transit electromagnetic

IEEE Transactions on Power Delivery, Vol. 9, No. 4, October 1994

TRANSIENT ELECTROMAGNETIC INTERFERENCE IN SUBSTATIONS

C. M. Wiggins, D. E. Thomas, F. S. Nickel, Senior Member, IEEE

T. M. Salas Student Member, IEEE

BDM International, Inc. 1801 Randolph Road, S.E. A1 buquerque, NM 87106

Abstract - Electromagnetic interference levels on sen- sitive electronic equipment are quantified experimen- tally and theoretically in air and gas insulated substations of different voltages. Measurement techniques for recording interference voltages and currents and electric and magnetic fields are reviewed and actual interference data are summarized. Conducted and radiated interference coupl ing mechanisms and levels in substation control wiring are described using both measurement results and electromagnetic models validated against measurements. The nominal maximum field and control wire interference levels expected in the switchyard and inside the control house from switching operations, faults, and an average lightning strike are estimated using high frequency transient coupl ing models. Comparisons with standards are made and recomnendations given concerning equipment shielding and surge protection.

Keywords - Electromagnetic interference, EMI, switching transient, electric field, magnetic field, fault, lightning, substation, shielding, surge protection.

INTRODUCTION

Increasingly, electronic equipment is being used in switchyards and inside control rooms. Substation switching operations, spontaneous faults, or lightning strikes inside the substation, can cause potentially damaging levels of high frequency electromagnetic interference (EMI). This EM1 can couple into low voltage control circuits and electronic equipment unless it is suitably protected. This transient EM1 environment needs to be fully characterized by waveforms and spectra for the highest expected levels both in the switchyard and inside the control house. These EM1 environment levels may then be compared with equipment suscepti bil i ty levels (if they can be determined) for upset and damage, and also with applicable surge withstand capability (SWC) test levels to assess their mutual compatibility and adequacy.

Substation EM1 issues have been investigated in a number of studies, such as [l] - [7]. Switching transient currents, voltages and fields were measured in

S. E. Wright Senior Member, IEEE

Electric Power Research Institute 3412 Hillview Avenue Palo Alto, CA 94303

(now at Huddersfield Polytechnic School of Engineering

Huddersfield, England HD13DH)

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[l] - [6]. Potential EM1 impacts and suggested protection remedies on solid state relays were discussed in [l] and [7]. In [7], several types of EM1 sources were identified including the fast transient and the "walkie talkie" transient which have contributed to improved test standards [81,[91.

Investigations under project RP 1359-2 by Texas A&M for the Electric Power Research Institute [I] are particularly relevant to the work reported here. Some of the conclusions of this earlier work indicated: an expectation that radiated EM1 would become more important as new distributed automation systems were intro- duced in substations; a need for further analysis of the nature of radiated EM1 transients in substations; recognition that no appropriate standard existed for determining equipment susceptibility to transient electromagnetic fields, particularly for equipment lo- cated in switchyards; and a need for pre-purchase susceptibility testing. Improved equipment design along with the use of surge suppression devices and shielding enclosures was recommended to mi tigate EM1 effects.

In 1985 EPRI initiated project RP2674-1 with BDM International to further address these concerns. Un- der RP2674-1 there has been a strong attempt to broaden the substation EM1 environment characterizations and their understanding by developing validated high frequency traveling wave models as well as by gathering additional detailed measurements of EM1 phenomena in substations. This paper summarizes the major findings o f this study as reported in [lo]. The emphasis here is on presenting the highest expected EM1 levels in the switchyard, in the control house and on shielded control wires in substations up to 500 kV. These estimates are based on the results of all measurements and model predictions for switching transients and for faults and lightning strikes occurring in the substation. The manner in which EM1 couples from sources on the high voltage bus to wires inside shielded control cables is discussed qualitatively and quantitatively. Control wire EM1 levels are compared with oscillatory and fast transient SWC test waveforms (IEEE/ANSI C37.90.1-1989). EM1 fields in substation switchyards have now been characterized, but currently there is no standard with which to c o m a r e them. Possible test waveforms for switchyard fields and for control wires are discussed. 94 WM 146-1 PWRD by the IEEE Substations Committee of the IEEE Power

A paper recommended and approved

Engineering Society for presentation at the IEEE/PES 1994 Winter Meeting, New York, New York, January 30 - MEASURED SUBSTATION EM1 CHARACTERISTICS

Measurement Techniaues

A number of different types of transient measurements were required to completely describe first how typical substation EM1 arises and then how it couples

February 3 , 1994. Manuscript submitted September 1, 1992; made available for printing December 15, 1993.

0885-8977/94/$04.00 0 1994 EEE

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to other circuits. Switching a disconnector or circuit breaker, for example, produces a complex sequence of high frequency traveling wave current and voltage transients on each phase of the high voltage bus. Travel- ing wave bus current transients excite the three- dimensional bus structure which acts as a complex an- tenna, radiating energy into the substation as transient electric and magnetic fields. Bus current transients can also couple into low voltage circuits that are connected directly to the bus. The net transient EM1 at some point in the substation is therefore the superposition of both conductive and radiative coupling components. The net transient electric and magnetic fields at a point include contributions from the ground and from all three bus phases, as well as scattering from nearby conducting structures.

In general, the required types of transient EM1 measurements included currents on the high voltage bus, electric and magnetic fields at various locations on and above ground, and currents and voltages on substation control wiring. Switching operations, faults, and flashovers from lightning occurring on the high voltage system all cause abrupt arcing discharges between affected substation conductors when their potentials are sufficiently different. Peak transient current amplitudes produced during arcing depend on the surge impedances of the conductors and the peak instantaneous phase-to-ground system voltage. Voltages in transmission substations generally range between 115 kV and 500 kV. Surge impedances over this voltage range also tend to be relatively constant at approximately 350 ohms. As a result, arcing discharges produce traveling wave currents proportional to the substation voltage. Since field and control wiring EM1 originate from the transient bus currents and voltages, their relative amplitudes are also generally proportional to substation voltage. Therefore, the levels of EM1 in substations should be expected to be higher at higher substation voltages.

Overall risetimes of transient EM1 phenomena caused by arcing discharges on the high voltage system are governed by the effective charging time constants of the circuit driven by the arc, typically the entire substation bus structure, and are on the order o f 200 nanoseconds or more in air-insulated substations (AIS). Because gas insulated substations (GIs) have relatively small dimensions, EM1 risetimes can be up to 10 times faster than in A I S . Bus charging times should generally be expected to increase as the physical size of the circuit excited by the arc increases, and vice- versa.

In short, substation EM1 measurements were found to require:

(1) wideband sensors for measuring bus current tran- iients at levels from 300A to lOkA, electric and magnetic fields up to 100kV/m and 300 A/m, and control wire transients of 10 kV and 100 A peak amp1 i tudes.

(2) at least four data channels to correlate different EM1 measurements simultaneously.

(3) each channel should have a bandwidth of at least 100 MHz and a dynamic range of at least 45 dB.

( 4 ) 100 m to 200 m long, remotely controllable, analog fiber optic data links for immunity from EM1 environment and installation flexibility.

(5) high speed, large memory transient digitizers to record individual transients at high resolution as well as bursts of consecutive transients from restriking switches.

selected events including the most significant.

and control system.

(6) versatile triggering t o ensure capture of

(7) fast, large capacity data acquisition, processing

For more detailed discussion of these specific requirements, and the mobile transient EM1 measurement system developed to meet them, see [lo] - [12]. Characteristic EM1 Data

During project RP2674-1, detailed transient EM1 measurements of bus currents, electric and magnetic fields, and control wiring currents and voltages were made in 12 separate tests at 8 different air and gas insulated substations at voltages of 115 kV - 500 kV addressing the requirements and using the techniques outlined above. In all, over 800 separate events of up to 4 transient measurements per event totaling 700 megabytes of data were recorded during the tests occurring between 1985 and 1991. Many results from these tests have already been reported. For example, switching transients measured in 115 kV substation are reported in [5]. Many detailed attributes and characteristics of switching transient electric and magnetic fields in 115 kV, 230 kV, and 500 kV A I S and in 230 kV and 500 kV GIs are summarized in [ 6 ] . A complete discussion of the results o f all types of substation EM1 measurements of staged faults and switching transients, including line energizations, can be found in [13] and [14]. These references also describe the details of the substations, the switching operations performed, and the various measurement configurations used to record the data.

Here we present only representative data to characterize switching transient EMI, emphasizing measurements in 500 kV A I S and GIs where the levels were found to be the highest. Waveshapes of bus currents, fields, and control wire EM1 at lower voltage substations are the same as those provided here; however, the dominant frequency components tend to increase somewhat as the substation voltage decreases r 6 1 ~ 1 3 1 .

When high voltage switches are operated in AIS, bursts of up to 5,000 to 10,000 individual transients varying 60 dB or more in amplitude are typically produced during the complete arcing sequence as the switch contacts open or close. Similar bursts, but of fewer total transients, occur under similar conditions in GIs. These complex sequences of events during a switching operation give rise to an overall macro- scopic, or macroburst, characteristic of a switching transient. Depending on the type of switch and how fast it operates, macrobursts may last from 40 ms up to 2 seconds. Any single arcing component of the macroburst produces one transient having a specific waveform and i s termed a micropulse. The specific micropulse waveform is mostly determined by the length of excited bus and its traveling wave properties [IS] and typically has a time-varying amplitude duration of no more than 10 to 15 microseconds in A I S and less than 4 microseconds in GIs. The most severe transients, those of maximum burst size and amplitudes, occur during operation of relatively slow hand-cranked disconnect switches. While they do generate much higher frequency components than disconnect switches, circuit breaker operations were found to produce much fewer and lower amplitude transients [6] .

Within a macroburst, the highest amplitude micropul se transients are those that are produced when the switch contacts are as widely separated as possible and still arc. Such a condition only occurs once per half-cycle (at 60 Hz), but can occur many times over consecutive half-cycles for a switch with relatively slowly moving contacts. Thus, a substantial number (20 or more) o f these highest amplitude transients can be produced during a single switch operation, depending on its speed. When characterizing the maximum EM1 en-

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vironment of a particular substation switching operation, it is necessary to measure these highest ampl itude transients. During disconnect switch operation, these highest amplitude micropulses occur first on closings 'and last on openings.

Representative micropul se switching transient EM1 characteristic of the highest levels found in 500 kV substations are presented below. Since the highest EM1 levels were recorded in 500 kV substations, these levels are also representative of the highest levels found in all substations investigated in this project. This is generally true of bus current, field, and control wiring transients. However, peak field amplitudes can vary significantly within a substation due to a number of factors as discussed below.

Electric and magnetic fields are vector quantities; different orthogonal components have different peak ampl i tudes. Normally one component will dominate the other two in amplitude when the field is measured relatively close to its source. Thus, the vertical electric field between the bus and ground and the magnetic field component perpendicular to the bus and horizontal to the ground usually have the highest amplitudes. In this study, these polarization components were measured on the ground directly beneath the section of excited bus to standardize the measurement at all substations for comparison purposes, but is an arbitrary choice otherwise. Amplitudes of electric and magnetic fields, which generally increase with in- creasing substation voltage for a given polarization component and measurement geometry, will a1 so increase strongly as the distance between the field point and the excited bus (source) decreases. As a result, electric and magnetic field amplitudes at other locations within the substation can easily be much higher or much lower than those referenced to the ground directly below the excited bus. Moving the field measurement point only a few meters closer or farther away from the bus relative to the arbitrary ground reference location can easily cause peak field amplitudes to change by a factor of 2 or more [6]. For this reason, the peak electric and magnetic field ampl itudes reported below are only representative of those on the ground (approximately 8 meters below the excited bus) in 500 kV substations; higher amplitudes could be measured closer to the bus.

Bus Current Transients. Figure l(a) shows the sensor installation used to measure the transient bus currents from switching operations in a 500 kV AIS. The sensor and fiber optic transmitter can be seen on the closest phase conductor about midway between the column CT on the left side of the photo and left side of an .air- break motor-operated disconnect switch on the right. The fiber optic cable providing remote control to the transmitter and bringing the bus current transient signal to the receiver inside the measurement van is the vertical line seen dropping from the transmitter. The fiber optic cable also provides the required electrical isolation from the 500 kV bus potential.

A bus current transient (typical of the highest ampl itude bus transients) produced during a disconnect switch operation that excited a short section of 500 kV bus is shown in Figure 2. The zero-to-peak amplitude and risetime are about 2.3 kA and 400 nanoseconds. The transient damps out to zero amplitude in about 10 to 15 microseconds; it has a dominant frequency component near 0.5 MHz, with other significant components up to about 3 MHz.

Electric and Mametic Field Transients. Figure l(b) shows the location of the vertical component electric field sensor (on the left) and horizontal component

Bus current transient measurement configuration.

(b) Ground-plane electric and magnetic field transient measurement configuration under bus.

Figure 1. Switching Transient Measurement Geometry in 500 kV AIS

magnetic field sensor (on the right) on the ground directly below the location of the bus current sensor shown in Figure l(a). The magnetic field transient shown in Figure 3(a) was measured at the same time as the bus transient measurement shown in Figure 2 and has a zero-to-peak amplitude of about 92 A/m. Note the strong similarities in the bus'current and magnetic field waveforms and spectra. This is because this component of the magnetic field is directly proportional to the current on the bus,

H = l / (xh) 9

where H is the magnetic field in A/m, I is the bus current in A, and h is the height of the bus above the ground (7.58 m). The factor of 2, normally present in the denominator of this expression, does not appear

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+ Q

L U

ul a m

-6 -2 .50 ' ' ' ' ' ' ' '

. 0 2 . 8 4.0 6 . 0 0 . 0 10.0 *le

Time (seconds)

Figure 2. Measured Bus Current Transient from Opening a Disconnect Switch in a 500 kV AIS. (Event C2.628)

8.9 10.0 nI0-6 .e 2.0 4 . 0 6 .0

Time (seconds)

(a) Magnetic field transient.

C4.628 Open DSS4818 Uatleu Sub S08kU OCD-60 10"s +6dB

c, y 2.6u .3 ... ....................................................................... c ~ , , , ~ , , , , , 1 , , , , , , , , , 1 , , , , ~

-6 -2.0

.e 4.0 8 . 0 12.0 16.0 20.0 -10

Time (seconds)

(b) Electric field transient (early-time only).

Figure 3. Magnetic and Electric Field Transients Measured on the Ground from Opening a Disconnect Switch in 500 kV AIS (Event C3.628)

since the horizontal component ot the magnetic field amplitude tends to double upon reflection from the ground at normal incidence.

The electric field transient, measured simultaneously with the bus current and magnetic field,is shown in Figure 3 (b). The electric field rises to a peak amplitude of about 13.2 kV/m in about 1

microsecond. The electric field transient exhibits a waveshape that is characteristically different from the magnetic field and bus current transients. This is because the electric field is proportional to the charge on the bus, i.e., the time integral of the bus current. Since the charge is equal to the product of the peak phase-to-ground voltage (408 kV for 500 kV system) and the capacitance of the transmission line (the latter being further related geometrically to the surge impedance), the vertical electric field below an excited bus can be expressed as [16]:

377 E = - F [ z, h]

where E is the electric field in V/m, V is the phase-to-ground voltage, 2 , is the surge 'Qmpedance (about 328 ohms for this location), and h is the bus height above the electric field sensor. If the simple, single transmission line over a ground plane is used to calculate the line capacitance or surge impedance, rather than the more complicated (and exact) three phase model, the simple expression given above overpredicts the observed peak electric field somewhat. Again, the factor of 2, normally present in the denominator of the E-field expression, has been suppressed since the vertical electric field component tends to double upon ground reflection. The electric field waveform is characterized by damped oscillatory and quasi-static components. The oscillatory behavior is due to the damping of the time-varying charge, whereas the quasi-static component results from the fact that the switched section of isolated bus is left in a trapped charge (or non-zero voltage) state until the next transient (arc) occurs. For this reason, electric fields do not damp to zero amplitude over time; rather, they step from one quasi-static voltage state to another at the rate at which arcs occur during the switching operation. Since the rate at which arcs occur can vary from 40 kHz to 120 Hz during disconnect operations, the duration of the electric field quasi- static component steps can vary from 25 microseconds up to about 10 milliseconds. The highest amplitude electric field transients, such as shown in Figure 3(b), occur at a repetition rate 120 Hz, and therefore typically have durations of 10 milliseconds. The quasi-static component of the last electric field transient produced on opening a disconnect can obviously persist for longer than 10 milliseconds. Examples of quasi -static (1 ate-time) electric field measurements may be found in [6].

Representative electric and magnetic fields measured on the ground beneath the gas enclosure near the gas/air bushing in a 500 kV GIs produced by a disconnect switch operation are shown in Figure 4. Four things are readily apparent when comparing GIs fields with those of AIS at the same voltage and for the same type of switching operation. First, the principal frequencies of GIs transients are at least 10 times hi her than those found in AIS, undoubtedly due to the smafler substation dimensions. Second, the peak field amplitudes in GIs are somewhat lower (E-field a factor of nine, H-field a factor of two) than those of AIS; the gas enclosure probably acts as a shield. Third, the durations of the GIs transients are much shorter than those o f AIS. Fourth, the GIs electric field no lon er has a quasi-static component and damps to zero ampqitude as fast as the magnetic field. Both of the last two observations are probably due to the fact that the gas enclosure (the source of most of observed transient fields) was grounded at many different locations in these substations.

Control Wirinq Current and Voltaqe Transients. Figure 5 presents control wire data representative of the highest levels measured on terminal strips near relay

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H < t > < W m > Event C3.523

20.0

10.0

-10.0

-20.0

-30.0 .e a.0 4.0 6.0 8 .0 10.0 xle-.

Time (seconds)

(a) Magnetic field transient.

E < t > Event C4.523

Time (seconds)

(a) Control wire voltage transient.

(b) Control wire current transient. (b) Electric field transient. .

Measured on Ground Under Gas Enclosure Near Gas/Air Bushing from Opening a Disconnect Switch in a 500 kV G I s (Event C3.529) disconnect switching).

Figure 4. Magnetic and Electric Field Transients Figure 5. Control Wire Transients Measured Between C Phase CT Wire and Ground on the Terminal Strip Near Relay in a 500 kV A I S Control House (Event C2.643, from

equipment inside of the control house of a 500 kV A I S . A (zero-to-peak) voltage of about 4 kV measured between a CT wire and ground is shown in Figure 5(a). This transient rises to peak in about 200 nanoseconds and continues to ring at low amplitude for up to 25 to 30 microseconds. Figure 5(a) shows significant frequency components as high as 20 MHz. The corresponding current transient measured at the same time i s shown in Figure 5(b). The current on the C phase CT wire rises to a peak o f 10.5 A in about 3 microseconds (it actually rises to an initial, slightly lower, peak of 9.5 A in 700 nanoseconds) and damps out in 25 - 30 microseconds. Major frequency components occur near 0.25, 0.50, and 1.0 MHz. The frequencies of the two transients are very different. This suggests that the impedance varies with frequency.

Summarv of TvDical Highest Measured EM1 Levels In the substations examined in this project, EM1

levels were found to be the highest in 500 kV substations (the highest tested). Disconnect switching was found to produce higher amp1 itude transients than circuit breaker switching. Table I summarizes how peak EM1 levels were found to scale at other voltage levels of AIS for both disconnect and circuit breaker switching. The peak levels reported are nominally the highest observed for each type of transient at substation voltages of 115, 230, and 500 kV, that is, they are averages over many measurements made at each sta-

tion under exactly the same measurement conditions while trying to record the highest EM1 levels produced by a particular switching operation. Table I summarizes the peak EM1 levels measured for bus current transients, principal component electric and magnetic fields on the ground beneath the excited bus, and field-induced current on a 26.5-111 open circuited test cable. This latter measurement was performed to aid in the validation of the coupling models described in [17]. In Table I, measured bus currents, magnetic fields, and test cable currents are reported in terms of their peak-to-peak values. Zero-to-peak amplitudes may be estimated (because of the asymmetry of the waveforms) by taking 70% of the bus current and magnetic field peak-to-peak values and 50% of the reported f i el d-driven cab1 e shield peak-to-peak currents. Electric field amplitudes are already quoted as zero- to-peak, since their waveforms are unipolar. Peak fields for several locations above ground are also reported in Table I and indicate how amplitudes increase closer to the bus. Peak levels of control wire current and voltage transients measured in all' substations are summarized in [17] and in Table V. Generally, EM1 levels were found to scale linearly with substation voltage, thus EM1 levels in substations at voltages above 500 kV are expected to be higher than those reported here.

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SYSTEM BUS ELECTRK: MAGWETIC VOLTAGE CURRENT FIELD FIELD

(hV/mJ ( h l 0 (4

Table I

SUMMARY OF NOMINAL PEAK AMPLITUDES OF VARIOUS EM1 TYPES AS FUNCTIONS OF SYSTEM VOLTAGE AND SWITCHING OPERATION

(measured, peak-peak)

TESTCABLE Open Circuiied

(AI

115

230

488 'h-d88m 15.4 h - 2 3 m 56.1 h-2.3m 14.0 h-l.Om 45.2 h-l.Om

1040 h - L l 8 m 82.2 h-2.3 m 7.0 h-Om 36.7 h-Om 25.8 h-Om

I 5.5 h-Om 70.4 h-Om I 52.3 h-Om

PREDICTIVE SUBSTATION EM1 MODEIS

115

230 500

Model Requirements

Substation EM1 models were developed to aid in understanding the measurements and to provide a means for estimating EM1 levels beyond measurement limitations. The goal for models was to replicate all details, including waveshape and peak amplitudes, of the transient EM1 measurements. Since fields arise from transient bus currents, and since both can produce transient EM1 in control wiring, models describing each different type of substation EM1 were required. Thus, traveling wave models of air and gas insulated substations were developed to predict transient bus currents and voltages and their radiated fields as a function of a given excitation (switching operation, fault, lightning strike). The outputs of the bus current and field models were designed to be the inputs to various conducted and radiated cable coupling models. In this way, control wire current and voltage transients were linked back to their causes and predicted.

Initial conditions for the source excitations on the bus determine the levels of all resulting EM1 produced. Disconnect switching transient calculations were often made for the condition that twice the peak phase-to-ground voltage appears initially across the switch gaps. This situation can occur when switching an isolated section of bus,,with trapped charge. ;n this paper, the statement; 2PU !nitial condition , 2PU switching transient , and 2PU prediction" all

refer to this condition, not to the amplitude of a transient that may result from it (which could be 3PU or higher). In this way estimates of EM1 produced from, for example, 2PU initial conditions across a switch, or from a 10 kA lightning strike to the bus, were made relatively easily using the models, whereas their measurement under precisely these conditions could be very difficult. All the EM1 models were tested to verify their ability to generate known results before using them to make estimates in regimes where comparisons with test results were not possible. Model details and their validations are discussed in a companion paper [17], and in [lo] and [18].

EM1 Control Wirinq Couolina Modes

Important conductive and radiative mechanisms for coupling substation EM1 into shielded and unshielded CT and CCVT control wiring have been investigated and quantified in some detail using measured data and transfer functions [17]. These couplinq mechanisms are

11.8 h-4.88m 9.1 h-2.3m 3.7 h-Om 0.62 h-Om 0.36 h-Om

76.8 h - C l 8 m O.9h-Om 5.3 h-Om 2.4 h-Om ,132 h-8.3Bm 5.6 h-Om 18.9 h-Om 5.3 h-Om

briefly reviewed here to illustrate how estimates of control wire EM1 levels presented later are the sum of contributions from each distinct coupling mode. Each mode contributes only a portion of the total control wire transient amp1 i tude and frequency content.

Conducted EM1 CouDlinq. When a shielded or unshielded control cable connects to an EHV CT or CCVT, there is a deliberate coupling of the control circuit to the hiah voltaae circuit. At sufficiently high frequencies, the CT and CCVT becomes conductively coupled to the high voltage bus via, e.g., the parasitic capacitance between the primary and secondary and their Faraday shields. When this coupling mode is present, a portion of the bus current transient couples directly to the conductors inside the shielded or unshielded CT and CCVT cable. This conducted coupling mode is particularly significant because it is not reduced by shielding the control cable. This type of coupling can be reduced using surge suppression devices.

Radiated EM1 Couplinq. In unshielded cables, high frequency radiated electric and magnetic fields couple directly to individual control wires and produce current and voltage transients at cable loads with little or no attenuation. Transient field coupling to wires inside shielded cables is a two step process. In the first step electric and magnetic field coupling induces voltages and currents to flow on the shield. The tan- gential component of the electric field over a short length of the shie)d acts as a local voltage source driving the cables impedance [19]. Similarly, the component of the magnetic field normal to the effective loop area formed by the cable shield and any nearby impedance-coupled conductors (e.g. CT and CCVT ground straps and their grounded pedestals), causes currents to flow in the loop [lo], [17]. Once field-induced current transients are present on cable shields, they can then couple onto the cable conductors. Shield-to- conductor coupling can occur several ways. One way is via the transfer impedance that exists between the shield and each wire [19]. The other way is through the mutual inductance between the shield pigtail and a conductor at the shield terminations [17]. The effectiveness with which EM1 couples to the conductors from both o f these coupling modes increases as the frequency increases, particularly at frequencies above about 1 MHz. Studies [17], [ZO], [21] indicate that pigtail coupling will be minimized by making the pigtail length parallel to the conductors as short as possible. Transfer impedance coupling can be minimized by using high-quality cable shields, but there is little ad- vantage in providing any better cable shielding than is required to reduce transfer impedance coupling to a level below that of pigtails.

When summed, the three control cable EM1 coupling modes just discussed were found to generally account for the observed control wire EM1 voltage and current waveforms and levels. It is interesting to note that, individually, conducted and pigtail coupling were found to each contribute up to 70 % of the observed coupling [lo]. However, both of these when added together with the 20% transfer impedance contribution, produced 100 X of the total wire EMI. The reason for this was that each different coup1 ing mode contributed EM1 components with different waveshapes, frequency content, and relative phase.

ComDarina Model Predictions With Measurements

In [17] and [18], overlays of predicted waveforms with those actually measured for transient bus currents, electric and magnetic fields, and control wiring currents and voltages for substation disconnect switching unambiguously illustrate the capabilities of the models described above. Here, we simply compare measured peak amplitudes with 1PU and 2PU predictions.

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(kV) 115 230

The measurement and prediction geometries are identi - cal. Bus current transients are nominally referenced to the center of the section of bus being excited; fields are referenced to locations on the ground directly below the bus sensor; control cable geometry and coup1 ing (including source impedances) are represented by the actual CT and CCVT cable installations found i n the 500 kV substation measured and modeled; and control wire current and voltage transients are referenced to the inputs of the actual load impedances measured inside the control house [17]. Predictions for 230 kV and 345 kV substations are based on linear extrapolations between 115 kV and 500 kV models, since the former substation voltages were not modeled.

Table I1 compares predicted and measured bus current transients produced by disconnect switching in 115 kV - 500 kV AIS. For each substation voltage, measured bus current peak amplitudes are given both for the average of all peak amplitudes measured at the substation and for the highest peak amplitude measured there. Comparing the two types of measured results may indicate that, while capture of the highest amplitude transient was always attempted, it was rarely success- ful. In general, measured bus currents tend to fall between the 1PU and 2PU model predictions and indicate, that on average, somewhat less than 2PU is measured. Peak measurements are much closer, but somewhat less than, the 2PU predictions for all substation voltages. This indicates that the 2PU model predictions provide a reasonable upper bound to the measured transient bus current level s.

(") (Nm) (Nm) (Ah) 19.2' 26.2 34.7 38.4' 40.2" 45.3 54.4 80.4"

Table I1

COMPARING 1PU AND 2PU PREDICTED WITH MEASURED BUS CURRENT TRANSIENTS FOR SEVERAL SYSTEM VOLTAGES

(zero-peak)

I I System (1 PU (Average (Peak Voltage 1 Predicted) I Measured) I Measured) I P s z d ) (kv) (A) (A) (A) (A) 115 295' 330 -496 590' 230 735 860 1203"

** - Interpolated

Tables I 1 1 and IV provide similar comparisons for magnetic and electric field transients produced by the bus currents. Magnetic field predictions at 2PU initial conditions are seen to just bound the peak measured in all substations. Electric field comparisons are presented slightly differently. Predicted and measured amplitudes are reported for both the peak as well as the DC, or late-time (quasi-static), level since they differ substantially. The peak electric field is governed by the transient charge re- distribution on the bus caused during arcing, and is further affected by the way the excited bus causing the field is loaded (or terminated). The DC (or quasi- static) electric field is governed by the phase-to- ground bus potential. Also, a separate column of peak measured electric field is not provided; the minimum and maximum values are indicated as adjustments to the average peak amp1 itudes. Peak measured electric fields are seen to fall much closer to the 1PU model predictions. This suggests the possibility that either the measurements are too low or the model predictions are too high. Both possibilities have been considered, but not resolved at this time. For example, it has been shown in [6] that measured electric field amplitudes

Table 111

COMPARING 1PU AND 2PU PREDICTED WITH MEASURED MAGNETIC FIELD TRANSIENTS FOR SEVERAL SYSTEM VOLTAGES

(Horizontal component fields measured on the ground bel ow excited bus)

(zero-peak)

System I (1 PU I (Average I I Voltage Predicted) Measured) I

~~

I 345 I 61.2'. I 69.0 I 69 I 122.4" I _ _ _ _ - .. 5 0 0 1 89.5' I 97.4 1 131.5 I 179.0' * - Predicted by TRAFIC model *. - Interpolated

Table IV

COMPARING 1PU AND 2PU PREDICTED WITH MEASURED ELECTRIC FIELD TRANSIENTS FOR SEVERAL SYSTEM VOLTAGES

(Vertical component fields measured on ground below excited bus)

(zero-peak)

actually double at a height of just 1 meter above ground. At 1-meter above ground 2PU predictions and measurements agree much better. This suggests that perhaps ground plane electric field amplitudes were suppressed somehow. Since vertical metallic structures (insulator posts, switchgear supports, etc.) were often within a meter or so of the measurement location, it is possible that they could have produced scattered fields which tended to lower the vertical field component measurement. In [6], enhancements of the electric field strength by more than a factor of 4 were reported near grounded structures protruding 2 meters above an otherwise flat ground plane. An insulator bus support column is such a structure that protrudes even higher; perhaps the field near the top of these posts is greatly enhanced and reduced at locations further away, such as where sensors were placed.

Alternatively, only a single phase of the three phase bus was modeled. It may be possible that the presence of other two phase conductors could change the line capacitance (surge impedance) enough to lower the electric field. Such a tendency was noted earlier during the discussion of the simple scaling formula for calculating the peak electric field. Also, the models currently ignore bus supports and possible effects caused by them. For now, the matter is unresolved. Using the data in table IV, the 2PU predictions may be overpredicting ground-plane electric fields by almost a factor of 2. On the other hand, when comparing 2PU predictions with peak measurements at heights of 1-m and 2-m above ground the relative agreement is much

1876

closer and, of course, the field ampiitudes are much higher. It is interesting to note that magnetic field measurements and predictions were found to agree closely at all locations, and showed no tendency to double in amplitude over their ground-plane value at just 1-m above ground.

Table V compares predicted and measured voltage and current EM1 transient peak amplitudes on conductors inside partially-shielded CT and shielded CCVT control cables at a point near the equipment load in a 500 kV control house. To validate model predictions, it was necessary to use actual frequency-dependent load impedances. This is because common-mode and differential-mode measurements at the inputs of a solid state transformer differential relay reported in [20] have shown that relay impedance can vary over two or- ders of magnitude (from 10s to 1,000s of ohms) over a frequency range from 100 kHz to 100 MHz, and gives different frequency responses dependent upon the input measured. The CT inputs for the relay in [20], for example, while having an average impedance of a few hundred ohms over all frequencies, exhibited a high Q impedance resonance of about 2,000 ohms at a frequency of about 1 MHz. Most control wire switching transients observed during this project have strong spectral components near this frequency and could be strongly affected by it. In general, load impedances are complex functions of frequency. This point was made evident during the tests by observing the dramatically different wire transient current and voltage waveforms that were simultaneously measured at the rack terminal block near the relay [lo]; but, since these measurements were not made at the relay inputs, one cannot conclude that observed impedance variations represent actual relay impedances, some of which were protected with surge suppression devices. The effect of measured frequency-dependent loads on predicted wire EMI, is discussed further in [lo] and 1211.

TYPE OF PREDICTION CONTROL CIRCUIT 1 p u I 2PU

MAXIMUM MEASURED

CT(1)

ccvr(2)

For model validations, the effective frequency- dependent load impedance was obtained by calculating the ratio V(w)/I(w) where V(w) and I(w) are the Fourier integral transforms of the simultaneously measured voltage and current waveforms, respectively. A smoothed fit for the calculated impedance was developed and used for the model validations. Once models have been validated, open or short circuit impedances can be used to generate the extreme load conditions for control wire EMI.

5.9A 11.8A 10.3 A 2.1 KV 4.2 KV 3.9 KV 0.48 A 0.96 A 0.81 A 1.76 KV 3.4 KV 2.4 KV

Table V compares 1PU and 2PU model predictions to the maximum measured CT and CCVT control wire current and voltage transients for a 500 kV substation. In each case, the model predicts the total wire transient EM1 from all important coupling modes based on the actual cable configurations (see Table V footnotes). For both CT and CCVT cables, the 2PU initial condition only slightly overpredicts the maximum measured control wire EMI, indicating that this model provides a reasonable upper bound in terms of peak amplitudes. Had the load been an open circuit, the predicted voltage would have increased (perhaps doubled) from the values cited in Table V, while the predicted current would have gone to zero. In the case of a short circuit load, the effect on predicted currents and voltages would be reversed.

ESTIMATED MAXIMUM SUBSTATION EM1 LEVELS

Using the validated models, estimates of the maximum (zero-to-peak) EM1 levels expected in substations from switching operations, faults, and lightning are summarized in Table VI [lo]. Maximum levels are given for EM1 types including: bus current transients, switchyard electric and magnetic field transients, and current and voltage transients on unshielded control wiring as seen from inside the control house near the cable loads. Electric and magnetic fields are representative of those found on the ground directly beneath a section of excited bus. Both vertical and horizontal electric field components are reported; the vertical component gives the highest overall peak amp1 itude, but the horizontal component is typically important for electric field coupling to cables. Ground currents (CT and CCVT groundstrap and pedestal currents) are very strong sources of magnetic field coup1 ing to nearby control wiring. Voltage and current EM1 estimates are given for both open circuit and short circuit loads on unshielded control wiring and include conducted and radiated EM1 coupling modes.

Table VI

COMPARISON OF SWITCHING, FAULT, AND LIGHTNING TRANSIENTS EXPECTED IN SWITCHYARDS

AND ON CONTROL WIRING INSIDE CONTROL HOUSE (zero-peak)

Switching transient predictions are given for disconnect operations under 2PU initial conditions across the switch gaps, nominally the worst case. Switching transient EM1 is reported for 115 kV and 500 kV AIS and for 500 kV GIs. A phase-to-ground fault occurring at a closed 500 kV circuit breaker, where it and the local bus both happen to be charged to, say -1PU, at the instant an incoming traveling wave arrives from having closed a remote switch charged to, say tlPU, was simulated as a worst case, 2PU initial fault condition. The fault was simulated by inserting a short, low im-

1877

a r e based s o l e l y on t h e s h i e l d i n g e f f e c t i v e n e s s measurements obta ined f o r a metal c o n t r o l house w i t h windows i n a 115 kV substat ion [13], and do no t inc lude any e x p l i c i t con t r ibu t ions from f i e l d s generated i n s i d e the cont ro l house by inadver tent (ungrounded) conductor penetrat ions. For example, penetrat ions by bundles o f telephone cables, etc., have been observed t o cause l o - c a l l y h i g h f i e l d l e v e l s i n s i d e of o therwise w e l l - sh ie lded communications rooms [lo] which, i n some cases, could be h igher than those given i n Table V I I .

pedance transmission l i n e t o ground a t the c i r c u i t breaker. Only the h igh frequency EM1 occur r ing dur ing t h e f i r s t 20 microseconds o f the f a u l t was predicted, r e s u l t i n g i n a f a u l t cur ren t o f 15.9 kA; no attempt was made t o inc lude normal load cur ren ts o r t o p r e d i c t the la te - t ime, low frequency (60 Hz) f a u l t EM1 which may a lso have very l a r g e amplitudes. EM1 produced by a l i g h t n i n g s t r i k e t o a 500 kV A I S was a lso estimated. I n t h i s case, a 10 kA st roke was d i r e c t l y attached t o a transmission l i n e j u s t outs ide the substat ion t o represent t h e maximum t h a t might be expected from a sh ie ld - i n g f a i l u r e . This l i g h t n i n g cur ren t waveform r i s e s t o a peak o f 10 kA i n 200 ns, f a l l s t o 525 A a t 100 microseconds, f a l l s t o 75 A a t 5 mil l iseconds, and i s f i n a l l y zero a t 10 mil l iseconds. A s t r i k e on a 500 kV substat ion was considered t o be worse than one on a 115 kV substat ion. It could be argued t h a t the same 10 kA on a 115 kV s t a t i o n would produce much h igher f i e l d s on the ground due t o the shor te r bus height, bu t the f lashovers would occur a t lower voltages, thus lower ing the surge cur ren t on the bus which dr ives the f i e l d s .

The da ta i n Table V I show t h a t t h e 2PU c i r c u i t breaker f a u l t produces t h e h ighes t l e v e l s f o r a l l EM1 types. The l i g h t n i n g s t r i k e produces l e v e l s comparable t o those o f 500 kV A I S d isconnect swi tch ing. Overa l l EM1 i s worse a t 500 kV than a t 115 kV, as expected. Levels i n 500 kV A I S are genera l l y h igher than those i n 500 kV G I s , t h e l a t t e r be ing lower presumably due t o sh ie ld inq provided by the gas enclosure. (However, the GIs models a re a l s o more complex and n o t as w e l l va l ida ted as those f o r AIS.) The p r o b a b i l i t y o f occurrence o f EM1 f rom t h e f a u l t and l i g h t n i n g s t r i k e scenarios should be low, whereas from swi tch ing opera- t i o n s i t i s r e l a t i v e l y h igh. These f a c t o r s should be considered i n choosing whether t o use t h e f a u l t o r the d isconnect sw i tch ing t r a n s i e n t as t h e upper bound on substat ion EMI. I n a l l cases, the EM1 l e v e l s i n a 115 kV substat ion are considerably lower than those o f 500 kV .

Table V I g ives t h e maximum est imated EM1 l e v e l s expected i n t h e swi tchyard and on unshie lded CT w i r e s i n s i d e the cont ro l house near equipment loads.

Peak e l e c t r i c and magnetic f i e l d EM1 l e v e l s i n - s ide o f c o n t r o l houses r e s u l t i n g from t h e same EM1 sources and scenar ios j u s t discussed were a l s o es- t imated and are repor ted i n Table V I I . These estimates

Table V I 1

PEAK TRANSIENT FIELD LEVELS INSIDE CONTROL HOUSES

(zero-peak) (Based on cont ro l house sh ie ld ing ef fect iveness only)

COMPARING MAXIMUM EM1 LEVELS TO STANDARDS

Estimates o f t h e maximum EM1 l e v e l s expected i n subs ta t ions o f d i f f e r e n t types and vo l tages are summarized i n Table V I i n terms o f peak ampli tudes. For comparing w i t h standard t e s t waveforms, i t i s more i n - format ive t o compare waveform ampl i tudes versus t ime o r spect ra l ampl i tudes versus frequency.

F igure 6 i s an over lay o f three predic ted 500 kV A I S v e r t i c a l e l e c t r i c f i e l d spect ra: 1) 2PU phase- to- ground c i r c u i t breaker f a u l t (bo ld curve) , 2) 10 kA l i g h t n i n g s t r i k e t o bus due t o s h i e l d i n g f a i l u r e ( d o t t e d curve) , and 3) 2PU disconnect sw i tch ing t r a n - s i e n t ( t h i n curve) . S ince t h e r e i s c u r r e n t l y no swi tchyard f i e l d t e s t standard, t h e envelope o f these spec t ra has been approx imate ly f i t t e d w i t h t h e smooth dashed curve as a suggested upper bound. The waveform o f t h e dashed curve i s g iven by a simple double ex- ponent ia l o f the form:

E ( t ) = A[exp(-at) - exp(-bt ) ] where

E = e l e c t r i c f i e l d ampli tude a t t ime t A = 100 kV/m, the peak e l e c t r i c f i e l d ampli tude a = 1.OE5 Hz, descr ib ing the l a t e t ime decay r a t e b = 3.5 E7 Hz, descr ib ing the r i s e t ime

The (10% - 90%) r i s e t ime o f t h i s e l e c t r i c f i e l d i s 58.6 ns and i t s ac tua l peak ampl i tude i s 99.6 kV/m. Decay t o 50% of peak ampl i tude occurs i n about 10 microseconds as seen i n F igure 7. A s i m i l a r waveform can a lso be generated f o r the magnetic f i e l d simply by sca l ing the e l e c t r i c f i e l d by an impedance o f about 300 ohms. This w i l l generate a peak magnetic f i e l d of about 332 A/m. A method f o r genera t ing these waveforms f o r t e s t purposes i s described i n [lo].

1 U

8 U

N -1

\ E -'

18

\ U v) -a z l 8 = - 4

U 3

l8

-6

l8' l8. US 18' U' ;me i m

FREQUENCY (Hertz)

Figure 6. Comparing Highest Expected Levels o f E l e c t r i c F i e l d Spectra i n 500 kV Switchyard w i t h a Suggested Test Envelope

1) 2PU c i r c u i t breaker f a u l t 2) 10 kA l i g h t n i n g s t r i k e (sh ie ld ing f a i l u r e ) 3) 2PU disconnect swi tch ing t r a n s i e n t 4) Envelope

1878

TOTAL VERTICAL ELECTRIC FIELD by the fast transient (curve 4) at all frequencies. The fast transient spectrum was generated using the

V(t) = A[exp(-at) - exp(-bt)l ~ double exponential waveform:

2PU CIRCUIT BREAKER FAULT

where

V = Voltage amplitude at time t A = 4.57 kV, the peak fast transient

a = 5.0 E6 Hz, describing the late

b = 1.75 E8 Hz, describing the rise time

SWC test wave open circuit voltage

time decay rate

With these choices of coefficients, the peak Voltage is .0 1.0 8.0 a.. 4.0 1.0 a.. 7.0 0.0 ,.. 4 kV. The rise time is about 10 ns and the fall time

(to one-half maximum) is about 150 ns. TIME (microseconds)

Figure 7. Suggested Electric Field Test Waveform Compared wi th Faul t , Lightning. and Disconnect Switching Transient Waveforms

DISCUSSION AND CONCLUSIONS

The highest expected levels of several types of Similarly, 500 kV control wire Spectra Were EM1 inside substations of up to 500 kV have been es-

compared with the fast transient and OSCillatOry swc timated using measurements and models based on measured test waves of IEEE/ANSI C37.90.1-1989 1151. TO make transient interference data. In most cases, the 2PU this conuparlson, open-circuited and unshielded cables switching transient models only slightly overpredict were assumed. Figure 8 shows an overlay of 5 curves Of the highest levels of EM1 actually measured. EM1 from voltage spectral anplitudes for the fault, lightning. switching operations, faults, and 1 ightning have been and disconnect switching transient data summarized compared. The 2PU phase-to-ground circuit breaker fault earlier in Table VI. The other two (solid bold) curves occurring during remote switching produced the highest are the fast transient and the oscillatory SWC Voltage EM1 levels overall. The probabilities of occurrence of spectra. All five curves are labeled and identified in EM1 from the fault and lightning scenarios considered the figure. The model predictions do not include loss. here are believed to be very low compared to that o f Based on measurements in a 500 kV GIs. losses Will 2PU disconnect switching. However, protection against would be expected to decrease predicted Spectral EM1 effects from a 2PU circuit breaker fault may be re- amplitudes above 7 MHz by a factor of 2 or 3. For quired. Lightning arrestors, which may reduce EM1 control cables that are shielded, analysis shows that levels in the bus, were not included in the model. spectral amplitudes will also be lower at all frequencies shown for the environments in Figure 8. Generally, EM1 levels from routine disconnect The alnount of decrease is greatest at low frequencies switching are shown to increase linearly with substa- and least (factor of 2) at high frequencies, i.e.. tion voltage. Measured and predicted bus current tran- above 1 MHz. However, even for shielded cables, the sients from disconnect switching are presented and fault, lightning, and disconnect switching transient agree well at all substation voltages. Bus current spectral voltages still exceed the envelope presented transients are also shown to be the origin of observed

transient electric and magnetic fields; simple peak a OIPCCIIIWIT I-- ALPONCL. mn UIWILWI, Q- amp1 itude field scaling formulas have been provided.

cI'xalIu* u-xWGc.OI1)* lyLT(- -3. - W. SNWOLD) Maximum predicted transient electric and magnetic field levels (from disconnect switching) on the ground under the bus in 500 kV substations are predicted to be 30 kV/m and 179 A/m, (zero-to-peak). (Faults produce peak field levels about three times higher.) The ground- plane electric field level estimate was found to be about a factor of 2 higher than the measurements; agreement improves at above-ground locations. Several possible explanations for the discrepancy were discussed; it was also indicated that electric fields near grounded structures can even be higher. The 2PU predicted electric field from disconnect switching is close to the value reported in [l], but both the measured and 2PU predicted magnetic field levels found in this study were 80 times higher than those reported in [I].

FREQUENCY (Hertz) Maximum control wire EM1 inside the control house has been described as the sum of contributions from

Amplitudes for Predicted CT Cable EM1 principal coupl ing pathways in unshielded cables: con- from a 2PU Fault, 10 duction Of bus current transients parasitically through 2PU Disconnect Switching with the SWC instrument transformers, and coupl ing of radiated Fast Transient and Oscillatory Test electric and maqnetic fields.

Fault, lightning, and disconnect switching tran- Voltage Spectra in 500 kV Substation. sient EM1 coupling to shielded and filtered control cable loads was also investigated using the models described earlier [lo]. In shielded cables, fields couple first to control cable and then to the inner conductors via piqtails and transfer impedance.

Figure 8 . Comparing Open Circuit Voltage spectral

lightning, and

1) Predicted fault EM1 2) Predicted lightning strike EM1 3) Predicted di"nect switching transient EM1 4) Fast transient SWC test wave 5) Oscillatory SWC test wave

1879

Comparison of the data in Table VI11 with that of Table IX illustrates the effectiveness of using shielded control cables.

Table VI11

PEAK EM1 COUPLING BY MODE IN AN UNSHIELOED CT CABLE (2PU disconnect switching transient; 150 ohm load)

I COUPLING MODE. 0-PEAK MPLITUDE I CURRENT I VOLTAGE I

1.bb I 0 . Z W 4 I). c ! .)n I. , c ee

Table IX

(2PU disconnect switching transient; 150 ohm load)

I COUPLING MODE. 0-PEAK AnPLITUDE I CURRENT I VOLTAGE 1

PEAK EM1 COUPLING BY MODE IN A SHIELDED CT CABLE

Table X

EFFECT OF SHIELDING AND FILTER CAPACITANCE ON LOAD AND FILTER EM1 STRESS LEVELS IN CT CONTROL CABLES

(EM1 source: 2PU circuit breaker fault; load: 150 ohms) (500 kV substation)

I load remonse I I I

Table V I 1 1 shows that magnetic field coupling to the CT ground strap and pedestal is the most important coupling mode, nearly twice that of conducted coupling, when the CT cable is unshielded. However, Table IX shows that once the cable is shielded then conducted coupling dominates all other modes.

High frequency conducted coupl ing through bushing and column type CTs and through CCVTs was found to be a significant contributor to control wire EMI. High frequency transfer function measurements of instrument tranformers are recommended to fully understand this mechanism because coupling is directly from the bus to the control wire and is not reduced by cable shields. In CTs and CCVTs employing shields, their effectiveness should be evaluated between 100 kHz and 150 MHz, unless known. From the CT transfer function data measured in project RP2674-1 [lo], coupling efficiency was found to increase at higher frequencies. Furthermore, because of the short time constants associated with arcing in the interrupter gaps, switching transients from circuit breakers, while of much lower overall amplitude than those of disconnects, do tend to produce EM1 having much higher frequency components; these frequencies may couple more efficiently through the CT to control wiring.

Surge suppressors such as filters and MOVs are also very effective in reducing EM1 levels, especially the conducted EM1 mode. Simple capacitors of 0.5 uF, 0.05 uF, and 0.01 UF are typically used to protect relay equipment, with the latter two values more typical for protecting solid state and digital relays, respectively. When appropriate cable shields and filters are used, even EM1 levels from the severe 2PU circuit breaker fault can be reduced significantly at relay equipment inputs. This is illustrated in Table X for various shield and filter combinations. Proper sizing of EM1 protection devices such as filter working voltage is also important to prevent damage from the EMI. Information concerning single transient peak voltage, peak current, peak power, average power, and peak energy delivered to cable shields, filters, and equipment loads can be easily calculated using the models; many of these quantities are reported in Table X.

Disconnect switching transient macrobursts imply a quasi-random repetitive EM1 stress of up to 5,000 varying amplitude micropulses on cable loads per switch operation. Up'to 20 - 50 of these transients occur at maximum amplitude (2PU) at a rate of 120 Hz per operation. At a maximum control wire EM1 level of 10 A and 3 kV per pulse, this amounts to 30 kW peak power per pulse. At 120 Hz this amounts to about 90 watts average power. For 50 pulses of 25 microsecond duration, this represents an energy of about 38 joules delivered to the cable load (or to a surge protection device) per disconnect switching operation. Cable loads and protection should be designed to withstand these EM1 levels.

The shielding and protection recommendations from [l] are reconfirmed here. Because of the possibility of conducted coupl ing, surge protectors are recommended in addition to cable shielding. Surge protectors may be required anyway to limit pigtail coupling to compatible levels.

Analytic descriptions of possible test waves characterizing bounding levels of electric and magnetic field EM1 in switchyards up to and including 500 kV have been presented. Comparisons of expected control wire EM1 levels with standards suggest possible increases in the fast transient test wave at certain frequencies below 7 MHz be considered. The oscillatory test wave was not found to provide an effective bound on the expected maximum EM1 levels for substations greater than 115 kV.

Susceptibility levels for control-critical substation equipment, not just relay equipment, should be determined, if not known, and compared with expected maximum EM1 stress levels and SWC test criteria to ensure that adequate margins exist. During substation switching transient tests, many upsets and damage o f non-relay equipment occurred [lo].

ACKNOWLEDGMENTS

The authors are grateful for the guidance provided by F . M. Phillips, S . L. Nilsson and L . L. Mankoff of EPRI. Public Service Company of New Mexico and Virginia Power have been strong supporters throughout this project.

1880

REFERENCES

B.O. Russell, S.M. Harvey, and S.L. Nilsson, "Sub- station Electromagnetic Interference, Part 1: Characterization and Description o f the Transient EM1 Problem", IEEE Trans. on Power Apparatus and Systems, vol.

S. A. Boggs, F.Y. Chu, N. Fujimoto, A. Krenicky, A. Plessl, and D. Schlicht, Disconnect Switch In- duced Transients andII Trapped Charge in Gas- Insulated Substations , IEEE Trans. PA&S, vol. PAS-101, pp. 3593- 3602, October 1982.

PAS-103, no. 7, July 1984.

D.E. Thomas, C.M. Wiigins, F. S. Nickel, C. D. K O , and S. E. Wright, Prediction of Electromagnetic Field and Current Transipts in Power Transmission and Distribution Systems , IEEE Trans. on Power pelivery, vol. 4, no. 1, pp. 744-755, January 1989.

E.F. Vance, Coupling to Shielded Cables, New York: Wi 1 ey- Interscience, 1978, pp ., 108-176.

HEMP-INDUCED TRANSIENTS IN TRANSMISSION SUBSTA-. m. Oak Ridge, TN: Oak Ridge National Laboratory, June 1990, ORNL/Sub-88-SC863.

[21] D. E. Thomas, C. M. k/iggins, T. M. Salas, and P. R. Barnes, On the HEMP Environment for Protective Relays", presented at the IEEE/PES 1993 Winter Meeting in Columbus, OH.

J. Meppelink, K.Diederich, K. Feser, and P. Pfaff, "Very Fast Transients in GIs", IEEE Trans. on Power Delivery, vol. 4, pp. 222-233, January 1989.

S.Ogawa, E.Haginomori, S.Nisiwaki, T. Yoshida, K. Terasaka, "Estimation of Restriking Transient Overvoltage on Disconnecting Switch for GIs", IEEE Trans. on Power Delivery, vol.1, p. 95 (1986).

C. M. Wiggins, F. S. Nickel, A.J. Haney, "Measure- ment of Switching Transients in a 115 kV Substation", IEEE Trans. on Power Delivery, vol. 4, pp. 756-759, January 1989.

C.M. Wiggins and S.E. W;ight,"Switching Transient Fields in Substations , IEEE Trans. on Power Delivery, vol. 6, pp. 591-600, April, 1991.

W. C. Kotheimer and L.L. Mankoff, "Electromagnetic Interference and Solid State Relays", IEEE Trans, on PAE, vol. PAS-96, no. 4, July/August 1977.

ANSI/IEEE C37.90.1-1989 "IEEE Standard Surge With- stand Capability ISWC) Tests for Protective Relays and Relay Systems , (P472/D9, January 6, 1987). ANSI/IEEE C37.90.2-1987 "Withstand Capabil i ty of Relay Systems to Radiated Electromagnetic Inter- ference from Transceivers".

[ 101 ELECTROMAGNETIC TRANSIENTS IN SUBSTATIONS, VOLUME I : PROJECT SUMMARY AND RECOMMENDATIONS, Pal o A1 to, CA: Electric Power Research Institute, April 1993, EPRI TR-102006.

[ll] C.M. Wiggins, F. S. Nickel, and A.J. Haney, "Mob- ile Transient Measurement System", 1987 IEEE Inter- national Symposium on EMC, vol 87CH2487-7, pp. 42- 54, August 1987.

[12] Proceedings: Telephone Lines Enterinq Power Sub- stations. Palo Alto, CA:Electric Power Research In- stitute, August 1988, EL-5990-SR, Section 5.

[131 See [lo], section 4. [ 141 ELECTROMAGNETIC TRANSIENTS IN SUBSTATIONS, VOLUME

UI: TEST REPORT. Palo Alto, CA: Electric Power Research Institute, EPRI TR-102006, April 1993.

[15] See [lo], pp 5-13 - 5-18.

[16] See [lo], section 6.

[17] D.E. Thomas, C.M. W!ggins, T.M.Salas, F.S. Nickel, and S. E. Wright, EMI-Induced Control Wire Tra!- sients in Substations: Measurements and Models , presented for the IEEE/PES 1994 Winter Meeting in New York.

Carl M. Wiqqins (M'74, SM'89) was born in Jackson, MS on August 5, 1941. He received the B.S. degree from Lamar State College of Technology, Beaumont, TX and the M.S. degree from Sam H o u s t o n S t a t e C o l l e g e , Huntsville, TX in 1964 and 1966, respectively, both in Dhvsics. From 1966 to 1973 he studied postgraduate physics at New Mexico State University in Las Cruces, NM. In 1973 he joined BDM

International, Albuquerque, NM. His work has been in the areas of transient electrodynamics, lasers, and op- tics. Currently, he is a senior principal scientist investigating electromagnetic interference phenomena. Mr. Wiggins is a senior member of the IEEE EMC and Power Engineering Societies and has authored over 38 pub1 ications.

I

7 curved mirrors was published in the December 1980 issue of Scientific American.

Currently, Mr. Thomas is a Principal Staff Member in

the Advanced Electromagnetics Group of BDM Interna- tional, Inc. His research areas include assessment of electromagnetic effects on aircraft, ships, and on electric power transmission and distribution systems.

1881

Frank S. Nickel was born in Salina, Kansas in 1962. He received the B . S . degree (1984) in physics engineering and a minor in mathe- matics from Southwestern Ok- lahoma State Universitv (SWOSU) .

Mr. Nickel joined BDM In- ternational, Inc., Albu- querque, NM in 1984.. He has concentrated his work in the areas of transient signal anal ysi s , data acqui sit i on and processing, electromagnetic and electrical network

model development, systems simulation, and hardware and software engineering. Currently, he is an engineer for an electromagnetics test and analysis group, and is a project manager responsible for the operation, main- tenance, enhancement, and application of a high bandwidth, versatile data acquisition and processing system.

and has worked in the areas of electromagnetic compatibility and transient analysis. He is a student

member o f the IEEE Electromagnetics Compatibility Soci etv.

Selwvn E. Wrisht (M86, SM88) received the B.S. degree in physics from North S t a f f s P o l y t e c h n i c in England, the M . S . degree in electronics in 1964, and the Ph.D. degree in acoustics in 1 9 6 9 from S o u t h h a m p t o n University. He became a chartered engineer in the United Kingdom in 1965.

Dr. Wright was Scientific Advisor with the French Government (ONERA) from 1976-1978. In 1978 he joined

Stanford University to start a laboratory in acoustics. Dr. Wright joined the Electric Power Research Institute in 1984, as a project manager in the Electrical Systems Division. His specialties include acoustic and electromagnetic fields, control s , and instrumentat ion. Dr. Wright has authored over 50 principal publications.

1882 Discussion

Steven A. Boggs (Electrical Insulation Research Cen- ter, University of Connecticut, Storrs, CT 06269-3136 and Department of Electrical Engineering, University of Toronto). The authors are congratulated on an interesting and timely study of transient electromagnetic interference in substations. My comments will be restricted to the subject as related to GIS (SF6 Gas-Insulated Sub- stations). The authors might have done well to separate their results into two papers, as phenomena in GIs differ SufEciently from those in AIS that treating both in a single paper is almost certain to result in confusion. For example, the authors state that Voltages in transmission substations generally range between 115 kV and 500 kV. Surge impedances over this voltage range also tend to be relatively constant at approximately 350 ohms. This statement is correct of AIS but incorrect for GIS, where the impedance ranges from about 45 to 65 ohms. In the next paragraph, the authors relate the risetimes of transient EMI phenomena caused by arcing discharges on the high voltage system to the effective charging time constants of the circuit driven by the arc, typically the entire substation bus structure. Again, this may be an appropriate description for AIS but is inappropriate for GIS, where the coaxial structure forms a system of relatively clean high frequency transmission lines capable of supporting ns risetime travelling waves and reflections thereof. Given that the typical time for collapse of the voltage across a disconnect switch is in the range of 3 to 5 ns, the frequency spectrum resulting from switching of GIS is related to details of station structure more complex than simply overall bus capacitance.

The authors state that the peak field amplitudes in GIS are somewhat lower (E-field a factor of nine, H-field a factor of two) than those of AIS; the gas enclosure probably acts as a shield. Where did the authors measure the fields in GIS? One might expect the maximum field to occur immediately under the gas-to-air termination. The intercontact breakdown of a disconnector in GIS creates a travelling electromagnetic wave within the GIS which reflects and refracts within the GIS until it reaches a gas- to-& termination, at which point part of the wave is re- flected back into the GIS, while part is refracted out the overhead line. The gas-to-air termination represents the junction of three transmission lines, viz., (1) the overhead line-to-earth transmission line, (2) the GIs conductor-to- enclosure transmission line, and (3) the GIS enclosure- to-earth transmission line. Part of the refracted wave is coupled into the GIS enclosure-to-earth transmission line. The duration and waveform of this wave depends strongly on the length and proximity of any ground straps, as the base of the bushing is often grounded to the station ground mat [l-31. The field below the bushing will normally be greatest when a line disconnector is op-

erated, when the enclosure at the base of the bushing is high above the earth, and when a ground strap is not present near the base of the bushing. Did the authors measure the field under bushings, and, in particular, did they measure the field under a bushing which represented such a worst case condition?

The authors relate many of their measured data to a 2 pu breakdown across a disconnect switch. he operation of GIs disconnectors has been understood for over a dec- ade, and manufacturers take great pains to minimize the likelihood of anything approaching a 2 pu intercontact breakdown, as such a breakdown implies both large transients and a very long intercontact arc, both of which reduce the reliability of the GIs. The intercontact breakdown voltage is basically controlled through careful at- tention to both the rate of contact separation and the asymmetry between breakdown voltage in the two direc- tions across the disconnector [4,5]. Thus while relating data to a 2 pu intercontact breakdown provides a reasonable basis for normalizing reported data, the resulting fields are substantially greater than can reasonably be expected in a GIs. The authors bottom line is that the measured fields are sufficiently great that the IEEE/ANSI standard C37.90.1- 1989 does not assure reliable operation. However, the authors measured transients in substations based on the standard practice which has grown up around the use electromechanical relays. In many early GIS, control wiring practice was so poor that even electromechanical relays could be damaged, especially by breakdowns during commissioning tests, which represent the highest normal exposure to EMI and control wiring transients. The usual reason for such damage was that control wiring shields were grounded at only one end in order to avoid circulating currents in the shields. When the shield was grounded at both ends, damage to electromechanical relays was eliminated. However, the resulting wiring practice was still poor. Control cable shields were grounded through long pigtails, including such pigtails within local cabinets. Such practice, however imperfect, has been adequate for substations based on electromechanical relay technology. More recently, a number of utilities have started to use microprocessor-based relaying in GIS, and at least three manufacturers have provided GIS which have incorpo- rated or been used with such technology. To implement such systems, control wiring had to beimproved substantially, with coaxial termination of control cable shields on the outer surface of wiring cabinets, elimination of pigtails, etc. The net result is a control wiring system in which the control wires travel through a continuous faraday cage from their point of origin to their point of termination. As noted above, such technology is now well-established with at least three manufacturers having imple-

mented GIs incorporating such technology on a commer- cial basis. Thus the authors' measurements do not relate to the easily achieved and established state-of-the-art in GIS control wiring but rather to earlier generations of technology which was, and is, adequate when electromechanical relays are employed but which is clearly an inadequate basis for implementation of microprocessor-based relaying. The soon-to-be-published revision of IEEWANSI docu- ment on GIS includes a revised Specification and numer- ous Guides, some of which cover Fast Transients in GIs, Transient Groundrise in GIs, and Control Wiring Prac- tice for GIs.

None of the above comment detracts from the very substantial contribution of the present authors. Hopefully, my comments serve to place these contributions in the context of the present state-of-the-art as it relates to GIs. 1. Ford, G.L. and S.A. Boggs. Transient Groundrise in SF6

Substations Investigated. Transmission & Distribution Magazine, Vol. 31, No. 8, Aug. 1979.

2. Ford, G.L., S.A. Boggs, and N. Fujimoto. Transient Groun- drise in GIs. Transmission & Distribution Magazine Vol. 34, No. 4, April, 1982, p. 42.

3. Fujimoto, N., E.P. Dick, S.A. Boggs, and G.L. Ford. Tran- sient Ground Potential Rise in Gas Insulated Substations- Experimental Studies. IEEE Trans. PAS-101, October, 1982.

4. Boggs, S.A., F.Y. Chu, N. Fujimoto, A. Krenicky, A. Plessl, and D. Schlicht. Disconnect Switch Induced Tran- sients and Trapped Charge in Gas-Insulated Substations. IEEE Trans. PAS-101, October, 1982.

5. Boggs, S.A., N. Fujimoto, M. Collod, andE. Thuries. The Modeling of Statistical Operating Parameters and The Computation of Operation-Induced Surge Waveforms for GIS Disconnectors. 1984 CIGRE, paper 13-15.

Manuscript received February 15, 1994.

Carl M. Wiggins, David E. Thomas, Frank S. Nickel (BDM Federal, Inc., Albuquerque, NM 87106) and Selwyn E. Wright (Electric Power Research Institute, Palo Alto, CA 94303). The authors would like to thank Mr. Boggs for his thoughtfbl comments. There are indeed significant differences between transient electromagnetic interference in air-insulated substations (AIS) and gas-insulated substations (GIs). He is correct in pointing out the differences in surge impedances in AIS and GIs. While the high fiequency traveling waves from an arcing switch do excite the entire bus structure,

1883 we have also shown that it is the large impedance discontinuities on either side nearest the arcing switch, and to some extent the details of the switch itself, that are predominately responsible for the observed transient waveshape [ 101 (references in paper). Typically such discontinuities are provided by the nearest instrument transformers and (open) circuit breaker bushings in the case of disconnect switching in AIS which offer low impedance paths to ground at frequencies in the MHz range. In GIs, the bus structure is more complicated and often may be viewed as three different intersecting transmission lines as Mr. Boggs points out. The surge impedances for the three lines are all different. High frequency electromagnetic interference produced by arcing from a switch operating inside the coaxial line couples onto all three lines. Circuit breakers and disconnects (with its associated breaker open) were operated to produce the measured transients; no external line switch operations were allowed. Due to the compactness of the coaxial structure, switching in GIs produces transients with much higher frequency components than in AIS, as we point out in the paper (e.g., Fig. 4) and in [6,10, 141. We measured transient electric and magnetic fields at 14 different locations between the gas enclosure and ground and under the air- insulated line on both the line and transformer sides [lo], but not between the bus and gas enclosure (i.e., not inside the gas enclosure), as mentioned in the paper. The peak fields may be much lower at these locations than would be found inside gas enclosure. In some cases fields were measured close to aidgas bushings [ 141 and were found to be somewhat higher at these locations, particularly the magnetic field. It should also be mentioned that the field measurement geometry between the enclosure and earth ground, which is often a three- dimensional structure of steel wakays , is not "clean". Thus, field sensors will tend to measure the net field scattered from these structures which may cause local peak amplitude enhancements or cancellations. Of course, the bottom line is: how do transients arriving at equipment in GIS compare with those found in AIS? Many results in [lo, 141 show that, generally, voltage and current transients have about the same peak amplitudes in GIS and AIS where the same type of switching is performed at the same voltage levels, and similar interference coupling to equipment (cable practices) are employed. The major differences were that GIS control cable transients measured near equipment loads exhibited major fiequency components up to 20 MHz (vs. up to - 3 h4Hz in AIS), and they damped out faster ( 4 p vs. 20 p).

The 2 pu initial condition across the disconnect switch arises when its the line-side has been left charged to +1 pu on opening and then it is closed where the initial arc

1884

occurs at the instant the source side is at amhimum, -1 pu. The correlation between peak MeasuTemenfs and travehg wave simulation model predictions set for 1 pu and 2 pu initial conditions c~njirm the possibility of occurrence of 2 pu transients (see for example, Chapter 5 of [ lo]).

Our observations of cable shielding practices found in the 10 AIS and GIs visited during project RP2674-01 show: only some use of shielded cables; sometimes cable shields were grounded; and when grounded, pigtails were used. Rarely did we encounter a substation where all cables were .&ielded. We have not seen a substation using high frequency coaxial shield terminations, but we are pleased to learn that apparently some now do. We would also like to point out that a key conclusion in the paper is that cable shielding alone, without proper surge

protection, will not necessarily safely limit the peak interference levels from switching, hults, and lightning transients that can occur at inputs to protection equipment. The paper characterizes quantitatively some of the peak EMI levels for normal and abnormal high frequency sources in AIS and GIs, their effects, and the effectiveness of various mitigation procedures such as cable shieldmg and surge suppression. The purpose of this research is to provide high fiequency EMI data that could be used to better understand and improve protection of modem (digital and microprocessor) substation electronic equipment placed either inside the mtro l house ot in the high yard (AIS and GIs).

Manuscript received April 19, 1994.

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