This characteristic is applicable only for SPDs with spark gap technology. The autoextinguish follow current Ifi must always be greater than the prospective short-circuit current Isc at the point of installation.

Selection of a Type 2 SPD

The maximum discharge current Imax is defined according to the estimated exposure level relative to the building’s location.

The value of the maximum discharge current (Imax) is determined by a risk analysis (see table in Figure J32).

Fig. J32 – Recommended maximum discharge current Imax according to the exposure level
Selection of external Short Circuit Protection Device (SCPD)

• The protection devices (thermal and short circuit) must be coordinated with the SPD to ensure – reliable operation, i.e.
– ensure continuity of service:
– withstand lightning current waves
– not generate excessive residual voltage.
• ensure effective protection against all types of overcurrent:
– overload following thermal runaway of the varistor;
– short circuit of low intensity (impedant);
– short circuit of high intensity.

Risks to be avoided at end of life of the SPDs

Due to ageing

In the case of natural end of life due to ageing, protection is of the thermal type. SPD with varistors must have an internal disconnector which disables the SPD.

Note: End of life through thermal runaway does not concern SPD with gas discharge tube or encapsulated spark gap.

Due to a fault

The causes of end of life due to a short-circuit fault are:

• Maximum discharge capacity exceeded.
This fault results in a strong short circuit.
• A fault due to the distribution system (neutral/phase switchover, neutral disconnection).
• Gradual deterioration of the varistor.

The latter two faults result in an impedant short circuit.

The installation must be protected from damage resulting from these types of fault: the internal (thermal) disconnector defined above does not have time to warm up, hence to operate.

A special device called “external Short Circuit Protection Device (external SCPD)”, capable of eliminating the short circuit, should be installed. It can be implemented by a circuit breaker or fuse device.

Characteristics of the external SCPD

The external SCPD should be coordinated with the SPD. It is designed to meet the following two constraints:

Lightning current withstand

The lightning current withstand is an essential characteristic of the SPD’s external Short Circuit Protection Device.

The external SCPD must not trip upon 15 successive impulse currents at In.

Short-circuit current withstand

• The breaking capacity is determined by the installation rules (IEC 60364 standard):
The external SCPD should have a breaking capacity equal to or greater than the prospective short-circuit current Isc at the installation point (in accordance with the IEC 60364 standard).

• Protection of the installation against short circuits
In particular, the impedant short circuit dissipates a lot of energy and should be eliminated very quickly to prevent damage to the installation and to the SPD.

The right association between a SPD and its external SCPD must be given by the manufacturer.

Installation mode for the external SCPD

Device “in series”

The SCPD is described as “in series” (see Fig. J33) when the protection is performed by the general protection device of the network to be protected (for example, connection circuit breaker upstream of an installation).

Fig. J33 – SCPD “in series”

Device “in parallel”

The SCPD is described as “in parallel” (see Fig. J34) when the protection is performed specifically by a protection device associated with the SPD.

• The external SCPD is called a “disconnecting circuit breaker” if the function is performed by a circuit breaker.
• The disconnecting circuit breaker may or may not be integrated into the SPD.

Fig. J34 – SCPD “in parallel”

Note: In the case of a SPD with gas discharge tube or encapsulated spark gap, the SCPD allows the current to be cut immediately after use.

Guarantee of protection

The external SCPD should be coordinated with the SPD, and tested and guaranteed by the SPD manufacturer in accordance with the recommendations of the IEC 61643-11 standard. It should also be installed in accordance with the manufacturer’s recommendations. As an example, see the Schneider Electric SCPD+SPD coordination tables.

When this device is integrated, conformity with product standard IEC 61643-11 naturally ensures protection.

Fig. J35 – SPDs with external SCPD, non-integrated (iC60N + iPRD 40r) and integrated (iQuick PRD 40r)
Summary of external SCPDs characteristics

A detailed analysis of the characteristics is given in section Detailed characteristics of the external SCPD .

The table in Figure J36 shows, on an example, a summary of the characteristics according to the various types of external SCPD.

Fig. J36 – Characteristics of end-of-life protection of a Type 2 SPD according to the external SCPDs
SPD and protection device coordination table

The table in Figure J37 below shows the coordination of disconnecting circuit breakers (external SCPD) for Type 1 and 2 SPDs of the Schneider Electric brand for all levels of short-circuit currents.

Coordination between SPD and its disconnecting circuit breakers, indicated and guaranteed by Schneider Electric, ensures reliable protection (lightning wave withstand, reinforced protection of impedant short-circuit currents, etc.)

Fig. J37 – Example of coordination table between SPDs and their disconnecting circuit breakers (Schneider Electric brand). Always refer to the latest tables provided by manufacturers.
Coordination with upstream protection devices

Coordination with overcurrent protection devices

In an electrical installation, the external SCPD is an apparatus identical to the protection apparatus: this makes it possible to apply selectivity and cascading techniques for technical and economic optimization of the protection plan.

Coordination with residual current devices

If the SPD is installed downstream of an earth leakage protection device, the latter should be of the “si” or selective type with an immunity to pulse currents of at least 3 kA (8/20 μs current wave).

Source URL: https://www.electrical-installation.org/enwiki/Design_of_the_electrical_installation_protection_system

The Dangers of Stray Current Damage in Electrical Systems & Arc Flash Protection

Published by Jim Galloway, International Association of Electrical Inspectors (IAEI) Magazine, Electrical Fundamentals – The Dangers of Stray Current Damage in Electrical Systems & Arc Flash Protection, July 3, 2019

The new extension to the factory is finally finished. The engineers have produced a state-of-the-art design; the qualified electricians have completed a professional installation with quality hardware; and the thorough electrical inspections verifying that everything is up to Code are complete. Now that the power is on, the machinery installation crews are busy getting everything ready for production.

But suddenly there is trouble in the plant. Hundreds of amperes are flowing in an uncontrolled fashion, surging through metallic conduit and panels, melting EMT connectors, arcing the bonding wires in a 3-phase receptacle, and burning through the insulation of the power cord to the brand new, million-dollar machine from Europe. If it continues, soon a live phase conductor will be in contact with a red-hot grounding wire, and an unsuspecting worker could be seriously hurt or even killed when they investigate the problem.

What is going on? Did a something come loose? Could it be a lightning strike? Is the foreign-made machine incompatible with our electrical network? Is the damage being caused by 1st, 3rd, 5th, 7th, or 9th level harmonics? Why hasn’t the over-current protection system reacted and tripped a breaker? There’s a very good chance that the problem has something to do with the common arc welding process being used to attach mechanical hardware as part of the new machine installation. But how?

Stray welding current

A stray current is a flow of electric current through unintended conductors such as building structures, electrical grounding or bonding conductors, or other equipment due to electrical supply system imbalances or improper equipment hookup. Often in industrial or construction environments, this trouble occurs due to a very simple error in setup by a welder. In Ontario, Canada, there has recently been a well-documented case of an electrical explosion and at least one fatal electrocution due to damage in a facility’s electrical system where the root cause was identified as stray welding current (SWC).1-3

Arc welding machines are designed to provide alternating or direct current on the order of hundreds of amperes for industrial, commercial, and residential or hobbyist applications. These specialized power sources are designed to provide a low-resistance electrical circuit (typically less than 0.25 Ohms) on the secondary (or output) side; from an electrical engineer’s perspective, this is essentially a short-circuit condition.

This secondary welding circuit is intended to be completed as an isolated, closed-loop system with two cables (see Image 1). It consists of the following parts:

Electrode lead — is the secondary circuit conductor transmitting energy from the power source to the electrode holder, gun, or torch.

Workpiece lead — is the secondary circuit conductor that is attached to the workpiece by the return current clamp and completes the welding circuit (these devices are commonly and incorrectly referred to as the ground cable and the ground clamp).

Image 1. A correct arc welding setup.

The SWC faults can occur with simple setup errors or minor system faults, which can introduce current on the order of hundreds of amperes into building structures, electrical networks, and machinery. Two common examples of SWC are illustrated in Images 2 and 3. However, there are many other possibilities as the welding current will always find the path of least resistance to its source. Often, the SWC fault involves other machinery such as fabricating equipment, power tools, other welding machines, and cranes or hoists. These SWC faults can also originate from portable engine-driven or battery-powered welding equipment that do not even derive power from the grid, and they can damage equipment even when powered off.

Image 2. A stray welding current fault due to an operator error, which involves other electrically powered equipment.
Image 3. A stray welding current fault due to a minor welding equipment malfunction.

Surprisingly, there is currently no foolproof system available to stop these damaging stray current faults. They are not interrupted by the overcurrent protection devices installed even in modern electrical networks, and they are not detected or interrupted in circuits protected by ground-fault circuit interrupter (GFCI) devices. The focus has always been on measuring current flow in the intended conductors, and the assumption is that the grounding or bonding conductors would never see current levels that exceed the level where the overcurrent protection devices would trip.

The scenario described in the introduction should not be happening. However, it is far too common of an occurrence in any facility where welding has been performed to be ignored. An example of the sort of damage that can occur is shown in Image 4. Beyond the electrical systems damage, SWC can cause arcing and fires in unexpected locations in a facility, overheating of lifting chains or slings (leading to arcing or annealing), damage to machinery bearings, and arc strikes leading to undesirable metallurgical transformations in certain alloys. Cases of accelerated corrosion caused by stray direct current electrolysis is also a concern on marine structures and buried metallic infrastructure. Stray current damage, therefore, can also be considered a problem from an economic standpoint since much of the damage may be initially hidden from view—even before it becomes an immediate safety issue. Tens of thousands of dollars of damage can occur in a facility or to machinery and systems from one SWC event, which may not even be immediately detected.

Image 4. Electrical arcing and damage in a 600-V electrical outlet from stray welding current.1
What can be done?

Stray currents from welding operations can be avoided through the strict adherence to rules of the applicable welding safety standards (e.g., ANSI Z49.1:2012 or CSA W117.2-19). These practices—also spelled out in equipment manuals—include locating the workpiece lead return attachment point as close as practicable to the arc, using well-maintained welding cables of sufficient ampacity, and ensuring that the work return current clamp is firmly attached on an intentionally cleaned spot (free of mill-scale, paint, etc.)4 All professional welders should be following these rules; however, it must be noted that anyone can buy a welding machine capable of producing hundreds of amperes and misapply it.

Image 5. A screenshot from the video The Problem of Stray Welding Current.5

Conestoga College in Cambridge, Ontario, recently produced educational videos to explain the SWC problem; to show just how easy it is to inadvertently create these damaging fault conditions; and to demonstrate the serious damage that can occur in a machine’s power cord. These videos are now publicly available on the college’s YouTube channel. (See screenshots in Images 5 and 6). The goal of the project is to better inform welding educators, welders, and the welding industry, in general, of the SWC hazard. The electrical engineering and inspection community also needs to be more aware of the SWC problem, what to look for, and how it happens.

Image 6. A screenshot from the video Stray Welding Current Damage to Power Cords.6

This work was sponsored by EnerDynamic Systems Inc. (ESI) of Brantford, Ontario. ESI partnered with Conestoga College to assist them in extending a patent-pending design for a stray current interrupter device from renewable energy systems into arc welding applications. The partnership had the goal of developing an engineering control device that can make the damage from SWC a thing of the past.

References

1. D. Hisey, “Stray Current Goes to College,” Canadian Welding Association Journal, vol. 14, pp. 20-25, 2016.
2. Office of the Chief Coroner of Ontario, DOKIS, Kelly (Inquest), Kitchener, Ontario, Canada: Queen’s Printer for Ontario, June 16th–19th, 2003.
3. D. Hisey, “How to Prevent Stray Welding Current Damage in Your Electrical System,” 17 November 2017. [Online]. Available: https://www.ecmweb.com/shock-electrocution/how-prevent-stray-welding-current-damage-your-electrical-system. [Accessed 27 December 2018].
4. Canadian Standards Association, CAN/CSA-W117.2-19 – Safety in Welding, Cutting, and Allied Processes, Toronto: Canadian Standards Association, 2019.
5. Conestoga College & Enerdynamic Systems Inc., “The Problem of Stray Welding Current,” 24 January 2019. [Online]. Available: https://www.youtube.com/watch?v=80ehl2nDXUk. [Accessed 24 January 2019]
6. Conestoga College & Enerdynamic Systems Inc., “Stray Welding Current Damage to Power Cords,” 24 January 2019. [Online]. Available: https://www.youtube.com/watch?v=kIVH5V9ntrY. [Accessed 24 January 2019]

Source URL: https://iaeimagazine.org/2019/2019may/the-dangers-of-stray-current-damage-in-electrical-systems-arc-flash-protection/

Current-Carrying Capacity Parallel Single-Core LV Cable

Published by Lech BOROWIK1, Artur CYWIŃSKI2
Politechnika Częstochowska, Wydział Elektryczny (1), Pracownia Projektowa omega-projekt (2)

Abstract. The paper presents the problems related to the selection of parallel single-core low voltage cables in terms of their current carrying capacity and accordance with PN-IEC 60364-5-523 national standard, exemplified by electric power transmission from transformer MV/LV 1600kVA, with proximity effects also taken into consideration.

Streszczenie. Przedstawiony poniżej tekst opisuje problemy związane z doborem wielowiązkowych linii kablowych niskiego napięcia pod względem obciążalności długotrwałej zgodnie z normą PN-IEC 60364-5-523 na przykładzie wyprowadzenia mocy z transformatora SN/nN 1600kVA z uwzględnieniem wpływu zjawiska zbliżenia. (Obciążalność długotrwała wielowiązkowych linii kablowych nN).

Keywords: Multi-conductor parallel cables, current carrying capacity, proximity effects.
Słowa kluczowe: wielowiązkowe linie kablowe, obciążalność długotrwała, zjawisko zbliżenia.

Introduction

While designing networks and electrical installation, a problem that often appears in the practice is how to select properly low voltage cables, taking into account their current-carrying capacity. The main document that is used by the designers is the Polish national standard “Low-voltage electrical installations – Current-carrying capacity” PN-IEC 60364-5-523:2001 [1] (translation from IEC 60364- 5 part 52 International Standard). The above-given standard along with other publications [2],[3] refer in detail to the configurations of cables, distances between them and their surroundings. The standard introduces various reduction factors as well as other calculative tools that allow for the optimal selection of wires and cables and protection. It seems, however, that the authors of the above mentioned standard and relative studies, in their calculation or research into the mutual influence of parallel cables have focused principally on thermal phenomena. Thus they have neglected the electromagnetic field generated by the flow of currents, which are significant in particular for parallel core which forms one circuit. This situation requires an attempt to study the effects of uneven load in the parallel line and its effect on current-carrying capacity of the cable. In the present paper the authors focus on the description of the problem on a real-life example of selecting multi-core low voltage cable which is the power lead from 1600kVA transformer. The tests have been conducted in accordance with PN-IEC 60364-5-523:2001 Polish National Standard.

In the example analyzed here, it was necessary to select a low-voltage cable which was connecting a 1600kVA transformer to the main low-voltage switchboard. Due to the location of the transformer station and the building specifications, it was impossible to make bus-ducts connections. Financial limitation made it impossible to build insulated bus-ducts.

A cable line made from single core cables of the YAKXS 1X240 type, laid out on cable tracks, with the temperature kept below 200C has been selected. The distance between the cores of one phase was equal to the diameter of the line, whereas the subsequent phases were laid out analogically at distances considerably exceeding twice of the diameter of a single line.

The nominal transformer current on the low-voltage side, with the power factor cos φ =0,9 (system working with reactive power compensation) for a single phase, under the symmetrical load equals 2576A.

According to the Standard [1], the cables were laid out in accordance with the 52-C12 Table (reference method of installation – G horizontal) . While calculating the current-carrying capacity, the correction factor were taken into consideration – according to 52-D1 Table of coefficients and exponents, taking into account the ambient temperature different from 30oC (coefficient 1,08) – the increasing coefficient has been skipped, and in accordance with 52-E1 Table (reduction factor 0,79 for a group of more than one circuit). After meeting all the conditions, included in the 523.6 Chapter of Standard [1], the following electrical system has been designed.

Fig.1. Diagram of connecting low-voltage switchboard

For each phase, 6 parallel lines were selected, made from XPLE insulated aluminum cable with the cross-section of 240 mm2 of the YAKXS type. The current-carrying capacity of each core, according to 52-C12 Table of norm [1] was 521A. However, if one accepts the values provided by the manufacturer i.e. those from Telefonika Kable catalogue [4], the current-carrying capacity will be 480A. The calculated current value of a single core for the nominal transformer load was 429A. It can be assumed then, that the cables have been selected correctly, with a substantial safety margin up to 10%. The designed system was made in accordance with the specification and put into operation. After a few days of operation, the system suffered a fault, which resulted in a total destruction of the cable line.

Measurements and registration of currents in the cable line

After rebuilding the power system, current measurements of each phase in the cable line were made in order to exclude possible asymmetries of the load and system over-current. We have also made check-up measurements of the circuit breaker, equipped with integrated protection system. The check-up of the circuit breaker excluded the possibility of its malfunction while the measurements confirmed that the system load was symmetrical.

Fot.1 Cables after fault

The maximum recorded effective value of the current did not exceed the current-carrying capacity in amperes. However, we have noticed a considerable discrepancy between the effective values of current in particular lines being a part of one phase. The courses of effective current values in extremely loaded lines for a selected period of time are presented below.

Fig.2. RMS current extremely loaded core of line

The difference in the load of extremely loaded lines is nearly double. For line L1, the effective current value of a maximum of 528A was recorded, while this value was 280A for line L3 – the proportion of load for extremely loaded lines was 1.9. It is worth recalling that the lines were made from the same material and had the same length whereas differences in the resistance of cable clamps were excluded from the analyses. Due to the significant differences in the load of particular lines as well as the possibility of another fault, we have investigated the causes of the uneven load of the lines. For his purpose, a computational model of the system was developed.

The computational model of the system

The simulations were made in FEMM program which is a suite of programs for solving low frequency electromagnetic problems on two-dimensional planar and axi-symmetric domains. The program currently addresses linear/nonlinear magnetostatic problems, linear/nonlinear time harmonic magnetic problems, linear electrostatic problems.

Six cores with the cross-section of 240mm2 and total length of 30m were entered into the model. Then, a sinusoidal current flow (amplitude 3632 A, frequency 50 Hz) was forced as the excitation source, with The cores were assumed to be made of AL 1100 aluminum, with electric conductivity γ = 34,45 MS/m. The cable insulation was not considered in the simulation.

The calculations were made for various types of line configuration – vertically (two variants), spherical layout, parallel layout. For each case, two coefficients kAS –unbalance and kPZ – overload were defined

(1) kAS=Imax/Imin,
(2) kPZ=6*Imax/IC,
Imax – current amplitude of the highest load core,
Imin – current amplitude of the lowest load core,
IC – current amplitude of the circuit (one phase).

Parallel single core in the flat layout In the first variant, the core of the circuit being one phase were laid out as flat – vertically laid, with a distance of 17.5 mm between each one, which means that the gap between each line was equal to the line diameter. The simulation is a model of a real system. The mesh being used for calculations consisted of 44736 nodes and 88748 triangle elements. Table 1 presents the results of the total current calculations for each core, whereas Figure 3 presents the real component of magnetic vector potential A, B = rot A, rot A=0 for this section and the distribution of current density of each core. Fig. 4 presents the eddy and source current density – cross-section of cores (X-Y Plot).

Table 1. Total current of each core

.
Fig.3 Real component of magnetic vector potential A, distribution of current density
Fig.4 Eddy and source current density – cross-section of cores

The values of the asymmetry and overload coefficients are as follows: kAS=1.78, kPZ=1.32 The next variant features the lines of cores that were laid out flatly, with a distance of 35 mm between each one, which is equal to the double diameter of the line. The mesh being used for calculating had 44865 nodes and 89006 triangle elements. Table 2 presents the results of the total current calculations for each core, Fig. 5 presents the real component of magnetic vector potential A for this section and the distribution of current density of the each core. The values of the asymmetry and overload coefficients are as follows: kAS=1.73, kPZ=1.33.

Table 2. Total current of each core

.
Fig.5 Real component of magnetic vector potential A, distribution of current density
Parallel single core in the spherical configuration

The cores were laid out spherically – around a circle. The lines did not come in contact with each other. The mesh being used for calculations had 42953 nodes and 85182 triangle elements. Table 3 presents the results of the total current calculations for each core, Fig. 6 presents the real component of magnetic vector potential A for this section and the distribution of current density of the each core. The values of the asymmetry and overload coefficients are as follows: kAS=1.03, kPZ=1.1

Table 3. Total current of each core

.
Fig.6 Real component of magnetic vector potential A, distribution of current density
Current bus bar in a parallel layout

The lines were laid out parallelly in two layers, with a distance of 17.5 mm between each line and each layer. The mesh being used for calculating had 44701 nodes and 888678 triangle elements. Table 4 presents the results of the total current calculations for each core, Fig. 7 presents the real component of magnetic vector potential A for this section and the distribution of current density of the each core. The values of the asymmetry and overload coefficients are as follows: kAS=1.39, kPZ=1.1

Table 4. Total current of each core

.
Fig.7 Real component of magnetic vector potential A, distribution of current density
Summary and final conclusions

The simulations conducted on the prepared model proved to be compatible with the measurements recorded in the real system. The calculated asymmetry coefficients are, respectively, kAS=1.9 for the real system, and kAS=1.73 for the model. The analysis of mutual interaction between parallel lines that formed one circuit showed a sign influence of the proximity effect on the load layout which is also reported in publications [5],[6]. The configuration of core layout shown in PN-IEC 60364-5-523:2001 National Standard as an optimal one (horizontal layout, with the double diameter distance between the lines), seems to be the worst solution if field phenomena are to be taken into consideration. An asymmetry coefficient above 1.7 and overload coefficient above 1.3 practically render it impossible to construct multi-conductor low-voltage parallel cables. An optimal solution from the simulations conducted proved the spherical layout with the coefficient values of kAS=1.03, kPZ=1.1. It seems advisable to conduct further studies into the current carrying capacity of multi-conductor low-voltage parallel cables, including three-phase systems, harmonics and thermal phenomena in order to work out an optimal way of constructing multi-conductor cables and make corrections to norm [1]. The authors are planning to make a simulation using another tool (Maxwell from Ansoft), construction of a real-life model and taking measurements in real objects with transformer rated power over 1000KVA that include multi-conductor cable lines. Currently, designing multi-conductor cable lines in accordance with the Standard [1], without considering the proximity effect and performing additional calculations is both erroneous and hazardous.

REFERENCES

[1] PN-IEC 60364-5-523:2001 Instalacje elektryczne w obiektach budowlanych – Dobór i montaż wyposażenia elektrycznego – Obciążalność prądowa długotrwała przewodów
[2] Skibko Z, Obciążalność prądowa przewodów ułożonych wielowarstwowo. Rozprawa Doktorska, Politechnika Białostocka, Maj 2008r
[3] Skibko Z., Analityczne metody wyznaczania obciążalnościprądowej długotrwałej przewodów ułożonych wielowarstwowo, Przegląd Elektrotechniczny 4 (2009), s. 190-194
[4] Telefonika Kable, „Solidna Energia – Katalog kabli i przewodów Elektroenergetycznych”, 2013 r., s. 156,172
[5] Piątek Z., Modelowanie linii, kabli i torów wielkoprądowych, yd. Politechniki Częstochowskiej seria Monografie nr 130, s.33-54
[6] Piątek Z., Jabłoński P., Podstawy teorii pola elektromagnetycznego, Wydawnictwo Naukowo-Techniczne Warszawa, 2010, s. 142, 277-282.

Autorzy: dr hab. inż. Lech Borowik, Politechnika Częstochowska Instytut Telekomunikacji i Kompatybilności Elektromagnetycznej, ul. Armii Krajowej 17, 42-200 Częstochowa, E-mail: borowik@el.pcz.czest.pl;
mgr inż. Artur Cywiński, Pracownia projektowa omega-projekt, ul. Topolowa 1, Tychy, E-mail: artur.cywinski@omega-projekt.pl

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 92 NR 1/2016. doi:10.15199/48.2016.01.16

The Measurement of Input Power of Power Supply in Network Disturbed by Low Frequency Distortions

Published by Stanisław GALLA, Arkadiusz SZEWCZYK,
Gdansk University of Technology Faculty of Electronics, Telecommunications and Informatics

Abstract. In the paper authors present results of observation of input power changes versus harmonics amplitude in supply voltage of low-power power supply device. In the study, the electrical measurements supported with thermal imaging were used. The input circuit elements of studied device responsible for input power increase are pointed.

Streszczenie. W artykule przedstawiono wyniki pomiarów zmian mocy wejściowej pobieranej z sieci zasilającej przez zasilacz małej mocy w funkcji harmonicznych napięcia zasilającego. Wytypowano elementy obwodu wejściowego badanego urządzenia odpowiedzialne za zwiększenie poboru mocy przy zasilaniu napięciem silnie zniekształconym. W badaniach oprócz pomiarów elektrycznych wykorzystano również obrazowanie termograficzne. (Pomiar parametrów mocy wejściowej zasilacza zasilanego przez napięcia odkształcone zaburzeniami niskich częstotliwości).

Keywords: power supply, distortions, measurements.
Słowa kluczowe: zasilacze, zaburzenia, pomiary.

Introduction

Main parameters that describe electronic equipment declared by producer are, among of others, supply voltage, current consumption, working frequency and efficiency. Producer by declaring values of parameters guarantees that in standard working conditions their values will be in specified range. For equipment intended for domestic use, values of those parameters are determined for standard supply conditions, that means the supply voltage is sinusoidal with specified frequency and amplitude (RMS value). However, in real supply nets the voltage is disturbed by low and high frequency components, what was described e.g. in [1, 2, 3, 4]. For low frequency range, the one of parameters that describes the quality of supply voltage is total harmonics distortion (THD). The methodology of THD estimation and its maximum values are briefly described in several international standards [5, 6]. Currently, it is presumed that in public low voltage supply nets the THD value should not exceed 8%. Unfortunately, significantly higher values of THD are registered in several supply subnets and can have harmful influence on electric and electronic equipment that is supplied with such a voltage. The example influence of disturbed supply voltage on electronic apparatus, which is switching power supply is described in this report.

Issue consideration

The considered problem was reported by one of the electronic equipment producer. The user of low-power (<70 W) switching power supply reported significant discrepancy between supply current value he recorded and the value declared in technical note. The user pointed also differences in values of other parameters measured for nominal load. During the work on the problem, it was recognized that even though the supply voltage was 230 VRMS (+5 VRMS / -7 VRMS) it was strongly distorted by harmonics. Conducted measurements showed that THD value reached 11% and both even and odd harmonics was observed. Atypically, the dominant 4th and 8th order harmonics were observed. It’s worth to mention here, that for low-power power supply sources the power ratio correction circuit is not required. The mentioned power supply source was then examined in laboratory conditions.

Experiments

In order to examine the influence of each particular harmonic on device behaviour, the device under test was supplied with voltage distorted with harmonic (one at a time) which amplitude exceeded acceptable value specified e.g. in standard [4] and the current consumption was monitored. The device under test was 60 W switching power adapter with following parameters: nominal supply voltage Un = 230 V, nominal power P = 60 W, input current InRMS = 0.6 A, nominal output current, Io = 3 ADC, nominal output voltage Uo = 18 VDC. Producer guarantees 20% of parameters accuracy, according to appropriate standards. Figure 1 shows the input circuit of the device.

Fig.1. Input circuit of the device under test
Measurement set-up

Examined device was supplied by CHROMA 61502 programmable AC source. The output RMS voltage of the AC source was URMS = 230 V (+/- 1%). The 50 Hz component of the output voltage was U1RMS = 228,8 V and was constant during the test. The RMS value of each harmonic component was UnRMS = 23 V, for n = 2 to 40 and the phase shift of the component was 90 degree. Figure 2 shows example test waveform for n = 4 (main component with 4th harmonic).

Fig.2. Test waveform for n = 4

During the test, the device was loaded with the constant current Io = 3 A which gives output power Po = 54 W by electronic load Array 3721A. Under all test input and output parameters was monitored. Measured input parameters are: root mean square value of input voltage in frequency range 50 Hz – 2 kHz (URMS), root mean square value of input current in frequency range 50 Hz – 2 kHz (IRMS), peak current (Ip), real power (P), reactive power (Q), apparent power (S), peak current (IP), power factor (PF), crest factor (CF).

During the measurement, the device under test was also observed by thermographic camera VIGO V50. The camera has spectral range of 8 μm to 14 μm and 384 x 288 pixels resolution. The camera was equipped with 15o x 11o lens [7]. Registered thermal images allowed for identification of circuits and components responsible for elevated power consumption by the device under test. First, the thermal image of the DUT was registered for pure, 50 Hz supply, with no harmonics. Next this image was compared with thermal images taken for power supply disturbed by harmonics. Images was taken in thermal equilibrium state, when the temperature of the DUT’s components was stable.

Results and discussion

Figure 3 shows real, reactive and apparent power for each harmonic. The increase in reactive and apparent power is observed for low order harmonics. Moreover, the more significant increase is observed in case of presence of even harmonics.

Fig.3. Real, reactive and apparent input power

The maximum increase, up to 210 VA is observed for 8th harmonic, while for not disturbed voltage the value is about 120 VA. For higher harmonics, above 25th , the difference in influence of even and odd harmonics is not significant and both reactive and apparent power decrease.

Figure 4 and 5 show input current (RMS and peak value) and crest factor respectively. Both RMS value of current and crest factor corresponds to power consumed by tested circuit.

Fig.4. Root mean square value of input current, IRMS, for particular harmonics
Fig.5. Changing crest factor under the test

Additionally, the influence of phase shift between fundamental component and harmonic was studied. Figure 6 shows the change in value of real, reactive and apparent power for 3rd harmonic. Received data shows that the influence of phase shift of the harmonic on power lost in device is less than 20% (+/- 10%) and can be treated as being in tolerance of parameters specified by producer.

Fig.6. Real, reactive and apparent power for 3rd harmonic versus phase shift

During the test, the temperature of device’s circuit board was monitored using thermal imaging camera VIGO v50. Figure 7 shows thermal image of DUT’s circuit board when supplied with not disturbed supply voltage (with no harmonics added).

Figure 8 shows thermal image of DUT’s circuit board when powered with supply voltage disturbed by 8th harmonic, while figure 9 shows the one for 11th harmonic. As is seen in figures 3, 4 and 5, for the 8th harmonic the maximum input power and for 11th one the maximum value of crest factor are observed.

Comparing thermal images stored for various supply conditions allows to point out regions and components, where the energy is lost in a form or heat radiation.

In our case the main temperature increase is observed in compensated choke, L (2 x 15mH). Increase of temperature of input filters circuit points, that the energy is lost there and for even harmonics the temperature of filter elements is twice higher than for odd harmonics.

Fig.7. Thermal image of circuit board of tested device supplied with supply voltage with no harmonics (fundamental component only) 1 – elements of input circuit shown in figure 1
Fig.8. Thermal images of circuit board of tested device. Supply voltage disturbed with 8th harmonic 1 – elements of input circuit shown in figure 1
Fig.9. Thermal images of circuit board of tested device. Supply voltage disturbed with 11th harmonic tested device. Supply voltage disturbed with 8th harmonic 1 – elements of input circuit shown in figure 1
Conclusions

Conducted measurements confirmed observation results reported by the device user. In specific supply conditions when the supply voltage is strongly disturbed, device parameters value can be significantly out of range given by producer. It’s worth to mention here, that in standard supply conditions device parameters are within the specified range. In most supply nets, the voltage is distorted mainly by odd harmonics, for which the observed increase in reactive and apparent power is much lower than for even ones. The device user should take appropriate steps to decrease disturbances in his supply net and its adverse influence on various electric devices.

Observed enormous increase in power consumption caused by even harmonics results in increase of input circuit elements temperature. That thermal exposure can lead to decease of device reliability. It should be taken into consideration by producers if it’s worth to evaluate device parameters for maximum negative supply conditions, e.g. for maximum distortion values specified in appropriate standards.

Nevertheless, it should be stated here, that the input filter circuit of examined supply device is not chosen fortunately. It meets restrictions pointed in standard [6] for class B devices in terms of conducted emission, while used inductive element (inductive choke) shows increased thermal emission and cause increase in input power, especially for even harmonics distortions.

REFERENCES

[1] Łuszcz J., Oddziaływanie przekształtników energoelektronicznych dużej mocy na jakość energii elektrycznej, Zeszyty Naukowe Wydziału Elektrotechniki i Automatyki, Nr 31/2012, 211-214
[2] Hanzelka Z., Jakość dostaw energii elektrycznej. Zaburzenia wartości skutecznej napięcia. Akademia Górniczo Hutnicza, 2013
[3] Shafiul I. M, Chowdhury N., Sakil A. K., AtifIqbal K.A., Abu-Rub H., Power Quality Effect of Using Incandescent, Fluorescent, CFL and LED Lamps on Utility Grid, 978-1-4673-6765-3/15
[4] Blackledge J., O’Connell K., Barrett M., Sung A., Cable heating effects due to harmonic distortion in electrical installations, International Association of Engineers: ICEEE12, London, 2012
[5] PN-EN 50160:2010, Parametry napięcia zasilającego w publicznych sieciach elektroenergetycznych
[6] EN 61000-4-30:2015 Electromagnetic compatibility (EMC) – Part 4-30: Testing and measurement techniques – Power quality measurement methods
[7] http://www.vigo.com.pl/pub/File/PRODUKTY/Thermal-imagingsystem/v50.pdf, dostęp z sieci PG: 2016.06.13

Authors: dr inż. Stanisław Galla, Politechnika Gdańska, Wydział Elektroniki i Informatyki, Katedra Metrologii i Optoelektroniki, ul. Narutowicza 11/12, 80-233 Gdańsk, E-mail: galla@eti.pg.gda.pl;
dr inż. Arkadiusz Szewczyk, Politechnika Gdańska, Wydział Elektroniki i Informatyki, Katedra Metrologii i Optoelektroniki, ul. Narutowicza 11/12, 80-233 Gdańsk, E-mail: szewczyk@eti.pg.gda.pl

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 92 NR 11/2016. doi:10.15199/48.2016.11.07

Thermal Problems in HTS Transformer due to Inrush Current

Published by Grzegorz KOMARZYNIEC1, Tadeusz JANOWSKI2, Grzegorz WOJTASIEWICZ2, Michał MAJKA2, Politechnika Lubelska, Instytut Podstaw Elektrotechniki i Elektrotechnologii (1), Instytut Elektrotechniki w Warszawie (2)

Abstract. With switching on a superconducting transformer to the energetic network, a unidirectional current of high amplitude may appear. Maximal values of the current impulses following shortly one after the other may be several times higher than the critical current values for the superconductor used for making the transformer’s windings. Because of the resistive area propagation overheating of the superconducting tape occurs and that might lead to switching the transformer off working. In the present article the results for transformer HTS of 10 kVA power were presented.

Streszczenie. Włączeniu transformatora nadprzewodnikowego do sieci energetycznej może towarzyszyć prąd jednokierunkowy o dużej amplitudzie. Wartości maksymalne następujących krótko po sobie impulsów prądu mogą przekraczać wielokrotnie prąd krytyczny nadprzewodnika, z którego wykonano uzwojenia transformatora. W skutek propagacji strefy rezystywnej dochodzi do przegrzania taśmy nadprzewodzącej, co może skutkować wyłączeniem transformatora z eksploatacji. W artykule przedstawiono wyniki badań transformatora HTS o mocy 10 kVA. (Problemy cieplne w transformatorze HTS spowodowane przepływem prądu włączania)

Keywords: inrush current, transformer, superconducting, temperature.
Słowa kluczowe: prąd włączania, transformator, nadprzewodnictwo, temperatura.

Introduction

In certain conditions, with switching the transformer onto the energetic network a high value unidirectional current will flow in the windings [1]. Its first impulse may be 40 times higher than the transformer’s rated current. The following impulses are attenuated by the transformer’s windings resistance and the powering circuit’s resistance [2]. Depending on the transformer’s power, this current’s duration may be equal to from several to more than ten thousand periods of the feeding voltage.

Inrush current, in the transformer’s power circuit, causes a series of disadvantageous phenomena [3]. In superconducting transformers (HTS) the flow of this current may cause loss of the superconductive windings state as a result of exceeding critical superconductor values: density of the current, field intensity, temperature.

Due to the heterogeneous structure of the superconductor, the weakest area in the windings becomes the origin of resistive area. In the place of that area’s occurring, the inrush current flows through the stabilizer generating heat by power losses on its resistance, according to the Jouel’s law. The resistive area propagation onto neighboring regions of the superconductor depends on the heat spread speed in the tape, cooling system efficiency and the inrush current wave’s shape.

The windings overheating threatens with the superconductor’s failure and the transformer’s switching off of working [4][5]. Continuing degradation of a second generation superconductor is observed at the temperature of over 600K. Keeping windings at cryogenic temperature is one of the most complex issues in the HTS transformers exploitation [6][7].

Object of research

The trials were executed on a one phase HTS transformer of 10kVA power. The core was done as wound and cut with metal sheet PN ET52-27 of induction B=1.75 T with H=10 A/cm and loss of P=0.8 W/kg at B=1T and f=50 Hz. The transformer’s primary and secondary voltage windings were made of superconducting tape Super Power SCS4050 (RE)BCO of effective value of critical current of 80 A in temperature of 77K in self field. The tape’s bulk was 0.1 mm, the width 4 mm. The windings were isolated by wrapping the superconducting tape with kapton. The windings geometry was shown in the Figure 1. The transformer’s windings are being cooled by liquid nitrogen to reach the temperature of 77K. The transformer’s core works at room temperature. The rated parameters were given in the Table 1.

Fig. 1. The geometry of superconducting transformer

Table 1. The rating of the transformer

.
Measurement

Measurements were taken in the system presented in Figure 2. The current’s flow was registered indirectly, by measuring the voltage drop on the shunt 60A, 60mV and exactness class 0.5. The data acquisition was realized with measuring card National Instruments USB-6212, using application written at LabVIEW. Switching the transformer on was done with a thyristor system switching the transformer on at the moment of passing of the power network’s voltage the zero value.

Fig. 2. The measuring system
Fig. 3. The inrush current of HTS transformer
Fig. 4. The initial pulses of inrush current

In Figure 3 we presented registered inrush current flows i(t) and voltage upw(t) measured at the HTS transformer’s primary windings. Since the transformer’s switching on, the voltage changes sinusoidally. The highest measured peak value of the TrHTS transformer’s inrush current is 178 A.

Table 2. The parameters inrush currents

.

HTS transformers inrush current decay time is longer as compared with transformers with copper windings [8][9]. The examined transformer’s registered unidirectional current decay time was 350 ms.

The sufficient condition of going of the superconducting tape into resistive state is exceeding the value of superconductor’s critical current. Exceeding critical current value of tape SCS 4050 (Ic=80 A) occurs after t0=4 ms (Fig. 4) as measured since switching the transformer onto the power network. The time tg in which the exceeding occurs is 5 ms. In this period it is expected that the tape loses its superconductive properties and its temperature grows.

The circuit with HTS transformer showed rapid attenuation of unidirectional impulses down to current values lower than the superconductor’s critical current. The first impulse’s amplitude exceeds the critical current (80 A) of the superconducting tape SCS4050 (Fig. 4) by 98 A and the transformer’s rated current (44 A) by 134 A. The second impulse coming after 0.02 s is comparable to the critical current value (80 A) and the peak value of the third impulse (50 A) (after 0.04 s) is by 30 A lower than the tape’s critical current value.

Endurance of the SCS4050 tape for thermal failure depends on distribution of the current’s density in each layer at resistive condition. The construction of the tape are shown in Table 3.

Table 3. Layer structure of tape SCS4050

.

Based on the resistivity of the individual materials, you can calculate the percentage of the current dispersed into individual layers. Assuming the materials are homogeneous the tape’s structure was pictured by parallel connections between resistances representing each layer (Fig. 5) [10].

Fig. 5. Equivalent circuit of SCS4050 tape

In the superconducting state all the current flows through superconductor. Calculations show, that in resistive state of the SCS4050 tape, at temperature of 77K, 89% of current flows through Cu layer, 9.4% through the Ag layer and 1.5% through the Hastelloy layer. With temperature growth, the resistance of layers materials changes. In addition, the reflow of current changes. At 293K we get 88.7%, 9.5%, 1.5% respectively.

The main conductor at the resistive state is copper. At moment tm (Fig. 4) when the inrush current reaches its maximal value, a current of 158 A flows through the copper layer and that gives the momentary current density of 987 A/mm2 (Tab. 4). This value is 318 times higher than the maximal acceptable value of current density for copper in the air (3.1 A/mm2). A higher momentary density is observed for the silver layer. At moment tm it is 1112 A/mm2.

Return of the HTS transformer’s winding to its superconducting state happens when three conditions occur simultaneously: (1) intensity of the outer magnetic field is lower than its critical value; (2) the inrush current maximal value is lower than the critical value; (3) superconductor’s temperature is lower than the critical temperature.

The first condition, because of great critical values of the 2G superconductors’ field intensity, is met for the entire inrush current lasting time. The second condition is met for the time ts=18 ms (Fig. 4) between consequent impulses of the unidirectional component. By this time the cooling process of the superconductor to cryogenic temperature takes place. If the third condition is met depends on the cooling system efficiency in tome ts. The cooling intensity strongly depends on temperatures difference between liquid nitrogen and the cooled surface. In case of the smallest heat currents and the smallest temperatures differences the heat is transferred due to natural convection. With raise in superconducting tape’s temperature vacuolar boiling of the cryogenic liquid occurs until in the peripheral layer the gas form of nitrogen appears. This worsens conditions for transformer’s windings cooling and lowers isolation durability against breakthrough.

Table 4. The instantaneous current density in layers for time tm (resistive superconducting state)

.

The third condition is hard to measure by simply measuring the temperature. The tape’s temperature can be estimated by electrical rates measuring. The medial resistivity of the SCS4050 tape in normal state is interesting, estimated for different temperatures. It can be calculated from the equation:

where: pi, Si–resistivity and cross sectional area of the i-th material of the superconducting tape, S – total cross-sectional area of the superconducting tape. Knowing the characteristic p=f(T) the superconducting tape’s temperature can be established by measuring electrical rates.

.

The medial resistivity of the SCS4050 tape at 77K, calculated from the (2) equation is:

.

The resistivity at 293K known from the direct measurements is:

.

That gives the conclusion, that change in SCS4050 tape’s temperature by 216K, results in more than 10 times’ raise in medial resistivity. During lasting of inrush current that kind of raise in resistance have not been noted along the whole of the winding (55 cm). Because the current at time tg exceeds critical current it can be expected, that a hard to spot, local loss of superconductivity occurs.

The examined HTS transformer went through numerous switching on trials. Despite exceeding the critical value by inrush current and a great exceed in acceptable current density for the copper stabilizer no failure of the SCS 4050 tape have been noted.

Conclusions

The methods in designing conventional transformers, with copper or aluminum windings, do not take into account the inrush current phenomenon. This experiment showed that the phenomenon may cause problems in switching on the superconducting transformer. The single unidirectional current impulse of great momentary density observed during switching on a transformer HTS of 10 kVA power, exceeding critical current, may lead to thermal damage in superconducting windings. The experiment showed, that it is difficult to establish the maximal temperature value and spot its appearing in the windings.

The loss of HTS transformer’s superconducting state (with raise in windings temperature in safe limits) may limit negative impact of inrush current. Raise in circuit resistance due to raise in windings resistance lowers the inrush current’s amplitude and its duration [11][12].

HTS transformers should be designed in the way that they stand certain inrush current for a certain period of time without exceeding the acceptable temperature for the transformer’s superconducting windings. There are no respective norms for the time being.

The research was conducted in scope of the project “Analysis of inrush current phenomenon and the phenomena related in superconducting transformers”. The project was financed with means of National Science Center given with the decision no. DEC- 2012/05/D/ST8/02384

The superconducting transformer was constructed in scope of research project no: N510526439, “Elaborating the model design for 1-phase superconducting transformer with windings made of HTS tape of 2nd generation”.

REFERENCES

[1] E. Jezierski, Transformatory, WNT, Warszawa (1983)
[2] M. Jamali, M. Mirzaie, S. Asghar Gholamian, Calculation and analysis of transformer inrush current based on parameters of transformer and operating condictions, Electronics and Electrical Engineering, Electrical Engineering, no. 3(109), (2011)
[3] R. A. Turner, K. S. Smith, Resonance Excited by Transformer Inrush Current in Inter-connected Offshore Power Systems, IEEE Industry Applications Society Annual Meeting, Edmonton, Canada, October (2008)
[4] G. Wojtasiewicz, T. Janowski, S. Kozak, J. Kozak, M. Majka, B. Kondratowicz-Kucewicz, Tests and Performance Analysis of 2G HTS Transformer, IEEE Transactions on Applied Superconductivity, vol. 23, issue: 3, part: 2, (2012)
[5] G. Wojtasiewicz, T. Janowski, S. Kozak, J. Kozak, M. Majka, B. Kondratowicz-Kucewicz, Experimental Investigation of the Model of Superconducting Transformer With the Windings Made of 2G HTS Tape, IEEE Transactions on Applied Superconductivity, vol. 22, issue: 3, (2012)
[6] S. S. Kalsi, Applications of High Temperature Superconductors to Electric Power Equipment, John Wiley and Sons Ltd, April (2011)
[7] J. K. Sykulski, C. Beduz, R.L. Stoll, M.R. Harris, K.F. Goddard, Y. Yang, Prospects for large high-temperature superconducting power transformers: conclusions from a design study, Electric Power Applications, IEE Proceedings, vol. 146, issue: 1, (1999)
[8] G. Wojtasiewicz, G. Komarzyniec, T. Janowski, S. Kozak, J. Kozak, M. Majka, B. Kondratowicz-Kucewicz, Inrush Current of Superconducting Transformer, IEEE Transaction on Applied Superconductivity, vol. 23, issue: 3, June (2013)
[9] G. Komarzyniec, T. Janowski, G. Wojtasiewicz, M. Majka, J. Kozak, S. Kozak, B. Kondratowicz-Kucewicz, Prąd włączania transformatora nadprzewodnikowego, Przegląd Elektrotechniczny, no. 9, (2013)
[10] M. Sjöström, B. Dutoit, and J. Duron, Equivalent Circuit Model for Superconductors, IEEE Transactions on Applied Superconductivity, vol. 13, no. 2, June (2003)
[11] H. Shimizu, K. Mutsuura, Y. Yokomizu, T. Matsumura, Inrush-Current-Limiting with high Tc superconductor, IEEE Transactions on Applied Superconductivity, vol. 15, no. 2, June (2005)
[12] S. Hun-Chul, K. Chul-Hwan, R. Sang-Bong, K. Jae-Chul, H. Ok-Bae, Superconducting Fault Current Limiter Application for Reduction of the Transformer Inrush Current: A Decision Scheme of the Optimal Insertion Resistance, IEEE Transactions on Applied Superconductivity, vol. 20, no. 4, August (2010)

Authors: dr inż. Grzegorz Komarzyniec, E-mail: g.komarzyniec@pollub.pl, Politechnika Lubelska, Instytut Elektrotechniki i Elektrotechnologii, ul. Nadbystrzycka 38a, 20-618 Lublin,
prof. dr hab. inż. Tadeusz Janowski, E-mail: t.janowski@pollub.pl,
dr inż. Grzegorz Wojtasiewicz, E-mail: g.wojtasiewicz@iel.waw.pl,
dr inż. Michał Majka, E-mail: m.majka@iel.waw.pl, Instytut Elektrotechniki, ul. Pożarskiego 28, 04-703 Warszawa

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 90 NR 2/2014. doi:10.12915/pe.2014.02.04

Principle of lightning Protection

Published by Electrical Installation Wiki, Chapter J. Overvoltage protection – Principle of lightning protection

General rules of lightning protection

Procedure to prevent risks of lightning strike

The system for protecting a building against the effects of lightning must include:

• protection of structures against direct lightning strokes;
• protection of electrical installations against direct and indirect lightning strokes.

The basic principle for protection of an installation against the risk of lightning strikes is to prevent the disturbing energy from reaching sensitive equipment. To achieve this, it is necessary to:

• capture the lightning current and channel it to earth via the most direct path (avoiding the vicinity of sensitive equipment);

• perform equipotential bonding of the installation;
This equipotential bonding is implemented by bonding conductors, supplemented by Surge Protection Devices (SPDs) or spark gaps (e.g., antenna mast spark gap).

• minimize induced and indirect effects by installing SPDs and/or filters.

Two protection systems are used to eliminate or limit overvoltages: they are known as the building protection system (for the outside of buildings) and the electrical installation protection system (for the inside of buildings).

Building protection system

The role of the building protection system is to protect it against direct lightning strokes.

The system consists of:

• the capture device: the lightning protection system;
• down-conductors designed to convey the lightning current to earth;
• “crow’s foot” earth leads connected together;
• links between all metallic frames (equipotential bonding) and the earth leads.

When the lightning current flows in a conductor, if potential differences appear between it and the frames connected to earth that are located in the vicinity, the latter can cause destructive flashovers.

The 3 types of lightning protection system

Three types of building protection are used:

The lightning rod (simple rod or with triggering system)

The lightning rod is a metallic capture tip placed at the top of the building. It is earthed by one or more conductors (often copper strips) (see Fig. J12).

Fig. J12 – Lightning rod (simple rod or with triggering system)

The lightning rod with taut wires

These wires are stretched above the structure to be protected. They are used to protect special structures: rocket launching areas, military applications and protection of high-voltage overhead lines (see Fig. J13).

Fig. J13 – Taut wires

The lightning conductor with meshed cage (Faraday cage)

This protection involves placing numerous down conductors/tapes symmetrically all around the building. (see Fig. J14).

This type of lightning protection system is used for highly exposed buildings housing very sensitive installations such as computer rooms.

Fig. J14 – Meshed cage (Faraday cage)
Consequences of building protection for the electrical installation’s equipment

50% of the lightning current discharged by the building protection system rises back into the earthing networks of the electrical installation (see Fig. J15): the potential rise of the frames very frequently exceeds the insulation withstand capability of the conductors in the various networks (LV, telecommunications, video cable, etc.).

Moreover, the flow of current through the down-conductors generates induced overvoltages in the electrical installation.

As a consequence, the building protection system does not protect the electrical installation: it is therefore compulsory to provide for an electrical installation protection system.

Fig. J15 – Direct lightning back current
Lightning protection – Electrical installation protection system

The main objective of the electrical installation protection system is to limit overvoltages to values that are acceptable for the equipment.

The electrical installation protection system consists of:

• one or more SPDs depending on the building configuration;
• the equipotential bonding: metallic mesh of exposed conductive parts.

Implementation

The procedure to protect the electrical and electronic systems of a building is as follows.

Search for information

• Identify all sensitive loads and their location in the building.
• Identify the electrical and electronic systems and their respective points of entry into the building.
• Check whether a lightning protection system is present on the building or in the vicinity.
• Become acquainted with the regulations applicable to the building’s location.
• Assess the risk of lightning strike according to the geographic location, type of power supply, lightning strike density, etc.

Solution implementation

• Install bonding conductors on frames by a mesh.
• Install a SPD in the LV incoming switchboard.
• Install an additional SPD in each sub-distribution board located in the vicinity of sensitive equipment (see Fig. J16).

Fig. J16 – Example of protection of a large-scale electrical installation
The Surge Protection Device (SPD)

Surge Protection Devices (SPD) are used for electric power supply networks[1], telephone networks, and communication and automatic control buses.

The Surge Protection Device (SPD) is a component of the electrical installation protection system.

This device is connected in parallel on the power supply circuit of the loads that it has to protect (see Fig. J17). It can also be used at all levels of the power supply network.

This is the most commonly used and most efficient type of overvoltage protection.

Fig. J17 – Principle of protection system in parallel

SPD connected in parallel has a high impedance. Once the transient overvoltage appears in the system, the impedance of the device decreases so surge current is driven through the SPD, bypassing the sensitive equipment.

Principle

SPD is designed to limit transient overvoltages of atmospheric origin and divert current waves to earth, so as to limit the amplitude of this overvoltage to a value that is not hazardous for the electrical installation and electric switchgear and controlgear.

SPD eliminates overvoltages

• in common mode, between phase and neutral or earth;
• in differential mode, between phase and neutral.

In the event of an overvoltage exceeding the operating threshold, the SPD

• conducts the energy to earth, in common mode;
• distributes the energy to the other live conductors, in differential mode.

The three types of SPD

Type 1 SPD

The Type 1 SPD is recommended in the specific case of service-sector and industrial buildings, protected by a lightning protection system or a meshed cage.

It protects electrical installations against direct lightning strokes. It can discharge the back-current from lightning spreading from the earth conductor to the network conductors.

Type 1 SPD is characterized by a 10/350 µs current wave.

Type 2 SPD

The Type 2 SPD is the main protection system for all low voltage electrical installations. Installed in each electrical switchboard, it prevents the spread of overvoltages in the electrical installations and protects the loads.

Type 2 SPD is characterized by an 8/20 µs current wave.

Type 3 SPD

These SPDs have a low discharge capacity. They must therefore mandatorily be installed as a supplement to Type 2 SPD and in the vicinity of sensitive loads.

Type 3 SPD is characterized by a combination of voltage waves (1.2/50 μs) and current waves (8/20 μs).

SPD normative definition

Fig. J18 – SPD standard definition

Note 1: There exist [T1] + [T2]SPD (or Type 1 + 2 SPD) combining protection of loads against direct and indirect lightning strokes.
Note 2: some [T2] SPD can also be declared as [T3].

Characteristics of SPD

International standard IEC 61643-11 Edition 1.0 (03/2011) defines the characteristics and tests for SPD connected to low voltage distribution systems (see Fig. J19).

Fig. J19 – Time/current characteristic of a SPD with varistor

Common characteristics

• Uc: Maximum continuous operating voltage
This is the A.C. or D.C. voltage above which the SPD becomes active. This value is chosen according to the rated voltage and the system earthing arrangement.

• Up: Voltage protection level (at In)
This is the maximum voltage across the terminals of the SPD when it is active. This voltage is reached when the current flowing in the SPD is equal to In. The voltage protection level chosen must be below the overvoltage withstand capability of the loads. In the event of lightning strokes, the voltage across the terminals of the SPD generally remains less than Up.

• In: Nominal discharge current
This is the peak value of a current of 8/20 µs waveform that the SPD is capable of discharging minimum 19 times[2].

Why is In important?
In corresponds to a nominal discharge current that a SPD can withstand at least 19 times[2]: a higher value of In means a longer life for the SPD, so it is strongly recommended to chose higher values than the minimum imposed value of 5 kA.

Type 1 SPD

• Iimp: Impulse current
This is the peak value of a current of 10/350 µs waveform that the SPD is capable of discharging of discharging at least one time[3].

Why is Iimp important?
IEC 62305 standard requires a maximum impulse current value of 25 kA per pole for three-phase system. This means that for a 3P+N network the SPD should be able to withstand a total maximum impulse current of 100kA coming from the earth bonding.

• Ifi: Autoextinguish follow current
Applicable only to the spark gap technology. This is the current (50 Hz) that the SPD is capable of interrupting by itself after flashover. This current must always be greater than the prospective short-circuit current at the point of installation.

Type 2 SPD

• Imax: Maximum discharge current
This is the peak value of a current of 8/20 µs waveform that the SPD is capable of discharging once.

Why is Imax important?
If you compare 2 SPDs with the same In, but with different Imax: the SPD with higher Imax value has a higher “safety margin” and can withstand higher surge current without being damaged.

Type 3 SPD

• Uoc: Open-circuit voltage applied during class III (Type 3) tests.

Main applications

• Low Voltage SPD

Very different devices, from both a technological and usage viewpoint, are designated by this term. Low voltage SPDs are modular to be easily installed inside LV switchboards.

There are also SPDs adaptable to power sockets, but these devices have a low discharge capacity.

• SPD for communication networks

These devices protect telephone networks, switched networks and automatic control networks (bus) against overvoltages coming from outside (lightning) and those internal to the power supply network (polluting equipment, switchgear operation, etc.).

Such SPDs are also installed in RJ11, RJ45, … connectors or integrated into loads.

Notes

1. find out more in application examples such as SPD for EV charging application, SPD for photovoltaic applications and SPD application example in Supermarket
2. Test sequence according to standard IEC 61643-11 for SPD based on MOV (varistor). A total of 19 impulses at In:
• One positive impulse
• One negative impulse
• 15 impulses synchronised at every 30°on the 50 Hz voltage
• One positive impulse
• One negative impulse
3. for type 1 SPD, after the 15 impulses at In (see previous note):
• One impulse at 0.1 x Imp
• One impulse at 0.25 x Imp
• One impulse at 0.5 x Imp
• One impulse at 0.75 x Imp
• One impulse at Imp

Source URL: https://www.electrical-installation.org/enwiki/Principle_of_lightning_protection

How to Select the Right Current Transformer for Your Application

Published by Accuenergy, November 4, 2021.

Help me choose the Right CT

If you have a power measurement project coming up, chances are you have narrowed down your search for a power meter to a few choices. Whether it’s a multi-circuit application or a high-precision metering in an industrial setting, the next step in project preparation is selecting the right current transformer to maximize the performance of your power meter. When going through the selection process, it can be helpful to answer a few, basic application questions to reach a decision and consider several parameters including current transformer output, conductor size, amperage range, and accuracy. If you need help deciding, reach out to your power meter manufacturer so they can help guide you to the CT that best meets your project’s measurement objectives and budget.

CT Output:

What current transformer output is your power meter compatible with?

Current transformers are available with several output options, some of the most popular of which include 333mV, 5A, or 80mA. A critical question in the current transformer selection process, it is important to note which output your metering equipment is compatible with. While it is possible that the meter may work with multiple output options, it may not be possible to make in-field adjustments to this setting or it may need to be configured by the factory.

Unlike typical split-core or solid-core current transformers, Rogowski coils have a unique output that is generally rated at a low AC voltage (e.g., 150mV or less) per 1000A. In addition, there is an inherent 90-degree phase shift. Many meters and other measurement devices require a higher signal than what a Rogowski can provide on its own and are not configured to compensate for the phase shift, so it is important to work with your meter manufacturer to determine whether this specialized CT is directly compatible with your device.

Conductor Size:

Are you measuring around large busbars/conductors or small branch circuits?

The dimensions of the conductor are a critical consideration and can be one of the leading deciding factors in CT selection. Any CT that is used needs to be able to physically fit around the conductor you plan to measure. At the same time, oversizing a CT to accommodate a small conductor may not make sense in terms of both cost and the space required in the electrical panel, which may not have enough room to accommodate a large, rigid current transformer. In this situation, a flexible Rogowski coil can make it easier to measure in crowded electrical panels or switchgear because they can easily slip around oversized bus bars in tight spaces, making them an ideal compromise between large window size and flexible functionality.

Load Size:

How many amps will you be measuring?

Like the physical dimensions, the size of the load under measurement is a key consideration. All current transformers have a current input range, or amperage range, specification which indicates the size of the load they can effectively measure. If the load fluctuates throughout the day—for example, when occupancy is low during evening hours—it can be helpful to choose a current transformer with a broad current sensing range, such as a flexible Rogowski coil. It is also important to note that, if a load goes outside the sensor’s range, the meter may not be able to measure the load accurately, so it is important to always choose a sensor with a range that matches what you intend to measure.

Accuracy Rating:

Does the project involve billing tenants for their consumption?

When it comes to tenant billing, selecting equipment with the highest accuracy is of the utmost importance. In fact, in any application where “money changes hands,” power monitoring equipment must meet certain accuracy requirements and is often labeled “revenue grade” to indicate its conformance to accuracy standards. What does revenue grade accuracy mean? It is generally understood to be better than 1% accuracy and, more often, in the range of 0.5% accuracy or better. Before selecting a revenue grade sensor, be sure to check which industry accuracy standards they meet to ensure the accuracy class matches your project requirements. A common revenue grade accuracy standard is IEC 60044-1 0.5 Class.

On the other hand, if you’re simply collecting overall consumption trend data for a facility, a 1% accuracy sensor may be sufficient, and you may not need to upgrade to a revenue grade model.

Form Factor:

Will the project be new construction or a retrofit application?

This question may also be framed as, “Will a split-core or solid-core current transformer be a better for my application?” Although either sensor type may be used for any job, it is almost always easier to use a split-core, or a Rogowski coil, current transformer for a retrofit application because it can easily open to fit around a conductor and does not require wire disconnection as part of the installation process. Alternatively, while a facility is still under construction, installing a solid-core CT does not require much additional work since facility shutdowns or wire disconnection are not yet disruptive. Another consideration is cost: Although the up-front price of a solid core CT is lower, the initial savings is negligible when compared to the largely uncalculated installation cost which must include shutdowns and disconnections, adding time and labor to the overall project.

Regulatory Requirements:

Does your application require a sensor that meets UL or other regulatory certifications?

A UL Listed current transformer has undergone rigorous testing to ensure that it complies with nationally recognized safety standards. Unlike a current sensor that is a UL Recognized Component, which means it is intended to be a component within a complete system or product, a UL Listed sensor can be sold as an end-user product and is designed to minimize installation hazards such as shock or fire. It may be that your application mandates a UL Listed current sensor to meet safety code requirements. If this is the case, be sure to look for CTs with a UL Listed marking that indicates they comply with XOBA UL2808 and CSA C22.2 61010-1.

Another key regulatory requirement is a CE mark. This mark is required for products used in the European Economic Area (EEA) which includes countries like Germany, France, Spain, Italy, and others. Unlike other quality marks, such as UL, the CE mark on a product means that it conforms with European safety, health, and environmental standards. The CE mark should be visible on product labeling and documentation.

A third regulatory requirement you may encounter concerns Measurement Canada approval. Tenant billing applications in Canada may require both a Measurement Canada approved meter and current transformers, each of which must meet rating, design, accuracy, testing, and other requirements. For example, a few characteristics of Measurement Canada approved CTs include that they must be solid core, meet an accuracy class of 0.6% or better, and be either 5A, 80mA, or 100mA output devices. The nature, scope, and location of your project will dictate whether Measurement Canada approval is required. Check the product labeling and documentation to determine whether a sensor meets the regulatory requirements.

Using Rogowski coils:

My power meter does not work directly with Rogowski coils. Is there a way that I can still use a rope CT?

Nearly any project can benefit from Rogowski coil current transformers which offer many advantages including a large window size, broad amperage range, lightweight flexibility, and no saturation point. However, if your power meter only accepts 333mV, 5A, 1A, or another standard output, it will not directly work with a Rogowski coil. Fortunately, there is a simple solution to this challenge and that is to use an integrator. An integrator is an electronic device that makes it possible to change the output of a Rogowski coil to a commonly accepted output, such as 333mV or 5A, so that it can work with a host a power meters, protection relays, or other devices. By adjusting the input ranges to match nearly any system, an integrator is a simple solution to solve a common compatibility dilemma and bridges the divide between Rogowski coils and industrial metering equipment.

Source URL: https://www.accuenergy.com/articles/current-transformers/how-to-select-the-right-current-transformer/

Electrical Shock and its Effects

Published by Alex Roderick, EE Power – Technical Articles: Electrical Shock and its Effects, August 17, 2021.

Anyone working on electrical equipment should have respect for all voltages, have knowledge of the principles of electricity, and follow safe work procedures.

An electrical shock occurs when a body becomes a part of an electrical circuit. The effects of an electrical shock vary from a moderate sensation to paralysis to death. Also, severe burns may occur internally and where the current enters and exits the body. The quantity of electric current flowing through the body in milliamps (mA), the amount of time the body is exposed to the electric current, the route the current takes through the body, and the physical condition of the body through which the current flows; all influence the severity of an electrical shock.

Table 1. Electrical shock results in any time a body becomes part of an electrical circuit.

* in mA
**effects vary depending on time, path, amount of exposure, and condition of the body

Prevention is the best medicine for electrical shock. Anyone working on electrical equipment should have respect for all voltages, have knowledge of the principles of electricity, and follow safe work procedures. All technicians should be encouraged to take a basic course in cardiopulmonary resuscitation (CPR), so they can aid a coworker in emergency situations.

A person’s body becomes a part of an electrical circuit during an electrical shock. The body of a person offers varied resistance to the flow of current. Sweaty hands have less resistance than dry hands. The resistance of a wet floor is less than that of a dry floor. The lower the resistance, the higher the current flow. As the current flow increases, the severity of the electrical shock increases.

If a person is receiving an electrical shock, power should be removed as quickly as possible. If power cannot be removed quickly, the victim must be removed from contact with live parts. Action must be taken quickly and cautiously. Delay may be fatal. Individuals must also avoid being a casualty while attempting to rescue another person. If the equipment circuit disconnect switch is nearby and can be operated safely, shut OFF the power. Excessive time should not be spent searching for the circuit disconnect. In order to remove the energized part, insulated protective equipment such as a hot stick, rubber gloves, blankets, wood poles, plastic pipes, etc., can be used if such items are accessible.

After the victim is freed from the electrical hazard, help should be called, and first aid (CPR, etc.) begun as needed. The injured individual should not be transported unless there is no other option and the injuries require immediate professional attention.

Grounding

Grounding is the connection of portions of the distribution system to earth in order to establish a common electrical reference and a low impedance fault path to facilitate the operation of overcurrent protective devices. Grounding provides an electrically conductive path designed and intended to carry current under fault conditions from the point of a fault on a wiring system to the electrical supply source. Grounding facilitates the operation of overcurrent protection devices.

For systems that are solidly grounded, grounding provides a means to limit the voltage to the ground during normal operation and to prevent excessive voltages due to lightning, line surges, or unintentional contact with higher-voltage lines and to stabilize the voltage to the ground during normal operation.

The non-current-carrying metal parts of a transformer installation are required by the NEC® to be effectively grounded. Conductive materials enclosing conductors or equipment are grounded to prevent a voltage or difference of potential on these materials. Circuits and enclosures are grounded to allow overcurrent devices to operate in the event of insulation damage or ground faults.

The electrical distribution system is grounded by connecting it to a metal underground water pipe, a building’s metal frame, a concrete-encased electrode, or a ground ring. To prevent problems, a grounding path must be as short as possible and of sufficient ampacity, never be fused or switched, be a permanent part of the electrical circuit, and be continuous and uninterrupted from the electrical circuit to the ground.

The ground is provided at the main service equipment or at the source of a separately derived system (SDS). A separately derived system (SDS) is a system that supplies electrical power derived or taken from transformers, storage batteries, solar photovoltaic systems, or generators. See Figure 2. The majority of separately derived systems are produced by the secondary of a distribution transformer.

Figure 1. A separately derived system (SDS) is a system that supplies electrical power derived or taken from transformers, storage batteries, solar photovoltaic systems, or generators. Image Courtesy of Sask Power

The neutral ground connection must be made at the transformer or at the main service panel only. The neutral ground connection is made by connecting the neutral bus to the ground bus with a main bonding jumper. The main bonding jumper (MBJ) is a connection at the service equipment that connects the equipment grounding conductor, the grounding electrode conductor, and the grounded conductor (neutral conductor). The purpose of the main bonding jumper is to bond the neutral and equipment grounding conductor together with the enclosure to create a common reference potential.

An equipment grounding conductor (EGC) is an electrical conductor that provides a low-impedance ground path between electrical equipment enclosures within the distribution system and takes current back to the source. A grounding electrode conductor (GEC) is a conductor that connects grounded parts of a power distribution system (equipment grounding conductors, grounded conductors, and all metal parts) to the grounding system.

A grounded conductor is one that has been intentionally grounded. The grounded conductor is commonly a neutral conductor. However, not all electrical distribution systems use the grounded conductor as a neutral. For example, corner-grounded delta systems contain a grounded conductor that is not a neutral conductor. Therefore, it is not correct to refer to all grounded conductors as neutral conductors, although that is the case in the majority of electrical distribution systems.

Ground Fault Circuit Interrupters

A ground fault circuit interrupter (GFCI) detects an imbalance of current in the normal conductor routes and opens the circuit to safeguard against electrical shock. A GFCI opens the circuit when the current in two conductors of an electrical circuit differ by more than 5 mA. A GFCI is designed to trip quick enough (1/40 of a second) to avoid electrocution (1/4 of a second).

A potentially dangerous ground fault is any quantity of current above the level that might cause a harmful shock. Any current more than 8 mA is regarded as potentially harmful — depending on the path the current follows, the physical state of an individual receiving the shock, and the length of time the individual is exposed to the shock. GFCIs are therefore necessary for places like homes, hotels, resorts, industrial sites, receptacles around swimming pools, and other places where a person may encounter a ground fault.

A GFCI compares the current flowing through the ungrounded (hot) conductor with the current flowing through the neutral conductor. A ground fault occurs if the current in the neutral conductor falls below the current in the hot conductor. The missing current is returned to the source via some path other than the intended one (fault current).

GFCI protection can be installed at various points throughout a circuit. Ground fault protection is provided at the point of installation with direct-wired GFCI receptacles. GFCI receptacles can also be used to protect all other receptacles installed downstream along the same circuit.  When implemented in a load center or panel board, GFCI circuit breakers offer GFCI protection as well as conventional circuit overcurrent protection for all branch-circuit elements connected to the circuit breaker.

Plug-in GFCIs

Plug-in GFCIs protect against ground faults for devices that are plugged into them. These plug-in devices are frequently used by personnel working with power tools in areas without GFCI receptacles.

Portable GFCIs are designed to be easily moved from one location to another (see Figure 2). Portable GFCIs commonly contain more than one receptacle outlet protected by an electronic circuit module. Portable GFCIs should be inspected and tested before each use. GFCIs have a built-in test circuit to ensure that the ground fault protection is operational.

Figure 2. A portable GFCI can be used on a job site to protect workers. Image Courtesy of Cable Organizer

A ground fault circuit interrupter (GFCI) safeguards against the most common type of electrical shock hazard, the ground fault. Line-to-line contact hazards, such as a technician holding two hot wires in each hand, are not protected by GFCIs. GFCI protection is mandatory in addition to NFPA grounding requirements.

Author: Alex earned a master’s degree in electrical engineering with major emphasis in Power Systems from California State University, Sacramento, USA, with distinction. He is a seasoned Power Systems expert specializing in system protection, wide-area monitoring, and system stability. Currently, he is working as a Senior Electrical Engineer at a leading power transmission company.

Source URL: https://eepower.com/technical-articles/electrical-shock-and-its-effects/