PROTECTION OF DISTRIBUTION TRANSFORMER AGAINST ARISING OR TRANSFERRED FAST-FRONT OVERVOLTAGES: EFFECTS OF

SURGE ARRESTER CONNECTION CONDUCTORS LENGTH

Z. G. Datsios1*, P. N. Mikropoulos1, Z. Politis2, A. G. Kagiannas2 and T. E. Tsovilis2 1High Voltage Laboratory, School of Electrical & Computer Engineering,

Faculty of Engineering, Aristotle University of Thessaloniki, 54124 Thessaloniki, Greece 2Raycap Corporation, 14 Telou and Petroutsou, 15124 Athens, Greece

*Email: zdatsios@auth.gr Abstract: The effects of the length of the surge arrester connection conductors on the lightning surges impinging on a typical wood pole-mounted 50 kVA, 20/0.4 kV transformer of the Hellenic distribution system are investigated through detailed ATP-EMTP simulations. By considering both first and subsequent direct lightning strokes to a connected overhead distribution line, the effectiveness of the common practice transformer protection scheme and of an alternative one utilizing shorter surge arrester connection conductors in suppressing fast-front overvoltages was evaluated. A shorter length of the surge arrester connection conductors results in a reduction in the amplitude of the overvoltages arising at the medium-voltage terminals of the transformer and in a slower rate of increase of the overvoltage amplitude with lightning return-stroke current. The overvoltages transferred to the low-voltage terminals of the transformer are practically not affected by the length of the surge arrester connection conductors. Protection against transferred overvoltages was provided by surge protective devices installed at the low-voltage terminals of the transformer. By utilizing shorter surge arrester connection conductors the transformer failure rate, estimated through risk assessment, is reduced by approximately 11%.

1 INTRODUCTION

The fast-front overvoltages arising at distribution equipment utilizing non-self-restoring insulation, such as transformers or cables, may cause permanent failure resulting in system outages and economic losses. Therefore, distribution equipment is most commonly protected against impinging lightning surges by surge arresters and surge protective devices. It is well known that the efficiency of the afforded protection is affected by the length of the connection conductors; the latter should be as short as possible in order to achieve optimum protection [1-6].

According to common practice, distribution transformers are protected against impinging lightning surges by surge arresters installed close to their medium-voltage bushings. It can be shown theoretically that the fast-front overvoltages arising at a protected transformer increase in amplitude with the steepness of the incoming lightning surge and that the greater the separation distance between surge arresters and transformer the less is the effectiveness of the provided protection. Actually, analytical methods for the estimation of the lightning overvoltages arising at a protected transformer can be found in [7, 8]. However, as theoretical analysis is based on several simplifications, analytical results should be considered as conservative yet acceptable estimates for the protective distance and the safety margin provided by surge arresters. A more

accurate evaluation of the protection afforded by surge arresters to the transformer can be made with the aid of detailed simulations using an electromagnetic transient analysis program.

In this study the effects of the length of the surge arrester connection conductors on the fast-front overvoltages arising at a typical pole-mounted 50 kVA, 20/0.4 kV transformer of the Hellenic distribution system are investigated through detailed ATP-EMTP [9] simulations. By considering both first and subsequent direct lightning strokes to the connected medium-voltage overhead line, the common practice transformer protection scheme and an alternative one utilizing shorter surge arrester connection conductors were evaluated. A better lightning performance of the transformer is achieved by implementing the alternative than the common practice protection scheme; the transformer failure rate, estimated through risk assessment, is reduced by approximately 11%.

2 MODELLING OF THE EVALUATED SYSTEM

Figure 1a shows the common practice protection scheme of a typical wood pole-mounted 50 kVA, 20/0.4 kV transformer of the Hellenic distribution system. An alternative configuration is shown in Figure 1b; surge arresters are mounted on an additional wood crossarm, considerably closer to the transformer terminal bushings. The total length of the surge arrester connection conductors is 6 m

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and 1.1 m for the common practice and the alternative protection scheme, respectively. It must be noted that, in accordance with [5, 10], surge durable fuses coordinated with the surge arresters should be used for the alternative protection scheme. Both configurations were employed in detailed ATP-EMTP [9] simulations to assess the effect of the surge arrester connection conductors length on the fast-front overvoltages arising at the transformer. Figure 2 shows the schematic diagram of the system employed in simulations. A medium-voltage (MV) overhead line terminates at a 20/0.4 kV transformer. The transformer, sharing the same grounding system with the surge arresters, feeds through a short low-voltage (LV) overhead service line a 3-phase symmetrical load; the latter uses a separate grounding system.

Simulations were performed by assuming negative lightning flashes to the MV overhead line; as a worst case scenario, the overhead line terminates at the transformer (dead-end configuration). Lightning strikes to the outer phase conductor of

Figure 1: Wood pole-mounted 50 kVA, 20/0.4 kVdistribution substation; protection schemes of thetransformer: (a) common practice in the Hellenicdistribution system, (b) alternative configuration

Figure 2: Schematic diagram of the evaluated system

the line positioned at the greatest distance from the middle phase conductor (Figure 3), at the time instant of negative power-frequency voltage peak of the struck phase and at a distance of 100 m from the wood pole-mounted substation. This distance corresponds to the maximum span length along rural overhead lines of the Hellenic distribution system. Both first and subsequent lightning return-strokes to the overhead line were considered. Lightning stroke was represented by a current source producing a current with front upwardly concave [11]. According to CIGRE [11], the median values of the front time and maximum current steepness are a function of the lightning first return-stroke peak current; the median of the time to half value is equal to 77.5 μs. For the subsequent return-stroke current, the median values of the front time and time to half value are equal to 0.67 μs and 30.2 μs, respectively, and the median value of the maximum current steepness is a function of the peak current [11].

The MV overhead line (Figure 3) and the LV service line were represented by J.Marti frequency-dependent models considering line geometry. Line parameters were calculated for a soil resistivity of 200 Ωm. The distribution transformer (50 kVA, 20/0.4 kV, Dyn1, 4%) was modelled by a capacitance π-circuit together with a BCTRAN model. The 3-phase symmetrical load connected to the 20 m long LV line was simulated according to [12]. Surge arresters were represented by a frequency-dependent model [13] (Figure 4), with parameters calculated based on the surge arrester characteristics given in Table 1. Surge protective devices (SPD) were modelled as nonlinear current dependent resistors taking into account their

Figure 3: Overhead line (20 kV) of the Hellenic distribution system; wood crossarms support the porcelain pin-type insulators (BIL=125 kV)

Figure 4: Frequency-dependent surge arrester model [13]; parameters calculated based on the surge arrester characteristics given in Table 1

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Table 1: Surge arrester (SA) and surge protective device (SPD) characteristics

SA SPD

Rated1 and Nominal2 voltage: 121 kV 2240 V

Continuous operating voltage: 16.8 kV 300 V

Nominal discharge current: 5 kA, 8/20 μs 10 kA, 8/20 μs

High current impulse: 65 kA, 4/10 μs 200 kA, 8/20 μs

Residual voltage for nominal discharge current:

58 kV 800 V

characteristics (Table 1). Connection conductors with lengths shorter than 3 m were modelled as lumped parameter inductances of 1 μH/m [14]. Lengths longer than 3 m were modelled as lossless distributed parameters line segments with surge impedance of 400 Ω. Flashover of line insulation was modelled by employing the integration method [1, 3, 15], which was implemented in ATP-EMTP by the ISF object [16]. The latter is connected between phase conductors, since for the particular line (Figure 3) the flashover path with the lowest critical flashover voltage (CFO) is between phase conductors along the wood crossarm. The parameters of the integration method were determined according to [17, 18] considering the total CFO of line insulation; the latter was estimated respectively as 482 kV and 357 kV between left and middle phase conductors and right and middle phase conductors (Figure 3), by using the extended CFO-added method [19]. The arc channel following flashover of line insulation was represented by an inductance of 1 μH/m [20] for simplicity.

The concentrated grounding systems of the pole-mounted substation and the load were modelled as current-dependent resistances, considering thus, the decrease of the grounding impedance to values lower than the initial low current and low frequency grounding resistance caused by soil ionization. Weck’s simplified grounding system model, adopted by CIGRE [11], was employed in simulations, implemented in ATP-EMTP by the TGIR object [21]. The low current and low frequency grounding resistance was taken 10 Ω for the load and 100 Ω for the substation. Such high transformer grounding resistance values, resulting in more severe stress for the LV side of the transformer [22], have been reported in the Hellenic distribution system and associated with transformer failures due to lightning [23].

3 SIMULATION RESULTS AND DISCUSSION

3.1 Overvoltages arising at the MV terminals of the distribution transformer

Figure 5 shows typical waveshapes of the overvoltages arising at the MV terminals of the transformer due to lightning first and subsequent strokes to the MV overhead distribution line. There is a damped oscillation superimposed on the arising fast-front overvoltages, associated with

Figure 5: Typical waveshapes of overvoltages arising at the MV terminals of the transformer due to lightning (a) first and (b) subsequent strokes to the MV overhead line; 9 kA lightning peak current

voltage reflections due to the different surge impedances between the line and the transformer. This oscillation has a higher frequency and occurs earlier in a faster rising wavefront of the arising overvoltage for subsequent strokes due to the higher steepness of the subsequent return stroke current. The effect of such oscillations, occurring around the peak of an impulse overvoltage, on the dielectric behaviour of insulation is frequency dependent. The peak value of the oscillation affects gradually lesser the dielectric strength with increasing frequency; this effect becomes negligible for frequencies higher than approximately 3 MHz [24]. Thus, in order to compare the arising overvoltages with the basic insulation level (BIL) of the transformer, the “equivalent” overvoltage that the insulation is subjected to under lightning impulse voltage was estimated by adopting the procedure for manual calculation of impulse parameters from graphical waveforms according to [24].

Figure 6 shows the variation of the amplitude of the highest “equivalent” phase-to-phase, UMeq, and phase-to-neutral, VMeq, overvoltages at the MV terminals of the transformer with lightning peak current for both first and subsequent lightning strokes. Obviously UMeq is significantly higher than the VMeq. In addition, both overvoltages increase with increasing lightning peak current however with a different rate of increase depending on the length of the surge arrester connection conductors. A shorter length results in lower overvoltage amplitudes (up to 27%), especially with increasing lightning return-stroke current, and in a slower rate of increase of the overvoltage amplitude with lightning peak current. These effects are more pronounced for subsequent than first lightning strokes to the MV overhead distribution line.

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lent

ove

rvol

tage

pea

k, k

V

Lightning peak current, kA

UMeq common practice

UMeq alternative

VMeq common practice

VMeq alternative

BIL=125 kV

Figure 6: Equivalent overvoltages at the MV terminals of the distribution transformer as a function of lightning first and subsequent return-stroke peak current, depicted respectively by continuous and dashed lines

For the common practice protection scheme of the transformer, the “equivalent” overvoltage exceeds the corresponding BIL (125 kV) for direct lightning flashes to the MV line with first and subsequent return-stroke current higher than about 8 kA and 6 kA, respectively (Figure 6). It is important that for the case of the alternative configuration protection scheme the corresponding limiting currents are higher, about 17 kA and 12.5 kA for first and subsequent return-strokes, respectively.

3.2 Overvoltages transferred to the LV terminals of the distribution transformer

Transferred overvoltages to the LV side of the transformer are mainly associated with the potential rise of the transformer grounding due to the current flowing through the MV surge arresters and to the electromagnetic coupling between MV and LV transformer terminals. Figure 7 shows typical waveshapes of the overvoltages transferred to the LV side of the transformer when SPD are not installed at the LV terminals. There is a superimposed damped oscillation associated with voltage reflections due to the different surge impedances between transformer and LV service line. Thus, in order to compare the transferred overvoltages with the BIL of the LV side of the transformer (30 kV), the “equivalent” overvoltage that the insulation is subjected to under lightning impulse voltage, was estimated according to [24].

Figure 8 shows the variation of the computed amplitude of the highest “equivalent” phase-to-neutral overvoltage, VLeq, transferred to the LV terminals of the transformer with lightning current. The “equivalent” phase-to-phase overvoltages were omitted from this graph as they were found significantly lower than both VLeq and BIL of the LV side of the transformer. From Figure 8 it is obvious that VLeq, increasing with lightning peak current, is practically not affected by the reduction of the length of the surge arrester connection conductors. It is must be noted that for both protection schemes when SPD are not installed VLeq exceeds

Figure 7: Typical waveshapes of overvoltages transferred to the LV terminals of the transformer due to lightning (a) first and (b) subsequent strokes to the MV overhead line; 9 kA lightning peak current, SPD not installed

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with SPD

Figure 8: Equivalent overvoltages at the LV terminals of the distribution transformer as a function of lightning first and subsequent return-stroke peak current, depicted respectively by continuous and dashed lines

the BIL (30 kV) of the LV side of the transformer for lightning flashes to the MV overhead distribution line with first and subsequent return-stroke currents higher than about 5.5 kA and 4.5 kA, respectively. It is important, however, that the lightning stroke peak current causing a LV side transformer failure depends on load and transformer grounding resistances, increasing as the latter decreases [22]. Nevertheless, for the evaluated system there is a need for protection against transferred overvoltages by installing SPD at the LV terminals of the transformer. In this case, the transferred overvoltages are greatly reduced to values significantly lower than the BIL of the LV side of transformer (Figure 8). The amplitude of the transferred overvoltage, being primarily dependent on the characteristics of the SPD, is practically not affected by variations in lightning return-stroke current.

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4 RISK ASSESSMENT

The failure rate of a distribution transformer, R (failures/yr), due to direct lightning flashes to incoming overhead lines can be estimated as the sum of the failure rates due to lightning first and subsequent strokes, RF (failures/yr) and RS (failures/yr), respectively. By assuming no correlation between the first and subsequent return-stroke current amplitudes, RF and RS can be expressed as

F S FR n N L P I I (1a)

1S S F SR n N L P I I P I I (1b)

where n is the number of the lines connected to the substation and LS is the limit distance from the substation along the overhead line within which a lightning flash may cause failure to the transformer, that is, an incoming surge exceeding the BIL of the transformer. N (flashes/100km/yr) is the annual number of lightning flashes to an overhead line; for lines in open ground N can be estimated by (2) [19].

0.60.1 28gN N h b (2)

In (2) Ng (flashes/km2/yr) is the ground flash density, h (m) is the height of the uppermost conductor at the pole and b (m) is the separation distance between the outer phase conductors. It must be mentioned however that the estimation of N would depend on the lightning attachment model adopted for the evaluation of the lightning performance of the overhead distribution line. P(I>IF) and P(I>IS) are the probabilities of the prospective lightning return-stroke current being greater than IF and IS, respectively. IF and IS (kA) are respectively the minimum first and subsequent return-stroke peak currents of all possible lightning flashes that may terminate within LS causing a failure to the transformer. P(I>IF) and P(I>IS) can be estimated as

F

FI

P I I f I dI ,

S

SI

P I I f I dI (3)

where f(I) is the log-normal probability density function of the lightning peak current distribution given as [11, 19, 25]

2

2lnln

ln ln1exp

22

I If I

I

(4)

where Ī, σln are respectively the median value and the standard deviation of the natural logarithm of the lightning peak current distribution. According to [25], Ī=30.1 kA and σln=0.76 for first and Ī=12.3 kA and σln=0.5296 for subsequent return-strokes.

Table 2 summarises the minimum first and subsequent return-stroke peak currents causing failure to the distribution transformer, derived from Figure 6. This table also shows the corresponding

Table 2: Lightning stroke parameters causing transformer failure

Common practice

Alternative configuration

IF, kA 8 17

IS, kA 6 12.5

P(I>IF) 0.959 0.774

P(I>IS) 0.912 0.488

probabilities of the prospective lightning return-stroke current being greater than these values, calculated according to (3) and (4). Figure 9 shows the estimated failure rate of the distribution transformer due to direct lightning flashes to the MV overhead distribution line as a function of the limit distance. It is evident that the increase in failure rate with limit distance is more pronounced for the common practice than the alternative configuration protection scheme. Also, from Figure 9 it can be deduced that the reduction in the transformer failure rate due to the shorter surge arrester connection conductors is about 11%. As a rough approximation, for the evaluated system the transformer failure rate per km of limit distance due to direct lightning flashes to the incoming overhead line can be expressed as 0.1Ng and 0.09Ng for the common practice and alternative configuration, respectively. It is noteworthy that the contribution of subsequent strokes to the total transformer failure rate was found less than 12%.

It must be mentioned that as simulations were performed without considering corona damping effects on the lightning surges propagating along the overhead distribution line, the transformer failure rate shown in Figure 9 should be considered as upper limit. Moreover, the risk assessment results would depend on the lightning peak current distribution, which varies seasonally and geographically, and on the extent of shielding against direct lightning strokes provided by nearby structures in the region which the distribution line crosses [26].

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Figure 9: Failure rate of distribution transformer, R, due to direct lightning first and subsequent strokes to the MV overhead distribution line as a function of limit distance, Ls; Ng = 4 flashes/km2/yr, line height 8 m

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5 CONCLUSIONS

Detailed ATP-EMTP simulations have been performed to investigate the effect of the length of the surge arrester connection conductors upon impinging lightning surges on a pole-mounted 20/0.4 kV distribution transformer. Both first and subsequent direct lightning strokes to a connected MV overhead line were considered. A shorter length of the surge arrester connection conductors, utilized in an alternative to the common practice transformer protection scheme, results in a reduction in the amplitude of the overvoltages arising at the MV terminals of the transformer and in a slower rate of increase of the overvoltage amplitude with lightning return-stroke current. These effects are more pronounced for subsequent than first lightning direct strokes to the overhead distribution line. The overvoltages transferred to the LV terminals of the transformer are practically not affected by the length of the surge arrester connection conductors.

By implementing the alternative than the common practice transformer protection scheme the transformer failure rate, estimated through risk assessment, is reduced by about 11%. Surge protective devices provided the necessary protection to the distribution transformer against transferred overvoltages. To further improve the lightning performance of the transformer additional measures should be considered such as reducing the transformer grounding resistance and installing line arresters; this calls for further work.

6 REFERENCES

[1] R. L. Witzke and T. J. Bliss: “Surge protection of cable-connected equipment”, Trans. AIEE, vol. 69, no. 1, pp. 527-542, 1950

[2] T. J. Carpenter, I. B. Johnson and L. E. Saline: “Evaluation of lightning-arrester lead length and separation in co-ordinated protection of apparatus against lightning”, Trans. AIEE, vol. 69, no. 2, pp. 933-944, 1950

[3] R. L. Witzke and T. J. Bliss: “Co-ordination of lightning arrester location with transformer insulation level”, Trans. AIEE, vol. 69, no. 2, pp. 964-975, 1950

[4] S. S. Kershaw, Jr. and C. R. Clinkenbeard: “Discharge voltage of arrester connecting lead wires”, IEEE Trans. Power App. Syst., vol. PAS-93, no.1, pp. 226-232, 1974

[5] G. L. Goedde, L. A. Kojovic and J. J. Woodworth: “Surge arrester characteristics that provide reliable overvoltage protection in distribution and low-voltage systems”, in Proc. IEEE Power Eng. Soc. Summer Meeting, Seattle, USA, 2000, vol. 4, pp. 2375-2380

[6] F. D. Martzloff and K. Phipps: “Lingering lead length legacies in surge-protective devices applications”, IEEE Trans. Power Del., vol. 19, no. 1, pp. 151-157, 2004

[7] IEC 60071-2, Insulation co-ordination, Part 2: Application guide, 1996

[8] CIGRE-CIRED WG C4.402: “Protection of medium

voltage and low voltage networks against lightning. Part 2: Lightning protection of medium voltage networks”, Tech. Bul. 441, 2010

[9] Canadian-American EMTP Users Group: “ATP Rule Book,” 1997

[10] C. L. Smallwood, H. S. Regina and C. J. Cook: “Nuisance operations of distribution fuse links due to lightning-induced current surges”, IEEE Trans. Ind. Appl., vol. 43, no. 1, pp. 196-201, 2007

[11] CIGRE WG 33.01: “Guide to procedures for estimating the lightning performance of transmission lines”, Tech. Bul. 63, 1991

[12] H. K. Hoidalen: “Lightning-induced voltages in low-voltage systems and its dependency on voltage line terminations”, in Proc. 23rd ICLP, Birmingham, U.K., 1998, pp. 287-292

[13] P. Pinceti and M. Giannettoni: “A simplified model for zinc oxide surge arresters”, IEEE Trans. Power Del., vol. 14, no. 2, pp. 393-398, 1999

[14] IEEE Task Force: “Modeling guidelines for fast front transients”, IEEE Trans. Power Del., vol. 11, no. 1, pp. 493-506, 1996

[15] M. Darveniza and A. E. Vlastos: “The generalized integration method for predicting impulse volt-time characteristics for non-standard wave shapes-A theoretical basis”, IEEE Trans. Electr. Insul., vol. 23, no. 3, pp. 373-381, 1988

[16] Z. G. Datsios, P. N. Mikropoulos and T. E. Tsovilis: “Insulator string flashover modeling with the aid of an ATPDraw object”, in Proc. 46th UPEC, Soest, Germany, 2011

[17] W. A. Chisholm, J. G. Anderson, A. Phillips and J. Chan: “Lightning performance of compact lines”, in Proc. X SIPDA, Curitiba, Brazil, 2009, pp. 45-64

[18] A. R. Hileman: “Insulation coordination for power systems”, Boca Raton, Florida, CRC Press, 1999

[19] IEEE Guide for improving the lightning performance of electric power overhead distribution lines, IEEE Std 1410-2010, 2011

[20] M. Kizilcay and C. Neumann: “Lightning overvoltage analysis for a 380-kV gas-insulated line”, in Proc. IPST, Delft, The Netherlands, 2011

[21] Z. G. Datsios, P. N. Mikropoulos and T. E. Tsovilis: “Impulse Resistance of Concentrated Tower Grounding Systems Simulated by an ATPDraw Object”, in Proc. IPST, Delft, The Netherlands, 2011

[22] P. N. Mikropoulos, T. E. Tsovilis, Z. Politis and A. G. Kagiannas: “Evaluation of fast-front overvoltages arising at a 20/0.4kV distribution transformer”, in Proc. 7th MedPower, Agia Napa, Cyprus, 2010

[23] D. P. Agoris, A. Stamatelos, E. C. Pyrgioti, D. Vasileiou and S. Dragoumis: “Post mounted distribution transformer failures due to lightning correlating to the grounding resistance”, in Proc. 28th ICLP, Kanazawa, Japan, 2006, paper no. VI-26

[24] IEC 60060-1, High-voltage test techniques-Part 1: General definitions and test requirements, 2010

[25] Lightning and Insulator Subcommittee of the T&D Committee: “Parameters of lightning strokes: A review”, IEEE Trans. Power Del., vol. 20, no. 1, pp. 346-358, 2005

[26] P. N. Mikropoulos and T. E. Tsovilis: “Statistical method for the evaluation of the lightning performance of overhead distribution lines”, IEEE Trans. Dielectr. Electr. Insul., vol. 20, no. 1, pp. 202-211, 2013

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