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December 2002 

Surge Protection Anthology 

All Papers – Part 4 

Propagation and coupling of surges 

(Experiments and numerical simulations) 

Surges Happe

FOREWORD 

The papers included in this part of the Anthology provide basic and tutorial information on the propaga

of surges in low-voltage AC power circuits. The subject was approached by a combination of experim

and theoretical considerations, including numerical simulations. One important distinction is m

between voltage surges and current surges. Historically, voltage surges were the initial conc

because the objectionable consequence of surges was a breakdown of insulation – a voltage-rela

phenomenon. In particular, the breakdown phenomena were significant in low-voltage circuits for 

clearances (wiring devices), solid insulation (windings) and surfaces (semiconductor edges). After 

introduction and widespread use of current-diverting surge-protective devices at the point-of-use, 

propagation of current surges became a significant factor. Another significant aspect of the investigat

was to note that many earlier studies on the propagation of surges were conducted from the point of v

of transmission line theory with impulses having durations shorter than the travel time along the relati

long lines. For most of the surge waveforms occurring in AC power circuits and typical installations, 

travel time is shorter than the impulse duration, so that the classical expectation of (characteri

impedance mismatches effect in traveling waves are irrelevant. The papers included in this part re

this dual dichotomy of voltage versus current and impedance mismatch effects versus simple ci

theory. 

Industry interest in the matter also grew, to the point that the IEEE Surge Protective Devices Comm

formed a subcommittee focused on low-voltage concerns, which became an additional forum for m

papers contributed by the researchers cited in Annex A. For obvious copyright limitations, these 

papers from other researchers cannot be reprinted here. The pre-1985 papers in this Part 4 w

copyrighted by the publisher, or were proprietary to General Electric; both graciously gave permission 

reprinting in this anthology. The post-1985 papers, written thanks to the resources of the National Inst

of Standards and Technology, are in the public domain. 

CONTENTS 

Surge voltage suppression in residential power circuits (1976) 

Lightning protection of residential AC wiring (1978) 

Coordination of surge protectors in low-voltage AC power circuits (1979) 

The propagation and attenuation of surge voltages and surge currents in low-voltage AC circuits (1982

Surge and high-frequency propagation in industrial power lines (1985) 

Varistor versus environment: Winning the rematch (1986) 

Fast transient tests – Trivial or terminal pursuit ? (1987) 

Coupling, propagation, and side effects of surges in an industrial building wiring system (1990) 

Electrical fast-transient tests: Applications and limitations (1989) 

Partial discharges in low-voltage cables (1990) 

Diverting surges to ground: Expectations versus reality (1991) 

On the propagation of old and new surges (1991) 

Surging the Upside-Down House: Looking into upsetting references (1994) 

Surging the Upside-Down House: Measurement and modeling techniques [(1995) 

Driving high surge currents into long cables: More begets less (1996) 

The effect of neutral earthing practices on lightning current dispersion in a low-voltage installation (199

On the dispersion of lightning current for a direct flash to a building (2000) 

The role and stress of surge-protective devices in sharing lightning current (2002) 
VOLTAGE

SUPPRESSION

IN

RESIDENTIAL

POWER

CIRCUITS
1

 1967 IEEE Conference papers were available at the meeting for $0.75 and later for $1.00. 

 However, they became no longer available some time thereafter.

Filename: Residential surges

All rights reserved IEEE

Reprinted, with permission, from 

IEEE Conference paper 31 CP 67-4301

Significance:

Part 3 – Recorded occurrences

One of the first papers reporting surge recording in low-voltage ac power circuits.

Both waveform information and frequency of occurrence marked the beginning of the IEEE effort tow

characterizing surge voltages in low-voltage ac power circuits, a departure from the traditional unidirectio

and separate 1.2/50 and 8/20 waveforms in use for high-voltage equipment testing.

In the mid-sixties, when these recordings took place, there were very few, if any, surge-protective dev

(SPDs) installed in residential power circuits. After the emergence of ubiquitous consumer-type SPDs

results of subsequent monitoring campaigns were drastically changed (See “Galore” in Part 2) but 

information is still valid for the frequency of occurrences to the extent that the origins of surges have 

changed, only the observed levels are now considerably reduced by the proliferation of SPDs.
Paper No. 

31 CP 67-430 

Conference Paper 

SWE VOLTAGES IN RESID?~T~KI POWER CIRCUITS 

F. D. Martzloff 

Menber IEEE 

Both of: 

General Electric Cunpany 

Schenectady, N. Y. 

A paper recmended by the IEEE Pmtective Devices Carmittee o

IEEE Power Group for presentation at the IEEE Scnmer Fbwer Me

Fbrtland, Ore. , July 9-14, 1967. Manuscript subnitted April 3

made available for printing May 3, 1967. 

- Price: $1.00 All Rights Resewed by (759 at meeting) The Institute of Electrical and Electronics E

Inc., 

(15C per copy additional if 345 East 47th Street first class mailing desired) New York, N. Y. 10017 

(254 per copy additional if 

air mail is desired) Litho in USA 
SURGE VOLTAGES IN RESIDENTIAL PWER CIRCUITS 

by 

G.J. Hahn and F.D. Martzloff 

Research and Development Center 

General Electric Company 

Schenectady, New York 

Abstract 

Special instrumentation was developed for 

monitoring the magnitude and frequency of occurrence of surge voltages in residential circuits. 

Over a period of two years, more than 400 

locations in 20 cities were surveyed. Monitoring 

was accomplished by automatic recording cathode 

ray oscilloscopes and simple surge counters. 

Two significant sources of surge voltages 

were identified: load switching within the house, 

and external surges, most likely associated with 

lightning, coming through the service drop. 

Internally generated surges as high as 2500 

volts were recorded during this test program, and 

six of the 250 households (i.e., 2.4 percent) in 

the sample experienced surges of this type in excess of 1200 volts once or more per day. A statistical evaluation indicated that a 90 percent 

confidence interval for the true proportion of 

homes with such surge voltages is between 1.0 and 

4.7 percent. 

Surges due to lightning reaching 5600 volts 

have been recorded on a 120-volt overhead distribution line. Independent evidence shows the 

occurrence of a number of surges in excess of 

2000 volts. 

In this study, a total of. 8 surges in excess 

of 1200 volts were caused by lightning and recorded on 39 counters at room outlets in 14 localities over the summer months. 

Introduction 

The successful operation of semiconductors 

and new insulation systems in appliances and consumer electronic devices may be adversely affected 

by transient overvoltages occurring on the 120- 

volt power supply. Detection and measurement of 

these transient surges permits a designer to provide suitable built-in tolerance, or at least to 

recognize that there is a need for appropriate 

suppress ion or protect ion: 

Special instrumentation has been developed 

for detection and measurement of surge voltages. 

In the first phase of testing automatic recording cathode ray oscilloscopes provided complete 

information on wave shape, and on magnitude and 

frequency of occurrence at a small number of 

arbitrarily chosen locations. These data, though 

statistically restricted, were adequate to demonstrate the existence of the problem, define some 

sources, and indicate typical wave shapes. It 

also became clear that there was a need fo

larger sample that would be more statistic

valid. 

In the second phase of testing, one h

surge counters, simple in design and easy 

stall, were developed especially for this 

These counters were installed in several h

homes in various cities located in the Ea

middle West of the United States. 

The purpose of this paper is to discu

cordings obtained during the two phases of 

program and the statistical aspects of the 

It is hoped that publication of this pape

encourage other investigators to publish t

findings. This would provide a broader st

tical base for verification of the frequen

magnitude of surge voltages. 

Instrumentation 

Recording surge voltages in the field 

special problems. For example, the instru

cannot be continually attended and the sig

be recorded are unpredictable in magnitud

ity, and frequency of occurrence. 

In cooperation with Tektronix, Incorp

a type 515 cathode ray oscilloscope was m

to record transients on film, with automa

vance of the film after a recording. 

The modifications involved a change 

trigger and sweep circuits, as well as rem

all knobs (an important detail for an inst

to be left in all types of locations). 

Most oscilloscope trigger circuits a

ized. This often causes a loss of record

the polarity of the initial signal is not 

For surge recording, the trigger circuit w

ified by feeding to the time base trigger 

two OR logic diodes the push-pull signal 

applied to the input of the oscilloscope 

line. With this modification, the sweep 

started for either polarity, while the sig

delayed before reaching the deflection pl

the CRT. In addition, the conventional b

circuit holds the beam in low intensity b

the sweep starts. 

A second modification changed the sw

a constant speed to a logarithmic speed, 

the start and slow toward the end. This 

the recording of an impulse front with goo

lution, while providing sufficient duratio

the sweep to record a long tail. 
Finally, a relay was added with its coil 

energized whenever the beam sweep is triggered. 

Through a set of contacts on this relay, the film 

advance sequence in the camera is activated. 

The camera, attached to the scope bezel, is 

a Beattie-Coleman system, accommodating 100 feet 

of 35 mm film, with no shutter, and recording on 

a 24 x 36 nun frame the trace displayed on the CRT. 

Following the sweep, the contact mentioned above 

closes the motor circuit, advancing the film to 

the next frame in a few seconds. This film transport duration is matched in the oscilloscope trigger circuit by a hold-off so tliat no sweep will 

occur during film advance. 

In order to prevent the film from being 

fogged by extended exposure to the faint glow on 

the phosphor (caused by imperfect blanking of the 

beam, cathode glow or light leaks) a built-in 

timer in the oscilloscope advances the film by one 

frame every hour. This provides an approximate 

method for timing the occurrence of surges. Since 

the surge-voltage tests were conducted, an automatic time recording feature has been added to 

give a more precise recording of time at each 

event. 

Figure 1 is a photograph of this oscilloscope 

camera system. The relatively high cost of this 

system, its conspicuous presence (e.g., size, 

camera noise, etc.) in a home, and the amount of 

film to be scrutinized frame by frame limit the 

number of locations at which recordings can be 

made. However, for conclusions to rest on a sound 

statistical basis, it is necessary to investigate 

a large number of locations. To be acceptable to 

cooperative home owners, a transient detector 

should therefore lend itself to inconspicuous installation and simple data handling. Thus, a device which can merely be plugged into a wall 

receptacle, and that has a digital counter indicating the number of surges occurring above a 

specified threshold is satisfactory for widespread 

recording. 

The circuit shown in Figure 2 was developed 

for the purpose, and packaged as shown in Figure 

3. This device is connected to a wall receptacle 

by a conventional appliance cord, not a special 

probe. This connection is simple, establishes a 

typical final path for a surge impinging upon any 

appliance connected to this particular wall receptacle, and provides power for the high impedance rectifier circuit and a signal input to 

the trigger-sensor circuit. 

A crudely compensated divider (RlCl/R2C2) 

attenuates the incoming surge before application 

to the Silicon Symmetrical Switch (SSS). This device has the characteristic of turning on abruptly 

whenever the voltage across it reaches a threshold, such as 250 volts. This threshold remains 

constant for durations as short as 0.1 s. When 

the SSS turns on, the attenuated surge is applied 

to the gate of the very sensitive, fast switching 

Silicon Controlled Switch (SCS). If the polarity 

of this surge is positive, the SCS turns on, 

allowing the capacitor C to discharge thro

coil of the solenoid-actuated counter S. 

impedance (50,000 ohm) divider and half-wa

tifier provide a trickle charge 15-volt su

the capacitor C. The resistance of this d

represents a negligible load for the house 

system, yet is sufficiently low to allow s

leakage in the electrolytic capacitor C. 

charging time (several seconds) of the cap

is not objectionable, since the transients 

detected do not have a high repetition rat

Calibration of the circuit is obtaine

justing the value of the divider (RlCl/R2C

these tests, the threshold level was set a

volts for one group of counters and at 200

for another group. A bench circuit was se

whereby a surge with adjustable level and 

could be superimposed on a 60 Hz, 120-volt 

supply. Some consideration was given to t

to segregate, or filter out, the surge fro

60 Hz power voltage; in this case, howeve

the minimum level at 1200 volts, the base 

crest represented a small contribution and 

cluded in the signal detected by the circu

Surges with approximately flat top and dur

from several microseconds to 0.2 microseco

applied to determine the variation of the 

old voltage as a function of surge duratio

typical response curve is shown in Figure 

response of this crest-indicating counter 

considered satisfactory, since the oscillo

recordings in homes, as will be discussed 

have indicated that the shortest half-peri

oscillatory surges is in the order of 2 m

seconds. 

Recording Procedure 

In the first phase, oscilloscopes wer

stalled at the basement service entrance o

in the Schenectady, New York area. These 

do not represent a statistician's ideal sa

but, during the initial phase of the progr

nature of transients rather than statistic

racy was the major subject of interest. L

the program, the range of locations was br

to include homes in other urban and rural 

particularly in Florida and South Carolina

In the second phase, surge counters w

stalled at unspecified outlets in the home

engineers in 20 cities in the northeast an

west. Two distinct recording periods were 

uled: winter, because there is usually a 

of lightning activity at that time of year

summer, which ordinarily includes a mziximu

lightning activity. 

Since the oscilloscope input circuit 

at a level of a few volts, it is somewhat 

to direct radiation of electromagnetic noi

the preamplifier, even if there is no sign

the probe. In order to discriminate again

type of spurious indication, the recording 

cedure included a period during which the 

scope probe was shorted with the ground te

remaining connected, so that spurious sign
to ground loops would be detected. In this manner, any noise entering the system could be recorded and later recognized for what it was in 

the analysis of the recordings made with the short 

circuit removed from the probe. 

The surge counters record only one polarity; 

in order to minimize record keeping by the homeowner, the connection of the cord was not polarized. In the event of an oscillatory surge (which 

is always quickly damped in these systems) two 

cases of counting can occur. For instance, the 

high side of the divider can be connected to the 

"hot" terminal of the receptacle (Figure 5). If 

the first half-cycle is positive and exceeds 1200 

volts, a count is registered. If the first halfcycle is negative, and the second, positive half 

is highly damped, no count is registered. If the 

first half-cycle is well above 1200 volts causing 

the second half, in spite of some damping, to exceed 1200 volts, a count is registered. For the 

reverse connection, the same would occur for reverse polarity. Thus, the actual number of surge 

occurrences of both polarities in excess of 1200 

volts is between 1 and 2 times the number indicated by the counter. The objective of the recording in this test series was only to determine 

whether or not surges occurred rather than their 

exact number. Consequently, the value of the 

multiplying factor is not very significant. 

Results Of The Recordings - Oscilloscopes 

The first recording analysis revealed that 

some homes were subject to frequent surge voltages, some experienced only a few isolated surges, 

and others did not experience any surges in excess 

of the trigger level (300 to 400 volts). Furthermore, among those installations where surge voltages frequently occurred, the surges at some 

houses were relatively low (rarely in excess of 

800 volts), while other houses had surges in the 

range of 1200 to 2500 volts. 

Rates of surge occurrence ranged from zero 

(no surge in one to two weeks) to 0.5 per hour, 

with peak values from 400 volts (trigger threshold) to 5600 volts. 

At the conclusion of the recording program, 

a total of 16 homes and two overhead distribution 

poles had been monitored for a total of 7000 

hours. The peak of the surges was distributed as 

shown in Figure 6a; the most frequent and severe 

surges were distributed as shown in Figure 6b. 

Further analysis of some recordings was made 

by deliberate switching of loads in the houses 

where frequent surges had been observed. In some 

cases, the operation of a specific device (e.g., 

oil burner, fluorescent lamp, pump motor, refrigerator, food mixer, etc.) was found to be the 

cause of the surges. In other cases, no amount of 

deliberate load switching could reproduce surges 

such as those recorded during unattended monitoring. The home owner was occasionally able to 

correlate surge recordings with lightning or 

power system disturbances. 

A pattern emerged from all this infor

showing two definite causes of surge volta

the homes: 

1. Load switching in the house occas

produced transient surges; these affe

only that particular house. For any 

ular house, these transient surges ha

shape which was consistently repeated 

variations in amplitude along the ent

recording. This probably resulted fro

bination of switch characteristics an

impedance of the house wiring system; 

variations in amplitude were probably 

by variations of the switching angle 

connected loads. Typical wave shapes 

recorded surges are shown in Figures 

and 7c. 

In one case, it was possible to i

an oscilloscope in a house adjacent t

where repetitive surges up to 1700 vo

being recorded. The service drops we

nected to the same pole, yet no surge

ring simultaneously were recorded in 

second house. 

2. Surge voltages not associated wit

switching within the house were assoc

with lightning storms. In some cases

home owner was able to correlate the 

vance counts with the storm. Althoug

corded during lightning storms at two 

ent locations, the surges shown in Fi

8a and 8b present the interesting cha

istic of being oscillatory at a frequ

the range of 300 kHz. The surge show

Figure 8a was recorded-at the overhea

tribution line (oscilloscope mounted 

pole), while the surge of Figure 8b w

corded at the service entrance in a h

The first exhibits far less damping t

second; this might be explained by th

damping due to lower resistance of th

tem at the pole than at the end of a 

drop. 

Both of these surges, as well as most 

other surges recorded during lightning sto

exhibit this oscillatory characteristic at 

quency which,is nearly constant for a part

locality. This constancy suggests that th

lation of the system followed an excitatio

by the lightning stroke. A number of surg

the range of 800 to 1200 volts were observ

ing several storms. The maximum surge vol

recorded was 5600 volts; several other sur

corded during the same period were in exce

4000 volts. 1 

Results of The Recordings - Counters 

With the two causes of transients ide

by the oscilloscope measurements, the surg

counters wee applied in two separate prog

First, during the winter months, the count

installed in a large number of houses for 

of one or two weeks; the objective was to 
determine how many houses sustained repetitive 

surges. Second, during spring, summer, and fall, 

each counter was left at one or two locations for 

periods of 9 to 48 weeks; it was known from the 

first test that these homes were not subject to 

load switching surges. The objective was to determine the frequency and characteristics of externally generated surges, presumably caused by 

1 ightning . 

The first period produced the data shown in 

Table 1, and the results obtained during the 

second period are shown in Table 2. 

Statistical Analysis of The Recordings 

Internally Generated Surges 

The data in Table 1 show that 6 houses, or 

2.4 percent, were subjected to repetitive surges 

from among a total of 250 homes sampled. The 

true percentage doubtless differs somewhat from 

2.4 percent due to statistical variations. However, one may be 90 percent confident that the 

true percentage is between 1.0 percent and 4.7 

percent." This band could be narrowed by taking 

additional samples. However, it should be noted 

that the precision is proportional to the square 

root of the sample size. Thus, four times the 

number sampled, or 1000 homes, would be necessary 

to cut the size of the statistical error band by 

approximately one-half. Since the above results 

refer to statistical variations only, they do not 

take into account possible biases due to such 

factors as restrictions in selecting members of 

the sample (principally engineers in a number of 

designated locations) or the time of year (winter 

months). 

The probability of internally generated surges 

undoubtedly varies among e.conomic groupings (i.e., 

the devices in use in homes probably vary with the 

economic status of the resident); however, devices 

found to generate surges (i. e. , furnaces, refrigerators, etc.) exist in most homes. 

Externally Generated Surges 

Results from the second testing period that 

was concerned with externally generated surges are 

show in Table 2. From 39 counters installed in 

a total of 91 homes in 14 localities, a total of 

8 occurrences in 6 separate localities were observed during an equivalent exposure time of 846 

weeks. A ninth occurrence was disregarded in this 

analysis because it occurred in the same home during the same storm. Of the two pairs of occurrences which took place in the same location, one 

pair occurred during the same storm and the second 

involved two occurrences at different times. 

s his result is obtained by the well-known method 

of setting confidence intervals for a sample percentage. Further details may be found in statistical texts .293s4 

Analyses could be conducted based on 

following alternative assumptions: 

1. Voltage surges above 1200 volts o

only during the period of the year th

counters were installed in the homes. 

although the counters were in homes f

part of the year, the time involved (

the summer months) was so chosen that 

ther surge8 would have been noted eve

each counter had been run for 52 cons

weeks. The average number of surges 

would then be 8/39, or 0.205. 

2. Voltage surges occur randomly at 

stant rate throughout the year. Thus

counter which was in use 9 weeks, on 

average, would be subject to a third 

surges as a counter in use for a perio

weeks. Under this assumption, a tota

surges observed in a total time equiv

16 years (846 weeks) yields an averag

surges per year. 

Using these two extreme assumptions, 

could be established for the estimated num

surges per year in a larger sample, based 

specific sample. The above calculations r

only to single polarity surges. If all su

being considered, the given values must be 

plied by a value corresponding to the addi

proportion of opposite polarity surges abo

volts which do not also result in positive 

above 1200 volts. This multiplying factor 

probably about 1.6; it accounts for the da

between the first and second 1/2 cycle of 

latory surge. 

Lightning-induced surges are likely t

more than one house when they occur. Loca

graphical and meteorological conditions ar

cal influences on these surges; however, th

factors could not be considered in this pr

investigation. 

To relate the above data to risk of a

failure, the given values must be modified 

probability that a surge above 1200 volts 

cause failure of operating appliances. Fa

effects would vary with different applianc

Possible Further Analysis 

A more refined analysis for a probabi

measurement of voltage surges per lightnin

is possible if the geographical location o

homes, the occurrence rate of lightning sto

during the testing period, and exact dates 

voltage surges occur is considered. The r

values can then be used in conjunction wit

formation given in Reference 5 to calculat

probability of a voltage surge in any spec

geographic area, during a particular part 

year. Such an analysis would remove the n

making one of the two alternative assumpti

stated above and lead to a single set of e

However, this would require more detailed 

could be collected in this program. 
Finally, it is noted that if one has knowledge of the actual voltage surges, rather than 

merely the information concerning whether or not 

a surge above 1200 volts occurred, a more sensitive analysis is possible. In this case, for 

example, probability plotting and other techniques 

based upon the statistical theory of extreme 

values4 3 might well be applicable. 

Effects of Surge-Voltages 

On Connected Electrical Devices 

The surges which have been recorded in this 

program occurred during normal operation of the 

household, with no knowledge of the connected load 

situation at the time of occurrence. 

The question of energy involved in the surge 

is related to the impedance of the system, since 

these recordings provide voltage data only. Surge 

impedance measurements of house wiring cable indicate a value of 100 to 300 ohms for a typical 

branch circuit, so that the surge impedance at the 

service entrance could be in the order of 5 to 10 

ohms. However, this low value exists only for the 

travel time (i.e., a fraction of a microsecond). 

Connected loads will have a lower impedance than 

that of the branch circuit. This value will be 

dependent upon frequency whenever inductive components are present. These loads will absorb part 

of the energy of the surges and thus lower their 

peak. 

Ikvices such as motors and transformers have 

solid insulation and such a long history of successful application that their performance is not 

in question. Perhaps unusual failures can be explained by extreme values of surges as indicated 

by the data. 

Defective wiring practices (e.g., pinched 

insulation, reduced air clearances in wall boxes, ' 

etc.) will cause air flashover with or without 60 

Hz power follow. In fact, one house was brought 

to our attention because of complaints of sparking 

in a light fixture. With the switch in the ground 

wire and the frame attached to a grounded pipe, 

flashover at 1700 volts was observed in correlation with the start of an oil burner in the house. 

This defective light fixture was acting as a voltage limiting gap for the house. 

Appliances containing semiconductors and directly exposed to the line transients may be more 

vulnerable. Actually, the 1200-volt threshold 

level was selected as the result of this consideration. It is interesting to note that, although a 

number of surges above 2000 volts were recorded by 

the oscilloscopes, the few surge counters calibrated for 2000 rather than 1200 volts did not 

produce any recording above 2000 volts. 

An independent study of clock motor failures 

produced information on failure rates versus withstand levels. This study was very pertinent to 

the surge counter program, since thousands of 

clocks are connected a t a l l times to the power 

system. Over a period of three years, failure 

rates were correlated with insulation leve

coils. A very significant 100 to 1 drop in 

ure rate resulted from an increase in withs

voltage from approximately 2000 to approxim

6000 volts. This shows that, even though n

surges over 2000 volts were recorded by the 

volt counters, surges in excess of 2000 vo

indeed occur. 

Conc lus ions 

1. Residential power circuits are sub

to surge voltages due to two distinct 

load switching within the house, and e

nally generated surges that are most l

associated with lightning. 

2. Internally generated surges caused 

load switching are likely to be repet

They can generally be associated with 

specific device, probably operating e

ically or exciting some natural freque

the wiring system. They are not relat

lightning or disturbances from the ut

Peaks as- high as 2500 volts have been 

served. The best single statistical e

is that 2.4 percent of households of t

sampled experience these internally g

repetitive surges in excess of 1200 v

However, because of the statistical v

ity in the sample, this value may be a

as 1 percent or as high as 4.7 percen

Surges may be repeated several times a 

3. The frequency of surges caused by 

ning is not affected by household ele

devices, but rather by local geograph

and meteorological conditions. The li

data in this program reveal several li

ning-caused surge occurrences above 3

volts with one reaching 5600 volts. 

4. Independent evidence shows that a 

icant number of surges above 2000 vol

occur periodically in residential pow

Acknowledgments 

The contribution of the hundreds of in

uals involved in the data collection and th

tribution of J.E. Lenz for the data on ove

distribution lines are gratefully acknowled

References 

Lenz, James E. , "Basic Impulse Insulati

Levels of Mercury Lamp Ballast for Out

Applications,'' (Discussion), Illuminat

gineering, February 1964, pp. 133-140. 

Dixon, W.J. and Massey, F.J., Jr., Intr

tion to Statistical Analysis, McGraw-H

New York, 1957. 

Bowker, A.H. and Liberman, G.J., Engin

Statistics, Psentice-Hall, New Jersey, 
P. Chowdhuri, "Portable Fast Response Transient 

Voltage Counter," Institute of Electrical and 

Electronics Engineers Transactions Power Appliances and Systems, Vol. 84, May 1965, pp. 417-422. 

J. Czech, Oscilloscope Measuring Techniques, New 

York, Philips Technical Library, Springer-Verlag, 

1965. 

S.F. Pearce and J.H. Bull, "High Frequency 

sients in Low Voltage Supply Systems,'' The 

cal Research Association, Leatherhead, Su

1962. 

R.D. Goldblum, "Measuring DC Relay Coil Tr

sients," Electronic Industries, Vol. 24, A

1965, pp. 68-71. 

O.H. Davie and J. Wood, "An Automatic Cathode-ray 

Oscilloscope Beam Brightening Device for Transient 

Recordings ,It Journal of Scientific Instruments, 

Vol. 33, No. 5, May 1956, p. 203. 
Figure 1. Oscilloscope and Camera 

Figure 3a. Surge Counter--Cover Removed 

Figure 3b. Surge Counter Elements--Remove

Shield 

Figure 4. Response of Surge Counter 

(L W 

I2 5 3 *-" 

- P 

'V 

% 1400- 

E 1200-- -- - - -- 

Y 

'L 2 600 

u 

"Y 3 200 

P 

- 

- 

- 

- I 1 1 1 I 1 

0.2 0.4 0.6 0.8 1.0 I .2 

DURATION OF FLAT TOP PULSE . microseconds 
Figure 5. Recording of Surge with Single Polarity 

Counter 

0) MOS

61 MOST 

Figure 6a, Distribu

Surge 

Figure 6b. Distribu

Surge 
kV 

I 

0 

I 

0 5 I0 20 30 40 prec 

Figure 7a. Internal Switching Transient Recorded 

at a Farmhouse (15 such transient surges in 1500 

hours) 

Figure 7b. Composite Record Showing Surges for a 

24-hour Period at a Suburban Home. (Oscilloscope is triggered for each surge, plus once every hour, 

resulting in 2 165-volt band of steady state 60 Hz 

voltage. Recordings above 1800 volts are blanked 

out by the oscilloscope.) 

Figure 7c. Si

corded at the 

5-day period s

corded as foll

1500 volts,21 

volts. Source 

a high-voltage 
0 5 10 20 30 40 75 100 

Figure 8a. Transient Surges Recorded on Overhead 

Distribution Systems in Charleston, South Carolina 

(3800-volt crest, 300 kHz oscillation.) 

Figure 8b. Transient Surge Recorded at 

vice Entrance of a Home in Florida 
OF

RESIDENTIAL

AC WIRING
Keith E. Crouch

Lightning Technologies *

Pittsfield, MA

François D. Martzloff

General Electric Company

Schenectady NY

f.martzloff@ieee.org

Reprint, with permission, of declassified General Electric memo report MOR-78-095

Significance:

Part 4 – Propagation and coupling of surges

Part 8 – Coordination of cascaded SPDs

Laboratory tests on the coupling of lightning current (flowing in the service drop grounded neutral

conductor) onto the phase conductors, inducing overvoltages that were limited by candidate surge

suppressors.

While the injected lightning-simulation current was unidirectional, the induced voltages in the house

wiring circuits had oscillatory components. This observation was used in support of the development 

of the “Ring Wave” concept that was adopted by IEEE 587 (now C62.41).

Three possible types of service entrance SPD of 1960-1970 vintage were investigated

• The then-commercially available silicon carbide/gap arrester

• Metal oxide varistors mounted external to the load center

• Metal oxide varistors fitted in a panel breaker housing for easy plug-in connection

The branch circuit SPD consisted of a simple MOV disc incorporated in a modified plug-and-receptac

combination, probably the first attempt at packaging an MOV for residential surge protection.

____________________________

*The experimental work, reported by F.D. Martzloff, involved performing the tests , recording of nearly 300 Polaroid oscillograms, 

was conducted by K.E. Crouch at the General Electric High Voltage Laboratory prior to his change to Lightning Technologies, Inc.
;ENERAL ELECTRIC COMPANY 

ZORPORATE RESEARCH AND DEVELOPMENT 

ichenectody, N.Y. 

Number MOR-7

LIGHTNING PROTECTION OF RESIDENTIAL AC WIRING 

K.E. Crouch* and F.D. Marttloff 

Automation and Control Laboratory 

ABSTRACT 

New transient suppressors using metal oxide varistors offer improved 

protection of appliances and consumer electronics against overvoltages. 

This improvement, however, could be at the risk of imposing excessive 

duty on the suppressor in case of a very severe lightning stroke near the 

house where these suppressors are installed. 

A simulated house wiring system was subjected to three levels nf !Ig.h.tging 

currents injected into the ground wires (moderate, severe, extremely 

severe), with various combinations of suppressors installed alone or in 

a coordinated com bination. 

I---+ ---..I*- a*.. *L-* -- -xz--*:..- --A --I- ---- kc3L I C3UlL3 S~UW LII~L a11 C~LCCLIVC a11u ~d~e cor~~bination oi devices can 

be specified for full protection of the loads in the house. 

"Lightning Technologies, Inc., Pittsf ield, Massachusetts 

. - <. . -. 

GENERAL @ fLCC1RIC I CLASS 3 - LIMITED AVAILABILITY INFORMA

Or F 0 UarIz1oII 

Room 8162. BuslJmg 220 

Natrcnel Bureau 01 Slsndmlr 

G~olharrbur~, t.40 2-59 

I information. 

Subsequent Company availability requires originat

approval. 

) Disclosure 0-utside General Eiectric Company requ

I of orlglnatlng component. - 

I 1 
LIGHTNING PROTECTION OF RESIDENTIAL AC WIRING 

K.E. Crouch* and F.D. Martzloff 

INTRODUCTION 

The development of metal oxide varistors has opened new opportunities for transient suppression in residential power circuits. The Wiring Device Department of the 

General Electric Company has introduced the VSP-I protector, which contains a 14 mm 

'n 

GE-MOV varistor. The HLP (Home Lightning Protector) has been available for many 

years, but the hot-line work required for its installation has been a deterrent; and, consequently, this protector has not been very widely applied. The new 32 mm GE-MOV @ 

varistor offers higher capacity than the 14 and 20 mm discs. Prior to reassignment of 

the product scope to the Distribution Transformer Department and later the Circuit Protective Devices Department, tests made in Pittsfield by J.S. Kresge had demonstrated 

that this 32 mm disc could meet the ANSI secondary requirements. By different packaging, the hot-line work might be eliminated and performance improved, opening the 

opportunity for greater acceptance. 

Therefore, the possibility of a coordinated protection system in residential power 

circuits meeting ANSI requirements became a more likely prospect than an earlier investigation had predicted for coordination between the present design of the HLP and 

the VSP-1.") While there is little evidence that extremely high currents caused by lightning strokes enter far into the house wiring, it seemed worthy of investigation to postulate a condition of "severe" lightning discharge near the house and to attempt recording 

on a simplified model wiring system how the currents and voltages would be distributed. 

This report describes the assumptions, test procedures, results, and conclusions of such 

an investigation. 

"Lightning Technologies, Inc., Pittsfield, Mass. 

"Registered trademark of the General Electric Company 
ground conductor of the service drop, without direct injection into the phase wires, is 

sufficient to induce voltage in excess of the clearance withstand of wiring devices. The 

transmission characteristics of the model and the relative sparkover levels were such 

that internal devices (receptacles) flashed over before the watt-hour meter gaps could 

flash over. 

Coordination between a centrally located surge arrester and an outlet-connected 

protector is possible; substantial, but within rating, currents flow in the outlet protector 

(VSP-1) when coordinated with a Home Lightning Protector (silicon carbide and gap) or 

its candidate successor, the 32 mm GE-MOV " varistor. 

For extreme strokes (1 00 kA at the pole), current in excess of rating can flow in 

VSP-1 protectors located close to the service entrance without other arresters. While they 

could fail there, the protectors do not present a greater hazard than an air clearance, which 

would flash over were there no protector; and, in fact, the presence of the VSP-1 is more 

likely to reduce the hazard of a flashover with subsequent 60 Hz power-follow. 

The addition of a 32 rnm varistor to the system, either in a plug-in (inboard) version 

or a s a n external addition (outboard) to the load center, will provide protection consistent 

with the ANSI requirements for secondary arresters. 

1.0 ASSUMPTIONS 

1.1 Current Magnitudes 

It was postulated that a lightning stroke attaching to the primary side of an overhead distribution system would produce a branching of the current flow into the ground, 

following sparkover of the surge arrester, which was presumed connected at the polemounted distribution transformer. Figure 1 shows the assumed circuit and the division 

of current flow. 
LA : PRIMARY ARREST 

XF DIST. TRANSFORMER 

M ' METER 

Figure 1. Division of Current Assumed for a 100 kA Stroke 

In their study of lightning environments, Cianos and indicate that only 

5% of all ground strokes exceed a peak current of 100 kA. The frequency of the strokes 

is quite dependent upon geographic location (isokeraunic levels),(3) as well as upon local 

configurations. An average expectation of a stroke involving the utility pole near a house 

with no adjacent tall trees or buildings may be in the order of one per 400 years for most 

of the U.S. Thus, for a 5% probability, the likelihood is one stroke in excess of 100 kA 

per 8,000 years. With nearby tall objects, this likelihood can be reduced I0 times; in 

areas of high lightning activities, this likelihood can be increased 10 times. The level 

of 100 kA, then, represents an expectation of being exceeded at one location only one 

time in perhaps 10,000 years (but there are millions of poles in the U.S.). 

From these assessments, the maximum current to be injected for the house model 

under discussion was selected to be 30 kA. From this maximum of 30 kA injected into 

the ground wire of the house service drop, two more values were used during the test 
series: 10 kA, corresponding to the requirement for the ANSI high-current, short-duration 

test; and 1.5 kA, corresponding to the requirement for the ANSI duty-cycle test - both 

specified by ANSI Standard C 62.1 for secondary valve arresters. (4) 

Another reason for selecting this low level (1.5 kA) is that no sparkover occurs 

in the wiring at this level. For the 10 and 30 kA levels, multiple flashovers would occur 

at variable times and locations, making exact duplication of tests impossible. By staying 

below sparkover levels, repeatability of the results was ensured, allowing comparisons 

among several alternate circuit configurations. 

1.2 Waveshape 

From ANSI Standard C 62.1, a waveshape of 8 x 20 ps would have been desirable. 

However, limitations in the test circuit required for driving 30 kA in the model loop forced 

a compromise of 10 x 25 ps as the test wave. 

1.3 Lightning Current Path 

It should be noted that, in this test series, the assumption was made t h a t t h e lightning current, applied first to the distribution primary (the highest wire on the pole) is 

transferred to the ground system by sparkover of an assumed surge arrester on the primary 

at the pole. In fact, if there were no arrester, an equivalent effect by direct flashover 

could be expected. 

For the secondary side, however, the assumption was made that both sides (phases) 

of the center-tapped (grounded) secondary remained uninvolved in conducting the direct 

lightning current, while the ground wire (messenger) from pole to house carried its share, 

as defined in Figure 1. 

1.4 Induced Voltages 

The generation of transient voltages in the house is attributed to electromagnetic 

coupling of the field established by the lightning current flowing in the messenger into 

the loop formed by the two phase wires encircling the messenger. In addition, there is 

some capacitive coupling between the wires (Figure 2). 
30 kA 

DISTRIBUTION 

TRANSFORMER 

BY 

= GROUND PLANE 

Figure 2. Voltages Induced in the House Wiring Systems 

2.0 TEST CIRCUIT AND TEST PROCEDURE 

2.1 Power Circuits 

The test circuit consisted of a high-current impulse generator, a distribution transformer with service drop, a simulated simplified house wiring system, and the necessary 

shielded instrumentation (Figure 3). Details of the catalog numbers, characteristics, etc., 

are given in the Appendix. 

The service drop connection between the distribution transformer and the meter 

socket was made with three AWG 416 wires, twisted at a pitch of about 5 turns/m (1.5 

turnslft), 13 m (45 ft) long. This service drop was folded in a loose "St shape, at about 

0.5 m (1.5 ft) above the ground plane serving as the return path for the lightning current, 

in order to reduce the loop inductance seen by the generator. This configuration does 

not influence the coupling between the messenger and the wires wrapped around it, 

coupling which has been identified as the voltage-inducing mechanism. 
GENERATOR I AND SERVICE SHOP 

I 

I 

I 

I 

I 

I 

I 

LIGHTNING I MODEL DISTRIBUTION I MODEL HOUSE I SHIELDED 

CURRENT I TRANSFORMER I INSTRUMENT I ROOM 

METER WITH I 

GAPS I 

I 

Dl FFERENTIAL cR,o#J l 

I \SURGE ARRESTER 

I OR JUMPER WIRE 

I 

L 

I -- I 

I 

I 

I 

I 

I 

I TO VOLTAGE 

I GROUND I , 

MEASUREMENT 

POINTS 

PLANE , 

I 

I 

I I 

Figure 3. Schematic Representation of Test Circuit 

The simulated house wiring started at the meter socket and continued to a load 

center by a 3 m (1 0 ft) length of AWG #6 aluminum entrance cable. The meter socket, 

watt-hour meter, and load center were typical General Electric Company hardware (see 

Appendix), except as noted in the detailed procedure description. From this load center, 

four "branch circuits" connected to the load center breakers were established, each terminating at a wall receptable mounted on the same 1.2 by 2.4 m (4 by 8 ft) plywood panel 

on which the watt-hour meter and load center were also mounted. The branch circuits' 

lengths were (one each) 6, 12, 24, and 48 m (20, 40, 80, and 160 ft), the wire being loosely 

coiled between the load center and receptacles (Figure 4). 

2.2 Instrumentation 

Recordings of currents and voltages were made at several points on the wiring 

system with cathode ray oscilloscopes (CRO); differential measurements were made for 

the voltages with especially built 100:l probes. These probes were built by placing a 

5000 52 resistor in series with a terminated 50 fi coaxial cable - all of these contained 
Figure 4. Connections at Load Center 

in a shield tied to the ground plane part of the shielded instrument room. Currents flowing 

in the suppressors were measured by means of a Pearson Model 110 A wide-frequencyband current transformer. The oscilloscopes were located inside the shielded control 

room adjacent to the test area, providing satisfactory protection against spurious signals 

(see Figure 6 in Section 2.4). 

2.3 Candidate Suppresson 

Four candidate suppressors were installed at various locations in the system, for 

various comparisons of performance: 

1. One Home Lightning Protector (HLP, GE Cat. 9L15DC 8002) was installed 

at the bad center; when connected to the circuit, the connection was at 
the incoming lugs of the load center, as it would normally be when connected 

by an electrician. 

2. Two V250HE8O varistors were mounted near the load center and connected 

to the incoming lugs of the load center. This connection required about 

45 cm (18 in) of #I0 copper wire. The return to ground was common to 

the two discs, as it is for the HLP device. 

(The varistor package contains a 32 mm disc with characteristics suitable 

for secondary arrester duty.(4) It is the candidate metal-oxide varistor 

substitute for, or successor of, the Thyrite "- gap combination currently 

used in the HLP, and has an RMS voltage of 250 V.) 

3. Two 32 mm varistor discs of the same characteristics as (2) (above) were 

installed by the Circuit Protective Device Department in a breaker housing 

so that they could be connected to the load center bus with a minimum 

of lead (10 cm, or 4 in). This connection can be made while the load center 

is energized without requiring "hot work," in the same manner as inserting 

additional breakers on the load center. 

4. VSP-I spike protectors, produced by the Wiring Device Department, were 

inserted in the receptacle at the end of the branch circuits. (The VSP-I 

protector contains a 14 mm GE-MOV" varistor with a voltage rating of 

170 V RMS.'~) 

5. In addition, the meter contained its standard gaps rated for a 10 kV sparkover. 

2.4 Test Procedure 

Preliminary tests indicated that flashover at the receptacles would occur with 

10 kA injected into the ground messenger, but no sparkover of the meter gaps was apparent. 

Therefore, a first test series was conducted at only 1.5 kA in order to provide consistent 

patterns of wave propagation undisturbed by flashover (Figure 5). 

It was also found that the auxiliary impulse generator used to trigger the main 

gap of the high-current generator induced voltages into the test circuit that could exceed 

those induced by the main discharge. A mechanical switch for closing the circuits was 

then substituted for the triggered gap. 

"Registered trademark of the General Electric Company for molded composite dielectric 

material 
VERTICAL - 500 A/di v 

HORIZONTAL - 5 us/div 

Figure 5. Applied Current Waveshape - 1.5 kA Crest, 10 x 25 LI s 

Noise checks were made for the voltage measurement system by shorting the probes 

together and attaching them to the neutral point on the circuit under test. Similarly, 

the center conductor of the cable to the current transformer was removed from the transformer output and connected to its sheath. No significant voltages (greater than 5% of 

measured signal) were measured. A typical noise check oscillogram is shown in Figure 6. 

VERTICAL - 5 V/div 

HORIZONTAL - 2 us/div 

Figure 6. Typical Noise Response 

of Measurement System 

with 1.5 kA Injection 

3.1 Test Results 

A large number of tests werg performed to investigate the effects of various combinations. From several hundred recorded oscillograms, a selection was made, as shown 

in this section, to illustrate these effects. The results are presented in the form of a 

matrix of oscillograms with corresponding commentary, generally providing a comparison 

of voltage and currents with or without protectors installed. First, a qualitative summary 

is presented, then some comparative oscillograms are shown to illustrate various effects. 

Figure 8 gives a qualitative summary of the effects obtained by installing a single 

protector at various locations in the system. The oscillograms are arranged in horizontal 

rows corresponding to the circuit configuration indicated in the legend. The vertical 

columns correspond to the location at which the oscillograms were recorded. From left 

to right appear Lines I and 2 of the load center, and the ends of the branch circuits at 

6, 12, 24, and 48 rn (20, 40, 80, and 160 ft), which will be referred to as B20, B40, B80, 

and B160. Quantitative information will be given in subsequent figures. 

In the first row, open-circuit voltages are shown. Note that the voltages at three 

locations of the Line 1 conductor are very similar, while there is a small difference between Line 1 and Line 2. 

The installation of a protector in Line I of the load center (second row of oscillograms) clamps the voltage on all Line 1 points, with some oscillations induced at the 

end of the B160 branch. While the initial peak of the Line 2 points is not changed, subsequent oscillations have lower frequency than in the open-circuit mode. For the oscillograms corresponding to the location where a protector is installed, the upper trace shows 

the current flowing in the protectors. 

The installation of a protector in Line 2 of the load center (third row) produces 

results analogous to the Line 1 case. Installation of a protector at the end of a branch 

rather than at the load center (last four rows) produces clamping of the voltage at the 

point of installation. At the other points of the same line, the effectiveness of the clamping decreases a s t h e protector is farther away. For the line with no protector, there 

is a minor voltage reduction and a frequency change similar to that noted in the first 

two rows. 
! ~rn ~ine *I /@ @I 

~ine*l-lZO~-ti20~ -Line 9 Outlet Box 

i 

! I - - 

Outlet Bnr 

_-( from Line *2 10 a] from Line4i 10 @I ~- . -. 

,- - 80 ft. Branch 1 I 1 1.4 from ~ine #2 1-1Outlet Box AWG "14-2W/Grd 

L------ 40 ft. Brunch 

AWG"i2-2W/Grd - - I-240~-I 20 ft. Bronch 

AWG *14-2W/Grd 

- - Outlel Box 160 ft. Branch 

AWG 812-2W/Grd 

A VSP-I in 

Load Center 

Line #I anly 

A VSP-l in 

Load Center 

Line *2 only 

Simuloted Household Wiring System With Measurements at Various Branches With and 

Without a VSP-I Protector Installed at a Single Paint in the System. 

1500 Amperes, 10 x 25 Microsecond Current Pulse Applied to Service Neutral. 

Figure 8. Summary of Protector Effects 

VOLTAGE AT BUS WITH H1 OF 

TRANSFORMER TIED TO 

NEUTRAL (GROUND) 

Current in HLP: 40114/div 

Voltage Across Bus: 

501) V/div 

VOLTAGE AT BUS WITH AIR GAP 

BETWEEN H1 OF TRANSFORMER 

AND NEUTRAL 

Current in HLP: 400 A/div 

Voltage Across Bus: 

500 V/div 

VOLTAGE AT BUS WITH 

DISTRIBUTION ARRESTER ON 

TRANS FORMER PRIMARY 

Current in HLP: 400 A/div 

Voltage Across Bus : 

500 V/div 

- - 

CURRENT INJECTED FOR 

TEST 157 

(The arrester affects 

the front of wave of 

the current) 

Current i n Arrester : 

2 kA/di v 

(Sweep: 5 vs/div) - 

Test Condition: 10 kA injected. 

A1 1 sweeps : 2 ps/div, except 7 571 

Figure 9. Comparison of Protector HLP Response for Various Protective Devices at the Primary of the Distribution Transformer 

15 

(LINE 2) 

265/267 

VOLTAGE AT OUTLET WITH 

(1) OUTBOARD PROTECTOR 

ONLY 1 AT EACH BUS 

Voltage at Outlet: 

500 V/div - 

VOLTAGE AT OUTLET WITH 

(1) OUTBOARD PROTECTOR 

AT EACH BUS, PLUS (1) 

"100 W LOAD" AT 8-1 60 

Voltage at Outlet: 

500 V/div 

266A 

VOLTAGE AT OUTLET WITH 

(1) OUTBOARD PROTECTOR 

AT EACH BUS, PLUS (1 ) 

VSP-1 AT EACH 8-80 AND 

B-160 OUTLET 

266/268 

VOLTAGE AT OUTLET WITH 

(1) OUTBOARD PROTECTOR 

AT EACH BUS, PLUS ( 1 ) 

VSP-1 AT EACH B-80 AND 

B-160 OUTLET 

Current in VSP-1 : 500 A/div 

voltage at Outlet: 

500 V/div 

Test Condition: 70 kA injected - Protectors and loads. 

All sweeps: 2 us/div, except as noted. 

Figure 10. Comparisons of Performances with Various Devices at Outlets, All with Protectors at Load Center 
visible during the first loop in each oscillogram are spaced apart at a distance which 

increases with line length. Thus, one can conjecture that these may be caused by 

reflections. 

Loading the line termination with a 130 S-2 resistor (Oscillograms 243A, 245, 247, 

and 249) eliminates the later oscillations and reduces the first peak to about 60% of the 

value without load. From this reduction, a Thevenin's calculation of circuit parameters 

(Figure 12), if applicable in an oversimplified form, would show that 130 nis 60% of the 

total loop impedance. Hence one can conclude that the source impedance is four/sixths 

of 130, or about 85 n. 

When a VSP-1 protector is added to the 130 fi resistor (Oscillograms 244, 246, 248, 

and 250), the clamping action of the varistor limits the voltage at the outlets to about 

400 V, which is consistent for the currents of about 20 A flowing in the varistor. 
Voltage at Outlet 

Indicated: 500 V/div 

VOLTAGE WITH (1) "100 W" 

INSTALLED AT OUTLET 

INDICATED 

I voltage at Outlet 

Indicated: 500 V/div 

VOLTASE AND CURRENT !IITH I 

(1) vSP-1 and (I) 

"100 W" INSTALLED AT 

OUTLET INDICATED 

Current in VSP-1 : 

23 A/div 

Voltage Across VSP-1: 

-- 

Test Condition: 1 .5 kA injected - 

~rotectors and load installed at outlets as shown. 

All sweeps: 2 us/div. 

Figure 11. Effect of Load (1 00 W Light Bulb) on Voltages at Branch Outlets 

Figure 12. Thevenin's Equivalent for Oscillogram 202 
lograms 261 and 262 of Figure 13. 

With outboard protectors rather than inboard protectors, the peak voltages are 

in the 1000 to 11 00 V range (Oscillograms 263 and 264). These higher voltages are attributable to the longer leads required to connect the outboard protector, compared to the 

inboard protector. (Figure 16 shows a comparison of lead length effects, which removes 

any question that the difference between inboard and outboard protectors might have 

been the result of an intrinsic difference in the varistors.) 

While not a recommended installation location, two VSP-1 were also installed 

directly at the load center (on the bus) in an arrangement that approximates the "inboard 

protector" geometry. Oscillograms 255 and 256 show the clamping voltage at 500 to 600 V 

with current crests at 11 00 to 1200 A. Scaling up these varistor current values for higher 

lightning currents than the 10 kA injected would indicate probable excessively large currents in the 14 mm varistor used in the VSP-l protector. 

Finally, a HLP protector was installed at the center, as shown for one bus on 

Oscillogram 153. The voltage is higher and the initial rise before sparkover of the gap 

takes place at about 2.2 kV. The current crest, after the sparkover, is of the same magnitude (1 100 A) as that of other tests. 
BUS 1 

(LINE 1) 

BUS 2 

(LINE 2) 

OPEN-CIRCUIT VOLTAGE 

Voltage at Bus Indicated: 

2 kV/div 

VOLTAGE AND CURRENT WITH 

(1 ) INBOARD PROTECTOR AT 

EACH BUS 

Current in Protector: 

500 A/div I 

Voltage Across Bus: 

500 V/div 

- 

VOLTAGE AND CURRENT WITH 

(1 ) OUTBOARD PROTECTOR 

AT EACH BUS 

Current in Protector: 

500 A/div 

Voltage Across Bus: 

505 V/div 

VOLTAGE AND CURRENT WITH 

(1 ) VSP-1 AT EACH BUS 

Current in VSP-1: 

500 A/div 

Voltage Across VSP-1 : 1 500 Y/div 

I VOLTAGE AND CURRENT WITH 

(1) HLP ON LOAD CENTER 

Current in HLP Line 1 : 

400 A/div 

i I 

Voltage Across Bus: ) 500 V/div 

Test Condition: 10 kA injected - All sweeps: 2 ~s/div. 

Figure 13. Comparison of Alternate Protectors at Load Center 

2 1 
those of the envelope, being at 400 to 600 V, with initial bursts at 800 to 1100 V. 

In the case of Oscillogram 284, flashover of the wiring at the load center limited 

the current impressed on the VSP-1 protector. This is a result of an unintentional wiring 

flashover, which occurs frequently. 
(F1 ashover at B-20 

1 imi ts the current.) 

VSP-1 AT OUTLET SHOWN WITH 

OUTBOARD PROTECTOR AT 

LOAC CENTER 

Current in VSP-1 : 

200 A/div 

Voltage at Outlet: 

500 V/div 

Test Condition: 30 kA injected 

A1 1 sweeps : 2 us/div. 

I 

Figure 14. Comparison of Duty Imposed on VSP-1 Installed at Outlets 

for Various Load Center Protections 

NO OTHER PROTECTOR 

Current in VSP-1 : 

1000 A/div 

Voltage at 0ut:et: 

500 V/div 

500 V/div 

(1) OUTBOARD PROTECTOR ON 

EACH BUS 

Current in Protector 

1000 A/div 

Voltage Across Bus 

500 V/di v 

Test Condition: 30 kA injected - 

protectors at load center. 

All sweeps: 2 vs/div. 

Figure 15. Comparison of Performances Between Inboard and Outboard 

Protectors Installed at Load Center 
The total lead length (60 crn, or 24 in) was approximately equal to that involved in 

separately connecting the outboard protector (45 crn, or 18 in) and the inboard protector 

(15 cm or 6 in) at the load center of the simulation circuit. Current pulses of constant 

magnitude (3.2 kA crest, lops rise time) were injected in the loop, and voltages across 

the protectors and their corresponding leads were recorded as shown in Figure 16. 

Oscillogram 1 shows a 1000 V maximum voltage across the outboard protector 

and its associated 45 cm (1 8 in) lead, compared to only 600 V for the inboard protector 

and its 15 cm (6 in) lead (Oscillogram 2). Changing the lead of the inboard device to 

45 cm (1 8 in) (Oscillogram 3) raised the voltage to 1000 V, demonstrating that the difference is attributable to lead length, not disc characteristics, and illustrating the benefits obtainable by making the protector an integral part of the load center. 
I Vol tage: 500 V/di v I I Current: I000 A/div I 

VOLTAGE ACROSS "INBOARD 

SUPPRESSOR" PLUS 6" 

LOOP OF WIRE 

Vol tage: 500 V/di v 

Current: 1000 A/div 

VOLTAGE ACROSS SAME 

SUPPRESSOR AS (1 ) BUT 

18" LOOP OF WIRE 

I Voltage: 500 V/div I I Current: 1000 I Aldi v 

Test Condition: Laboratory bench, not house simulation. 

Current injection in the two suppressors 

connected in series: 2.7 kA. 

All sweeps: 2 us/div. 

Figure 16. Effect of Lead Inductance on Clamping Voltage 

Test Condition: Injected current 10 kA. 

A1 1 sweeps : 2 vs/di v . 

Figure 17. Voltages at Load Center with VSP-1 and "100 W load," at Remote 

Outlets, No Protector at Load Center 
3.2 Discussion of the Results 

From the oscillograms collected during the test series, a summary of maximum 

values has been compiled as shown in Table 1. The three levels of current injection are 

included in this table, illustrating a mild, severe, and extremely severe lightning incident 

near the house. For the sake of simplicity in this table, only one protector is included 

in the arrangement matrix. A subsequent discussion will address the case of a coordinated scheme involving more than one protector. 

Injection of the maximum values recorded on the protector current shows that 

no rated values are exceeded, even at 30 kA injection. Voltages observed are consistent 

with the corresponding currents, from the V-I characteristics of the varistors. 

In the first group at 1.5 kA injection, voltages that are particularly damaging to 

appliances (2500 V) are observed throughout the system. Installation of a protector (on 

both lines) at the load center eliminates the hazard and would suffice to protect all of 

the house. Installation of a VSP-1 at only one close or remote outlet provides protection 

at that outlet and moderate protection on all locations of the same line. The other line 

is not protected. 

The unsymmetrical load (diode) of the TV input circuit behaved in a predictable 

mar,ner: when the polarity of the voltage was such that a forward bias was applied, the 

diode clipped the voltage, with the series resistance limiting the current. With reverse 

bias polarity, the diode failed when the 2500 V transient occurred at that outlet. 

In the second group, representing a severe incident, flashover can be expected 

throughout the system in the absence of protection, with the associated fire hazards as 

well as damage to electronics during the initial voltage rise. Installation of a HLP at 

the load center eliminates the flashover hazard but does not lower the voltage sufficientl

to assure protection of sensitive electronics, nor does an outboard installation of varistor

assure protection. Installation of an inboard set of varistors is effective, for the voltage 

is limited at the load center (and consequently on the whole system) to 700 V. Installatio

of a VSP-1 at a close outlet (producing the maximum current flow, hence highest voltage

is effective for that outlet only; on the basis of the differential observed at 1.5 kA, one 

can presume that the voltage at the load center would be too great to consider any other 

point but that outlet as being protected. 

g

None 1.5 2500 2500 - Inboard Protector at Load Center 1.5 500 500 100 

Outboard Protector at Load Center 1.5 700 700 70 

VSP-1 at B-20 1.5 800 400" 30 

VSP-1 at B-160 1.5 1100 350* 20 

None 10 8800 F/O - Inboard Protector at Load Center 10 700 700 1200 

Outboard Protector at Load Center 10 1200 2000 11 00 

HLP at Load Center 10 2200 NR 1100 

VSP-1 at B-20 10 NR 700" 950 

VSP-1 at B-160 10 NR 600" 600 

Inboard Protector at Load Center 30 950 NR 3500 

Outboard Protector at Load Center 30 2200 NR 3500 

VSP-1 at 8-20 30 NR 800 2500 

VSP-1 at B-160 30 NR 500 250 

Notes - 

NR- No Record 

* "Protected outlet" is on same line as protector. Voltages on outlets of the 

other line close to unprotected values. 

t Only one protector at a line in this table. 
ter bus bars with very short leads. 

Until such an integral package is marketed for new systems, a coordinated protection scheme can be implemented, as a retrofit, that would still provide reliable protection for millions of sensitive appliances in existing systems. 

The coordination involves a protector at the load center, either the commercially 

available HLP or a packaged 32 mm disc set (two lines) with reasonably short leads in 

a package similar to the HLP. This protector will limit the voltage at the load center 

to about 2200 V. This 2200 V level is below the flashover level of the wiring but can 

still cause damage to sensitive appliances. The currents passing through the protector 

at that location will not exceed the protector capability. In addition, VSP-1 protection 

should be installed at those outlets where a sensitive appliance is plugged. The voltages 

allowed by the VSP-1, typically 400 to 600 V, will be low enough to assure survival of 

all but excessively sensitive appliances, while the VSP-I will not be exposed to currents 

that can lead to a failure in case of frequent exposure to severe lightning incidents. 

Thus, a coordinated protection scheme is technically feasible. The cost should 

be acceptable to do-it-yourself homeowners, although it might be a deterrent to those 

owners who have to call in an electrician to install a protector at the load center. Based 

on increasing awareness in the technical and regulatory agencies community of overvoltage protection, the incorporation of protection to load centers offers the best approach 

to new installations. 


COORDINATION

OF

SURGE

PROTECTORS

IN

LOW-VOLTAGE

AC POWER

CIRCUITS
François D. Martzloff 

General Electric Company 

Schenectady NY 

f.martzloff@ieee.org

Reprinted, with permission, from IEEE Transactions on Power Apparatus and Systems, PAS-99, Jan 1980 

First presented at IEEE Summer Meeting, Vancouver, July 1979

Significance: 

Part 4 – Propagation and coupling of surges 

Part 8 – Coordination of cascaded SPDs 

This paper presents a summary of two earlier and detailed proprietary General Electric reports describing 

experiments conducted in Schenectady NY and in Pittsfield MA, respectively by Martzloff and Crouch. (These 

have now been declassified by General Electric and are included in this Anthology – see Coordination 1976

and Propagation 1978.) The prime purpose of that paper at the time was to report in a non-classified 

platform experimental results that could be useful for the development of IEEE Std 587 (later known as IEEE 

Std C62.41).

In the first experiment, a simple test circuit of two branch circuits originating at a typical service entrance paper 

was subjected to relatively high-energy unidirectional impulses, with various combinations of surge-protective 

devices installed at the service panel and/or at the end of the branch circuits. That 1976 experiment was the 

beginning of recognition of the “cascade coordination” issue that became the subject of intense interest in the 

80’s and 90’s (see the listing of contribution by many authors in Part 1, Section 8). 

In the second experiment, the coupling and subsequent propagation of surges was investigated in a more 

complex circuit that included a distribution transformer, service drop, entrance panel, and several branch 

circuits. The surge was injected in the grounding system, not into the phase conductors. This 

experiment thus brought new evidence that ring waves can be stimulated by unidirectional surges. 

Nevertheless, the threat was considered at that time as a surge impinging onto the service entrance from the 

utility, not resulting from a direct flash to the building grounding system. On that latter subject, see Dispersion

and Role of SPDs. 

This paper received the 1982 Paper Award from the Surge-Protective Devices Committee. 

Filename: Coordination 1980
COORDINATION OF SURGE PROTECTORS IN LOW-VOLTAGE AC POWER CIRCUITS 

F.D. Martzioff, Member, IEEE 

General Electric Company 

Schenectady, NY 12345 

- Surge protectors can be installed in low-voltage ac power systems 

to limit overvoltages imposed on sensitive loads. Available devices offer a 

range of voltage-clamping levels and energy-handling capability, with the 

usual economic trade-off limitations. Coordination is poss~ble between lowclamping-voltage devices having limited energy capability and high-clampingvoltage devices having high energy capability. The paper gives two examples 

ofcoordination, as well asadditionalexperimental resuitson surge propagation. 

1. INTRODUCTION 

Surge voltages occurring in low-voltage ac power circuits have two origins: 

external surges, produced by power system switching operation or by lightning, 

and internal surges, produced by switching of loads within the local system. 

Typical voltage levels of these surges are sufficient to cause the failure of sensitive electronic appliances or devices, and high surges can cause the failure 

of rugged electromechanical devices (clocks, motors, and heaters) [I ,2]. 

For many years secondary surge arresters from a number of manufacturers 

have been available. These arresters are effective in protecting nonelectronic 

devices against the high-voltage surges associated with lightning or power system switching. However, the voltage allowed by an arrester is still too high for 

sensitive electronic devices. Furthermore, installation requires an electrician 

to connect thedevice on hot terminals. 

The advent of the metal oxide varistor packaged as a convenient plugin 

device or incorporated into the appliances makes possible a voltage clamping 

which is more effective than that of the conventional secondary arrester. However, ihe energyhandling capability of such packages is lower than that of an 

arrester, so that large currents associated with lightning strikes cannot be 

handled by these packages. 

The availability of these two different types of suppressors now makes it 

possible to obtain a coordinated protection of all the appliances in a home or 

all the equipment in an industrial environment. Improper coordination, however, could force the lower voltage device to assume all the current, leaving the 

high-energy protector uninvolved; this situation could then cause premature 

failure of the low-voltage suppressor. This paper discusses the elements of a 

coordinated protectivesystem based on experimentation. 

11. SECONDARY ARRESTERS AND LOW-VOLTAGE 

SUPPRESSORS 

Typical secondary arresters for 120 V service consist of an air gap in series 

with a varistor made of silicon carbide. The device is generally packaged with 

two arresters in the same housing; the physical arrangement is designed for installation on the outside of a distribution panel, through a knockout bole of 

the panel enclosure or at the entrance to the building. 

Limitations on the gap design imposed for the purpose of reliable operation and clearing after a high current discharge(10 kA, 8 x 20)do not allow the 

sparkover of the gap to be less than about 2OOO V. This sparkover and the 

time required to achieve it allow injection of a potentially damaging surge into 

the "protected" power system downstream from the arrester: While this 

2000 V level provides better protection than the protective characteristics indicated in ANSI standards [3], lower voltage clamping is desirable for the 

protection of sensitive electronics. 

*In this paper the high-energy suppressor, typically installed at the service 

entrance, will be called arrester. The low-energy, low-voltage suppressor, 

typically installed at an ourlet or incorporated into an applianceor connected 

load, will be called suppressor. 

F 79 635-4 A paper recamended and approved by the 

IEEE Surye Protective Devices Cornnittee of the IEEE 

Power Engineering Society for presentation at the IEEE 

PES S-r Meetinq, Vancouver, British Coldia, 

Metal oxide varistors suitable for I20 V line applications can clamp surge 

voltages at less than I000 V, typically at 500 to 600 V for surge currents of less 

than 1000 A. These varistors provide excellent protection for electronic systems. The economics of device size, however, limits the wlde use of large varistors, especially since smaller varistors can do an acceptable job if they are not 

exposed to excessive currents. Proper coordination among the devices used 

is required to obtain a reliable protect~on system. 

111. PROTECTION COORDINATION 

While the installation of surge protective devices functions effectively for 

high-voltage utility systems coordinated by centralized engineering, tbecurrent 

trend toward regulatory installation in low-voltage systems, because they are 

seldom centrally engineered and coordinated, can result in damaged equipment and system failure. The successful application of protective devices to a 

low-voltage system demands a perspective of the total system, as well as a 

knowledge of individual device characteristics. Where such knowledge and 

coordination are lacking, a low-voltage suppressor installed in conjunction 

with an arrester can prevent the voltage at the terminals of an arrerter from 

reaching its sparkover level. As a result, all of the surge current may be forced 

into the suppressor, wh~ch may not have been intended to withstand extreme 

conditions. 

Proper coordination in an arrester/suppressor system requires some impedance between the two devices. This impedance is generally provided by the 

wiring: at the beginning of the surge, the rapidly changing current produces 

an inductive voltage drop in this wiring, in addition to the drop caused by the 

resistanceof the wiring. Thus, thevoltageat theterminalsof thearrester during 

the current rise of the surge is equal to the clamping voltage of the suppressor, 

plus the voltage drop in the line (tests reported below indicate that this voltage 

drop is indeed appreciable). This voltage addition can then raise the terminal 

voltage of the arrester sufficiently to reach sparkover. In this way the arrester 

will divert most of the surge current at the entrance, rather than permitting it 

to flow in the suppressor. 

The application of a suppressor alone is likely to occur because electronic 

appliance manufacturers increasingly provide suppressors incorporated into 

their products. With no arrester at the service entrance, the wiring clearances 

can become a voltage-limiting device, thus establishing a clearance/suppressor 

system. The suppressor would again tend to assume all of the surge current 

flow. The voltage drop in the line, in a manner similar to that of the arrester/ 

suppressor system, would raise the voltage at upstream points to levels that 

may spark over the clearances of wiring devices, providing unplanned rehef 

for the suppressor. When sparkover of the clearances occurs. there are three 

possible results: 

a. A power-follow current occurs, with destructive effects on the 

components. 

b. A power-follow current occurs, but overcurrent protection (breaker 

or fuse) limits the damage. The system can be restored to operation 

after a mere nuisance interruption. 

c. No power-follow current takes place; the overvoltage protective 

function of the system can be considered as accomplished. 

The concept of protecting solid insulation by allowing clearances to spark 

over first is actively promoted by the Low Voltage Insulation Coordination 

Subcommittee of the International Electrotechnical Commission [4]. 

Further discussion of it is outside the scope of the present paper; nevertheless, 

the concept is worth attention because cost reductions and system reliability 

could be obtained through its proper application. 

Two examples of protection coordination will now be discussed in detail. 

These examples represent two scenarios on surge injection; they are based on 

experiments involving an arrester and suppressors in simulated lightning surge 

conditions. In the first scenario the surge is assumed to be ~njected between 

one of the phase wires and the center conductor (ground) of the service entrance. In a second scenario thesurgecurrent is assumed to be injected directly 

into the ground system of a service entrance only. Both experiments show the 

benefits and importance of proper coordination. In both tests the arrester was 

a gap-silicon carbide combination (Fig. I) and the suppressor, a metal oxide 

varistor in a plugin package (Fig. 2). 

IV. SURGE APPLIED BETWEEN PHASE AND GROUND 

Test Circuits 

The test circuit (Fig. 3) consisted of a terminal board from which two lines, 

one 7.5 m (25 ft )long and the other 30 m (100 ft) long were strung in the test 

-. area. A short, 3 m (10 ft), line simulated the service dro~. All of these lines 

Canada, July 15-20, 1979.~anuscri~t submitted February 6, were made of three-conductor, nonmetallic. #I2 AWC sheath wire. The neu1979; made available for printing April 3, 1979. tral and ground wires of the three lines were connected together at the terminal 

board and from there to the reference ground of the test circuit. 

0018-9510/80/0100-0129$0000750 I980 IEEE 
(a) (b) 

Fig. 4. Open-circuit voltage and short-circuit current (without any protector). 

Fig. 1. Typical arrester for service entrance installation. 

Fig. 2. Typical suppressor for plugin installation. 

GR~EN 

IGROUNDl 

Fig. 3. Test circuit. 

All surge currents were applied between the line conductor (black) at the 

end of the service drop and the reference ground (green and white). These 

impulses were obtained from a 5 pF capacitor charged at a suitable voltage 

and discharged into the wiring system by an ignitron switch. The resultant 

open-circuit voltage waveform, a unidirectional wave of 1 ps rise time x 50 ps 

to one/half value time, corresponds to the standard test wave in utility systems. 

Fig. 4 shows typical open-circuit voltage and short-circuit current waveforms. 

Voltages were recorded by a storage oscilloscope through an attenuator probe 

(1000:l); currents, through a current probe and a current transformer. Thus, 

the calibrations displayed on the oscillogram are to be multiplied by 1000 for 

the voltage. The current traces show the 50 mV setting corresponding to the 

rated output of the current probe, with the amperes per division shown in 

parentheses corresponding to the current transformer ratio and current probe 

input setting for a direct reading. The sweep rate is also shown on the oscillograms, at IOps/div. for all the tests. 

Test Results 

Fig. 5a shows the voltage across the arrester when subjected to the surge defined by Figs. 4a and 4b. Note that the sparkover voltage reaches 2200 V, with 

several oscillations, before the voltage settles down to the impulse discharge 

voltage at about 2000 V at its start. 

Figs. 5b and 5c show, respectively, the voltage and current across the varistor in the suppressor. Note that the maximum voltage is 600 V for a 550 A 

current on the varistor. (The current in the suppressor is lower than the available short-circuit current as a result of the reduced driving voltage, because 

the varistor holds off 600 V. 

Fig. 6 shows several osc~llograms indicating how the surge propagates in 

the wiring in the absence of any wppressor. Fig. 6a shows the open-circuit 

voltage at the service box. At the open-ended 7.5 m (25 ft) line, the voltage is 

substantially the same as at the box (Fig. 6b). However, at the end of the 30 m 

(100 ft) line with a 50 Q termination, a significant decrease of the slope is noticeable, while the crest remains practically unchanged (Fig. 6c). 

(a) open-circuir volrage-ai box 

(b) open-clrcult vollage-7.5m (25 ill (cl open-clrcuil voltage - 30m (IW fl) 

Fig. 6. Propagation of surge 

With voltage limitmg at the box provided by the installation of a suppressor, even at a remote outlet, an arrester connected at the service box would not 

reach its sparkover voltage until rubstantial surge currents were involved. A 

larger current was requ~red for a short distance between the service box and 

the suppressor than for a greater distance. The value of the current required 

to reach sparkover as a function of the distance is therefore of interest. 

For a distance of 7.5 m (25 ft) the threshold condition for sparkover of the 

arrester is shown in Fig. 7. In Figs. 7a and 7b the open-circuit voltage and 

short-circuit current are shown for this threshold setting of the generator. Inspection of the oscillograms shows an open-circuit voltage of 8.1 kV, with a 

calculated equivalent source impedance of 4.2 R. This low value of the source 
impedance, compared to proposed values 151, provides a conservative evaluation of the system performance. For the same setting as Figs. 7a and 7b, the 

oscillograms of Figs. 7c and 7d show the case in which the arrester has sparked 

over, as indicated by its voltage (7c) and current (7d) traces. In Figs. 7e and 7f, 

the traces show the voltage (7e) and current (70 in the suppressor for a case in 

which the arrester did not spark over (as a result of the scatter of sparkover or 

a slight difference in the output of the surge generator). This case represents 

the most severe duty to which the suppressor would be exposed, for a distance 

of 7.5 m (25 ft) 

(a) opncircult voltagc (b) short-circuit current 

(c) voltageat arrester when arrester dm sparkover - suppressor at 7.5 rn (25 It) 

Id) rurrcnl marrcsIcr allcr ,prrk~,rr - 

wpprrsrur at 7 I m (25 It) 

(e) voltageal suppressor when arrestor does not sparkover - suppressor at 7.5 rn (25 It) 

(0 current in suppressor when arrester docs not sparkover - suppressorat 7.5 m (25 ft) 

Fig. 7. Transfer of surge conduction 

From these tests it is apparent that the I200 A flowing in the line to the 

suppressor (70 and establishing 1000 V at the varistor terminals (7e) causes an 

additional 1OOO V drop in the line. The resulting 2000 V appearing at the arrester terminals may cause sparkover of the arrester (7c). 

For a case in which there is no arrester installed at the box but only the 

suppressor installed at an outlet, the voltage rise in the wiring and the meter 

coils will most likely result in a flashover of the system, which would then divert the excessive energy away from the suppressor, just as the arrester did in 

the test. Of course, this diversion may be destructive, a result that the arrester, 

when ~nstalled, is prec~sely designed to prevent. 

For greater distances between the suppressor and the arrester, the transfer 

of the surge wtll occur at lower currents. For instance, w~th the suppressor 

installed at the end of the 30 m (100 ft) line, only 700 A were required in the 

suppressor to reach sparkover of the arrester. 

Discussion 

The tests on simulated high-energy surges indicate that a transfer occurs 

from the suppressor to the arrester at a current level which depends on the distance between the two devices. Even for a short length of wire, the suppressor 

is relieved from the surge by sparkover of the arrester before excessive energy 

can be deposited in the varistor of the suppressor. At lower current levels, 

where the voltage in the system is clamped by the suppressor and thus prevents 

sparkover of the arrester, the suppressor absorbs all of the surge energy. 

In all instances, the voltage level at the suppressor is held low enough to 

protect all electronic appliances having a reasonable tolerance level (MX) V in 

most cases, IOOO V in some cases). Furthermore, the installation of only one 

suppressor in the house provides substantial protection for other outlets, 

although optimum protection requires the use of a suppressor at the most sensitive appliance, with additional suppressors for other sensitive appliances. 

V. SURGE INJECTED INTO GROUND SYSTEM 

Assumptions 

For this experimenl it was postulated that a lightning stroke attaching ta 

the primary side of an overhead distribution system would produce a branching of the current flow into the ground after sparkover of the pole-mounted 

utility's surge arrester (which was presumed connected at the pole-mounted 

distribution transformer). Fig. 8 shows the assumed circuit and the division 

of current flow. 

'k,- " :"'" 

Fig. 8. Division of current assumed for a 100 kA stroke. 

In their study of lightning environments, Cianos and Pierce [61 indicate 

that only 5% of all ground strokes exceed a peak current of 100 kA. The fre

quency of the strokes is dependent upon the geographic location (isokeraunic 

levels) (71, as well as upon local configurations. The probable mcurrence 

of a stroke involving the utility pole near a house with no adjacent tall trees o

buildings is I per 400 years for most of the U.S. For a 5% probability

the likelihood can be reduced 20 times; in areas of high lightning activity, thi

likelihood can be reduced 10 times. A stroke exceeding 100 kA at one loca

tion, therefore, can be expected to occur only once in 10,000 years (but ther

are millions of poles in the U.S.). 

From these assessments, the maximum current to be injected for the hous

model under discussion was selected to be 30 kA. From this maximum o

30 kA injected into the ground wire of the houseservice drop, two more value

were used during the test series: 10 kA, corresponding to the requirement fo

the ANSl high-current, shortduration test; and 1.5 kA, corresponding to th

requirement for the ANSl duty-cycle rest -both specified by ANSl Standar

C 62.1 for secondary valve arresters (31. All had wavethapes of 8 x 20 ps. 

Another reason for selecting this low level (1.5 kA) was that no sparkove

occurs in the wiring at this level. For the 10 and 30 kA levels, multiple flash

overs occur at variable times and locations, making exact duplication of test

impossible. By limiting current to below sparkover levels, repeatability of th

results was ensured, allowing comparisons among several alternate circui

configurations. 

The generation of transient voltages in the house is attributed to electro

magnetic coupling. The lightning current in the messenger establishes a fiel

that couples into the loop formed by the two phase wires encircling the mes

senger. In addition, there is some capacitive coupling between the wire

(Fig. 9). 

Fig. 9. voltages induced in the house wiring system. 

Test Circuit 

The test circuit consisted of a high-current impulse generator, a distribu

tion transformer with a service drop, a simulated simplified house wiring sys

tem, and the necessary shielded instrumentation. 

The service drop connection between the distribution transformer and the 

meter socket was made with three 13 in: (45 ft-) long AWC #6 wires, twisted 

at a pitch of about 5 turns/m (1.5 turns/ft). This service drop was folded in a 

loose '5" shape at about 0.5 m (1.5 ft) above the ground plane serving as th

return path for the lightning current, in order to reduce the loop inductance 

seen by the generator. This configuration does not influence the coupling be

tween the messenger and the wires wrapped around it, coupling which has been 

identified as the voltage-inducing mechanism. 
The simulated house wiring started at the meter socket and continued to a 

load center over a distance of 3 rn (10 ft). From this load center four "branch 

circuits" connected to the load center breakers were established, each terminating at a wall receptacle. Individual lengths of the branch circuits were 6, 

12.24. and 48 m (20.40. 80, and 160 ft). 

Test Results 

Many tests were performed to investigate the effects of various combinations. A selection was made from ceveral hundred recorded oscillograms to illustrate these effects. The results are presented in the form of oscillograms 

with corresponding commentary, generally providing a comparison of voltages 

and currents with or without protectors installed. 

The first striking result noted was that the injection of a unidirectional impulse into the ground system produces oscillatory voltages between the phase 

and ground wires. Inspection of the no-load oscillogram (Fig. IOa) reveals 

two interesting phenomena. First, the frequency of the major voltage oscillation is constant for all branch circuit lengths (period = 2 ps). Thus, we can 

conclude that this frequency is not affected by the line length and that other 

circuit parameters, rather, are responsible for inducing this 500 kHz oscillation 

from a 8 x 20 ps current wave. Second, the minor oscillations visible during 

the first loop in each oscillogram are spaced apart at a distance that increases 

with line length. One can conjecture that these may be caused by reflections. 

Loading the line termination with a 130 Q resistor (Fig. lob) eliminates the 

later oscillations and reduces the first peak to about 60% of the value without 

load. From this reduction. a Thevenin's calculation of circuit oarameters. if 

applicable in an oversimplified form, would show that 130 51 is 60% of the 

total loop impedance, while the source impedance* is 40% of the total loop 

impedance. Hence. one can conclude that the equivalent source impedance 1s 

in the order of four-sixthsof 130, or about 85 R, in this scenario. 

VOLTAGE WITH 1300 

COhNFCTFn ATOUTI FT 

INDICATED 

Sweep lpr dn 

h, VO tascat outkt mdrated SW V dn 

Fig. 10. Open-circuit voltages and effect of terminal ~mpedance. 

Injected current: 1.5 kA. 

With no protectors at the load center nor at any outletc, the wiring flashes 

over at I0 kA injected current, but not before crests in the range of 8 kV have 

been reached (Fig. 1 la). With an arrester installed at the load center, voltages 

are limited to 2.2 kV, with about I kA current discharge in the arrester 

(Fig. Ilb). While eliminating the hazard of a wiring flashover or the failure 

of a typical electromechanical device, this 2.2 kV protective level may still be 

excessive for sensitive electronics. 

VOLTAGE AND CURRENT WITH 

ARRESTER ON LOAD 

CENTER 

Sweep: 2 rddiv 

Current In arrescer 400 A/dw 

Voltage Across Bur 

5W V/d~v 

Fig. 12 shows the recordings made during a 30 kA current injection. This 

extreme condition is capable of producing a 3500 A current in an arrester installed at the service entrance (Fig. 12a). If now we postulate a pessimistic 

situation where there is no arrester at the service entrance, but only a suppressor at an outlet, there are two possible outcomes. When no wiring sparkover 

occurs, as discussed in Section Ill, all the surge is indeed forced upon the suppressor (Fig. 12b). This current may be excessive for some suppressors, but 

this example is certainly a limited case. The more likely scenario is illustrated 

in Fig. 12c, where sparkover of the wiring upstream of the suppressor limits 

the current in the suppressor. In this last scenario, protection is obtained 

downstream from the suppressor. It is important to note that no additional 

hazard is created by installing the suppressor: the undesirable sparkover 

would occur even without the suppressor; in fact, without the suppressor, 

sparkover would be even more likely to occur. 

Current In 

suppressor: I kAldlv 

Voltage at 

outlet: 5W V/dw 

Sweep: 2 @r/diu 

Current in 

suppressor: I kA/dlv 

Voltage a1 

outlet: SWV/div 

Sweep: Zps/div 

(c) Wiring Flashover 

Fig. 12. Duty imposed on stngle suppressor with 30 kA injection 

VI. CONCLUSIONS 

Coordination of surge protectors is feasible with existing devices, even if 

device characteristics vary. The experiments reported in the paper show three 

facts from which conclusions can be drawn: 

Fact I. Where a unidirectional current is injected into the ground system 

only, the responseof the system is an oscillatingvoltage, at 500 kHz 

for the system described. 

Fact 2. The equivalent source impedance, as determined by loading the 

system, is in the range of 50 to 100 R for the particular system 

investigated. 

Fact 3. Without substantial connected loads in the system, the open-circuit 

surges appearing at the service entrance propagate along the branch 

circuits with very little attenuation. 

Concl. 4. Coordination of surge suppressors requires a finite impedance to 

seoarate the two devices. enabline the lower voltaee device to oerConcl. 5. 

Concl. 6. 

fokm its voltaee-clamoine func~i~n while the hieher voltage debice . - performs the energydiverting function. 

The concept that surge voltages decrease from the service entrance 

to the outlets is misleadmg for a lightly loaded system. Rather, the 

protection scheme must be based on the propagation of unattenuated voltages. 

Indiscriminate application of surge protector? may, at best, fail to 

provide the intended protection and, at worst, cause disrupt~ve 

operation of the suppressors. What is needed ir a coordinated approach based on the recognition of the e%ential factors governing 

devices and surge propagation. 

Fig. 11. Protection provided by arrester at service entrance. VII. ACKNOWLEDGMENT 

Injected current: I0 kA. 

The contribution of K. E. Crouch in obtaining the current injeclion lest 

'Not to be confused with the surge impedance (L/C)'/' of the line. results i5 gratefully acknowledged. 
REFERENCES 

1. Martzloff. F.D., and G.J. Hahn. "Surge Voltage in Residential and 

IndustrialPowerCircuits," IEEE PAS-89,6, luly/Aug. 1970, 1049-1056. 

Lenz, J.E., "Basic Impulse Insulation Levels of Mercury Lamp Ballast 

for Outdoor Applications," Illuminating Engrg., February 1964, 

pp. 133-140. 

Guidefor Applicalion of Valve-Type Lighrning Arresters ForAlternaling 

Current Systems. ANSI Standard C62.2-1969. 

Insularion Coordination Within Low- Volrage Systems Including Clearances and Creepage Distances for Equipmenl, International Electrotechnical Commission Report SC28A (Central Office) 5, to be published 

in 1979. 

Guideline on Surge Voltages in AC Power Circuits Rated Up ta 600 V. 

IEEE Project P587.1, First draft, 1978. 

Cianos. N. and E.T. P~erce "A Ground-Lightning Environment for 

Engineering Usage," Sranford Research Institute, Menlo Park, CA 

94205, August 1912. 

Elecrrical Transmission and Distriburion Reference Book, Westinghouse 

Electric Corporation, 4th Edition, East Pittsburgh, Pa., 1950. 

Francois D. Martzloff (M, 1956) was born in 

France and received his undergraduate degree at 

the Ecole Sp6ciale DeM6caniqueet d'Electricit6 

in 1951; he received the MSEE degree from 

Georgia Tech in 1952 and the MSIA degree 

from Union College in 1971. 

Since 1956 he has been with the General 

Electric Company, where he first gained experience in the Transformer and Switchgear 

Divisions. Upon joining General Electric 

Corporate Research and Development in 1961, 

, he became involved in power semiconductor 

c~rcuits and overvoltage protection. He has 

oartici~ated in the introduction and aoalication . . bf metal oxide varistors since 1971. 

In IEEE Mr. Martzloff has been active on the Surge Protective Devices 

Committeeand chairman of the Working Groupon Surge Voltagesin AC Power 

Circuits Rated 6W V or Less. He is also a member of the Ad Hoc Advisory 

Subcommittee of the USA Advisory Committee on IEC S/C 28A. He has 

been awarded 10 U.S. patents, primarily in the field of varistors and transient 

protection. 
PROPAGATION

AND

ATTENUATION

OF SURGE

VOLTAGE AND

CURRENT IN

LOW-VOLTAGE

AC CIRCUITS
Schenectady NY 

f.martzloff@ieee.org 

Reprinted, with permission, from IEEE Transactions on Power Apparatus and Systems, PAS-102, May 198

First presented at IEEE Summer Meeting, San Francisco, July 1982 

Also reprinted as “High Interest Paper” in IEEE Power Engineering Review, January 1984 

Paper Award, 1985, Surge-Protective Devices Committee

Significance: 

Part 4 – Propagation and coupling of surges 

Examples are given showing the propagation of voltage and current surges in low-voltage wiring systems. The 

difference between surge impedance (characteristic impedance Z0) of a transmission line and the impedance t

the surge of wire runs is pointed out and illustrated. 

The relationship between front time/duration of a voltage surge on the one hand, and the travel time (lengthrelated) along the circuit, on the other hand, is placed in the perspective of transmission line theory and makes 

clear the point that the classical doubling of an impulse at the end of an open line requires a travel time greater 

than the front time of the impulse. 

A comparison is made between the propagation of a surge through isolating transformers and through a ferroresonant line conditioner. The isolation transformers do not provide effective attenuation of voltage surges in th

differential mode but the ferro-resonant line conditioner does, in addition to its prime function of voltage regulat

For current surges of the type encountered in AC power circuits (not short pulses), their propagation is impede

– as in “impedance” – not by the characteristic impedance of the line nor appreciably by skin effects, but most

by the inductance of the line for a frequency spectrum in the range of 5 kHz to a few hundred kHz. This provid

some relief for SPD connected at the end of branch circuits. The issue was revisited and confirmed several ye

later in the 1995 “Upside-Down House” experiments (see file “ Upsdown measure” in this Part 4) 

The effects of connection options are shown for one, two or three SPDs connected at the end of a 3-wire line. 

Filename: Propagation 1980 
IEEE Transactions on Power Apparatus and Systems, Vol. PAS-102, No. 5, May 1983 

THE PROPAGATION AND ATTENUATION 

OF SURGE VOLTAGES AND SURGE CURRENTS 

IN LOW-VOLTAGE AC CIRCUITS 

F.D. Martzloff, Senior Member 

Corporate Research and Development 

General Electric Company Schenectady, 

Schenectady, New York 

Abstract - Examples are given showing the propagation of voltage and current surges in low-voltage wiring systems. The 

difference between surge impedance (characteristic impedance) of 

a transmission line and the impedance to the surge of wire runs is 

pointed out and illustrated. A comparison is made between the 

propagation of a surge through an isolating transformer and a 

ferro-resonant line conditioner. The effect of connection options 

are shown for one or several surge protective devices connected at 

the end of a 3-wire line. 

INTRODUCTION 

Considerable progress has been made during the last decade 

toward recognizing the occurrence of surge voltages in low-voltage 

circuits, particularly in ac power circuits; at the same time, 

improved protective devices have become available. New standards and guides have been published on the subject [I-51, but 

practical information is still scarce on the propagation of these 

surges in circuits. In fact, misconceptions are sometimes encountered, such as an expectation that surges will always attenuate substantially as they propagate in the wiring system of a building or 

through transformers. 

This paper provides concrete examples on the propagation of 

surges and on some possible means to divert or attenuate them, 

from which some conclusions can be drawn and sound practices 

recommended. 

The tests reported here have been performed with the voltage 

and current waveshapes recommended by the recently published 

"IEEE Guide for Surge Voltages in Low-Voltage AC Power Circuits," IEEE Std 587-1980. 

Starting from the basic propagation of a pulse along a transmission line, the examples show how real wiring systems differ from 

an idealized transmission line, how complex systems transform 

pure standard waves, and how some connection practices for surge 

protective devices can introduce adverse effects by producing residual voltage surges between conductors. Merits and misconceptions regarding isolating transformers are compared with ferroresonant line conditioners which, in addition to their regulating 

function, block the propagation of fast transients, such as the 

IEEE 587 ring wave. 

This paper does not propose to present a comprehensive treatment of surge suppression, but rather to show how the propagation of surges affects the voltages appearing at the loads. The 

examples reported are given in order to alert the reader against 

some pitfalls resulting from occasional misconceptions observed by 

the author during discussions and reviews of many surge-related 

problems. 

82 SM 453-9 A paper recomended and approved by the 

IEEE Surge Protective Devices Comittee of the IEEE 

Power Engineering Society for presentation at the IEEE 

PES 1982 Summer Meeting, San Francisco, California, 

July 18-23, 1982. Manuscript submitted February 2, 1982; 

made available for printing May 17, 1982. 

TEST CIRCUITS AND TEST GENERATORS 

The test circuits selected to represent typical low-voltage ac c

cuits include runs of nonmetallic jacketed 600 V wire such as t

found in residential wiring and runs of rigid conduit with wi

pulled in the conduit, such as that found in industrial installatio

In these tests, all use two-conductor configurations with a th

grounding wire, AWG #12 size (2.05 mm dia.). 

The test waves, in accordance with IEEE Std 587-1980, inclu

the O.Sps - 100 kHz voltage ring wave, the 1.2/50 ps volta

impulse, and the 8/20 ps current impulse. The 0.5 ps - 100 k

wave is produced by a KeyTek 424 surge generator, capable 

superimposing the voltage pulse at a controllable time of the 60 

line voltage, with crest up to 6 kV (Figure 1). The 1.2/50 impu

is produced by a Haefely P6R generator, capable of supplying 

to 6 kV (Figure 2). The 8/20 impulse is produced by the sa

KeyTek 424 with a different plug-in unit (Figure 3) or by a labo

tory circuit using storage capacitors and an ignitron switch. T

initial test involving transmission line propagation is made with 

narrow pulse obtained from a Velonex 350 pulse generator. 

Figure 1. Test waves of 0.5 ps - 100 kHz, open circuit 

loaded, produced by KeyTek 424 generator 

PN281LSC plug-in 

IIIIIIII 

Figure 3. Test waves of 8/20 ps, short circuit produced 

KeyTek 424 pulse generator with PN247 plug-in 

IIIi;lllJ 

0018-95 10/83/0500-1163$01 .00 @ 1983 IEEE 

I 

Figure 2. Test wave of 1.2/50 WS, open circuit, produced 

Haefely P6R pulse generator 
The oscillograms were recorded with a Tektronix 7633 oscilloscope. Voltage measurements were made with two 1000:l 

Tektronix P6015 probes in differential configuration, so that the 

display calibration of the oscillograms is to be multiplied by 1000. 

Current measurements were made with a 0.010 T&M Research 

coaxial shunt, so that the display calibration is to be multiplied by 

100 A/V. 

The oscillograms are shown together with a schematic of the 

circuit configuration. The three wires of the various lines are 

shown as B (black?, W (white?, and G (green), with the usual 

convention on color. The impulse was applied between the point 

shown by the lightning bolt and the ground symbol. The points of 

connection of the differential probes and the resultant recording 

are connected by the arrows. 

PROPAGATION OF VOLTAGE SURGES 

Transmission line behavior 

To establish the baseline of the propagation characteristics, a 

75 m line of 2-wire plus ground nonmetallic plastic jacket cable is 

subjected to pulses of 100 ns duration, with the voltages measured 

at the sending end and at the receiving end of the line. Figure 4 

shows the oscillograms recorded at the sending end, with the outgoing pulse and the reflected pulse appearing 740 ns later, from 

which a propagation speed of 150 m/740 ns, or 200 m/ps, can be 

computed. As a side experiment, the classical nonreflection 

obtained by terminating the line with a resistance equal to the line 

surge impedance* is also observed for a terminating resistance of 

100 D. The slight mismatch indicated by the small reflection 

remaining, even with the optimum value of 100 D termination, is 

attributable to the connection of the line with closely spaced wires 

fanning out to the ends of the noninductive wire-wound resistor 

card. Resistor cards of 90 D and 110 D produce a positive-going 

or a negative-going reflection, respectively, indicating that the 

matching impedance is between these two values. 

-75rn-I I I I I I I I I I I NON-METALLIC SHEATHED CABLE 

XI2 AWG WITH GROUND. 

Fignre 4. Transmission line behavior of non-metallic, 

plastic jacket wire 

From this first test, we can draw the conclusion (predictable, 

but too often not recognized in qualitative discussions of 

reflections in wiring systems? that it is not appropriate to apply 

classical transmission line concepts to wiring systems if the front 

of the wave is not shorter than the travel time of the impulse. 

For a 1.2/50 ps impulse, this means that the line must be at least 

200 m long before one can think in terms of classical transmission 

line behavior. In the next example, we can observe reflections on 

the front of the impulses, but they are not significant to the final 

voltages at the crest values of the impulses. 

Short lines behavior 

The response of lines shorter than the 200 m limit identified 

above is illustrated in Figures 5 and 6. Figure 5 shows the 

response of a 25 m line of nonmetallic plastic jacket wire to a 

1.2/50ps impulse; Figure 6 shows the response of conduitenclosed wires to the same impulse, for the same line length. 

A difference between surge impedance (also known as "characteristic 

impedance") and impedance to the surge will be discussed in the section dealing with current surges propagation. 

Figure 6. Response to a 1.2150 FS impulse of wiring 

in metal conduit 

(A) Effect of grounding options at the sending end 

(B) Sending end and receiving end voltages 

~----25rnIn both Figures 5 and 6, the front of the wave (SD side

slower than 1.2 ps (OC); this effect is caused by the 

impedance loading the impulse generator. It is even more not

able in Figure 6(A), which shows the voltages at the sending 

for the connection with white, green, and conduit tied togeth

The greater capacitance of this configuration, compared to w

only at ground, produces a greater load on the generator, includ

the multiple reflections occurring at the mismatches produced b

conduit fitting at the mid-point of the conduit run; hence, 

jagged appearance of the oscillogram. 

i - 

1 

- -4.- 

IIIIIIII 

The front of the impulse is further expanded in Figure 5(

showing the difference between the sending and receiving end

the line. Nevertheless, for an open-end line, as shown 

Figure 5(B), the final voltage crest is not affected by 

reflections occurring during the rise time of the impulse. L

wise, the slight differences between the sending end and 

Figure 5. Response to a 1.2/50 ps impulse of wiring in no

metallic jacket 

(A) Output of unloaded generator and with 

line connected 

(B) Sending end and receiving end, 

open receiving end 

(C) Expanded trace of oscillogram (B) 
receiving end occurring early in the rise do not affect the final 

voltage at the receiving end of the conduit line. 

These tests show that the propagation of voltage surges in 

open-end lines of lengths likely to be encountered in buildings does 

not procwe appreciable attenuation of the surges, nor does it 

cause a \ ltage buildup by reflection of the surges. This fact is 

significant for the case of an appliance or industrial equipment with 

a control circuit drawing very little load while the power circuit 

being controlled is off. During the on cycle of the power circuit, 

there will be some attenuation of the surge by the combined effect 

of the line impedance and load impedance, but that beneficial 

effect is not available during the off-cycle of the power circuit with 

standby condition of the control circuit, the very circuit containing 

the most sensitive electronic components of the appliance or 

equipment. 

Branched circuits 

Departing further from the simple and sanitary behavior of a 

transmission line, a still simplified branch circuit behavior is illustrated in Figure 7. In this semi-idealized case of a real circuit, a 

10 m line feeds 4 branches, each 10 m long. Three of the 

branches are left open ended, and the fourth has a heavy load - a 

short circuit. The interaction of this circuit with the impulse generator, set for the 1.2/50ps impulse of Figure 2, produces the 

wave shown at the sending end, while the voltage at one of the 

open receiving ends goes through oscillations that only vaguely 

resemble the sending-end wave shape; of course, the idealized 

unidirectional impulse has vanished. 

tllllIIIII 

- iFigure 7. Response to a 1.2150 ps impulse 

of a branched circuit 

This simple branch circuit behavior demonstrates why it would 

be an illusion or fallacy to cling to the concept that nature can be 

simulated with simple test waves.* 

This dichotomy between simulating nature and performing standard tests has 

been recognized l6,7l, but still needs to be emphasized. A test wave is 

applied to a device, not to demonstrate that it can survive any of the waves 

that it will encounter in nature, but only to demonstrate for the benefit of 

both manufacturer and purchaser that the device can survive an agreed-upon, 

arbitrary, simple, clean impulse. From surviving the test impulse, the inference is made, subject to cor&mation by Jield experience, that the device does 

have the capability to survive the infinite variety of surges that it will 

encounter during its life in the real world. In other words, simple (and clean) 

test waves are useful because they can be reproduced over a period of time at 

the same facility, and between different facilities, providing a common 

language and a standard of comparison that is essential to conduct orderly 

transactions. Test waves should not, however, he misconstrued as representing natural phenomena. They are "realistic" (which is not the same thing as 

"representing reality") only to the extent that the conclusion drawn from 

surviving the test wave is validated by better survival in the field than for 

those devices that do not survive the test wave. 

Does an isolating transformer help? 

The author has witnessed and engaged in many discussions o

the merits of isolating power transformers, sparked by the misco

ception indicated by statements such as "spikes are attenuated b

transformers" or "spikes do not pass through transformers." Fi

ures 8 through 12 are offered to support the position that the

quotations are misconceptions. When properly applied, isolati

power transformers are useful to break ground loops, but they 

not by themselves attenuate surges that occur line-to-line or in t

normal mode. 

Figure 8 shows the propagation - or worse, the enhanceme

- of a voltage impulse in a 1:I isolating power transformer. T

6 kV impinging ring wave appears as 7 kV crest on the seconda

side of this "isolating" transformer. 

1 A I lus I 

1 kVA GENERAL PURPOSE 1201120 

TRANSFORMER 

Figure 8. Propagation of a 0.5 ps - 100 kHz ring wave 

through an isolating transformer 

Figure 9 shows similar behavior in a transformer offered as 

"line isolator." This product is intended to provide ground loo

isolation and low effective capacitance between primary and seco

dary windings, but here again, the author has observed that use

of this device expect attenuation of surges. The response of th

isolator, due to its internal construction, is different from that 

the simple two-winding transformer of Figure 8, but we also no

that a crest of 8 kV occurs on the secondary side, during th

second half-cycle. Hardly an improvement. 

1 kVALOWCAPAClTANCE 1 I 1 1 I "LINE ISOLATOR" TRANSFORMER 

Figure 9. Propagation of a 0.5 ps - 100 kHz ring wave 

through a "line isolator" transformer 

Figures 8 and 9 were recorded with no load on the transform

secondary, which represents the extreme case of a low-power ele

tronic control in the standby mode. Figure 10 shows the prima

and secondary voltages of the transformer with a 10 W (1500 a

and a 100 W (150 a) load on the secondary side, at the same ge

erator setting as Figure 8. With the 10 W load that might be typ

cal of an electronic control in standby mode, the combined seri

reactance of the transformer and shunt resistance of the load pr

duce the output shown in Figure 10(A), still slightly higher tha

the input. 

With the 100 W load shown in Figure 10(B), the attenuation 

now apparent, but is only 2:l. Capacitive loads would, of cours

produce a greater attenuation than resistive loads for the inductiv

series impedance of the transformer, at the frequency spectrum o

this fast 2 bs-wide surge. For surges of longer duration, th

attenuation would be smaller. 
Figure 10. Effect of loading on the secondary side 

These examples show that, unless a well-defined load is connected to the transformer, expecting attenuation from the 

transformer may prove to be hazardous to the health of low-power 

electronics connected on the secondary side of the transformer. 

In contrast, decoupling of the surge is possible with a ferroresonant line conditioner, which is primarily intended for line voltage regulation, but which also provides a high degree of surge 

suppression. Figure 11 shows a 6 kV impinging ring wave 

attenuated to 60 V (100:l) on the secondary side of the unloaded 

line conditioner, and to 40 V (150:l) with a load of only 10%; at 

full load, less than 10 V was observed. The nature of the ferroresonant line conditioner is such that the decoupling improves 

with loading, while the simple transformers of Figures 8, 9, and 10 

can only act as linear dividers with load changes. Conversely, the 

decoupling between primary and secondary sides of the line conditioner is further seen on the oscillogram recorded on the input 

side of the line conditioner. This os~illogram is, in fact, a 

photograph of two successive measurements, one with no load on 

the line conditioner and one with a 100 W load. The input waves 

are exactly superimposed. Compare this with the regulation of the 

generator output voltage noticeable in Figure 1, where a 100 

resistor is connected directly at the terminals of the generator. 

Figure 11. Decoupling of a 0.5 fis - 100 kHz ring wave 

by a ferro-resonant line conditioner 

This decoupling reflects the nonlinear behavior of the ferroresonant l~ne cond~t~oner, wh~ch 1s sgn~ficant In this case, compared to the hear behav~or of transformers: for surge sources of 

lower Impedance than the generator used in these tests, or for frequenaes lower those than contained in the 0 5 ps - 100 kHz ring 

wave, the transformer attenuatlon would become lower, In d~rect 

proportion to the correspond~ng impedance change, while the 

ferro-resonant transformer would keep the decoupling unchanged 

The two oscillograms of the output were recorded with th

surge timed to occur at the peak of the 60 Hz line voltage, f

worst-case demonstration. The peak-to-peak amplitude of the lin

voltage is indicated by the gray band recorded on the oscillogram

by photographically superimposing repetitive traces of the lin

voltage. For timings other than the peak, the small voltage oscill

tion on the output voltage would be completely contained with

the normal peak-to-peak band of the 60 Hz line voltage. 

While these measurements were being taken, an addition

observation was made. Figure 12 shows the response of the lin

conditioner to surges occurring at different times in the 60 H

cycle, as indicated by the different vertical position of the traces 

their beginning. The ferro-resonant mechanism is responsible f

this different response. In itself, this is not very important in th

present context, but it does provide another example of th

importance of performing surge testing at different angles alon

the power-frequency cycle (as recommended in the discussio

presented in IEEE Std 587) because the outcome of the test m

be influenced by the timing of the surge. 

Figure 12. Effect 

u 

of timing of the surge with respect to th

power frequency voltage 

PROPAGATION OF CURRENT SURGES 

Line impedance: surge impedance or impedance to the surge? 

As mentioned in the transmission line behavior, a distincti

has to be made between two concepts that unfortunately can 

confused because of the language: surge impedance of t

transmission line, and impedance of the line to the surge. T

first is the classical transmission line parameter, also cal

"characteristic impedance", 2, = m, and applies for lo

lines and short pulses. It is independent of the line length and f

quency. The second, impedance to the surge, is indeed depende

on the line length, and is the impedance of the complex (real a

imaginary) network of distributed parameters, R,L,C, of the w

ing configuration. This impedance is also dependent on the f

quency, so that rigorous analysis would involve computation ov

the frequency spectrum of the impulse of interest. For practi

applications, it would be more convenient, although not rigoro

to define the impedance of a line to the surge as the ratio of vo

age to current, stating the current wave form. 

Thus, inspection of Figure 13 shows a current crest of 400 

flowing in the line with shorted end and a voltage crest of 1700 

at the sending end, with a current wave form of 25/70 ks It 

noteworthy that the short-circuit impulse of 8/20 ps produced 

the generator has been stretched out by the effect of the l

impedance. This impedance is mostly inductive, as shown by t

fact that the crest of the voltage occurs during the initial curre

rise where di/dt is large, with a resistance detectable by a fin

voltage at the time dud1 is zero - that is, at the crest of t

current, not counting the capacitance. Thus, one might define 

impedance to the surge of this 75 m line as being 1700 V1400 

for a 25/70ps wave, or 4.25 'ohms', a far cry from the 100 

characteristic impedance determined by the first measureme

reported in this paper. This impedance is essentially proportio

to the line length, in contrast to the constant value of the char

teristic impedance. 
Figure 13. Current and voltage in 75 m line, 

unidirectional impulse 

The complex nature of the line configuration is also evident in 

the voltage observed between the shorting jumper at the receiving 

end of the line and the green ground wire: the voltage is not just 

half of the sending voltage but, rather, the superposition of that 

half-voltage and higher frequency components which are not seen 

in the sending-end voltage. 

Likewise, Figure 14 shows a first current crest of 48 A in the 

shorted 75 m line with a sending-end voltage of 5000 V when the 

0.5 ps - 100 kHz generator is driving the line. This corresponds 

to an impedance to the surge of 100 'ohms', not very different 

from the characteristic impedance, For the second crest, however, 

the current crest is 30 A with a voltage crest of 1300 V, or about 

45 'ohms' for the significant frequency of the second crest. 

Figure 14. Current and voltage in 75 m line, 

oscillatory wave 

The addition of the grounding wire (Figure 15) to the circuit of 

Figure 14 does not considerably change the crests of voltage and 

current but introduces the added complexity of a secondary oscillation superimposed on the driving oscillation. 

I-SD 

I /'- 

Figure 15. Effect of added grounding wire for 75 m line, 

oscillatory wave 

Figure 16, simllar to Figure 13, shows the propagation of 

current surge in conduit-enclosed lines. For the same generat

short-circult wave form of 8/20 ps, the resultant current and vo

tages are shown for a 25 m conduit run. The ratio 

voltagelcurrent yields a value of 1500 V/850 A for the 20150 p

current wave form, or an impedance of 1.8 'ohms'. 

Figure 16. Current and voltage in 25 m conduit run, 

unidirectional impulse 

Figure 17 shows the impedance of the 75 m line as a functi

of frequency, as measured by an impedance vector meter. T

values of impedance defined as approximations for impulses a

also shown on this graph. 

Therefore, as a first approximation, a more useful view of t

relative impedance values in a wiring system can be derived fro

this concept of "impedance to the surge" than from the use 

characteristic impedance, provided that the user of this appro

mate concept does not lose sight of the approximations implied 

the concept. 
Figure 17. Impedance versus frequency of 75 m line 

For instance, consider the case of a wiring system where a 

decreasing "staircase" of voltage surges is expected as the wiring 

progresses within the building, starting from the servlce entrance. 

Such a staircase is described in the IEC recommendations on Insulation Coordination [2]. The staircase is obtained by using the 

interface effect of the series and shunt impedai~ces of the wiring, 

including line impedance, transformer (if any) series impedance, 

and shunt impedance of connected loads. While the series 

impedances are likely to constant for a given system, the shunt 

impedances will vary with the loads. Alternatively, the decreasing 

voltages can be the result of installing surge protective devices at 

interfaces between sections of the wiring where the voltage 

decrease is to take place Such a coordinated wiring system will 

require careful consideration of the line impedances for the various ranges of surge-effective frequency to be expected, so that 

proper coordination can be ensured between the successive protective devices installed at the interfaces [a]. 

The pitfalls of unsanitary wave forms 

While an impulse generator is essentially an energy storage element (capacitor, line or inductance) discharged into the test specimen through some wave-shaping network, producing a clean wave 

shape as described in standards specifications is not a trivial undertaking. Unless precautions are observed, the stray inductance of 

capacitors or the stray capacitance of inductors as well as the wiring impedances can introduce unwanted oscillations - an "unsanitary" wave form. 

Figure 18 gives an example of the problems that an unsanitary 

wave form can introduce. An attempt was made to apply a laboratory surge generator (built for energy deposition testing) to force a 

8/20 ps current into the 75 m line, since the 8/20 ws short-circuit 

wave becomes stretched as discussed previously. The reasonably 

clean current wave form of Figure 18(A) would be quite acceptable as a test wave where total current, crest current, or energy are 

Figure 18. Effect of unsanitary current wave on line voltage 

the significant parameter. However, when applied for the purp

of evaluating line impedance by measuring and computing V/I, 

small ripple occurring on the current rise produces the L(di

oscillations seen in Figure 18(B). Thus, such an unsanitary w

form is totally useless for that purpose. 

The effect on residual surges 

of connections options for suppressors 

A noticeable lack of agreement has been observed among v

ous application information sources on the most effective transi

suppression configuration to be applied. Taking, as an exam

the task of specifying the protection of an appliance or equipm

connected at the end of a line with no opportunity to divert 

transient closer to the source (for instance, at the serv

entrance), the options would be to connect one, two, or th

surge suppressors between the three wires (black, white, 

green) at the end of the line. However, more needs to be kno

Will the impinging surge be in the normal mode (black to wh

or in the common mode ([black-and-whitel-to-green)? Where 

the equipment is the most sensitive component: line-to-line (m

likely) or line (black OR whitel-to-green? Clearly, the situatio

confusing, and there will not be a single, simple answer applic

indiscriminately to all cases. The National Electrical Code 

specifically allows the connection of surge arresters (Article 280-

if the interconnection occurs only by operation of the surge ar

ter during the surge. Since the standby current of a varistor or 

leakage current of an avalanche diode suppressor is very low, 

intent of this requirement can be met. Furthermore, there 

not be any interference with the operation of Ground Fault Cir

Interrupters if the total number of suppressors does not result 

large current. 

The set of measurements recorded in Figure 19 shows 

example of these many options with increasing protection, albe

increasing cost, from a single suppressor to three suppress

The selection would depend on the vulnerability level and loca

of the equipment to be protected. The impinging surge 

assumed to be black-to-[white and green], since white and gr

are tied together at the service entrance. The line is the 75 m 

previously investigated, and the surge is that available from 

generator set for a 2000 A 8/20 rs short-circuit impulse. Rat

than attempt to modify the setting of the generator for each c

in order to maintain constant current crest for the vari

configurations (an impossible task if wave form is also to be m

tained), the generator was left unchanged, to discharge a cons

total energy in the system - not a bad hypothesis for the 

world. This test was performed with 20 mm diameter varis

rated for 130 V rms line voltage, as an example. Similar res

would be obtained with other types of clamping suppressors. 

point is not so much the clamping voltage measured, as it is 

relative differences for the various options shown. The curr

crests are all in the range of 300 to 380 A, which is no

significant change for comparing clamping voltages. 

If only one suppressor is allocated to protect the equipm

the black-to-white suppressor connection affords maximum pro

tion for the electronics which are also likely to be connected bla

to-white. However, the voltages between either black or w

and green are large; this is the stress that will be applied to 

clearances of the equipment. This example shows a current sur

which might seem relevant only to surge suppressor applicatio

becomes a voltage surge issue, which is relevant to insulat

coordination of clearances. 

The configuration with suppressor black-to-green does 

afford very good protection for components connected black

white; therefore, it should be used only if there is a special n

to clamp black-to-green at a low voltage. 

An improved protection is obtained with a suppressor black

white complemented by a second suppressor white-to-gre

Another option, not investigated here but often used in appl

tions of three-electrode gas tubes, would be the connection 
CIRCUIT 

Figure 19. The effect on residual voltage surges of connections options for one, two, or three suppressors 

tv 1ous 

380 A 

suppressors, one between a line and the grounding conductor, the 

other between the second line and the grounding conductor. The 

ultimate protection is, of course, one suppressor in every position, 

but this should be required only for exceptionally sensitive loads. 

CONCLUSIONS 

The examples of surge propagation described provide the basis 

for several practical conclusions that should provide guidance in 

designing or evaluating surge protection schemes. 

1. Surge propagation in wiring systems should be considered as a 

case of classical transmission lines only if the lines are long 

enough to contain the surge front. 

2. For typical voltage or current surges produced by lightning or 

switching, the surge impedance (characteristic impedance) is 

not the significant parameter. Rigorous analysis requires considering the frequency spectrum of the impulse and the line 

impedance at the significant frequencies of that spectrum. 

Approximations can be made for specific current surge wave 

forms. 

3. Isolating power transformers are intended to serve as ground 

isolators, or ground-loop breaks. They do not provide appreciable attenuation of line-to-line transients unless they are 

operating with their series reactance combined with a welldefined shunt load on the secondary. 

4. Ferro-resonant line conditioners can provide attenuation of 

fast line-to-line transients with ratios of 100:l or higher. 

Adding a small fixed load on the output side can raise this 

attenuation to 150:1, or more. 

5. The connection options for surge suppressors must b

matched to the protection requirements for optimum prote

tion at minimum cost. Universally applicable solutions alwa

tend to be more expensive. 

6. Careful design is required for impulse generators. Improvis

tion can lead to meaningless results and wasted time. 

7. In testing for surge protection evaluation, the timing of th

surge with respect to the power line frequency can b

significant. 

8. The pure and sanitary test waves specified by test standard

are intended to obtain reproducible results rather than 

duplicate surges occurring in reality. Complex wiring syste

(within a building or within equipment) will prompt

transform the pure wave form into a distorted form, but th

does not prevent consistent results, since an agreement exis

on the initial test wave. 

mv ?mfTT lous rwlv nKIdl 'OH 

360 V max 380 V max 

ACKNOWLEDGMENTS 

rwlv mT iw 100 V max 

Discussion of surge protection techniques among members 

the IEEE Working Group on Surge Characterization (for instanc

the effect of connection options) has provided valuable insigh

into the problems and has motivated the examples presented 

this paper. The contributions made by reviewers of the paper pr

vided further insight on the subject and are gratefully ac

nowledged. 
IEEE Std 5871980, IEEE Guide for Surge Voltages in 

Low-Voltage AC Power Circuits," The Institute of Electrical 

and Electronics Engineers, Inc., New York. 

IEC Publication 664, "Insulation co-ordination within lowvoltage systems including clearances and creepage distances 

for equipment," International Electrotechnical Commission, 

Geneva, Switzerland, 1980. (Available in the United States 

from ANSI.) 

IEEE Std 465.1-1977, "IEEE Standard Test Specifications for 

Gas Tube Surge Protective Devices," The Institute of Electrical and Electronics Engineers, Inc., New York. 

IEEE Std C62.32-1981, "IEEE Standard Test Specifications 

for Low-Voltage Air Gap Surge-Protective Devices," The 

Institute of Electrical and Electronics Engineers, Inc., 

New York. 

IEEE Std C62.33-1982, "IEEE Standard Test Specifications 

for Varistor Surge-Protective Devices," The Institute of 

Electrical and Electronics Engineers, Inc., New York. 

F.A. Fisher and F.D. Martzloff, "Transient Control Levels," 

IEEE Transactions on Power Apparatus and Systems 95 (I), 

JanuaryIFebruary, 1976. 

F.D. Martzloff and FA. Fisher, "Transient Control Level 

Philosophy and Implementation," paper presented at the 2nd 

EMC Symposium on Electromagnetic Compatibility, Montreux, June 28-30, 1977. 

F.D. Martzloff "Coordination of Surge Protectors in LowVoltage AC Power Circuits," IEEE Transactions on Power 

Appararus and Systems 99 (I), JanuaryIFebruary, 1980. 

National Electrrcal Code 1981, National Fire Protection Association, Boston, Massachusetts. 

in France, and received his undergra

uate degree at the Ecole Spiciale D

Micanique et D'Electricitt: in 1951; 

received the MSEE degree from Geo

gia Tech in 1952 and the MSIA degr

from Union College in 1971. 

Since 1956 he has been with t

General Electric Company, where 

I gained experience in the Transform

and Switchgear ~ivisions. Upon joining General Electric Co

porate Research and Development in 1961, he became involv

in power semiconductor circuits and overvoltage protection. H

has participated in the introduction and application of me

oxide varistors since 1971. 

In IEEE, Mr. Martzloff is active on the Surge Protecti

Devices Committee. He is chairman of the Working Group 

Surge Characterization in Low-Voltage Circuits. He is also 

member of the Ad Hoc Advisory Subcommittee of the US

Advisory Committee on IEC SIC 28A and ANSI C.62 Subcom

mittee on Low-Voltage Surge Protective Devices. He has be

awarded 10 U.S. patents, primarily in the field of varistors a

transient protection. 

Discussion 

A. C. Liew (National University of Singapore, Kent Ridge, Singapore): 

The author has once again presented us with an interesting and practically useful paper. 

Regarding the application of classical transmission line concepts to 

wiring systems and the concept of surge impedance, it is well known 

that surge impedance or characteristic impedance is applicable directly 

only until the time of arrival of the first reflection. After that, a lumped 

equivalent circuit is usually used or great effort in keeping track of the 

reflected and transmitted waves must be taken. This is evident in figure 

14 of the paper. 

The author's comments of the following observation made by us are 

appreciated. 

We have found that for nearby lightning strokes, the induced 

voltages on the wiring system of a building (even when supplied by an 

underground cable at the service entrance) does not have to be very 

large to cause operation of sensitive Ground Fault Circuit Interrupters 

or Earth Leakage Circuit Breakers (current-operated type). Even with 

the installation of low voltage lightning arresters (500 V or 380 V type 

for a 415 V system) before the Ground Fault Circuit Interrupter, sufficient unbalance surge currents to ground can flow to cause its operation. This was traced to the distributed capacitances to ground of the 

wiring system. With the liberal connection of suppressers to the input 

terminals of sensitive equipment, the situation is likely to be further aggravated. Thus, while no damage occurs as with successful surge suppression, this nuisance tripping may be intolerable in certain cases. 

Manuscript received August 9, 1982. 

F. D. Martzloff: Indeed, as pointed out by Liew and myself, there are 

limitations to the application of classical transmission line concepts to 

wiring systems. However, while Liew states that these are well known, 

my experience in discussing the topic has shown me that in many instances, in the heat of a discussion, or under the pressure of a postmortem, some erroneous or misapplied concepts can surface. 

To avoid these situations, some repetition of known facts, prese

with concrete examples, may he helpful and provide useful guida

Thus, the purpose of the paper is not to report discoveries, but to m

better known the limitations and pitfalls cited in the paper, in the 

text of concerns on surge propagation and attenuation. 

E. K. Howell: While Liew reports operation of Ground Fault Ci

Interrupters coincident with nearby lightning strokes and attributes 

operation to surge currents resulting from distributed capacitance o

wiring system, the information provided is not sufficient to warrant 

specific conclusions, explanations, or recommendgions. 

Most Ground Fault Circuit Interrupters (GFCI) today use electr

signal processing and provide limiting and integration of the fault 

rent signal, which tends to prevent operation by the fast surge curre

However, a flash-over in the wiring system may initiate a sub-c

follow-through current, at power system frequency, having suffic

magnitude and duration to require operation of the interrupter. 

thermore, spark-gap types of low voltage ligntning arresters have

herently, a follow-through current which is limited by a varistor 

may be large enough to properly cause GFCI operation if that curre

allowed to pass through the ground fault current sensor. 

There is also the possibility that the electronic circuit was suscep

to the surge voltage, rather than the current, as the result of insula

breakdown or some parasitic high-frequency coupling within the G

device. The present Underwriters Laboratory GFCI Standard No. 

requires immunity (no tripping) to 3 kV crest of the 0.5 p-100 

voltage surge waveform. No current surge response requirement e

today. If tripping in response to a surge of voltage or current is s

ciently intolerable, then specifications defining acceptable performa

should be considered, for either general use or special-purpose dev

We both thank Professor Liew for the opportunity to clarify this 

ject. 

Manuscript received October 12, 1982. 
SURGE

AND

HIGHFREQUENCY

PROPAGATION

IN

INDUSTRIAL

POWER LINES
First presented at IEEE ICPS, May 1985

Significance: 

Part 4 – Propagation and coupling of surges 

The propagation and attenuation of surges or high-frequency disturbances in power lines has been described in diffe

terms by workers hailing from time-domain or frequency-domain schools of thought. Nature, of course, recognizes 

neither one in particular, and the phenomena are the same. This paper is an attempt at unification of the description

reporting measurements made on the same specimen by the two different techniques. 

Unidirectional pulses with duration ranging from 200 ns to 50 μs, and the 0.5-μs – 100-kHz ring wave were injected 

a metal-enclosed line as well as a non-metallic jacketed line. Data are presented in graphical form for the continuous

frequency measurements and as typical oscillograms for the pulse measurements. 

From the time-domain surge measurements, it becomes apparent that long lines will attenuate single-shot impulses f

very short duration (less than 1 μs), but no appreciable attenuation can be expected for longer pulses. An open-end 

will produce the classical doubling effect when the line length is sufficient to contain the surge front. If the line if shor

than this value, reflections occur while the surge front is still rising, so that the doubling effect produces steps on the 

but no doubling of the ultimate peak of this surge.
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. IA-22, NO. 4, JULYIAUGUST 19

Surge and High-Frequency Propagation in 

Industrial Power Lines 

Abstract-Laboratory measurements were made on a three-conductor 

line, with or without steel conduit shesth (typical of single-phase 120/240-V 

systems), to determine the attenuation of surges and the response to 

steady excitation at frequencies in the range of 100 Hz-10 MHz. Line 

length ranged from 75 to 225 m. Impedance matching was maintained at 

the sending end, while at the receiving end various loads (matched or 

mismatched) were included. A nonlinear load was also used, illustrating 

the side effects of connecting a surge protective device at the end of the 

line. 

INTRODUCTION 

T HE PROPAGATION and attenuation of surges or 

high-frequency disturbances in power lines has been 

described in different terms by workers hailing from timedomain or frequency-domain schools of thought. Nature, of 

course, recognizes neither one in particular, and the phenomena are the same. This paper is an attempt at unification of the 

description, reporting measurements made on the same specimen by the two different techniques. In a second phase of this 

work, numerical methods would be applied to correlate and 

convert the data from one domain to the other. 

A typical industrial power line was deployed in the 

laboratory, consisting of three conductors in a steel conduit. 

The line was folded into a zig-zag arrangement to allow both 

ends and two intermediate points to be within short reach of 

the instruments. The parameters of a three-conductor nonmetallic jacket line were also examined. Pulses of different shapes 

and continuous swept-frequency signals were applied at one 

end of the line. Measurements of voltages or currents were 

made to characterize the propagation and attenuation (or 

enhancement) of the signals. 

From the measurement results, data are presented in 

graphical form for the continuous-frequency measurements 

and as typical oscillograms for the pulse measurements. 

Implications are discussed for application to typical situations. 

The line consisted of 225 m of 20-mm (314-in) steel conduit, 

arranged in a zig-zag configuration to bring the start, the end, 

Paper lPCSD 85-34, approved by the Power Systems Engineering 

Committee of the IEEE Industry Applications Society for presentation at the 

1985 Industrial and Commercial Power Systems Technical Conference, May 

13-16. Manuscript released for publication March 7, 1986. 

F. D. Martzloff was with the General Electric Company, Schenectady, NY. 

He is now with the Electrosystems Division, National Bureau of Standards, 

Building B 162, Room 220, Gaithersburg, MD 20899. 

H. A. Gauper, Jr., was with the General Electric Company, Schenectady, 

NY. He is now at 826 Sanders Avenue, Scotia, NY 12302. 

IEEE Log Number 8608868. 

and two intermediate points at 75 m and 150 m close to eac

other, allowing connection of recording instruments withou

change of probe length. Three wires of 3.3 mm2 cross sectio

(12 AWG) were pulled through the conduit, representing th

line, neutral, and grounding conductors of a typical singl

phase installation. In all measurements, the neutral an

grounding conductors were bonded to the conduit at th

sending end, representing a typical service entrance configur

tion. At the intermediate points and at the far (receiving) en

the neutral was not connected to the grounding conductor. A

the receiving end, the grounding conductor was also bonded 

the conduit. Limited measurements were also performed on 

75-m-long line of nonmetallic three-wire conductor to invest

gate the differences between the "open line" configuratio

and the "quasi-coaxial" configuration of the conduit-enclose

line. 

Because the conduit-enclosed line is practically insensitiv

to its surroundings, it could be zigzagged in a steel-walle

room without undue effects. In contrast, the nonmetallic jacke

line might be sensitive to the proximity of walls and, ver

likely, to its own adjacent turns if not stretched out in "fre

space." To evaluate the importance of this concern, measur

ments were made on the three-conductor wire left in i

original coiled package (50-cm average diameter) before th

line was stretched out in a long hairpin shape, with only th

sending and receiving ends inside the building, the rest of th

line being held at least 4 m away from the building an

ground. In these measurements the grounding conductor an

neutral conductor were bonded at the sending end. 

Surges and signals were generally injected and measure

between the line conductor and the neutral conductor, acro

which the sensitive loads are expected to be connected. In th

experiment concerned with connection effects at the line en

measurements were also made between the neutral and th

grounding conductors at the end of the 75-m line, whe

various configurations of surge suppressors were installed. 

Surges were generated by a Velonex 350 or KeyTek 71

generator, monitored by a Tektronix 7633 oscilloscope wi

two P6015 probes in differential mode, allowing measur

ments at the four access points of the line. For the fa

repetitive pulses generated by the Velonex 350 generator, 

trigger signal allowed synchronized sweep of the oscilloscop

for all four measurements, so that propagation time would b

displayed on the recordings. For single pulses generated wi

the KeyTek 711 generator, internal triggering of the oscill

scope in single-sweep storage mode provided the recording 

the pulses. Continuous frequency measurements were pe

0093-9994/86/0700-0634$01 .OO @ 1986 IEEE 
MARTZLOFF AND GAUPER: SURGE AND HIGH-FREQUENCY PROPAGATION 

TEK 

7633 

VELONEX \ 

STEEL CONDUM 

Fig. 1. Recordings at sending and intermediate points and receiving end of line, terminated with matching impedance, for 200-11swide pulse. 

formed with a Hewlett-Packard system consisting of a Model 

8553A spectrum analyzer and a 8443A tracking generator. 

Four types of surges were injected into the line, with 

different waveforms: 

a 200-ns-wide unidirectional pulse, 

a 2-ps-wide unidirectional pulse, 

a 1.2150-ps surge, as described by [3], 

a 0.5-ps 100-kHz ring wave, as described by [3]. 

A. 200-ns Pulse 

Fig. 1 shows schematically the test line and a composite 

oscillogram of the voltages at the sending end, intermediate 

points, and receiving end. Note the impedance-matching 

network between the generator and the line, providing an 

optimum wavefonn for the generator output (200 0) and 

minimizing reflection of the surges returning from the line 

(100 Q). A 100-0 matching termination was connected at the 

receiving end for this recording. 

The attenuation of this pulse along the line is quite apparent, 

with an average 0.7 ratio between the voltages of points 

separated by 75 m of line. The propagation time for the 225 m 

can be seen as 1.1 ps. 

Fig. 2 shows the propagation of the same 200-ns-wide pulse 

with open-ended line. The voltage enhancement at the receiving end (3.75 divisions compared to the 1.9 divisions in Fig. 1) 

illustrates the classical doubling effect at the open end. In spite 

of this doubling effect, enough attenuation of the short spike 

occurs over the total 225 m of the line that the pulse at the 

receiving end is lower than that at the sending end. For a line 

length of 150 m, the doubling effect would result in a 

receiving-end crest of 2 x 2.5 = 5 divisions, i.e., equal to the 

sending-end pulse. For a line length of 75 m, the doubling 

effect would result in a receiving-end crest of 2 x 3.1 = 6.2 

divisions, i.e., a 20-percent enhancement of the sending-end 

pulse. Note also that, while the amplitudes are attenuated, the 

rise times tend to be increased, so that the volt-time integral of 

the pulse (and therefore its potential for damaging energysensitive components) is not attenuated as quickly as the 

amplitude. 

B. 2-p Pulse 

Fig. 3 shows the propagation characteristics of a 2-ps-wide 

pulse along the same line, with matched termination at the 

receiving end. For this slower pulse the attenuation is lower, 

with a ratio of 0.94 between the voltages of points separated by 

75 m of line. The change of slope, which was quite apparent 

for the 200-11s pulse, is hardly noticeable in this recording, the 

four pulse traces having essentially parallel rises and falls. 

This lesser distortion is easily explained by the lesser effect of 

the line shunt capacitance on the lower frequencies associated 

with the 2-ps-wide pulse. 

C. 1.2/50-ps Surge 

Fig. 4 shows the test circuit for surge tests with matched 

terminations. The R, series resistance at the sending end 

provides matching for the low-impedance (0.6 or 12 0) 

generator plug-in networks, and the R, terminating resistance 

at the receiving end eliminates reflections. 

Fig. 5 shows three recordings for the 75-m conduit line 

excited by a 1.2150-ps surge, with open end and with a 

matching termination at the receiving end. Fig. 5(a) shows the 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. IA-22, NO. 4, JULYIAUGUST 198

Fig. 2. Recordings at four points of line, open end, for 200-ns-wide pulse. Fig. 3. Recordings at four points of line, matched termination, for 2-p

wide pulse. 

Z pi Rs Rt 

1.2/50,bnduit 0.6 100 100 

1.2/50,Non-metal I ic 0.6 125 125 

Ring Wave,condui t 12 50 100 

Fig. 4. Test circuit for recordings with 1.2/50-ps surge and with ring wave 

Fig. 5. Recordings on 75-m conduit line with 1.2150-ps surges. All oscillograms: vertical = 1 kvldiv, sweep = 500 nsldiv. (a) 

Openzircuit voltage produced by generator. @) Voltage at sending end. Top: open at receiving end. Bottom: 100 Cl at receiving 

end. (c) Voltage at receiving end. Top: open at receiving end. Bottom: 100 fl at receiving end. 

open-circuit voltage delivered by the surge generator. Because 

a series resistor is present at the sending end, a 2: 1 voltage 

divider effect is produced at the sending end for the long-term 

voltages when a terminating resistor is connected at the 

receiving end. Therefore, the terminated-end trace (bottom) 

shows half of the voltage of the open-end trace (top). For the 

first 750 ns of the sweep, the reflection from the receiving end 

has not arrived, and, consequently, the two traces are identical. 

Fig. 5(c) shows the voltages at the receiving end, with and 

without termination. Observe that the long-term voltages 

match the sending-end voltages of Fig. 5(b). Only the rising 

part of the top trace of Fig. 5(c), for the first 1500 ns, shows a 

doubling of the voltage applied at the sending end, but this 

doubling would not change the stress at the receiving end, 

except for the steeper rate of rise. 

The behavior of the nonmetallic (stretched) 75-m line shown 

in Fig. 6 is similar to that of the conduit line, but the traces are 

not quite as smooth. From the differences in high-frequency 

response of the conduit and nonmetallic line shown by the 

continuous-frequency measurements discussed later, these 

irregularities in the nonmetallic line can be expected. 

D. 0.5-ps 100-kHz Ring Wave 

Fig. 7 shows the recordings for the 75-m conduit line 

excited by a ring wave, with open end and with matched 

termination. The voltages at the sending end (Fig. 7(a) and (c)) 

(a) (b) 

Fig. 6. Recordings on 75-m nonmetallic stretched line with 1.2150-p

surges. Vertical: 1 kV1div. Sweep: 500 nsldiv. (a) Voltages at sending end

Top: open at receiving end. Bottom: 125 Q at receiving end. (b) Voltages a

receiving end. Top: open at receiving end. Bottom: 125 flat receiving end

differ by the reflection during the first loop, producing th

double peak in Fig. 5(a). Furthermore, the voltage divide

effect of the series resistance also lowers the sending-en

voltage when a terminating resistor is connected at th

receiving end. At the receiving end the difference between a

open or terminated condition is apparent. 

1) With a matched termination, the receiving-end voltage o

the first loop is 2.5 kV for a 3-kV sending-end voltage

indicating attenuation of the high frequencies as previously 

noted in the case of the 200-11s pulse. Later loops show les

attenuation. 

2) With an open end the reflection effect is less tha
MARTZLOFF AND GAUPER: SURGE AND HIGH-FREQUENCY PROPAGATION 

( c) (d) 

Fig. 7. Recordings on 75-m conduit line with 0.5-ps 100-kHz ring wave. All 

oscillograms: vertical = 1 kvldiv, sweep = 2 psidiv. (a) Voltage at 

sending end, line open at receiving end. (b) Voltage at receiving end, line 

open at receiving end. (c) Voltage at sending end, 100-Q termination at 

receiving end. (d) Voltage at receiving end, 100-0 termination at receiving 

end. 

doubling-5 kV against 3 kV-as the result of the highfrequency attenuation and again little attenuation in later 

loops. 

Two types of frequency domain measurements were performed over the frequency range of 100 Hz-10 MHz with the 

spectrum analyzer and tracking generator. 

1) The input impedance versus frequency of the line was 

determined by measuring the voltage applied to the line 

and then the current drawn by the line. 

2) The transfer ratio of the line versus frequency was 

determined by measuring the voltage applied to the line 

and the voltage at the far end of the line. 

To assure adequate decoupling of the tracking signal 

generator output, a resistive pad was connected to the output. 

This pad consisted of a 47.5-Q resistor in series with a 2.54 

resistor to a coaxial shield ground. The test line was thus 

driven by the 2.54 source. All data were recorded with an 

X-Y recorder, and then calculations were made to establish 

the reported data. 

The transfer ratio measurements required ungrounded 

instrumentation in order to eliminate ground loops and the 

influence of undesired grounding at the remote end of the test 

line. Therefore, a relatively high resistance of a 7500-i'l 

noninductive resistor was connected at the remote end of the 

test line and a Genistron current transformer was used to 

measure the current flowing in the resistor, thus establishing 

the voltage at the quasi-open-ended line. The voltage at the 

sending end was obtained through a Hewlett-Packard 1121A 

ac voltage probe. 

Fig. 8 illustrates the input impedance of the conduit line 

under this type of excitation. The two curves labeled 275 and 

MHz 

Fig. 8. Impedance of 225- and 75-m tines and VJV, ratio of 75-m line as 

function of frequency. 

2225 show the impedance of the line seen from the sending 

end when a 100-Q matching resistance is connected at the 

receiving end, as a function of the frequency of the applied 

signal. 

Referring to the 2225 curve, the behavior of this 225-m line 

shows the standing wave effects at 114, 112, 314, and 1 

wavelength. At 0.2 and 0.6 MHz the line impedance is low, 

typical of 114 and 314 wavelengths, while it is high at 0.4 and 

0.8 MHz, typical of half and full wavelengths. For the 0.8- 

MHz frequency, the period is 1.25 ps; with a line length equal 

to one full wavelength, the speed of propagation determined 

by this measurement is 225 ml1.25 ps = 180 dps, which is 

close to the result obtained by the pulse propagation time of 

1.1 ps shown in Fig. 1, i.e., 225 m1l.l ps = 204 mlps. The 

difference between these two results is well within the limits of 

calibration and reading of the different instruments used for 

the measurements. 

Referring to the 275 curve of Fig. 8, the first low point 

occurs at three times the frequency of the first low on the 2225 

curve, which is consistent with the 1:3 ratio of line length. 

Note also that this shorter line involves a lower total 

resistance; hence the IN-wave resonance effect is larger than 

the 314-wave resonance. This resonance is quite apparent on 

the VJV, 75 curve, which shows the ratio of the voltages at 

the receiving end to the sending end, with a quasi-open (7500 

0) receiving end. 

Fig. 9 is a plot of the input impedance versus frequency of 

three configurations of open-ended 75-m lines: the conduitencased line, the nonmetallic coiled line, and the nonmetallic 

stretched line. The differences of input impedance for the 

three configurations should come as no surprise. The conduit 

line appears well controlled, with more damping that the 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. IA-22, NO. 4, JULYIAUGUST 

IMPEDANCE 1 1 I I I I 1 I I I I 

dB ohms 

60 1000 

40 100 

20 10 

- 

- 

NON-MET STRETCH NON-MET COILED 0 1 I 

0 2 4 6 8 10 

FREQUENCY, MHz 

Fig. 9. Input impedance for open-ended 75-m lines. 

0 2 4 6 8 10 

FREQUENCY, MHz 

Fig. 10. Ratio of receivinglsending voltages for open-end nonmetallic 75-m line, coiled and stretched. 

nonmetallic line of amplitude between the 114-wave minima 

and 1/2-wave maxima of the lines. The simple response of the 

conduit line is also apparent, compared to the parasitic effects 

that appeared in the coiled configuration. 

Fig. 10 shows a comparison between coiled line and 

stretched line for the ratio between sending-end and receiving-end voltages. The voltage multiplication seen at the far end 

at about 0.6 MHz agrees with the Vo/K 75 curve of Fig. 8 for 

the 114-wavelength condition. Both stretched and coiled lines 

show a multiplication of about 24 dB (15 times) with little 

difference between the responses. The second peak begins to 

show some shift between the two lines, and the differences 

increase at higher frequencies, showing the need to stretch out 

the line for obtaining results free from the parasitic effects of 

the coiled configuration. The ratio drops to a minimum of 5-6 

dB rather than zero (no voltage enhancement) because the 

measurement system used had a noise background which 

precluded recording of any lower current measurements at the 

receiving end. 

Fig. 11, reproduced from [I], shows an earlier measurement of line impedance in the low-frequency region made with 

a Hewlett-Packard Model 4800A vector impedance meter, 

with the line shorted at the far (receiving) end, which is the 

converse situation of the measurements reported here, where 

the voltages at the receiving end are recorded with the line in 

quasi-open condition. 

RELATIONS BETWEEN IMPULSE AND CONTINUOUS WAVE 

MEASUREMENTS 

The time-domain impulse measurements illustrate directly 

the propagation of a surge of the type encountered in actual 

Fig. 11. Input impedance of 75-m line, nonmetallic type, with sho

receiving end. 

power systems. The effects of an open end, or a h

impedance at the far end, are quite apparent. 

The frequency-domain measurements indicate the imp

tance of standing wave effects for specific frequencies, wh

may be frequencies of low-level communication carriers or 

frequency of a particular point of the transform spectrum o

single impulse. To predict performance of a particu

configuration, it may then be possible to model the line, wh

would then become a simple case of conventional netwo

analyses by numerical methods [2]. On the other hand, sett

up a real line or network of lines may be cumbersome 
MARTZLQFF AND GAUPER: SURGE AND HIGH-FREQUENCY PROPAGATION 

yields immediate results, such as those illustrated in the next 

section on the effect of connecting surge suppressors at the end 

of the line. 

Comparisons of impulse measurements sometimes give 

considerable emphasis to crest values, neglecting the effect on 

waveform, as noted in connection with Fig. 2. A continuousfrequency spectrum, however, will readily identify the 

change. 

With the advent of increasingly sophisticated electronic 

devices, a proliferation of add-on surge suppressors has been 

offered for insertion at the point of use of these devices, that 

is, at the receptacle located at the end of a line, such as those 

discussed in the preceding sections. The actual configuration 

of these suppressors is generally not disclosed, but they may 

be classified into three categories, for the case of a singlephase two- or three-wire power supply. A first category 

involves only one protective device, which may be connected 

line-to-neutral or line-to-ground. A second category would 

involve two devices, and the third, three devices. The 

differences among these three categories is shown in [I]; 

however, ring-wave surges [3] may present a more insidious 

situation than the unidirectional surges used in [I]. The 

insidious nature of the effect is associated with the inductive 

characteristics of the line discussed here, when carrying fastrising surge currents associated with the presence of a surge 

suppressor at the end of the line. 

Fig. 12 shows four possible combinations of protective 

devices that may be connected at the user's end of a branch 

circuit when the user has no control on what protection could 

be added at the service entrance. Fig. 12(a) shows a single 

protective device connected between the line and neutral 

conductors, under the perception that this is the location of the 

most sensitive components of the connected load (L). Fig. 

12(b) shows again a single device, connected between the line 

and grounding conductors, under the perception that the surge 

current should be returned to "ground." Fig. 12(c) shows two 

devices, one between line and neutral conductors, the second 

between neutral and grounding conductors. Fig. 12(d) shows 

three protective devices. The effects of various comections 

when a slow-front current is involved were described in [I]. 

With faster surges, while lower energies are involved, 

ironically, a greater voltage spike is produced, as revealed by 

Fig. 13. 

Fig. 13 shows the result of connecting a surge suppressor 

device that provided clamping only between the line and 

neutral conductors, as shown in Fig. 12(a). Fig. 13(a) shows 

the voltage set at a peak of 3 kV at the sending end of the line. 

With that setting of the generator, the current in the line is 

shown in Fig. 13(b). The generator is capable of delivering 6 

kV open-circuit and 500-A short-circuit. The load-dependence 

of this type of circuit is evident in Fig. 13(b) where the current 

peak is limited to 40 A. Fig. 13(c) shows the voltage between 

line and neutral which, predictably, is correctly clamped at 

about 400 V. However, the voltage between neutral and 

grounding conductors (Fig. 13(d)) reaches a peak of 2300 V 

during the initial current rise and 1300 V during the fall of the 

: NeutrelConductor BondedTopaLer.ItheSendinp End 

G : Grounding Conductor 1 (Service Entrance) 

(L) : Load to be Protected 

Fig. 12. Connection options for protective devices at end of branch circui

Fig. 13. Effect of connecting single surge suppressor between line and 

neutral at end of 75-m line, with neutral and grounding conductor bonded a

origin of line. (a) Voltage at sending end while line is carrying current o

Fig. 12(b). Vertical: 1 kVidiv. Sweep: 2 psldiv. (b) Current in line with 

ring-wave surge generator excitation. Vertical: 20 Aldiv. Sweep: 2 psldiv

(c) Voltage between line and neutral with surge current of Fig. 12(b)

Vertical: 500 Vidiv. Sweep: 2 psidiv. (d) Voltage between neutral and 

grounding conductor with surge current of Fig. 12(b). Vertical: 500 Vldiv

Sweep: 2 psldiv. 

current at the end of the first loop. The correspondence 

between peaks of voltage during fast current change and zero 

voltage during peaks of current is quite apparent in a 

comparison of Fig. 13(b) and (d). This correspondence i

attributable to the inductive voltage drop in the neutra

conductor, which elevates the neutral terminal of the line-end 

receptacle with respect to the grounding terminal. The 

grounding terminal remains at the potential of the ground a

the sending end. 

A load connected at the receptacle and "protected" by the 

single device will then be exposed to substantial peaks o

voltage between its neutral and chassis, if any. Because of the 

expectation that neutral and grounding conductors should no
640 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. IA-22, NO 4, JULYIAUGUST 1986 

Fig. 14. Effect of connecting suppressor having clamping devlces between 

both line-neutral and neutral-ground pairs, at end of same line and same 

applied surge as in Fig. 13. (a) Voltage between l~ne and neutral with surge 

current of Fig. 13. (b) Voltage between neutral and grounding conductor 

with surge current of Fig. 13. 

drift far apart in voltage, being bonded at the service entrance, 

the designer of the load equipment might not have provided for 

this 2300-V peak in the insulation between neutral and chassis. 

If the function of the equipment can be disturbed by fast 

voltage changes, there is also a good probability that disturbances may occur. Thus what was intended to be a simple 

protective scheme can result in unwanted and unexpected 

probiems. 

An effect similar to that shown in Fig. 13 would be obtained 

for the connection of Fig. 12(b), but with the 400-V clamping 

voltage occurring between line and grounding conductors and 

the uncontrolled voltage of 2300 V produced between line and 

neutral conductors, placing any components connected across 

the line in severe jeopardy. 

In contrast, the performance of a surge suppressor that 

provides clamping between line and neutral and between 

neutral and grounding conductors is shown in Fig. 14, for the 

same surge-injection conditions as those of Fig. 13(a) and (b). 

Both voltages are clamped below 400 V, and while the initial 

rates of voltage changes are not affected by the clamping 

action, the total excursion is limited. Clearly, this mode of 

protection with added clamping between neutral and grounding conductors offers a substantial improvement over the case 

of Fig. 12(a). ' Further protection of equipment sensitive to 

fast voltage changes may be obtained by insertion of a filter 

that will decrease the rate of voltage changes shown in Fig. 14. 

The data obtained by these measurements should be a source 

of useful information for the application of preventive measures for equipment connected at the end of lines, such as 

those reported here. 

1) From the time-domain surge measurements, it becomes 

' This improvement is obtained from the point of view of equipment 

protection, where a surge of 1-ps duration as seen in Fig. 13(d) can cause 

failure. The National Electrical Code, In Sectlon 280-22, specifically permits 

the connection of a surge arrester between any two conductors of a three-wire 

slngle-phase system. However, with the surge current now returning to the 

service entrance by the parallel combination of the neutral conductor and the 

grounding conductor, a brief surge will appear between the equipment chassis 

connected to the grounding conductor and other local grounded structures, 

unless a bond is installed at that point to bridge the chassis and the other local 

grounded structures. The occurrence of this brief aurge would be lmited to the 

duration of high rates of current change, but its implication on personnel safety 

has not yet, to our knowledge, been evaluated In this context. 

apparent that long lines will attenuate single-shot impulses of 

very short duration (less than I ps), but no appreciable 

attenuation can be expected for longer pulses. 

2) An open-end line (e.g., a line feeding a low-power 

control circuit) will produce the classical doubling effect when 

the line length is sufficient to contain the surge front. If the line 

if shorter than this value, reflections occur while the surge 

front is still rising, so that the doubling effect produces steps 

on the front but no doubling of the ultimate peak of this surge. 

3) While conduit-enclosed lines may be folded for convenience in making indoor testing, nonmetallic lines need to be 

stretched away from surrounding conductors to obtain valid 

results at high frequencies. 

4) Connecting surge protective devices at the end of a line is 

a typical action undertaken by prudent users; however, unless 

the behavior of the circuit is well understood, simple protective schemes can backfire unless adequate connections are 

provided. 

111 F. D. Martzloff, "The propagation and attenuation of surge voltage and 

surge currents in low-voltage ac power circuits," IEEE Trans. Power 

App. Sysl., vol. PAS-102, pp. 1163-1170, May 1983. 

121 H. W. Dommel , "D~gital computer solution of electromagnetic 

transients in single and multiphase networks," IEEE Trons. Power 

App. Syst., vol. PAS-88, pp. 388-399, Apr. 1969. 

131 IEEE Guide for Surge Voftages in Low-Voltage AC Power 

Circuits, ANSIIIEEE Standard C62.41, 1980. 

Fran~ois D. Martzloff (M'56-SM'80-F'83) was 

born in Strasbourg, France, on June 9, 1929. He 

received the degree from Ecole Speciale de Mecanique et d'Electricitk, the M.S.E.E. degree from the 

Georgia Institute of Technology, Atlanta, and the 

M.S.1 A. degree from Union College, Schenectady, 

NY, in 1951, 1952, and 1971, respectively. 

He was with the General Electric Company for 

many years, working on transient measurements, 

surge protection of electronics, and applications of 

varistors. He is presently with the National Bureau 

of Standards, Gaithersburg, MD, focusing on conducted electromagnetic 

interference. 

Mr. Manzloff has received numerous Awards and he is active in writing 

standards with the IEEE, IEC, and ANSI. 

Harold A. Gauper, Jr. (M'49-SM'55) received 

the B.S. degree in electrical engineering from the 

University of Wisconsin, Madison, in 1943. 

He joined the General Electric Company shortly 

after graduation. Since his assignment to the field of 

Electromagnetic Compatibility in 1950, he has 

continued to specialize in the development of 

measurement methods, prediction and control of 

EMI, NEMP, and immunity associated with a broad 

range of consumer, industrial, electrical power, and 

mllitary products and programs. He holds three 

patents on EMC-related control techniques. Among his assignments for many 

years was the responsibility to represent the General Electric Company 

interests in EMIIEMC technical standards and regulations. As a result he was, 

and continues to be, involved with Federal Communications Commission 

rulemaking with regard to Pan I8 (ISM Equipment) and Part 15 (Radio 

Frequency Devices). He was Chairman of the NEMA Electromagnetic 

Compatibility Committee and actlve in the EMC standards work of ANSI 

C63, IEEE EMC Society, and other IEEE standards committees, as well as 

the IEC-CISPR and TC-77. He continues to be involved in lEEE EMC 

standards matters since h~s recent retirement from the General Electric 

Company. 

Mr. Gauper has served on the IEEE EMC Society Board of Directors, most 

recently in 1978-1981, and served as a Charter Member of the Federal 

Electromagnetic Radiation Management Advisory Council which advised on 

RF biohazard matters. 
VARISTOR

VS.

ENVIRONMENT

:

WINNING

THE

REMATCH
National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

© 1986 IEEE

Reprinted, with permission, from IEEE Transactions on Power Systems, Vol.PWRD No.3, April 1986

Significance:

This paper is listed under four categories of the Annotated Bibliography as it bears on the corresponding topics. 

The multiple listing reflects the sections in which this paper is cited as supporting material for IEEE Std C62.41.

and C62.41.2. Therefore, it can be found in the following four parts of the Anthology:

Part 2 Development of standard – Reality checks

Provides an example of the need to recognize capacitor switching transients when characterizing the surge

environment

Part 3 Recorded occurrences, surveys and staged tests

Provides an example of monitoring and staged tests motivated by field failure, leading to a better understanding 

the environment in which SPDs were expected to perform.

Part 4 Propagation and coupling of surges

Provides an example of how far (3000 meters) the low-frequency transients generated by capacitor switching ca

propagate, unabated, in a path involving two step-down transformers.

Part 7 Mitigation techniques

Provides an example of improved mitigation design based on field experience
Fraqois D. Martzloff, Fellow IEEE 

Corporate Research and Development 

General Electric Company 

Schenectady, New York 12345 

Abstract - An unusual case of difficult application of surge 

prntective devices was solved by field measurements with retrofit 

of protective devices suitable for the particular environment. Onsite measurements indicated that capacitor switching transients 

were causing excessive current surges in the varistors and fuses 

protecting the input to a thyristor motor drive. Knowledge of the 

environment gained by the measurements allowed understanding 

of the probieiti aiid ~p&biioii of maiching siirge ijioi&i~a 

devices. 

SUMMARY 

During the initial startup of a solid-state motor drive in a 

chemical processing pian:, dL!m!:ies a:me with !he vxis!~: and 

its protective fuse at the input of the thyristor circuits. Frequent 

blowing of the fuse was observed, with occasional failure of the 

varistor. On-site measurements of the voltages and currents at the 

input to the drive indicated that switching transients associated 

with the operation of a remote substation capacitor bank and the 

relatively low clamping level of the varistor were producing current 

above the fuse and varistor ratings; hence the short lives of these 

two components. When the actual conditions at that site were 

determined by measurements, it became possible to specify surge 

protective devices capable of withstanding that environment. Immediate relief was secured by the installation of a larger varistor at 

the same point of the circuit; long-term protection was obtained by 

the addition of a gapless metal-oxide varistor arrester on the primary side of the step-down transformer feeding the drive. The 

situation has been changed from failures occurring every few days 

to no further problems in the 3 years since the larger varistor was 

insiaiied. 

INTRODUCTION 

This paper presents a case history illustrating how surge protective devices that are successfully applied for the majority of cases 

an ~ccasima!!y sn&: fzi!ure when ~YPOECC! to exceptiom!!~ 

severe surge environments. This paper also shows how little 

attenuation occurs, at the frequencies produced by switching 

surges, between the distribution level (23 kV) and the utilization 

level (460 V), even though a long line and two step-down 

transformers exist between the source of the transient and the 

point of measurement. 

85 SM 365-2 A paper recommended and approved 

by the IEEE Surge Protective Devices Committee of 

the IEEE Power Engineering Society for presentation at the IEEE~PES 1985 Summer Meeting, Vancouver, 

B.C., Canada, July 14 - 19, 1985. Manuscript submitted February 1, 1985; made available for printing April 22, 1985. 

The problem involved a 460 V power supply t

drive circuit in a chemical processing plant extending 

square miles. During the initial startup, difficulties a

varistor and its protective fuse at the input of the thy

Frequent blowing of the fuse was observed, with occa

of the varistor. The plant substation, fed at 23 kV f

utility, included a large capacitor bank with one-third 

jwitchec! on aiid GK ff: provide pwe: fact=: an6 sy

regulation. These frequent switching operations were 

generating high-energy transients that might be the 

failure of the fuses and varistors, because literally 

similar drive systems have been installed in oth

without this difficulty. On-site measurements per

repeated blowing of fuses and occasional failure of v

nected at the input to the thyristor drive indicated th

devices were not matched to their environment. Fro

on, specifying larger sizes, sizes appropriate to the 

[ll, solved the problem. 

POWER SYSTEM AND SWITCHING TRANS

Figure 1 is a simplified one-line diagram of th

elements of the power system causing the varistor 

incoming 115 kV power is stepped down to 23 kV. 

of 5400 kVAR capacitors are connected to the 23 kV 

operating conditions involve two banks connected 

with the third bank switched on or off automatical

voltage regulation. Power distribution throughout th

at the 23 kV level. 

The various drive systems which experienced the 

supplied at 460 V by a 2300/460 V transformer in 

house. A substation close to the control house 

2300 V power from the 23 kV distribution system. 

i SiTE 

SUBSTATION DISTRIBUTION 

SYSTEM HOUSE i cONTROL 

I 

I 

VARIST

0885-89/77/86/0004-0059$01.00@1986 IEEE 
gppp

input. In the original circuit. the thyristor modules were protected 

by varistors at the power input of the 1250 hp drive, where the 

measurements were made. A 6 line inductance, L1, was 

inserted between the bus and the thyristor modules; 20 mm 

varistors rated 510 V were connected in a delta configuration, in 

series with a current-limiting fuse in each line. The varistor confieciioii via3 about 80 ciii loag, in:iodi;cing an e3tima:ed ! jiH 

inductance into each lead. 

TO CRO 

11- 

THYRISTOR 

MODULE 

[ IDENTICAL CIRCUIT TO THYRISTOR MODULE 

Figure 2. Simplified three-phase schematic 

Instrumentation and Measurements 

Oscilloscopes were used to measure voltage across one varistor 

and its connection and currents through all varistors. Voltage 

measurements were made phase-to-phase on the floating delta 

460 V bus bars with Tektronix P6015 1000:l probes, connected to 

a Tektronix 7633 storage oscilloscope in differentia mode. Current measurements were made with a Tektronix CT5 20:l current 

transformer coupled with a P6021 current probe and connected to 

a second Tektronix 7633 storage oscilloscope. 

The trigger modes used during a two-hour monitoring period 

inciuaed posiiive or negaiive siopes for both siow ac and highfrequency modes. For the various modes, the level was adjusted 

to produce a trigger for a voltage exceeding the normal line voltage crest by about 20°/0, or a varistor current in excess of 2 A. No 

trigger occurred during the monitoring period. A low-frequency 

voltage recorder installed by plant personnel produced a recording 

characterized as representative of an unusually quiet day in the 

power system operations. 

Manual off-on switching of the 5400 kVAR capacitor bank at 

the 23 kV utility substation was the next step in the measurement 

procedures because the switching of a capacitor bank is always a 

prime suspect for producing transients. Measurements were performed with one oscilloscope monitoring the line voltage upstream 

of the line inductors (Figure 2) and another oscilloscope monitoring the sum of the currents in the three varistors (Figure 3). 

Figure 3. CT connection for recording 

all three varistor currents 

-. 1ne osciiiogiams of Figure 4 show typicai voitage 

made during this sequence. The voltages are not open-c

sient voltages. They are instances of the voltage appea

bus entrance point. This voltage is the sum of the varis

ing voltage, the voltage drop in the varistor connection

voltage across two L1 inductances. 

A typical total event recorded on one of the phase

capacitor bank closing is shown in Figure 4A. A low

oscillation with a period of 3 ms (330 Hz) and initial pe

amplitude of 450 V decayed in about 10 ms. The high

oscillations are resolved in the recording of Figure 4R 

during a similar switching sequence). This high frequen

initial peak-to-peak amplitude of 2000 V, decaying in ab

The period is 180 ks (5.5 kHz). A similar, third event i

Figure 4C. For scaling the amplitudes, the steady-state 

shown in Figure 4D. 

A Sweep: 2 ms/div Vertlcal: 500 Vfdlv 

B Sweep: 0.5 msld

Verticsl: 500 Vl

C Sweep: 1 msldlv 

Vertlcal: 500 Vldiv 

D Sweep: 2 maldlv 

Vcrticd: 500 V

Figure 4. Capacitor switching transients 

and steady-state voltage 

Observe that, depending on the time of clos~ng with 

the 60 Hz voltage, the 5.5 kHz oscillation varies in 

f,lrtherrr?ore, !he modc!atic)n by the 330 Hz osc!!!ntlc

crests of the 5.5 kHz oscillation above the 1000 V level 

after the beginning of the trace, at a time when the 5.5 k

tude is already lower, producing a burst of pulses 

1000 V level. 

The significance of this finding will be discussed n

reference to Figure 5, which shows recordings of tran

rents in all of the three varistors. The 510 V varistor h

nal voltage at 1 mA [21 in the range of 735 V to 970 

varistor with a nominal voltage in the middle of this ran

rent in the order of tens to hundreds of amperes will 

..-B.~-- -r jnnn TT :- ---$:-A vwrrat;E; wi lwww r 13 app~lcu io ih2 iiiiii~i~i. Figure 5A 

train of current pulses in the range of 10 to 40 A. In th

Figure 5B, the recorded current pulses range from 5 A 

The current and voltage traces are not simultaneo

because each of the two oscilloscopes was triggered by 

circuit. The nearly symmetrical appearance of this bu

compared to the symmetry of the voltage peaks exce

1000 V level in Figure 4, the one correlating with the oth

The oscillograms of Figures 4 and 5 were selecte

severe from a series of 20 capacitor switching sequenc

sequences could not even produce a current or volta

tour sequences produced bursts w~th the central peak 

120 A, two of these reaching 200 A peaks. 
Table 1 

FIVE HIGHEST TRANSIENTS 

IN SEQUENCE OF 10 SWITCHINGS 

Without Varistors I With Varistors 

A Sweep: 0.2 ms/di* 

Vertical: 20 A/dlv 

I3 Sweep: 8.5 msldlv 

Vertlcsl: 40 A/div 

Figure 5. Current surge bursts during capacitor switching 

These recordings establish the nature of the current surges that 

are conducted by the varistors, with an estimate of 10% reaching 

200 A maxlmum crests and another 10% reach~ng 120 A crests, 

,--- ,, -----:*..- L..-,. :Ap'-:-- 

IUI 211 &apa&nul uaun SWILLIIIII& 

In F~gures 4 and 5, we note that the characteristic appearance 

of the voltage and current usually observed during a switch 

restrike is absent [3], indicating a clean switching action of the 

vacuum interrupters used for switching the capacitor bank. Restrikes are most likely to occur during de-energizing. In all the offon switching sequences of this test series, no significant transient 

was observed during de-energizing; all occurred during energizing. 

The oscillograms of Figures 4 and 5 establish and explain the 

pattern of current pulses. The voltages of Figure 4 are not the 

open-circuit voltages impinging the drive input but, rather, the 

voltages resulting from the clamping action of the varistors. To 

better evaluate the magnitude of the switching transients, opencircuit voltages were recorded in a next sequence, with all fuses to 

the drive open, thus disconnecting both the varistors and all sensitive loads. Figure 6 shows two typical recordings of open-circuit 

voltages and two of voltages resulting from varistor clamping, 

recorded during a series of 10 switching sequences for each condition. Tabie i snows tne recorded crests of tne five nighest voitages in each coix!i:ion; :he di!Te:ence be!ween the :wo gioiips, 

with due allowance for the imperfect statistical basis of the observations, indicates that the 510 V varistors reduced the peaks from 

a typical high of 1450 V to a typical high of 1100 V. 

Open-Circuit Voltsges 

With 20 mm Varistors 

All Traces: Sweep: 0.5 nts/div 

Vertical: 500 Vldiv 

Figure 6. Capacitor switching transients 

DISCUSSION 

Nature of the Transients 

The absence of any transient (over 120% of 

during ihe 2-hour moniroring period was somewhat 

the cn,r?!ex? n,f enr!ier rep~:?s ef high msns :e-,-,ideb 

disturbance analyzers. Frequent checks of thresho

variations of the possible trigger modes were made

the chance of catching an overvoltage, but indeed n

This unusual quiet was also reflected in the chart re

by the plant personnel, so that the absence of rand

for that period can be accepted at face value. 

Therefore, conclusive evidence was obtained th

current pulses were absorbed by the varistors du

switching. The magnitude and duration of these 

excessive for the capability of a 20 mm disc; many 

installed elsewhere do not experience the failures e

that particular location. 

Another significant finding from these measure

fact that the switching transients, generated at the 

propagate down to the point of utilization at the 

Numerical discussion of this finding is given later in 

Effect of Transients on Varistors 

Published varistor pecifimtions indude the "pnh

family of curves that define, for each varistor type, t

isolated pulses that a varistor can absorb until it

reached i4i. The curves show lines relating amplit

and total number of pulses. Figure 7 shows this fam

for the original 20 mm varistor. 

Figure 8 shows the same curves for a proposed 

tor. It should be noted that the pulse rating does n

strophic failure of the varistor at the end of this rati

10% change in the varistor nominal voltage. A

change is indicated, the varistor is quite capable of 

voltage and of clamping surges. 

Figure 7. Pulse ratings of 20 mm varisto
e [ I$:Z 

, 20.030 5 lo: 

8 - 

1MO 

! 

Z 200 

1m 3 50 

i : 

5 

2 

20 50 100 200 500 1OOO 2000 SaaO 10.000 

IMPULSE DURATION -pa 

Figure 8. Pulse ratings of 32 mm varistor 151 

A careful examination of the pulse rating curves will show that 

the duration of the pulses has a strong influence on the number of 

permissible pulses. Furthermore, the relationship between the 

increased duration of the pulses and the decreased number of permissible pulses is not proportional. For instance, consider pulses 

of 106 i\ peak eiid 100 CL3 diiia~oii (Figure 9A): the ci;ives shew 

5000 pulses allowed. Now increase the duration of the pulses to 

1000 ps (a ten-fold increase), while keeping the amplitude at 

100 A: the curves show the permissible number as one pulse 

only. Thus, the ten-fold increase in duration does not result in a 

ten-fold decrease in the number of permissible pulses; the reduction in that number is much greater than the inverse of the 

increase in duration. Conversely, taking a pulse duration of 

1000 ps, and seeking the amplitude allowable for the same 5000 

pulses, Figure 9B shows that the current is 20 A, which is five 

times less than the original 100 A, not ten times less. Therefore, 

it wou!r! he incnrrect to treat the multiple pulses of Figure 5 as 

five separate short pulses; rather, one equivalent long pulse has to 

be defined. 

The five-pulse burst of Figure 5 has been redrawn in Figure 10 

in order to plot an equivalent continuous pulse of approximately 

equal duration, with a crest such that the i.t integral of the burst 

and the i.t of the equivalent pulse are approximately the same. 

The use of i.t rather than the i2.t integral typically used for fuses or 

other linear loads is justified by the fact that heat deposited in the 

varistors is the significant parameter because the nominal voltage 

change process is temperature related; this heat is the product of IL- -.--:-Ll- --A 4h --..- .. -nnotnnt ., Iton- nrrnae the nnntin~nr rrle valravlr; i auu llle I,ral!j curlarsur rOarvbv ar.vuu I.v......rr. 

varistor during the burst. 

The equivalent pulse of Figure 10 can then be used to evaluate, from the pulse ratings of Figure 7, the number of highamplitude switching transients that will consume 100% of the 

varister pu!se ra!ing. !nspec!ior? of Figure 6 shows that for a 

800 ps duration and 100 A amplitude, the pulse rating of the 

20 mm varistor (6 kA rating at 8/20 ps) is reached with two such 

events. With a probability of about 10% that this highest switching 

transient would occur during random timing of the switching (the 

effect decreases rapidly for transients other than the highest) and 

with 2 to 4 swirching operaiions each day, iht piiise iaiiiig of the 

varistors could be reached with 20 operations, failure perhaps 

starting at 40 to 50 operations, or after about 10 days of exposure 

to that power system environment. This estimate is unavoidably 

imprecise because the pulse rating curves represent a conservative 

minimum; actual failures will occur only for amplitudes or numbers of pulses exceeding the rating by a large but imprecise margin 

to allow for manufacturing variations. However, the order of 

magnitude of this estimated time to failure is in accord with the 

observations made at that installation. 

I i - i 

20 1W loo0 

IMPULSE DURATION-us 

A. Same current, increasing duration 

I 1. 1 " - 20 1W 1000 

IMPULSE DURATION-pa 

B. Same number of pulses, increasing duration 

Figure 9. Reading pulse ratings curves 

ACTUAL 

Figure 10. Single-pulse equivalent pulse for multipl

Note that two types of events occur. One is the p

blowing of the fuse, which is not caused by a varistor i

by the i2.t capacity of the fuse being exceeded by the env

ill. The other is the fuse blowing caused by the varisto

life ultimate failure. 
Alternate Varistor Selection 

An obvious remedy would be to use a varistor with greater 

energy-handling capability. The 32 mm size offers such a possibility. Inspection of Figures 7 and 8 shows that the equivalent 

pulse of Figure 10 (800 ps and 100 A) corresponds to a permissible occurrence of 100 pulses for the 32 mm varistor, in contrast to 

the two for the 20 mm varistor. The improvement in the number 

of pulses is 50 times more pulses until pulse rating is reached. The 

improvement in the number of pulses until varistor failure occurs, 

however, is not necessarily 50 times more pulses. Because of the 

imprecision mentioned previously in the margin between end of 

pulse rating and ultimate failure, that margin is not necessarily the 

same for the two sizes, 20 mm and 32 mm, but it is reasonable to 

expect the same order of magnitude improvement in the ultimate 

failure as in the pulse rating. This expectation of a 50 times 

improvement would change the time he!weet? fai!ures frez the 

few days observed with the 20 mm size to perhaps one year with 

the 32 mm size, providing immediate relief and time to make fur- -. ther changes for the iong term. 'l'herefore, the change to a 32 mm 

size, cannec:eb a: :ha saiiie poiii: of the ciiciilt, was immadiateiy 

implemented for that particular environment. 

Further gains could be obtained in the length of time between 

varistor failures by increasing the clamping voltage of the varistors. 

This increase would result in lower current pulses for the same 

open-circuit transient voltage. A 510 V rating had been selected 

by the designer of the drive as the result of a trade-off between 

varistor clamping voltage and the withstand voltage of the thyristors protected by the varistors. If thyristors with higher voltage 

withstand were used, the solution would be easy. 

Of course, the standard varistor product line has a certain tolerance band, reflecting normal production lot variations. In principle, a selection could be requested from the manufacturer that 

varistors with a narrower band be supplied for this application. 

The maximum clamping voltage allowed by the drive specifications 

would be retained, but those varistors in the lower half of the distribution, which draw larger current pulses for a given open-circuit 

transient voltage, would have been removed from the population 

of varistors. For instance, tne range of nominal voltages for a 

575 V, 32 mm ua:iatc: (the next higher vakage ofkred is 805 10 

1005 V for 1 mA dc, while the maximum nominal va!tage of the 

same diameter but rated 510 V is 910 V for 1 mA dc. Thus, for a 

normal distribution of nominal voltages of the 575 V varistor, 50% 

of the devices could theoretically be used without exceeding the 

upper limit of the 510 V varistor that is consistent with the drive 

specifications. To achieve this end, it would be necessary for the 

supplier or user to make a careful determination of the nominal 

voltage on a population of 575 V varistors in order to retain only 

the lower half of the distribution (Figure 11). 

ACCEPTABLE LEVEL FOR THYRISTORS I 

SPECIFIED BY DRIVE DESIGNER 

PRESENT USE 

RANGE OF NOMINAL 

910 

RANGE OF NOMINAL 

VOLTAGES FOR 575 V 805 I 1005 

I 

Figure 11. Tolerances bands of 510 V 

and 575 V varistors 

Other Remedies 

In addition to the proposed upgrading of pr

460 V level, three other remedies could be cons

tion of surge arresters at the 2300 V level, insta

arresters at the 23 kV level, or a change in the circ

the capacitor switching, designed to reduce the 

transients at their origin. 

In general, the protection available from surge 

to improve when the arresters are installed at hig

ages. Thus, it is quite possible that arresters 

2300 V primary of the 23001460 transformer could 

effective clamping (and at the same time reliev

energy stress) than the varistors at the 460 V 

course implied that these would be the zinc-oxi

arresters.) The full benefit of these arresters d

cofiguration of the 2300 V system and its gro

grounded iieiiirai iii a wye sysiem, resisiance-gro

Boating &!?a! when the rrresters "re mnnected 

tional line-to-ground mode. In a second phase 

described here, 2300 V arresters were installed at 

primary. A discussion of their expected performan

the success of the retrofit, is given later on. 

Likewise, arresters on the 23 kV side could be 

23 kV substation to mitigate the capacitor switchi

their origin, or at the primary of the 23 kVl2300 V 

the control house, where they would also ser

protection for the overhead 23 kV incoming pow

arresters, again, must be of the gapless type to 

effective protection. 

The final remedy in the list of alternatives, but 

in effectiveness when the opportunity exists, woul

reducing the severity of the capacitor switching tr

origin. Series inductors or damping resistors may 

the effectiveness of which would be predictable if 

the power system behavior were performed by c

ing. 'wh:l- 

.. +hn. --...- A.. ^....lA ^^' %- ---,:-> .- .c:- llllu LII-L tGueuJ WUIU IIUL G appcu LO LIII

tion, it is a remedy that should be considered ffor 

exceptionally severe environment. 

EXPECTED PE~F~~~~VOF TEE 2380 -"

The measurements made first with open-circui

20 mm, 510 V varistors on the 460 V side have sh

of maximum voltage from 1450 V to 1100 V (T

current of approximately 200 A is flowing in the li

(Figure 5). 

We can assume that the voltage drop in the lin

tation and two step-down transformers is mos

5.5 kHz, and that the voltage in the varistors can b

voltage across a resistor at the time of the crest 

wave. The diagram of Figure 12 shows the relat

the three voltages VoC, VL, and VV, respectively, 

voltage generated by the capacitor switching act

drop in the line and two transformers, and the va

the current peak. Treating this highly nonlinear c

circuit is an appraxlmatinn !ha! ail! pmvide P! ea

full range of voltage and current conditions a valid 

tude for the purposes of this discussion. Numeri

avaiibie for rigorous treatment at any instant ov

of conbi:;oiis :6!. W-'L *L - ----:"--- -nirrl LLI~ sirrqmying assumpiio

mine the order of magnitude of the 5.5 kHz cur

flow in an arrester installed at the primary termin

V/460 V transformer as follows. 

CONCLUSIONS 

Voltage and current measurements made on the 460 V input to 

a thyristor motor drive, during staged capacitor switching operations, showed current surges in the varistors originally used in the 

system that could consume the pulse rating life of these varistors 

in a few days of typical operation. Short- and long-term remedies 

were achieved. 

For the short term, the change to a larger varistor connected 

on the 460 V side of the system was readily implemented to maintain the originally specified protective level, while the fuse-blowing 

nuisances were eliminated by use of a larger fuse. Available 

devices for this 460 V circuit may still have a relatively short life 

(a few hundred days) in the prevailing environment of the site, 

but they offered immediate relief and therefore allowed successful 

startup of the system. 

For the long term, further protection was obtained by the 

installation of conventional station-class surge arresters, of the 

zinc-oxide, gapiess type, at the 2300 V ievei. T'ne system has now 

eperated fer 3 years \vi!heut p:eh!ems. 

This case history also illustrates the low attenuation of the 

switching transient between the distant source at 23 kV (about 

3000 m, or 2 miles) and the point of utilization at 460 V. 

ACKNOWLEDGMENTS 

J.S. Kresge and B.I. Won provided information and guidance 

on the surge arrester and the varistor characteristics; C.L. Fisher 

contributed advice in clarifying and unifying the presentation of 

the concepts. Their contributions are gratefully acknowledged. 

REFERENCES 

[l] F.D. Martzloff, "Matching Surge Protective Devices to their 

Environment," Proc. IEEE/IAS Meeting. October 1983, pp. 

387-392. (Also scheduled for IAS Transactions, JanIFeb 

1985.) 

(21 ANSI/IEEE Std C62.33-1982, IEEE Stan

Spec8cations for Varistor Surge Protective Devices

tute of Electrical and Electronic Engineers, Inc., 

[31 A. Greenwood, Electrical Transients in Power Sys

Interscience, New York, 1971. 

141 Transient Voltage Suppression Manual, Fourth Ed

eral Electric Company, Auburn, NY, 1983. 

151 Transient Voltage Suppression Manual, Third Ed

eral Electric Company, Auburn, NY, 1983. 

I61 H.W. Dommel, "Digital Computer Solution of E

netic Transients in Single and Multiphase Netwo

Transactions on Power Apparatus and Systems, V

pp. 388-399, April 1969. 

joining General Electric Corporate Research and Developm

he became involved in power semiconductor circuits and ove

tection. He has participated in the introduction and applica

oxide varistors since 1971. 

In IEEE, Mr. Martzloff is active on the Surge Protec

Committee. He is chairman of the Working Group on Surg

zation in Low-Voltage Circuits. He is also a member of 

Advisory Subcommittee of the USA Advisory Committee 

28A, ANSI C.62 Subcommittee on Low-Voltage Surg

Devices, and Chairman of the NEMA Low-Voltage Arreste

Committee. He has been awarded 13 U.S. patents, primarily 

of varistors and transient protection. 
Discussion 

J. L. Koepfinger (Duquesne Light Company, Pittsburgh, PA): The author 

has addressed one of many mechanisms for producing repetitive overvoltages on low-voltage circuits. In this partic& instance, ii was possible 

to obtain controlled conditions so that a measurement could be made 

of the voltage and currents resulting from the capacitor switching. It 

would be useful if there was an analytical method presented that cor- related the generation of the 5.5-kHz pulses with those measured. Did 

the author attempt to make such a correlation? 

This paper points out the need to know the characteristic of the surge so that proper sizing of the protection can be achieved. Therefore it would 

be desirable to be able to have some analytical tool to permit calculation of the frequency of the surge due to remote capacitor switching. 

Manuscript received July 24, 1985. 

Frnncois D. Martzloff: The paper reported a case history from which 

usefui informaiion may be derived on reirofiiiing correciions of simiiar 

pmb!ems or, he??er, on .voiding ?he pmh!em hy foresight. The d?lntions confronting the author was the need for immediate corrective action rather than complete investigation and mutual validation of analytical 

methods ad fieid measureiiieiits. 

The literature is fairly rich in both theoretical and practical papers on 

the problems associated with capacitor switching, both for energizing 

and for de-energizing, the latter involving the risk of restrikes. Because 

of this availability and the limited space available in the TRANSACTXONS on one hand, and because of the limitations in scope of the field retrofit 

mission on the other hand, no attempt was made to correlate the 

measurements with the power system parameters (which were not readily available to the author). In response to Mr. Koepfinger's suggestion, 

abstracts are cited below to provide references to both analytical tools 

and practical results published by other workers. 

Bibliography, 1970-1985 : REFERENCES 

[I] M. F. McGranaghan, W. E. Reid, S. W. Law, and D. W. Gresham, 

"Overvoltage Protection of Shunt-Capacitor Banks Using MOV Arresters," IEEE Trans. Power App. Syst., PAS-103, No. 8, Aug. 1984, 

pp. 2326-2333. 

Protection reauirements and surge arrester duties are analyzed for 

boih digiiai and transient neiwork mdyzer (Tu'Aj shuiations. 3hn!c an~n_a!ytira! ex~ressions develoned for evaluating arrester duty 

is a function ofcapacitor bank size: Guidelines and Emitations for 

applying arresters at grounded- and ungrounded-wye capacitor banks 

ae deiieloued based oii overvokzge c~zrzc~er;spics &-ies;er 

capabilities. - 

121 - J. H. Brunke and G. G. Schoeckelt "Synchronous Energization of 

Shunt Capacitors at 230 kV," IEEE P& (Power Eng. sic.) Winter 

Meeting, New York, NY; Jan. 20-Feb. 3, 1978; Publ. by IEEE 1977, 

Paper A78-148-9, p. 4. 

This paper reports on the application of synchronous switching 

to reduce inrush transients when switching a 230-kV shunt capacitor 

bank. Computer studies determined the required switch performance. 

[3] J. D. Cuffman, John Linders, M. A. Zucker, and S. Willima , "Power Factor Correction Capacitors and Their Side Effects," IEEE 

Conf. Rec., 28th Ann. Conf. Electr. Ens. Probl. Rubber Plast. Ind., 

Akron, OH, April 5-6, 1976, pp. 37-49. 

The major reason for applying capacitors to an electrical distribution system is to correct poor power factor. In the greater majority 

of installtions it is a routine procedure. In some instances capacitors 

may cause problems with other in-plant equipment and in other instances they may suffer undesirable side effects that originate in other 

equipment. Among these problems are switching surges, voltage un- baiance due to fuses biowiiig, and iimiioiiics geiiergted by SCRcnn?m!!cd en,r?ipmen?= 

[4] H. J. Yelland, and C. P. Yellna, "Vacuum Contactors: Latest 

Development in Their Design and Application," Certif. Ens. V 54, 

No. 1. Jan. 1981, pp. 824-841. 

The paper includes the following topics: Design of vacuum bottles (glass and ceramic); design of the complete vacuum contactor 

and their panels; panels with on-load isolation, and panels with offload isolation vacuum contactor applications (moto

capacitor switching, arc and other (furnace-switching, tr

switching, mine-winder-reversers). Surge generation by va

tactors is considered under the following heads: basic e

and de-energization transients; transients generated when 

inductive loads; assessment of a vacuum contactor fro

generation point of view; types of surge suppression de

ching of capacitive loads. An extensive discussion of th

appended. 

Jack R. Harbaugh and John E. Harder, "Important Con

for Capacitor Applications in the Petroleum and Chemi

Industries," IEEE Pet. Chem. Ind. Conf., 27th Ann. Re

Pap., Houston, TX, Sept. 8-10, 1980; IEEE, #80CH1

Piscataway. NJ, pp. 157-167. 

The location of capacitors may have a significant ef

(I**2R losses within the plant transformers and conduct

is wasted energy. Transients generated by capacitor swi

require attention in the selection of arresters, system ins

ofier equipmi. Ti ofcap&iors artly require s

attentinn to large motors during system r=!&ng or 

This study addresses each of these considerations, prov

guidelines for effective, reliable capacitor application. In 

a c:l&iisi is provided for indiisirid capaciior ap

J. F. Burser, R. J. Santoro, J. W. Stolle. R. E. Owen 

Clinkenbeard, "Comparative Evaluation of Field Test 

Computer Results on Capacitor Switching Transients.

Minutes PA Electr. Assoc., Eng. Sect., Trans. and Distr

West Middlesec, PA, May 15-16,1979; PAElectr. Assoc., 

Harrisburg, PA 1979. 

An arrester failure case was analyzed occurring at t

substation capacitor bank switching with the use of a tra

work analyzer (TNA) and digital computer techniques. T

volved a 12.5-kV ground-wye distribution sytem. Results o

study, which were validated by field tests, showed the effect

configuration magnitude of transient over voltages. 

Eldon J. Rosers, and Don A. Gilles, "Shunt Capacitor 

EM1 Voltages, Their Reduction in Bonneville Power Adm

Substations," IEEE Trans. Power Appar. Syst. PAS-

Nov.-Dec. 1974, pp. 1849-1860. 

Back-to-back switching of grounded wye shunt capac

high frequency, high-magnitude current flow in overhe

___..->___. .,___._____PI_>..__> ___l.__

~~UUIIU mat GUIIUUSLU~S. IVLS~SU~CIIICIILS UL IIIUUSCU vu~ta

trol cables and receptacles; transverse voltages on fuse 

secondaries and personnel intercept voltages are repo

methods of confining transients to capacitor areas and 

*--I.-: --..:-.*."a LGcIlluqusJ as Isvlswsu. 

Paul C. Krause and William C. Mauser, "On-Line Tran

trol of Capacitor Switching to Improve System Stabili

Trans. Power Appar. Syst. PAS-92, No. 1, Jan-Feb. 

321-329. 

A simplified voltage-reactance equivalent of the one-m

fhte bus system was used. The material presented in this p

that this simplified model does not predict the performanc

machine infinite bus system with the accuracy necessary to 

the capacitor switching times so as to achieve the control 

However, it is shown that computation accuracy may be 

by including system losses. It appears that in order to app

control techniques it will be necessary to develop mor

models of the power system components. Also, faster tha

iterative, on-line computation techniques as simulated in 

should be implemented and used to calculate the switch

Until these obstacles are overcome, optimal control theor

little impact upon the power industry. 

D. 0. Wiitanen, J. D. Morgan, and G. L. Gaibrois

Tr -_-- :I-- "...1.-LI-- IF..--^:^-*^ 1- -,..I-, --A r?.

uqxutul JWILSIIIII~ I I iulwslrLa, fulayrrcal 611u JX

Results;" IEEE Trans. PowerApp. Syst. PAS-90, No. 4; 

1971, pp. 1639-1645. 

Station capacitor bank energization transients predicte

cuit model are compared to field-test results. Selection o

model is discussed. A computer solution of the model is 

Manuscript received September 19, 1985. 
FAST

TRANSIENT

TESTS :

TRIVIAL

OR

TERMINAL

PURSUIT?
immunity of electronic equipment against fast transients such as those that can be produced by contact 

bouncing in power circuits. Such transients could propagate from their source – the bouncing contact – to 

the power port of equipment, or be coupled by proximity into control cables connected to the equipment of 

interest. The issue was not so much the potential for damaging the equipment – as in surge occurrences – 

but rather the possibility of disturbing proper operation of the equipment. 

Motivated by some skepticism on how far such fast transients can propagate, measurements, augmented by 

theoretical numerical simulations were conducted on representative power wiring configurations. The 

findings, showing good agreement between theory and experiments, validated the expectation, that is, 

substantial attenuation occurs when these fast transients travel more than a few tens of meters away from 

their source. See also file “Propag EFT2 1990” for additional tests on a variety of cable configurations. 

This paper was recognized by the EMC Symposium award. 

Filename: Propag EFT 1987 
283
284
285
286
287
288
COUPLING,

PROPAGATION

SIDE EFFECTS

OF SURGES

IN INDUSTRIAL

BUILDING

WIRING

SYSTEM
Reprinted, with permission, from IEEE Transactions on Industry Applications IA-26, March/April 1990 

First presented at the IEEE-IAS Annual Meeting, Pittsburgh, 1989

Significance: 

Part 4 – Propagation and coupling of surges 

The paper reports a rare opportunity for injecting surges in a full-size building, before and after it became populated w

manufacturing and information technology equipment. The surges, of the unidirectional type or the ring-wave type 

described In ANSI/IEEE Standard C62.41-1980, were injected at one point of the system and the resulting surges 

arriving at other points were measured. 

The results show how unidirectional surges couple through transformers and can produce a ring wave component in 

response of the system. Once again, it was observed that even in this relatively large building, a 40-m long branch c

produces the transmission line reflection effect of an open-ended line only on the front part of the 0.5 μs – 100 kHz r

wave. (At 200 m/μs propagation speed, the travel time for a 40-m long line is only 0.2 μs.) 

Limited tests on the injection of the 5/50 ns EFT burst verified again the loss of steepness in the front of a nominal 5 

arriving as a 100 ns front after traveling along 95 m of branch circuits. (See pdf files “Propagation EFT1 1987” and 

“Propagation EFT2 1990” in this Part 4.) 

An unexpected side effect of these surges, applied to the power lines only, was the apparent damage suffered by the 

data line input components of some computer-driven printers. That particular finding became significant in developin

the concept of “surge reference equalizers “ – a surge protective device through which both power wires and data w

are routed, also more recently known as “multi-port surge protector.” 

Filename: Propagation 1990
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 26, NO. 2, MARCHIAPRIL 1990 193 

Coupling, Propagation, and Side Effects of Surges 

in an Industrial Building Wiring System 

Abstract-Measurements were made in an industrial building to 

determine the propagation characteristics of surges in the ac power wiring 

of the facility. The surges, of the unidirectional type or the ring-wave type 

described in ANSI/IEEE Standard C62.41-1980, were injected at one 

point of the system and the resulting surges arriving at other points were 

measured. The results show how unidirectional surges couple through 

transformers and produce a ring wave component in the response of the 

system. An unexpected side effect of these surges, applied to the power 

lines only, was the apparent damage suffered by the data line input 

components of some computer-drivep printers. 

P REVIOUS MEASUREMENTS have been reported on the 

propagation of surges in the lines used for industrial and 

residential power systems. These measurements were made in 

the laboratory on a point-to-point line, isolated from the 

building wiring system or grounds. These measurements 

indicate little attenuation of slow-front surges as they propagate along the line [I], [2]. In contrast, the propagation of the 

fast-front surges follows the behavior expected from classical 

transmission line analysis [3]-[5]. For fast-front surges, the 

amplitude changes at interfaces where an impedance mismatch 

exists. The difference between slow and fast is relative and is 

only a way to relate the duration of the surge rise time to the 

travel time of the surge along the line. In the wiring of an 

actual building, the configuration is more complex, involving 

multiple branch circuits, transformers, and changing loads. 

An opportunity arose to perform new surge propagation 

measurements during two stages of a new building project: 

before any loads were connected, and after the owner had 

moved in and various loads had been installed in the building. 

Surges representative of the types encountered in low-voltage 

power circuits were injected at various points of the wiring 

system. The resulting surges appearing at other points of the 

system were measured and synchronized with the injected 

surges by several disturbance monitors, in conjunction with 

two storage oscilloscopes. The results of these measurements 

give new insights on the coupling of surges into other parts of 

the wiring system not directly connected to the part being 

surged, as well as on their propagation along the various 

branch circuits in the building. Some of the more important 

Paper IPCSD 88-24, approved by the Power Systems Engineering 

Committee of the IEEE Industry Applications Society for presentation at the 

1988 Industry Applications Society Annual Meeting, Pittsburgh, PA, October 

2-7. Manuscript released for publication May 11, 1989. 

The author is with the National Institute of Standards and Technology, 

Building 220, Room B344, Gaithersburg, MD 20899. 

IEEE Log Number 8932121. 

480 L-L/277 L-N 

T 

30 35 40 40 

-'--[1) -nh 

+-?? 

I 

I 

I I I 55 I I 

I I I I I 

I I I I I 

191 191 El la m PANEL L PANEL HVAC PANEL 1 PANEL 2 PANEL 3 

LEGEND 

Tl - 480A/208A-l20Y 75 kVA - -40 m CONDUIT LENGTH 

T2 - 4BOA /208A -120Y 75 kVA 40 

T3 - 4BOA /208A-120Y 45 kVA - NUMBER OF BRANCH CIRCUITS 

Fig. 1. Simplified one-line diagram of building power wiring. 

findings will be described: 

coupling of the surges through the system stepdown 

transformers; 

conversion of unidirectional surges into oscillatory 

surges; 

propagation of the surges in branch circuits; 

difference in the propagation of fast-front versus slowfront surges; 

(these four types of measurements were the intended prime 

objective); and 

measurement of the coupling into adjacent but not 

connected wiring; 

anecdotal discussion of component failures in data ports 

of printers during surge tests on the power lines. 

Building Power System 

Fig. 1 shows a simplified one-line schematic of the 

building power system. Three-phase service is provided at the 

4801277-V level by a step-down transformer outside the 

building. A 4801277-V bus provides power directly to a 

lighting panel and a heating-ventilation-air-conditioning 

panel. Other loads in the building are supplied at 2081120 V 

through three transformers, each with a distribution panel 

feeding the individual branch circuits. Special efforts were 

0093-9994/9010300-0193$01.00 O 1990 IEEE 
194 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 26. NO. 2. MARCHIAPRIL 1990 

SURGE GENERATOR 

Fig. 2. Coupling of surge through delta-wye step-down transformer. 

made during the construction to record the actual length of the 

branch circuits for accurate documentation of the system. 

Surges were injected and measured at various points on the 

system that were accessible at the panels or at the end of 

branch circuits. The resulting surges arriving at other points of 

the system were measured simultaneously with, the injected 

surge, to characterize the propagation of the surges. These 

points of injection (sending end) and arrival (receiving end) 

are shown by the numbers in parentheses in Fig. 1. The length 

of conduit between the panel and the end of the branch circuit 

is also shown in the figure. These numbers will be used in 

describing the various configurations of the tests reported in 

this paper. 

Instrumentation 

Three types of instruments were used in the tests: 

a surge generator capable of delivering either the 1.2150- 

8120-ps combined surge or the 100-kHz ring wave, as 

described in [6], [7] ; 

storage oscilloscopes for monitoring the surge at the 

sending end and at the receiving end; 

disturbance monitors with auto-ranging and graphic 

output capability. 

A complete description of the instruments and the surge 

coupling method is given in the Appendix, together with some 

practical suggestions on making field tests. 

Test Schedule 

Widely varying configurations were explored; those reported here are the most illustrative of the propagation 

characteristics of typical surges in this typical installation. The 

tests were performed in two phases: first when the building 

was unoccupied, so that a no-load condition existed; and 

second, with the building operational and under representative 

loading conditions. Two types of surges were applied: 

Unidirectional Surge Propagation 

The unidirectional 1.2150-ps surge is generally considered 

representative of lightning surges on the incoming service 

entrance 161, [8]. A scenario equivalent to having such a surge 

impinge on the building was created by injecting the surge at 

an accessible point (I) of the lightning panel L in Fig. 1. From 

that point, the surge propagated to the 480-V bus, through the 

4801208 V transformers, to the distribution panels and the 

branch circuits. In a first set of measurements, the surge 

arriving at point (2) of the panel fed by transformer T1 was 

recorded for a surge injected at point (1). These measurements 

were made at the initial stage, in the empty building, before 

any of the load equipment was connected. 

The surge was injected between two lines on the primary 

side and coupled to the corresponding line-to-neutral pair of 

lines on the secondary side of the 4801208-V delta-wye 

transformer (Fig. 2). To find which of the three secondary 

phases is coupled to the primary phase of surge injection, the 

auto-ranging disturbance monitor was connected to all three 

phases on the secondary side. 

Fig. 3 shows the graphic display obtained with that 

instrument. Phase B has the highest amplitude and thus 

corresponds to the direct coupling on one leg of the transformer. Phases A and C correspond to the coupling by the two 

other legs of the transformer, effectively connected in series 

on the delta-coupled primary side. Note how the auto-ranging 

feature of the instrument changes the time scale of the graphics 

to display a longer time sweep for phase B, where the 

disturbance remains high for a longer time than for the smaller 

disturbances on phase A and C. The direct coupling path 

having been identified, the storage oscilloscope used for the 

recordings of Figs. 4-7 was then connected across phase B. 

One probe was connected to the line conductor, the other to 

the neutral conductor. A differential connection was thus 

obtained, according to the recommended practice for surge 

measurements [9]. 

Figs. 4 and 5 show the recordings for a circuit condition 

with minimum wiring connected to the 480-V bus. Only 

breakers at points (I) and (2) in Fig. 1 are closed on panels 

L and 1. Transformers TI, T2, and T3 have their primaries 

connected to the 480-V bus, but their secondaries are 

connected only to input lugs of the distribution panels. With 

the main service circuit breaker open, the utility connection is 

severed. 

In Fig. 4, a 600-V 1.2150-ps surge is generated and injected 

at point (1) (Fig. 1). In this configuration, the connection from 

the lighting panel is used as an equivalent to the service 

entrance connection for bringing the surge to the 480-V bus. 

Fig. 4(a) shows the open-circuit voltage of the surge generator 

before connection to the sending end point (1). Fig. 4(b) shows 

a 1.2150-ps unidirectional surge representative of a the voltage recorded at point (I), indicating interaction 

conventional lightning surge, applied to the service between the surge generator and the circuit under test. Fig. 

entrance and impacting the whole building; 4(c) shows the voltage recorded at the receiving end point (2), 

a 100-kHz ring wave representative of a conventional where the unidirectional surge has acquired an oscillatory 

switching surge, applied at some point of the building and component. This conversion of a unidirectional surge into an 

propagating in the system, for each of various combina- offset oscillatory wave illustrates the basis for introducing the 

tions of the connected loads. concept of ring waves into [6]. 
MARTZLOFF: SURGES IN INDUSTRIAL WIRING SYSTEM 

m. mco mm . . ammm 

d - 

1600U1 ' "' '' '' '' I 800 U/diu uertical 12.8 us/diu horiz. 

-250U '"""" 

125.8 U/diu uertical 12.5 ms/div horiz. 

R NEUT IMPULSE 

335 LRKESIDE RCU Mar 21 1987 5:53 PH 

125.8 U/diu uertical 12.5 rs/div horiz. 

B NEUT IMPULSE 

...... ; 1 125.0 U/diu uertical 12.5 ms/div horiz. 

I l. 

I IQ. m E 

I- IL A 

I LO 111 I- m c E IN . 0 c I M ."O 

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1600U- ......... ..... 

nu' 

.... 

a,,-1600U ' "' " " " 

880 U/diu uertical 51.2 us/diu horiz. 

Fig. 3. Graphic display of surges on three phases at receiving end (2). Note highest impulse on phase B, indicating direct coupling, 

and auto-ranging of two sweep rates. 

la I- IJ 3 IW Im Z 

Fig. 5 shows the recordings at the sending end and at the 

receiving end for an open-circuit voltage setting of 3000 V by 

the surge generator. A comparison of Figs. 4 and 5 yields 

interesting results. First, on a qualitative basis, the voltage at 

the receiving end in both cases contains a unidirectional 

component with superimposed oscillation. A line has been 

marked on the oscillograrns to show the unidirectional 

component at the center of the oscillation. 

Second, quantitative inspection of the results yields further 

insight into the coupling of the surge through the transformer. 

For the 600-V sending-end condition (Fig. 4), the resulting 

unidirectional component is 150 V, or a 4: 1 ratio, which is 

C NEUT IMPULSE 

335 LRKESIDE RCU Hay 21 1987 5:53 PM 

precisely the 4801120 turns ratio of the step-down transformer. 

For the 3000-V sending-end condition (Fig. 5), the unidirectional component is 750 V, again a 4: 1 ratio. This constant 

ratio demonstrates the linearity of the transformer in coupling 

unidirectional surges over a 5: 1 range of overvoltages. 

The response of the combined transformers TI, T2, and T3, 

and the 480-V bus is a 370-kHz ringing overshoot, peaking, 

respectively, at 380 V (Fig. 4) and at 1900 V (Fig. 5). The 

corresponding overshoot ratios are 3801150 = 2.53 and 

19001750 = 2.53, again showing the linearity of the response. 

Thus tests for propagation only (no nonlinear protective 

devices in the system) could be performed at low surge 

~n I n u 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 26, NO. 2, MARCHIAPRlL 1990 

Fig. 4. Recordings at sending end (1) and receiving end (2). for 600-V 

unidirectional surge. Minimum wiring connected 480-V bus (breakers to 

277 V branch circuits and 120 V branch circuits open). Voltage: 100 Vldiv; 

sweep: 2 psldiv. (a) Open-circuit output of surge generator. (b) Voltage at 

sending end. (c) Voltage at receiving end. Note line showing unidirectional 

component. 

voltages and p;oduce valid results for higher surges. However, interest in evaluating the effects of surge-protective 

devices was a motive for making further tests with relatively 

higher surge levels. This paper, however, is primarily 

concerned with the propagation aspects. Surge-protective 

devices will by fully discussed in a later paper. 

Figs. 6 and 7 show the recordings, for the same injected 

unidirectional surge, when more of the wiring system is 

connected to the circuit under test, with the system still 

isolated from the utility supply. In the test illustrated by Fig. 6, 

lighting circuits were connected. Although the fluorescent 

lights were not operating, the built-in capacitors of the ballasts 

(b) 

Fig. 5. Recordings at sending end (1) and receiving end (2), for 3-kV 

unidirectional surge. Minimum wiring connected 480-V bus (all breakers to 

277 V branch circuits and 120 V branch circuits open). Voltage: 500 Vldiv; 

sweep: 2 psldiv. (a) Voltage at sending end. (b) Voltage at receiving end. 

Note line showing unidirectional component. 

Fig. 6. Recordings at sending end (1) and receiving end (2), for 3-kV 

unidirectional surge. All 277-V lighting branch circuits connected to 480-V 

bus. Voltage: 500 Vldiv; sweep 2 psldiv. (a) Voltage at sending end. (b) 

Voltage at receiving end. Note line showing unidirectional component. 
MARTZLOFF: SURGES IN INDUSTRIAL WIRING SYSTEM 

(b) 

Fig. 7. Recordings at sending end (1) and receiving end (2), for 3-kV 

unidirectional surge. All 120-V branch circuits connected to 480-V bus 

through three 4801208-120-V step-down transformers. Voltage: 500 Vldiv; 

sweep 2 psldiv. (a) Voltage at sending end. (b) Voltage at receiving end. 

were effectively added to the circuit. At the receiving end, a 

unidirectional component of 700 V with a 150-kHz ringing 

overshoot with a ratio of about 1.25 can be observed. 

In the test illustrated by Fig. 7, all of the circuit breakers at 

panels 1, 2, and 3 of Fig. 1 were closed, adding the 

corresponding branch circuits, but without loads at the 

receptacles. The response at the receiving-end, point (2) in 

Fig. 1, still shows a unidirectional component of 700 V, but 

the ringing overshoot is now quickly damped out. 

These three circuit conditions illustrate the wide range of 

ring waves that occur when a wiring system is stimulated by a 

unidirectional surge. The general prevalence of ring waves in 

building wiring systems was confirmed by these initial tests, 

so that most of the other tests were made with the 100-kHz 

ring wave. 

Ring- Wave Surge Propagation 

Further measurements were made with the ring waves 

injected at the sending-end, point (3) in Fig. 1. Amving surges 

were recorded at the receiving-end, point (4). This scenario 

corresponds to a surge being generated by some equipment 

within the building, or the surge resulting from system 

stimulation by lightning. A few measurements were also made 

to investigate propagation from point (3) to point (5), to 

illustrate coupling through two cascaded wye-delta and deltawye transformers, but these did not produce remarkable 

results. 

The ring wave used for these measurements is defined in 

[7]; Fig. 8 shows the open-circuit voltage of a generator 

Fig. 8. 0.5 ps-100 kHz ring wave defined in 161. 

producing such a wave. The use of a 100-kHz ring wave with a 

0.5-ps rise time for the tests introduces an additional parameter: the reflections at the ends of the lines become significant 

for the lengths of some branch circuits found in this building. 

For a velocity of propagation on the order of 200 m/ps, a 

length of line greater than 50 m means a travel time longer 

than the rise time of the surge, so that reflections become 

significant [2]. 

The measurements covered a series of tests; the propagation 

path between the sending and receiving ends was progressively modified. Starting with simple point-to-point configurations, branch circuits were added by closing all breakers at 

Panel 1, connecting the secondary of transformer TI, and 

terminating the line at point (4) with various loads. 

Figs. 9, 10, and 11 show typical recordings for three of 

these configurations. Fig. 13 will present a summary of the 

various configurations with resulting propagation characteristics for the ring wave, after first discussing in detail the 

features of the three recordings of Figs. 9-1 1. 

Simple Line from Point to Point 

For the case of Fig. 9, no breakers other than 2, 3, and 4 are 

closed at Panel 1 ; the configuration is a simple point-to-point 

line from point (3) to (4). Fig. 9(a) is the recording at the 

sending end, located 40 m from Panel 1. The waveform shows 

the interaction occurring between the surge generator and the 

wiring system. Fig. 9(b) shows the recording at intermediate 

point (2) and presents some interesting features, as follows. 

1) The initial peak of 1100 V at the sending end has been 

attenuated to 750 V, a ratio of 0.68, after 40 m of travel from 

point (3) to point (2). 

2) A second peak is visible, about 0.6 ps after the first: the 

pulse arriving at the open end of point (4) is reflected with 

twice the amplitude and travels back toward Panel 1. It amves 

at Panel 1 after a 2 x 55-m travel, requiring about 0.6 ps for 

the round trip at a velocity of 200 mlps; this expected time 

matches the time observed between the two peaks. 

3) The returned pulse, arriving at Panel 1 after 110 m 

round-trip travel in the line, has lost some its higher frequency 

components [3] and is less sharp than that coming from the 

sending point. 

4) The returned pulse travels further back to sending point 

(2), where it arrives to produce a discontinuity in the trace 

visible at the first zero crossing of the trace in Fig. 9(a). 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 26, NO. 2, MARCHIAPRIL 1990 

Fig. 9. Recordings at sending end (3), intermediate point (2) of Panel 1, and 

receiving end (4) for 100-kHz ring wave and point-to-point propagation in 

simple line. No additional branch circuits have been connected between 

points (3) and (4). Voltage: 500 Vldiv; sweep: 2 psldiv. (a) Voltage at 

sending end. (b) Voltage at Panel 1. (b) Voltage at receiving end. 

5) The second peak of the ring wave (negative peak) and the 

third peak (positive peak) do not show very significant changes 

in amplitude among the three oscillograms. These peaks do not 

contain the high-frequency components associated with the 

initial fast rise, and thus are not significantly attenuated by the 

40- or 95-m travel along the line. 

Finally, Fig. 9(c) shows the initial peak, which traveled 

95 m from the point of origin and, therefore, might be 

expected to have been attenuated about twice as much as 

between points (3) and (2), a 40-m travel. (The attenuation is 

not quite proportional to distance, as the higher frequency - -- 

components suffer greater attenuation in the early parts of the 

Fig. 10. Recordings at sending end (3), intermediate point (2) of Panel 1, 

and receiving end (4) for 100-kHz ring wave. Transformer T1 is connected 

at Panel 1 and supplies power to branch circuits 3 and 4. Voltage: 500 VI 

div; sweep 2 psldiv. (a) Voltage at sending end. (b) Voltage at Panel 1. (c) 

Voltage at receiving end. 

travel; the waveform contains fewer of them, and is therefore 

less attenuated in the end parts of the travel [4] .) The expected 

amplitude of the attenuated pulse arriving at point (4) should 

be about 1100 V, the initial value, attenuated by about two 

times (twice the distance) the 0.68-ratio noted between points 

(3) and (2): 1100 x 0.6g2 = 5 10. The oscillogram shows, on 

the contrary, a 1050-V peak, or close to twice the expected 

arriving pulse value-the doubling effect at the open-ended 

transmission line. 

To summarize the observations from Fig. 9 data, the first 

peak, with a duration shorter than the travel time in the line, 

exhibits all the behavior of pulses traveling in a transmission 

line. Later peaks, with durations longer than the travel time, 
MARTZLOFF: SURGES IN INDUSTRIAL WIRING SYSTEM 

Fig. 11 

wave. 

loads 

. Recording at receiving end point (4) for 100-kHz ring 

Transformer T1 is connected at Panel I and supplies power to all 

connected to branch circuits of that panel. Voltage: 100 Vldiv; 

sweep: 2 psldiv. 

exhibit little attenuation for the 40- and 95-m travel distances 

in this building. 

Adding Complexity to the System 

Fig. 10 shows the recordings at the same three points 

discussed for Fig. 9 but with transformer T1 connected to 

Panel 1 and supplying power to that panel. All other branch 

circuits of that panel are still left disconnected. 

Fig. 10(a) shows the sending end to be little different from 

Fig. 9(a), except that later reflections have modified subsequent peaks. Fig. 10(b) shows again the double peak at Panel 

1, resulting from the initial pulse arriving from the sending 

end, followed by a reflected pulse coming back from the 

receiving end. Fig. 10(c) shows the initial pulse arriving at the 

receiving end, doubled by the effect of the open-ended line in 

the same manner as discussed for Fig. 9(c). The same 

conclusions as those drawn from Fig. 9 apply here: the initial 

peak behaves as a pulse in a transmission line, and subsequent 

peaks are not significantly attenuated. The inspection of 

subsequent peaks is no longer as simple as it was for Fig. 9, 

because with the added elements in the transmission path, the 

cumulative effect of the reflections distorts these peaks. 

One further stage in the reconstitution of the building wiring 

system is illustrated by Fig. 11, where all branch circuits at 

Panel 1 are now connected, supplying their normal loads with 

power obtained from transformer T1, which is energized. For 

the same ring wave of 1100 V applied at the sending end, point 

(3) in Fig. 1, the first peak at the receiving end, point (4), is 

only 220 V, down from the 1000 V of Fig. 10. This large drop 

is caused by the impedance mismatch resulting from connecting the many parallel transmission lines at Panel 1. Even the 

subsequent peaks show a reduction (down to 100 V from the 

400 V of Fig. lo), but the waveform is so distorted by the 

cumulative effect of the multiple reflections that a simple 

analysis by inspection is no longer possible. Computer 

modeling might be an interesting task for specific cases where 

a rigorous prediction would be of interest [lo], [l 11 . Summarizing now all the observations from the ring-wave 

measurements, three major conclusions emerge. 

1) In a building with no or few loads, fast-front surges 

propagate in a manner that can be readily predicted by 

classical transmission line analysis. 

Fig. 12. Recording at receiving end point (4) for fast-front surge, injected at 

sending end (3) by generator with 5 ns nominal rise time. 

2) Adding many parallel branch circuits produces amplitude 

reductions, predictable by the impedance mismatch at the 

point of parallel connections. 

3) Surges having durations longer than the travel time 

within the building propagate with very little attenuation for a 

relatively simple configuration. In a building of complex 

configuration, cumulative effects of reflections from branch 

circuits with random lengths tend to reduce subsequent peaks 

by averaging out their interaction. This observation, however, 

leaves open the possibility of a random combination that could 

enhance subsequent peaks, or at least not reduce them by the 

averaging effect. 

Propagation of Fast Transients 

The unplanned availability of a surge generator having a 

rise time of a few nanoseconds made possible incidental tests 

of the propagation of fast transients in the building. However, 

suitable probes were not available to record accurately the fast 

rise time at the sending end. Only the receiving-end surges 

could be observed with confidence, because the rise time 

observed there was indeed longer than the response time of the 

probes. For a nominal 5-ns rise time available at the generator 

output and injected at the sending end, point (3), the rise time 

observed at the receiving end, point (4), was 100 ns (Fig. 12). 

This increase in rise time confirms the findings reported 

earlier in [4] and [S]. This observation also confirms that 

concerns over the presence of surges with rise times of a few 

nanoseconds are not generally applicable to power systems for 

locations remote from the source of the fast transients [12]. 

Summary of Ring Wave Propagation 

Fig. 13 presents a summary of the attenuation observed for 

various configurations of the building wiring. The many 

values obtained show how difficult a prediction would be for a 

complex system, but such data are helpful in providing some 

upper and lower bounds. In this figure, the transmission path 

is represented by a branch circuit from the sending end (left 

line) to Panel 1 (box at center) and an emerging branch circuit 

to the receiving end (right line). 

From the top to the bottom of Fig. 13, various combinations 

of added circuits are shown at Panel 1 and at the receiving end, 

labeled A through K. The numbers shown at the right of each 

configuration, A through K, are the observed values at the 

receiving end. For the first column, the numbers are the value 
200 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 26, NO. 2, MARCHIAPRIL 1990 

SENDING END 

RING WAVE 

(100%) 

-0 

1000 OHMS 

100 OHMS 

(FIGURE 9) 

f 0- 

RECEIVING END 

PERCENTAGES 

FIRST THIRD 

PEAK PEAK 

1.9 NSR 

1.7 NSR 

1.1 NSR 

95 100 

70 NSR 

90 75 

65 NSR 

90 90 

25 15 

10 10 

5 NSR 

NSR - NO SIMPLE 

READING 

POSSIBLE 

'-1 BRANCH CIRCUITS 

CONNECTED 

Fig. 13. Schematic representation of configurations for propagation path 

from sending end (3) to receiving end (4). Numbers at right are receivingend percentages of first and third peaks of arriving surges for ring wave of 

100 percent applied at sending end. 

of the first peak; for the second column, the numbers are the 

value of the third peak (second positive peak of the ring wave). 

These numbers are given as percentages of the sending-end 

values of the first peak. In some cases, the response of the 

circuit was so complex that no simple reading could be made at 

the time of the third peak, hence the entry "NSR" in the 

second column. 

Configurations A, B, and C show all breakers open at Panel 

1, so that no direct metallic connection exists between the 

sending and receiving ends. However, the proximity of the 

wires in common conduits shared for part of the distance is 

sufficient to couple some of the surge by the stray capacitance. 

Adding load at the receiving end between the line and the 

neutral (configurations B and C) provides a measure of the 

transfer impedance of that coupling. 

Configurations D to G show increasing complexity in the 

transmission path, with added branch circuits at Panel 1 E, 

added transformer F, and both added G, but still no power 

applied to the primary of the transformer. For configurations 

H to K, the transformer primary is energized, supplying 

power to Panel 1 and its loads. In H, only the transformer is 

connected; in I the branch circuits have been added. In J and 

K, two different types of loads have been connected at the 

receiving end, between line and neutral, to show how a lowimpedance termination reduces the reflections observed with 

higher impedance terminations. 

An unexpected side effect of the surges was found only upon 

the resumption of normal weekly operations, after the tests 

(which were performed on a weekend). Two laser printers 

shared by several personal computers and thus linked by a data 

cable had become inoperative. Repair by the service organization of the manufacturer diagnosed failed components in the 

data line port of the printer. Yet, all the surges had been 

injected into the power lines, and at levels initially deemed 

high but not hazardous to equipment designed for connection 

to typical ac power systems. 

The exact nature of the failure is unknown because the 

service organization performed repairs as a routine operation, 

not associated with this surge measurement project, and its 

repair records were not available. However, with hindsight, 

this side effect might well be explained by the following 

hypothetical, but plausible scenario. (Anecdotal information 

on problems encountered at other facilities indicates that this 

type of problem is not at all unusual. A detailed description of 

a scenario that could lead to this failure of information 

technology systems should be useful for understanding the 

problem-a first step toward avoiding it in the future.) 

Many applications of information technology equipment 

require that peripheral terminal devices (video displays, data 

entry keyboards, printers) be located at some distance from the 

computer. Frequently, terminal devices are powered by a line 

cord plugged into a wall receptacle which is on a different 

branch circuit from that of the computer. Sometimes, the 

branch circuits are supplied by different transformers (Fig. 

14). This situation produces multiple ground references and 

loops of communication cables, with the possibility of 

substantial differences in the potentials of points expected to be 

at the same zero-reference potential. 

The general practice is to have the device chassis bonded to 

the grounding conductor of the power cord. Also as a general 

practice, the zero reference of the signal circuit is bonded to 
MARTZLOFF: SURGES IN INDUSTRIAL WIRING SYSTEM 20 1 

DATA LINK 7 

Fig. 14. Typical arrangement for powering computer and its peripherals. 

\ 

REFERENCE LINE 

6 LOGIC REFERENCES 

(CHASSIS) - 6 

A - SINGLE-ENDED CIRCUIT 

SIGNAL LINES 

+- LOGIC REFERENCES 

(CHASSIS) 

B - DIFFERENTIAL CIRCUIT 

Fig. 15. Links between driver and receiver of computer and remote 

terminal. 

the chassis. Fig. 15 is a schematic of the receiver and driver 

circuits used to communicate between the computer and its 

terminals. This figure shows that the designer has the option of 

using single-ended transmission, where the signals are developed with respect to a zero-reference plane, or a differential 

transmission, where signals are developed as a voltage 

difference between the two signal lines. 

Typically, a twisted-pair cable or a coaxial cable is used 

between the computer and the terminal. Various types of 

shielding may be included with these cables. Intuitively, one 

might expect that the receivers and drivers linked by a shielded 

cable should be least susceptible to interference or damage. 

However, the problem encountered here does not involve 

transients coupled capacitively or inductively onto the data 

communication cables. Rather, the problem is a difference in 

the reference potentials of the chassis of the two devices, 

resulting from surges being diverted in the separate grounding 

connections of the two devices. 

Fig. 16 shows a schematic of a system consisting of a 

computer and a terminal, powered by different sources and 

linked by a shielded two-wire differential data line with zero 

reference bonded to the chassis of each device. In each power 

supply, an electromagnetic interference (EMI) filter or surge 

suppressor has been provided between the ac lines and the 

grounded chassis. 

Let us now examine the scenario unfolding when surges 

arrive on the power supply line of the computer. The 

expected role of the EM1 filter is to divert these surges to 

ground-that is, by passing the current i through the filter, 

returning it to the power source by the grounding conductor G. 

This path unavoidably has a finite inductance L. Consequently, the fast-changing current i produces a voltage drop 

L.di/dt between the computer chassis (which is the signal 

reference at the sending end) and the grounding connection of 

the power supply. At the terminal, no surge event is occurring 

at that time; the terminal chassis (which is the signal reference 

at the receiving end) stays at the potential of the grounding 

connection G' of its own power supply. The two chassis, 

therefore, the two references of the data link ports are now at 

different potentials. This difference is the L - di/dt voltage 

from the surge, plus any other voltage that might develop in 

the undefined path linking the two grounding points G and G ' . This scenario explains how surges, initially limited to the 

power lines, can impact the data port components of a system 

powered without the installation of coordinated surge suppressors on both power port and data port. 

Thus, with hindsight, this anecdote illustrates the need for 

developing and applying means of surge protection that will 

avoid the potential difference between ends of the data link. 

This need presents a challenge to the designers of an 

information technology system and an opportunity for the 

manufacturers of packaged surge-protective devices: coordinating protection of the power supply port with protection of 

the data link port in a single unit, with the same return path for 

surges diverted by the two protective devices. Such coordinated surge suppressors built for protecting both ports and 

using a single ground reference are now commercially 

available. They are sometimes referred to as "local ground 

window" [12]. 

1) The response of the step down transformer and its 

associated bus wiring to stimulation by a 1.2/50-ps unidirectional surge contains two components: 

a unidirectional component matching the stimulation, and 

a ringing overshoot at a frequency dependent upon the 

circuit characteristics. 

2) The unidirectional surge couples through the transformer 

according to the turns ratio, with negligible attenuation. The 

ringing overshoot frequency depends on the circuit parameters; its peak can exceed twice the peak of the stimulus. 

3) The existence of multiple branch circuits in the building 

wiring reduces the overshoot and affects its frequency but does 

not change the unidirectional component. 

4) A ring wave with a rise time shorter than the travel time 

in a simple point-to-point line produces the expected enhancement of the surge at an open-circuit receiving end. Adding 

loads at the end of the line reduces the amplitude of the surge 

at that point in a predictable manner, according to the classical 

transmission line theory. 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 26, NO. 2, MARCHIAPRIL 1990 

COMPUTER 

7 

TERMINAL 

L ' 

N' 

G' 

SEE NOTE 

I- L . d~/dt + OTHER INTERFERENCE 9 E SOME ORGANIZATIONS IN-SIST ON HAVING ONE END OF THE SHIELD FLOATING 

Fig. 16. Scenario for creating difference of potential between signal references by diverting surge through grounding conductor of 

one side of the system. 

5) Adding branch circuits and other circuit elements along 

the propagation path introduces mismatches in the line 

impedance, reducing the amplitude of the initial peak of the 

surge arriving at the receiving end. Subsequent parts of the 

surges, however, are less affected. 

6) Providing protection against power line surges at the 

power line interface of devices linked by a data communication circuit does not guarantee that surges occurring in the 

power line environment will not cause damage to the devices. 

A more comprehensive protection scheme, coordinating both 

the power line and the data line, is required to ensure 

protection. 

For complete documentation of the test procedure, the 

instrumentation used in the tests is listed in this Appendix1. 

Also, some practical aspects of the procedure and logistics are 

recited here; some were planned, some resulted from hindsight. These are offered as helpful hints to anyone planning 

this type of field test. 

Instrumentation 

Surge Generator: A KeyTek Model 711 was used, with 

F31 coupling network and plug-in pulse waveshaping networks: P7 for 1.2150 ps unidirectional waveform, and P1 for 

100-kHz ring wave as described in [6]. 

Storage Oscilloscopes: Two Tektronix 7934 with 500- 

MHz bandwidth were used, each with pairs of matched 1000: 1 

P6015 probes or 100: 1 P6009 probes. The probes were 

connected differentially with a 7A26 vertical preamplifier 

performing the differentiation. 

Auto-ranging Monitors: A BMI 4800 Powerscope with 

three-phase multimode monitoring capability, graphic display, 

auto-ranging of sweep rate, and adjustable threshold was 

employed. 

Certain commercial instruments are identified in this paper to specify the 

experimental procedure adequately. Such identification does not imply a 

recommendation or endorsement by the National Institute of Standards and 

Technology, nor does it imply that these instruments are necessarily the best 

available for the purpose. 

Procedure 

Surge Coupling: The surges produced by the generator 

were injected into the building wiring by a 2-pF coupling 

capacitor for the initial unpowered tests. For subsequent tests 

with power on, a coupling circuit was used so that the surge 

was applied at a fixed phase of the sine wave. 

Differential Measurements: All measurements were made 

with a differential connection of two matched probes. Admittedly, the 7A13 differential preamplifier would be more 

appropriate for this type of connection, but it was not available 

for these field tests. The common-mode rejection and noise 

background were checked by connecting the two probes 

together and then to the high and low side of the point of 

measurement. 

One of the planned measurements was to measure the 

voltage between the building grounding system and one of the 

so-called "isolated separate dedicated ground" ground rods, 

deliberately installed in the building as a tutorial example of 

questionable practice. For this measurement, however, the 

differential connection with the 7A26 preamplifier and unshielded oscilloscope was not successful. The levels of 

voltages observed across the desired points of measurement 

had the same order of magnitude as the noise background 

observed in the check just described. Instrumentation less 

sensitive to this interference was not available for these field 

tests, but new attempts might be made in the future. 

HELPFUL HINTS 

While these hints might seem trivial, they are offered after 

successful planning or hindsight and as a counterpoint to the 

more fundamental conclusions of the paper. If they can save 

time or reduce problem? for future experimenters, including 

them in this paper will have been worthwhile. In random order 

of importance, here is a list. 

Whenever field measurements involve several people at 

different locations, walkie-talkies are far superior to 

direct voice communication (attenuated by distance), 

fast-footed couriers, or an internal telephone network. 

Whenever the work is carried on outside of normal 

working hours, be sure that the team includes one 
MARTZLOFF: SURGES IN INDUSTRIAL WIRING SYSTEM 

individual thoroughly familiar with the facility, and 

custodian of all keys to rooms, cabinets, and interlocks. 

The availability of monitoring instruments that combine 

logging and timing of events provides a very useful 

correlation of recordings made by independent oscilloscopes. 

Always bring at least twice as much film as you think you 

will need. The same remark applies to materials for filing 

oscillograms at the site, and to extension cords. 

Bring some spare fuse holders for the instruments. These 

protruding accessories have a unique capability of coming loose during shipment and getting lost in the packing 

material. 

When scheduling the rate of progress in the tests, 

rememb r that Murphy was an optimist! fi 

The measurements and the analysis of the results were made 

possible by the cooperation and contributions of many 

organizations and individuals, as follows: 

Building Industry Consulting Service International, for 

coordinating an informal consortium of private sector 

companies in support of this work; 

Alex McEachern, for making the facilities of Basic 

Measuring Instruments available as a measurement site; 

Joseph Connolly, Robb Gould, Jim McDaniels, and 

Christine Mihelich, for their participation in the measurements; 

Maurice Tetreault, for participating in the measurements 

and suggesting the principles underlying the analysis of 

the surge side effects on the data ports; 

Catherine Fisher, for suggestions on reporting these 

results. 

These contributions to improved understanding of the problems, and thus greater system reliability, are most gratefully 

acknowledged. 

[I] F. D. Martzloff, "The propagation and attenuation of surge voltages 

and surge currents in low-voltage ac power circuits," IEEE Trans. 

Power App. Syst., vol. PAS-102, no. 5, May 1983. 

F. D. Martzloff and H. A Gauper, "Surge and high-frequency 

propagation in industrial power lines," IEEE Trans. Ind. Appl., vol. 

IA-22, no. 4, JulyIAug. 1986. 

L. V. Bewley, Traveling Waves on Transmission Lines. New York: 

Dover, 1963. 

F. D. Martzloff and P. F. Wilson, "Fast transient tests: Trivial or 

terminal pursuit?" in Proc. Int. Zurich Symp. Electromagnetic 

Compatibility, 1987. 

F. D. Martzloff and T. F. Leedy, "Electrical fast transient tests: 

applications and limitations," IEEE Trans. Ind. Appl., vol. 26, no. 1, 

pp. 151-159, Jan./Feb. 1990. 

IEEE Guide on Surge Voltages in Low- Voltage AC Power Circuits, 

ANSIIIEEE Standard C62.41-1980. 

R. B. Standler, "Equations for some transient overvoltage test 

waveforms," IEEE Trans. Electrornagn. Compat., vol. EMC-30, 

no. 1, Feb. 1988. 

Insulation Coordination within Low-Voltage Systems, Including 

Clearances and Creepage Distances for Equipment, IEC Pub. 664- 

1980. 

IEEE Guide on Surge Testing for Equipment Connected to LowVoltage AC Power Circuits, ANSIIIEEE Standard C62.45-1987. 

A. K. Agrawal, H. M. Fowles, and L. D. Scott, "Experimental 

characterization of multiconductor transmission lines in inhomogenous 

media using time-domain techniques," lEEE Trans. Electromagn. 

Compat., vol. EMC-21, no. 1, Feb. 1979. 

A. H. Paxton and R. L. Gardner, "Application of transmission line 

theory to networks with a large number of component wires," in Proc. 

Int. Zurich Symp. Electromagnetic Compatibility, 1987. 

F. D. Martzloff, "Protecting computer systems against power transients," IEEE Spectrum, April 1990. 

Fraqois D. Martzloff (M'56-SM'80-F'83) 

completed his undergraduate studies in France and 

received the M.S.E.E. degree from the Georgia 

Institute of Technology, Atlanta, and the M.S.I.A. 

degree from Union College, Schenectady, NY. 

After a long career at General Electric, he joined 

the staff of the National Bureau of Standards (now 

renamed National Institute of Standards and Technology) to work on conducted electromagnetic 

interference issues. His early experience covered 

high-voltage fuses, high-voltage bushings, and electronic power conversion; the latter marked a turning point to issues of 

overvoltage effects on semiconductors, surge suppression, metal oxide 

varistor applications, and interference mitigation. In addition to his research 

work on surge propagation and mitigation, he is contributing to the 

development of standards on surge environment and surge suppression within 

IEEE, ANSI, and IEC by documenting his measurements, as in the present 

paper, by presenting tutorials, and by drafting new standards. 
FASTTRANSIENT

TESTS:

APPLICATIONS

AND

LIMITATIONS
Significance: 

Part 4 – Propagation and coupling of surges 

In the early 80’s, considerable interest arose in developing an IEC standard test method to assess the immunity of 

electronic equipment against “fast transients” such as those that can be produced by contact bouncing in power circ

 Such transients could propagate from their source – the bouncing contact – to the power port of equipment, or be 

coupled by proximity into control cables connected to the equipment of interest. The issue was not so much the 

potential for damaging the equipment – as in surge occurrences -- but rather the possibility of disturbing proper 

operation of the equipment. 

Motivated by some lingering skepticism among US industry on the necessity to require these tests across-the-board 

industrial equipment, measurements, augmented by theoretical numerical simulations were conducted on representa

power wiring configurations to help determine how distance of propagation will reduce the threat. The findings, show

good agreement between theory and experiments, validated the expectation, that is, substantial attenuation occurs w

these fast transients travel more than a few tens of meters away from their source. See also file “Propag EFT1 1987

this Part 4 for earlier tests on different cable configurations and numerical modeling technique. 

Filename: Propag EFT2 1990 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 26. NO. I. JANUARYIFEBRUARY 1990 151 

Electrical Fast-Transient Tests : Applications and 

Limitations 

Abstract-According to a new standard of the International Electrotechnical Commission (IEC), a fast-transient test must be applied to 

the connecting cables of electronic equipment. The purpose of the test 

is to demonstrate equipment immunity to fast transients resulting from 

switching. Tests and simulations of the propagation and attenuation of 

these fast transients in typical connecting cables are described, placing 

the IEC requirements in perspective. 

C ONTINUING technical progress in the performance of 

electronic control systems has been made possible by 

fast logic devices operating at low levels of voltage and 

power, with high component density and interactive peripherals. These undisputed advantages, however, can make the circuits increasingly sensitive to external interfering overvoltages 

resulting in amplitude-related hardware failures or rate-ofchange-related system upsets. To guard against such effects, 

immunity to such failure must be demonstrated by realistic 

tests which validate the design rules and installation practices 

of the equipment. However, a balance must be found between 

overdesign and cost limitations. 

In contrast with traditional surge testing aimed at validating 

insulation levels or the energy-handling capability of surge 

protection devices, recent proposals advocate fast-transient 

tests demonstrating interference immunity. The new test waveforms represent electrostatic discharge effects or reignitions 

occurring during switching sequences. The basis for such proposals is an increasing awareness of the significance of fast 

transients. This new awareness results from two factors: 1) 

increases in the bandwidth and writing speed of oscilloscopes 

which makes possible the detailed observation and measurement of these phenomena, and 2) field failures of equipment 

which pass traditional tests but fail when subjected to the new 

fast-transient tests. 

The rationale for requiring the new tests is based on the 

assumption that interfering transients generated by power circuit switching will couple into adjacent power or signal lines. 

These interfering transients will then propagate toward susPaper PID 89-09, approved by the Petroleum and Chemical Industry Committee of the IEEE Industry Applications Society for presentation at the 

1989 Petroleum and Chemical Industry Committee Technical Conference, 

San Diego, CA, September 11-13, and at the 1989 IEEE Industry Applications Society Annual Meeting, San Diego, CA, October 1-5. Manuscript 

released for publication February 22, 1989. 

The authors are with the Electricity Division, National Institute of Standards and Technology, Building 220, Room B-344, U.S. Department of Commerce, Gaithersburg, MD 20899. 

IEEE Log Number 8931459. 

TIM 

(MICROSECONDS) 

Typical transient produced by contact opening. 1: voltage recovery 

of circuit (kHz frequency); 2: reignition (ns collapse); 3: complete recovery 

(ps duration). 

ceptible equipment. The fast transients contain high-frequency 

components that are expected intuitively to suffer greater attenuation than low-frequency components as they propagate 

in the lines. Verifying and quantifying this intuitive expectation provides a perspective on the severity of the situation 

and helps define realistic test requirements. To that end, this 

paper describes specific measurements conducted for typical 

low-voltage power line configurations. An attenuation model 

is proposed which provides a tool for understanding the significance of the line parameters and extends the usefulness of 

the results to general cases. 

THE IEC ELECTRICAL FAST TRANSIENT TEST 

The Technical Committee TC65 of the International Electrotechnical Commission (IEC) has completed a document [l] 

requiring immunity of industrial process control equipment to 

fast transients be demonstrated. According to that document, 

fast transients must be applied to the incoming power lines as 

well as to the inputloutput data lines of this equipment. 

The origin of fast transients is the well-known behavior of some air contactors when they interrupt an 

inductive-capacitive circuit: as the contacts part, the arc drawn 

between the contacts becomes unstable and the current is then 

interrupted, only to be re-established promptly by reignition 

across the gap. A race occurs between the dielectric strength 

of the increasing gap, and the voltage recovery of the circuit. 

Consequently, a sequence of clearing and reignition ensues, 

with an eventual full clearing of the circuit. As shown in 

Fig. 1, each of the clearings produces the beginning of an 

oscillation at relatively low frequency depending on the local 

circuit parameters. As the voltage across the gap increases 

at a rate corresponding to this frequency, reignition occurs if 

the gap withstand voltage is exceeded. The waveform of each 

0093-999419010100-015 1$01.00 O 1990 IEEE 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 26, NO. I. JANUARYIFEBRUARY 1990 

VOLTAGE 

t 

VOLTRGE 7 Pulse 

~Lacing of 5ingle pulses depends ' on the test voltage level 

VOLTAGE /-- Burst 

1 Burst length 1 

1- Burst period - 300 ms -4 

Fig. 2. Fast-transients specified by IEC 801-4. Top: waveshape of single 

pulse for 50-0 load. 

sequence is then a series of pulses, each with a relatively slow 

voltage rise followed by collapse in a few nanoseconds. 

Thus the circuit behaves as a generator of transient pulse 

bursts. Each pulse has a slow rise ending in an abrupt collapse. 

Because the pulses occur in bursts, these single pulses have 

the same effect as pulses with fast rise time and slower decay. 

For standardization purposes, the IEC document specifies an 

electrical fast transient (EFT) 5150-11s waveform as shown in 

Fig. 2, where the rise time is 5 ns and the duration (full width 

at half maximum) is 50 ns. The repetition rate of the pulses 

within a burst is not specified. However, the duration of the 

burst and the interval between bursts are specified. 

Three limitations exist in the severity of the situation, resulting from the fast transient nature of the test and the conditions 

of application. 

1) As described in the IEC test procedure, the test transients are coupled to the equipment under test (EUT) by a 

capacitive "clamp" through which the power or data cable is 

inserted. This clamp is a mechanical test fixture consisting of 

two hinged metallic plates designed to provide a fixed and repeatable capacitance between the output of the pulse generator 

and the input cables. Discrete capacitors can also provide the 

coupling. This coupling method produces a capacitive divider: 

the high side of the divider is the coupling capacitor, and the 

low side is the internal capacitance to ground of the EUT (Fig. 

3). This effect, therefore, is influenced by the design of the 

DATA OR 

POWER 

CABLE 

I I I 

GROUND PLANE 

COUPLING CLAflP 

\ 

- J PULSE 

GENERATOR 

1- T 

Fig. 3. Coupling electrical fast transient (EFT) pulses into equipment cables. 

EUT 

I 

A 

In 

I 

I 

I 

I 

'r 

equipment and becomes a design parameter controlled by each 

manufacturer. 

2) The freedom left to those who perform the test on setting the repetition rate of the pulses leaves some ambiguity 

on the total number of the pulses applied to the EUT. The 

test presumably aims at identifying an interference with the 

logic of the circuit. That interference is likely to occur during 

critical state transitions of the semiconductor circuitry of the 

equipment, malung timing and coincidence very significant. 

Advocates of electrostatic discharge tests make a case that up 

to 10000 pulses should be applied at random to produce a 

high degree of confidence that the issue of pulse timing has 

been addressed 121. 

3) In an actual installation, attenuation between the point of 

coupling and the victim equipment will reduce the severity of 

the impinging transient. The line propagation characteristics 

reduce the peak amplitude and increase the rise time of the 

pulse as it travels down the line. This effect, therefore, is 

influenced by the line regardless of this equipment design. 

Although our measurements and modeling address this third 

limitation, the two other limitations should also be kept in 

mind. 

Two line configurations were selected as typical of lowvoltage power lines: a three-conductor line in a steel conduit, 

and a three-conductor cable with a nonmetallic jacket. The 

steel conduit line is easy to set up in the laboratory by arranging conduit sections in a zig-zag pattern. This arrangement 

allows placing the start and the end of the line next to each 

other to minimize the distance between the oscilloscope used 

for measuring the signals at either end of the line. At one end 

of the line, called the "sending end," single pulses with the 

IEC EFT waveform were injected. At the other end, called 

the "receiving end," the arriving pulses were recorded for 

various line terminating impedance configurations. The same 

oscilloscope was used to measure the waveforms at the two 

ends of the line. Thus the voltage ratios of the waveforms do 

not depend on the absolute accuracy of the instrumentation. In 

addition, the differential connection of the oscilloscope probes 

allows measurement without introducing an additional ground 

in the line, which is already grounded by its connection to the 

pulse generator. This arrangement also removes any question 

of variations of the stray capacitances to ground, provided that 
MARTZLOFF AND LEEDY: ELECTRICAL FAST TRANSIENT TESTS 153 

the instruments stay in the same position for the two measure- loO1 I 

ments, at both the sending and receiving end. The experi- 90 

MEASURED INPUT PULSE mental details of the measurements are discussed in , oo a MOOEL - WIRES NEAR CWTER OF CMlWIT 

Appendix I. !4 

a 70 

Because the conduit contains the electromagnetic fields as- $ 

z 60 sociated with the propagating pulse, there is no need to be 

concerned with coupling between adjacent sections of the line : " 

in the zigzag, nor with losses by radiation from the line act- a * 

ing as an antenna. For the nonmetallic jacketed line, however, 9 XJ 

these effects must be addressed. - .$ 20 

To investigate the effect of radiation losses from the non- ' ODEL - WIRES NEAR WALL OF CONDUIT 

metallic jacketed line, three experiments were performed outdoors. First, measurements were made with the line arranged o 10 20 30 40 60 70 80 so 

in a wide loop, elevated from ground and away from metallic TIME (NANOSECONDS) 

objects. TO assess the importance of the proximity of the line Fig. 4. Composite plot of measured and modeled pulses, as percent of 

to ground, the line was later dropped to ground level and a amplitude, for wires enclosed in metal conduit. 

second set of measurements was made. Last, a coil of identical cable was also subjected to the same measurements. The 

observed difference of the three sets of measurements was 

negligible for practical purposes. 

For the case of the conduit-enclosed line, a previous paper 

reports in detail the experimental results and theoretical expectations from a model solely based on the geometry of the 

line [3]. As a convenience for the readers and for comparison 

with the new results on nonmetallic jacket line, the results of 

the first configuration are summarized here. For the nonmetallic line, we used a new model based on measurable electrical 

parameters of the line, rather than the geometry of the line. 

Appendix I gives details on instrumentation considerations; 

Appendix I1 gives details on the line propagation model. 

Steel Conduit Line 

A representative low-voltage indoor single-phase power line 

was installed in the laboratory, with three conductors in a steel 

conduit. The line length was 63 m (20 thin-wall conduit sections, each the standard 10-ft length), using typical industrial 

components such as wire nuts at the splices in the wire. Three 

insulated (600 V) copper conductors of number 14 AWG 

(1.6 mm diameter), pulled in the 20-mm diameter conduit (112 

EMT), represent the phase, neutral, and grounding conductors. At the sending end, the neutral conductor was bonded to 

the grounding conductor and to the conduit. 

At the receiving end, the neutral conductor was not bonded 

to the grounding conductor, nor to the conduit. This type 

of connection is typical U.S. practice for a branch circuit 

originating at a service entrance. 

For modeling purposes, it is much easier to assume that 

the line is uniform. In practice, the location of the conductors 

vary within the conduit due to sag, bends, or crossed wires. 

Therefore, we considered two limiting cases: 1) the wires are 

located at the center of the conduit, separated from each other 

by the thickness of the wire insulation, and 2) the wires are 

located against the wall of the conduit. In the numerical analysis, a model parameter k accounted for these two positions. 

A detailed description of the analysis is given in [3]. 

Fig. 4 shows the composite results of the measurements and 

modeling for the IEC EFT pulse arriving at the receiving end. 

The results are expressed as a percentage of the pulse applied 

Fig. 5. Oscillograms of measured pulses, as percent of input amplitude 

(0 m), for various lengths of wires enclosed in metal conduit (25, 38, and 

63 m). 

at the sending end. All the measurements were made with a 

injected pulse of 2.5 or 5 V, to avoid the problems of electromagnetic interference or safety associated with the levels 

of 500-4000 V specified by the IEC standard. The linearity 

of the circuits allows this simplification. The pulse, measured 

at the receiving end, is 50 percent of the sending-end pulse 

and lies between the two limiting cases of the model. Qualitatively, the model pulses look very much like the measured 

pulses. From the position of the measured pulse with respect 

to the two limiting cases, we can then assign a value to the 

model parameter k such that attenuation for the 63-m line is 

accurately matched. Given this empirically determined value, 

the model will show the effect of varying the line length. 

Measurements for several line lengths (Fig. 5) show 

receiving-end peak amplitudes of 76 and 65 percent of the 

sending-end peak amplitude for 25- and 38-m lines, respectively. Corresponding peak amplitudes predicted by the model 

are 75 and 67 percent (Fig. 6). For the greater line lengths, 

where measurements would be cumbersome, the model can 

now readily provide the answers. Fig. 6 includes examples 

for 120- and 240-m lengths, showing the continued attenuation and smearing of the pulse as the line length increases. 

Nonmetallic Jacketed Line 

For this type of line, the insulation holds the conductors in 

a fixed and uniform geometry relative to each other, in contrast with the random location of the conductors pulled in a 

conduit. It is possible to determine, by a simple set of measurements or calculations, the parameters of the transmission 
154 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL 26. NO I. JANUARYIFEBRUARY 1990 

A----5150 IEC INPUT PULSE I 

0 10 30 50 70 90 110 130 150 

TIME INANOSECONDSI 

Fig. 6. Model prediction, as percent of input amplitude, for various lengths 

of wires enclosed in metal conduit. 

line: the series inductance L; the shunt capacitance C; and the 

series resistance R. The shunt admittance, normally included 

in the model of a transmission line, is negligible in a wellinsulated power line. For the high-frequency components of 

the fast-transient pulse, skin effect influences the effective series impedance. The model includes an approximation of this 

effect. 

The 75-m line (250 ft) consisted of a neoprene-jacketed cord 

with three number 14 AWG (1.6-mm diameter) neopreneinsulated copper conductors. The line was suspended 2 m 

above ground. The line was arranged in a triangular loop 

with the sending end and the receiving end at one corner of 

the triangle. In this manner, either end could be connected 

to the oscilloscope. As described in Appendix I, a constant 

impedance load was maintained at the sending end by using a 

dummy load when switching the oscilloscope from the sending 

end to the receiving end. In all the measurements, two of the 

conductors were bonded at the sending end, representing the 

neutral and the grounding conductors. At the receiving end, 

the arriving pulse was recorded for various terminating load 

combinations. Fig. 7 shows an EFT signal with 2.5 V peak 

and 5150-11s waveform injected at the sending end, arriving as 

a 3011 10-ns signal at the end of the line. This peak amplitude 

of 48 percent is very close to the 50-percent value found for 

the steel conduit. The small difference between the two values 

is due to the oscillogram reading uncertainty rather than to 

a significant difference between the characteristics of the two 

lines. The question raised in the introduction, concerning radiation losses, is answered in the negative-or at least in the 

form, "not discernable at this level of significance." 

Need to Consider Loading at Receiving End 

Classical transmission line theory points out that reflections 

occur at the receiving end of the line when the terminating 

impedance differs from the characteristic impedance of the 

cable: 1) doubling the arriving pulse for an open end (and thus 

reflecting a pulse of the same polarity); 2) maintaining zero 

for a shorted end (and thus reflecting a pulse of reverse polarity); and 3) maintaining the level unchanged for a terminating 

Fig. 7. Pulses for nonmetallic jacketed cable. (a) Sending-end pulse. 

Vertical: 1 Vldiv, sweep: 20 nsldiv. Signal parameters: amplitude = 5 V 

(100 percent), rise time = 5 ns, FWHM = 50 ns. (b) Receiving-end pulse. 

Vertical: 0.5 Vldiv, sweep: 20 nsldiv. Signal parameters: amplitude = 2.4 

V (48 percent), rise time: 30 ns, FWHM: 110 ns. 

impedance that matches the line characteristic impedance (and 

thus having no reflection). 

Measurements made with "slow" 1.2150-ps impulses have 

shown that, while correct, this analysis does not produce useful results if applied to line lengths that do not contain the 

whole impulse front 141. For a 1.2-ps front in a line where 

the propagation velocity is approximately 213 the speed of 

light in vacuum (300 mlps), this lower limit on line length 

is approximately 200 m. For the 5150-ns IEC EFT pulse, the 

63-m line produces at the receiving end a 30-ns front time. 

According to the model, a 240-m line would produce a 60-ns 

front time (Fig. 6). Clearly, then, for this short an impulse, 

the receiving-end terminating impedance effects must be considered according to classical transmission line theory. 

Because the measurement system has an impedance of 

100 R while the characteristic impedances of the lines investigated are approximately 60 R, the case of an open-ended line 

cannot be measured. In actual installations and in the absence 

of a low-impedance instrument, the pulse would have twice 

the amplitude shown in Fig. 7. An open end is the limiting 

case for a very light load such as a control circuit during the 

off cycle of the power circuit. Adding even a small admittance 

at the end of the line substantially reduces the doubling effect. 

Common-Mode and Normal-Mode Coupling 

The method specified by IEC 801-4 for coupling the test 

transients to the power or the data cables introduces them as 

a common-mode signal in the cable (Fig. 3). This situation 

is not explicitly pointed out in the IEC document, but it must 

be clearly understood. Another standard advocating fast transients [5] specifies the application of the test transients in any 
MARTZLOFF AND LEEDY: ELECTRICAL FAST TRANSIENT TESTS 

Fig . 8. 

small 

Superimposed oscillograms of pulses at receivingdifference between overhead line (left) and coiled li~ 

OUTER NEOPRENE JACKn -\ 

INNER NEOPRENE 

INSULATION 

STRANDED COPPER 

CONDUCTOR 

FILLER MATERIAL - 

end, showing 1.4 - 1 

1e (right). Fig. 9. Section of three-conductor, nonmetallic jacketed cable. 

possible combination of the conductors for all input and output ports. While space and scope limitations for this paper of the 5150-ns EFT pulse as it propagates along a length of 

prevent a comprehensive discussion of the subject, alerting typical low-voltage power line. The low-voltage cable conreaders to the issue is worthwhile. sists of three number 14 (1.6-mm diameter) conductors, with 

The IEC EFT test focuses on installed equipment rather individual neoprene insulation and covered with a neoprene 

than on stand-alone equipment. The coupling clamp method jacket. Fig. 9 shows a cross section of this cable. 

for injecting the transients into the circuits is a very conve- A 63-m length of cable was simulated as 63 1-m lengths 

nient procedure when cables are accessible. Practitioners of connected in series (Fig. 10). Each 1-m length was represented 

this test also report useful results by coupling the transients by the equivalent twoport subcircuit shown in Fig. 11. The 

with foil wrapped around installed bundles of cables. The doc- subcircuit values represent the series and parallel elements of 

ument provides detailed descriptions of the configuration for the two-conductor line, to which a capacitance from each wire 

a laboratory test in an effort to ensure reproducible results. to a hypothetical ground plane was added. Numerical values 

However, propagation, coupling, and attenuation effects asso- were assigned to these parameters as described in Appendix 

ciated with the capacitive divider nature of the test application 11. 

will be strongly affected by small changes in the test circuit Electrical circuit analysis simulations were performed using 

geometry, making the test difficult to describe and reproduce the SABER software system' which allows the 1-m subcircuit 

in absolute terms. of Fig. 11 to be described only once as a "template." The 

These remarks should not be construed as a criticism of the software template is then replicated as many times as necessary 

test. Anecdotal field experience shows that equipment with to represent the transmission line. To investigate the effects of 

demonstrated immunity to the IEC 801-4 EFT procedure is a change of a parameter, such as a resistance Or 

capacitance, the parameter need be changed only once in the less likely to be disturbed than equipment not meeting this 

requirement. The addition of the new test procedure has con- data set. The 5150-11s IEC EFT pulse was digitized from an 

tributed to improved reliability. Experience is the ultimate Oscillogram recorded during the measurements 

so that a piecewise linear voltage source could be constructed idation of the usefulness of any performance standard. for the circuit simulation. 

No Practical Need for Outdoor Tests 

After completing the tests on the suspended aboveground 

line, the line was lowered to ground (dry grass) level, and one 

set of measurements was made under that condition. The same 

measurements were also made on a control coil of identical cable, left in its original packing carton. These two control measurements show detectable but insignificant differences 

(Fig. 8) in the pulse arriving at the receiving end. Thus concern over line geometry and coiled versus "free space" configurations, while technically correct, is a negligible factor in 

practical applications. Therefore, future experimenters need 

not go outdoors to perform measurements, even at the higher 

frequencies since the coiled versus free space configurations 

appear to produce similar results. This finding extends the 

conclusion previously reached for 100-kHz ring waves 161. 

Line Model 

An electrical circuit analysis model was devised to investigate the changes in the voltage amplitude and waveform shape 

Modeling Results 

Fig. 12 shows the comparison of the observed (solid line) 

and the simulated (dashed line) waveforms at the end of 

63 m of cable in response to the 5150-ns IEC EFT pulse. 

The pulse simulated by the simple model described demonstrates the same general waveshape, rate of voltage rise, and 

amplitude as the observed pulse. The simulated pulse, however, has a shorter time duration than the observed pulse. In 

addition, the propagation velocity of the pulse, as predicted 

by the model, is slightly greater than the propagation velocity 

measured in the cable. 

The discrepancies between the simulated and the observed 

waveforms and the velocities may be explained by several limitations of the model. First, the model did not consider the skin 

effect of the conductors. The skin effect would raise the resistance and inductance of the conductors for the high-frequency 

components of the waveform. Secondly, the cable consisted of 

I SABER is a product of Analogy, Inc., Beaverton, OR 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 26. NO. I, JANUARYIFEBRUARY 1990 

TERMINATION SOURCE 

- - 63 SECTIONS REPRESENTING 63m CABLE - 

t 

Fig. 10. Building line with repeated sections 

A pragmatic conclusion on the quantification of the EFT 

pulse propagation is that equipment likely to be installed close 

to sources of such disturbances would benefit from immunity 

demonstration; equipment intended to be located away from 

such sources, however, might not be required to be subjected 

to a complete set of EFT tests. 

g ACKNOWLEDGMENT 

P. F. Wilson (former staff member of the National Bureau 

Fig. 11. Circuit parameters for individual sections. of Standards, Boulder, CO) developed the model [3] for the 

conduit line, which is only summarized here. The nonmetallic jacket measurements were funded in part by a consortium 

of private sector companies coordinated by Building Industry 

Consulting Service International of Tampa, FL. 

APPENDIX I In 

5 

9 INSTRUMENTATION 

The voltages were recorded with a 1-GHz bandwidth oscilloscope and two 50-0 input vertical preamplifiers. The mainframe of the oscilloscope performed the differential combination of the two signals. In this configuration, the bandwidth of 

TIME (ns) the system is only 500 MHz, compared to the 1-GHz capabilFig. 12. Comparison between measured and modeled pulses at receiving ity the measurements with One preamplifier. 

end of nonmetallic iacketed line. Solid line: measured data; dashed line: This reduction in performance is an acceptable trade-off to 

predicted data. 

three conductors of which only two were modeled. The third 

conductor was connected to ground at the pulse generator end 

of the cable and left open at the receiving end of the cable. 

The presence of this third conductor would change the propagation characteristics of the cable slightly. Finally, a simple 

loss model (a series RC network) has been assumed for the capacitance associated with the neoprene dielectric of the cable. 

The response at high frequencies of organic dielectrics may 

be considerably more complicated than this model reflects. 

The two line configurations and their models yield similar 

results for measurements and for computations. Some difference might be expected between the quasi-coaxial configuration of a conduit-enclosed line and the open-line configuration 

with unshielded conductors of a nonmetallic jacket. 

The effects of propagation of an EFT pulse in the lines are 

a reduction in the amplitude, a decrease in the steepness of 

the front, and an increase in the duration of the pulse. The 

first two effects reduce the severity of a pulse arriving at the 

victim equipment; the third effect is less significant for victim 

equipment sensitive to rate-of-change disturbances. 

The two models provide the tools for analysis of other configurations. The choice of the geometry model or the electrical parameter model will depend on the available software 

and preferences of the investigator. 

obtain the differen;ial connection of the probes. In the experiments, the conduit, oscilloscope chassis, and low-side output 

of the generator are bonded together and hence ensure minimum noise coupling. Where high voltages are involved, this 

configuration also provides improved safety. This was not a 

concern here as the measurement on this linear system were 

conducted with signals in the 5-V range. A check on the common mode rejection of the system indicated less than fivepercent residual signal, an insignificant effect compared to 

the attenuation levels measured. 

The preamplifiers have a 50-0 input impedance, so that the 

effective impedance of the differentially connected probe is 

100 R. At the sending end, the voltages of the line and neutral 

conductors with respect to chassis ground were each fed to 

one of the preamplifiers (Fig. 13). Thus the pulse delivered 

by the generator impinges on an impedance consisting of the 

two amplifiers in series, in parallel with the line. To make 

measurements at the receiving end, the same oscilloscope was 

used with its differential probes. However, the pulse arriving 

at the receiving end would experience a reflection due to any 

impedance mismatch. To obtain an impedance match, timedomain reflectometry was used to trim the terminating load 

until a minimal reflection occurred. 

Good matching occurred with a 75-f2 termination for the 

steel conduit line (300-0 trim in parallel with the 100-0 

probes). For the nonmetallic-jacketed cable, a 50-0 termination gave best result (100-0 trim and 100-f2 probes in parallel). For the steel conduit line, a dummy load was used, made 
MARTZLOFF AND LEEDY: ELECTRICAL FAST TRANSIENT TESTS 

MATCHING 

PULSE 

GENERATOR 

RESISTOR 7 

OSCILLOSCOPE 

(b) 

Fig. 13. Configuration for constant-load measurements with dummy load 

and single oscilloscope. (a) Measurement at sending end. (b) Measurement 

at receiving end. 

,- SPECTRUM ANALYZER INPUT 

6 23m TEST LINE 

Fig. 14 

1 '- --i-i LBNc,,T,, 

COAXIAL-TO-BINSING 

POST ADAPTER 

COAXIAL CABLE 

- 40 MHz 

TRACKING GENERATOR OUTPUT 

. Configuration for determining velocity of propagation in nc 

lic jacketed line. 

of two 50-R coaxial terminators to provide the same matching impedance as that of the oscilloscope probes, as shown 

in Fig. 13. For the nonmetallic-jacketed cable measurements, 

two preamplifiers of the same type as the measuring preamplifiers, in a compatible mainframe, were used as the dummy 

load. In this manner, when recording the receiving-end waveforms with the measurement oscilloscope, the dummy load 

connected in parallel with the line at the sending end maintained a constant impedance load on the pulse generator. 

The propagation velocity in the cable was measured as reported previously: a 40-MHz spectrum analyzer, with a tracking generator output, was used to determine the frequencies 

at which the input impedance of a length of cable is minimum 

[6]. Fig. 14 shows schematically the test line, the spectrum 

analyzer, and a resistive attenuator to improve the decoupling 

of the tracking generator output from the test line. 

Fig. 15 shows the spectrum for a 6.23-m length of opencircuited test line. Voltage minima occur at 5.64, 19.4, 

and 30.4 MHz, corresponding to wavelengths of hl4, 3M4, 

0 10 2 0 3b 40 

FREQUENCY (MHz) 

Fig. 15. Spectrum of 6.23-111 length of test line. 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 26. NO. I. JANUARYIFEBRUARY 1990 

Fig. 16. Circuit parameters for line sections. 

and 5hl4, respectively. The average propagation velocity, the 

product of h and frequency, is 155 mlps. 

The series inductance of the 1-m section L, was calculated 

using the formula for the inductance of two parallel conductors in close proximity to each other, with current in opposite 

directions in each conductor [7]. For such a situation, the effective inductance L, is given in engineering handbooks as 

where 

Le 

1 

d 

r 

@ 

6 

effective inductance (nH), 

length of the cable (in), 

distance between the conductors (in), 

radius of the conductors (in), 

relative magnetic permeability (copper = 1.0), 

skin depth factor. 

The skin effect is a phenomenon affecting the conduction cf 

cables because current passing along the surface of the conductor encounters less inductance and resistance than current 

passing through the center of the conductor. The skin-depth 

factor 6 (over a range of 0 < 6 < 0.25) may be expressed as 

where 

K 

P 

f 

= 3168, a conversion constant, 

volume resistivity (0-in) ( = 0.68 . lop6 for copper), 

frequency (Hz). 

For a low-frequency approximation of the response of the cable to a 5150-ns pulse, the skin effect term can be neglected. 

There are two types of capacitance associated with the cable (Fig. 16). The first is associated with the conductor-toconductor capacitance, C, . The other type of capacitance consists of equd valued capacitors C, I and Cg2, which represent 

the two conductor-to-ground capacitances. The equation of 

the capacitance between two parallel cylinders representing 

the conductor-to-conductor capacitance may be written as [8] 

where C, is the capacitance per unit length (pF11000 ft) and 

E, is the relative dielectric constant of the material filling the 

space between the two conductors. The equation for the capacitance between a wire and a ground plane is similar; this 

capacitance is 

where C, is the capacitance per unit length (pFI1000 ft) and 

His the distance between the conductor and the ground plane. 

The resistance of the conductor is represented by resistance 

R, in series with each inductance L,. A resistance, R, is 

in series with the capacitance C, to simulate the losses encountered in the dielectric surrounding the two conductors. 

One additional resistance R, simulates the resistance associated with the conductor-to-conductor leakage to complete the 

model. 

Using the equations given, the following values were computed for a 1-m length of number 1413 Type SO cable: 

L, 261 nH, the series inductance of each conductor where 

d = 0.45 cm (0.177 in) and r = 0.016 cm (0.032 in); 

C, 53 pF, the conductor-to-conductor capacitance assumes er = 3 for neoprene dielectric; 

C, 17 pF, the conductor-to-ground capacitance; 

R, 0.003 R, the series resistance of conductor (obtained 

from wire tables); 

R, 10 R, the estimated resistance in series with C,; 

R, 10 MR, the estimated leakage resistance. 

The velocity of propagation predicted by the model is approximately 170 mlps for the 5150-ns pulse traversing the 

neoprene-jacketed cable. In contrast, the higher velocities of 

propagation would be expected in the line enclosed by the 

steel conduit where the conductors have only a thin insulation 

surrounded by air. Measurements of the velocity of propagation confirm the difference between the two lines: 210 mlps 

for the air-surrounded conductors in the conduit and 155 mlps 

for the neoprene covered conductors. 
MARTZLOFF AND LEEDY: ELECTRICAL FAST TRANSIENT TESTS 159 

REFERENCES tions, and interference mitigation. In addition to his research work on surge 

,electrical F~~~ ~~~~~i~~~~~~i~~~~IEC publication 801-4, propagation and mitigation, he is contributing to the development of stanGeneva, Switzerland, 1988. dards on surge environment and surge suppression within IEEE, ANSI, and 

R. B. Bush, "Statistical considerations of electrostatic discharge evalu- IEC by documenting his measurements, as in the present paper, by resent in^ 

ations," in Proc. Int. Zurich Symp. Electromagnetic Compatibility, and by new standards. 

1987. 

F. D. Martzloff and P. F. Wilson, "Fast transient tests: Trivial or terminal pursuit?" in Proc. Int. Zurich Symp. Electromagnetic Compatibility, 1987. 

F. D. Martzloff, "The propagation and attenuation of surge voltages 

and surge currents in low-voltage ac power circuits," IEEE Trans. 

Power App. Syst., vol. PAS-102, no. 5, May 1983. 

Standard Surge Withstand Capability (SWC) Tests for Protective 

Relays and Relay Systems, ANSIIIEEE C37.9O.l. 1989. 

F. D. Martzloff and H. A. Gauper. "Surge and high-frequency propagation in industrial power lines," IEEE Trans. Ind. Appl., vol. IA-22, 

no. 4, JulyIAug. 1986. 

P. M. Rostek, "Avoid wiring-inductance problems," Electron. Design, Dec. 6, 1974. 

H. Pender and W. A. Del Mar, Electrical Engineers' Handbook. 

New York: Wiley (Engineering Handbook Series), 1949, p. 14-33. 

Franqois D. Martzloff (M'56-SM'80-F'83) completed his undergraduate studies in France and received the M.S.E.E. degree from the Georgia Institute of Technology, Atlanta, and the M.S.I.A. degree from Union College, Schenectady, NY. 

After a long career at General Electric, he joined 

the staff of the National Bureau of Standards (now 

renamed National Institute of Standards and Technology) to work on conducted electromagnetic interference issues. His early experience covered highvoltage fuses, high-voltage bushings, and electronic 

power conversion; the latter marked a turning point to issues of overvoltage 

effects on semiconductors, surge suppression, metal oxide varistor applicaThomas F. Leedy (S'64-M'65) received the B.S. 

degree in physics from the American University, 

Washington, DC, and the M.S. degree in electrical 

engineering and the E.Ad. degree in engineering 

administration from George Washington University, 

Washington, DC, in 1965, 1972, and 1984, respectively. 

From 1965 to 1980 he did research in microelectronics at the National Bureau of Standards where 

he developed and validated measurement methods 

for determining the effects of transient high-energy 

radiation on integrated circuits and for evaluating materials and processes of 

microelectronic device fabrication. Since 1980, he has been a member of the 

technical staff of the Electronic Instrumentation and Metrology Group of the 

Electrosystems Division where he has participated in the research program 

directed toward assuring the accuracy of automatic test equipment. 

Mr. Leedy has served on numerous standards committees in microelectronics, computers, radiation effects, and electronic instrumentation for IEEE, 

ASTM, and the U.S. Department of Commerce International Trade Administration 
LOW-VOLTAGE

CABLES
Significance 

Part 4 – Propagation and coupling of surges 

This paper is only indirectly related to the topic of surge propagation in the context of surgeprotective devices, but is included here as it might be a gateway to further information on puls

propagation – the basic technique used in the investigation summarized by this paper 

Filename: PD in cables 
Conference Record of the 1990 IEEE International Symposium on Electrical Insulation, Toronto, Canada, June 3-6, 1990 

PARTIAL DISCHARGES IN LOW-VOLTAGE CABLES 

J.P. Steiner 

Biddle Instruments 

Blue Bell, PA 19422 

ABSTRACT-Testing of high voltage apparatus for partial discharges has long been recognized as an important part of quality 

control for these devices. Recently, interest has been focused on 

methods for testing low voltage cables to determine their integrity 

under adverse operating conditions such as a loss of coolant accident. A new method, utilizing partial discharges, is presented 

which has the potential for locating breaches in the insulation of 

in situ, low voltage, multi-conductor cables. 

INTRODUCTION 

Cables used in power plants are selected on the basis of qualification tests [l] which provide a sufficient degree of confidence 

in the ability of the cables to maintain operational readiness for 

a well-defined "life" duration. As the cables age under the various environmental stresses prevailing in a power plant, nuclear 

plants in particular, the question arises of how much "residual 

life" exists in the cables as they approach their rated, qualified 

life. There is considerable interest in assessing this readiness, and 

a number of investigators have proposed various test methods 121. 

In particular, a method that could allow nondestructive, in situ 

assessment would be extremely valuable. The detection of partial discharges at identifiable sites along a cable offers a method 

for assessing the likelihood of a fault to occur during a loss of 

coolant accident on a cable containing an incipient defect. Such 

an incipient defect might be a breach of the insulation that would 

remain undetected under normal operating conditions. 

Low-voltage cables, in contrast with cables designed for highvoltage service, arc not provided with an insulation structure 

aimed at making them free from partial discharges under modelate overvoltage conditions such as the application of a test voltage. Therefore, the occurrence of a partial discharge at an incipient defect site has to be differentiated from the background of the 

expected partial discharges that will occur over the length of the 

cable in the absence of any significant defect. The test method 

described in this paper offers the opportunity to make this differentiation and locate the discharge sites associated with insulation 

breaches. Three representative types of cables in which artificial 

defects had been created were subjected to the test; every defect 

was successfully located. 

PRINCIPLE OF THE METHOD 

Partial discharges will occur along the entire length of a cable 

when voltages greater than the partial discharge inception voltage 

are applied. To develop a strategy for analyzing experimental 

data from a cable that also contains defects resulting in partial 

discharges at specific sites, several assumptions need to be made. 

The first is that there are no preferred partial discharge sites in 

the cable under test. Therefore, it would be expected that, ideally 

and in the absence of any defect, the partial discharges would be 

uniformly distributed along the length of the cable. 

F.D. Martzloff 

National Institute 

of Standards and Technology 

Gaithersburg, MD 20899 

The second assumption is that, in the absence of a defect, the 

partial discharge pulses occur as random events in time, i.e., the 

observed process is a marked random point process [3]. Using 

these assumptions causes the observed partial discharge process 

to resemble, asymptotically, a Poisson shot noise process [3,4], 

with pulses emanating from all positions along the cable. If a 

sample of data were analyzed by calculating a histogram of the 

positions from which each pulse in the sample emanated, then, 

using this model, the histogram would correspond to a uniform 

distribution. 

The third assumption is that amplitudes of the pulses, expressed 

in units of charge at each partial discharge site, are gammadistributed [4] with the parameters of the probability density 

being independently and identically distributed. The observed 

probability density of the charge will then have the form 

where p(q) is the probability density of the observed charge in 

the entire cable, p(qjz, w) is a conditional probability density of 

the charge at a single site, p(z) is the probability density of the 

locations of the partial discharge sites along the cable and p(w) 

is the probability density of the parameter w, such as the degree of imperfection in the insulation (local field enhancements) 

that control the most probable discharge amplitude q at that 

site. By assu~nption, there is no preferred site and experimental evidence indicates that the physic~l behavior at each site is 

similar. This suggests that the superposition of these infinitesimal processes will yield a seemingly well behaved, predominantly 

unimodal probability density for the charge and this has been 

observed experimentally. 

Now consider the modcl for a defect. The goal of this measurement technique is to detect an insulation flaw which is dcfincd, 

for these purposes, as a coniplete crack through the insulation to 

the conductor. Partial discharge in electrode arrangements that 

consist of a conductor-conductor interface or conductor-dielectric 

interface is often more active than when the electrodes are a 

dielectric-dielectric interface. The implication is that the partialdischarge sites at the flaws will be more active than the sites in 

the rest of the cable and consequently cause significant deviations 

from the ideal model. 

The basic idea of the measurement technique is to identify whether 

the data significantly deviate from t,hc behavior defined by cquation (1). In this case, equation (1) is modifird by adding a term 

in which p(w) is replaced by p(wlz) 

where w depends on z and is highly localized to particular points 

z,. Admittedly, thc model is a gross idcalization of the cxprcted 

U.S. Government work not protected by U.S. copyright 
Figure 1. Block diagram of partial discharge detection and 

recording. 

behavior, however, the general concept has merit if the defects do 

behave differently from the normal background partial discharge 

in the cable. In practice, there will be naturally occurring, preferred partial discharge sites that are not flaws but cause variations in the expected background count rates. In fact two such 

preferred locations have already been identified, the ends of the 

cable. 

EXPERIMENTAL APPARATUS 

The experimental apparatus implemented is a general purpose 

system designed for acquiring broadband partial discharge waveforms and processing the acquired signals as shown in the block 

diagram of Figure 1. The key elements are the high-voltage system, the coupling circuit and the data acquisition and analysis 

system. The high-voltage system used to apply the necessary 

voltages to the cables under test consists of a 60 Hz high-voltage 

transformer with appropriate noise filtering. Measurements on 

the cables are made at voltage levels exceeding the partial discharge inception voltage by 10, 20, and 40 percent. The inception 

level is defined as that voltage which causes at least one partial 

discharge pulse above a specified charge level q per second. 

The partial discharge signals are coupled from the cable under 

test to the measurement system through a coaxially mounted, 

1 nF capacitor. The capacitor is the first element in a high-pass 

filter having a lower cutoff frequency of 30 kHz and a 1000 fl input impedance. Initial testing with sensitivities ranging down to 

0.05pC did not produce useful results because the partial discharge magnitudes at inception are in excess of 10 pC. Since the 

predominant partial discharge levels are greater than 10 pC, an 

attenuator and input protection are included in the circuit which 

limits the input sensitivity to 1 PC. The capacitance of the input 

protection limits the bandwidth of the system to approximately 

10 MHz. This input bandwidth is adequate for these experiments 

since the cables under test are lossy and limit the usable bandwidth to within this range. A variable attenuator is included 

between the input preamplifier and the gain block of the system 

to provide amplitude scaling which prevents saturation of the 

amplifier for larger partial discharge amplitudes. 

Partial discharge pulses are recorded using a LeCroy 9400A1 digital oscilloscope which has an 8-bit digitizer that can sample 32 000 

'The identification of commercial materials and their sources is made to 

describe the experiment adequately. In no case does this identification imply 

rcommendation by the National Institute of Standards and Technology, nor 

does it imply that the instrument is the best available. 

points at a 1OOMHz rate. The advantage of using this oscilloscope 

is its ability to segment its 32 kB acquisition memory so that it 

can acquire up to 250 separate waveforms at its maximum sampling rate before having to transfer them to a computer over an 

IEEE 488 interface. The computer is a PC-AT class computer. 

The ASYST' scientific software package is used for acquisition 

control and data analysis. Up to 20000 waveforms can be collected before the available disk space limits the computer's ability to process the data; typically 5000 waveforms are collected. 

Active data collection occurrs only during a few seconds of the 

voltage application to the cable under test, however, voltage is 

typically applied for 8 minutes since the high voltage system is 

not under control of the computer and time is needed to archive 

the data. Since the system uses a computer with only moderate 

computational ability, all processing has to be performed off line. 

MEASUREMENTS 

Two fundamental estimates are made from each recorded waveform: an estimate of the charge, and an estimate of the location 

of the partial discharge site in the cable. It would be desirable to 

calculate estimates for all detectable pulses; however, since the 

measurement system is general purpose, trade-offs for cost, speed, 

and data handling capabilities have been made. With these constraints, only a limited number of waveforms can be collected and 

analyzed. 

The method used to calculate the charge in the partial discharge 

waveform is the same as that used in [4] and has been shown to be 

more accurate than the usual measurement of the peak value of 

the pulse. The technique uses the pulse energy derived from the 

waveform to develop an estimate of the charge and assumes that 

the measured partial discharge waveform is similar to calibration 

pulses injected into the cable. The calibration pulses are generated by applying a fast-rise step of known amplitude through a 

known capacitance into the cable under test. Further corrections 

in the estimates to account for the cable attenuation were not 

deemed necessary. 

If the flaw is detectable, then the histogram of the charge will 

no longer be unimodal. Using this fact, the measurement can be 

enhanced by setting a lower discrimination level so that the instrument measures only those partial discharges which appear to 

be anomalous. The lower discrimination level is also referred to as 

the trigger level, the level which the partial discharge pulse must 

exceed to initiate the recording cycle of the digital oscilloscope. 

The basic principle used to determine the location of the partial 

discharge can be understood by referring to Figure 2. When a 

partial discharge pulse originates at some point interior to a cable 

it splits evenly into two plses which travel in opposite directions. 

One pulse, referred to as the direct pulse, proceeds toward the 

detector end arriving at time tl. The other pulse proceeds toward 

the open end of the cable and is reflected from the open circuit 

back toward the detector end, arriving at the detector at time 

t2. Knowing the velocity of propagation, v, for the cable, the 

location, e, is given by 

where the location is referenced from the open circuit end of the 

cable. 

To save time in processing, a simple but accurate method is used 

to calculate the locations. The technique applies an approximate 

matched filter to the recorded pulses whose shape resembles a 
DISCHARGE 

DETECTOR SITE 

EN0 -A- OPEN 

EN0 

Figure 2. Position versus time diagram of direct pulse and reflected pulse propagation. 

partial discharge pulse. In effect, this assumed pulse is crosscorrelated with the pulses in the measured data. The time delay 

between the two largest pulses in the waveform is determined 

and converted into a location using a measured propagation velocity. More accurate delay estimators are not justified since the 

increased computational effort does not provide enough increase 

in accuracy. When the two largest pulses in the data record are 

from two distinct discharge sites, the estimate is referred to as a 

phantom delay. The probability of calculating a phantom delay 

is small at the voltage levels used, with less than 1 percent of the 

estimates being phantoms. These phantoms are not a problem 

since the time delay between the pulses is random and does not 

lead to a count rate that favors any particular site. A limitation 

of this technique is the existence of a measurement blind spot in 

the cable: when the arrival times of the direct pulse and reflected 

pulse are nearly the same, it becomes difficult to differentiate 

between a single pulse and two closely spaced pulses. Without 

special, computationally intensive processing, these sites all appear to be at a distance range near the end of the cable, in effect 

creating a blind spot. 

TEST RESULTS 

In this case, the measurement technique is intended for use in 

situ on 600V class nuclear power plant cables and so representative cable types were used in the experiments. These cables 

can be grouped into three categories: shielded cables, unshielded 

multiconductor cables, and unshielded single conductor cables. 

The following samples were used in a series of experiments: 

An unaged, unshielded, two conductor, #14AWG, 600V 

class, neoprene jacketed cable of the same type used for 

control cabling in nuclear power plants. 

An unaged, 50 ohm, coaxial cable of the same type used for 

instrumentation cabling in nuclear power plants. 

0 An aged, unshielded, 15 conductor, #18 AWG cable of the 

same type used for instrumentation cabling in nuclear power 

plants. 

No experiments were conducted on unshielded single conductor 

cables since the techniques to be discussed require that the cable 

have some transmission line properties; a single conductor, in the 

absence of ground plane, does not. 

Vertical: 

Number of pairs 250, 

counted per bin of 

the histogram I 

150. 

Horizontal: 

Distance in feet 50.0 

from the open end 

Note the scattering of counts over the length of the cable 

Figure 3. Histogram of delay times between the two major 

pulses in each event of a 5000-element record containing true 

pairs and random pairs. 

Positive PD 

Vertical: 

Numberof 24.9 

pulses per 

bin of the 8.99 

histogram 79'.9 ' 21'9. ' 35'9. ' 49b. ' 630.. * 

Horizontal: 

Pulse charge ii:i [ egative PD 

in pC 

4.. 9 

Figure 4. Histogram of the distribution of partial discharge 

pulses (PD) as a function of their charge transfer magnitude 

( "pulse height"). 

Tests were performed on 100-m (330ft) length of unaged, twoconductor cable with artificial defects introduced at three locations. A small portion of the jacket was removed at three locations and the insulation was carefully pierced through to the conductor using a #60 twist drill in two of the locations; a small razor 

knife was used to slit the wire insulation in the third. In neither 

case was any of the insulation removed; only a break was created 

from the conductor surface to the surrounding atmosphere. The 

partial discharge inception voltage of the cable was 4500V and 

measurements were made at voltages levels 10, 20, and 40 percent above inception. With a small lower discrimination level, 

the low-level background partial discharges have dominant count 

rates indicating that the partial discharges are distributed along 

the length of the cable as shown in Figure 3. If the lower discrimination level is increased, then the count rates of the partial 

discharges from the defect sites dominate the background count 

rate as indicated by the charge histograms of figure 4; note that 

the histograms are not unimodal. The locations of the individual damage sites were found using those delay values which 

corresponded to data near the peaks of the charge histograms. 

Each of the damage sites was correctly identified as can be seen 

in figures 5, 6 and 7. Other measurements were made on a 

20m (66ft) length of coaxial cable with artificial damage inflicted 

by slightly abrading the shield. The partial discharge inception 

voltage was 5200V. In this case, the defect was immediately 
Vertical: 

Number of pairs 218. 1 )( counted per bin of 

the histogram 

12~. t II A Horizontal: 

Distance in feet 48.8 -- 

from the open end 

4.8 10 200. 28b. t 

Note the emergence of a defect signal at 70.2 feet , location of an actual defect. 

Figure 5. Histogram of time delays for 5000 V and pruned data. 

Vertical: 

Number of pairs 

counted per bin of 

the histogram lZea 

Horizontal: 

Distance in feet 

from the open end 

Note the emergence of a defect signal at 34 feet (actual defect), 

with minor indications at 70 feet (already detected in Figure 5), 

and at 178 feet (no known defect at that point). 

Figure 6. Histogram of time delays at 5200V and pruned data. 

Note the emergence of a defect signal at 240 feet (actual defect), 

with substantial signal at 70 feet (already detected in Figure 6), 

small signal at 34feet (as in Figure 5), and new signal at 280 feet. 

Figure 7. Histogram of time delays at 6000 V and pruned data. 

Vertical: IZI. 

Number of pairs 

counted per bin of 

the histogram ,I.e -- 

Horizontal: 

Distance in feet 

from the open end 

identified because the partial discharge activity at the damage 

site was much greater than the background partial discharges. 

Similar results were also obtained while measuring a 12-m (40 ft) 

length of aged multiconductor cable containing an artificial defect. The artificial defect was created by piercing the insulation 

on one of the 15 conductors with a small razor knife, similar to 

the slit described above. The partial discharge inception voltage 

was 2300V. Similar to the previous case, the partial discharge 

activity at the damage site was greater than the background of 

the rest of the cable allowing easy identification of the flaw. 

CONCLUSIONS 

3w.a ' 3d.e ' ~sb. , ztb. ' 27b. ' 

The test method based on partial discharge detection using signal 

processing makes possible the identification and location of insulation breaches, undetectable under normal service conditions, 

that may become a fault under additional environmental stresses, 

or may progress into a fault under further aging. 

#. 

L 

-- 

Tests performed on three representative types of low-voltage cables have demonstrated that the method can differentiate between signals emanating at incipient defect sites and the background of partial discharges occurring over the length of the cable. 

-- 

In order to induce partial discharges, the test voltage must clearly 

be raised above the inception voltage. Depending on the cable 

structure, this test voltage will be on the order of a few thousand 

volts, a subject of possible concern, but not a fatal limitation 

for a test method offering the reward of definite identification of 

defects that could become fault locations. 

At\,, , IS 

The demonstration was obtained with a general-purpose PC-type 

computer, using a commercial software package augmented by 

moderately complex custom-designed software. In a more dedicated system, with further software enhancement, additional benefits are possible, such as the detection of incipient defects with 

only brief exposure of the insulation to the test voltage, automated scanning of the data for most significant content, and optimization of the test procedure into an expert system approach. 

ACKNOWLEDGEMENTS 

The authors acknowledge the contributions of S. Clevenger and 

D.R. Flach in the development and implementation of the menudriven software used in the application of the principle of the 

method. The signal processing concepts were developed by J.P. 

Steiner as described in Reference 4 and implemented prior to 

his new position with Biddle Instruments. The implementation 

and demonstration tests were performed by J.P. Steiner and F.D. 

Martzloff with funding provided by the USNRC Nuclear Plant 

Aging Research Program. 

REFERENCES 

[I] IEEE Std 383-1974 Standard for Type Test of Class 1E Electrical Cables, Field Splices, and Connection for Nuclear Power Generating Stations. 

(21 Proceedings: Workshop on Power Plant Cable Condition Monitoring 

EPRI EL/NP/CS-5914-5R, July 1988. 

[3] Snyder, D.L., Random Point Processes, John Wiley and Sons, 

New York NY, 1975. 

[4] Steiner, J.P. "Digital Measurement of Partial Dischargen, Ph.D. Thesis, 

Purdue University, W. Lafayette IN, May 1988. 

[5] Steiner, J.P., Weeks, W.L. and Furgason, E.S., =Digital Estimation of 

Partial Dischargen, 1987 Annual Report, Conference on Electrical Insulation and Dielectric Phenomena, Gaithersburg MD, October 1987. 
SURGES TO

GROUND:

EXPECTATIONS

VERSUS

REALITY
, , p, y, p

down the drain of a grounding conductor only makes them reappear within a microsecond about 200 meters 

some other conductor. 

This paper presents a brief review of some of the fallacies, with illustrative measurement results, and propose

approaches for remedy, rather than counterproductive grounding practices based on misconceptions. 

Filename: Divert surges 
125
126
127
128
129
130
131
132
ON THE

PROPAGATION

OF

OLD

AND

NEW

SURGES
Reprinted from Proceedings, Open Forum on Surge Protection Application, NISTIR-4654, August 1991 

Significance

Part 4 – Propagation and coupling of surges 

The 1991 revision of the IEEE Recommended Practice on Surge Voltages C62.41 introduced a new 

generation of surge waveforms; how they travel in low-voltage power systems will affect some of the 

earlier tenets on surge propagation characteristics. The emergence of cascaded surge-protective devices 

also raised a new set of concerns in which propagation characteristics play an important role, where 

system designers rely on the inherent impedance of the wiring between the two devices to provide the 

electrical separation necessary to obtain coordination. 

During the development of the revised IEEE Recommended Practice in the late eighties, some reluctance 

was encountered in deleting the mention of wire diameter for the branch circuits. The wire size was 

included in the definition of the 'Location Categories' given in the 1980 version of the IEEE Std 587 Guide 

The paper presents a review the propagation characteristics of the old and the new generation of surges 

waveforms encountered in low-voltage ac power systems. To complement information developed on this 

subject over the last ten years, measurements results are reported for the new 10/1000 μs waveform, and 

the effect (or, rather, the lack of significant effect) of wire diameter is documented by a simple experimental 

demonstration. 

Filename: Propagation 1991
19
20
21
22
23
24
25
SURGING

THE

 UPSIDE-DOWN

HOUSE:

LOOKING

INTO

UPSETTING

REFERENCES
Electronic equipment with two input ports  power and communications  can be exposed to damaging 

differences of voltage across the two ports during surge events. Two exposure scenarios of producing suc

differences of voltages are explained and illustrated by measurements performed in a replica of a residenti

or light commercial installation of power, telephone, and cable TV wiring. 

Several mitigation methods are described, and one possible retrofit solution is shown. In a later paper, (se

the pdf file “ Upsdown measure”) numerical simulations were performed on a model of the system in order 

expand the range of conditions and identify significant variables. Nevertheless, there are still very few 

published data on quantifying the stress that can be produced by these scenarios, and hopefully mitigated 

by “surge reference equalizers” -- also known as “multi-port surge protectors.” 

Filename: Upsdown surging 
Thomas S. Key and Frawois 0. Martzloff ' 

Abstract - Electrvnb equipment with two input ports 

- power and communications - can be exposed to 

damaging differences of voltage across the two 

ports during surge events. Two exposure scenarios 

of producing such differences of voltages are 

explained and illustrated by measurements 

performed in a replica of a residential or light 

commercial installatian of power. telephone, and 

cable N wiring., Several mitigation methods are 

described, and one possible retrofit solution is 

shown. It is planned that in a further phase of this 

research, numerical simulation will be performed on 

a model of the system in order to expand the range 

of conditions and identify significant variables. 

INTRODUCTION 

As more and more electronic equipment enter 

the home and business environment. these often 

invoke a communications port as well as their usual 

power cord port. In this paper, we will use the term 

'twoport appliance' or "appliance' for short. being 

understood that it covers two-port information 

technology equipment. Examples of such two-port 

appliances include fax machines, telephone 

answering machines, personal computers with 

modem communications or printer connections. and 

cableconnected TV receivers. Although each of the 

power and communications systems may include a 

scheme for protection against surges, the surge 

current flowing in the surged system causes a shift 

in the voltage of its reference point while the other. 

non-surged system reference point remains 

unchanged. The difference of voltage between the 

two reference points appears across the two ports 

of one appliance, or between the communications 

ports of two appliances linked by a data cable. 

Depending-on the nature of the appliance and its 

immunity, which is not often defined. this difference 

of voltage may have some upsetting or damaging 

consequences. In this paper,-we will present just 

two examples of measurements illustrating the broad 

variety of possible exposure scenarios. 

To identify and quantify the significant variables 

and their effects, a representative configuration of 

the circuitry in a residence (metallic cold water pipe. 

power and grounding conductors. telephone and 

coaxial cable N wiring) has been set up in the 

laboratory, according to US practice. The circuits 

have been hung from the laboratory ceiling, to decouple them from nearby metallic masses and get 

them out of the way of laboratory personnel. hence 

the name 'Upside-Down House' given to the project

To evaluate the threat of impinging surges in an 

actual installation. surges of various types, as 

defined in standards covering AC power circuits and 

communications, can be injected at various points 

of the Upside-Down House circuits. Combinations 

of surgeprotective devices (SPDs) can also be 

placed at various locations of the Upside-Down 

House. corresponding to a variety of real-world 

exposure scenarios. A measurement can then be 

made of the resulting differences of voltage 

appearing between the power and communications 

ports of a single appliance. or between the 

communications ports of two appliances installed at 

some location within the Upside-Down House. No 

conclusions are drawn in this paper on the withstand 

capability of any particular appliance for this type of 

threat, because the manufacturers typically do not 

provide immunity data for any exposure scenario of 

this type of interaction. However, some of the 

voltages thus recorded in the Upside-Down House 

confirm the suspicion derived from field failures that 

damaging differences of voltages can occur. 

APPROACHES 

Various mitigation schemes have been proposed 

by researchers and industry. but not quantified. to 

remedy upsetting or damaging voltage differences. 

The most effective is likely to be a fiber optic 

decoupling inserted in the communications link. bu

the expense may be objectionable for residentia

Thomas Key is with the EPRl Power Electronics Applications Center (PEAC), 1 O52l Research Drive, 

Knoxville TN 37932, USA; Franpois Martzloff is with the National Institute of Standards and Technology 

(NIST), Gaithersburg MD 20899. Contributions from NlST are not subject to copyright in the U.S.A. . 
offer some degree of remedy, but leaves out all 

existing installations. Increasing the appliance 

withstand capability. may raise objections of market 

economics. and may not be practical for some of the 

voltages that can appear. 

Many different exposure scenarios can be 

identified. even in a simple residential circuit replica 

Reference voltage shiis are a multi-dimensional 

problem in the real world. In this paper, the problem 

has been simplified to looking at the effect of only 

twovariables: spatial relationships of conductors and 

effectiveness (including some side-effects) of SPDs. 

Other important variables that were identified but not 

addressed at this stage of the research are cited in 

the discussion section of this paper. 

In this paper, two simple exposure scenarios are 

illustrated: a voltage difference occurs between the 

ports of an appliance connected to two systems 

when a surge impinges on one of the systems, and 

a voltage difference occurs between the communications ports of two appliances powered by separate 

circuits of the same power system when a surge 

impinges on that system. In the final discussion. we 

will look into some ways of expanding these results 

and perhaps identifying a recurring set of variables 

that can be mitigated or avoided. 

SETUP AND MEASUREMENTS 

Measurements reported in this paper describe 

exposure scenarios leading to voltage differences 

being developed during surge events across the 

power and telephone systems of the Upside-Down 

House. as well as between the signal reference 

points of two interconnected appliances linked by 

their communications ports, such as a personal 

computer (PC) and associated printer. For each 

case, the Upside-Down House circuits may include 

some form of upstream surge protection on the 

telephone service entrance or appliance port, as well 

as on the power service entrance or appliance port. 

It is planned to continue the project with similar 

measurements involving the cable TV port. 

Figure 1 shows an isometric of the Upside-Down 

House configuration. with the arrangement of the 

three tiers of conductors shown in Figure 2. The 

power wiring includes two tiers of 3conductor cable 

(2.05 mm dia - #12 AWG, non-metallic jacket). 

~i~urel- Schematic representation of the 

Upside-Down House conductors 

typical of residential wiring, and one tier of three 

2.05-rnrn dia. conductors in a steel conduit, typica

of commercial or office installation. A 4conductor

two-pair telephone cable and a 70-R N coaxia

cable also run along the 3conductor power cables

To illustrate the expected benefit from goo

wiring practice (cables routed close to the eart

reference -- the copper water pipe in the Upside

Down House). one tier has been lashed to the 

copper pipe. Of course. such idealized practice i

not practical. but will serve here as baseline an

illustration of EMC principles [Van Deursen. 19931

In an actual installation, the system would exist in a

three dimensions. For the sake of simplification. the 

Upside-Down House has been reduced to only two 

dimensions, one horizontal run spanning the house

and the vertical separation indicated in Figure 1

For the purpose of accessing both ends when injec

ting surges and measuring voltages and currents

the horizontal span has been folded into a hairpi

with both ends accessible in junction box JB 14. 

3 wnductors. NM jacket 2#12 +G - 

Telephone pairs 

70Qa~xcaMe 

3 wnductors. NM jacket 2#12 +G 

Telephone pairs 

70 Q coax cable 

2 conductors. bose. Y12 AWG 

Figure 2 - Vertical arrangement of conductors 

in the Upside-Down House (X-X of Figure 1) 
mediate point of the span. Short cable runs (3 m), 

not shown on the diagram, provide for appliances 

located dose to the senrice entrance. 

Neglecting the vertical separation of the three 

tiers. the length of the span from end to end is 

36 m. (This number is cled to give an idea of the 

size of the house. Any numerical computations will, 

of course, use the exact values.) A typical service 

entrance breaker panel and revenue meter have 

been provided at one end. upstream of junction box 

JB 1-4. A Network Interface Device (NID). typical of 

the US. practice for entry of the telephone service. 

has been installed next to junction box JB 1-4. 

By connecting the NID grounding conductor 

(US. code terminology) to one or the other end of 

the copper pipe. it is possible to represent the 

scenario where telephone and power service enter 

at the same end of the house (the preferred 

practice) or at opposite ends (not preferred, but 

often encountered). All of the conductors are 

insulated from the existing earthing arrangement of 

the laboratory building, making it possible to 

represent various configurations of the earthing 

arrangement of the Upside-Down House. 

Surges were injected into the power system in 

the line-to-ground (L-G) mode. Note that the U.S. 

practice of bonding the neutral and grounding 

conductors at the service entrance makes any 

impinging L-G surge become also a line-to-neutral 

surge. Surges injected into the balanced-pair 

telephone system were in [tip & ring]-to-ground 

mode, with the NID acting to divert them to the 

common earthing point of the laboratory building and 

Upside-Down House via the copper pipe. 

The waveform and amplitude of the injected 

surges were selected to harmonize with the values 

cited in industry standards. Because of the different 

values of the impedance of the various circuits into 

which -the surges were injected. the resulting 

waveforms reflect the interaction of the surge 

generator and load impedances and do not exactly 

duplicate the familiar standard waveforms. Nevertheless. the resulting waveforms are representative 

and provide examples of the threat and needs of 

mitigation. These results will provide experimental 

data for later validation of computer modeling, so 

that the modeling can then expand the results to 

other waveforms and circuit impedances. 

TWO SYSTEMS SERVING ONE APPLIANCE 

In this exposure scenario, a modern-equippe

PC is connected by its power port to a branc

circuit. and by its modem port to the telephon

service of the house. For a worstcase scenario. th

power and telephone services enter the house 

opposite ends (Figure 3). 

Figure 3 - Power and telephone services 

entering the house at opposite ends, with PC 

connected across the two systems 

An open loop is formed by the copper pipe. th

protective conductor (international symbol "PF) 

the branch circuit feeding the PC. and the telephon

wires from the NID to the PC. If a surge impinge

on the external telephone plant, it is diverted by th

NID via the copper pipe to the common earthin

point of the house, at the power service entrance

The surge current in the copper pipe creates 

changing magnetic flux around the pipe. whic

induces a voltage in the loop. This voltage w

appear between the two PC ports if they ar

separated by a high impedance (of unknown surg

voltage withstand capability). 

With the telephone wires routed away from th

copper pipe -- which can be expected in residenti

wiring -- a large loop is formed, embracing the flu

produced by the surge current flowing in the copp

pipe. Wflh the telephone wires lashed to the coppe

pipe -- a theoretical more than practical routing 

the loop embraces less flux and one can expect 

lower induced voltage across the two ports. 

Figure 4 shows the recording obtained with th

telephone wire routed away from the pipe. For 

rate of change in the surge current of 75 Nps, 

peak of 4.3 kV is induced in the loop and appea

between the two ports. For the same injectipn 

current. a peak of only 1.3 kV was noted with th

telephone wires lashed to the copper pipe. 
Figure 4 - Voltage difference recorded 

with telephone and power services 

entering at opposite ends of the house 

A relatively simple retrofit solution is to equalize 

the difference of voltage between the two systems 

by a device designed for the purpose and inserted 

in both communications and power links just before 

they enter the appliance. This device, defined in 

IEEE standards [IEEE Std 1100-19921 as a "Surge 

Reference EqualizeF is commercially available in 

the U.S. as a unit featuring a plug and receptacle for 

the power link, as well as a pair of telephone jacks 

or TV coaxial fittings for the communications link. 

However. its necessary effectiveness has not yet 

been quantified in any performance standard. 

To illustrate the effectiveness. Figure 5 shows 

the reduction of the voltage obtained by inserting a - 

typical surge reference equalizer in the power and 

telephone lines at the point of connection of the PC. 

The generic design of such a device includes 

insertion in the two telephone wires of two matched 

gas tubes. two series resistors. and two silicon 

avalanche diodes. with a shared earthing reference. 

Figure 5 shows the immediate clamping effect of 

the diodes down to 200 V. followed by a further 

reduction of voitage as the gas tube sparks over. 

power service entered at the same end of t

house. the recommended practice. With such 

configuration. a reduction in the voltage difference 

about 75% of the large loop value was found in t

test series. Available space limits the number 

records that can be shown in this paper for vario

combinations. but a more comprehensive report w

be prepared and published later. 

SECOND EXPOSURE SCENARIO: 

ONE SYSTEM SERVING TWO APPLIANCES 

VIA TWO BRANCH CIRCUITS 

In this scenario. a PC and the associated prin

are connected by the usual communications cab

and each is powered by a separate branch circu

This situation is often encountered when a printe

shared among several users. or when an installat

has been deliberately configured to provide 

separation of the 'clean' branch circuit supplying t

PC from the "noisy" branch circuit supplying t

printers and other peripherals (Figure 6). Bo

branch circuits originate at the service panel. b

might not have the same length. 

In a first case of this scenario, a slight differen

may occur in the time of arrival at the two ports o

surge originating outside of the building ("EX" 

Figure 6). A greater difference in the time of arri

would occur if the surge were internally genera

("IN") along a branch circuit. propagating directly 

that branch toward the PC and in a roundabout pa

via the service entrance and the other branch circ

toward the printer. Wih the internally genera

surges having steeper fronts that the externa

generated surges [Martzloff. 19901. the difference 

arrival time would be significant since the volta

spike occurs upon the initial current rise, not at t

peak of the current surge. 

.-.. 

.... -... .... .... 

.... 

.... 

Other noisy equipnent 

Figure 5 - Mitigation obtained by Figure 6 - Personal computer and printer 

inserting a surge reference equalizer in . linked by a data cable and powered from 

the power and telephone lines two separate branch circuits 
severe situati~n is created by mismatched protective 

devices at the power ports of the two appliances. A 

shift in the reference voltage can occur if one SPD 

provided in either the printer or the PC invites a 

disproportionate surge current in its PE conductor. 

These can be built in the appliance or be a plug-in 

device installed by the user. Such a device, if it 

includes an SPD connected L-G, will return the 

surge to the service entrance through the protective 

conductor PE ("equipment grounding conductor is 

the US. term) and produce a shift in the voltage of 

the corresponding chassis. The resulting difference 

of voltage between the two chassis will be applied 

across the communications link with possible 

upsetting or damaging consequences. 

The effect of such difference in SPDs -is 

illustrated in Figure 7. for the worst case scenario of 

one SPD connected L-G in one appliance, and none 

in the other. An oscillatory difference of voltage 

peaking at 3.2 kV, with a spike in the nanosecond 

range. occurs at the time of the initial rise. for the 

400 Alps rate of current change corresponding to an 

8/20 ps 1400 A peak surge injected at the service 

entrance. Note the decay of the voltage to a low 

value at the time of the curtent peak 

DISCUSSION 

The two exposure scenarios described in this 

paper represent the mechanisms most likely 

responsible for many of reported. but seldom welldocumented, field failures of two-port amliances. 

The variables considered in these two scenarios 

are marked in the cells of Table 1. (1) for the first 

scenario, (2) for the second. The columns in the 

Figure 7 - Difference of reference 

voitages caused by dierent protections 

in the two separate branch circuits 

table correspond to spatial variables; only a few 

all the possible variables are shown. The rows 

the table show a few of the possible variable

corresponding to the nature and combinations of th

SPDs. This table is a beginning toward defining th

mufti-dimensional matrix of all possible variables. 

Many other variables need consideration, suc

as the presence of more than two ports in th

appliances, different types of ports (serial RS23

Ethernet ...), dierent power system configuration

(single-phase 120/240 V or three-phase 120/208 

in the US.. three-phase systems in other countries

. wiring errors and poor practices, lack of coordinatio

between upstream and downstream SPDs. larg

or higher buildings, separate buildings, immun

levels of equipment and consequences of insufficie

immunity (upset vs. failure, failure modes), and. la

c-. ..-. I---., uvvm.vmmmr ncauu-urlu. uy IGVlGW 

these many variables. it may be possible to ident

a limited number of scenarios and thus defin

effective mitigation means. 

Table 1 

Equipment power port SPD I (2) (1) (2) I 

Telephone service entrance SPD 1 (3 ) (1 

Two-dimension matrix of variables considered for reference voltage shifts in two scenarios 

1 . . I . . Equipment telephone port SPD ' (1 ) 1 (1) (2) 

Other SPD locations and types 

. . .. . . . . . 

Built in the appliance, or external, userprovided plug-in SPD 

Spatial variables 

SPD variifes 

Sile entrance 

of services 

("ground whdour) 

Multiple points of 

entrance of power (L 

telephone service 

W~ring routing 

and 

pradices 

Other spat'd 

variables of the 

installation ,. 
One solution. illustrated in this paper, is the insertion 

of a properly designed surge reference equalizer. 

One cause of the problem h the flow of large surge 

currents in the wiring system of the building. Wih a 

telephone senrice entrance located at the opposite 

end of the power service entrance, the required 

bonding of the NID unavoidably involves the surge 

current in the long bonding connection. hence the 

need for preventive mitigation for this type of nonrecommended telephone service installation, unless 

adequate betweenports immunity of the appliance 

is documented. While these examples of two such 

exposure scenarios have illustrated the mechanisms. only a computer-driven model might cover 

all possible combinations of the many variables 

that could be encountered in all existing or future 

installations. Hence, it is essential that a comprehensive and well-documented experimental data 

base be established for validation of the model. 

For surges impinging on the power service 

entrance. the problem is associated *th large surge 

currents flowing in the branch circuits. if the surge 

current were diverted at the service entrance by a 

suitable SPD. the problem would be reducedHowever. the specification of a "suitable SPD' at the 

service entrance involves the issue of coordinating 

cascaded SPDs [Martzloff-Lai. 19921. The ongoing 

program of measurements at the Upside-Down 

House will include measurements and numerical 

simulation of cascaded SPDs. A joint Working 

Group of the IEC is developing guidelines for 

cascade coordination. based on the work of many 

-researchers [Goedde, 19901; [Standler. 19911; 

[Hostfet et at.. 19921; [Hasse et at.. 19931. 

CONCLUSIONS 

1. Quantitative measurements in the Upside-Down 

House clearly show objectionable differences in 

reference voltages. These &ur even when, or 

perhaps because. surgeprotective devices are 

present at the point of connection of appliances. 

2- Accounting for all the variables may be done in a 

multi-dimensional matrix, a task for computer 

analysis. the next step of this project. 

3. The analysis should be directed toward obtaining 

a limited set of typical scenarios resulting from 

the many combinations of many variables. 

The laboratory measurements were made wi

the kind assistance of Doni Nastasi and Kerm

Phipps. Glenn Grotz and Donald Worden provide

peer-review comments on the test procedure an

interpretation of the results. 

This project was made possible by a sabbatic

granted by NlST to Martzloff. allowing him 

experiment with his concept of the 'Upside-Dow

House.' Thk Upside-Down House was installed 

PEAC. with support of the Electric Power Researc

Institute. as a permanent facility of the Pow

Quality Testing Network 

The authors also explicitly invite their peers 

conduct similar experiments and corrputations 

build a consensus on the threat represented b

shifting reference voltages. Such a consensu

would be an important step toward the developme

of standards for system compatibility. 

REFERENCES 

Goedde, G.L.. Marz. M.B., and Henry, D.C., Coordinating 

Lightning Stroke Protedion From the Wii System to Load 

Devices. Proceedings, Second lnternatronal Power Qualify 

AS0 Conference. October 1990. Philadelphia. 

Hasse, P.. et al., Lightning current arrester containing spark 

gaps for AC supply mains. EleMrie. June 1993. 

Hostfet, O.T.. et al., Coordination of surgeprotective device

in power supply systems: Needs for secondary protedion. 

Proceedings, international Conference on Lghtning 

Prdedbn. Berlin, September 1992 

EEE Std 1100-1992, lEEE Recommended Pradice for 

Powevrng and Grounding Sensitive Elecbonic Equipment 

CEmerald Boor of the IEEE Cdor Books series). 

Martzloff, F.D., Coupling, Propagation, and Side Effeds of 

Surges in'an Industrial Building Wiring System. IEEE 

Transactions /A-26, MdApril 1990, pp. 193-203. 

Martzloff, F.D. and Lai, J.S., Cascadiig Surge-Protedive 

Devices: Coordination vs. the IEC 664 Staircase. 

Proceedings. PQ4'91 Conference. pp 191-198. 

Standler. 4.0.. Coordination of Surge Arresters and 

Suppressors for Use on Low-Voltage Mains. Proceedings, 

199 1 Zirrich EkfC Symposium. pp 51 7-524. 

Van Deursen. A.PJ., Electromagnetic Compatibikty, Pat 5

Installation and MRigation Guidelines. Section 3, Cabling an

Wiring, EUT Report 93-E-275. Eindhoven University of . 

Technology. July 1993. 
SURGING

THE

UPSIDE-DOWN

HOUSE:

MEASUREMENT

AND

MODELING

RESULTS
Electronic equipment with two input ports - power and communications - can be exposed to damaging 

differences of voltage between the two ports during surge events. To identify and quantify the significant 

variables and their effects during surge events in residential or commercial facilities, a representative 

configuration of the circuitry in a residence (metallic cold water pipe, power and grounding conductors, 

telephone and coaxial cable TV wiring) was set up in the laboratory, under the name of “Upside-Down 

House.” 

To evaluate the threat of surges impinging upon an actual installation, surges of various types, as defined in 

standards covering AC power circuits and communications circuits, can be injected at selected points of the 

Upside-Down House. Typical surge-protective devices (SPDs) can be placed at suitable locations of the 

Upside-Down House, corresponding to a variety of real-world exposure scenarios. Preliminary experimental 

results of two exposure scenarios were reported in a PQA'94 paper (see pdf file “Upsdown surging”). 

Additional measurements and parametric variations are reported here to characterize the impedance of the 

various components of the wiring system and the source impedance of the resulting overvoltages appearing 

between the ports. 

Filename: Upsdown measure
1
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8
9
10
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15
DRIVING

HIGH SURGE

CURRENTS

INTO

LONG CABLES

MORE

BEGETS

LESS
Power Electronics Applications Center

Knoxville TN 37932 USA

Amansoor@epri-peac.com

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

© 1996 IEEE

Reprinted, with permission, from IEEE Transactions on Power Delivery, Vol.12, No.3, July 1997

Significance:

Part 2 Development of Standards – Reality checks

Part 4 Coupling and propagation of surges

In the propagation of a surge current injected at the service entrance of a building, two significant factors can

prevent the propagation of a postulated “large” surge current to the end of the branch circuits of the facility. 

1. The combination of the inherent inductance of the wiring and the high rate of current change for such a

current to begin flowing into the branch circuit results in a high voltage at the driving end (V = L x di /

dt).

2. In the absence of a surge=protective device at the service entrance, the withstand voltage of the wiring

devices at the driving end – the service entrance – is very likely to be exceeded by the voltage that this

rising current will develop along the branch circuit. 

The resulting flashover will abort further propagation of the surge current toward the far end, thus

establishing a limit to what is physically possible. If there is a surge-protective device at the service

entrance, the scenario becomes a matter of cascade coordination.

Provides quantitative information on this limitation, as a function of wiring length and current rate of rise.

Filename: More begets less
1176 IEEE Transactions on Power Delivery, Vol. 12, No. 3, July 1997 

High Surge Currents into Long Cables: 

Arshad Mansoor, Member, IEEE 

Power Electronics Applications Center 

Knoxville, TN 37932 USA 

Abslract - Reality checks can and should be applied to proposals 

for characterizing the surge environment and application of surgeprotective devices (SPDs) to end-user, low-voltage power systems. 

One such check is the fact that driving a large current with steep 

front toward an SPD installed at the far end of a branch circuit 

cable could require such a high voltage that the connections at the 

near end of the cable will flashover, limiting the stress applied to the 

far-end SPD. Tests and numerical modeling were performed to 

support this thesis. The results of real-world measurements and 

modeling, presented in the paper, are in good agreement and 

validate each other. From that point on, the model allows 

parametric variations of cable length and surge current amplitude 

and waveform, of which several examples are presented. 

I. INTRODUCTION 

In the never-ending quest for better data on the frequency of 

'occurrence and level of threat of overvoltages, we should not 

overlook some "reality checks" that can be applied to proposals 

for characterizing the surge environment. One such check is the 

fact that forcing a large surge current with steep fiont toward a 

surge-protective device (SPD) installed at the far end of a branch 

circuit cable could require such a high voltage that the wiring 

device connections at the near end of the cable will flashover, 

limiting the stress applied to the far-end SPD. 

' Electricity Division, Electronics and Electrical Engineering 

Laboratory, Technology Administration, U.S. Department of Commerce 

Frangois D. Martzloff, Life Fellow, IEEE 

National Institute of Standards and Technology ' 

Gaithersburg, MD 20899 USA 

Large surge currents considered by standards-writing bodies 

and discussed in this paper are presumed to impinge from the 

outside of a building, as a result of a direct or indirect lightning 

flash. These involve postulated rise times in the order of a few 

microseconds, with a duration ranging from a few tens to a few 

hundreds of microseconds. While there are different propositions 

made on what duration should be considered as "representative" 

waveforms, there is a consensus on rise times ranging from about 

4 ps to 20 ps [I]. However, consensus on what value to select 

for "representative" amplitude(s) has been challenged by 

proposals to increase the current surge capability of devices 

intended for installation at the end of branch circuits. 

A growing trend in the application of SPDs to residential or 

commercial installations is to provide " whole-house protection" 

with an upstream SPD connected at the service entrance, and 

downstream SPDs in the form of plug-in devices installed at 

receptacles. Selecting the ratings for these two devices is the 

subject of some debate. The voltage rating of the devices 

introduces the issue of cascade coordination which has been 

addressed at length in the literature [2]-[8] and will not be 

discussed here. At this point in time, the vast majority of 

installations do not include an upstream SPD intentionally 

connected at the service entrance, other than a gap in the 

revenue-meter socket. This gap is provided by the meter 

manufacturer to protect the meter more than the downstream 

installation. Nevertheless, there are other "gaps" at the service 

panel -- the clearances of the wiring devices, which have some 

limits to their voltage withstand capability. 

II. SURGE PROPAGATION IN WIRING 

The possibility of a clearance flashover is the basis of our 

96 SM 399-6 PWRD A paper and by the lEEE thesis: If a large surge current is postulated as propagating Surge Protective Devlces Comrn~ttee of the IEEE Power Engmeering 

Soc~ety for presentatm at the 1996 IEEEIPES Summer Meetmg, July 28 (and then taken as a requirement for the 

-August 1, 1996, In Denver, Colorado Manuscript subm~tted December SPD), the propagation characteristics of this surge current would 

28, 1995; made available for printing June 19, 1996. result in high voltages at the service entrance, upstream. In turn, 

the high voltage would cause flashover of upstream clearances, 

acting as a relief valve for the surge energy headed for the 

downstream SPD. This relief action would then contradict the 

proposed requirement for high energy-handling capability of the 

downstream SPD. Thus, appropriate selection of current ratings 

for the downstream SPD, in the light of our thesis, should take 

into consideration this reality check that defines an upper limit 

for the current rating required for the downstream SPD. 

0885-8977/97/$10.00 O 1996 IEEE 
The surge propagation characteristics mentioned in the 

preceding paragraph are controlled by three parameters: the 

impinging surge, the impedance of the wiring from the service 

entrance to the downstream SPD, and the I-V response of the 

downstream SPD. The impinging surge could be considered 

either as a voltage source or as a current source. The present 

consensus is to consider it as a current source, resulting from the 

coupling and subsequent division of a lightning surge, part of 

which impinges on a given service entrance. 

The impedance of the wiring is that of two parallel wires of 

known dimensions and separation. It can be represented either 

by lumped parameters -- series R and L and parallel C -- or by a 

"short" transmission line. The reason for placing quote marks 

around the qualifier of "short" is that the term is to be viewed by 

comparing travel time over the length of the transmission line and 

duration of the traveling pulse -- another subject discussed in the 

literature [9] that we will not discuss here, with the exception of 

a brief comparison of results obtained when modeling the 

propagation with lumped parameters or with a transmission line. 

When using the lumped RLC model, during the rise of the 

surge current, the significant parameter of the wiring impedance 

is its inductance, L. The voltage at the upstream end resulting 

from driving the surge current into such an impedance is primarily 

L x dildt, with dildt determined by the amplitude and rise time. 

By performing surge measurements on real-world wiring 

components, followed by numerical modeling with the Electrolmagnetic Transients Program (EMTP)' [lo], this proposition can 

be verified and applied to a range of postulated surge waveforms 

and typical configurations found in the premises wiring of lowvoltage systems. These results will allow developing realistic 

recommendations for the rating of SPDs offered for surge 

protection at the equipment location -- either as plug-in additions 

by the end-user, or as permanently wired devices at the end of 

typical branch circuits. The measurement results also show the 

need to consider the possibility of "blind spots" in the protection 

schemes, and illustrate our title paradox of "more begets less." 

Measurements were conducted on a simple circuit consisting 

of 9 meters of nonmetallic jacket cable typical of residential 

installations, with a metal-oxide varistor connected downstream 

at the far end. A Combination Wave surge generator, suitable 

for producing the waveform described in IEEEIANSI C62.41- 

1992 [I] was used to inject a surge current at the upstream end 

of the cable. Current and voltage waveforms were recorded. 

The current waveform resulting from this injection was 

duplicated in a closed-form equation to be applied as the 

postulated surge current injected into the EMTP model of the 

circuit, allowing computation of the corresponding voltages. 

' Certain commercial instruments and software packages are identi$ed 

in this paper in order to adequately specify the experimental procedure. 

Such identzjication does not imply recommendation or endorsement by 

\the National Institute of Standards and Technology, nor does it imply 

that these are necessarily the best available for the purpose. 

111. MEASUREMENTS AND MODELING 

A. Characterizing the varistor 

First, the varistor to be connected at the far end was tested to 

determine its I-V response and demonstrate that the model to be 

used for this highly nonlinear component would be adequate to 

simulate its behavior in the circuit when connected at the downstream end. Figure 1 shows the test circuit used for making that 

measurement. The surge generator used for the tests was the 

KeyTek 7 1 1 with a P7 wave-shaping output network. 

The varistor used in these tests was a 20-mm diameter metaloxide varistor (MOV) disc, rated 130 V rms (200 V at 1 mA dc). 

The inductance Lp shown in series with the varistor is not a 

deliberate addition of a real component, but is the representation 

of the coupling between the loop where the surge current flows 

and the voltage measurement loop formed by the varistor leads 

and the two probes used for the differential measurement. That 

inductance is included in the model as a discrete series 

inductance, with a value of 0.5 pH selected to emulate the 

observed voltage at the point of measurement -- which is not the 

"pure" varistor voltage, as discussed in the narrative of Figure 2. 

Figure 2 shows the recording obtained for a particular setting 

of the surge generator, and Figure 3 shows the result of modeling 

the circuit shown in Figure 1 for an injected current surge corresponding to the actual current surge recorded in Figure 2. The 

equation used for the modeling is a damped sine wave that 

allows a close approximation of the current delivered by typical 

Combination Wave generators into inductive loads [7]. It is 

known that actual generators tend to produce an "undershoot" 

when connected to an inductive load, and this test was no 

exception. However, computational artifacts occur when using 

a simple damped sine wave because its dudt derivative (a cosine) 

is not zero at time zero. Furthermore, we know that nature does 

not allow an instantaneous jump of current from zero to a steep 

rise. By adding a multiplier term [l-e'-t)], these artifacts are 

eliminated and the waveform has a "gentle toe" which is a better 

model of reality. This improved equation is then: 

= 4200 * sin((). 12fjt) * e(-t'28.') * [ 1 - e( -1 (1) 

with I in amperes and t in microseconds. 

Digital 1 signal 

analyzer printer 

Figure 1 - Test circuit for determination of 

the I-V characteristics of the varistor 
DSA 602A DIOITIZINQ SIGNAL ANSYZER dsto: 91-*"a-95 tlmr: 14:38:47 

(Center trace: inactive) 

Bottom trace: Current, 500 Ndiv 

Sweep: 10 psldiv 

Figure 2 - Real-world recording 

Inspection of Figures 2 and 3 clearly shows the agreement 

between real-world measurements3 and model, and thus merits 

some observations. One might have expected a flat-top voltage 

waveform reflecting the clamping action of the varistor. Instead, 

a drooping waveform is observed. This droop is caused by the 

parasitic inductance Lp in series with the ideal varistor. At the 

time of current peak (di/dt = O), the "true" varistor voltage is 

seen on the oscillogram. Before the peak, the positive Lp x di/dt 

adds a spurious voltage to the recording. After the peak, the 

/negative Lp x di/dt subtracts the spurious voltage. 

These observations are significant in appreciating the allimportant inductive effects during the rise and fall of a surge 

current in the wiring of branch circuits. The issue of the 

importance of inductance versus other circuit parameters [I11 

hopefully has been put to rest by the surge and impedance 

measurements with corresponding computations performed in 

the so-called "Upside-Down House" [12], a real-world replica of 

a typical residential wiring system. In [12], it was shown that 

inductive effects prevail, so that rate of rise of the surge current 

and circuit inductance, more than any other parameter, are the 

significant parameters for the voltage necessary at the upstream 

end to drive a given current into the branch circuit. 

The model used in the simulation for the varistor is derived 

from the published varistor I-V characteristic (general shape and 

slope of the curve) with one specific point defined by the "true" 

varistor voltage read from the oscillogram of Figure 2 at the 

point of zero Lp x di/dt contribution. In turn, this varistor model 

will be used for the modeling of a varistor connected at the 

downstream end of a branch circuit, as discussed in the following 

reported measurements and simulations. 

The measurements reported in this paper have been made with 

instrumentation for which the cumulative uncertainty should not exceed 

5 to 6%. Given the process of applying the measurement results to the 

response of surge-protective devices exposed to environment with 

)charactenstics that are at best known within an order of magnitude, 

this level of uncertainty does not affect the practical conclusions. 

Note: the voltage trace has been expanded by a factor of 2 to 

enhance resolution on the vertical scale. 

Figure 3 - Modeling the circuit of Figure 1 with the impinging 

current set to match the test current, as shown in Figure 2 

B. Measurement and modeling with varistor installed 

at the downstream end of a branch circuit 

The circuit of Figure 4 shows the varistor characterized by 

the test and modeling in the preceding paragraphs, connected at 

the downstream end of a "branch circuit" consisting of two 

copper conductors of 2-mm2 cross-section (#I2 AWG) with solid 

insulation and a separation of 6 mm between centers. The first 

current transformer monitors the total current impinging at the 

upstream end. The second current transformer monitors the 

current flowing toward the downstream end, which will be 

imposed on the varistor. The clearances at the upstream end, 

such as clearances in a service-entrance panel, are represented by 

a discrete gap that will be set to produce sparkover at some given 

voltage during the test as well as in the model. 

Figure 5 shows the recording obtained with the circuit of 

Figure 4, with the surge generator left at the same setting as that 

used for Figure 2. To determine the response of the circuit 

without the clearance limitation, the gap setting was adjusted for 

this test so that no sparkover occurred at the upstream voltage 

developed for the current delivered by the generator. 

9 meters 

- analyzer prlnter 

Figure 4 - Test circuit for determination of the voltage 

necessary at the sending end to drive a given current 

into the far-end SPD 
Top trace: Gap voltage, 500 Vldiv 

Center trace: Total current, 500 Ndiv 

Bottom trace: MOV current, 500 Afdiv 

Sweep: 10 psldiv 

Rgure 5 - Real-world recording of sending-end 

voltage with gap set for no sparkover 

Comparing the traces of Figure 5 and Figure 2, the addition 

of the inductance of the 9 meters of branch circuit changes the 

load on the surge generator, reducing the current peak from the 

2.8 kA in Figure 2 down to 2 kA in Figure 5. 

The two current traces of Figure 5 are identical. Since there 

is no current diverted by the gap, the current in the branch circuit 

I is the same as the current delivered by the surge generator. 

Another effect of the added inductance is the increase in the 

time from origin to the first current zero, 33 ps in Figure 5, 

compared to 25 ps in Figure 2. In the subsequent model, that 

change of the actual impinging current surge is taken into 

consideration by modifying the current equation as follows: 

I = 3571 * sin(0.095 t) * e(-t'26.') * [I-e'"] (2) 

with I in amperes and t in microseconds. 

Turning to the modeling, Figures 6 and 7 show the 

waveforms of the impinging current, as defined by Eq. (2), and 

the resulting voltage at the upstream end. To address some 

concerns expressed by colleagues in discussions of this subject, 

the EMTP modeling was also done with the transmission-line 

model which is readily available in the EMTP code. Figure 6 

was obtained with the lumped-parameter circuit model, and 

Figure 7 was obtained with the transmission-line model. 

Inspection of the two figures reveals no difference in the 

results. The only difference is in the consumption of computing 

time: with the transmission line model, the computation timestep has to be significantly shorter (0.02 ps in this case) than the 

travel time for the reflections, while in the case of the lumped 

model, the time-step can be longer (0.1 ps in that case). The 

result is that the simulation of Figure 6 took 43 seconds on a 

486-based PC, compared to 263 seconds for Figure 7. 

Therefore, the lumped-parameter model is perfectly adequate to 

represent reality, and performing a transmission-line analysis [5] 

is an unnecessary consumption of computing time and resources. 

Figure 6 - Impinging current and resulting upstream 

voltage as computed with lumped-parameters model 

Figure 7 - Impinging current and resulting upstream 

voltage as computed with transmission-line model 

In both Figures 6 and 7, the effect of the branch circuit 

inductance on the resulting voltage is apparent as the peak voltage 

occurs at the beginning of the rise (as soon as the "gentle toe" 

effect ceases), not at the peak of the current. The step change in 

the voltage trace corresponds to the reversal of the current in the 

varistor, showing the relative contributions of the varistor effect 

and of the inductive effect as seen from the upstream end. 

Table 1 below shows the results of such computations for the 

waveform of Figures 5, 6 and 7. As mentioned above, the 

insertion of an inductance in the load connected to the surge 

generator increased the rise time beyond the standard 8 ps. In 

making the parametric computations, we chose to stay with this 

10 ps value to maintain continuity with the test/model validation. 

TABLE 1 

Upstream voltage (in kV) necessary to drive a current of the peak 

value shown (columns) and rise time of 10 ps into a branch circuit 

of length as shown (rows), terminated with a 130-V rated varistor 

Length\Peak 2kA 3kA 5kA 7kA 10kA 
Figure 8 - Three surge current waveforms with different rise 

times used to compute the values of Table 2 

Figure 8 shows three waveforms of same amplitude, with 

nominal rise time of 5 ps, 10 ps, and 20 p, obtained by taking 

half or double of the frequency used in Eq. (2). The actual rise 

time [1.25 x (time from 10% to 90%)], as opposed to the nominal 

rise time used to describe the waveforms, was computed as well 

as the maximum rate of rise for each wave. The maximum rate of 

rise (which is obtained when the second derivative of the current 

is equal to zero) occurs initially, once the gentle toe is over, and 

determines the maximum resulting voltage produced by the 

inductive effect. Table 2 shows the corresponding values of the 

rise time, maximum rate of rise, and resulting voltage for a 

branch circuit length of 10 m and amplitude of 5 kA. Note that 

for a 1-to-4 increase in nominal rise time, the maximum di/dt 

decreases only by one half, with the same decrease appearing in 

the resulting voltage, showing once again that initial rate of rise 

is more important than rise time and amplitude. 

TABLE 2 

Effect of the rate of rise of the postulated current on the 

resulting voltage at the upstream end of the branch circuit 

Nominal rise time, ps 5 10 20 

Actual rise time, ps 4.3 9.5 13.5 

Maximum di/dt, A!ps 1250 850 630 

Resulting voltage, kV 7.0 5.2 3.6 

In the scenario tested and modeled so far, no flashover 

possibility was considered. Nevertheless, the values shown in 

Table 1 clearly indicate that some real-world circuit lengths and 

surge parameters postulated in some SPD application standards 

under development can produce high upstream voltages that will 

cause a flashover of the upstream wiring devices. 

C. The paradox of "more begets less" 

Common-sense intuition might lead the unwary to expect that 

higher surge currents would impose a greater stress on the circuit 

components, including the downstream varistor. Also, a longer 

branch circuit, with its corresponding higher inductance, could 

be expected to have the capability of storing more energy during 

build-up of the surge current toward the downstream varistor, 

into which that stored energy ultimately has to be dissipated. 

Cascade coordination studies [4], [6], [a], have shown that in 

some cases, the downstream varistor continues to carry current 

long after the impinging surge current has gone past its peak. 

To explore the validity of such expectations, we performed 

tests and modeling, with an actual gap in the test circuit, and a 

switch in the model circuit, to bypass the current at the upstream 

end when sparkover voltage is attained. By measuring the 

current that flows in the branch circuit toward the downstream 

varistor and the voltage across the varistor, the energy deposited 

in the varistor during the total surge event can be determined. 

Likewise, the modeling can determine the current in the varistor, 

hence the voltage across it, and allow computation of the energy. 

In [4], agreement was reported between, on the one hand, 

computing the deposited energy through actual measurement of 

the current and voltage, followed by computation of the energy 

by means of the digital signal analyzer used for measurements 

and, on the other hand, the model computations. Therefore, in 

the tests reported here, we were satisfied to verify waveform 

agreement between the actual varistor current measurement and 

the computed varistor current, and let the model alone compute 

the energy deposited in the downstream varistor. 

Figure 9 shows the real-world recording of the situation that 

develops for a "clearance" sparkover of 2 kV. This relatively 

low value, compared to the 6 kV to 10 kV level that we might 

expect from typical low-voltage wiring devices, is made necessary for the test case where only 9 meters of branch circuit were 

considered, and the setting of the surge generator was maintained 

at the same nominal 3 kA short-circuit current. The object, of 

course, is to demonstrate that the clearances are likely to flash 

over, as indicated by progressively higher values of the necessary 

upstream driving (or resulting) voltage shown in Table 1. 

Under the conditions of Figure 9, sparkover of the gap 

occurred at approximately 1 ps. After sparkover, the current 

delivered by the surge generator is the sum of the currents in the 

gap and in the branch circuit. Its peak (3.2 kA) is greater than 

those of Figures 2 and 5 because the generator does not need to 

overcome the varistor that reduced the voltage available for 

driving the current, nor the impedance of the 9 meters of cable. 

Top trace: Resulting voltage, 500 Vldiv 

Center trace: MOV current, 500 Ndiv 

Bottom trace: Total current, 500 Ndiv 

Sweep: 10 usldiv 

Figure 9 - Voltage and currents with gap sparkover at 2 kV 
Figures 10 and 11 show the results obtained by the model for 

voltages and current in the circuit. In the modeling, only one 1 current waveform was applied to the circuit, the one prevailing 

until flashover occurs, which the postulated current-source real 

world would maintain. In contrast, the surge current delivered by 

the surge generator (Figure 9) increases after the flashover, but 

that is not relevant to our consideration of what happens to the 

circuit before and up to the time of flashover. 

I Time (us) 

Figure 10 - Voltage across the gap set to sparkover at 2 kV 

Figure 11 - Current in downstream varistor 

The waveforms of Figures 10 and 11 are shown with an 

expanded scale, compared to that of Figure 9, that gives a better 

resolution for the gap voltage and current in the varistor. There 

is good correspondence between the waveforms of the two traces 

and the gap voltage and downstream current traces of Figure 9. 

In Figure 10, however, the gap voltage collapses to zero, while it 

does not in Figure 9. The difference is that the real-world circuit 

has a parasitic inductive voltage added to the true gap voltage, 

already discussed for the varistor of Figure 2. Figure 11 shows 

the linear ramps typical of current changes in an inductance. 

As mentioned above, we can expect that the energy deposited 

in the downstream varistor for a given impinging surge will be 

influenced by the length of the branch circuit. Using the model 

developed and validated according to Figures 5 and 6, the energy 

can be readily computed. In the case described by Figures 9, 10, 

and 11, the gap sparkover voltage was preset at 2 kV so that 

sparkover could indeed occur for the surge current available 

from the real-world generator and the resulting upstream voltage. 

Now that we are in the (validated) model-world, we can 

arbitrarily set the sparkover voltage at a level more typical of the 

flashover point of clearances, say 6 kV. Of course, we have the 

possibility of assessing energy for a wide range of parameters. 

In the example reported below, we kept the same three values 

of branch circuit length and performed the computations for the 

same five values of impinging current as those used for the 

computations of Table 1. Table 3 shows the energy deposited in 

the downstream varistor for these combinations of branch circuit 

length and peak current values, for the applied current waveform 

of Figure 5, and a 6 kV flashover point. 

TABLE 3 

Energy deposited into a 130-V rated far-end varistor 

as a function of the branch circuit length shown (rows), 

current peak (columns) of waveform shown in Figure 5, 

and flashover of the clearances set to occur at 6 kV 

The results shown in Table 3 merit close examination as they 

reveal some counter-intuitive trends: we might have expected 

that for higher impinging current values, the resulting energy 

deposited in the downstream varistor would be higher. Likewise, 

we might also have expected that for a longer branch circuit, the 

greater inductance would store more energy, ultimately to be 

deposited in the varistor. In fact, the opposite occurs. The table 

also reveals the interesting finding that the first three lowercurrent, short-line cases (bold face type in the table) produce 

larger energy deposition, compared to the other cases. Actually, 

the explanation that follows is simple and might be anticipated 

(especially with hindsight, illustrating that intuitionis a hazardous 

process when dealing with nonlinear circuit components). 

Starting with the second observation (more joules at lower 

threat levels), we have a beautiful illustration of the blind spot 

effect -- not limiting tests and designs to the maximum stress of 

a worst-case scenario -- [13]: for 10 meters of circuit and at the 

lower current levels, the resulting voltage at the clearance is not 

sufficient to cause flashover, and all the energy has to go to the 

downstream varistor. At the higher threat level of 7 kA, the 

voltage produced in the inductance of 10 meters of line, added to 

the varistor voltage, is sufficient to sparkover the 6 kV gap, 

relieving the varistor from further involvement beyond that of 

discharging the energy stored in the line. In the case of the 30-m 

long line, this transition occurs between 2 kA and 3 kA. 

Turning now to the first observation, that higher current or 

greater inductance result in less stress, this apparent paradox is 

caused by the fact that with the higher values of di/dt and L, the 

voltage at the clearance rises more quickly to the flashover point. 

Consequently, the build-up of energy in the line inductance is 

shut-off earlier so that the current level in the line reached at that 

point is lower and, in spite of the greater inductance, the stored 

energy '/z L iZ is lower for higher applied current peaks and 

longer branch circuits. 
IV. CONCLUSIONS 

The development of a validated EMTP model using existing 

computational tools allows us to look into all scenarios of surge 

propagation and surge mitigation schemes. The reality check 

proposed by the measurements and modeling reported in this 

paper should be useful in the process of selecting stress levels to 

be specified in the application of SPDs downstream from the 

service entrance, from the point of view of successful cascade 

coordination as well as integrity of electromagnetic compatibility. Specific conclusions can be drawn: 

1. Realistic surge current amplitudes and rise times can be 

defined for SPDs installed at the end of branch circuits, with 

upper limits set by the laws of physics applied to real-world 

conditions. 

2. The general practice for describing surge waveforms is to cite 

"rise time" or "front time", followed by duration, as in 8/20. 

However, when the effects of circuit inductance are assessed, 

in particular by numerical modeling, the maximum rate of 

rise must be considered, not an average over the rise time. It 

is especially important to define the conditions at the origin 

of the waveform, such as inclusion of a gentle toe. 

3. The importance of looking for blind spots is, once again, 

demonstrated by the parametric computations, a much 

simpler task than exhaustive equipment-exhausting tests. 

4. Reliable computational tools makeit possibleto obtain a wide 

range of parametric assessments, and thus avoid recourse to 

intuition when dealing with nonlinear circuits, where blind 

reliance on common-sense may lead to flawed conclusions. 

5. The parametric computations offered in the paper point out 

the need to consider a balance or trade-off among several 

critical factors in the design of branch circuit protection, in 

particular the uncontrollable length of branch circuits in 

actual installations. 

V. ACKNOWLEDGMENTS 

The measurements reported in this paper were conducted by 

Kermit Phipps at the EPRI Power Electronics Applications 

Center. Support for the modeling was provided by the Electric 

Power Research Institute. Support for the development of this 

reality check, motivated by the discussions and contributions of 

In Proceedings, Second International Power Quality ASD Conference, 

October 1990, Philadelphia. 

[4] Lai, J.S. and Martzloff, F.D., "Coordinating Cascaded Surge 

Protection Devices: High-Low versus Low-High," In Conference 

Record, IEEE-IAS 1991 Annual Meeting, pp 1812-1819. Also in IEEE 

Transactions, IAS-29 No.4, July/August 1993, pp 680-687. 

[5] Standler, R.B., "Coordination of Surge Arresters and Suppressors 

for Use on Low-Voltage Mains." In Proceedings, 1991 Ziirich EMC 

Symposium, pp 517-524. 

[6] Hostfet, O.T., Hervland, T., Nansen, B. and Huse, J., "Coordination 

of of surge-protective devices in power supply systems: Needs for 

secondary protection," In Proceedings, 20th International Conference 

on Lightning Protection, 1992. 

[7] Hasse, P., Wiesinger, J., Zahlmann, P. and Zischank, W., "Principle 

for an advanced coordination of surge protective devices in low voltage 

systems," In Proceedings, 22nd Inter-national Conference on Lightning 

Protection, 1994. 

[8] Rousseau, A. and Perche, T., "Coordination of Surge Arresters in 

the Low-Voltage Field," In Proceedings, Seventeenth International 

Telecommunications Conference (INTELEC 95),95CH35824, pp 119- 19: 

I&>. 

[9] Martzloff, F.D., "Coupling, Propagation, and Side Effects of Surges 

in an Industrial Building Wiring System," IEEE Trans-actions IA-26, 

No.2, MarchIApril 1990, pp 193-203. 

[lo] EPRI, "Electromagnetic Transients Program (EMTP), Version 2.0; 

Volume 1; Main Program; Volume 2: Auxiliary Routines" EPRI Report 

EL-6421-L, July 1989. 

[ll] Standler, R.B., 'Transmission Line Models for Coordination of 

Surge-Protective Devices," In Proceedings, IEEE Inter-national 

Symposium on Electromagnetic Compatibility, 1993, pp 358-362. 

[12]Martzloff, F.D., Mansoor, A., Phipps, K.0, and Grady, M. 

"Surging the Upside-Down House: Measurements and Modeling 

Results," In Proceedings, PQA'95 Conference, EPRI, 1995. 

[13] IEEEYANSI C62.45-1992 "Guide on Surge Testing for Equipment 

Connected to Low-Voltage AC Power Circuits." 

Arshad Mansoor (M' 1995) Is Electrical Systems Engineer at the EPRI 

Power Electronics Applications 

Center (PEAC). He received his MS 

and Ph.D. in electrical engineering 

hom the University of Texas, Austin, 

in 1992 and 1994 respectively. His 

areas of interest include power 

quality, power systems transients, 

harmonics, surge protection, and 

EMTP model development. 

IEEE and IEC colleagues, was provided by Delmarva Power & 

Light and by Pacific Gas & ~lectric. 

VI. REFERENCES 

[lIIEEElANSI C62.41-1991, "Recommended Practice on Surge 

Voltages in Low-Voltage AC Power Circuits." 

[2] Martzloff, F.D., "Coordination of Surge Protectors in Low-Voltage 

AC Power Circuits," IEEE Transactions PAS-99, No.1, January1 

February 1980, pp 129-133. 

[3] Goedde, G.L., Marz, M.B., and Henry, D.C., "Coordinating 

Lightning Stroke Protection From the Utility System to Load Devices." 

FranqoisMartzloff (M'1956, F'1983) 

Born and educated in France, with 

additional MS degrees from Georgia 

Tech and from Union College, 

worked at General Electric for 29 

years and now ten years at the 

National Institute of Standards and 

Technology. He is contributing to 

several committees for development 

of standards on EMC and surge 

protection in the IEEE and the IEC. 
Discussion 

M. Darveniza (University of Queensland, Australia 4072): 

The authors are to be congratulated for drawing attention to an important feature in surge 

protection. Namely, the magnitude of the voltage up-line from a surge protective device 

will exceed the protective level of the SPD, the extent of the over-voltage depending on 

distance and on waveshape of the incident surge. Because of this, flashover at an 

upstream device (for example, another SPD) will limit the severity of the surge stressing 

the downstream SPD. 

Two examples are offered which support the author's statement "more begets less" and 

which can he rephrased by saying that in some cases, a less onerous surge may impose 

more severe overstress than a more onerous incident surge. The two examples are: 

1. A cable-entry substation protected by an upstream SPD connected at the 

overhead line-tocable junction. The most severe stress at the substation occurs 

when the surge incident from the line onto the cable is just not large enough for 

operation of the SPD at the line-cable junction. 

2. A hybrid surge protection system for low-voltage and electronic equipment 

involving two SPD's coordinated by an intervening series impedance. The series 

impedance is selected to ensure that the downstream SPD is not overstressed, by 

virtue of operation of the upstream SPD caused by the voltage drop in the 

impedance (which adds to the clamp voltage of the downstream SPD). However, 

if the voltage drop is not large enough to "turn-on" the upstream SPD, either 

because the magnitude or the steepness of the incident surge current is not 

sufficiently large, then the downstream SPD may still be overstressed if the 

duration of the surge current is too long. Paradoxically, a more severe incident 

surge will "turn-on" the upstream SPD, thus protecting the downstream SPD 

from excessive overstress. 

Manuscript received October 4, 1996. 

Franqois Martzloff : 

We are glad that the message we were presenting has found a 

favorable echo with Professor Darveniza, and appreciate his kind 

words. The two examples he cites are indeed good illustrations 

of the "more begets less" theme which we have expressed in the 

manner of a paradodepigram to make it easy to remember. This 

reality check on the likelihood of a stress-limiting flashover 

should be applied whenever a scenario is proposed that involves 

the propagation of surge currents. We hope that our colleagues 

involved in standards development will remember this well and 

assess some of the proposals now under consideration for SPD 

requirements in the light of that epigram. 

Manuscript received November 13, 1996. 
EFFECT OF

NEUTRAL

EARTHING

PRACTICES

ON

LIGHTNING

CURRENT

DISPERSION
Amansoor@epri-peac.comf.martzloff@ieee.org

© 1996 IEEE

Reprinted, with permission, from IEEE Transactions on Power Delivery, Vol.12, No.3, July 1997

Significance:

Part 2 – Development of standards – Reality checks

Part 4 – Propagation and coupling of surges

In the case of a direct lightning stroke to a building, the earth-seeking current is dispersed among all

available paths to earthing electrodes, including intentional made electrodes and opportunistic electrodes. 

A substantial part of that current will exit the building via its connection to the power distribution system.

The initial part of the stroke current dispersion is dominated by the relative inductance values of the

multiple paths and the tail of the stroke current dispersion is dominated by the relative resistance values of

the multiple paths.

The configuration of this power distribution system (daisy chain from the transformer or radial from the

transformer) influences the sharing of the current among these possible paths.

From simulations performed with a 10/350 waveform, the paper provides quantitative information on these

effects.

Filename: Dispersion
IEEE Transactions on Power Delivery, Vol. 13, No. 3, July 1998 

The Effect of Neutral Earthing Practices on 

Lightning Current Dispersion in a Low-Voltage Installation 

Arshad Mansoor, Member, IEEE 

Power Electronics Applications Center 

Knoxville TJX 37932 USA 

Abstract - Computer modeling with the EMTP code has been 

applied to several configurations and earthing practices in use in 

various countries to show the effect of any differences in the 

dispersion (sharing) of a lightning stroke current among the available 

paths for the earth-seeking lightning current. Simplifying assumptions 

have been made to some details of the configurations to focus on the 

main difference -- earthing practices. Identifying such differences 

provides the necessary perspective on their significance and the strong 

need to take them into consideration when developing international 

standards on surge-protective device applications. 

I. INTRODUCTION 

When designing a lightning protection scheme for a lowvoltage power system within a building, several scenarios must 

be considered for the point of termination of the lightning 

stroke. Common wisdom classifies these by decreasing order 

of severity: directly to the building, directly to overhead lowvoltage distribution lines (or other utilities) outside of the 

building, to other objects near the building, distant cloud-toearth strokes, and finally perhaps cloud-to-cloud discharges. 

Several standards-writing projects are underway, at the IEEE 

and at the IEC, based on present knowledge of the lightning 

flash characteristics and on assumptions about the way the 

lightning current divides among the many paths available for 

distributing (dispersing) this current to the ill-defined "earth" 

which is the termination of the cloud-to-earth strike. 

The purpose of our paper is to show the effect that different practices for neutral earthing in the low-voltage distribution 

system can have on the relative dispersion of the lightning 

current which is seeking the path of least impedance to earth. 

* Electricity Division, Electronics and Electrical Engineering Laboratory, Technology Administration, US. Department of Commerce. 

PE-968-PWRD-0-05-1997 A paper recommended and approved by the 

IEEE Surge Protective Devices Committee of the IEEE Power 

Engineering Society for publication in the IEEE Transactions on Power 

Delivery. Manuscript submitted January 2, 1997; made available for 

printing May 23, 1997. 

Franqois Martzloff, Life Fellow, IEEE 

National Institute of Standards and Technology* 

Gaithersburg MD 20899 USA 

To accomplish this purpose in an eight-page paper, and to 

concentrate on the essential difference, the models we present 

are simplified from the detailed reality, so that one of our first 

tasks will be to explain and justify the simplification. To avoid 

confusion in the meaning of the word "distribution" which can 

relate to the distribution of electric power by the utility or to the 

distribution of the lightning current among the available paths, 

we will use the term "dispersion" for the second meaning, 

lighting current dispersion. Another term used by some authors 

to convey the concept is "sharing" (among available paths). 

Note that the actual return stroke actually goes from earth to 

cloud in the majority of cases, but the scenario is generally 

described as if the stroke "t~erminated" on earthbound objects. 

In the case of a low-voltage power distribution system, 

different countries have adopted different practices on earthing 

the neutral conductor, and writing a history of why that is so 

would give an interesting insight into the development of power 

systems. The fact is thai today, two approaches are well 

entrenched in their respective territories, the so-called TN 

system and TT system where the difference lies in the mode of 

earthing the neutral. We will give a brief overview of the 

differences in ,a following section. Our purpose is to show how 

the difference in these practices affect the sharing, or 

dispersion, of the lightning current among the available paths to 

earth, and coiisequently affect the rating of surge-protective 

devices which may be included in these paths. We used the 

EMTP simulation code [I] to model several scenarios in each 

of the TN and TT systems, with small but possibly significant 

differences in the configuration. By postulating a direct stroke 

to one building, and requesting EMTP to compute currents in 

the (simplified) complete power system, we obtained results for 

the two most severe cases of lightning termination: the case of 

a direct stroke to one building, and the case of a nearby stroke 

which propag,ates and impinges at the service entrance of many 

buildings on that part of the low-voltage distribution system. 

The literature and draf standards contain many examples 

of such scenarios, but it seems that each is confined to a 

specific appr~oach or power system configuration with fairly 

detailed arrangements of load connections. The result is that 

from this plurality of examples, it is difficult to extract a clear 

perception of the significant parameters in the dispersion of the 

lightning current resulting from different earthing practices. In 

this paper, we: will simplify the scenarios to concentrate on the 

fundamental difference between the neutral earthing practices. 

0885-8977/98/$10.00 O 1997 IEEE 
IT. THE TT AND THE TN SYSTEMS 

The IEC has promulgated a letter code system describing 

the arrangement of the neutral earthing in single-phase and 

polyphase power systems [2]. For the purposes of our paper, 

we can summarize the TT system as being a distribution system 

where the neutral is earthed only at the distribution transformer 

secondary, and the protective earth in a building is obtained 

from a local earth electrode. This system is used in some 

countries. The TN system has its neutral earthed at any 

available opportunity outside of a building, including the 

distribution transformer secondary, some or all poles, and the 

service entrance. In the United States, an "Equipment 

Grounding Conductor" (EGC) is created at the service 

entrance, bonded to the incoming power system neutral and to 

the common local earthing point, after which the neutral 

conductor and the EGC are carefully (and by mandate from the 

National Electrical Code [3]) kept separate from one another. 

111. NECESSARY SIMPLIFICATIONS 

Another difficulty in making a detailed comparison of 

results Erom different authors is that different models are often 

used. When apparently different results are reported, a 

lingering question is that of differences attributable to the 

simplifying assumptions and possible modeling artifacts. We 

have used the well-known EMTP code [l] for which our 

previous experience in cross-validation between the computer 

model and full-scale experimental measurements [4], [5] gave 

us great confidence in the validity of the results. 

The literature offers many contributions on the system 

simulation but our purpose is not literature review -- again, our 

purpose is only to focus on the neutral practices considerations. 

However, to support some of our postulates, we will cite some 

papers to show that in the maze of assumptions, 

simplifications, and simulations, we are not alone. 

A. Down-conductor representation 

Some authors have included in their modeling a downconductor feeding the stroke current to the common bonding 

point of the building [6]. In our model, since we postulate that 

the current is delivered from a current source, the impedance of 

the down-conductor has no effect on the current being injected 

at the common bonding point which is the point at which 

dispersion (sharing) begins. Therefore, we did not include a 

down-conductor in our models. 

B. Earthing impedance as a function of time and current 

Some authors consider the fact that the exact value of the 

earthing impedance is variable as a function of time and current 

level. For instance, [6] initially proposes a model involving 

resistance, capacitance, and inductance, with some dependency 

on time or current, or both. But after studying the problem 

closer, the authors of [6] conclude that a reasonable approximation is merely a fixed 10-8 resistance. We have used this 

value in our models of the building earthing, and postulated an 

improved, lower 5-8 resistance at the earthing electrode of the 

distribution transformer. 

C. Other available current paths 

Some standard proposals include telephone, water and gas 

connections as possible paths for the earth-seeking lightning 

current. Considering that the telephone service is a balanced 

system normally isolated from earth (until a network interface 

device becomes involved), that some water and gas services can 

include a cathodic-protection isolation or be implemented with 

plastic pipes, we chose a conservative approach of not 

including these as additional paths to earth. 

D. Actual Circuit Configuration for Service Entrances 

Figure 1 shows a schematic of a single-phase 3-wire TN 

1201240-V service to a building. One surge-protective device 

(SPD) is connected between each of the two lines and the 

common earth at the service entrance, ignoring any SPDs within 

the building under the assumption that in a well-coordinated 

cascade [7] the majority of the current is carried by the service 

entrance SPD which has the lowest limiting voltage in the 

installation. The stroke current, postulated to have terminated 

on a point of the earthing system of the building, can seek a 

path to earth in two ways: directly through the earth electrode 

of the building, and by means of the three conductors back 

toward the power system. 

Figure 1. Service connections in a 3-wire TN system 

Figure 2 shows a schematic of a three-phase 4-wire TT 

2301400-V service to a building. A dedicated protective earth is 

created and connected to a local earth electrode, while the 

incoming neutral of the power distribution system is not bonded 

to this protective earth. At the service entrance, SPDs are 

connected between the local earth and each of the incoming 

lines and the neutral. 

Figure 2 - Service entrance connections in a 3-phase, 4-wire TT system 
E. Postulated lightning stroke current 

While some authors propose a 200 kA, 101350 ps surge 

[S], [9], others suggest that even a 100 kA peak might already 

be too high a value 161, [lo], [Ill. In agreement with the latter 

three references, we postulated a 100 kA peak, 1013 50 ps surge 

current. This selection also offers the convenience that when 

we report current levels in kiloamperes in the various circuits, 

the numbers also represent the percentage of the sharing, 

making it easier to follow the process. Since many standards 

for surges impinging on SPDs (at the service entrance) are still 

based on an 8/20 ps current waveform, we will also show one 

example of the energy deposition in the SPDs when such an 

8120 ps surge is postulated. 

The surge currents are modeled using the EMTP Type 60 

Slave Source. Using the "Freeform FORTRAN expression, 

any surge current waveform that can be expressed as a closedform equation can be used as signal source in the main EMTP 

program. The equations for the 101350 ys and 8120 ps waveforms with a 100 kA peak are respectively (1) and (2) below: 

10/350 ,us: I(t) = [IJq] [exp(-t/zJ - exp(-t/zJ] (1) 

where I, = 100 kA 

q = 0.9542 

z, = 480 

2, = 4 

8/20ps: I(t)=AI,fexp(-t/z) (2) 

where I, = 100 kA 

A = 0.01243 

z =3.911 

(In both equations, t and z's are in ps; I(@ is in same units as I,) 

F. Influence of Distribution Transformer Simplzj?cation 

The presence of distribution transformers has been 

included in many models in the literature, but their characteristics are not the same among authors. Some authors have used 

a coupled inductor with parasitic capacitor to represent the 

inter-winding capacitor in the transformer model [12]. 

While these models are more accurate in studying transformer failure modes due to low-side surges, for our main 

focus which is current dispersion among available paths, we 

have chosen the simple model postulated in [8] of a simple 

inductor to represent the winding. As results show, the 

presence of a transformer at the far end of a daisy-chain lowvoltage distribution system does not have considerable effect 

on the results. Therefore, we felt justified in adopting the same 

transformer model as described in [8] for all of our circuit 

configurations. 

G. Simplifying the Circuit for Modeling 

The circuit impedances have been modeled in EMTP using 

discrete components. The wiring between buildings and from 

building to transformer is modeled as a series inductance with 

the following parameters: R = 1 mO/m and L = 1 pWm, typical 

values for aluminum conductors of 34 mm2 cross section 

(#2 AWG) [13]. The SPDs are modeled using the EMTP Type 

92 Nonlinear Element model. Because of the simplified nature 

of the model, we performed parametric variations on factors 

such as line impedance and transformer inductance, and found 

that their influence on current dispersion is not large enough to 

warrant concern on the somewhat arbitrary values we have 

postulated in the baseline wenario. 

IV. MODIELING RESULTS 

In this section, we present selected results of EMTP runs 

for each of three TN or TT system configurations with points 

of lightning 'termination next to the distribution transformer 

("first" case) or at the opposite end of the transformer ("last" 

case), for a total of seven scenarios. We postulated a separation 

of 100 m between buildings and 20 m from the transformer. 

For each scenario, a pair of figures is given. The first 

figure of each pair is a schematic showing the configuration and 

point of stroke, together vvith indications of the peak current 

values in the circuit branches. The second figure of each pair 

shows selected current waveforms, generally currents leaving 

the house by way of the earthing electrode and the service 

conductors. Note that the peaks can occur at different times 

so that the sum of peak brimch currents shown on the figures, 

Kirkhoff notwithstanding, is not always exactly zero. 

A. TN-Radial, strike on one of the buildings 

A distribution transformer supplies three buildings in a 

radial arrangement where all the service drops originate at the 

pole where the transformer is installed (Figure 3). This 

configuration is a typical U.S. residential configuration. The 

lightning stroke is postullated to terminate on the earthing 

system of one of the three buildings. Figure 4 shows the 

current waveforms. 

"Peak: occurs very 

late In event 

I 

A - 

1 

Figure 3 - Radial TN configuration with three buildings supplied by one 

distribution transformer, one building struck by a 101350 ps, 100 kA surge, 

showing peak values of currents shared among available paths. 
I Time (vs) I 

SPD - Current into each line of service drop, through SPDs 

GND - Current into local building earth electrode 

Nout - Current into neutral conductor of service drop 

Figure 4 - Waveforms of currents leaving Building 1, as defined in Figure 3, 

for a 100 kA, 101350 ps surge terminating on the building earthing system 

B. TN-Daisy clzain, strike on first building 

Another typical arrangement uses a distribution 

transformer which supplies several buildings along a street, 

with short service drops from the poles to each building. The 

lightning stroke is postulated to occur upon the first building, 

next to the transformer. Figure 5 shows the circuit 

configuration and the peak currents in the branches; Figure 6 

shows the waveforms of the currents leaving the building. 

Note the early peak of the current in the neutral -- directly 

connected to earth at the pole, thus a lower inductance 

compared with the inductance of the line conductors that 

include the transformer winding. 

Buildlng 1 Buildlng 2 Buildlng 3 -* - 20 m 100 m - - 100 rn * 

Figure 5 - Daisy chain TN configuration with building next to transformer 

struck by a 101350 ps, 100 kA surge, showing peak values of currents 

C. TN-Daisy clzain, strike on last building 

This is the same configuration as B, but the building being 

struck is at the opposite end (Figure 7). The difference, if any, 

would give insight on the relative importance of modeling the 

presence of a specific transformer. In fact, the difference in the 

SPD stress for a strike on the first building (20 kA) compared 

with a strike on the last building (26 kA) is small, showing the 

small effect of transformer position. In the building earthing, 

where there are no SPDs, a strike on the last building produces 

42 kA compared with 23 kA for a strike on the first building. 

- - 

SPD -Current lnto each lme of servlce drop, through SPDs 

GND - Current mto local bulldlng earth electrode 

Nout - Current lnto neutral conductor toward the transformer earth 

Flgure 6 - Waveforms of currents leavmg Bu~ld~ng 1, as defined ~n F~gure 5, for 

a 100 k.4, 101350 ps surge terminating& the building earthing system 

' Peak occurs early In event 

transformer struck by a 101350 ps, 100 kA surge, showing peak currents 

0 zw - Xx) 

Time (ps) 

SPD - Current into each line of service drop, through SPDs 

GND - Current into local building earth electrode 

Nout - Current into neutral conductor toward the transformer earth 

Figure 8 - Waveforms of currents leaving Building 3, as defined in Figure 7, 

for a 100 kA, 101350 ps surge terminating on the building earthing system 

For Figures 7-8, the greater distance (inductance) from the 

transformer earth electrode forces initially more current flow in 

the building earth than in Figures 5-6 for a closer transformer. 
D. TT Zwire, strike on first building 

A transformer (single-phase or one phase of a three-phase 

transformer) supplies several buildings along a street, with 

short service drops from the street poles to each building. The 

lightning stroke is postulated to occur upon the building next 

to the transformer (Figure 9). The waveforms of the currents 

leaving the building are shown in Figure 10. 

Buildlng 1 Buildlng 2 Building 3 -* - - 20 rn loom - loom * 

10 kA ID kA 9.5 lc4 I T T I 

* Peak occurs early in the event I L 

Figure 9 - Daisy chain TT 2-wire configuration with building next to 

distribution transformer struck by a 101350 ps, 100 kA surge 

Time (ps) 

N-E SPD - Current through neutral-to-earth SPD 

L-E SPD - Current through line-to-earth SPD 

GND - Current into building earthing electrode 

Figure 10 - Waveforms of currents leaving Building 1, as defined in Figure 9, 

for a 100 kA, 101350 ps surge terminating on the building earthing system 

E. TT 2-wire, strike on last building 

The configuration is the same as in D, but the lightning 

stroke is postulated to strike the building at the opposite end of 

distribution line, away from the transformer (Figure 11). 

Figure 12 shows the waveforms of the currents leaving 

building 3. 

Where end-users are provided with three-phase service, a 

three-phase transformer supplies several buildings along a 

street, with short service drops from the street poles to each 

building. In this configuration, the difference from a 2-wire, 

single-phase service is that four conductors instead of two are 

available as exit paths for the lightning current postulated to 

have stmck the building of interest (first or last building). 

To conserve space, we do not present two pairs of figures 

for that configuration, but the summary of Table 1 includes the 

current values computed by EMTP for the two scenarios in that 

configuration. 

Bulldlng 1 Building 2 Building 3 ---+-----+< 20 rn 100 rn 100 m 

IT. 

Q kA 28 W 38 1cA 

100 kA 

I "Peak occurs late In the event -t 

Figure 11 - Daisy chain TT 2-wire configuration with building at opposite end 

of distribution transformer struck by a 101350 ps, 100 kA surge 

N-E SPD -Current through neutral-to-earth SPD 

L-E SPD - Current through line-to-earth SPD 

GND - Current into building earthing electrode 

Figure 12 - Wavt:fonns of current; leaving Building 3, as defined in Figure 11, 

for a 100 kA, 101350 ps surge terminating on the building earthing system 

G. Compariwn of the seven scenarios 

Results of our model runs for the seven scenarios (Table 1) 

show that, contrary to some speculations or intuitive 

considerations on the sharing among service conductors, the 

earthing connection of the building does not carry anywhere 

near the 50% quoted in some proposed standards [9]. 

The most severe stress, for the parameters postulated, 

occcurs in the neutral SPD in Scenario D (TT 2-wire, first 

building struck) for whicZl the configuration has the lowest 

impedance to earth and thus invites the largest share. Other 

scenarios generally reflect ]primarily the number of service-drop 

wires available for the current exit. 
TABLE 1 

SUMMARY OF CURRENT SHARING AMONG CONDUCTORS FOR THREE CONFIGURATIONS IN SEVEN SCENARIOS FOR 100 kA STROKE 

Scenario: Most severe but rare - Building being directly struck Less severe but more frequent 

Configuration: Building See Currents leaving building via building earthing Currents impinging onto 

Distribution being figures and service conductors (peak kA or %) * adjacent buildings (peak kA or %) 

system struck -- 

-- Building Service SPD in SPD in SPD in SPD in 

earthing neutral the neutral the lines the neutral the lines 

TN Radial Any 3 -4 2 1 33 NIA 23x2 NIA 10 x 2 

TN Daisy First 5-6 23 27 NIA 20 x 2 NIA 7x2 

TN Daisy Last 7-8 42 26 NIA 26x2 NI A 8x2 

TT 2-wire First 9-10 26 N/A 5 1 28 x 1 10 lox 1 

TT 2-wire Last 11-12 48 NIA 3 8 38x 1 13 13x1 

TT 4-wire Fmt -- 22 NIA 32 16 x 3 5 5x3 

TT 4-wire Last -- 3 8 NI A 20 20 x 3 6 5x3 

* Peak values do not occur at the same time in the different paths so that totals of numbers shown may be more than the impmging 100 kA peak 

V. ENERGY CONSIDERATIONS 

In the model parameters, to start the iterative process, we 

have postulated that the SPD consists of a metal-oxide varistor 

(MOV) with relatively large cross-section that might be capable 

of absorbing the energy involved in diverting the 101350 ps 

surge. For the TN configurations, we selected a 150 V rms 

rating, and a 300 V rms rating for the TT configurations. For 

the cross section, we postulated an area equal to ten 20-mm 

discs in parallel because available manufacturer's data [14] 

readily gives the 20-mm disc characteristic. Such a combination 

would have a total one-shot joule rating of 800 joules for a 

1011000 ps surge in the 150 V rating. 

Because we suspect that even this array of ten discs might 

not be capable of dissipating the energy involved in a 101350 ps 

surge, the next step in this iterative process is then to compute 

the energy that would be deposited in the SPDs, under the 

current distribution patterns computed in the seven scenarios. 

As one example, Table 2 shows the energy deposited in the 

MOVs, computed for the case of the TN Radial configuration 

where one SPD is connected between each of the two lines and 

the earth point of the installation (Figure 3). Two waveforms 

are shown in the table, the 101350 ys and the 8/20 ps surges. 

TABLE 2 

ENERGY DEPOSITION IN SERVICE ENTRANCE MOVS 

FOR THE TN RADIAL CONFIGURATION AND TWO WAVEFORMS 

Rating for ten Energy deposition Waveform 

20-mm discs Direct strike Nearby strike 

101350 ps 800 J 3500 J 840 J 

8/20 !AS 800 J 200 J 805 

For the 101350 ps waveform, the rare scenario of a direct 

strike (energy deposited is 3500 J) would require a very large 

varistor at the service entrance -- four times the ten discs we 

postulated, while this ten-disc array would be sufficient in the 

less rare scenario of a nearby strike (840 J). 

On the other hand, if we were to stay with the 8120 ps as a 

postulated waveform, even the large 100 kA peak would be 

handled with comfortable margin by the ten-disc array. These 

results provide quantitative data which we will discuss further 

under the CostlRisk heading. 

VI. PARAMETRIC VARIATIONS 

We performed several parametric variations for the purpose 

of exploring the typical "what if' questions, and also to show 

whether or not our postulated values might be viewed as too 

arbitrary because of their influence on the results. 

A. Line impedance and building separation 

The value of 1 pHlm for conductors has long been used by 

many researchers as a typical value. To investigate the 

significance of that postulated 1 pH1m combined with the 100- 

m separation, we ran two cases, one with half the value and one 

with double the value. The first case corresponds to either half 

the separation for the same unit impedance or half the unit 

impedance with the same separation. For the second case, one 

of the parameters is doubled while the other is held constant. 

Table 3 shows a comparison of the baseline case with these two 

parametric variations. 

TABLE 3 

EFFECT OF LINE IMPEDANCE ON SHARING - TN DAISY LAST 

Baseline Half Double 

Percent of 100 kA peak 100 m, 1 uHlm baseline baseline 

Current into building earthing 42 32 53 

Current in service neutral 26 27 25 

Current in SPDs 26 26 25 

This comparison shows no significant differences in the 

current sharing for each of the three available conductors (there 

are two line conductors, each with an SPD) when the postulated 

unit impedance or building separation is varied over a 1:4 

range, so that our selection for these two parameters should not 

be a matter of concern. 

B. Transformer pole earthing resistance and building 

earthing system resistance 

By their relationship, these parameters can be expected to 

have an influence on the outcome. In the baseline case, we 

postulated a 5-Q pole earthing resistance and a 10-8 building 

earthing resistance. Table 4 shows the comparison of the 

baseline case with the reversed relationship between the pole 

earthing resistance and the building earthing resistance. 
TABLE 4 

EFFECT OF POLE EARTHINGBUILDING EARTHING - TN RADIAL pp - - -- 

Percent of 100 kA peak Baselme Reverse baseline 

5-Q pole, 10-Q bldg. 10-Q pole, 5-Q bldg. 

Current in building earth 2 1 3 1 

Current in serv~ce earth 33 14 

Current in SPD 23 22 

Indeed, the relationship of pole versus building earthing 

resistance has a significant effect on the current carried by the 

neutral, but not on the current carried by the SPDs. This is 

particularly true, although not obvious in the table (where only 

the peak values are shown, reflecting the inductive effect on 

initial current dispersion), for the tail of the 101350 ps 

waveform where the subsequent sharing is determined by the 

resistance ratios [6], [12]. 

C. Length of circuit (more buildings along a street) 

Postulating a greater number of buildings along the daisy 

chain, while keeping the resistance of the building earthing 

constant, can be expected to offer a path of lesser impedance to 

the currents exiting the building, because of the greater number 

of available earth electrodes. Table 5 shows the effect of going 

from 3 buildings (baseline) to 9 buildings, still with the last 

building being struck. 

TABLE 5 

EFFECT OF NUMBER OF BUILDINGS IN TN CIRCUIT 

ON SHARING, DISTANT HOUSE STRUCK 

Percent of 100 kA peak Baseline 3 buildings 9 buildings 

Current in building earth 42 42 

Current in service earth 39 14 

Current in SPD 26 27 

Again in this case, a difference is noticeable in the neutral 

conductor current, but not in the SPD current. Thus, this parametric variation shows that the number of buildings between the 

building being struck and the distribution transformer, while 

affecting the neutral current, does not affect the stress imposed 

on the SPDs in this TN configuration. 

VII. DISCUSSION 

A. Effect of posiulaied waveform 

While we have adopted for our baseline the 101350 ps 

waveform, many SPD standards cite an 8/20 ys or a 4/10 ps 

surge waveform as an SPD capability requirement [15], [16] or 

as a surge environment description [17]. To explore the effect 

on sharing of the stroke current with different waveforms, in 

particular during the initial part of the 101350 ps surge where 

inductive effects dominate for the circuit parameters selected, 

we made one run with a 8120 ys surge instead of the 101350 ps 

used in the baseline case of the TN Radial. Predictably, given 

the small difference between a 10 ps and an 8 ys rise time, little 

effect was noted in the sharing during the first 20 ys. Of 

course, the energy involved for the total surge duration is 

another matter, already discussed in Section V. 

B. Selection of SPDs 

If the design objective is to provide protection for a direct 

stroke to the building (a topic which will be the subject of the 

costhisk analysis mentioned below), the SPDs must be selected 

with sufficient current-handling capability to survive the surges 

resulting from the postulated surge. 

Alternate proposals have been made to use a spark gap as 

service entrance SPD. Such a gap must then be capable of 

clearing the resulting follow current, which may be an issue for 

systems having a large available fault current, such as the 10 kA 

rms specified For U.S. installations [3], [la]. 

We have ]made one run with a spark gap model instead of 

a varistor model. From the sharing point of view, the difference 

is small, which can be readily explained by the fact that 

inserting in the lightning current paths a varistor with a limiting 

voltage of 1 kV or so, or a gap with an arc voltage of 100 V or 

so, should have a very small effect on the sharing because of 

the many kilovolts developed by the lightning current flowing 

in the inductances and resistances of the line conductors and 

earthing connections. 

An essential aspect of designing an effective surge 

protection system is to perform a costlrisk analysis involving 

the probability of a building being struck by a large surge, such 

as 100 kA, versus the cost of ensuring survival of the service 

entrance SPDs to be installed. This analysis introduces factors 

such as the flash density in the locale, the randomness of the 

distribution of the flashes over the area of attraction of the 

building which depends in part on the height of the building, 

and the distribution of peak amplitudes of lightning strokes. 

For instance, [l I] reports statistics [19] whereby an 80 kA 

amplitude is exceeded for less than 5% of the strokes. Such an 

analysis is beyond the scope of our paper, but it must be 

mentioned here to keep the situation in perspective and remind 

developers of SPD application guides or standards to include it 

in their recommendations. 

D. Applying-field experience lo standards 

The ultimate test of the usefulness of a standard is that 

equipment manufactured according to that standard has 

satisfactory fi~ald experience, while being produced at a cost that 

users are willing to accept. Very low field failure rates can be 

seen as overdiesign, high failure rates obviously as underdesign. 

It is the dream of one of the authors to establish a clearinghouse 

where field experience of manufacturers could be collected and 

applied to optimize the definition of the environmental stress 

[20]. Given the competitive nature of the industry, this is likely 

to remain only a dream. However, many U.S. utilities are now 

offering to their customers the installation of a meter-base 

adapter SPD. The field experience for these SPDs might be 

collected from utilities ..- with safeguards on proprietary 

information -- and become an input to the process of 

moderating slome proposals for high-stress requirements, on the 

basis of the successful field experience of SPDs with 

capabilities below those irnplied in proposed standards. 
VIII. CONCLUSlONS [4] Martzloff, F.D., Mansoor, A., Phipps, K.O., and Grady, W.M., "Surging 

the Upside-Down House: Measurements and Modeling Results," 

Modeling several typical TN and TT configurations of Proceedings, EPRI PQA'95 Conference, 1995. 

neutral earthing practices and scenarios of lightning strike point [5] Mansoor, A. and Martzloff, F.D., "Driving High Surge Currents into 

provides insights on significant effects, which should lead to Long Cables: More Begets Less," Paper 96 SM399-6 PWRD, IEEE PES 

more effective application of surge-protective devices (SPDs). Summer Meeting, Denver CO, 1996. 

[6] Rakotomalala, A,, Auriol, Ph., and Rousseau, A,, "Lightning Distribution 

A direct lightning stroke to a building can produce high 

stresses on the service entrance SPDs as the earth-seeking 

current will exit in part by way of the utility service drop. 

SPDs in that building will be strongly affected, while nearby 

buildings will be impacted by much lower surge currents. 

The major difference among the scenarios we have modeled 

appears in the current carried away from the building by the 

neutral conductor. 

In a TN system where the neutral is bonded to earth at the 

service entrance, there is no SPD in that path, and thus no 

concern about neutral SPD integrity. In typical residential 

single-phase U.S. systems, the line SPDs can carry about 

25% of the stroke current. 

In a TT system where there is an SPD in the neutral path, 

a single-phase two-wire configuration can have 50% of 

the stroke current being carried by the neutral SPD. In a 

three-phase TT system where there are four conductors to 

carry away the stroke current, the neutral SPD can carry 

up to 30% of the stroke current. 

For line conductors, the difference reflects primarily the 

total number of conductors in the system, which can be two, 

three or four. The earth-seeking lightning current will divide 

(but not always equally) among these conductors. While the 

initial dispersion (during the first 20 ps) is controlled by the 

inductances, the later dispersion is controlled by the relative 

values of the earthing resistances. 

If the postulated stroke is as high as some of the proposed 

standards suggest, modeling the behavior of service entrance 

SPDs of the type installed in increasing numbers by U.S. 

utilities shows that some failures could be expected. As 

field experience seems to indicate an acceptable failure (if 

any) rate, one can question the need for imposing such 

severe requirements, unless the mission of the,facility is 

such that even a rare failure would be unacceptable. 

IX. ACKNOWLEDGMENTS 

Support and encouragement for this work was provided by 

the parent organization of each author. Additional support was 

received from Delmarva Power Company, from Pacific Gas & 

Electric Company, and from PECO Energy Company. Gerald 

FitzPatrick and Roger Witt contributed comments on the draft. 

X. REFERENCES 

[I] EPRI, "Electromagnetic Transient Program (EMTP), Version 2.0; Volume 

1: Main Program: Volume 2: Auxiliary Routines, EPRIReport EL-6421-L, 

July 1989. 

[2] IEC International Standard 364-1, Electrical installations ofBuildings - 

Part I: Scope, Object and Fundamental Principles, 1992. 

[3] NFPA 70, Natzonal Electrical Code, 1996 

Through Earthing System." Symposium Record, IEEE International 

Symposium on EMC, Chicago IL, USA, August 1994, pp 419-423. 

[7] Lai, J.S. and Martzloff, F.D., "Coordinating Cascaded Surge Protection 

Devices: H~gh-Low versus Low-High," IEEE Transactions IA-29 No.4, 

JulyIAugust 1993, pp 680-687. 

[8] Birkl, J., Hasse, P., and Zahlmann, P., "Systemgerechter Einsatz von 

Ableitern in Niederspannungsnetzen," Elektrotechnische Zeitschrift 

N0.17, 1994, pp 964-971 

[9] Flisowski, Z., and Mazzetti, C., "Efficiency of Lightning Protection 

System as Screening Measure Against LEMP," Proceedings, EMC'96 

ROMA, September 1996. 

[lo] Darveniza, M., Sargent, Limbourn, Liew, Caldwell, Currie, Holcombe, 

Stillman & Frowd, "Modelling for Lightning Performance Calculations" 

IEEE Transactions PAS-98, No.6, NovIDec 1979, pp 1900-1908. 

[l I] Kirder, E.P. and Uman, M.A., "Cloud-to-Ground Lightning Mechanisms 

of Damage and Methods of Protection," Seminars in Neurology, Vol. 15, 

No.3, September 1995. 

[12] Dugan, R.C. and Smith, S.D., "Low-Voltage-Side Current-Surge 

Phenomena in a Single-Phase Distribution Transformer System," IEEE 

Transactmns PWRD-13, No. 2, April 1988. 

[13] Steigenvald, R.L., Ferraro, A,, and Tompkins, R.E, "Investigation of a Family of Power Conditioners Integrated into the Utility Grid," Sandia 

ReportSAND81-7031, 1981, page 20. 

[14] Transient Voltage Suppression Devices, Harris Corp., 1991. 

[15] ANSILEEE C62.11-1991, IEEE Standard for Metal-Oxide Surge 

Arresters for AC Power Circuits. 

[16] ANSIIIEEE C62.34-1996, IEEE Standard for Performance of LowVoltage Surge-Protective Devices (Secondary Arresters). 

[17] ANSILEEE C62.4 1 - 1991, IEEE Recommended Practice on Surge 

Voltages in Low-Voltage AC Power Circuits. 

[18] UL Std. 1449, Transient Voltage Surge Suppressors, 1996 

[19] Berger, K., Anderson, R.B., and Kroninger, H., "Parameters of Lightning 

Flashes," Electra, 1975-80, pp 1548-1565 

[20] Martzloff, F.D., "Keeping up with the Reality of Today's Surge 

Environment," Power Quality Solutions, September 1995 Proceedings, 

pp 243-249. 

Arshad Mansoor (M' 1995) is an 

Electrical Systems Engineer at the EPRI 

Power Electronics Applications Center 

(PEAC). He received his MS and Ph.D. in 

electrical engineering from the University of 

Texas, Austin in 1992 and 1994 

respectively. His areas of interest include 

Power Quality, power systems transients 

analysis, harmonics, surge propagation and 

protection, and EMTP model development. 

Franqois Martzloff (M' 1956, F' 1983) 

Born and educated in France, with 

additional MS degrees from Georgia Tech 

and Union College, worked at General 

Electric for 29 years and now 12 years at the 

National Institute of Standards and 

Technology. He is contributing to several 

committees for the development of standards 

on EMC, surge protection and Power 

Quality in the IEEE and the IEC. 
Discussion 

P. Hasse and J. Birkl (Dehn + Sohne, 92318 Neumarkt, 

Germany): The problem of lightning current distribution 

depending on the different kinds of current distribution 

systems and under the influence of possible variables has been 

very clearly represented with this contribution. 

The curve development gained by the EMTP program 

matches very well with the simulation calculations conducted 

in Germany with the PSPICE program. 

In respect to the described results, however, a few additions 

are necessary from our point of view: 

1. Section 111. G. and VI. A.: 

The induction of a conductor system always results out of 

the geometric system of the slip-knot consisting of coming 

and going conductor. A separation in coming conductor 

impedance and going conductor impedance is not realistic. 

In particular, in case of multiple conductor cables it is to be 

observed that in case of the same flow direction of the 

lightning current, the inductivity of the total system 

differentiates to a single conductor system. 

2. Section VI. Schedules 3 - 5: 

In particular, in case of longer connecting cables between 

buildings and between building and transformer a change 

in waveform of the surge flowing through these cables. 

Only the observance of the amplitude factor of the flowing 

lightning currents is not sufficient. In this situation, it 

would be more meaningful to consider also the energy 

distribution. 

Section VII. B.: 

For decades now, in Europe, spark gap arresters, with a 

mains follow current quenching capability, are being 

installed successfully as lightning current arresters at the 

building entrance. In particular, the high down-lead ability 

and impulse-time shortening of the rest impulse make a 

favourable co-ordination with connecting MOV's possible. 

Section VII. 4: 

The lightning protection necessity for a system, as well as 

the deduced lightning protection class resulting from this, 

is described in IEC 1662. At the same time the lightning 

protection class is determined, the layout of the lightning 

protection system necessary lightning parameters are 

defined (IEC 13 12-1). A deviating layout of protection 

measures on the basis of test currents 8/20 ps is therefore 

not permissible. 

Manuscript received November 3, 1997. 

Frangois D. Martzloff (National Institute of Standards and 

Technology, Gaithersburg MD) : 

We thank the two authors of the discussion for taking the time 

to review our ;paper and provide comments aimed at broadening 

the consensus on the subject. In particular, we are delighted to 

hear that our computations based on EMTP matches very well 

with the simulation calculations conducted in Germany with the 

PSPICE program. With respect to their specific four comments, 

we offer the following responses, preceded by the general 

remark that the purpose of our computations was to reveal the 

differences among various postulates for the circuit configurations, as influenced by the grounding practices for the neutral 

in effect in different countries, rather than the precise values for 

a particular set of parameters. We emphasize the concept of 

postulate, lest we fall into the trap of taking electromagnetic 

environment standards as an exact duplication of reality, while 

they are in fact only the documentation of an industry consensus 

on how reality might be represented '. 

1. Section I1 G and IV A 

Indeed, the concept of inductance is based on a conductive 

loop that carries the current in a closed circuit. However, in the 

circuits we postulated for our computations, the conductors in 

question - phase and neutral - may be considered as one part 

of the closed circuit and might be called "coming," according to 

the tenninolcgy used by our colleagues, while the path consisting of the earth, the distant return to the cloud, the lightning 

channel, and [even the down-conductor (see Section I11 A) may 

be considered as the other part of the loop and might be called 

"going" conductors. 

For this reason, we represented in our figures the phase and 

neutral as if they were separate, while in reality they can be at 

some finite distance from each other (the so-called "open 

wiring" used jn overhead lines) or in close vicinity, as in the case 

of an underground cable or an overhead "triplex." Aware of the 

differences, we started our computations for a given, postulated 

configuration - always the same for the variations in the neutral 

grounding -. but performed a parametric variation in the line 

inductance (taking twice or half the value used in the baseline), 

as stated in the subject paragraph, to convince ourselves that the 

influence on dispersion is not large enough to cause concern. 

Space limitations for the paper prevented us from providing 

detailed numerical results - as they also do here - and we 

were hoping that our simple statement that we did consider the 

issue and found little effect on the differences among neutral 

grounding scenarios might be acceptable. 

2. Section VI 3 to 5 

One of the results of our computations based on a postulated 

101350 ps waveform was to show that, for the distances we 

selected, the impedance of the cable between buildings - and 

therefore their length - has only a small influence on the longterm current waveform and dispersion among conductors, which 

is primarily influenced by the postulated values for respective 

earthing resistances. With the values selected for inductances, 
the current dispersion is substantially affected by the respective 

inductances only for the first 20 or 30 ps. 

We agree that additional information might be conveyed by 

reporting the energy distribution along the complex path of the 

lightning current, but here again space limitations intervene. We 

can offer the response, however, that in view of the large values 

of the earthing resistance compared to the other resistances in 

the circuit - cable resistances and dynamic "resistances" of the 

varistor or gap SPDs - the latter are not a priority in reporting 

results. The EMTP model of course has the capability of 

reporting any set of parameters if "asked" to do so. 

For specific applications of one type or another of SPD 

technology, the EMTP model can provide detailed information 

on the energy that will be deposited in these SPDs for the 

various scenarios to be considered. 

Section VII B 

We are aware that in some countries, the installation of a 

service-entrance arrester is a common practice, and that gapped 

arresters may be used for that purpose. The issue is one of cost 

vs. benefits for an arrester designed for the large lightning 

currents associated with a rare direct strike to the building. We 

have observed, during our interactions with several international 

or IEEE technical committees, that consensus has not been 

reached on what current waveform and peak amplitudes should 

be considered when making the cost vs. benefits analysis. 

Depending on the nature of the installation, the cost vs. benefits 

equations are different. Several proposals for "risk analysis" are 

currently under consideration in several standards-developing 

bodies, and consensus is clearly not achieved at this point. This 

lingering question is addressed in our response to the fourth and 

last comment after the present one. 

Our intention in making the remark on available fault 

current in the second paragraph of this section was not to contest 

the successful European experience cited by our colleagues, but 

to alert our readers at large to the importance of considering that 

requirement. The point that mains follow-current quenching 

capability is not trivial was confirmed in a comment by one of 

the reviewers of our forthcoming paper, "Gapped Arresters 

Revisited" (scheduled for presentation at the IEEE-PES Winter 

1998 Meeting and later publication in IEEE Transactions). 

Section VII 4 

We are aware of the work conducted in the IEC Technical 

Committee 8 1, the responsible body for development of the IEC 

61662 and IEC 61312 publications. We are also aware of some 

discomfort among other parties concerning the stipulations from 

that body which might result in less than fully cost-effective 

solutions to the question of real necessity for protection against 

worst-case scenarios. The footnote offered in support of our 

introductory remark applies here also. There is a long and 

successful history of application of surge-protective devices 

based on a postulated 8/20 ps surge current waveform, using the 

appropriate values of amplitudes. For that reason, we included 

in our paper as alternate postulate the 8/20 ps waveform. From 

the point of view of EC TC 81, their recommendations might be 

considered normative and thus non-negotiable, but protection 

measures in the various countries are typically determined - if 

at all - by bodies that promulgate codes based on a consensus 

drawn from experience based not exclusively on TC 81 recommendations. Therefore, the use of the term "not permissible" 

appears somewhat strong in the context of voluntary or even 

regulatory practices. 

In conclusion, we appreciate the opportunity to present more 

detailed background information on our computations and 

underlying postulates, thanks to the discussion contributed by 

our colleagues. 

1. Long ago, my mentor, Frank Fisher, taught me this concept 

which I recite in the following terms, well worth repeating in the 

present context: "The criterion of validity of an environment 

standard is not so much how closely it duplicates reality but 

rather how well equipment designed in accordance with this 

standard pe~orm in the field. If equipment designed in 

accordance with the standard pedorm well in the field, while 

equipment ignoring the standard do not perform well, the 

chances that the standard be a good standard are pretty good." 

Manuscript received January 7, 1998. 
ON THE

DISPERSION

OF 

LIGHTNING

CURRENT

FOR A DIRECT

FLASH TO

A BUILDING
Reprinted with permission from Proceedings, 25th International Conference on Lightning Protection, Rhodes, 2000

Significance:

Part 2 – Development of Standards – Reality checks

Part 4 – Propagation and coupling of surges

Comparison between two simplified modeling studies of the dispersion and a documented case of the complexity o

a direct flash to a residence.

Reservations on the justification of very high stress requirements for SPDs are expressed in a discussion, followed

by a proposal to encourage more information sharing on the subject.

Filename: Dispersion
25th International Conference on Lightning Protection 5.

18-22 September 2000 

ON THE DISPERSION OF LIGHTNING CURRENT 

AFTER A DIRECT FLASH TO A BUILDING 

Francois Martzloff 

f.martzloff@ieee.org 

National Institute of Standards and Technology* 

U.S.A. 

Abstract: Simplifications often made when simulating 

the dispersion of lightning current can yield results 

that depart too much from the complex reality of a 

direct lightning flash to a building. The unpredictable 

occurrence of side-flashes increases even more the 

complexity. Such simulations, if taken at face value, 

might lead to unrealistic specifications for serviceentrance surge-protective devices (SPDs). A real-world 

anecdote illustrates both the complexity and a case 

where an SPD with only modest ratings, compared 

wiih some present proposak, pyovided satisfactory 

protection on the power-port apphances of a residence. 

Key Words: Lightning current dispersion, sideflashes, surge-protective devices 

1. INTRODUCTION 

Recent discussions among participants in the 

development of surge-protection standards have shown a 

lack of consensus on-the~possiblescenarios concerningthe 

dispersion of the lightning current when a direct flash to a 

building is involved. Skepticism has been expressed both 

on the simulation of available paths for the dispersion, and 

on the magnitudes and waveforms postulated for the 

resulting currents flowing in the conductive elements of the 

installation - especially the service entrance surgeprotective devices (SPDs). Note that the skepticism does 

not aim at the parameters of the lightning flash itself, which 

have been accepted now for many years [I], but at the 

simulation conclusions concerning requirements for 

service-entrance SPDs with very high current-handling 

capability. 

Another objection has been that the distinction 

between a building equipped with a lightning protection 

system (air terminals, down-conductors and earthing 

system), on the one hand, and a building without such a 

system, on the other hand, might be misleading. 

Every building that contains electrical circuits (po

or communications), electrically-conducting mechan

elements, metallic structures, etc., has a de facto light

'protection' system of intended or unintended air term

and down-conductors - except that their connection to 

earthing system might have unpredictable and unwa

side effects. 

As an input toward developing consensus, this p

reviews in a first part two examples of simulations that 

been performed by others, and in a second part relat

real-world anecdote of a corroborated case of a direct f

to a residence. 

2. SIMULATING DISPERSION 

2.1 Examples of simulations 

A Joint Working Group of the International Ele

technical Commission (IEC) has recently develop

Technical Repad on surge pmntp,ctian[2]. This &-ye

effort involved the participation of five ZEC Tech

Committees interested in the subject. The data 

considered by the group included, among many sou

two published papers, identified in the Bibliography o

report, authored independently by members of the gro

In both studies, a 10f350 ps waveform was postul

and a time-invariant earthing resistance and induct

were postulated. Currents in the available paths to e

and voltages at selected points of the systems 

computed. For the purpose of this paper, three figures 

are reproduced here for a qualitative glimpse on the res

Figure 1, simplified from Ref [2], shows the natu

circuit components and configuration: two buildings an

distribution transformer linked by a cable in a l

arrangement. The point of strike is Building 1. The det

numerical values, which are given in the referenced p

are not significant for this comparison of the two stud

* Electricity Division, Electronics and Electrical Engineering Laboratory, Technology Administration, U.S. Department of Comm

Contributionsfrom the National Institute of Standards and Technology are not subject to U.S. Copyright. 

ICLP zooo 4
Figure 2 shows examples of the computed waveforms 

for currents at selected points of the Figure 1 circuit, from 

top to bottom : the postulated stroke current, the current 

exiting Building 1 via the power supply cable, the current 

in the earthing impedance of the building, and the current 

in the service-entrance SPDs of Building 2 resulting from 

the surge that is now, for Building 2, an impinging surge. 

The radial service drops also consist of a resistan

an inductance, not drawn in the figure, but modeled 

computation according to the 20-m length of each 

drop. The point of strike of the flash is the earthing s

of Building 1 (to which the neutral is bonded). 

Figure 2 - Current dispersion for Figure 1 

The numerical values are not significant, but the 

waveforms are. The current exiting & building (IwlM) 

has the same waveform as that of the stroke. The earthmg 

current (Iearthing) has an initial peak, due to the additional 

inductance of the power supply cable; in the long term, the 

inductive effect disappears, and the current division simply 

reflects relative values of the available earthing resistances. 

In contrast with Figure 1 where the buildings are strung 

along a power supply cable, Figure 3, from Ref [3], shows 

a radial configuration of three buildings, each supplied by 

its own service drop, with all three connected directly to the 

terminals of the common distribution transformer. 

Varistor-type service-entrance SPDs are provided for each 

building. The transformer and each building have their 

own earthing electrode connection, represented by a fixed 

resistance and an inductance. 

Figure 3 - Buildings configuration and resulti

currents, according to Mansoor (Ref [3]) 

The significance of citing these two indepe

studies, separated by an ocean, by two different lang

used by the authors, and by two different simu

programs, is that quite compatible conclusions 

reached after exchanging information, first across the 

during IEC working group meetings, then later on

intensive e-mail messages, as a working relatio

blossomed in spite of logistics barriers. This impo

point will be elaborated further in Section 7 of this p
With hindsight, it appears predictable that the initial 

current dispersion (the first ten microseconds) should be 

strongly influenced by the relative values of the postulated 

inductances, and the later dispersion by the relative values 

of the postulated (time-invariant) earthing resistances. 

2.2 Involvement of service-entrance SPDs 

The results of these simulations show that the stresses 

imposed on the service-entrance SPDs that become 

involved in dispersing the part of the lightning current 

toward remote earthing electrodes of the power distribution 

system will reflect the relative values of the earthing 

impedances. These stresses also vary with the postulated 

current waveform, ranging from the short 8/20 ps long-used 

for designing SPDs, to the more recent proposed 101350 ps 

and finally to the seldom-considered continuing current 

within a multiple-stroke flash. 

Performing these simulations was not a futile exercise 

but provided insight on the influence of significant 

parameters. However, among end-users, this complexity 

of postulates decreases the credibility of defining SPD 

requirements on the basis of simulations, giving a greater 

credibility to field experience of widely-used SPDs that 

have demonstrated satisfactory performance over many 

years. 

3. SIDE FLASH 

One event that contributes to the complexity and 

uncertainty of lightning current dispersion is the possible 

occurrence of a side-flash. A side-flash can establish 

unexpected paths to earth, with two consequences that 

extend beyond the consideration of service-entrance SPD 

stresses - the motivating concern for this paper. 

The side-flash itself can have hazardous consequences 

by acting as an igniter, as will be told in the anecdote 

of Section 4. 

The side-flash can cause currents to flow along 

conductive paths within the installation, thereby 

coupling transient overvoltages in the circuits of the 

installation, by common path or by induction. 

Of course, the latter has the same end-result as what 

the bonding applied to avoid the side flash will produce, 

except for its unpredictability. 

4. FROM SIMULATIONS TO REALITY 

4.1 Setting the stage 

To illustrate the credibility gap that separates reality 

from simplified representations, the following story should 

be narrated: 

This engineer had bought a house from the previo

owner who had lived many happy years there without a

problem, so that our engineer made the (unwarrant

assumption that the house and its electrical wiring were 

good order. The house was surrounded by several t

mature trees so he thought that the cone of protection fr

the trees would benefit the house. Alas ! All-knowing Z

recognized that this engineer needed to be taught a less

on reality and thus sent a downward stepped leader tow

the general area of the engineer's house ... 

By now, dear readers, you have guessed that o

mythical engineer is none but the author of this paper 

who will now offer this true story for your edificuti

First, the "where" : Figure 4 shows a simplified (here 

go again ...) topology of the house, a two-story woodji-a

with basement and attic. Utilities (power, telephone, a

cable TV), all entered, via overhead service drops, at 

rear of the house, while water and sewer undergrou

pipes were at the front of the house. The telephone sys

was not involved in the incident and therefore is not sho

in the figure. The power installation included the us

revenue-meter (outside) and service panel (inside) w

circuit breakers controlling a multitude of branch circu

Only three are shown in the figure: lightingfixture in 

attic, TV on the secondfloor, and a counter-top recepta

(via ground-fault interrupter) for the kettle (@in Figure 

sitting on the enamelled cast-iron kitchen sink. (

significance of this detail will sulface shortly.) 

-Firm bonds - Dubious bonds 

.... Once upon a time, in a far-away land (Upstate New 

York, U.S.A.) there lived an engineer who was recording 

surges, writing papers and presenting tutorials on surge 

protection, including the need for good bonding practices. Figure 4 - Simplified configuration of anecdotal st

ICLPLOOO 4

This arrester used only a 30-mm diameter silicon 

carbide disk as varistor, which most likely would be 

destroyed by the high-energy lightning surges presently 

considered or recommended by some IEC committees. 

(Sorry, the house has changed owners and an exhaustive 

test on that particular surge arrester, desirable as it would 

be in retrospective, is not possible.) 

The attic side-flash (@ in Figure 4) clearly indicated 

that the sewer vent pipe was the point of strike (a), raising 

the question- of why the tall trees failed in their expected 

mission of establishing an effective cone of protection. 

Perhaps one explanation might be that during the initial part 

of the rainfall, the still-dry trees could not emit a successful 

competing upward streamer, compared to the wellgrounded cast-iron pipe. Comments from lightning physics 

experts on this speculation would be welcome. 

Thus, our engineer had learned his lesson, and lived 

happily without further incident for fifteen more years in 

the far-away land. However one cannot say 'lived happily 

ever after' : After moving to a new home further South, 

one night a nearby lightning flash triggered a burglar alarm 

(which had to be pried open to silence the horn turned on 

by a failed semiconductor, at 02:OO am no less) and 

damaged a remote-control garage door opener: Zeus had 

still kept track of the battle-hardened surge-protection 

engineer, but that is another story ... 

5. FROM REALITY TO SIMULATION 

Among several investigations based on rockettriggered lightning, the ongoing effort at Camp Blanding in 

Florida, U.S.A. is aimed at injecting a lightning current at 

specific points of the replica of a residential power system. 

Initial results (1997) were inconclusive because of 

instrumentation problems, but as these are progressively 

overcome, more definitive information becomes available. 

Actually, the most recent report [8] provides so many raw 

measurement results that an effort of synthesis will be 

necessary to gain a better understanding of the issues. 

The major advantage of such systematic projects over 

a random recitation of anecdotes could be the possibility of 

goirrg frum a real-world cdguratio~~ to a sufficiently 

detailed numerical representation of the circuit parameters. 

A cross-validation of the measurement results and of the 

simulation results would then significantly increase the 

credibility of both, and lead to realistic designs and ratings 

for SPDs. 

The challenge, of course, will be to represent enough 

of the many, many parameters involved in the real world 

but not so many as to make the simulation model 

unmanageable. For instance, the real-world situation of 

the anecdote already simplified in Figure 4 - with the 

ill-defined bonds and side-flashes - would be difficult to 

turn into a manageable and credible simulation. 

6. DISCUSSION 

The simplified assumptions on lightning cu

dispersion illusPated in Section 2 have met with 

skepticism among the North-American surge-prote

community and perhaps others. Part of this skeptici

also based on the relatively rare occurrence of ma

failures for secondary arresters (distribution transfo

secondary terminals and residential service entra

designed to withstand the "classical" 8/20 ps or 4/1

surges, at crest levels of a few to a few tens of kiloamp

Furthermore, the two simulations cited in Section 2 

based on the assumption that earthing electrodes ha

constant resistance during the flow of the lightning cur

an assumption that is questioned on the basis of prelim

results of measurements made in Florida in connection 

triggered lightning experiments [8]. 

In contrast with these simplified scenarios, the 

word anecdote would be a challenge for any nume

modelling but demonstrates evidence of substa

overvoltages developed in the* installation (insul

puncture at the kettle) during the flow of this undeterm

lightning current dispersion among the complex avai

paths to earth. The anecdote also offers an example 

surge arrester with modest current-handling capab

surviving the scenario of a direct strike to a building. 

A symptom of the incomplete consensus is 

noticeable lack of a discussion of risk analysis in the re

developed by the IEC Joint Working Group [2]. This 

was initially included in the document outline, ra

high expectations, but, confronted with incompa

proposals, the group gave up on that initiative. 

proposed methodologies ranged from elaborate and det

mathematical formulae - which turned out to be u

somewhat arbitrary postulates - to common-sense, al

intuitive considerations. 

7. A PROPOSAL FOR THE DISPERSION 

OF LIGHTNING INFORMATION 

In a 1963 freedom-seeking speech that still reson

today, the mantra "l'hve a dream" was coined. On a m

more modest scale, the author has a dream of unfet

information-sharing on lightning. Having cited 

preceding examples of developing, but still incomp

consensus on the dispersion of lightning current, here i

proposal (or is it a challenge ?): Hopefully helpful ti

participation, on a world-wide basis via electronic 

could supplement - not compete with - the establi

routes for information sharing, at a much accelerated p

We are still mostly in a mode of developing standard

a notoriously slow process - by volunteers or deleg

often hampered by travel budgets, or of publishing p

reviewed papers - unquestionably a wise process

entailing long delays between generation and ulti

publication of the information. 
This process of information dispersion might take one 

of the many forms by which the Internet has revolutionised 

information sharing. Should this paper be accepted for oral 

presentation at the Conference, the author would propose 

to make only a very brief summary of the paper itself - available to all in print - and make use of the scheduled 

presentation time for a cross-pollination of ideas among the 

attendees (much superior to the one-on-one poster process) 

on how to implement the proposal, bringing reality to the 

dream. Pessimists will point out hurdles such as the 

requirement of ''previously unpublished information" for 

later acceptance of an archival paper reporting research on 

the subject, or the understandable modesty of researchers 

who want to be sure that the work is complete before 

publishing even preliminary results, and so forth. Optimists 

will find ways to by-pass these hurdles and broaden an 

early consensus. 

CONCLUSIONS 

While there is no disagreement, or at least very little 

skepticism, on the specific parameters of the lightning 

discharge, consensus on the implications of lightning 

current dispersion for the rating of surge-protective 

devices has not yet been reached. 

Anecdotal information offered in many countries on 

their experience with service entrance surge-protective 

devices having moderate handling capability suggests 

that the proposed ratings for very high duty levels 

might be unnecessary and not cost-effective, unless a 

convincing risk analysis demonstrates otherwise. 

Information dispersion on these issues could be greatly 

enhanced by establishing an informal and timesensitive world-wide site (in parallel, not in conflict 

with more formal procedures), which the author is 

prepared to undertake if encouraged and supported by 

colleagues in the lightning-protection community. 

REFERENCES 

9.1 General 

Many publishing organizations, in their instructions for 

the peer-review process, raise the question "Are references 

adequate to show knowledge of work by others?" or words 

to that effect. While undoubtedly a valid question, the 

result is sometimes a lengthy recitation of up to several 

hundred citations, which seems an overkill. 

Standard-writing organizations have evolved the 

concept of differentiating between, on the one hand 

"References" - a listing of documents that are made an 

integral part of the standard by a ritual introductory 

statement, and, on the other hand, citatio~~ - in the form 

of a "Bibliography" with or without annotations. 

For this paper, "References" are limited to the 

minimum necessary to support a particular point 

made. To illustrate where extensive listing of "refere

might lead, a likrature search was conducted 

"lightning" as a leading key word, and next wit

additional word. The results are listed below, showi

number of "hits" found for the period of just 1969-19

the accessible on-line data base did not include Ben

Franklin's seminal letters to the Royal Society on lig

protection of houses and the Purfleet munitions storag

Lightning 15 791 

Lightning + surge 2348 

Lightning + current 3306 

Lightning + damage 1130 

Lightning + protection 6349 

Lightning + arrester 1816 

Lightning + earth + electrode 139 

These numbers show that it would be unrealisti

single researcher to examine in detail the contents of f

thousand papers. Injecting the concerted filterin

sharing action of today's active researchers into a r

accessible data base - the author's dream - woul

great improvement. 

9.2 Specific references 

[I] Anderson, R.B. and Eriksson, A.J.. "Lightning param

for engineering applications." ELECTRA, No.69, Ma

1980. 

[2] IEC 62066: 2000, General basic information regard

surge overvoltages and surge protection in low-volta

ax. power circuits. 

131 Mansoor, A. and Martzloff, ED., 'The Effect of Neu

Earthing Practices on Lightning Current Dispersion 

Low-Voltage Installation," IEEE Transactions PWR

July 1998. 

[4] Key, T.S., Nastasi, D., Phipps, K.O., and Martzloff, 

"Some enlightening case histories on lightning dama

Submitted for presentation at ICLP 2000. September 

[5] Key, T.S., and Martzloff. F.D., "Surging the Upside-

House: Looking into Upsetting Reference Voltages," 

Proceedings, PQA '94 Conference, Amsterdam, Octo

1995. 

[a] IEEE Std I 100-1999 ("Emerald Book") IEEE 

Recommended Practice for Powering and Groundin

Electronic Equipment. IEEE, 1999. 

[71 IEEE PC62.41.1-2000 - Drafi IEEE Guide on the S

Environment in Low-Voltage (I000 Vand less) AC P

Circuits. 

[8] EPRI "Investig?tion of lightning entry into a second

service using rocket-triggered lightning," EPRI Repo

TR-110418, April 1998. 

[9] Cavendish. Watson. Franklin and Robertson, "A rep

of the Committee appointed by the Royal Society 

to consider a Method jor securing the Powder 

Magazines at Purjleet," Philosophical Transact

of the Royal Society, London, 1772. 
ROLE AND

STRESS OF

SURGEPROTECTIVE

DEVICES

IN SHARING

LIGHTNING

CURRENT
Most simulations performed to investigate the sharing (dispersion) of lightning current for the case of a direct flash to 

a building have focused on the role and stress of surge-protective devices (SPDs) installed at the service entrance o

a building and their involvement in that part of the lightning current that exits the building via the power supply 

connection to the energy supply. 

The numerical simulations performed for this paper, based on a postulated waveform and amplitude suggested by 

current standards, include downstream SPDs, either incorporated in equipment or provided by the building occupan

The results show that a significant part of the exiting lightning current can involve those downstream SPDs with som

likelihood that their surge withstand capability might be exceeded. Such a possibility then raises questions on the 

validity of the postulated amplitude in the face of the relatively rare occurrence of reported failures. 

Filename: Role of SPDs
THE ROLE AND STRESS OF SURGE-PROTECTIVE DEVICES 

IN SHARING LIGHTNING CURRENT

François D. Martzloff 

National Institute of Standards and Technology 1

Gaithersburg MD 20877 USA 

f.martzloff@ieee.org

Arshad Mansoor 

EPRI PEAC Corp. 

942 Corridor Park Blvd 

Knoxville TN 37932 USA 

 

1 Contribution of the National Institute of Standards and Technology; not subject to copyright in the United States.

Abstract – This paper examines the sharing of lightning current associated with a direct flash to a 

building. This sharing involves not just those surgeprotective devices (SPDs) that might be installed at the 

service entrance, but also all SPDs involved in the exit 

path of the lightning current. Such sharing might 

involve built-in SPDs of some equipment located close to the service entrance, but heretofore not included in 

numerical simulations performed by many researchers. From the numerical simulations reported in this paper, 

conclusions are offered that may influence the design 

and EMC testing of equipment, as well as the risk 

analysis associated with lightning protection. 

I. BACKGROUND AND RATIONALE 

This paper offers additional information to the body of 

knowledge accumulated on how the lightning current of a direct flash, injected into the earthing system of a building, is shared among the many available paths 

towards intended or opportunistic earthing electrodes. 

Recent developments in the International Electrotechnical Commission (IEC) and the Surge-Protective 

Devices (SPD) Committee of the Institute of Electronics 

and Electrical Engineers (IEEE) have focused on the role of SPDs connected at the service entrance of a 

building in the case of a direct lightning flash to the 

building. This scenario is described in IEC 61312-3 

(2000) [9], IEEE PC62.41.1 [12] and PC62.41.2 [13]. 

Prior to this new focus, most of the considerations on 

SPD applications were based on the scenario of surges 

impinging upon the service entrance of a building as 

they come from sources external to the building. The new (additional) focus addresses the scenario of the 

earth-seeking lightning current as it is shared among the 

many possible paths to earth, including the deliberate 

and opportunistic exit paths of the building earthing 

system, services other than the power system 

connection and, mostly, the power supply connection. 

Quite independently from these lightning protection 

considerations, the IEC Subcommittee SC77B had 

developed a series of documents on the electromagnetic 

compatibility of equipment, IEC 61000-4-5, Surge 

withstand capability [8] in particular. These documents 

were primarily concerned with immunity against typical 

disturbances, the rare case of a direct lightning flash to a 

building containing electronic equipment not included. 

Increasing recognition of the need to include the 

scenario of a direct flash to a building – rare as it might 

be – has motivated the formation of an IEC Joint Task 

Force TC81/SC77B for the purpose of considerin

surge stresses on equipment higher than those current

described in the IEC document 61000-4-5 on immuni

testing [8]. 

The purpose of the paper is to examine in detail th

sharing of lightning current, not just by the SPDs at th

service entrance, but also by all SPDs that might b

involved in the exit path of the lightning current. Suc

sharing might well involve SPDs incorporated in th

equipment located close to the service entrance, but n

always included in the numerical simulations that hav

been performed by many researchers (Altmaier et a

1992) [1]; (Standler, 1992) [23]; (Rakotomalala, 1994

[20]; (Birkl et al., 1996) [3]; (Mansoor and Martzlof

1998) [15]; (Mata et al., 2002) [19]. In its rece

development of a Guide and a Recommended Practic

on surges in low-voltage ac power circuits [13] th

IEEE has refrained from identifying SPDs as bein

those that may be connected at the service entranc

Instead, it refers to "SPDs involved in the exit path

without reference to their point of installation. 

Given the tendency of equipment manufacturers 

include an SPD at the equipment power input port, th

issue of "cascade coordination" arises. Several previou

papers (Martzloff, 1980) [17]; (Goedde et al., 1990

[5]; (Lai and Martzloff, 1991) [14]; (Standler, 199

[22]; (Hostfet et al., 1992) [7]; (Hasse et al., 1994) [6

have explored the concept of cascade coordinatio

involving two or more SPDs connected on the sam

power supply but at some distance from each other. 

The legitimate wish of the energy service providers 

specify robust SPDs at the service entrance results 

SPDs having a relatively high Maximum Continuou

Operating Voltage (MCOV). On the other hand, som

equipment manufacturers tend to select SPDs with 

low MCOV under the misconception that lower is bett

(Martzloff and Leedy, 1989) [18]. This dichotomy ca

result in a situation where the low-MCOV SPD

included in equipment might well become involved 

the "exit path" and thus become overstressed in the ca

of a direct flash to the building. This situation is mad

more complicated by the fact that commercial SPD

packages are assembled from typical distributor

supplies that can have an allowable tolerance band o

±10% on the voltage-limiting rating. 

To explore the possibility and implications of 

questionable coordination, numerical simulations we

performed on a simplified model of a building featurin

SPDs installed at the service entrance and SPDs th

may be incorporated in equipment connected inside th

building near the service entrance. 
II. NUMERICAL SIMULATIONS 

II.1 Basic circuit 

Figure 1 shows a simplified building power system that 

includes the key elements of this scenario: the building 

earthing system and all earthing electrodes, with the 

corresponding exit paths via the service-entrance SPDs 

and a built-in SPD provided at the power port of a 

typical item of electronic equipment. In this example, 

these SPDs are metal-oxide varistors (MOVs) with 

typical voltage ratings (150 V at the service entrance 

and 130 V in the equipment) selected for a 120/240 V 

residential power system. (The conclusions obtained 

for this type of power system will also be applicable to 

240/400 V systems.) 

Numerical analysis of the circuit behavior by EMTP [4] 

allows inclusion of the SPD characteristics as well as 

the significant R and L elements of the wiring, with 

injection of a stroke current of 100 kA 10/350 μs at any 

selected point – the earthing system in this case. The 

selection of a 100 kA peak is consistent with the 

postulate made in many published simulations, but 

might be questioned on the basis of field experience and 

lightning detection statistics, as will be discussed later 

in this paper. 

In Figure 1, the neutral is defined as part of a "multiplegrounded neutral" system (TN-C-S), with distributed R 

and L elements between its earthing electrode 

connections. The R and L values for the cables used in 

the numerical simulation, but not shown in the figure to 

avoid clutter, were selected to emulate the typical wire 

diameters used in low-voltage power distribution 

systems and building installations. 

Previous studies (Birkl et al., 1996) [3]; (Mansoor an

Martzloff, 1998) [15] have validated the intuitiv

expectation that the tail of the 10/350 μs wavefor

often postulated for simulations will be shared amon

the available paths simply according to the relativ

values of resistance in the paths leading to the earthin

electrodes. This fact is apparent in the results of Figu

2, for example at the 350 μs time: when inductiv

effects have dwindled, the current IH in the 10-

earthing resistance of the building is ten times small

than the total current exiting the building [IN+IL1+IL

toward the power distribution system in which multip

earthing electrodes offer an effective earthing resistanc

of only 1 Ω. It is also worthy to note that this sharing 

controlled by the relative values of the resistances, s

that any earth conductivity differences associated wi

local conditions will wash out. 

The combination of the service-entrance 150-V MO

on Line 2 and the 130-V MOV incorporated at th

power port of the equipment constitutes a so-calle

"cascade". When two such cascaded SPDs are to b

coordinated, a decoupling impedance must be provide

between the two SPDs so that the voltage drop cause

by the current flowing in the decoupling impedance – 

this example the impedance of the 2,5 mm2

 diamet

wires – and added to the limiting voltage of the 130-

MOV, will cause enough of the current to flow throug

the 150-V MOV to reduce stress on the 130-V MOV. 

The simulation was performed for three values of th

impedance (length) of the connection, i.e., 0,1 m, 1 m

and 10 m to assess the effect of this impedance fo

practical situations. Figure 3 shows the results fo

these three cases and Table 1 shows the resulting energ

deposition in the respective MOVs.

Figure 1 Simplified building schematic with service-entrance SPDs, one built-in equipment SPD, and 

multiple-grounded power distribution system in case of a direct lightning flash to the earthing system
0

20

40

60

80

100

120

0 100 200 300 400 500

IN

IH IL1 & IL2

Io (100 kA 10/350)

Legend 

IO: 100 kA, 10/350 μs stroke to the building earthing system 

IN: current exiting via the neutral of the power supply 

IL1, IL2: current exiting via the two lines of the power supply 

IH: current into the building earthing electrode(s) 

Vertical scale: current in kA – Horizontal scale: time in μs

Figure 2 – Sharing of the lightning current among 

available paths to earth electrodes 

In the traces of Figure 3, the total current in Line 2 (sum 

of the two currents in the two MOVs) remains 

essentially unchanged for the three combinations, but 

the sharing of the current between the two MOVs is 

significantly affected. 

Figure 3a, with only 0,1 m of separation, is not a 

practical example of connection of equipment that close 

to the service entrance – except perhaps an electronic 

residual current device incorporated in the service 

panel. The two other figures, 3b and 3c, show how the 

130-V MOV that took the largest part of the current in 

the case of Figure 3a, now takes on less as separation 

length increases. An interesting situation develops as 

the current flowing in the 10-m line to the 130-V MOV 

stores energy that will cause a stretching of the current 

in the 130-V MOV long after the 150-V MOV current 

has decayed. This is significant because the total 

energy deposited in the MOVs is the criterion used for 

coordination, even though the current in the 130-V 

MOV could be lower than the current in the 150-V 

MOV. Table 1 shows how this energy sharing changes 

with the length of the decoupling connection, according 

to the integration of the varistor currents and voltages 

obtained from EMTP. 

Table 1 – Sharing energy between MOVs 

for three different connection lengths 

Energy deposition (joules) SPD 0,1 m 1 m 10 m 

150-V MOV 620 1090 2470 

130-V MOV 2560 2030 890 

These energy levels might be acceptable for a 150-V 

MOV sized for service entrance duty, but the 890-joule 

deposition into the 130-V MOV incorporated in the 

equipment exceeds common-wisdom ratings for such 

0

2

4

6

8

10

12

14

16

18

0 500 1000 1500 2000

IL2

Ie

IS2

a) 0,1 m connection 

0

2

4

6

8

10

12

14

16

18

0 500 1000 1500 2000

IL2

IS2

Ie

b) 1 m connection 

0

2

4

6

8

10

12

14

16

18

0 500 1000 1500 2000

IL2

IS2

Ie

c) 10 m connection

Legend 

IL2: current exiting via the power supply phase conductor 

IS2: current into the service entrance SPD 

Ie: current into the equipment SPD 

All vertical scales: current in kA 

All horizontal scales: time in μs

Figure 3 – Sharing of lightning stroke current 

devices. This finding then raises a question on th

effectiveness of a cascade for the case of direct flash 

the building. In an actual installation, there would b

more than one piece of equipment, presumably eac

with a 130-V built-in MOV at the power port. On

might expect that some sharing among these multip

SPDs would reduce the energy stress imposed on the

devices. 
To explore this situation, an additional simulation was 

performed for three branch circuits, respectively 10 m, 

20 m, and 30 m, each of them supplying equipment 

incorporating a built-in 130-V MOV. Figure 4 shows 

the sharing of current among these three MOVs and the 

150-V service entrance MOV, and Table 2 shows the 

energy deposition. 

0

2

4

6

8

10

12

14

16

0 500 1000 1500 2000

IS2

Ie(10m) Ie(20m Ie(30m)

Legend 

IS2: current into the service entrance SPD 

Ie: currents in the three SPDs at end of 10, 20, and 30 m lines 

Vertical scale: current in kA – Horizontal scale: time in μs

Figure 4 – Sharing of current among MOVs 

Table 2 – Energy sharing among MOVs 

Branch circuit length and energy 

deposition into three 130-V MOVs 

10 m 20 m 30 m 

Service entrance 

150-V MOV 

620 J 370 J 280 J 1930 J 

 

II.2 Effect of manufacturing tolerances on 

commercial-grade metal-oxide varistors 

The simulations discussed so far were performed by 

postulating that both the 150-V MOV and the 130-V 

MOV had their measured voltage limiting at the 

nominal value as specified by typical manufacturer 

specifications. Such a postulate is of course difficult to 

ensure in the reality of commercial-grade devices. For 

instance, the nominal voltage-limiting value of MOVs 

rated 130 V rms is 200 V, with lower limit of 184 V and 

upper limit of 220 V. To check that aspect of the 

problem, an arbitrary lot of 300 devices rated 130 V rms 

was purchased from a distributor and the actual 

measured voltage-limiting value at 1 mA dc was determined in accordance with IEEE Std 62.33-1994 [11]. 

For this lot, the standard deviation (sigma) was found to 

be 8 V. 

 

On the basis on these measurements and to give an 

indication of the significance of tolerance effects, the 

computations reported for Figure 3c (10 m separation) 

were repeated, still with a 150 V MOV at the service 

entrance, but with varistors at ±1 sigma of the 130 V 

rms rating, that is, 122 V and 138 V rms. The results 

are shown in Table 3. 

Table 3 Energy sharing for three values of the 

equipment built-in MOV (10 m separation) 

 Equipment Energy deposited (J) 

MOV rating 

(V rms) 

Equipment 

MOV 

150-V service 

entrance MOV 

122 915 2320 

130 890 2890 

138 750 2650 

These results illustrate the significance of tolerances 

a situation where the difference between the two SPD

of the cascade is not large, because of the de fac

situation of low values of MCOV that the industry ha

unfortunately adopted. Of course, if tolerances we

also taken into consideration for the service entranc

MOV, the extremes of distributions for both MO

would make an effective coordination between 

nominal 150-V MOV and a nominal 130-V MOV eve

more problematic. 

 

II. 3 Nonlinearity of circuit elements 

 

Most of the reported simulations, as cited above, hav

been performed with a conservative postulate of a 10

kA 10/350 lightning discharge. The median of th

current peaks compiled in the seminal Berger et a

paper [2] is only 20 kA. Occasional reservations hav

been voiced on the validity of these data collected wi

technology dating back to the 1970’s. A recent (Ju

2000) actual case history was communicated to th

authors by a colleague for two major lightning storm

recorded in the area of Tampa in Florida by means o

the Lightning Detection System [24], during which ov

30 000 flashes were detected in a period of less than 1

hours, with only one at the 150 kA level, and a media

of 20 kA, confirming the Berger at al. data. 

One could expect that the dispersion of the lightnin

current that results from the combined action of line

elements (resistance and inductance) with nonline

components (MOVs) might produce a different sharin

of the current as the decoupling element is linear but th

SPDs are nonlinear. To explore this hypothesis, th

computations for the case of Figure 4 and Table 2 we

repeated, for peak currents of 100 kA (the original valu

of the computation), 50 kA, and 25 kA (about th

median of the statistics). Table 4 shows the results o

these computations. It is interesting to note that as th

applied stroke is decreased 4 to 1 (from 100 to 25), th

total energy deposited in the varistors is decreased by 

factor of 3200/610 = 5.2. This relative greater decrea

is caused by the larger portion of the current exiting v

the linear-path neutral, further relief for all the SPD

involved in the exit path. 

 Table 4 Nonlinear effects on current sharing 

Branch circuit length and 

energy deposited into 

three 130-V MOVs 

10/350

stroke 

(kA) 10 m 20 m 30 m 

Energy 

into 

service 

entrance 

150-MOV 

Total 

energy 

in the 

MOVs 

100 620 J 370 J 280 J 1930 J 3200 J 

50 329 J 215 J 179 J 700 J 1423 J 

25 170 J 120 J 90 J 230 J 610 J 
III. DISCUSSION 

 

We have made all these computations based on 

postulating that the insulation levels are sufficient to 

prevent a flashover that would drastically affect the 

continuing energy deposition in the downstream SPDs. 

We have not included the limits of energy handling of 

the devices, which of course should be compared with 

computed deposited energy levels in a practical case. 

Another set of readings from the EMTP computations 

confirmed that the presence of SPDs at the critical 

points prevents such overvoltages from occurring (as 

long as the SPDs can carry the resulting currents) 

 

Not surprisingly, the results of the simulation confirm 

that the sharing of the lightning current occurs in 

inverse ratio of the resistances leading to the earthing 

electrodes after the initial phase of the 10/350 μs stroke. 

Likewise, one can expect that inductances will limit the 

current flow so that low-inductive paths, such as 

intended and opportunistic earth electrodes of the 

building itself, compared to the longer lines of the 

power supply, will carry a larger share of the total 

current during the initial phase of the current. This 

effect is clearly visible on the IH of Figure 2, for the 

relatively slow rise time of 10 μs of a first stroke. One 

may expect that for the subsequent strokes, or the 

flashes associated with triggered lightning experiments 

that have shorter rise times (Rakov et al., 2001) [21], 

this effect will be even more apparent. 

 

An important finding – predictable on a qualitative basis 

but heretofore not quantified for the case of a direct 

lightning flash to buildings containing electronic 

equipment – concerns the cascade coordination of builtin SPDs in the equipment. From the simple examples 

presented, it appears that a cascade of a robust serviceentrance SPD and a built-in SPD sized for limited 

energy-handling capability, according to the commonwisdom practice, might well be a delusion. 

A solution to the difficult coordination could be to 

replace the all-MOV SPD at the service entrance with a 

combined series gap-varistor device (Mansoor et al., 

1998) [16]. Such a device would also alleviate the 

concerns about the temporary overvoltage problems 

associated with MOV-only SPDs. Sparkover of the gap 

during the initial rise of the lightning current (when the 

coordination by means of the decoupling inductance 

occurs) will invite the remainder (continuing rise and 

tail) of the surge current exiting via SPDs to use the 

service entrance SPD rather than the simple and less 

robust built-in MOVs downstream. 

Last but not least, the practical question remains open 

on the need to provide surge protection against worst 

cases – the combined worst case of a direct flash to the 

building and the high-level 100 kA stroke, which is 

only at the 4% probability, according to the Berger et al. 

data [2] and even lower in the yet-anecdotal case of the 

Tampa Bay lightning storm [24]. The nonlinearity effect 

presented in II.3 adds further credibility to the overall 

need to make reasonable risk assessments of costeffectiveness before specifying high surge level 

requirements, both for the service entrance SPDs and 

for built-in SPDs in connected equipment. 

IV. CONCLUSIONS 

1. When accepting the postulate that the referenc

parameter of a direct lightning flash to a building shou

be a 10/350 μs current with a peak of 100 kA, th

numerical simulations performed for a simplifie

system with one surge- protective device installed at th

service entrance, and one or more built-in SPD 

downstream equipment indicate that the downstrea

SPD is very likely to be overstressed and fail, mo

likely catastrophically. 

2. There are several possible explanations for th

apparent contradiction between a prediction of down

stream equipment failures based on this postulate

lightning parameters, and equipment field experienc

that does not report such frequent failures, although o

course anecdotes abound. 

• The occurrence of a direct flash to a building ca

cause such extensive damage that a post-mortem fo

investigating the specifics of a prevailing ineffectiv

coordination is not performed at that time and th

issue is ignored. 

• Enough uncontrolled clearance flashovers occur 

the installation to provide significant relief for any a

risk SPDs incorporated in downstream equipment. 

• In an installation where many built-in or plug-

SPDs are present, the sharing illustrated by Figure 

combined with a low probability of a 100 kA strok

might reduce the stress on downstream devices to 

value within their capability. In particular, man

commercial plug-in SPDs advertise capabilities o

hundreds of joules, unlike the 20 joules of a sing

MOV, which might be provided at the input port o

electronic equipment. 

• Insufficient field failure data have been obtaine

compiled, shared, and published to enable realist

assessment of frequency and severity of occurrence

involving an unsuccessful cascade coordination. 

 

3. It is impractical at this point to mandate high energ

handling capability for built-in SPDs. Such a mov

might meet with strong objections from manufacture

whose products have satisfactory field experience, and 

risk analysis might show it to be not cost-effective. 

4. Economic and political realities related to the typ

and mission of the installations to be protected shou

be kept in mind. Clearly, mass-market application

such as cost-conscious consumers, in a framework o

regulated or unregulated installations, are different fro

bottom-line-conscious industrial applications, and eve

more so in the case of national assets – be they cultur

or military. 

5. Another approach for manufacturers might be 

avoid placing low MCOV varistors at the input port o

their equipment. Rather, they should select an SP

with an MCOV and resulting surge-protective level a

high as their equipment can inherently stand. This is 

“selfish” approach which is mentioned here hal

seriously, half-facetiously: there are enough low MCO

SPDs installed by users or included in other equipme

in a typical system that those unfortunate low-MCO

devices will take up the stress, leaving unscathed th

equipment wisely provided with high MCOV SPDs! 
V. REFERENCES 

Note: The citations that appear in the text are listed here 

by alphabetical order of the lead author, not by 

chronological order of appearance in the text. 

[1] Altmaier, H., Pelz, D. and Schebe, K., “Computer 

Simulation of Surge Voltage Protection in LowVoltage Systems,” Proceedings, 21st International 

Conference on Lightning Protection, Berlin 1992. 

[2] Berger, K., Anderson, R.B. and Kröniger, H., 

“Parameters of Lightning Flashes,” ELECTRA

No.41, 1975. 

[3] Birkl, J., Hasse, P., and Zahlmann, P., “Investigation 

of the Interaction of Lightning Currents with LowVoltage Installations and Their Related Threat 

Parameters,” Proceedings, 23rd International 

Conference on Lightning Protection, Firenze, 1996. 

[4] EPRI, “Electromagnetic Transients Program,” 

(EMTP), Version 2.0, Vol. 1: Main Program; Vol.2: 

Auxiliary Routines, EPRI Report EL-6421-L, 1989. 

[5] Goedde, G.L., Marz, M.B. and Henry, D.C., 

“Coordinating Lightning Stroke Protection From the 

Utility System to Load devices,” Proceedings, 

Second International Power Quality /ESD 

Conference, October 1990. 

[6] Hasse, P., Zahlmann, P., Wiesinger, J., and 

Zischank, W., “Principle for an Advanced 

Coordination of Surge-Protective Devices in LowVoltage Systems,” Proceedings, 22nd International 

Conference on Lightning Protection, Budapest, 

1994. 

[7] Hostfet, O.T., Hervland, T., Nansen, B., and Huse, 

J., “Coordination of Surge Protective Devices in 

Power Supply Systems: Needs for Secondary 

Protection,” Proceedings, 21th International 

Conference on Lightning Protection, Berlin, 

September 1992. 

[8] IEC 61000-4-5 (2001) Electromagnetic 

Compatibility - Part 4: Testing and measurement 

techniques - Section 5: Surge immunity test.

[9] IEC/TS 61312-3 (2000) Protection against LEMP - 

Part 3: Requirements of Surge Protective Devices.

[10] IEC 61643-1 (2002) Surge protective devices 

connected to low-voltage power distribution systems 

- Part 1: Performance requirements and testing 

methods

[11] IEEE Std 62.33-1994 IEEE Standard Test 

Specifications for Varistor Surge-Protective 

Devices. 

[12] IEEE C62.41.1-2002 Guide On The Surge 

Environment In Low-Voltage AC Power Circuits

[13] IEEE C62.41.2-2002 Recommended Practice on 

Characterization of Surges in Low-Voltage AC 

Power Circuits 

[14] Lai, J.S. and Martzloff, F.D., “Coordinating 

Cascaded Surge-Protection Devices: High-Low 

versus Low- High,” IEEE Transactions IA-24, 

No.4, 1993. 

[15] Mansoor, A. and Martzloff, F.D., “The Effect of 

Neutral Earthing Practices on Lightning Current 

Dispersion in a Low-Voltage Installation,” IEEE 

Transactions PWRD-13, 1998. 

[16] Mansoor, A., Martzloff, F.D., and Phipps, K., 

“Gapped Arresters Revisited: A solution to 

Cascade Coordination,” IEEE Transactions PWRD

13, 1998. 

[17] Martzloff, F.D., “Coordination of Surge Protectors 

in Low-Voltage AC Power Circuits,” IEEE 

Transactions, PAS-99, 1980. 

[18] Martzloff, F.D. and Leedy, T.F., “Selecting 

Varistor Clamping Voltage: Lower is not Better !,” 

Proceedings, 8th International Zurich Symposium 

on EMC, 1989. 

[19] Mata, C.T., Fernandez, M.T., Rakov, V.A. and 

Uman, M.A., “EMTP Modeling of a Triggered 

Lightning Strike to the Phase Conductors of an 

Overhead Distribution Line,” IEEE Transactions 

PWRD, 2002.

[20] Rakotomalala, A., Rousseau, A., and Auriol, P., 

“Lightning Distribution Through Earthing 

Systems,” Proceedings, IEEE International 

Symposium on EMC, 1994. 

[21] Rakov, V.A., Uman, M.A., Fernandez, M.I., Mata

C.T., Rambo, K.J., Stapleton, M.V. and Sutil, R.R

“Direct Lightning Strikes to the Lightning 

Protection System of a Residential Building,” IEE

Transactions PE-032PRD, 2001. 

[22] Standler, R.B., “Coordination of Surge Arresters 

and Suppressors for Use on Low-Voltage Mains,” 

Proceedings, 9th International Zürich Symposium 

on Electromagnetic Compatibility, 1991 

[23] Standler, R.B., “Calculations of Lightning Surge 

Currents Inside Buildings,” Proceedings, IEEE 

EMC Symposium, August 1992. 

[24] Waterer, S.F., Personal communication on the 

Lightning Storm in Tampa Bay Area, July 15, 

2000.