1489670

©General Electric Company and Public Service Electric and Gas Company 2004.

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A Possible Mechanism for Neutral Grounding Resistor Failures

William Labos, and Antonio Mannarino, Member, IEEE

Goran Drobnjak, Member, IEEE, Satoru Ihara, Fellow, IEEE, and John Skliutas, Member, IEEE

GE Energy, Schenectady, New York

Abstract – The authors have found a mechanism that may account for recent failures of neutral grounding resistors in 26 kV substations owned by Public Service Electric and Gas Company (PSE&G). A neutral grounding resistor can fail due to a local, high-frequency oscillation involving the inherent inductance of the resistor and the transformer phase-to-neutral coupling capacitance. High voltages within the neutral resistor structure can be developed by the cable ring-down transients initiated by a single-line-to-ground fault on a 26 kV feeder circuit. This failure mechanism has not been previously reported in literature.

Keywords – neutral grounding resistor failure

I. INTRODUCTION

Resistance grounding is known to reduce point-of-fault damage, transient overvoltages, and the flash hazard in substations. It solves many of the problems associated with solidly grounded systems as well as those associated with ungrounded systems [1,2,3]. There are many neutral grounding resistors (NGRs) installed in 26 kV switching stations in the Public Service Electric and Gas Company sub-transmission system. A NGR limits the fault current and reduces the voltage dips during a single-line-to-ground (SLG) fault along a 26 kV feeder. However, the recently installed units are failing at a higher rate than expected. Digital fault recorder traces indicate that the NGR current abruptly changes 5 to 10 cycles after initiation of a feeder SLG fault and later disappears before the fault is cleared. The recorded NGR current values were below the rated value. None of the failed units had been tested as a full unit for the rated duty or for the impulse withstand capability.

This paper describes a possible failure mechanism and means to prevent future failures based on the ATP/EMTP simulation results.

“W. Labos and A. Mannarino are with Public Service Electric and Gas Company, Newark, NJ 07102-4194 USA. The content of this paper reflect the views of the authors who are responsible for the facts and accuracy of the data presented herein. The contents do not reflect the official views or policies of Public Service Electric and Gas Company.”

II. DESCRIPTION OF FAILED NGR UNIT

One of the failed NGR units, shown in Figure 1, consists of 180 resistor elements in two assemblies (neutral-side assembly and ground-side assembly) insulated from ground. Each resistor element is comprised of a flat stainless steel element spirally wound on a ceramic form. Three resistor elements are connected into a parallel set, and 60 parallel sets are connected in series. The neutral bus of the 26 kV

Neutral Bus

most stressedarea

Fig. 1. NGR Connection Diagram.

system is connected to the bottom parallel set of the neutral-side assembly. The neutral-side assembly has 11 resistor trays including the bottom seven trays and

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the top four with insulators between the two series groups.

The failed units were rated at 2,000 A, 8 Ω, for 14 seconds.

The inductance of the entire NGR unit is estimated to be 2.4 mH based on a 1 kHz bridge measurement on a single resistor element. This inherent inductance presents high impedance for high-frequency current (e.g. 125 kHz); and the stray capacitance of the structure and small capacitance of each resistor element determine the distribution of the high-frequency voltage stress along the structure. The area that is most stressed is between the neutral bus connection and the bottom tray of the neutral-side assembly (Figure 1). The design is based on the assumption of uniform voltage division by resistor elements but apparently, not for high-frequency voltage stress. The high-frequency voltage distribution can be improved and the basic insulation level (BIL) of the NGR unit will be substantially increased if the tray arrangement, tray grading, and connections are modified.

None of the failed units had been tested as a full unit for the rated duty or for the impulse withstand capability. Current IEEE Standards do not require such testing.

Upon examination, the ground-side assembly of a failed NGR unit looked clean and free of damage while the neutral-side assembly showed numerous burning marks made by high current as well as many spitting marks. One of the grounded corner steel posts showed burning marks and the neutral connection was blown open, indicating direct arcing to the ground involving high current on the order of 20 kA full short-circuit current (Figures 2 and 3). Direct arcing to the ground bypasses the current transformer for detection of the fault current. The spitting marks could have been made when the arc moved around and bridged some of the resistor elements.

Fig. 2. Burnt Resistors and Damaged Neutral Connection.

III. SWITCHING TRANSIENT STUDY

Figure 4 shows the 26 kV substation diagram. Four 230 kV/26.4 kV, 40 MVA step-down transformers are equipped with delta connected 11 kV tertiary windings. Two tap changers connected in parallel regulate the 26 kV bus. Five single and five double-cable circuits (fifteen 26 kV cables in total) exit the substation. The length of each cable is around 1,200 feet. Overhead lines extend from cable potheads into mainly residential, wooded areas. The 26 kV neutral bus is grounded through a NGR unit. A NGR breaker closes when the fault current is small and shunts out the NGR unit in order to increase the fault current for effective relay protection.

Fig. 3. Damaged Ceramic Forms.

230

KV

NGR

26.7 KVCables

CT

NeutralBus

Fig. 4. 26 kV Substation Diagram.


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In search of candidate mechanisms for the NGR failures, various circuit models were studied with different parameter values in the Alternative Transient Program/Electro-Magnetic Transients Program [www.emtp.org]. Figure 5 shows the study circuit model for a candidate mechanism identified and the major parameters for the base case simulation. The NGR unit is modeled with 8 Ω resistance in series with 2.37 mH inductance. The transformer is modeled with 2.3 nF capacitance in parallel with the Thevenin equivalence source for SLG faults (22.63 kV crest, 60 Hz source voltage behind 1.3 mH inductance estimated for four transformers). The 1,200 foot cables were modeled as distributed elements with 38 Ω surge impedance and 4 µsec travel time. No damping effects are modeled so that the highest theoretical overvoltage can be found.

1 Cable

14 Cables Fau

lt L

NG

R

Transformer

NGR: 8 Ω + 2.37 mH

Transformer: 1.3 mH, 22.63 kV, 2.3 nF 14 Cables: Z0 = 2.7 Ω, 4 µsec 1 Cable: Z0 = 38 Ω, 4 µsec Fault L: 0.4 µH/ ft

Fig. 5. Study Circuit Model – Base Case. Figure 6 shows the voltage transient across the NGR unit, initiated by a SLG fault at the cable pothead (Fault L=0). The maximum overvoltage across the NGR unit is 82 kV. The overvoltage magnitude decreases rapidly as the fault moves away from the cable pothead (Figure 7) and it is sensitive to the phase-to-neutral coupling capacitance value (Figure 8). The maximum NGR voltage does not exceed 83 kV. With system damping effects, the maximum NGR voltage is expected to be smaller. It could be around 50 kV and the NGR unit could have failed at this lower voltage.

(f ile G1R1.pl4; x-var t) v:N

0 4 8 12 16 20[ms]-100

-75

-50

-25

0

25

50

75

100

[V]

Fig. 6. NGR Voltage in kV – Base Case with Fault at Cable

Pothead. (No damping effects modeled.).

Fig. 7. NGR Overvoltage vs. Distance to Fault.

Fig. 8. NGR Overvoltage vs. Coupling Capacitance.

Fig. 9. Effect of Transformer Leakage Inductance.

The maximum NGR voltage is similar for two different transformer leakage inductance values corresponding to three and four transformer arrangements (Figure 9). The maximum NGR voltage remains similar as the length of the faulted cable is increased from 1,200 feet to 1,500 feet (Figure 10). The results in these two figures imply


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that regardless of the transformer parameter variations there is a matching cable length that can lead to undesirable NGR overvoltage when a fault occurs.

Two types of solutions were evaluated across the NGR: 1) a power capacitor and 2) a metal oxide varistor (MOV). A 16 kV, 300 kVAR power capacitor unit connected from the neutral bus to ground bypasses the 125 kHz current and effectively eliminates the overvoltage problem (Figure 11). The capacitor voltage overshoot is small because of the NGR resistance.

Fig. 10. Effect of Length of Faulted Cable.


0 4 8 12 16 20[ms]-100

-75

-50

-25

0

25

50

75

100

[V]

Fig. 11. NGR Voltage in kV with 300 kVAR Capacitor Solution.

A 24 kV rated MOV connected across the NGR unit limits the 125 kHz NGR overvoltage to 40 kV (Figure 12). This would be an alternative solution provided that the BIL of the NGR unit is 50 kV or higher. A 15 kV rated MOV limits the 125 kHz NGR overvoltage to 30 kV (Figure 13). Table 1 shows the MOV characteristics used. To determine the effect of a local ground fault on the NGR surge arrester, energizing into a local 10 MVAr ungrounded capacitor with a ground connected to the


0 4 8 12 16 20[ms]-100

-75

-50

-25

0

25

50

75

100

[V]

Fig. 12. NGR Voltage in kV with 24 kV MOV Solution.


0 4 8 12 16 20[ms]-100

-75

-50

-25

0

25

50

75

100

[V]

Fig. 13. NGR Voltage in kV with 15 kV MOV Solution.

TABLE 1

MOV CHARACTERISTICS Current (A) 15 kV MOV 24 kV MOV

1 22.5 kV 36.0 kV

10 23.8 kV 38.1 kV

100 25.7 kV 41.1 kV

500 27.9 kV 44.6 kV

1,000 28.5 kV 46.6 kV

2,000 30.6 kV 49.0 kV

15 sec TOV 15 kV MOV 24 kV MOV

without prior energy 17.3 kV rms 26.6 kV rms

with full prior energy 16.5 kV rms 25.3 kV rms

Phase C terminal was simulated (Figures 14, 15, 16). The calculated MOV energy duty is 0.5 kJ, much smaller than the 60 kJ capability.


5

NG

R

15 Cables

15 Cables

15 Cables

Phase A

Phase B

Phase C

Neu

tral

Bu

s

MO

V

Fig. 14. 10 MVAR Capacitor Energization with Ground

Connection on Phase C.

(f ile G1R5.pl4; x-var t) v:VA -XA v:VB -XB v:VC -

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5[ms]-50.0

-37.5

-25.0

-12.5

0.0

12.5

25.0

37.5

50.0

[V]

Fig. 15. Breaker Voltage in kV.

Energization of Capacitor Bank with Fault.

(file G1R5.pl4; x-var t) v:N

0 4 8 12 16 20[ms]-100

-75

-50

-25

0

25

50

75

100

[V]

Fig. 16. NGR Voltage in kV with 15 kV MOV

Energization of Capacitor Bank with Fault.

IV. CONCLUDING REMARKS

The probable failure mechanism is a local, high-frequency oscillation involving the NGR inductance and the transformer phase-to-neutral coupling capacitance. The NGR is excited by the cable ring-down transients initiated by a SLG fault on a 26 kV feeder circuit. This failure mechanism has not been previously reported in literature.

The frequency of the local oscillation changes with the transformer electrical characteristics. The coupling capacitance between the 26 kV bus to neutral is the critical parameter for verification of the failure mechanism conjectured. The frequency of the exciting cable transients changes with the cable electrical parameters (surge impedance and travel time) and fault location. The maximum NGR voltage is similar for two different transformer leakage inductance values corresponding to both a three and four transformer station arrangement. The maximum NGR voltage remains similar as the length of the faulted cable is slightly increased (by 25%).

The dielectric insulation of the failed units was not designed for high-frequency voltage stress. It could be substantially increased if connections are modified.

If the failure mechanism conjectured is valid, an effective solution is to connect a 16 kV, 300 kVAR power capacitor unit from the neutral bus to ground. Another solution is to determine the BIL of the NGR structure through an impulse test and to install a 24 kV or 15 kV surge arrester to protect it.

Upon completion of the study, a program was initiated to modify the NGR internal connections and to install a 24 kV surge arrester and a 300 kVAR power capacitor across each of the grounding resistors installed at several 26 kV substations. No plan is made for measurement of the coupling capacitance. Impulse test and rated duty tests will be performed in the near future. Several resistors of various types were retrofitted with a power capacitor and a surge arrester in parallel with the NGR and no recent failures have occurred.

REFERENCES

[1] Luke Yu and Rolf L. Henriks, ”Selection of System Neutral Grounding Resistor and Ground Fault Protection for Industrial Power Systems”, 91-CH3057-7/1/91/0000-0147, 1991 IEEE.

[2] Gary E. Paulson, “Monitoring Neutral Grounding Resistors,” 0-7803-5526-1/99, 1999 IEEE.

[3] Thomas Novak, Joseph Basar, Joseph Sottile, and Jeffery L. Kohler, “The Effects of Cable Capacitance on Longwall Power Systems,” 0-7803-7883-0/03, 2003 IEEE.


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William Labos is an Inside Plant Reliability Engineer with the Electric Delivery Asset Management Department of Public Service Electric and Gas Company (PSE&G). He has been with PSE&G for over 30 years and has held various technical and supervisory positions in both power generation and electric transmission and distribution. He received a BSEE from Newark College of Engineering in 1973, completed the General Electric Power System Engineering Course (PSEC) in 1995 and is a licensed Professional Engineer in the State of New Jersey.

Antonio Mannarino received a BSEE from New Jersey Institute of Technology in 1981 and completed the General Electric Power System Engineering Course (PSEC) in 1985. He joined Public Service Electric and Gas Company in 1981 and held various technical positions in the area of Switching/Substation Engineering and Equipment Standards. Mr. Mannarino has 23 years experience in the design, installation, operation and maintenance of electrical switching station and substations and is knowledgeable of various major equipment as well as power system studies. Part of his current responsibility is to manage and maintain the switching and substation assets. He is a member of the IEEE Switchgear Committee and is a registered Professional Engineer in the State of New Jersey.

Goran Drobnjak is a consulting engineer with GE Energy’s Energy Consulting business in Schenectady NY. His responsibilities include several different areas of power system analysis: electromagnetic transients, transient stability, and load flow. Mr. Goran Drobnjak received his BSEE and MSEE from University of Belgrade, Serbia and Montenegro. He is a member of the IEEE Power Engineering Society. He has authored several technical papers on different areas of power system analysis.

Satoru Ihara (F’01) recently retired from GE Energy where he was a senior research fellow at GE Energy. He received B.S.E.E. and M.S.E.E. degrees from Kyoto University, Japan and a Ph.D. in Applied Mathematics from Harvard University. Dr. Ihara is an IEEE Fellow and is past chairman of the IEEE load modeling working group and a member of IEE of Japan. He has authored over 40 papers on bus load modeling, series capacitor application, HVDC, fault current limiter, and power flow control.

John Skliutas is a consulting engineer with GE Energy’s Energy Consulting business located in Schenectady, NY. He conducts and manages transient, insulation coordination, and harmonic studies for utility, industrial, and GE client businesses. Mr. Skliutas received his BSEE from Worcester Polytechnic Institute, Worcester, MA and his MSEE from Rensselaer Polytechnic Institute, Troy, NY. He is a member of IEEE and has co-authored several IEEE and IEEE working group papers.


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