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Surge Protection Anthology 

All Papers – Part 2 

Development of Standards – Reality Checks 

Surges 

Happen! 

FOREWORD 

The papers included in this Part 2 of the Anthology report on the development of standards concerned 

with surges in low-voltage AC power, bringing “reality checks” based on field performance of surgeprotective devices, which in some cases raise questions on the validity of the requirements promulgated 

in these standards, or in other cases confirm the validity of these standards. The pre-1985 papers were 

copyrighted by the respective publishers who graciously gave permission for reprinting. The post-1985 

papers were published under the auspices of the National Institute of Standards and Technology and are 

therefore in the public domain. The citations of Annex A of Part 2 were collected for the working group 

that developed the IEEE SPD Trilogy (C62.41.1 TM –2002; C62.41.2 TM –2002; and C62.45TM-2002) – but 

do not purport to be an exhaustive listing. Although acknowledged and appreciated, for obvious copyright 

limitations, these 12 papers from other researchers cannot be reprinted here. 

CONTENTS 

Transient control levels: A proposal for insulation coordination in low-voltage systems (1976) 

Transient control level philosophy and Implementation – Part 1: Reasoning behind the philosophy (1977) 

A guideline on surge voltages in AC power circuits rated up to 600 V (1979) 

The development of a guideline on surge voltages in low-voltage AC Power (1979) 

Varistor versus environment: Winning the rematch (1986) 

Real, realistic ring waves for surge testing (1991) 

Incompatibility between the 100/1300 surge test and varistor failure rates (1991) 

Testing varistors against the VDE 0160 standard (1991) 

Standards: Transnational Aspects (1991) 

Validating surge test standards by field experience: High-energy tests and varistor performance (1992) 

Applying reality checks to standards on the surge environment (1996) 

Using incandescent lamp failure levels for assessment of the surge environment (1997) 

Driving high surge currents into long cables: More begets less (1997) 

Developing a consumer-oriented guide on surge protection (1997) 

The effect of neutral earthing practices on lightning current dispersion in a low-voltage installation (1998) 

The dilemma of surge protection vs. overvoltage scenarios: Implications for low-voltage SPDs (1998) 

The fallacy of monitoring surge voltages: SPDs and PCs Galore ! (1999) 

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

A new IEC standard on the measurement of power quality parameters (2000) 

The Trilogy update of IEEE C62.41 (2000) 

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

New IEEE standards foster next-generation system compatibility (2002) 
TRANSIENT

CONTROL

LEVELS:

A PROPOSAL

FOR INSULATION

COORDINATION

IN

LOW-VOLTAGE

SYSTEMS
Transient Control Levels:

A proposal for Insulation Coordination in Low-Voltage Systems

F.A. Fisher

General Electric Company

Pittsfield MA

François Martzloff

General Electric Company

Schenectady NY

f.martzloff@ieee.org

© 1976 IEEE

Reprinted, with permission, from

IEEE Transactions on Power Apparatus and Systems, Vol. PAS-95, No.1, Jan/Feb 1976

Significance:

Part 1 Development of standards – Reality checks

Part 6 Textbooks and tutorial reviews

One of the first papers addressing the issues of surge protection in low-voltage AC power circuits, making a

proposal for a departure from the traditional unidirectional and separate 1.2/50 and 8/20 waveforms, on the basis 

of the results of monitoring the occurrence of surges in these circuits. Nevertheless, the concept is emphasized that

surge test waveforms should not attempt to duplicate the environment, but only to apply “representative” waveforms

and levels that will demonstrate the equipment withstand capability.

The proposal also included the concept of establishing first a level of surges that will not be exceeded, thanks to 

the application of appropriate SPDs, and only then designing equipment that will withstand level higher than the

allowable level of surges. This was nothing new, having been applied successfully in the high-voltage utility

environment. However, the proposal was new for the low-voltage community. 

Unfortunately, the fait accompli of equipment being designed and placed on the market without such coordination

prevented application of that proposal. Thus, industry is left with the situation where equipment failures under

surge conditions can occur, after which remedies must be found as retrofits.

In 1976, the following statement appeared in the paper and should be kept in mind when questions arise on the

selection of “representative waveforms” in IEEE Std C62.41.2:

These BIL amplitudes, while assigned somewhat arbitrarily, were (and are) kept in touch with reality

by the fact that equipment designed in accordance with standards do not fail when exposed to surges

produced by lightning, in contrast to equipment designed prior to the development of the philosophy

of insulation coordination and the establishment of standard BILs.
ABSTRACT

Failure and circuit upset of electronic equipment

due to transients is a problem now and is one which

has promise of becoming more of a problem in the

future as trends continue toward miniaturization and

circuit complexity. Protection methods are used more

or less extensively and often haphazardly.

At present, there does not appear to be a clear approach

toward achieving compatibility between the transient withstand capability of devices and the transients to which such

devices are exposed. A more scientific approach is needed to

guide manufacturers and users of equipment.

The purpose of this paper is to promote a concept

of transient coordination for electronic and other lowvoltage equipment through the establishment of a system of Transient Control Levels, similar to the concept of Basic Insulation Levels so successfully used

for many.years in the electric power industry. Specific

suggestions for possible Transient Control Levels and

standard test wave shapes are made, in order to promote wide discussion as to whether these waveforms

and levels are the best that can be developed toward

good transient coordination for the electronic industry.

INTRODUCTION

Failure and circuit upset of electronic equipment due to

transients is a problem now and is one which has promise of

becoming more of a problem in the future as trends continue

toward miniaturization and circuit complexity. At present,

there does not appear to be a clear approach toward achieving compatibility between the transient withstand capability

of devices and the transients to which such devices are

exposed. This situation appears somewhat as illustrated

on Figure 1. A similar situation prevailed many years ago in

the electric power industry. Transients produced by lightning frequently caused failure of such vital and expensive

power equipment as transformers and generators. Those

transient problems were solved by engineering design

guided by the concept of insulation coordination and the

establishment of a series of Basic Insulation Levels (BIL’s).

Paper F 75 466-3, recommended and approved by the IEEE Surge

Protective Devices Committee of the IEEE Power Engineering Society for

presentation at IEEE PES Summer Meeting, San Francisco, Calif., July

20-25, 1975. Manuscript submitted February 3, 1975; made available for

printing April 28, 1975.

TRANSIENT CONTROL LEVELS

A Proposal for Insulation Coordination in Low-Voltage Systems

F. A. Fisher F. D. Martzloff

General Electric Company General Electric Company

Pittsfield, Mass. Schenectady, N.Y.

The purpose of this paper is to promote a conc

transient coordination for electronic and other low-v

equipment through the establishment of a system of 

sient Control Levels (TCL’s), similar to the conce

BIL’s so successfully used for many years in the e

power industry. In the following sections, specific su

tions for possible standard Transient Control Level

standard test wave shapes will be made. While the 

forms here suggested are chosen somewhat arbitrarily

are well grounded in physical reality. The purpose of 

ing such suggestions is to promote wide discussion 

whether these waveforms and levels are the best that c

developed, or if indeed the establishment of such stan

is the best way to promote good transient coordinati

the electronics industry. The ultimate purpose of any s

of transient coordination would be to achieve g

product reliability at minimum cost to the user.

AN EXAMPLE OF THE PROBLEM

TCL concepts would be of benefit to all use

electronic and other low voltage equipment, su

railroad, telephone, power, oil industry, air

and high frequency communications. The sour

transients to which equipment is exposed may be 

external (lightning and power system switching) 

ternal (switching of inductive loads, contactor res

or cross talk from adjacent circuits). While the 

cept of TCL’s is intended to apply to the full 

trum of frequencies and voltages (DC, 120 V, 

AC, 400 Hz) the problem of transient coordination 

Fig. 1. The present situation.

120
here be illustrated by discussion of 120 volt AC systems

intended for consumer and residential use. During the introduction of electronic equipment into consumer appliances

and other residential use, the importance of transient coordination was not always sufficiently recognized. In some

cases, excessive failure rates occurred as a result of transients having amplitudes greater than the withstand level of

the equipment.

In residential circuits, transients can occur from two main

sources: internally, from the switching of appliances, and

externally, most typically from the effects of lightning. One

study of internally generated transients1 has indicated that in

about three percent of U.S. households transients greater

than 1200 volts occur one or more times per week. Several

studies have been made of externally generated transients.

One such study2 indicates two percent of recorded transients

exceed 1500 volts. The data also indicate that at the location

studied, approximately two surges per year would exceed

1000 volts. Field experience1 indicated that a 100:1 drop

occurred in the failure rate of clock motors when the withstand level was increased from 2000 to 6000 volts. These

data indicate that the exposure rate to surges of 2000-volt

amplitude was sufficient to be of concern, but that surges

exceeding 6000 volts were quite rare, at least on a national

basis. Another study3 showed that during two weeks of

monitoring in a lightning-prone area, several surges exceeding 2000 volts were recorded, with the maximum recorded

being 5600 volts. Experience with field trials of Ground

Fault Circuit Interrupters sponsored by NEMA and the

Underwriters’ Laboratory4

, when correlated with the known

nuisance trip level of the devices and the observed number

of trips5

, would indicate an occurrence frequency of perhaps

one surge per 7 years above 2000 volts per household.

Most residential wiring systems are constructed in such a

manner that the various wiring boxes will flash over if they

are exposed to surges greater than 5 to 10 kV. This means

that the amplitude distribution will be chopped at 5 to 10 kV.

Based on these admittedly scattered and tentative

numbers, it appears that the typical residential circuit will be

exposed to surges of magnitude and frequency of occurrence

as illustrated in Figure 2.

The magnitude of the transients produced on 120

volt power lines, however, is not of importance except as it relates to the vulnerability level of the equipment connected to such lines. “Vulnerability” is defined

here as the level that causes an irreversible and undesirable change (usually failure) in a device. A

corollary term is susceptibility, or that level which

causes temporary malfunction of the device. The

susceptibility level cannot, by definition, be higher

than the vulnerability level. Rectifier diodes and

similar semiconductors do not have any particular

susceptibility level; they either fail or do not fail when

exposed to transients. Active semiconductor devices

or a control system operated by a mini-computer

system might be a different story. It is quite possible

Fig. 2. Exposure of residential circuits to surge (Number of su

highest surge at any one location)

that transients of a low level interfere with the o

tion of the mini-computer, causing it to give inc

results without causing permanent physical dam

The vulnerability level of such a mini-compute

be higher than the susceptibility level. Both l

must be higher than the normal operating level o

computer logic elements or input/output terminals.

The transient breakdown level or vulnerability of 

conductors is not presently a part of any industry acc

rating system. The vulnerability level is furthermo

inherently related to the normal operating voltage or 

inverse voltage (PIV) level. As examples, consider th

of Table I. During this investigation, power diodes 

subjected to unidirectional transient voltages crestin

few microseconds. The voltages at which failure occ

are seen to have little correlation to the nominal PIV r

Similar data have been accumulated for many 

conductors, particularly when semiconductor

exposed to very short transients, characterist

those produced by nuclear weapons (NEMP). Suc

formation has not been widely reported.

TABLE I

Transient Vulnerability Levels

Typical 1A Silicon Diodes

Diode PIV Failure Level Un

Number Rating Reverse Impul

Volts Volts

1 200 1100 – 1500

2 400 1400 – 1500

3 600 1400 – 160

*Breakdown observed when exposed to a unidirectional surg

at 1000 volts per microsecond.

121
Clearly, surges occur with amplitudes greater than the

vulnerability of the indicated semiconductors. The

frequency of occurrence of such damaging surges,

while small on an individual basis, may be unacceptably high on a product line. The transient amplitudes, of course, could be reduced by the use of suitable protective devices. Likewise, the vulnerability

levels of the diodes to transients could be raised. Some

questions now present themselves, all having to do with

the question of who should assume what part of the job

of providing transient coordination.

a) Should it be the responsibility of the user to control

transients to levels that do not damage equipment

supplied by vendors?

b) Should it be the responsibility of the manufacturer to

provide equipment that will not be damaged by the

naturally occurring transients?

c) If it is the responsibility of the user to control transients, to what level should he control them — the

published operating levels (in this case the published

PIV levels) or some other level higher than the

operating level but below the vulnerability level?

d) If it is the responsibility of the vendor to provide

surge-proof equipment, what level of transient

voltage and transient energy must he anticipate?

Similar questions can be asked for all product lines:

consumer, industrial, and military, and at all levels of

operating voltage.

INSULATION COORDINATION

IN THE ELECTRIC POWER INDUSIRY

Similar questions occurred many years ago during the

development of the electric power industry at a time when

the art of designing equipment to withstand the effects of

lightning was in its infancy. The nature of the transients, the

level of insulation to be used, or what should be expected of

the designers of transmission lines and lightning arresters

was not clear.

Those transient problems have largely been eliminated

today by proper engineering design on a system-wide basis.

The evolution of insulation coordination in the electric

power industry, while it can be only very briefly described

here, may be of benefit to the electronic industry.

First, the type of transients produced by lightning on

transmission lines, their magnitude and wave shape were

measured. This was not easy in the days of cold-cathode

oscilloscopes employing 50 kV accelerating voltages. Even

today with vastly improved instrumentation, such investigations are expensive and time-consuming to make.6 Yet, on

the basis of very limited testdata, a standard voltage test

wave was derived, the familiar 1.5  40 s wave. Similar

investigations in other countries led to the establishment in

Europe of the 1  50 s impulse wave. International

standardizing activities have now produced the 1.2  50 s

impulse wave, a test wave used throughout the world for

coordination of insulation protection. It was never p

tended, however, that naturally occurring surges were of 

type, only that the rise and fall times of the natural sur

were in the vicinity of the above values.

The next stage in the process of insulation coordina

was the establishment of a series of standard test and des

levels, BIL’s. For example, equipment designed for op

tion on 115-kV systems was assigned a BIL of 550 kV. 

designer of equipment to be used on 115 kV systems t

was required to provide an insulation structure that wo

withstand 550 kV. The level of 550 kV was derived on 

premise that existing lightning arresters could be used 

control the transients applied to that apparatus to less t

550 kV. The proper design of the insulation system was n

demonstrated by subjecting the apparatus in the laborat

to a surge of 1.5  40 s wave shape and a peak amplit

of 550 kV . Frequently it was part of the purchase agreem

that the equipment had to successfully pass the laborat

test. If the equipment failed, it had to be rebuilt or 

designed. Conversely, it became the responsibility of 

user to insure that no surge greater than 550 kV was e

applied to the apparatus.

As a result, power equipment achieves its resistance 

lightning-induced transients not so much by being desig

to the threat that might be posed by lightning, but by 

threat that will be posed by an acceptance test. This acc

tance test does not subject the equipment to transients h

ing the complex wave shapes produced by lightning, 

instead to transients having elementary wave shapes that 

be produced by basically simple test apparatus. Neither d

the acceptance test subject the equipment to transients of 

amplitude produced by lightning. However, it subjects 

equipment to transients of amplitude consistent with 

capabilities of existing surge-protective devices.

These amplitudes, the BIL’s while assigned somew

arbitrarily, were (and are) kept in touch with reality by 

fact that equipment designed in accordance with standa

does not fail when exposed to surges produced by lightn

in contrast to equipment designed prior to the developm

of the philosophy of insulation coordination and the es

lishment of standard BIL’s.

The test and design levels, the BIL’s, are not necessa

fixed. As better protective devices are developed, the lev

may be lowered so that reliable equipment can be buil

lower cost.

Electronic and control equipment, on the other hand, i

too often designed, built, and delivered before the existe

of a transient threat is recognized. If transients turn ou

endanger the equipment, there may be no adequate su

protective devices. There may, in fact, not be any satis

tory answer to the problem posed by transients.

THE TRANSIENT CONTROL LEVEL CONCEP

One way in which transient compatibility might 

achieved in the electronics industry is to establis

transient coordination system similar in concept to the B

122
system, but of a nature more adapted to the requirements of

electronic and control equipment.

In this paper, such a concept is called the Transient

Control Level (TCL)* concept. Specifically, it is hereby

proposed:

a) That there be defined for electronic equipment (and

other low-voltage equipment) a standard transient

voltage similar in concept to, but different in wave

shape from the 1.2  50 s wave used in coordination of insulation in high-voltage power apparatus.

b) That there be defined for electronic (and other lowvoltage) equipment a series of TCL’s similar in

concept to the BIL’s.

c) That a start be made on assigning one of these

standard levels to individual electronic components

and electronic devices.

d) That individual protective devices be rated in terms

of their ability to control transients to levels no

greater than, and preferably lower than, one of the

above levels.

e) That equipment and procedures be developed by

which equipment may be tested by vendors to

determine which TCL is appropriate to assign to

individual components and equipment.

f) That TCL’s begin to be used in purchase specifications.

g) That such equipment and procedures be used by

purchasers to evaluate vendor-supplied equipment to

determine its compliance with such purchase

specifications.

h) That such TCL’s begin to appear in regulatory

specifications for consumer apparatus in which the

consumers cannot make the appropriate tests or

prepare appropriate specifications.

Suggested TCL Voltage Wave Shape

The wave shape suggested for the TCL concept (with the

understanding that discussion and presentation of alternatives is actively encouraged) is shown on Figure 3. Shown

are both proposed open-circuit voltage and short-circuit

current waveforms, since the question of the impedance of

the source from which voltage surges derive must ultimately

be considered. These shapes are different from the longestablished 1.2  50 s wave employed in the BIL rating

system for electric power apparatus because none of the

recorded transients exhibited this type of wave shape on

120-volt AC circuits. The type of transient most frequently

recorded appeared of an oscillatory nature, very strongly

damped, and in a frequency range between 100 and

500 kHz.

Independent work on the resonant frequency of

power systems previously indicated a range of 150 to

500 kHz as being the natural frequency of typical residenial sytems. 7 Other investigations indicate that a

lower limit of 5 kHz might be more typical.8 Thus, it appe

that the observed transients are not at all typical of lightn

surges propagated directly into the system but are rather 

response of the power system to an initial excitation cau

by a nearby lightning stroke. The internally generated tr

sients due to switching operations typically are of the sa

basic type as those produced by the indirect effects of li

ning. The observed transients are in each case more ne

the result of the natural oscillatory response of the lo

wiring system, in this case the wiring system of typ

residences. Similar surge wave shapes have been enco

tered in a wide variety of other systems, ranging from 

planes to space booster rockets.9, 10 Typical examples 

recorded transient wave shapes are given in the Appen

The great bulk of the recorded transients exhibit a fa

front time and shorter decay time than do the transie

produced by lightning on high-voltage power lines, 

1.2  50 s type of wave.

Switching transients in air break contacts (intern

generated transients) can produce rise times in the orde

10 to 100 ns. Although this steepness attenuates rapidly w

distance, the typical front time is still less than 1.2 s. 

some types of devices (rectifier diodes) the wave shape i

secondary importance, with only the peak magnitude be

important. For other types of apparatus (inductive dev

such as motors), the front time, or more correctly the rat

change, is of importance equal to that of the peak magnitu

In still other types of devices (surge protective devices), 

total energy content of the surge is of most importan

Current Wave Shapes and Source Impedances

The characteristics of short-circuit current wave sha

are less well known than those of open-circuit voltage. 

short-circuit current is of importance both for evaluation 

surge protective devices and for equipment of low in

impedance such as lower voltage semiconductor devices

any discussion of test wave shapes and test levels, i

important to recognize the natural response of the devic

the test. It is inappropriate to prepare a specification 

implies that a specified voltage must be developed acro

device of low input impedance, such as a spark gap afte

has broken down, or to seemingly require that a speci

short-circuit current be produced through a high in

impedance, such as the line-to-ground insulation of a re

coil. The characteristics of short-circuit currents are po

defined because the impedance of the circuits from wh

transients are produced is poorly defined or unknown.

For purposes of discussion, it is suggested t

two different types of impedance be considered, 

independent of frequency (resistive source impeda

or classical surge impedance, Z = L/C), and one 

simple inductive source impedance. The wavefo

shown on Figure 3b assumes a source impedance 

* The TCL concept was first proposed by one of the authors (F. A. Fi

in regard to electronic equipment on the Space Shuttle.12

123
Fig. 3. Proposed TCL wave shapes.

10 H. Again, for purposes of discussion, it is proposed that

a resistive source have an impedance of 50 ohms, and an

inductive source have an impedance of 10 mH.

Voltage and Current Levels

Central to the success of the BIL system of insulation

coordination is the fact that only a limited number of BIL’s

were established, arranged in a generally geometric order of

progression. For purposes of discussion, we therefore propose that there be established a series of TCL’s progressing

in the approximate ratio of 3

10 or 3 values per decade.

Such possible TCL’s, as rounded to convenient voltages,

then appear as shown on Table II.

The subject of source impedance and short-circuit current

needs to be further discussed since the concept of constant

surge impedance, and particularly constant inductive surge

impedance, may not be valid. Transients of high voltage and

large energy content tend to be produced by physically large

systems, whose inductance tends to be larger than that of the

systems producing lower voltage or lower energy transients.

Proof Test Techniques

The generation of surge voltages in the laboratory is well

known to manufacturers and users of high power equipment.

However, producing a test wave of the shape and levels

proposed here may present some difftculty for the small

equipment manufacturer. To answer this need, a previously

developed circuit11, as shown in Figure 4, may be applicable.

Fig. 4. Test circuit for applying spikes on 120-volt. AC l

The objective of this design was to super-impo

120-volt, 60-Hz power line a transient having a rise 

first peak of 0. 5 uus, followed by a damped ri

100 kHz in which each successive peak is 60% 

preceding peak amplitude. The amplitude of the firs

adjustable f r o m 0 to 8000 volts. The source imped

the high-voltage transient is 50 ohms.

The 0.5 s rise characteristic is obtained by th

resonance of L1 and the capacitance of C1 an

series. Component values were selected to make 

approximately 50 ohms, and R1 was selected to 

heavy damping for a smooth transition to the fo

wave.

The 100 kHz damped ring results from the 

resonance of L2 with the parallel capacitance of C1 

Again, L/C is about 50 ohms. The series damping 

R2 was selected to produce the decay to 60% am

between successive peaks.

CONCLUSIONS

1. The present lack of transient coordination me

low-voltage systems does not allow the user of el

equipment to obtain the best reliability at lowest c

2. Manufacturers, vendors, and users coul

fit from a systematic approach to transient co

124
tion similar in concept to the BIL used for many years

in high-voltage systems. This is illustrated in Figure 5.

3. A concept of Transient Control Level (TCL) is

proposed by the authors. This involves discrete steps of

withstand level and proof tests based on the capability of

available s urge protective devices and reflecting the occurrence of surges in the real world.

Fig. 5. Well-coordinated low voltage system.

4. Discussion is earnestly invited on the parameters to

be considered in defining TCL’s such as:

• voltage waveform of the transients

• source impedance of the transients

• current waveform of the transients

• levels to be assigned — current and voltage

• proof-test techniques.

Successful application of the TCL concept will require

careful stud yof these factors, so as to develop a valid

consensus among all interested parties.

REFERENCES

(1) F. D. Martzloff and G. J. Hahn, “Surge Voltages in

Residential and Industrial Power Circuit, “IEEE

Trans. Power Apparatus Systems PAS-89, No. 6,

1049-1056 (July–Aug. 1970).

(2) P. M. Speranza (Bell Laboratories); Ora

munication to IEEE-SPD WG 3. 4. 4 (Oc

(3) J. E. Lenz, “Basic Impulse Insulation L

Mercury Lamp Ballast for Outdoor Appr

Illum. Eng. 133-140 (Feb. 1964).

(4) A. Martin and A. W. Smoot, “Fact Finding R

Ground Fault Circuit Interrupters,” Unde

Laboratories File E45269 (March 1972).

(5) D. W. Nestor (Rucker Electronics), private 

nication (1972).

(6) IEEE Committee Report, “Bibliography o

Voltages in AC Power Circuits Rated 600 

Less,” IEEE; Trans. Power Apparatus S

PAS-89, No. 6, 1056-1061 (July–Aug. 1970

(7) H. A. Gauper Jr., General Electric Co

communication to NEMA, Ground 

Protection Section (Aug. 1973).

(8) P. M. Speranza and L. H. Sessler (Bell Labor

oral communication to NEMA, Groun

Protection Section (Aug. 1973).

(9) J. A. Plumer, F. A. Fisher, and L. C. 

“Lightning Effects on the NASA F-8 Digital

Wire Airplane,” Contract NAS 4-2090, 

Scientific and Technical Information F

College Park, Md. 20740.

(10) J. A. Plumer, “Lightning Effects on Gener

tion Aircraft,” Report No. FAA-RD-73-99, 

Technical Information Service, Springfie

22151.

(11) E. K. Howell and F.D. Martzloff, “High 

Impulse Testers,” Report 75CRD039, 

Electric Company, Corporate Resear

Development Center, Schenectady, N. Y.

(12) “Space Shuttle Lightning Protection C

Document,” NASA, Lyndon B. Johnson 

Center, JSC-07636, September 1973.

125
TYPICAL WAVE SHAPESS

Fig. A1. Transient recorded during starting of a furnaceblower at service Fig. A3. Transient recorded during unidentified disturbance at se

box.

Fig. A2. Transient recorded during lightning storm on street pole. Fig. A4. Composite recording of furnace ignition transformer 

over 24 hours at service box.

Fig. A5. Typical transients recorded during lightning injection tests on

fighter-type aircraft (amplitudes are relative).

Fig. A6. Typical transients recorded during lightning injectio

small general aviation aircraft (amplitudes are relative).

126
Discussion

S.M. Harvey (Ontario Hydro Research Division, Toronto, Canada): This

paper provides a clear presentation of the case for a transient interference

immunity standard applicable to residential and, presumably, light commercial

electronic equipment. Designing transient or surge withstand compatibility into

low-voltage equipment is not, of course, a new concept. The telephone companies have been doing it for years. However, the authors have commendably

proposed their Transient Control Level concept in the context of a general and

down to earth philosophy of testing that should encourage informed discussion.

Following the establishment of Basic Insulation Levels, the electric power

industry has not been idle in the area of overvoltage testing of low-voltage

equipment. A number of committees, including the Power System Relaying

Committee of the IEEE Power Engineering Society and Technical Committee

No 41 of the International Electrotechnical Commission have been working for

years on the surge testing of static relays used for transmission line protection.

The Swedish Electrical Commission has prepared a draft proposal for interference withstand capability testing of apparatus used in power stations and

industrial installations. These committees have proposed a range of test waveforms including the familiar 1.2/50 impulse at peak voltages of 1, 3, and 5 kV,

a moderately damped 1 MHz oscillatory wave at peak voltages of 0.5, 1, and

2.5-3.0 kV, and a high-frequency spark test at 2 - 4 and 4 - 8 kV.

In 1974, Ontario Hydro introduced a uniform transient immunity test specification for relays and other equipment intended for substation relay or control

buildings. The test waveform is a moderately damped oscillatory transient

whose frequency ’can be specified in the range of 100 kHz to 2 MHz. One of

four test levels, specified in Table I, can be called for. The test is supervised

by our Supply Division and manufacturers are encouraged to supply their own

test equipment. However, it is still frequently necessary for Ontario Hydro to

make its own test generators available.

Table I

Transient Test Levels

Test Peak Amplitude (Volts) Source Impedance (ohms)

A 5000 100-500

B 2500 100-150

C 1000 30-50

D 500 30-50

Note that these levels when specified at I 00 kHz are very similar to tests 6 and

9 in Table II of the present paper. Level B, incidentally, when specified at I

MHz is equivalent to the IEEE Relay Test [1].

Our experience with the tests, although limited, suggests that minor circuit

deficiencies leading to operational upsets are common but that damage is

relatively rare. Probably the marginally greatest value of the tests at this time

lies in their potential for creating an awareness of the transient problem.

A number of questions being considered at this stage of our transient test

program can be rephrased to apply also to the proposals in this paper. Perhaps

the authors could comment on the following:

1. What is the advisability of introducing a new test waveform or test

procedure in addition to those already in circulation?

2. Would it be necessary to shield the test circuit of Fig. 4 or to locate it,

say, 4-6 meters from the equipment under test? In the latter case, should the

voltage and current waveforms be measured at the near end or the far end of

the connecting cable?

3. Can the test circuit of Fig. 4 correctly simulate transient disturbances

that occur when the white wire neutral and the green wire ground are connected

together a quarter wavelength from the device under test?

4. Can a reliable certification procedure, particularly in terms of energy

deliverable to a load, be established for test generators differing in design from

the one shown?

5. Finally, what is the incidence of damage or significant upset to equipment now used in resident at or light commercial environments and does it

justify the introduction of transient testing to this class o apparatus? If applied,

in view of the data contained in Fig. 2 of the paper, what criterion would be

used to select a test level of less than, say 500 volts?

REFERENCES

[1] ANSIC37.9Oa-1974(IEEE Std 472-1974)

Guide for Surge Withstand Capability (SWC) Tests.

Manuscript received August 13, 1975.

E.J. Cohen (U.S. Dept. of Agriculture, Washington, D.C.): We feel th

cept expressed in this paper is long overdue in the field of electrical pro

of electronic equipment. Experience within the telephone industry has a

demonstrated that, with present trends to ever smaller equipment, pro

problems can be severely aggravated. The over voltage and current tol

of microelectric circuits has decreased to the point where protection sho

major consideration in circuit design.

Added to this increased equipment vulnerability, we have fo

.,communications gap” between the manufacturers of electronic equip

and the producers of protection devices. When a protection defect is unco

we frequently encounter disagreements between the equipment and a

manufacturers. By establishing “Transient Control Levels,” as proposed 

paper, much of this “finger pointing” could be eliminated. As both equi

and arrester manufacturers -should know precisely what the other ad

protection should be minimized.

It is felt that while the concept expressed here is valid, further conside

should be given to the levels and waveshapes involved in the tests. As 

parameters may be critical to the workability of this proposal, every 

should be made to generate realistic values.

Manuscript received August 13, 1975.

Richard F. Hess (Sperry Flight Systems, Phoenix, Arizona): I agree tha

form of action is needed to properly assess and overcome the adverse 

of power transients on military and commercial equipment. Assuming 

sensus is reached concerning the need for transient control and the adop

Transient Control Levels (TCL), the following comments are intended to 

plement the proposal for transient control in low voltage systems.

The voltage specification is based upon measurements which are appro

to present and past equipment designs. For the most part these desig

devices which present a relatively high impedance to a source of tra

energy.

Damage occurs during a power transient when the device breaks dow

high to medium voltages are developed across the device while la

medium currents are flowing through it. Standard components are not no

tested under transient conditions, therefore it may be difficult to dete

whether they would break down or to assign a confidence level that they 

survive such a transient. When a device breaks down, either a voltag

current viewpoint could be assumed when describing the threat of the 

transient to the device.

If in order to conform to a specified TCL a device has been desig

withstand a specified voltage level, then the voltage specification is app

ate. However, a manufacturer designing equipment to meet a specifi

could adopt an approach which calls for the use of transient power suppr

devices (tranzorbs, metal oxide varistors, etc). In this case, transient 

surges are manifested as large current surges into equipment (throu

protection device) rather than a large voltage transient across the equip

Even when passing large currents, the network impedances (suppr

devices, etc.) will probably be significant enough to produce a natural 

current response within the total network. Thus, current measurement o

a network would contain a significant oscillatory component similar 

present in the voltage measurements.

Two types of TCL specifications should be provided:

1. Voltage

2. Current

Like the voltage specification, the waveform and magnitude of the c

specification at each TCL would be based upon the measurement of the c

response modes of networks containing power suppression devices and e

by a power transient.

With the two types of specifications, equipment could be designed and 

to withstand a power transient by safely withstanding specified voltage 

or by safely passing specified currents levels. The test equipment fo

voltage specification would be calibrated under open circuit conditio

would be designed to deliver current (in the event of device breakdow

level at least as large as that specified in the current specification. T

equipment for the current specification would be calibrated under 

circuit conditions and would be designed to provide voltage (in the eve

high impedance) at a level at least as large as that specified in the v

specification.

Manuscript received August 14, 1975.

127
Tests for semiconductor vulnerability (damage) levels using square pulse

waveform are common practice with the military. The damage level of many

discrete components has been determined an recorded. However, the damped

sinusoid pulse is more appropriate to susceptibility testing (transient upset).

Depending upon the type of equipment being tested and the frequency content

of expected transients, it may be desirable to test using more than one waveform. lower frequency, high amplitude sinusoid (100 KHz) would be used to

vulnerability testing and a higher frequency sinusoid (500 KHz, 1 MHz or

10 MHz depending upon the bandwidth of the equipment) would be used for

susceptibility testing. At each frequency the equipment shoul be subjected to at

least two pulses:

1. Maximum pulse is positive

2. Maximum pulse is negative

As a final observation, testing and test equipment should be kept a simple as

possible to avoid adding inordinate costs to the equipment ideally, the degree

of confidence obtained by such testing should result in a net reduction in

equipment costs (manufacture plus maintenance).

F.A. Fisher and F. D. Martzloff: We appreciate the response of the discussors

and will attempt to both respond to their questions and expan somewhat on the

protection philosophy we propose. First of all, it should be pointed out that

while this paper was written using household appliances as an example and

presented before a group largely concerned with utility relaying, the problems

of transients pervade the entire field of low voltage electrical and electronic

apparatus, including the communication (telephone) industry. One of the areas

where th authors have seen a great need for better transient compatibility is i

the Aerospace field. Much of the background upon which the TCL concept is

based comes from consideration of the transients induced in aerospace vehicles

by lightning and other energetic discharges. Designers in the Aerospace community tend not to have had the problem of transients brought as forcibly to

their attention as have the designers of relay devices intended to work in the

harsh electrical environment of a utility substation. With reference to Mr.

Harvey’s first question, we feel that it is advisable to introduce new test

procedures because th specialized test procedures adapted in the electric utility

field may no meet the needs of users in other fields.

Each of the discussors mentions the subject of levels and waveshapes. We

suggested the voltage waveshape of Figure 3 of the pape because measurements have indicated that most transients to which electronic equipment is

exposed are oscillatory in nature and generally of faster front and tail times than

the 1.2  50 microsecond test wave common in the electric power industry.

Several other factors influence our choice. One was that the proposed wave is

of long enough duratio that breakdown of semiconductor junctions would not

be greatly influenced by deviations from the specified waveshape. With much

shorter waveshapes, the resistance of semiconductor junctions to burn out

becomes strongly influenced by waveshape. Another is that transients of this

nature can be injected into wires by rather simple transformer-coupled pulse-injection generators, whereas transformer injection of higher frequency oscillatory voltages and currents is more difficult. Transformer injection of transients

has not been discussed in this paper but is sometimes an appropriate means of

evaluating the resistance of a device to circuit upset. Mr. Hess mentions the

need for two types of TCL specifications: voltage and current. We agree. We

have seen instances of groups worrying wastefully about specifications that call

for a specific voltage transient to be developed at the terminals of a device

when that device had properly been fitted with a low-pass filter, a low

impedance suppressor, or transient suppression spark gap Specifications that do

not recognize that one can neither develop a voltage across a short circuit nor

circulate a current through an open circuit are not only incomplete but mischievous and counterproductive.

With reference to more of Mr. Harveys questions, we feel that any test circuit

should be built in a sufficiently well-shielded cabinet so that there is no need

to physically separate the test circuit from any device under test. If a test circuit

must be located away from the device under test and an interconnecting cable

be used, we would think that the generator open-circuit voltage and shortcircuit current should be measured at end of the cable nearest the device under

test.

We do not really know what would be the interaction between a

white wire neutral and a green wire ground if the two were connected

together a quarter wavelength away from the generator. We take refuge

in the observation that transient coordination is more likely to be

Manuscript received October 10 1975.

achieved through the successful passing of even an imperfect test than it i

the avoidance of all but perfect tests.

We hold no special faith in the virtues of the test circuit shown on Figu

of the paper and show it only as one example of various test circuits that m

be produced. We feel that a reliable certification procedure not only can be, 

must be, based on specifications that are not unique to any one test circui

is for this reason that we propose specifications be written in terms of open-

cuit voltage and short-circuit currents; a concept that implies a fixed gener

impedance. Care must be taken that the voltage and current specifications 

be incompatible with the generator impedance. Since the writing of this pa

another paper discussing the impedance of AC wiring circuits has been p

lished [1]. Based on this paper, we would now propose that the inte

impedance of a transient generator be 50 ohms paralleled by 50 microhenr

Figure 1, reproduced from the referenced paper with the permission of 

author, shows how the impedance of the line (“the mains”) can be clo

approximated by the parallel combination of 50 ohms and 50 microhenr

Levels and waveshapes appropriate to such an impedance might then appea

shown in Figure 2 and Table 1.

As Messrs. Cohen, Harvey and Hess emphasize, the choice of appropr

levels is crucial to the successful implementation of a TCL philosophy. W

a TCL of 5000 or 6000 volts might be appropriate to high reliability ut

relays or a safety-oriented consumer product such as the Ground Fault Cir

Interrupter, it might impose an unnecessary economic hardship on a h

volume item intended for routine household use. Likewise, while a TCL of 

volts might be too low for residential purposes, it might be appropriate for 

power inputs of electronic equipment used in aircraft, and excessively high 

the signal inputs of data processing equipment intercommunicating thro

well-shielded signal wires.

Since of the major purposes of this paper is to promote discussion, 

appropriate to list some of the questions the authors have posed to themse

during the formulation of this proposal:

Fig. 1. Comparison of impedance measurements made by the Electrical

Research Association (ERA) on the impedance of power systems with a 

work of 50 ohm & 50 H in parallel

Fig. 2. Short-circuit current (ISC) resulting from a transient source with V

open-circuit voltage and 50 / /50 H source impedance.

128
TABLE 1

Proposed Transient Open-Circuit Short-Circuit

Control Level Voltage Level Current Level

Number (volts) (amperes)

1 10 0.68

2 25 1.7

3 50 3.4

4 100 6.8

5 250 17

6 500 34

7 1000 68

8 2500 170

9 5000 340

— Are there sufficient problems relating to transient coordination to warrant an

effort, likely to be major and long term, to achieve better coordination

between the transients to which equipment is exposed, and the ability of

equipment to withstand such transients?

— Would transient control level (or some other) specifications and sta

help achieve successful transient coordination between equipment ma

turers, utilities and equipment users?

— Should there be a limited number of fixed levels? The authors feel th

essential that the number of levels be limited, perhaps to 9-15 leve

tributed in a geometric progression over the range 10-5000 volt

assignment of the levels may have -to be done arbitrarily. This need 

cause for alarm. The electronic industry for years has worked succe

with resistor and capacitor values produced according to an arb

selected geometric progression.

— Should these levels reflect the system voltage, the expected reliability 

equipment function, the environment?

— What kind of source impedance is appropriate? As mentioned abo

impedance of 50 ohms paralleled by 50 microhenries may be appro

— Should open-circuit voltage and impedance be stated or, alterna

should open-circuit voltage and short-circuit current be specified?

— Is one impedance value suitable for the majority of the systems?

— What waveshape is appropriate, for voltage as well as current? For da

we are mostly concerned with energy and front-ofwave but if upset (in

ence) is to be included in TCL, then do we need to specify a freq

spectrum?

REFERENCE

[1] “Impedance of the Supply Mains at Radio Frequencies”, J. H

Proceeding of 1st Symposium on EMC, Montreux, May 1975.

129
TRANSIENT

CONTROL LEVELS

PHILOSOPHY

AND

IMPLEMENTATION

PART 1: THE

REASONING

BEHIND THE

PHILOSOPHY
Transient Control Levels Philosophy and Implementation

Part 1: The Reasoning behind the philosophy

François Martzloff

General Electric Company

Schenectady NY

f.martzloff@ieee.org

F.A. Fisher

General Electric Company

Pittsfield MA

Reprinted, with permission, from

Proceedings, 2nd International Symposium on Electromagnetic Compatibility, Montreux, 1977

Significance:

Part 2 Development of standards – Reality checks

Part 5 Monitoring instruments, laboratory measurements and test methods

Part 6 Textbooks and tutorial reviews

Presentation to the EMC community in a European forum of the Transient Control Level concept being proposed in

the US via the IEEE Power Engineering Society (See Fisher and Martzloff in IEEE Transactions PAS 95, 1976).

A companion paper on implementation is reprinted in Parts 5 and 6 (See Fisher and Martzloff in the same forum).

The proposal also included the concept of establishing first a level of surges that will not be exceeded, thanks to 

the application of appropriate SPDs, and only then designing equipment that will withstand level higher than the

allowable level of surges. This was nothing new, having been applied successfully in the high-voltage utility

environment. However, the proposal was new for the low-voltage community. 

Unfortunately, the fait accompli of equipment being designed and placed on the market without such coordination

prevented application of that proposal. Thus, industry is left with the situation where equipment failures under

surge conditions can occur, after which remedies must be found as retrofits.

In 1975, the following statement appeared in the paper and should be kept in mind when questions arise on the

selection of “representative waveforms” in IEEE Std C62.41.2:

These BIL amplitudes, while assigned somewhat arbitrarily, were (and are) kept in touch with reality

by the fact that equipment designed in accordance with standards do not fail when exposed to surges

produced by lightning, in contrast to equipment designed prior to the development of the philosophy

of insulation coordination and the establishment of standard BILs.
TRANSIENT CONTROL LEVEL PHILOSOPHY AND IMPLEMENTATION 

I. The Reasoning Behind the Philosophy 

F.D. Martzloff and F.A. Fisher 

General Electric Company, Corporate Research and Development, 

Schenectady, New York, and Pittsfield, Massachusetts 

Abstract 

This is the first of a pair of papers 

describing how better transient protection might 

be achieved through the use of a Transient Control Level (TCL) philosophy. The authors have 

developed and are proposing this TCL philosophy 

because damage to and upset of electronic and 

other low-voltage equipment by transients seems 

to be a never-ending problem, and one that is 

likely to get worse in the future as electronic 

controls permeate even more of the products which 

affect our lives. A number of proposals have 

been made - some already incorporated into standards - on various test wave shapes and specifications. The authors propose an approach integrating many of these proposals while focusing 

attention on significant parameters. 

Introduction 

An area where present standards do not seem 

to offer sufficient guidance to designers and 

manufacturers of electronic equipment is in what 

types of transients to consider and how to prove 

that equipment works in the presence of transients. This situation is perhaps under better 

control in the electric power field than it is in 

the fields of aerosapce, general indu-stry, housewares, and the military. For instance, the insulation of high-voltage apparatus is coordinated 

to the threats that nature provides to that 

insulation through the philosophy of insulation 

coordination as expressed in the Basic Insulation 

Level (BIL) system. The BIL system provides for 

a standardized series of levels being coordinated 

with the protective abilities of existing protective devices. On the other hand, electronic and 

control equipment is all too often designed, 

built, and delivered before the existence of a 

transient threat is recognized. If transients 

turn out to endanger the equipment, there may be 

no adequate surge protective devices. In fact, 

there may not be any satisfactory answer to the 

problem posed by transients. 

The authors1 TCL philosophy is aimed at 

achieving better coordination than now exists 

between the transients to which equipment is 

exposed and the abilities of equipment to withstand the transients. It is patterned after the 

BIL approach to insulation coordination so successfully used in the electric power field. 

The purposes of this first paper are t

explain the reasoning behind the different ele

ments of the BIL system of insulation coordina

tion, and to explain how similar reasoning ha

led to the formulation of the TCL philosophy

Some observations on how to perform TCL tests ar

given in a companion paper [I]. 

Proposal for TCL 

This proposal can be summarized by sayin

that we want to: 

Establish the concept that equipment shall b

rated in terms of its ability to withstand 

limited set of transient proof tests, rathe

than in terms of its ability to withstan

unknown "actual" transients. 

Establish the concept that transient specif

cations apply to power and signal lines. 

the past, only power mes have been con

s i dered . 

Establish a set of levels (limited in numbe

to which equipment is designed and tested. 

Establish a set of standard test waves (lim

ited in number) to which low-voltag

equipment will be subjected. 

Establish standardized relationships betwee

voltage and current (source impedance). 

Differentiate between the task of establish

ing the family of test levels and wa

shapes, and the task of actually selecting 

specific level. This means that: 

a We will propose to you a family 

levels and wave shapes 

0 You will select the specific level a

shape, based on your reliability goals

your costs, and your experience. 

This ~ro~osal ismade with awareness th

it may be one &re of an already confused array 

standards. However, if accepted by a large se

tion of industry and users, it could become 

unifying link and make the applications mo

successful. 
In the following paragraphs, we will attempt to present the background justifying our 

proposal, for each of the points listed above. 

1. Basis for ratinq equipment 

The concept that equipment be rated in 

terms of its ability to withstand a standard test 

rather than "actual" service conditions is not 

new. This is at the very heart of the system of 

BIL, which has been so successful in the field of 

electric utility equipment. 

Fortunately for the utilities, few parties 

were involved in making the decisions, and thus 

it was possible at an early stage to establish 

the BIL system and to enforce it because of the 

near total control of the engineering department 

of a utility over the system design. In the field 

of low-voltage systems, however, the selection 

and purchase of a multiplicity of components and 

equipment by a multiplicity of buyers from a 

multiplicity of vendors on behalf of a multiplicity of users have made it very difficult to 

maintain the organized systems approach which 

succeeded in the case of the electric utilities. 

A basic concept, which needs to be mutually 

accepted by users and manufacturers of equipment, 

is that it is impossible to simulate all possible 

transient overvoltages (and over-currents) that a 

product line might experience in service. However, by designing the equipment to a certain 

standard and controlling the level of transients 

by suitabrprotection, a much greater chance of 

successful operation in the cruel real world will 

be obtained. 

The task is then to establish a set of 

standard tests, acceptable to the vast majority 

of applications, reflecting the real world but 

not pretending to duplicate it, simple enough to 

be practical, conservative enough to ensure reliability, but realistic in terms of economics. 

Obtaining complete agreement from all is 

most unlikely an impossible goal, and thus the 

unsatisfactory situation endures. This stalemate 

can be broken by accepting a proposal which might 

not be perfect, but is better than many isolated 

standards or no standard at all. 

2. All lines subject to transient tests 

The existence of transients on power lines 

is by now a recognized and accepted fact, so that 

most applications will involve a certain amount 

of precautions in specifying transient withstand 

capability. However, in the case of signal 

lines, this recognition is less frequent, and 

there have been examples where a total lack of 

appreciation of the problem has led to the design 

and deployment of equipment that cannot be protected from transients. 

Transients can be introduced into a piece 

of equipment by the power lines from many 

sources, such as lightning, switching transients, 

fault clearing, and coupling from adjacent circuits. Signal lines, especially in the case of 

extensive systems covering a vast area, can also 

be subjected to induced transients by lightning, 

adjacent circuits, ground currents, etc. Since 

quite often the signal circuits tend to be 

lower voltage than the power circuits, the 

crepancy between the rated level in the cir

and the actual level of transients makes 

signal circuits more susceptible to trans

problems. 

A question related to which lines are 

subjected to transients is that of "comnon m

versus "transverse mode." This is not al

clear and must be addressed in a comprehen

specification. 

3. Test Levels 

An important feature of the BIL system 

that it involved a limited number of test le

graded to the operating voltage of the system 

which apparatus was being designed. A succes

TCL system should also be designed around a r

tively small number of levels. One who trie

establish levels is pulled in two directions

to avoid complexity by establishing a min

number of levels, and in another to pro

levels that accomnodate existing practices 

minimum disruption. 

One way to achieve this is through the 

of major and minor intervals in the lev

Figure 1 shows several possible level ser

The scales show the range 30 to 3000 volts di

ed into intervals based on 10lb, and 

The physical positioning of the numbers on 

figure shows how those numbers match the pro

tionate interval scales. In the past, we 

proposed that there be three levels per de

with the s acing between levels being appr

mately lo1/. The factors 1.5, 3, and 6 

appropriate, particularly since such a set c

include the voltage levels 600, 1500, and 

volts in some existing specifications. 

widely used specification MIL-704 includes 

600-volt level for transients, and it w

appear that this number, at least, should ap

in any set of TCL levels. Levels based on 

above progression appear in the left-hand col

TRANSIENT LEVELS CoNmoL SUGGESTED PROPOSED 

LEVELS 

PROPOSED IEC TC 28A 

7-n 

- 2000 

- 1200 

- BOO 

- 500 

- (3001 

- (200) 

- (1201 

- I801 

- (50) 

- (301 

- (201 

10"~ 

Fig. 1: Proposed levels for TCL voltages 

compared to existing level system
A progression proposed in IEC TC 28A, Low 

Voltage Insualtion Coordination, is shown in the 

right-hand column. The levels that have been 

proposed range from 500 to 12000 volts. On Fig. 

1, the levels in parentheses are inserted only to 

indicate the sequence. This progres ion, which 

seems to be based on the factor 10li, does not 

include the 600-volt level. 

Levels as arranged in the center column 

might appear to provide an appropriate compromise. We propose that the levels in boldface 

print be the recommended levels while those in 

lighter print be used, preferably sparingly, when 

intermediate levels are needed. Associated with 

each of these levels would be a short-circuit 

current level, the magnitude of which is related 

to the voltage levels through defined source 

impedances. Source impedance wi 11 be discussed 

further below. 

Some of the levels will seem very low, 

particularly to those accustomed to dealing with 

transients on power lines. They may not be 

unrealistic for some low-voltage signal circuits. 

A more important point, however, is that the 

establishment of a series of levels, from which a 

choice may be made, is a task separate and 

distinct from that of deciding to what level a 

piece of equipment should be designed. This 

latter point is discussed in more detail later. 

4. Wave shape 

Many test waves have been proposed in the 

past. Table I shows some that have been proposed. 

These wave shapes range from the very fast 

rise, short duration, to the slow-rise, long 

duration, with oscillatory or unidirectional 

voltages. Each of these is based on practical 

considerations for specific applications; but the 

total picture is then one of confusion and discouraging attempts at standardization. 

Observations of oscilToscope recordings 

and independent work on the resonant frequency of 

power systems [2] have shown that most transient 

voltages in low-voltage systems have an oscillatory wave shape, in contrast to the well-known 

and generally accepted unidirectional wave used 

in high-voltage insulations standards. Frequencies are typically in the range of 5 kHz to 

500 kHz, with the majority of the transients 

having frequencies above 100 kHz L31 . 

On the basis of these observations, the 

authors have proposed the voltage wave shape of 

Fig. 2, as being most representative of transients in low-voltage systems. 

This wave is a composite. One component is 

aimed at producing the effects associated with 

fast rise times. Coupled interference and the 

response of inductive devices are examples. 

Another component is aimed at producing the effects associated with the more slowly changing, 

and osci 1 latory, tai 1. Voltage summation in 

capacitive circuits coupled by rectifiers is an 

example. Energy handling capability of surge 

protective devices is another. 

Fig. 2: Proposed TCL voltage wave 

While this wave may then appear art

ficially contrived, it will subject test sample

to the two most significant effects of voltage

circuit upset, and circuit damage. Since th

wave may be produced by simple laboratory cir

cuits, comparison tests ma be easily done b

different organizations. [41 

This wave shape was first defined by 

consensus at a meeting of the Ground Fault Pro

tection Section of NEMA, in August 1973, and ha

since received increasing acceptance, notably a

the Underwriter's Laboratories. Recently

independent considerations [Slhave given furthe

support to a 0.5 ps rise time and 5 ps duratio

impulse. 

However, in all probability this on

osci 1 latory TCL wave will not meet the needs o

all users. Therefore, we propose that the wav

of Fig. 2 be supplemented by two unidirectiona

voltage waves: the classic ANSI 1.2x50p.

impulse wave and a 10 x 1000 us wave [6,7]. 

We believe that most applications can b

treated by one of these three wave shapes, onc

the concept is accepted that a perfect match o

"actual" wave shape and "test" wave shape is no

essential. The first wave, fast rise and 100 kH

ring, would be more applicable for circuits ex

posed to 'lightning remnants" (the natural osci

lation of a power system excited by a lightnin

discharge or switching transient at some remo

point) as well as control circuitry exposed t

induced transients. The second wave shape, th

familiar 1.2 x 50 unidirectional, would be appl

cable to circuits where direct exposure t

lightning. strokes is likely; while the thir

(long tail) would be applicable to situation

involving lightning current discharge on lon

cables. The second and third wave shapes ar

also representative of transients produced by th

switching of inductive circuits. 

Special applications, such as NEMP (Nuclea

Electromagnetic Pulse) hardening, or high-volt

age substation supervisory equipment, wou

rather retain their own well-documente

standards. 

5. Source impedance and enerqy 

In some types of tests, the object is t

determine what level of voltage will cause fail

ure (permanent or temporary) of insulation. Th
TABLE I - PARTIAL LISTING OF EXISTING OR PROPOSED TEST WAVES 

- ORIGIN 

ANSI, IEC 

DESCRIPTION 

Wave Shape Amp1 i tude 

Specified voltage Specified current 2.5 kV peak 

AP

Power ap

Low-volta

and co

substa

IEEE Std. 472 

Guide for Surge Withstand 

Capability (SWC) 

- 1.25 MHz repetitive 

at 60 Hz 

- 6 ps decay to 50X - 150n source impedance 

Specified 

levels 

Low-volta

and si

Fisher-Martzloff [a] - 0.25 ps rtse 

- 5 us to zero - Unspecified ring 

GE Transient Suppression Manual 141 - 500 kHz rise 

- 100 kHz ring 

- 40% decay 

Low-volt

Crouch-Fisher-Martzloff 1101 - 0.5 us rise 

- 100 kHz ring U.L. Ground Fault Interrupters - 2nd peak 2 60% first 

- 50n source impedance 

Specified 

levels Low-volta

IEEE Std. 465.1 Three requirements: 

Test Specifications for Gas Tube Surge Protective Devices : ~oxX2~o~~ ::rF:iFe"t Telephon

Linear loltage ramp 

of 100, 500, 5000, 10,000 V/gs 

unti 1 sparkover 

FCC Docket 19528 Comnunl

equipm

- 10 x 560 ps - 100 A short circutt current 800 V peak 

. Longitudinal 

- 10 x 160~s 

- 200 A short circuit current 1500 V peak 

Rural Electrlf ication 

Administration Spec. PE-60 

- 10 x IDDO ps voltage - 100 V/us rise 

3o of protector 

level 

Telephon

Rectangular Evaluati

pulse 3 ns to 10 ps NEMP Hardening 

* Damped sinewave 

10' to lo8 HZ 

NASA Space Shuttle Damped stnewave 

125 kHz 

a Unidirectional 

-2x100us 

Space sh

Mi 1 itary 

Envelope specif fed, 

max. duration 50 gs 
nature of the transient following breakdown is 

not of much concern. The typical test piece is of 

high impedance (except after breakdown), and thus 

does not load the generator. People have tended 

to overlook the source impedance of the generator, even in applications where that impedance 

is important. 

However, with the development of voltage 

suppression devices, the source impedance becomes 

an integral part of the suppression scheme. Some 

types of devices (spark gaps) function by 

switching into a low impedance state and reflecting the energy associated with the transient back 

from whence it came. Other devices (varistors, 

selenium, and Zener type diodes) clamp the 

voltage across their terminals while conducting 

the surge current and thus dissipate the surge 

energy in the protective device. The a b i l i t y o f 

the device to handle that energy becomes of importance. In either case, the test generators 

must be capable of supplying an appropriate 

amount of current, but should not supply too much 

current. 

Test specifications should reflect the fact 

that, in some cases, voltage is the appropriate 

measure of the transient, and in other cases 

current is the appropriate measure. Above all, 

they must avoid wording that leads the inexperienced to struggle valiantly, with everlarger 

surge generators, to develop a specified voltaqe 

across a correctly functioning spark gap or 

varistor. This has occurred. 

In the original formulation of the TCL 

concept, the authors proposed, and still do 

propose, that the generator impedance associated 

with the 100 kHz oscillatory test wave be an 

impedance representative of that measured on a-c 

supply mains. Such an impedance can be t-gW$~erIzy! as 50 ohms in parallel with 50 microThe ANSI specifications dealing with the 

long-established 1.2 x 50 ~s unidirectional wave 

do not treat source impedance directly, but 

recognize its existence by providing a separate 

current test wave for surge arresters or other 

surge protective devices. In the TCL concept as 

we now visualize it, this same approach would be 

followed: separate voltage and current levels. 

One of the applications where the 10 x 

1000 us unidirectional test wave might be appropriate would be those involving switching of 

inductive circuits. The impedance associated 

with such transients can vary over wide limits 

and may be quite low. We do not feel there is yet 

a sufficient engineering consensus as to what a 

suitable standard source impedance might be. Accordingly, we made no recommendations for such 

impedance, feeling that the evaluation of such 

impedance must be done on an individual basis for 

the specific application at hand. 

6. Selection of specific levels 

The task of selecting the transient control 

level appropriate for any one piece of equipment, 

or any one application, is one of engineering and 

cannot be fully dealt with in this paper. However, some discussion of the task Is necessary to 

show how that task fits into the overall TCL 

philosophy. The BIL system provides som

guidance. A fundamental tenet of the BIL system 

is that the insulation structure of apparatus i

not designed until after the required insulatio

level is agreed upon, and that this insulatio

level is not chosen until one is sure that ther

are voltage-limiting devices (surge arresters

that can control natural transients to level

lower than those to which the factory proof tes

will subject the apparatus under design. 

On the other hand, low-voltage and elec

tronic equipment is all too often designed without consideration of transients or whether pro

tective devices might even be available i

needed. One guideline is then that equipmen

should not be designed until an appropriate 

design level has been chosen. This choice shoul

be made after consideration of the distributio

of naturally occurring transients. 

The occurrence of transients is a statistical process, both in voltage levels and energ

content. LOW levels are comnon while high level

occur rarely. Figure 3 shows the relationshi

between voltage level and frequency of occurrence 

on 120-volt residential circuits, from 

observations made in the United States 191

While this type of information cannot serve t

predict the occurrences at individual locations, 

it is of interest if one is concerned with th

overall statistics of transients. For instance, 

a manufacturer can select a withstand level (o

conversely, a failure level) by trading off th

tangible and intangible cost of failures for th

cost of the added protection required to achieve 

that level. From the graph of Fig. 3, we can se

that decreasing the withstand level from, say, 4 

kV to 2 kV is l i k e l y t o increase the failure rat

of a product by a factor of 10. 

I I I I I I .I 

2 4 6 8 ID kV 

SURGE CREST 

Fig. 3: Exposure of residential circuits to 

surges (number of surges vs highest 

surge at any one location) 

Selection of the most appropriate level fo

a specific application should remain the preroga

tive of the parties directly interested. Thi

choice will be based on a number of factors suc

as the circuit rated voltage, the exposure of th

circuit to induced transients, the presence o

absence of a mandatory suppressor in the circuit, 

the risk analysis (probability of failure, conse

quence of a failure, cost-trade off), etc. 
TABLE I1 - PROPOSED IMPULSE LEVELS BY IEC-TC 28A 

I Rated Voltages 

Line-toEarth 

up to 

Volts 

1-phase 

or d.c. 

~ine-to-line 

L-M or L-L 

volts 

150- 300 

300- 600 

600-1200 

3-Phase 

L-N or L-L 

up to 

Volts 

System Voltages 

According to 

IEC-Publ. 38 

Volts 

Preferred Series of Impulse 

Withstand Voltages in Volts 

Category 

Note: The values of impulse withstand voltage given in columns I through V are a preferred series of values to be 

used by the Technical Cormittees for the purpose of insulation coordination. Products subjected in the field to 

the salne conditions of overvoltages or rated to withstand the same overvoltages are to be assigned values from 

the same column. While it might be useful to describe products and specify a preferred column for such products, 

SC 28A has refrained from doing so. 

An example of such a selection process is 

found in current proposals of IEC 28A for lowvoltage insulation coordination. This proposal 

includes a matrix of voltage levels depending on 

one hand on the system voltage and on the other 

hand on a level category, which is left to the 

users to choose but implies some recognition of 

exposure factors. This proposed table is reproduced here as Table I1 with the permission of 

the IEC TC 28A Chairman. 

Conclusion and Recomnendation 

Acceptance of the TCL concept by manufacturers and users of equipment, as well as standardizing and regulatory agencies, would be a 

great step toward simplification of specifications and toward more reliable system perf ormance. 

This paper has incorporated the feedback 

received after several proposals made at IEEE 

meetings, and at this point represents the position of the authors, supported and amended by 

the comnents received. Further feedback from the 

EMC comnunity is earnestly invited and welcome. 

To sumnarize our proposal, we recomnend 

consideration and eventual acceptance of the 

following: 

Major voltage levels of 300, 600, and 1250 

volts, with intermediate levels of 450, 850, 

and 2000 volts used if necessary; the levels 

to be scaled upwards or downwards by the 

appropriate powers of ten. 

A voltage wave shape of 0.5~s rise x 100 kHz 

ring with current related to voltage by a 

source impedance of 50n and 50 pH. This wave 

shape would be supplemented by 1.2 x 50 ys 

and 10 x 1000 t~s unidirectional waves. 

All terminals, power and signal, are to be 

subjected to TCL tests. 

For any particular piece of equipment, an 

appropriate level would be chosen from the 

above series, by mutual agreement between 

supplier and user. 

References 

Fisher, F.A. and Martzloff, F.D., "Transi

Control Level Philosophy and Implementati

Part 11. Techniques and Equipment for Mak

TCL Tests, " Proceedings, 2nd EMC Symposi

June 1977. 

Speranza, P.D. and Sessler, L.H. (Bell Te

phone Laboratories), Oral Comnunication 

NEMA, Ground Fault Protection Section, A

1973. 

Martzloff, F.D. and Hahn, G.J., 'Surge Vo

ages in Residential and Industrial Power C

cuits." 

Transient Voltage Suppression Manual, G

eral Electric Company, Electronics Pa

Syracuse, N.Y., 1976. 

Lerstrup, K., "Atmospheric Overvoltages 

Low-voltage Installations," Document E

28A/WG1 02/76. 

IEEE Standard 465.1, "Test Specifications 

Gas Tube Surge Protective Devices," 1977. 

Rural Electrification Administration (R

Specification PE-60, Washington, DC, 2025

Bull, J.H., "Impedance of the Supply Main

Radio Frequencies," Proceeding 1st EMC S

posium, 1975. 

Fisher, F.A. and Martzloff, F.D., "Trans

Control Levels, a Proposal for Insula

Coordination in Low-voltage Systems," 

Trans. Power Apparatus System PAS-95 No. 

120-129 (Jan.-Feb. 1976). 

Crouch, K.E., Fisher, F.A. and Martzlo

F.D., "Transient Control Levels, A Better 

to Voltage Ratings in Power Converter App

cations," Proceedings of 1976 Annual Meet

of the IEEE Industry Application Society
A GUIDELINE

ON SURGE

VOLTAGES

IN AC POWER

CIRCUITS

RATED UP

TO 600 V
A Guideline on Surge Voltages in AC Power Circuits Rated up to 600 V

François Martzloff

General Electric Company

Schenectady NY

f.martzloff@ieee.org

Reprinted, with permission, from

Proceedings, 3rd International Symposium on Electromagnetic Compatibility, Rotterdam, 1979 

Significance:

Part 2 Development of standards – Reality checks

Progress report to the European EMC community on the development of what became IEEE Std 587-1980.

Explains the proposition that a Ring Wave should be added to the traditional unidirectional impulses 

NOTE: A parallel paper was presented to the 1979 IEEE PES community at the Summer Power Meeting under the

title “The Development of a Guideline on Surge Voltages in Low-Voltage AC Power Circuits”.
A GUIDELINE ON SURGE VOLTAGES IN AC POWER 

CIRCUITS RATED UP TO 600 V 

F.D. ?Aa:tz!cff 

General Electric Company 

Corporate Research and Development 

Schenectady, New York 12345, U.S.A. 

Summary 

Surge voltages occurring in ac power circuits can be 

the cause of misoperation or product failure for residential as well as industrial systems. The problem has received 

increased attention in recent years because miniaturized 

solid state devices are more sensitive to voltage surges 

(spikes and transients) than were their predecessors. 

Although surge voltage amplitudes and their frequexcy =f eccgrrence on unprotected circnits =re we!! 

known, their waveshapes and energy content are less well 

known. Qn the his of measurements; statltlc~: and 

theoretical considerations, a practical guideline for outlining the environment for use in predicting extreme 

waveshapes and energy content can nevertheless be established. The Surge Protective Device Committee of the 

Institute of Electrical and Electronics Engineers has been 

developing such a guideline, the essential elements of 

which are presented in this paper. 

Surge voltages [ 11 occurring in ac power circuits 

rated up to 600 V can be represented by various waveshapes 

in an attempt to duplicate actual surge voltages. Two 

major types of surges reflecting differences in the environment are described to represent the situation realistically. 

Systems located inside a building and separated from 

the overhead lines by some line impedance experience 

surge voltages of waveshapes and energy levels that differ 

from those of the outdoor environment. Outside systems 

exposed to direct lightning strikes or lightning-induced 

surges--typically overhead lines--experience levels implied L.. ..-Pi- ---->--A- I-- ..----,..-- ",,:A*l:,.- uy lccc DLalluaI UD IVI JCLUIIUaL y a1 ~CJLSLJ. LlUJ 6YIUFllllr 

addresses particularly the hazards to these two types of 

-..- rqi J,JLCIIIJ LL J. 

Scope 

The guideline presented here primarily addresses ac 

power circuits with rated voltages up to 600 V, although some of the conclusions offered could apply to higher 

voltages and also to some dc power systems. Other 

standards have been established, such as IEEE 472, Guide 

for %Surge Withstand Capability (SWC) Tests, intended for 

the special case of high-voltage substation environments, 

and IEEE 28, Standard for Surge Arresters for ac Power 

Circuits, covering primarily the utilities environment. The 

guideline presented here intends to complement, not 

conflict with, existing standards, and to present a practical 

proposal for the selection of voltage and current tests to be 

applied in evaluating the surge withstand capability of 

equipment connected to these power circuits, primarily in 

residential and light industrial applications. 

Some guidance is also presented on how to proceed 

from the environment description to the selection of 

"standard" test waves. 

The surge voltages [I I considered in this guideline & rhc* exceeding ?we pr unit (or twice the peak 

operating voitagej and having durations ranging irom a fraction of a microsecond to a millisecond. Overvoltages 

of less than two per unit are not covered here, nor are 

transients of longer duration resulting from power equipment operation and failure modes. Because these lowamplitude and long-duration surges are generally not 

amenable to suppression by conventional surge protective 

devices, they require different protection techniques. 

1. The Origin of Surge Voltages 

Surge voltages occurring in low-voltage ac power 

circuits originate from two major sources: !oad switching 

transients and direct or indirect lightning effects on the 

power system. Load switching transients can be further 

divided into transients associated with (1) major power 

system switching disturbances, such as capacitor bank 

switching; (2) minor switching near the point of interest, 

such as an appliance turnoff in a household or the turnoff 

of other loads in an individual system; (3) resonating 

circuits associated with switching devices, such as thyristors; and (4) various system faults, such as short circuits 

and arcing faults. Measurements and calculations of 

lightning effects have been made to yield data on what 

levels can be produced, even if the exact mechanism of any 

particular surge is unknown. The major mechanisms by 

which lightning produces surge voltages are the following: 

A direct lightning strike to a primary circuit 

injects high currents into the primary circuit, 

producing voltages by either flowing through 

ground resistance or flowing through the surge 

impedance of the primary conductors. , 1:",.4,.:.." -*..:I,- +I...+ -:.--am +ha I;-- h,,+ h:+e ., rr ,,6"L""'6 JLLInr Lalo. ,111J-J ,#I= ,*I,& Y". ,I.,., a 

nearby object sets up electromagnetic fields 

*zhich can induce vo!?=pes on the conduc?ors of 

the primary circuit. 

The rapid collapse of voltage that occurs when a primary arrester operates to limit the primary 

voltage couples effectively through the capacitance of the transformer and produces surge 

voltages in addition to those coupled into the 

secondary circuit by normal transformer action. 

Lightning strikes the secondary circuits directly. 

Very high currents can be involved, exceeding 

the capability of conventional devices. 

Lightning ground current flow resulting from 

nearby direct-to-ground discharges couples onto 

the common ground impedance paths of the 

grounding network. 

Fast-acting protection devices, such as currentlimiting fuses and circuit breakers capable of clearing or 

beginning to part contacts in less than 2 ms, leave trapped 

inductive energy in the circuit upstream; upon collapse of 

the field, very high voltages are generated. 

Transient overvoltages [ I I associated with the 

:-itching of pe~~er facoy coyrffction cq=citors haye 

frequencies than the high-frequency spikes with which this 

dnmment is concerned; Their !eve!sj at !east in the case of 
restrike-free switching operations, are generally less than 

twice normal voltage and are therefore not of substantial 

concern here, but should not be overlooked. 

On the other hand, switching operations involving r,.rh .; +h,,, ,.-I ., A h.. . LsJL, LnCJ, D"C1I ,,lux y. YUUCCU YJ -,,+,,+,.. 

CYr,racrY, a 0: 

mercury switches, can produce, through escalation, surge 

voltages of complex waveshapes and of amplitudes several *: -.-. ---- *-- -L-- *L- 8 L-- TLillllca gr cairr LII~II LIE IIUIIII~I SYSLCIII VUILI~C. I IIC 

severest case is generally found on the had side of the 

switch and involves only the device that is being switched. 

While this situation should certainly not be ignored, in such 

a case the prime responsibility for protection rests with the 

local user of the device in question. However, switching 

transients can also appear on the line side across devices 

connected to the line. The presence and source of 

transients may be unknown to the users of those devices. 

This potentially harmful situation occurs often enough to 

command attention. 

2. Occurrence and Voltage Levels in IJnprotected Clrcljltr. 

limits, depending on the particular system. Prediction of 

the rate for a particular system is always difficult and 

frequently impossible. Rate is related to the level of the 

surges; low-level surges are more prevalent than high-level 

surges [ 31. Data collected from many sources (Appendix I) 

have led to the plot shown in Figure 1. This prediction 

shows with certainty only a relative frequency of occurrence, while the absolute number of occurrences can be 

described only for an "average location." The "high 

exposure" and "low exposure" limits of the band are shown 

as a guide, not as absolute limits [21, to reflect both the 

location exposure (lightning activity in the area and the 

nature nf the system! and the expos~e to stvitching sllrges 

created by other loads. Such data are useful in that they 

describe the maximum levels likely to be encountered and 

-.:.... -..-^ ..-a:--*.. ^S &La ....LA ..* ^^^..""^^^^ ^* "..^,. slvs ~u~nls saun#rars YL uns ~ars ul u~cut~st~s ut ~ucln 

surges. Of equal importance is the observation that surges 

in the range of 1 to 2 kV are fairly common in residential 

circuits. 

SURGE CREST kV 

F~gure 1. Rate of Surge Occurrence vs Voltage Level 

From the relative values of Figure 1, two typical 

levels can be cited for practical applications. First, the 

expectation of a 3 kV transient occurrence on a 120 V 

circuit ranges from 0.01 to I per year at a given location - a number sufficiently high to justify the recommendation 

of a minimum 3 kV withstand capability. Second, the 

wiring flashover limits indicate that a 6 kV withstand 

capability may be sufficient to ensure device survival 

indoors, but a 10 kV withstand capability may be required 

outdoors. 

2.2 Timing of Occurrence 

Surges - occur at random times with respect to the 

power frequency, and the iaiiure mode oi equipment may 

be affected by the power frequency follow current. 

Furthermore, the timing of the surge with respect to the 

power frequency may affect the level at which failure 

occurs 141. Consequently, surge testing must be done with 

the line voltage applied to the test piece. 

3. Waveshape of Representative Surge Voltages 

3.1 Waveshapes in Actual Occurrences 

Indoor - Measurements in the field, measurements in +h, 

rnns >auv~arvt 

1-I.---.+ --., 7, amw ..A *h 

rmlsvn ..--A sr~cal 

+:--I --,-..,-.:--- :-A: -...- calcularlulm Iltulcarc 

that most surge voltages in indoor low-voltage systems have oscillatory waveshapes, unlike the wellknown and generaliy accepted uiiidiieciionai waves 

specified in high-voltage insulation standards. A 

surge impinging on the system excites the natural 

resonant frequencies of the conductor system. As a 

result, not only are the surges typically oscillatory, 

but surges may have different amplitudes and waveshapes at different places in the system. These 

oscillatory frequencies of surges range from 5 kHz to 

more than 500 kHz. A 30 to 100 kHz frequency is a 

realistic measure of a "typical" surge for most 

residential and light industrial ac line networks. 

Outdoor - Surges encountered in outdoor locations 

have a!so been recorded, some being osci!htory !5!, 

others. being unidirectional. Because the overriding 

concern here is the energy associated with these c,,-"-e 3 ,-,.ncc.w.,3+:,,,3 h,,+ Fa31;.-+;c. .-I-C,--;-+;~- -6 +he 2". 6L2, V LV".7LL .Y.*.L UU. L LU".,L'C "Lac1 1pL.V" Y' L"G 

surges can be derived from the long-established 

specified duty of a secondary arrester, as detailed in 

Paragraph 3.2. While this specification is arbitrary, 

it has the strength of experience and successful 

usage. 

3.2 Selection of Representative Waveshapes 

The definition of a waveshape to be used as representative of the environment is important for the design of 

candidate protective devices, since unrealistic requirements, such as excessive duration of the voltage or very 

low source impedance, place a high energy requirement on 

the sq?pressor, with jl res"!ti~g most pem!ty tc the ed 

user. The two requirements defined below reflect this 

trade-off. 

Indoor - Based on measurements conducted by several 

indepei~deni uiganizaiions in iiO and 240 V systems 

(Appendix I), the waveshape shown in Figure 2 is 

reasonably representative of surge voltage in these 

power circuits. Under the proposed description of a 

"0.5 us x 100 kHz ring wave," this waveshape rises in 

0.5 p, then decays while oscillating at 100 kHz, each 

peak being about 60% of the preceding peak. 

Figure 2. The Proposed 0.5 us x 100 kHz Ring Wave 

(Open-circuit Voltage) 

The fast rise can produce the effects associated with 

nonlinear voltage distribution in windings and the 

dv/dt effects on semiconductors. Shorter rise times 

are found in many transients, but, as those transients 

propagate into the wiring or are reflected from 

discontinuities in the wiring, the rise time becomes 

longer. 

The oscillating and decaying tail produces the effects 

of voltage polarity reversals in surge suppressors or 

other devices that may be sensitive to poiarity 

changes. Some semiconductors are particularly sensitive to damage when being forced into or out of a 

conducting state, or when the transient is appiiea 

during a particular portion of the 60 Hz supply cycle 

(Appendix 11). The response of a surge suppressor can 

also be affected by reversals in the polarity, as in the 

case of RC attenuation before a rectifier circuit in a 

dc power supply. 
The pulse withstand capability of many semiconductors tends to improve if the surge duration is much 

shorter than one microsecond. For this reason, the 

first half-cycle of the test wave must have a -,.sz;-:--* ,4..-..*:-.. JUI11LASI11 UUI aLIYII. 

Outdoor - In the outdoor and service entrance -.--..- * er~virur~~r~n~r, weii as in iocations C~O% to the 

service entrance, substantial energy, or current, is 

still available. For these locations, the unidirectional 

impulses long established for secondary arresters are 

more appropriate than the oscillatory wave. 

Accordingly, the recommended waveshape is 1.2 x 

50 us for open-circuit voltage and 8 x 20 ps for 

short-circuit current or current in a low-impedance 

device. The numbers used to describe the impulse, 

1.2 x 50 and 8 x 20, are as defined in IEEE Standard 

28 - ANSI Standard C62.1; Figure 3 presents the 

waveshape and a graphic description of the numbers. 

(a) Open-circu~t 

Voltage Waveform 

I li 

10. (b) Short-circuit Current 

(or Current in LowImpedance C~rcu~t) 

I //I \ 

Figure 3. Waveshapes for Outdoor Locations 

4. Energy and Source Impedance 

4.1 General 

Thenergy involved in the interaction of a power 

system with a surge source and a surge suppressor will 

divide between the source and the suppressor in accordance 

with the characteristics of the two impedances. In a gaptype suppressor, the low Impedance of the arc after 

-.r-....-.-. cnrarlmvar fnrrnc --.--'- mnct ...- ~f ?he enerov to di~cin~t~CI I. 0, r---- 

elsewhere: for instance, in a power-follow current-limiting 

resistor that has been added in series with the gap. In an 

energy-absorber suppressor, by its very nature, a substantial share of the surge energy .is dissipated in the 

suppressor, but its clamping action does not involve the 

power-follow energy resulting from the short-circuit action 

of a gap. It is therefore essential to the effective use of 

suppression devices that a realistic assumption be made 

about the source impedance of the surge whose effects are 

to be duplicated. 

The voltage wave shown in Figure 2 is intended to 

represent the waveshape a surge source would produce 

across an open circuit. The waveshape will be different 

when the source is connected to a load having a lower :--A....r^ -.A +ha Aan..^s +,. .., h;-h ;+ ir I,...,c.r ir &,,nr+;rm L"'~Ua"CC, a1IU LllZ UF61 FL LY WII1LII LL 1- ."..La La O 1UIILLI"II 

of the impedance of the source 161. 

TL- >_____ A_ ___L:^L :-..^A^^^^ ' :----*....* 11rc UC~, cc LU WB~JC~I JUULCC IIII~~U~IICC i3 IIII~UL LantI 

depends largely on the type of surge suppressors that are 

used. The surge suppressors must be able to withstand the 

current passed through them by the surge source. A test 

generator of too high an impedance may not subject the 

device under test to sufficient stresses, while a generator 

of too low an impedance may subject protective devices to 

unrealistically severe stresses. A test voltage wave 

specified without reference to source impedance could 

imply zero source impedance - one capable of producing 

that voltage across any impedance, even a short circuit. 

That would imply an infinite surge current, clearly an I,nrsal;r+;r r;t,,~+;r.n ",,.LO..a..L a.L"-..Ymm. 

4.2 Proposed Approach ' Because oi the wide range of possibie source iiiipedances and the difficulty of selecting a specific value, 

three broad categories of building locations are proposed to 

represent the vast majority of locations [7,8 I, from those 

near the service entrance to those remote from it. The 

source impedance of the surge increases from the outside 

to locations well within the building. Open-circuit 

voltages, on the other hand, show little variation within a 

building because the wiring provides little attenuation 19 1. 

Table 1 outlines the three categories of building wiring. 

Table 2 shows open-cirtuit voltages and short-circuit 

currents for each of the three categories. The energy 

deposiied in a 500 'v' suppressor has been coiiipuied and is 

shown for each of the categories. 

TABLE l 

Locatlon Categories 

A. Outside and Service Entrance 

Service drop from pole to building ent 

Run between meter and distr~bution pa 'I 

Overhead line to detached build~ngs 

Y 

Underground lmes to well pumps 

B. Major Feeders and Short Branch Circuits 

nic+.;)...*in.. ..=.-I A-.,ir-c "La.. ."".."" y"ll.1 ".LC. 

n, 2nd feeder SYS?PmS in ind.s?ria! n~;lntr 

Heavy appliance outlets with "short" ronnertions 

to the service entrance 

Lightmg systems In commerc~al bu~ldlngs 

C. Outlets and Long Branch Circults 

All outlets at more than LO m (30 ft) from 

Category B with wires 1114-10 

All outlets at more than 20 m (60 ft) from 

Cateeorv A with wlres 1114-10 

TABLE 2 

Ranges of Voltage and Currents 

Energy DeposMax~mum lted ~n a 500 V , -,--..,.-... 

L"L-.."am mpu!s- $Unnrpssor 

A. Outdoor and 0 10 kV 1.2 x 50 Us tor 

Serv~ce h~gkmpedance c~rcults 

Entrance 10 kA 8 x 20 w for 

low-~mpedance crrcults 150 

B. Major Feeders 06 kV 1.2 x 50 Ps for 

and Short hlgh-impedance clrcults 

Branch 03 kA 8 x 20 Ps for 

Clrcu~ts ~mpedance clrcults 40 

l 6 kV 0.5 )IS x 100 kHz 

for htgh-impedance clrcults 

0500 A short clrcult for 

low-~mpedance clrcults 2 

C. Long Branch 06 kV 0.5 us x 100 kHz for 

Grcuris an; hgh-~iiipzdance CaiCuiiS 

Outlets 200 A short clrcult for 

low-impedance clrcults 0.8 

The values shown in the table represent the maximum 

range, corresponding to the "High Exprsure" situation of 

Figure 1. For less exposed systems, or when the prospect 

of a failure is not highly objectionable, one could specify 

lower values of open-circuit voltages with corresponding 

reductions in the currents. 
5. Conclusion 

The broad range of surge voltages occurring in lowvc!t=ge zc pcwer circnitr car! k simu!ated by a limited set 

of test waves, for the purpose of evaluating their effects 

on equipment. 

Field measurements, laboratory experiments, and 

calculations indlcate that two basic waves, at various opencircuit voltages and short-circuit current values, can 

represent the majority of surges occurring in residential, 

commercial, and light industrial power systems rated up to 

600 V rms. 

txceptions will be found to the simplification of a 

broad guideline; however, these should not detract from the 

benefits that can be expected from a reasonably valid 

uniformity in defining the environment. Other test waves 

of different shapes may be appropriate for other purposes, 

and the present guideline should not be imposed where it is 

not applicable. 

6. Acknowledgments 

The members of the IEEE Working Croup contributed 

the data, shown in Appendix I, that shaped this guideline. 

Heipiui comments, discussions, and reviews by ail members 

of the Surge Protective Devices Committee are also 

acknowledged. The author is particularly indebted to the 

contributions of Catharine Fisher and Peter Richman for 

the presentation and discussion of the concepts and 

recommendations made in this guideline. -- Appendix I - Data Base 

Recordings and surge counter data have been contributed from several sources, in addition to the surge 

counter data obtained by members of the working group. 

Representative oscillograms and summary statistics are ieqro&i;c& i:: this appendix, in snpport cf the ~c!?=.xa lo~ol~ a- -- ---- 

and oscillatory wave proposals. 

I. R~cordrngs by Bell Telephone Laborator~es 

(Data contributed by P. Speranza, internal report, 

unpublished to date) 

1.1 Typical Surge Counter Statistics 

120 V line at BTL facility in Chester, New Jersey, 

during 42 months of monitoring: 

I46 counts at 300 to 500 V 

14 counts at 500 to 1000 V 

3 counts at 1000 to 1500 V 

3 counts above 1500 V 

1-2 Typlca! Automatic Recordine Oscilloscopes - 

0 20 40 60 80 100 120 ps 

120 V OUTLET. LABORATORY BENCH 

277/480 V SERVICE ENTRANCE 

2. Recordings by General Electric Company 

(Data contributed by F.D. Martzloff [ 2 1 ) 

2.1 Surge Counter Statistics 

a) Three percsnt of a!! C.S. residences experience 

frequent occurrences (one per week or more) 

above 1200 V 

b) There is a 100:l reduction in the rate of device 

failure when the withstand level is raised from 

2 kV to 6 kV. 

2.2 Typical Automatic Recording Oscilloscopes 

0 5 10 20 ?O6OM75lOOUS . . FURNACE IGNITION - 24 HOUR PERIOD 

2.3 Simulated Lightning Strokes on a Residential Power 

eircuii iiaboratory ~odei oi jysiemi i 9 1 

,Recording of open-circuit voltage 

at a branch circuit outlet: 

2200 V peak 500 kHz oscillations 

By connecting a 130 R load at the same outlet (lA load) the voltage 

is reduced to 1400 V peak, with more damping. 

Coi~ciusiuns From This Test Series 

i. A curreni of i.5 kA imoderaie for a iighining 

discharge injected in the ground system) raises 

the wiring system of the house 2.2 kV above 

ground. Four kiloamperes (still a moderate 

value) will bring this voltage to 6 kV, the 

typical flashover value of the wiring. 

2. A natural frequency of 500 kHz is excited by a 

unidirectional impulse. 
3. In this example, the source of the transient 

(from the loading effect of 130 a) appears as 

3. Statistics By Land~s and Gyr Company 

Surge counter data on various locations in Swiss 220 V 

systems (Data contributed by L. Regez - unpublished to 

date) 

The General Electric counter statistics yield a 

point of 1200 V at about 1 occurrence per year (x). 

The General Electric clock data indicate a slope of 

iOO:i from i itit to 6 kv i- - - -j. 

The Regez data prov~de a band for the majority of 

- Inrstinnc - - - - (chnwn rrncc-hstrhorl\ with thn nvran- - - . . - . - . . - . . . . - . --I . . - I - . . - - , , . . . I. . ...- tion of the rural location with long overhead line, 

which has more occurrences. 

Working Group statistics (. -- . --) indicate a less 

steep slope, perhaps because of the influence of 

outdoor locations Included in the sample (similar 

to the rural data of Regez). 

The proposed curve, which is the center of the 210 

range of Figure 1, is shown In bold dashed lines (- - -). 

It has been drawn at the 100:l slope, passing near the Bell 

and General Electric points and located within the band of 

the Regez data. 

'" 102 2 5 10' 2 5 10' 2 

PEAK VOLTAGE V 

- Servrce entrance, 16-famdy house, underground system 

--- Same house, outlet third floor hvmg room 

- - . -- - - . 

Same house, ourlet fifth floor lmmg room 

krvlce entrance of bank burldrng m Barel --- Landlr and Gyr Plant. Zug, outlet m lab. 

Landlr and Cyr, Zug, outlet in furnace room 

Farmhouse rupplled by overhead lmes 

4. Working Group Surge Counter Statistics 

Surge counters with four threshold levels (350: 500: 

1000, and 1500 V) were used to record surge occurrences at 

various iocations. Members oi the Working Group instaiied 

these on 120 and 240 V systems of various types, including 

the following: outlets in urban, suburban, and rural 

residences; outlets in a hospital; secondary circuits on 

distribution system poles (recloser controls); secondary of 

padrmounted distribution transformers; lighting circuits in 

an industrial plant; life test racks at an appliance manufacturer; bench power supply in a laboratory. 

Summary Statistics of these measurements are as follows: 

Data base from 18 locations with a total recording 

time of 12 years spread over 4 calendar years, 

using 6 counters. 

Number of occurrences per year (weighted averages) at ?'average iocaiion.!' 

e 350 V: 22 

e 500 \.I: !! 

0 1ooov: 7 

0 1500 V: 3 

Significant extremes 

One home w~th large number of surges 

caused by washer operation. 

Four locations out of 18 never experienced a 

surge. 

0 One home experienced several occurrences 

above 1500 V, with none below that value. 

0 One industriai iocation iswitching oi a test 

rack! produced ?housands of surges in the 

350-500 V range, and several surges in 

excess of 1500 V. This location was left out ,-.f +hn ~.,.~=,na mrnnn~+=ti,-.n hn~t it nvnrnnli- V' L.'.. O.L.UbL ..V...,,U.Y..V.., 1". 1. --.....,.,..- 

fies a significant extreme. 

From the data base cited in the preceding pages, one 

can draw the chart below, including the following information on voltage vs frequency (rate) of occurrence: 

1. The Bell Laboratories data yield a point of 1000 V 

at about 2 occurrences per year (-). 

PEAK VOLTAGE - V 

Appendix 11 - Effect of Transient Polarity Reversals 

oii S~iiii~~iidii~t~i~ 

Breaitaown of sem~conaucrors under varlous conaltlons of load and translent overvoltage appl~rat~ons has 

been 1nvest1~ated.t Ev~dence 1s presented In the two 

mvestlgatlons c~ted that a reverse voltage apphed durlng 

the conduct~on per~od of the power frequency produces 

lower breakdown voltages than the apphcat~on of the same 

translent w~th no load or durmg blocking. Examples are 

glven below, taken from these two mvestlgatlons, showlng 

stat~stlcally slgnlf~cant d~fferences In the voltage levels. 

IN1 190 Diode* 

Average 

Breakdown (V) 

Transient at no load 1973 

Fast wave under load 830 

Slow wave under load 1097 

Transient at no load 2056 

Fast wave under load 894 

- Zlnw wave ~tnrler lnarl - - . . . - . - - . . - - . . - - - .--- I Inc. 

Translenr appiiea ar: 

- peak of reverse voltage 1766 

-25" after start of conductron 1181 

-90' after start of conduction 906 

-155' after start of conduction 1 115 

This effect is one of the reasons for selecting an 

oscillatory waveform to represent the environment: it will 

be more hkely to Induce sem~conductor fa~lures than a 

. unidirectional wave. Also, it shows the significance of the 

timing of the transient appl~cation with respect to the 

power frequency cycle. 

Appendix I11 - Notes and References 

1. Surge Voltage 

Deilnmons oi terms used in this guideiine are 

cc~jijtent with !EEE Sr=&=rd !aQ-!977,Dicti~q> cf &ctrical and Electronic Terms, 2nd ed.; however, some differences exist. For instance, IEEE Std 100-1977 defines a sijrge as a " -..--- :--* -' --*--A:-' -- ' 

LI~II~CIIL wave ur ~ur IFIIL, ~ULFIILIPI UI puwes in 

the electric circuits'--a definition broader than that used 

*Chowdhuri, P., "Transient-Voltage Characteristics of 

Silicon Power Rectifiers," IEEE IA-9, 5, September1 

October 1973, p. 582. 

t F.D. Martzloff, internal report, unpublished. 
here. Transient overvoltage is defined as "the peak voltage 

during the transient condition resulting from the operation 

of a switching devicef'--a definition more restricted than 

that of the present guideline. 

2. Amplitudes of Strikes, Worst Case 

The siiige .v.o;tages described in g.",;e;iile iilc;.u;e 

lightning effects on ~ower systems, mostly strikes in the 

vicinity of a power line, or at a remote point of the power -. system. Ine iiterature describes rhe irequency oi occurrence vs amplitude of lightning strikes, from the low levels 

of a few kiloamperes, through the median values of about 

20 kA, to the exceptional values in excess of 100 kA. 

Clearly, a secondary arrester rated for 10 kA can protect 

adequately in case of a mild direct strike, or of a more 

severe strike divided among several paths to ground. 

However, a very high and direct strike will exceed the 

capability of an ANSI-rated secondary arrester. 

References: 

Cianos, N. and E.T.Pierce, A Ground-Lightning Environment for Engineering Usage, Stanford Research Institute, 

ivienio Park, CA 94205, August i972. 

hdie, D.T., A.Z. Ghazi, Ivi. Syed, and 2.i. Toodslde, Characterization of the Electrical Environment, Toronto and 

Buffalo, N.Y.: University of Toronto Press, 1976. 

Martzloff, F.D. and G.J. Hahn, "Surge Voltage in Resldential and Industrial Power Circuits," ZEEE PAS-89, 6, July/ 

August 1970, 1049-1056. 

3. Level vs Rate of Occurrence 

The relations hi^ between the level and the rate of 

occurrence of surges is partly caused by the attenuation of 

the surges as they propagate away from the source of the 

surge and divide among paths beyond branching points. 

Equipment at a given point will be subjected to a relatively 

small number of high-level surges from nearby sources, but 

to a larger number of surges from more remote sources. 

timing with respect to the power frequency. Switching 

surges are iikeiy to occur near or after current zero, but 

variable load power factors will produce a quasi-random 

distribution. Some semiconductors, as shown in Appendix 11, exhibit failure levels that depend on the timing of 

the surge with respect to the conduction of power 

frequency current. Gaps or other devices involving a 

power-follow current may withstand this power follow with 

success, depending upon the fraction of the half-cycle 

remaining after the surge before current zero. Therefore, 

it is important to consider the timing of the surge with 1 

respect to the power frequency. In performing tests, either 1 

complete randomization of the timing or controlled timing 

should be specified, with a sufficient number of timing 

coditions to ievea! the most eriiicai timing. 

the "classical lightn s been established as 

1.2 x 50 ~s for a voltage wave and 8 x 20 ps for a current 

wave. Evidence has been coiiected, however, to show that 

oscillations can also occur. Lenz reports 50 lightning 

surges recorded in two locations, the highest at 5.6 kV, 

with frequencies ranging from 100 to 500 kHz. Martzloff 

reports oscillatory lightning surges in a house during a 

multiple-stroke flash. 

References: 

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

Lamp Ballast for Outdoor Applications," Illuminating Engrg., 

February 1964, pp. 133-140. 

Martzloff, F.D. and G.J. Hahn, "Surge Voltage in Residential a~d !nd"stria! PCI\WP~ Cir~cif ," IEEE PACg_O, 6, Ju!~! 

August 1970, 1049-1056. 

6, Surge !mpec!ance 2nd Socrce !mpedt.nce 

To Drevent misunderstandina. a distinction between 

source impedance and surge imped&ce needs to be made. 

Surge impedance, also called characteristic impedance, is a 

concept relating the parameters of a long line to the 

propagation of traveling waves. For the wiring practices of 

the ac power circuits discussed here, this characteristic 

impedance would be in the range of 150 to 300 Q, but 

because the durations of the waves being discussed (50 to 

20 ps) are much longer than the travel times in the wiring 

systems being considered, traveling wave analyses are not 

useful here. 

Source impedance, defined as "the impedance 

presented by a source of energy to the input terminals of a device, or network?! iiEEE Standard iOOi, is a more useiui 

concept here. 

7. Power System Sou~ceImpedance 

The measurements from whlch Fl~ure 1 was derived 

were of voltage only. Little was -known about the 

impedance of the circuits upon which the measurements 

were made. Since then, measurements have been reported 

on the impedance of power systems. Bull reports that the 

impedance of a power system, seen from the outlets, 

exhibits the characteristics of a 50a resistor with 50 pH in 

parallel. Attempts were made to combine the observed 

6 kV open-circuit voltage with the assumption of a 

50N50 pH impedance. This combination resulted in low 

energy deposlt~on capability, which was contradicted by 

field experience of suppressor performance. The problem 

led to the proposed defmition of oscillatory waves as well 

as high-energy unidirectional waves, in order to provide !-.*I. *k^ I*^^.- -$ ^^ ^-^:,,^*^".... --A *L- L:-L ----- -- uvut rats f rrccra ur at uacr~~a~u~ y WavC allu ill~llljll-CIICI gy 

deposition capability. 

Reference: 

Bull, J.H., "Impedance of the Supply Mains at Radio 

Frequencies," Proceedings of 1st Symposium on EMC, 

75CH1012-4 Mont., Montreux, May 1975. 

8. Installation Categories 

Subcommittee 28A of the International Electrotechnical Commission has prepared a report, referenced below, 

in which installation categories are defined. These 

installation categories divide the power systems according 

to the location in the building, in a manner similar to the 

location categories defined in this guideline. However, 

there are :=Ke significan: diffei?ncej be:w?eii the two 

concepts. First, the IEC categories are defined for a 

"Controlled Voltage Situation," a phrase that implies the 

presence of some surge suppression device or surge 

attenuation mechanism to reduce the volta~e levels from 

one category to the next. Second, the IEC report is more 

concerned with insulation coordination than with the 

application of surge protective devices; therefore it does 

not address the question of the coordination of the 

protectors, but rather the coordination of insulation levels 

- that is, voltages. Source impedances, in contrast to this 

guideline, have not been defined. Further discussion and 

work toward the application guidelines of both documents 

should eventually produce a consistent set of recommendations. 

Reference: 

ov la ti on Coordination Within Low-Voltage Systems Including Clearances and Creepage Distances for Equipment, 

internationai Eiectrotechnicai Commission, Report SCisA 

[Central Office) 5; to he puh!lshed in !979, 

9. Open-Circuit Voltages and Wiring Flashover 

Surges propagate with very little attenuation in a 

power system with no substantial connected loads. 

Measurements made in an actual residential system as well 

as in a laboratory simulation have shown that the most 

significant limitation is produced by wiring flashover, not 

be attenuation along the wires. Ironically, a carefully 

insulated installation is likely to experience higher surge 

voltages than an installation where wiring flashover occurs 

at low levels. Therefore, the open-circuit voltage specified 

at the origin of a power system must be assumed to 

propagate unattenuated far into the system, which is the 

reason for maintaining the 6 kV surge specification when 

going from the "B" location to the "C" !omtion. 

References: 

Martzloff, F.D. and K.E. Crouch "Coordination de la 

protection contre les surtensions dans les r6seaux basse : :: a, . . .. rcl~a~vtk iexucrlilrls, - rrucaearngs, is78 iEEE Canadian 

Conference on Communications and Power, 78CH1373-0, 

pp. 451-454. 

Martzloff, F.D. Surge Voltage Suppression in Residential 

Power Circuits, Report 76CRD092, Corporate Research 

and Development, General Electr~c Company, Schenectady, 

N.Y., 1976. 
DEVELOPMENT

OF A

GUIDELINE

ON SURGE

VOLTAGES IN

LOW-VOLTAGE

AC POWER

CIRCUITS
The Development of a Guideline on Surge Voltages 

in Low-Voltage AC Power Circuits

François Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

Reprinted, with permission, from IEEE Paper A 79 428-4, PES Summer Meeting, Vancouver, 1979

Significance:

Part 2 Development of standards – Reality checks

Progress report to the IEEE PES community on the development of what became IEEE Std 587-1980.

Explains the proposition that a Ring Wave should be added to the traditional unidirectional impulses

NOTE Parallel presentation of the subject made to the European EMC community under the title “A Guideline on

Surge Voltages in AC Power circuits rated up to 600 V”
py 

Schenectady, N.Y. 12345 

Abstract - Surge voltages In ac power circu~ts become more significant 

with the increased application of minia?urized electronics in consumer 

and ~ndustr~al products. A Worklng Group of IEEE is preparing a 

Guideline describing the nature of these surges in ac power circuits up 

to 600 V. 

The paper describes the data base and approach used by the 

Working Group and the recommendations proposed to represent typical 

surges, in order to obtain feedback before the final wriring of the 

Guideline. 

Two waveforms are proposed, one oscillatory, the other 

unidirectional, depending on the locatlon within the power system. 

Recommendations for source impedance or short-circuit current are 

also included. 

INTRODUCTION 

Surge voltages occurrlng in ac power circuits can be the cause 

of misoperation or product failure for residential as well as industrial 

systems. The prob!em has received increaed attention in recent yeais 

because miniaturized solid state devices are more sensitive to voltage 

surges isp~kes and transients) than were their predecessors. 

Although surge voltage amplitudes and their frequency of 

occurrence on unprotected circuits are well known, their waveshapes 

and energy content are less well known. On the basts of measurements, statistics, and theoretical considerations, a practical guideline 

for outlining the environment for use in predicting extreme waveshapes and energy content can nevertheless be established. A Working 

'Group of the Surge Protective Devices Committee is currently 

developing such a guideline; this paper reports the status of the 

Guideline, presents the considerations which led to the approach 

chosen, and provldes a possible vehlcle for discussion before the final 

writing and publication of the Gu~deline. 

SCOPE 

The C."ide!ine primarily addresses ac po~er circuits with rated 

voltages up to 600 V, although some of the conclus~ons offered could 

apply to higher voltages and also to some dc power systems. Other 

standards have been established, such as IEEE 472, Guide for Surge 

Withstand Capability (SWC) Tests, intended for the special case of 

hlgh-voltage substation environments, and IEEE 28, Standard for Surge 

Arresters for ac Power Circuits, covering primarily the utilities environment. The Guideline intends to complement, not conflict with, 

exist~ng standards. 

The surge voltages considered in the Guidehe are those 

exceeding two per unit (or twice the peak operating voltage) and 

having durations ranging from a fraction of a microsecond to a 

mill~second. Overvoltages of less than two per unit are not covered, 

nor are transients of longer duration resulting from power equipment 

operation and fa~lure modes. Because these low-amplitude and long, 

durat~on surges are generally not amenable to suppression by conventional surge protective devices, they require different protection 

techniques. 

Definitions of terms used In the Guideline are consistent with 

iEEt Standard 100-1977, Dictionary of Electrical and Electronic 

Terms, 2nd ed.; however, some differences exist. For instance, !CEE 

Std 100-1977 defines a surge as a "translent wave of current, potential 

or power in the electric circultl'-a definition broader than that used 

here. Transient overvoltage is defined as "the peak voltage during the 

transient condition resulting from the operation of a switching 

device4'-a definition more restricted than that of the Guideline. 

While the major purpose of the Gu~deline is to descr~be the environment, a secondary purpose is to lead toward standard tests, 

through an application guide that will be prepared in the fut

standard tests wi!! providc a realistic e~.aluation of ihe surg

capability of equipment connected to these power cir

necessity, the conlplex real s~tuation must be simpllfled to 

manageable set of standards. One must recognize the u

arbitrary character of any standard and be prepared to 

imperfect approach whlch can smplify matters, rather than 

perfect but unattainable match between the actual s~tua

standard. 

THE ORIGIN OF SURGE VOLTAGES 

Surge voltages occurrlng in low-voltage ac pow

originate from two major sources: load switching transients 

or indirect lightning effects on the power system. Load 

transients can be further divtded into transients assoc~a

-",-- !ann;va pwct >y>rclll -.,-LA- swiirhir~g disturbances, such as capa

switching; (2) minor switching near the polnt of Interest, 

appliance turnoff in a household or the turnoff of other 

,",i;l,;A3a=I c.,-+--. 1-2) ------- ...U...YUYI _lJJLLIII, \_IT 1~3~l1a~ijig ~it-cuiis db30ciaTed WIT

devices, such as thyristors; and (4) various system faults, su

circuits and arcing faults. Measurements and calculattons o

effects have been made to yield data on what levels can be 

even if the exact mechanism of any particular surge IS unk

major mechanisms by which lightning produces surge volta

following: 

(a) A direct lightning str~ke to a primary clrcult ~njec

currents into the primary clrcult, producing voltag

elther ilowlng through ground resistance or fl

through the surge impedance of the primary conduc

(b) A lightning strike that misses the !ine but hits a n

object sets up electromagnetic fields which can i

voirages on the conductors of the prmary circuit. 

icj The rapid coiiapse of voltage that occurs when a pr

arrester operates to hmit the prlmary voltage co

effect~vely through the capdcltdnce of the transf

and produces surge voltages In addit~on to those co

into the secondary clrcult by normal transformer ac

(d) Lightning strikes the secondary circults directly. 

high currents can be ~nvolved, exceeding the cap

ot conventional devlces. 

(e) Lightnmg ground current flow resulting from n

direct-to-ground discharges couples onto the co

ground impedance paths of the grounding network. 

Fast-acting protection dev~ces, such as current-lim

and circu~t breakers capable of clearing or beginning to pa

In less than 2 ms, leave trapped mductive energy in 

upstream; upon collapse of the field, very high voltages are 

Tra"sient overvo]tao- 0-1 assoc!at& w~fh the :-:.r!tching 

factor correction capacitors [I] have lower frequencies tha

frequency sp~kes wlth whlch this document IS concerned. T

at least In the rasp of restrikefree zw!tching operat!o.ns, ar

less than twice normal voltage and therefore are not of 

concern here, but should not be overlooked. 

On the other hand, switchmg operations involv~n

such as those produced by alr contactors or mercury sw

produce, through escalation, surge voltages of complex wav

of amplitudes several tlmes greater than the normal syste

The severest case is generally found on the load side of the 

involves only the device that is being switched. Wh~le th

should certainly not be ignored, in such a case the prlme re

for protection rests with the local user of the device i

However, switching transients can also appear on the l~n

devices connected to the line. The presence and source of 

may be unknown to the users of those devices. Thls 

harmful situation occurs often enough to command attention

A 79 428-4 A papr recatu~nrled and qoprwed by the 

IL'EE SWCJe Protective Wvices Committee of the IEEE 

PGG~~ Cngiiieerir~g Society for presentation at tile IRE 

PES Smwr FIeetFrq Vancocver, ~ritish Co?;-his, 

While the data have been recorded pr~marily on 120, 

--- .. -- ~///ria~ V systems, the general conclus~ons should be valid 

systems. To the extent that %large voltages are produced 5; 

amount of energy being dumped Into a power system, low-i

heavy ~r~dusrr~ai sysrems can be expected to experlence lo

Canada, July 15-20, 1979. ~~~~~~~i~

submitted ~~h~~~~~ 6, from surge voltages than 120 V resldentlal systems, but com

1979: made available for printing April 5, 1979. greater, amounts of energy potent~ally available for depo

surge suppressor. 

CH 1458-9/79/0000-4284$00.75 0 1979 l EEE 
The rate of occurrence of surges varies over wide limits, 

depending on the particular system. Prediction of the rate for a 

particular system is always difficult and frequently impossible. Rate 

is related to the level of the ,surges; low-level surges are more prevalent than high-level surges. The relationship between the level 

and the rate of occurrence of surges is partly caused by the 

attenuation of the surges as they propagate away from the source of 

the surge and divide among paths beyond branching points. Equipment 

at a given point will be subjected to a relatively small number of highlevel surges from nearby sources, but to a larger number of surges 

from more remote sources. 

Data collected from many sources have led to the plot shown in 

Fig. 1. This prediction shows with certainty only a relative frequency 

of occurrence, while the absolute number of occurrences can be 

described only for an "average location." The "high exposure" and "low 

exposure" limits of the band are shown as a guide, not as absolute 

limits, to reflect both the location exposure (lightning activity in the 

area and the nature of the system) and the exposure to switching 

surges created by other loads. 

The literature describes the frequency of occurrence vs 

amplitude of !ightning strikes, from the low levels of a few 

kiloamperes, through the median values of about 20 kA, to the 

exceptional values in excess of iOO kA [23. Clearly, a secondary 

arrester rated for 10 kA can protect adequately in case of a mild 

direct strike, or of a more severe strike divided among several paths to 

ground. However, a very high and direct strike will exceed the 

capability of an ANSI-rated secondary arrester [3 I. 

The voltage and current amplitudes presented in the Guideline 

attempt to provide for the vast majority of lightning strikes but should 

not be considered as "worst case," since this concept cannot be 

determined realistically. One should think in terms of the statistical 

distribution of strikes, accepting a reasonable upper limit for most 

cases. Where the consequences of a failure are not catastrophic but 

merely represent an annoying economic loss, it is appropriate to make 

a trade-off of the cost of protection against the likelihood of a failure 

caused by a high but rare surge. For instance, a manufacturer may be 

concerned with nation-wide failure rates, those at the upper limits of 

the distribution curve, while the user of a specific system may be 

concerned with a single failure occurring at a specific location under 

"worst-case conditions." Rates can be estimated for average systems, 

however, and even if imprecise, they provide manufacturers and users 

with guidance. Of equal importance is the observation that surges in 

the range of 1 to 2 kV are fairly common in residential circuits. 

From the relative values of Fig. I, two typical levels can be 

cited for practical applications. Fiisi, the expectation of a ?!:'I 

transient orrurrence on a 120 V circuit ranges from 0.01 to 1 per year 

at a given location -a number sufficiently high to justify the 

recommendation of a minimum 3 kV wiiirsiand capability. Second, the 

wiring flashover limits indicate that a 6 kV withstand capability may 

be sufficient to ensure device survlval indoors, but a 10 kV withstand 

capability may be required outdoors. 

TYPICAL 

OUTDOOR 

FLASHOVER 

Ftg. 1. Rate of surge occurrence vs voltage level. 

Timing of Occurrence 

Surges occur at random times with respect to the power 

frequency, and the failure mode of equipment may be affected by the 

power frequency follow current or by the timing. Consequently, surge 

testing must be done with the ac voltage apphed to the test piece. 

Lightning surges are completely random in their timing with 

respect to the power frequency. Switching surges are likely to occur 

in Appendix !!, exhibit failure levels that depend on the timing of 

surge with respect to the conduction of power frequency curre

Gaps or other devices involving a power-follow current may withsr

this power follow w~th success, depending upon the fraction of 

half-cycle remaining after the surge before current zero. Therefo

it is important to consider the timing of the surge with respect to 

power frequency. In performing tests, either complete randomizat

of the timing or controlled timing should be specified, with a suffici

number of timing conditions to reveal the most critical timing. 

WAVESHAPE OF REPRESENTATIVE SURGE VOLTAGES 

Waveshapes in Actual Occurrences 

Indoor - Measurements in the field, measurements in the laborato

and theoretical calculations indicate that most surge voltages in ind

low-voltage systems have oscillatory waveshapes, unlike the w

known and generally accepted unidirectional waves specified in hi

voltage insulation standards. A surge impinging on the system exci

the natural resonant frequencies of the conductor system. As a res

not only are the surges typically oscillatory, but surges may h

different amplitudes and waveshapes at different p!ace; In the syste

These osrillatory frequencies of surges range from 5 kHz to more th

500 kHz. A 30 to 100 kHz frequency is a realistic measure o

"typical" surge for most resideniiai and light industrial ac l

networksOutdoor - Surges encountered in outdoor locations have also b

recorded, some being oscillatory, others belng unidirectional. T

"classical lightning surge" has been established as 1.2 x 50 us fo

voltage wave and 8 x 20 ps for a current wave. Evldence has b

collected, however, to show that oscillations can also occur. Lenz 

reports 50 lightning surges recorded In two locations, the highest 

5.6 kV, with frequencies ranging from 100 to 500 kHz. Martzloff 

reports oscillatory lightning surges in a house during a multiple-str

flash. 

Because the overriding concern here IS the energy associa

with these surges, a conservative but realistic description of the sur

can be derived from the long-established specified duty of a second

arrester, as detailed below. While this specification is arbitrary, ~t 

the strength of experience and successful usage. 

Selection of Representative Waveshapes 

The definition of a waveshape to be used as representative 

the environment is important for the design of candidate protect

devices, since unrealistic requirements, such as excess~ve duration 

*c.. ~1,~ vvlragL or very lew source impedancej place a high energy requi

ment on the suppressor, with a resulting cost penalty to the end u

The two requirements defined below reflect this trade-off. 

Indoor - Based on measurements conducted by several inaepend

organizations in 120 and 240 V systems !.Appendix !!; the wavesh

shown in FI~. 2 is reasonably representative of surge voltages in th

power circuits. Under the proposed descrlptlon of a "0.5 US - 100 k

ring wave," this waveshape rises in 0.5vs, then decays w

oscillating at 100 kHz, each peak being about 60% of the preced

peak. 

FI~. 2. The proposed 0;5 gs - 100 Hz ring wave 

(open-circuit voltage). 

The fast iise can produce the effects associated with nonline

voltage distribution in windings and the dv/dt effects on semicondu

tors. Shorter rise times are found in many transients, but, as tho

transients propagate into the wiring or are reflected fr

discontinuities in the wiring, the rise time becomes longer. 

The oscillating and decaying tail produces the effects of volta

polarity reversals in surge suppressors or other devices that may 

sensitive to polarity changes. Some semiconductors are particular

sensitive to damage when being forced into or out of a conducti

state, or when the transient is applied during a particular portion 

the 60 Hz supply cycle (Appendix 11). The response of a su
suppressor can a!m be affected by reversals in the po!arity, as in the 

case of RC attenuation before a rectifier clrcuit in a dc power supply. -. Ine pulse withstand capabiiity 01 many sem~conductors tends to 

improve if the surge durati~n is m~h sherter than one microsecond. 

For this reason, the first half-cycle of the test wave must have a 

sufflc~ent duration. 

Outdoor - In the outdoor and service entrance environment, as well as 

in locations close to the service entrance, substantial energy, or 

current, is still available. For these locations, the unidirectional 

impulses long established for secondary arresters are more appropriate 

than the oscillatory wave. 

Accordingly, the recommended waveshape is 1.2 x 50 us for 

open-circuit voltages and 8 x 20 us for short-circu~t current (impulse 

discharge current) or current In a low-impedance device. The numbers 

used to descrlbe the impulse, 1.2 x 50 and 8 x 20, are as defined in 

IEEE Standard 28 - ANSI Standard C62.1; Fig. 3 presents the waveshape and a graphlc description of the numbers. 

Z2 x 1.25=8pr 

Flg. 3. Waveshapes for outdoor locations. 

ENERGY AND SOURCE IMPEDANCE 

General 

The energy involved in the interaction of a power system with a 

surge source and a surge suppressor will divide between the source and 

the suppressor in accordance with the characteristics of the two 

impedances. In a gap-type suppressor, the low impedance of the arc 

after sparkover forces most of the energy to be dissipated elsewhere: 

for Instance, in a reslstor added in series with the gap for limiting the 

power-follow current. In an energy-absorber suppressor, by its very 

nature, a substantial share of the surge energy is dissipated in the 

suppressor, but its clamping action does not involve the power-follow 

energy resulting from the short-circuit action of a gap. It is therefore 

essential to the effective use of suppression devices that a realistic 

assumption be made about the source impedance of the surge whose 

effects are to be duplicated. 

The voltage wave shown in Fig. 2 is intended to represent the 

waveshape a surge source would produce across an open circuit. The 

waveshape will be different when the source is connected to a load 

having a lower impedance, and the degree to which it is !ewer is a 

function of the impedance of the source. 

To prevent misundersranding, a distinct~on between source imwedance and surge impednnce needs to be made. krge impedance, 

also called characteristic impedance, is a concept relating the paramerers of a long iine to the propagation of traveling waves. For the 

wiring practices of the ac power circuits discussed here, this 

characteristic impedance would be in the range of 150 to 300Q, but 

because the durations of the waves being discussed (50 to 20 ps) are 

much longer than the travel times in the wiring systems being 

considered, traveling wave analyses are not useful here. 

Source impedance, defined as "the impedance presented by a 

source of energy to the input terminals of a device, or network" (IEEE 

Standard loo), is a more useful concept here. In the conventional 

Thevenins description, the opencircuit voltage (at the terminal

the network or test generator) and the source impedance (of the s

source or test generator) are sufficient to calculate the short-cir

current, as we!! a; any current for a specified ;uppiessoi impedanc

The measurements from which Fig. I was derived were 

voltage only. Little was known about the impedance of the circ

upon which the measurements were made. Since then, measurem

have been reported on the impedance of power systems. Bull 

reports that the impedance of a power system, seen from the out

exhibits the characteristics of a 50Q resistor with 50 pH in para

Attempts were made to combine the observed 6 kV open-cir

voltage with the assumption of a 508/50 uH impedance [ 71. 

combination resulted in low energy deposition capability, which 

contradicted by field experience of suppressor performance. 

problem led to the proposed definition of oscillatory waves as we

high-energy unidirectional waves, in order to provide both the eff

of an oscillatory wave and the high-energy deposition capability. 

The degree to which source impedance is important dep

largely on the type of surge suppressors that are used. The s

suppressors must be able to withstand the current passed through t

by the surge source. A test generator of too high an impedance 

not subject the devlce under test to sufficient stresses, whi

generator of too low an impedance may sub!ect protective device

unrealistically severe stresses. A test voltage wave speclfled wit

refeience to source impedance could imply zero source impedan

one capable of producing that voltage across any impedancei eve

short circuit. That would imply an infinite surge current, clearly 

.,--..-, :-*:- -:A..-L:.- 

UIII CaLIJLIC SILVdLIUtI. 

Because of the wide range of possible source impedances 

the difficulty of selecting a speciflc value, three broad categorie

building locations are proposed to represent the vast majority 

locations, from those near the service entrance to those remote f

it. The source impedance of the surge increases from the outs~

locations well within the building. Open-circuit voltages, on the o

hand, show little variation wlthin a building because the wi

provides little attenuation. Figure 4 illustrates the application of 

three categories to the wiring of a budding. 

I I 

Fig. 4. Locatlon categories. 

n n.,*_._I.. __A C^_..._^ C -A_-- -- , ,. VYIaLUc O1lY .,=I "*LC LIIIIaII~~ 8. ividjor Feeders and inort Brmrh Cc

Servlce drop from pole to bulldmg Dtstribmon pane! rlevres entrance Bus and feeder systems m jndurrrial iiun berween merer and atstrtbu- plants tlon panel Heavy appllan~e outletr with "rhort

Overhead lme to detached buildmgs connections to the service eniran

Underground lines to well pumps Llghting systems In commerctai bul

C. Outlets and Long Brarch Circurts 

All outlets at more than 10 m (30 it) front 

Category B with wires 114-10 
rrtission has prepared a Repoii i8 j, in which instailation categories are 

defined. These installation categories divide the power systems 

according to the location in the building, in a manner similar 'to the 

location caregories deiined in rhe Guideline. However, rhere are some 

significant differences between the two concepts. First, the IEC 

categories are defined for a "controlled voltage situation," a phrase 

that implies the presence of some surge suppression device or surge 

attenuation mechanism to reduce the voltage levels from one category 

to the next. Second, the IEC report is more concerned with insulation 

coordination than with the application of surge protective devices; 

therefore it does not address the question of the coordination of the 

protectors but, rather, the coordination of insulation levels - that is, 

voltages. 

Surges propagate with very little attenuation in a power system 

with no substantial connected loads. Measurements made in an actual 

residential system as well as in a laboratory simulation have shown 

that the most significant limitation is produced by wiring flashover, 

not by attenuation along the wires. Ironically, a carefully insulated 

installation is likely to experience higher surge voltages than an installation where wiring flashover occurs at low levels. Therefore, 

the open-circuit voltage specified at the origin of a power system must 

be assumed to propagate unattenuated far into the system, which is 

the reason fcr maintai~ing the 6 k?' surge specificatim when going 

from one category to an adjacent category farther into the building. 

Furthermore, source impedances are not defined in the IEC 

report. The Guide!:ne attempts to fill this need by specifying severs: 

levels of source impedance, or of short-circuit current, for the various 

categories. 

PROPOSED REPRESENTATION OF THE ENVIRONMENT 

On the basis of the preceding discussions, the Guideline 

proposes to reduce the infinite variety of actual conditions to three 

categories, from the outside service drops to the long branch circuits 

and outlets. 

For each category the most appropriate waveshape is indicated, 

an open-circuit voltage for high-impedance loads, or a short-circuit 

current for low-impedance loads. The tabulation that follows shows 

open-circuit voltages and short-circuit currents for each of the three 

categories. The energy deposited in a 500 V suppressor has been 

computed and is shown for each of the categories. 

The values shown in the table represent the maximum range, 

corresponding to the "High Exposure" situation of Fig. 1. For less 

exposed systems, or when the prospect of a failure is not highly 

objectionable, one could specify lower values of open-circuit voltages 

with corresponding reductions in the currents. IEC Category I, not 

represented in the Guideline, would correspond to line cord-connected 

devices in this context; 

CONCLUSIONS 

The broad range of surge voltages occurring in low-voltage ac 

power circuits can be simulated by a lim~ted set of test waves, for the 

purpose of evaluating their effects on equipment. 

Field measurements, laboratory experiments, and calculations 

indicate that two basic waves, at various open-circuit voltages and 

short-circuit current values, can represent the majority of surges 

occurring in residential, commercial, and light industrial power 

systems rated up to 600 V rms. 

he; however, these should nor aetracr from the benefits that can b

expected from a reasonably valid uniformity in defining the environ

ment. Other test waves of different shapes may be appropriate fo

orher purposes, and the present guideiine should not be imposed wher

it is not applicable. 

The Working Group is approaching the final phases of prepara

tion of the Guideline document; comments are solicited from th

engineering and user communities. However, readers must recogniz

the unavoidably arbitrary character of any standard and be prepared t

accept an imperfect approach, which can simpllfy matters and clarif

the issues as well as provide uniform evaluations of performance

rather than demand a perfect but unattainable match between th

actual situation and the standard. 

ACKNOWLEDGMENTS 

The concepts presented in this paper have greatly benefite

from the informed questions and discussions by members of th

Working Group on Surge Voltage in AC Power Circuits Rated 600 V o

Less, and from interested reviewers; particular appreciation fo

effective critiques from Catharine Fisher and Peter Richman 

acknowledged. The data base presented in Appendix I has bee

broadened by the contributions of the Bell Telephone Laboratories an

Landis h Gyr, Inc. 

REFERENCES 

"Switching Surges Due to De-Energization of Capacitiv

Circuits," AlEE Committee Report, AIEE Transactions, Augus

1957, pp. 562-564. 

Cianos, N. and E.T. Pierce, A Ground-Lightning Environmen

for Engineering Usage, Stanford Research Institute, Menlo Park

CA 94205, August 1972. 

Surge Arresters for Alternating-Current Power Circuits, IEEE 

Standard 28-1 974; ANSI Standard C62.1-1975; IEC Stan

dard 99-2. 

Lenz, 3.E., "Basic Impulse Insulation Levels of Mercury Lam

Ballast for Outdoor Applications," Illuminating Engrg., Februar

1964, pp. 133-1 40. 

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

and Industrial Power Circuits," lEEE PAS-89, 6, July/August 197

! 0+?-! 056. 

Bull, J.H., "Impedance on the Supply Mains at Radio Frequen

cies," Proceedings of 1st Symposium on EMC, 75CH1012-4 Mont

Msntreux, May 1975. 

Martzloff, F.D. and F.A. Fisher, "Transient Control Leve

Philosophy and Implementation: The Reasoning Behind the 

Philosophy," Proceedings 2nd Symposium on EMC, 77CH1224- 

5EMC, Montreux, June 1977. 

Insulation Coordination Within Low-Voltage Systems Includin

Clearances and Creepage Distances for Equipment, Internation

al Electrotechnical Commission Report SC28A (Centra

Office) 5, to be published in 1979. 

Recommended Surge Voltages and Currents Deemed to Represent the Environment 

TY pe Impulse of Specimen Energy 

Comparable to Waveform Maximum* or Load Deposited in 

Location IEC SC28A Amplitude Circuit a 500 V Suppressor 

Cateaorv Cateeorv (ioules) 

A. Outdoor and IV 1.2 x 50 ps 10 kV O.C. High Impedance -- Service Entrance 8 x 20~s 10 kA s.c. Low Impedance 150 

A 1 .- 7 x 50~s 6 !?I! O.C. High Impedance -- I ! 8. Major Feeders n x 20~~ 3 kA S.C. I Luw -... L~avcua< .....-A^--- ,LC 

I and Short 111 

Branch Clrcults U:-L 1-..- .J ---- O.C. rr,~ltIIIIIJC"allCt- 0.5 ps - 100 kHz 

500 A s.c. Low Impedance 

I Long 

Branch Circuits I1 

and Outlets 

6 kV O.C. High Impedance 

0.5~~ - 100 kHz { 200 A s.c. Low Impedance 

I *o.c .: open-circuit voltage s.c.: short-circuit current I 
Recordings and surge counter data have been contributed from 

several sources; in addition to the surge counter data obtained by 

members of the Working Group. Representative oscillograms and 

summary stat~stlcs are reproduced in this appendix, in support of the 

voltage levels and oscillatory wave proposals. 

1. Recordings by Bell Telephone Laboratories 

(Data contributed by P. Speranza, internal report, unpublished 

to date) 

1.1 Typical Surge Counter Statistics 

120 V line at BTL facility in Chester, New Jersey, during 

42 months of monitoring: 

146 counts at 300 to 500 V 

14 counts at 500 to 1000 V 

3 counts at I00 to 1500 V 

3 counts above 1500 V 

l .2 Typical Automatic Recording Oscilloscopes 

0 20 40 60 80 100 120 ps 

120 V OUTLET. LABORATORY BENCH 

2.1 Surge Counter Statistics - Martzloff, F.D. and G.J. Hahn

Voltage in Residential and lndustr~al Power Circuits, iEEE 

6, Ju!y/Augu:t ! 970, ! 049-! 056. 

a) Three percent of all US. residences experience 

occurrences (one per week or more) above 1200 V. 

b) There is a 100:l reduction in the rate of device failu

the withstand level is raised from 2 kV to 6 kV. 

Number of Houses with Repetit~ve 

Surge Activity Above 1200 Volts 

Providence, R. I Cleveland, Ohio Auburn. N. Y. 1,vnehbk Vs. Syracuse, X Y Chicago, Ill. Ashland. Mass 

Fort Wayne, Ind iJeiiali,, ill 

Surge Counter Recordings Above 1200 Volts 

(Sprmg, Summer, and Fall) 

- - - - - - -- 

Vlrmhcr of 

I,oc at >on llornri - 

Prowdenre, I< I 6 

. A.hh,,ro . . . . . , \ . C' . 1 { 

IkKalb, Ill. I I Somenworth, S H. { 

Chicago, Ill. I.' 

Cleveland, Ohm X 

Ikcatw, Ill. 12 

Holland. Mwh. 

o 200 400 600 800 lo00 1200 ps 

277/480 V SERVICE ENTRANCE 

I { 

lynch bur^, Va. { 

Total !I I 

-- - - -- - - 

2.2 Typical Automatic Recording Oscilloscopes 

277/480 V SERVICE ENTRANCE 

FURNACE IGNITION - SINGLE RECORDING 
SERVICE ENTRANCE, LIGHTNING STORM 

STREET POLE. LIGHTNING STORM 

2.3 Simulated Lightning Strokes on a Residential Power Circuit (Laboratory Model of System) - Martzloff, F.D. and K.E. Crouch, 

"Coordination De La Protection Contre Less Surtensions Dans Les 

~gseaux Basse Tension ~&~dentiels," Proceedings, 1978 lEEE 

Canadiun Conference on Communications and Power, 7dCH1373-0, 

pp. 451-454. 

1.5 kA current imoulse (8 x 20 us 

approx.) is injected in round wire 

02 of service drop. 

ents produce flashover of wiring.) 

Recording of open-circuit voltage 

at a branch circuit outlet: 

2200 V peak 500 kHz oscillations 

By connecting a 130 Q load at the 

same outlet (1 A load) the voltage 

is reduced to 1400 V peak, with 

more damping. 

Conclusions from this test series 

I. A current of 1.5 kA (moderate for a lightning discharge 

injected in the ground system) raises the wiring system 

of rhe house 2.2 k'v' above ground. Four kiioamperes krill 

a moderate value) will bring this voltage to 6 kV, the 

typical flashover value of the wiring. 

2. A natural frequency of 500 kHz is excited by a unidirectlonal Impulse. 

3. In this example, the source of the transient (from the 

loading effect of 130 Q ) appears as 

y 

y, 

Surge counter data on various locations In Swm 220 V system

(Data contributed by L. Regez - unpublished to date) 

PEAK VOLIAGE V 

Ser\/ice entrance, !6-fami!u hnrrce, under- I 

--- ground system 

--- Same house, outlet th~rd floor l~ving room 

- - - - - - - - Same house, outlet fifth floor living room 

Service entrance of bank building in Basel 

+- Landis and Gyr Plant, Zug, outlet in lab. ---_-, Landis and Gyr, Zug, outlet in furnace room 

--.- Farmhouse supplied by overhead lines 

Frequency of Voltage Transients per Year as a Function of the Pea

Value of the Voltage Transient for a 220 V, 50 Hz Distribution Syste

w~th Grounded Neutral 

4. Working Group Surge Counter Statistics 

Surge counters with four threshold levels (350, 500, 1000, an

1500 V) were made available to the Working Group by Josly

Electronic Systems, for recording surge occurrences at variou

locations. Members of the Working Group installed these on 120 an

240 V systems of various types, including the following: outlets 

urban, suburban, and rural residences; outlets in a hospital; seconda

circuits on distribution SYS~PEpole^ (recIoser Contr~!~)t secondary 

pad-mounted distribution transformers; lighting c~rcuits In an indu

trial plant; life test racks at an apphance manufacturer; bench pow

suppIy in a !aboratorv I - Limitations on the availability of personnel and cornrnunicat~o

made this sampling less than optimum from a statistical point of view

However, by computing weighted averages for each location, one ca

quote an acceptable overall average; this average has been included 

the graph drawn to establish the low and high exposure limits. 

Summary Statistics of these measurements are as follows: 

I. Data base from 18 locations with a total recording time of 

I2 years spread over 4 calendar years, using 6 counters. 

2. Number of occurrences per year (weighted averages) at 

"average location." 

350 V: 22 

3. Significant extremes 

One home with large number of surges caused by 

washer operarion 

Four locations out of I8 never experienced a surge. 

One home experienced severai occurrences above 

1500 V, with none below that value. 

One industrial location (switching of a test rack) 

produced thousands of surges in the 350-500 V range, 

and several surges in excess of 1500 V. Thls locatlon 

was left out of the average computation, but it 

exemplifies a significant extreme. 
Frorn the data base c~ied In the preceding pages, one can draw 

the chart below, including the following information on voltage vs 

frequency (rate) of occurrence: 

1. The Bell Laboratories data yield a point of 1000 V at 

about 2 occurrences per year (0 ). 

2. The General Electric counter statistics yield a point of 

1200 V at about I occurrence per year ( x ). 

3. The General Electric clock data mdicate a slope of 100:l 

from 2 kV to 6 kV (-.-.-). 

4. The Regez data provlde a band for the majority of 

locations (shown cross-hatched), with the exception of 

the rural location with long overhead line, which has 

more occurrences. 

5. Working Group statistics (0-0-) indicate a less steep 

slope, perhaps because of the influence of outdoor 

locations included in the sample (similar to the rural data 

of Regez). 

The proposed curve, which is the center of the '10 range of 

Fig. 1 In the Guideline, is shown in bold dashed lines (- -). It has 

been drawn at the 100:I slope, passlng near the Bell and General 

Electric p~int: and located -:thin the band of the Regez data. 

PEAK VOLTAGE - V 

APPENDIX I1 - EFFECT OF TRANSIENT POLARITY REVERSALS 

ON SEMICONDUCTORS 

Breakdown of semiconductors under various conditions of load 

5.6 transient o\~er\~oltage app]ication: h;; been inv?stigated.'t Evidence is presented in the two investigations cited that a reverse 

voltage applied during the conduction period of the power frequency 

produces lower breakdown voltage than the app!ica:ion of the same 

translent with no load or during blocking. Examples are glven below, 

taken from these two investigations, showing statistically significant 

differences in the voltage levels. 

iidi i 90 Diode" 

IN21 60 Diode* 

IN679 Diode t 

7- ..-.:-..*-A -- ,--J I r at IXCI I I a L I IU luau 

Fast wave under load 

Slow wave under load 

Transient at no load 

Fast wave under load 

Slow wave under load 

Transient applied at: 

- peak of reverse voltage 

- 25O after start of conduction 

- 90' after start of conduction 

Average 

Breakdown 

1766 

11 8! 

906 

- 155O after start of conduction 1115 

This effect .is one of the reasons for selecting an oscilla

waveform to represent the environment: it will be more llkely 

induce semiconductor failures than a unidrectional wave. Also

shows the significance of the timing of the transient apphcation 

respect to the power frequency cycle. 

*Chowduri, P., "Transient-Voltage Characteristics of Silicon Po

Rectifiers." IEEE IA-9, 5, September-/Giioter i973, p. j82. 

~F.D. Martzloff, internal report, unpublished. 

Franco~s D. Martzloff (M, 1956) was bor

France and received his undergraduate de

at the Ecole Sp&iale ~e- ~&ani~

d'Electricit6 in 1951; he received the M

degree from Georgia Tech in 1952 and 

MSIA degree from Union College in 1971. 

Since 1956 he has been with the 

eral Electric Company, where he first ga

experience In the Transformer and Sw

gear Divisions. Upon jolning General Ele

Corporate Research and Development in 

he became involved In power semicondu

circuits and overvoltaae Drotectlon. He 

", 

participated in the introduction and application of metal o

.,2r;e+n.c vuL La~ui a c:--a JL~~LL I,, la71 L. 

In IEEE Mr. Martzloff has been active on the Surge Protec

Devices Committee and chairman of the Working croup on S

\/,.T+3mec iR ,&,c D -.,,-- P; -,-,. :+- "-+..A LAA \I -- T --- 7VLC~b~= vrcl Ln uuu r "1 LC>>. He is ai

member of the Ad Hoc Advisory Subcommittee of the USA Adv

Committee on IEC SIC 28A. He has been awarded 10 U.S. pat

primarily in the field of varistors and transient protection. 
VARISTOR

VERSUS

ENVIRONMENT:

WINNING

THE

REMATCH
Varistor versus Environment: Winning the Rematch

François Martzloff, Life Fellow, IEEE

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 C62.41.1

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 of

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 can

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
VARISTOR VERSUS ENVIRONMENT: WINNING THE REMATCH 

Francois 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 probiciti aiid sp~ihiiioii of maiching siirge ijioi&i~a 

devices. 

SUMMARY 

During the initial startup of a solid-state motor drive in a 

chemicd processing pian:, dL!m!:ies arcse wI!h !he vxiste: ad 

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 ~ccasicna!!y snge: fzi!u:e when ~YPOECC! to exceptio??~!!~ 

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 to 

drive circuit in a chemical processing plant extending o

square miles. During the initial startup, difficulties aro

varistor and its protective fuse at the input of the thyris

Frequent blowing of the fuse was observed, with occasi

of the varistor. The plant substation, fed at 23 kV fro

utility, included a large capacitor bank with one-third o

"...:+A.c\* -r nr.4 .-.a- *A ...,...:A- ..,....a. Fnr+n. ,,-A ".ro

anlrnlru vu cuau VAL rv pvrnur yvnrn larrvr azru ojur

regulation. These frequent switching operations were s

generating high-energy transients that might be the c

failure of the fuses and varistors, because literally th

similar drive systems have been installed in othe

without this difficulty. On-site measurements perfo

repeated blowing of fuses and occasional la~lure of va

nected at the input to the thyristor drive indicated that 

devices were not matched to their environment. From 

on, specifying larger sizes, sizes appropriate to the en

[ll, solved the problem. 

POWER SYSTEM AND SWITCHING TRANSI

Figure 1 is a simplified one-line diagram of the 

elements of the power system causing the varistor fa

incoming 115 kV power is stepped down to 23 kV. T

of 5400 kVAR capacitors are connected to the 23 kV b

operating conditions involve two banks connected a

with the third bank switched on or off automatically 

voltage regulation. Power distribution throughout the 

at the 23 kV level. 

The various drive systems which experienced the d

supplied at 460 V by a 2300/460 V transformer in th

house. A substation close to the control house s

2300 V power from the 23 kV distribution system. 

SiTE 

SUBSTATION DISTRIBUTION 

SYSTEM HOUSE i 

I 

VARISTO

0885-89/77/86/0004-0059$01.00O1986 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 

A Sweep: 0.2 ms/div 

Vertical: 20 A/dlv 

R Sweep: 0.5 msldl, 

Vertlcsl: 40 A/div 

Figure 5. Current surge bursts during capacitor switching 

Without Varistors With Varistors ic 

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% reachmg 120 A crests, 

for a:: c?lpac;ior baiik swiic:liiig. 

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 131, 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 shows the recorded crests of the five 'nighest voitages i:, each condi:ion; :he di!Te:ence bekeen 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. 

Figure 8 shows the same curves for a proposed 3

tor. It should be noted that the pulse rating does no

strophic failure of the varistor at the end of this rating

10% change in the varistor nominal voltage. Alth

change is indicated, the varistor is quite capable of sta

voltage and of clamping surges. 

Open-Circuit Voitsges 

With 20 mm Varistors 

All Traces: Sweep: 0.4 ntsldiv 

Vertical: 500 Vldiv 

DISCUSSION 

Nature of the Transients 

The absence of any transient (over 120% of no

during ihe 2-hour moniroring period was somewhat s

!he cestex! ef ear!ier repc:!s sf high c~u:,S :cc~ibed w

disturbance analyzers. Frequent checks of threshold 

variations of the possible trigger modes were made, 

the chance of catching an overvoltage, but indeed no

This unusual quiet was also reflected in the chart reco

by the plant personnel, so that the absence of rando

for that period can be accepted at face value. 

Therefore, conclusive evidence was obtained tha

current pulses were absorbed by the varistors duri

switching. The magnitude and duration of these 

excessive for the capability of a 20 mm disc; many s

installed elsewhere do not experience the failures enc

that particular location. 

Another significant finding from these measurem

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 th

Effect of Transients on Varistors 

Published varistor specifications inr!ude the "p&e 

family of curves that define, for each varistor type, th

isolated pulses that a varistor can absorb until its 

reached i4i. The curves show lines relating amplitud

and total number of pulses. Figure 7 shows this fami

for the original 20 mm varistor. 

lYCULSE DURATION - u 

Figure 6. Capacitor switching transients Figure 7. Pulse ratings of 20 mm varistor 
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 A peak eiid 100 CLs (juia~oii (FigGFe 9A): the ci;rves 

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!d be incorrect 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 - --A 4 r-n- .. --notnnt .,nltona nrrnae the nnnlin~nr rrle valrawr; i auu rile I,ral!j curlarsur vvarvbv ar~vuu UA~ I.vl.....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 

vzris!~r 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 rnm 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. 

IMPULSE DURATION-us 

A. Same current, increasing duration 

20 1% I0 MX

IMPULSE DURATION-@n 

B. Same number of pulses, increasing duration 

Figure 9. Reading pulse ratings curves 

/L/ ACTUAL 

Figure 10. Single-pulse equivalent pulse for multiple pu

Note that two types of events occur. One is the prem

blowing of the fuse, which is not caused by a varistor iaiiu

by the i2.t capacity of the fuse being exceeded by the enviro

[I]. The other is the fuse blowing caused by the varistor e

life ultimate failure. 
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 !he time be!ween fel!ures fro.?? 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; :he saiiie of the circiiit, was immadiaieiy 

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 

01" varistors. For instance, the range of nominal voltages for a 

575 V, 32 m.r. ua:iatc: (the next higher vokage ofTered1 is 805 to 

1005 V for 1 mA dc, while the maximum nominal vo!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 

VOLTAGES FOR 510 V 

- - VOLTAGES FOR 575 V 805 1 1005 I 

Figure 11. Tolerances bands of 510 V 

and 575 V varistors 

In addition to the proposed upgrading of prote

460 V level, three other remedies could be consider

tion of surge arresters at the 2300 V level, installat

arresters at the 23 kV level, or a change in the circuit

the capacitor switching, designed to reduce the sev

transients at their origin. 

In general, the protection available from surge ar

to improve when the arresters are installed at higher 

ages. Thus, it is quite possible that arresters ins

2300 V primary of the 23001460 transformer could pr

effective clamping (and at the same time relieve 

energy stress) than the varistors at the 460 V leve

course implied that these would be the zinc-oxide 

arresters.) The full benefit of these arresters dep

configuration of the 2300 V system and its ground

grounded iieiiirai in a wye sysiem, resistance-groun

Boating de!?a! when the arresters "re mnnected i~ 

tional line-to-ground mode. In a second phase of 

described here, 2300 V arresters were installed at the 

primary. A discussion of their expected performance, 

the success of the retrofit, is given later on. 

Likewise, arresters on the 23 kV side could be in

23 kV substation to mitigate the capacitor switching 

their origin, or at the primary of the 23 kVl2300 V su

the control house, where they would also serve 

protection for the overhead 23 kV incoming power 

arresters, again, must be of the gapless type to obta

effective protection. 

The final remedy in the list of alternatives, but per

in effectiveness when the opportunity exists, would b

reducing the severity of the capacitor switching trans

origin. Series inductors or damping resistors may be 

the effectiveness of which would be predictable if a s

the power system behavior were performed by comp

..," 'Wk:l- +h,,+ ----A.. -,... IA --' k6. ,. Lll'LL ,GI.IGUJ wuu HUL be appiied io this pm

tion: it is a remedy that should be cnnsiderac! for n si

exceptionally severe environment. 

The measurements made first with open-circuit, t

20 mm, 510 V varistors on the 460 V side have show

of maximum voltage from 1450 V to 1100 V (Tabl

current of approximately 200 A is flowing in the line 

(Figure 5). 

We can assume that the voltage drop in the line fr

tation and two step-down transformers is mostly 

5.5 kHz, and that the voltage in the varistors can be t

voltage across a resistor at the time of the crest of 

wave. The diagram of Figure 12 shows the relation

the three voltages VoC, VL, and VV, respectively, the 

voltage generated by the capacitor switching action, 

drop in the line and two transformers, and the varist

the current peak. Treating this highly nonlinear circu

circuit is an apprnrimation !ha! wi!! prnvide P! each 

full range of voltage and current conditions a valid ord

tude for the purposes of this discussion. Numerical 

avaiibie for rigorous treatment at any instant over t

of coiidi:ioii~ :6!. With iht shpil'ying assumpiion, w

mine the order of magnitude of the 5.5 kHz curren

flow in an arrester installed at the primary terminals 

Vl460 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 ~i!heut p:eb!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 

(11 F.D. Martzloff, "Matching Surge Protective Devices to their 

Environment," Proc. IEEEIIAS Meeting. October 1983, pp. 

387-392. (Also scheduled for IAS Transactions, Jan/Feb 

1985.) 

[2] ANSVIEEE Std C62.33-1982, IEEE Stan

Spect$cations for Varistor Surge Protective Devices, 

tute of Electrical and Electronic Engineers, Inc., 

[3] 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. 

161 H.W. Dommel, "Digital Computer Solution of E

netic Transients in Single and Multiphase Netwo

Transactions on Power Apparatus and Systems, Vo

pp. 388-399, April 1969. 

Francois D. Martzloff (

was born in France, and 

undergraduate degree at 

SpCciale de MCcanique et 

in 1951; he received the 

,-:,."I c..":..a--:.." --- 

.Lb. b11(5111GGlllllj deg~

gia Tech in 1952 and 

Industrial Administration 

Union College in 1971. 

Since 1956 he has be

General Electric Compan

gained experience in the 

and Switchgear Divis

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 particular instance, it 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 reirofiiiina - corrections 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 aid 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-eneraizina. the latter involving the risk of restrikes. Because 

of this availabi<ty Gd the limited space available in the mucno~s on one hand. and because of the limitations in scoDe of the field retrofit 

mission on ihe 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, 

vv. 2326-2333. - - 

Protection requirements and surge arrester duties are analyzed for 

both diairai - and iransieni neiwurk arlaiy~er -.. ~'Tiu'Aj suiiuiaiioas. Suiip!c &n~n_a!pirA expressions =e developed for evaluating arrester duty 

as a function of capacitor bank size. Guidelines and limitations for 

applying arresters at grounded- and ungrounded-wye capacitor banks ae based oil overvokEge c~larac~er;spics &-ies;er 

capabilities. 

[2] J. H. Brunke and G. G. Schoeckelt "Synchronous Energization of 

Shunt Capacitors at 230 kV," IEEE PES (Power Eng. Soc.) 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 perfokce. 

[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 unbaiance due to fuses biowiiig, and iimiionics 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 (motor 

capacitor switching, arc and other (furnace-switching, tran

switching, mine-winder-reversers). Surge generation by vacu

tactors is considered under the following heads: basic ene

and de-energization transients; transients generated when s

inductive loads; assessment of a vacuum contactor from 

generation point of view; types of surge suppression devic

ching of capacitive loads. An extensive discussion of the 

appended. 

[5] Jack R. Harbaugh and John E. Harder, "Important Consid

for Capacitor Applications in the Petroleum and Chemical 

Industries," IEEE Pet. Chem. Ind. Conf., 27th Ann. Rec. 

Pap., Houston, TX, Sept. 8-10, 1980; IEEE, #80CH154

Piscataway. NJ, pp. 157-167. 

The location of capacitors may have a significant effec

(I**2R losses within the plant transformers and conductor

is wasted energy. Transients generated by capacitor switch

require attention in the selection of arresters, system insul

ofier equipment. ofcap&iors artly require som 

attentinn to !&rge motors during system r=!&ng or t

This study addresses each of these considerations, providi

guidelines for effective, reliable capacitor application. In a

a c:l&iisi is provided for indiisirid capaciior app&

[6] J. F. Burser, R. J. Santoro, J. W. Stolle. R. E. Owen an

Clinkenbeard, "Comparative Evaluation of Field Test D

Computer Results on Capacitor Switching Transients." 

Minutes PA Electr. Assoc., Eng. Sect., Trans. and Distrb. 

West Middlesec, PA, May 15-16,1979; PAElectr. Assoc., En

Harrisburg, PA 1979. 

An arrester failure case was analyzed occurring at the 

substation capacitor bank switching with the use of a trans

work analyzer (TNA) and digital computer techniques. The 

volved a 12.5-kV ground-wye distribution sytem. Results of t

study, which were validated by field tests, showed the effects o

configuration magnitude of transient over voltages. 

[7] Eldon J. Rosers, and Don A. Gilles, "Shunt Capacitor S

EM1 Voltages, Their Reduction in Bonneville Power Admin

Substations," IEEE Trans. Power Appar. Syst. PAS-93, 

Nov.-Dec. 1974, pp. 1849-1860. 

Back-to-back switching of grounded wye shunt capacito

high frequency, high-magnitude current flow in overhead 

___..->___. .,___._____PI_>..__> ___l.____ 

~~UUIIU ma~ WIIUUSLU~S. IVLS~SU~CIIICIILS UL IIIUUSCU vui~agc

trol cables and receptacles; transverse voltages on fuse bl

secondaries and personnel intercept voltages are reporte

methods of confining transients to capacitor areas and s

*--I.-: --..:-.*."a LGcIlluqusJ as Isvlswsu. 

[8] Paul C. Krause and William C. Mauser, "On-Line Transie

trol of Capacitor Switching to Improve System Stability,

Trans. Power Appar. Syst. PAS-92, No. 1, Jan-Feb. 1

321-329. 

A simplified voltage-reactance equivalent of the one-mac

fhte bus system was used. The material presented in this pap

that this simplified model does not predict the performance o

machine infinite bus system with the accuracy necessary to d

the capacitor switching times so as to achieve the control ob

However, it is shown that computation accuracy may be im

by including system losses. It appears that in order to apply 

control techniques it will be necessary to develop more 

models of the power system components. Also, faster than r

iterative, on-line computation techniques as simulated in th

should be implemented and used to calculate the switchin

Until these obstacles are overcome, optimal control theory w

little impact upon the power industry. 

[9] 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 ~A~~

Results;" IEEE Trans. PowerApp. Syst. PAS-90, No. 4; Ju

1971, pp. 1639-1645. 

Station capacitor bank energization transients predicted 

cuit model are compared to field-test results. Selection of a 

model is discussed. A computer solution of the model is pr

Manuscript received September 19, 1985. 
REAL,

REALISTIC

RING WAVES

FOR

SURGE

TESTING
Real, Realistic Ring Waves for Surge Testing

François Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

Guiseppe Pellegrini

Italian Electricity Board (ENEL)

Automatica Research Center

Milano, Italy

Reprinted, with permission, from

Proceedings, 9th International Zürich Symposium on Electromagnetic Compatibility, 1991

Significance:

Part 2 Development of standards - Reality checks

Part 5 Monitoring instruments, laboratory measurements, test methods

Standards for surge testing have a long tradition of using unidirectional waves (“impulses”), in particular a 1.2/50

voltage impulse and an 8/20 current impulse. Many surveys of surge activity in low-voltage AC power circuits have

shown that a large number of recordings actually show oscillatory surges rather than unidirectional surges. 

This paper provides examples of such waveforms, independently recorded by two organizations, one in the US and

one in Italy. These examples draw on field measurements as well as laboratory experiments and are offered to

make the case that Ring Waves should be included in the regimen of electromagnetic compatibility tests.

1






INCOMPATIBILITY

BETWEEN

THE 100/1300

SURGE TEST

AND

VARISTOR

FAILURE RATES
Incompatibility Between the 100/1300 Surge Test

and Varistor Failure Rates

Charles Fenimore and François Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

Reprinted, with permission, from

Proceedings, 9th International Zürich Symposium on Electromagnetic Compatibility, 1991

Significance:

Part 2 Development of standards – Reality checks

Demonstration ad absurdum: 

Accepting the premise of prevalent 100/1300 high-energy surges and modeling the response of typical metal-oxide

varistors leads to the conclusion that most of the billions of varistors in service should fail at alarming rates – but

we know they do not. Ergo, the premise is not valid. 

(See also paper “VDE 0160" in this Part 2 for an experimental demonstration.)




is based, a ready calculation using the capacitively 

stored energy, q2/2C (of order 5000 J), shows that 

no more than 10% of the stored energy is spent in 

the MOV in the simulations according to the two approaches discussed above. Yet, these two tests are 

aiready destructive or' the device. it seems iikeiy that 

a test that would meet the 80% criterion would provide an even more severe stress to the equipment, and 

provide a greater disparity between the model results 

and field experience. 

Thus, the authors suggest that a reexamination of the 

premises that led to the VDE 0160 Standard should 

be considered before incorporating a blanket requirement for such a test into new IEC surge immunity 

standards. The authors plan to perform actual tests 

on typical varistors to further support the computations presented in this paper. 

CONCLUSIONS 

1. A mathematical and derived computational model 

has been presented which permits the evaluation of 

many aspects of varistor performance over a range of 

conditions which are characteristic of the actual operating environment and also of the test environment 

contemplated by VDE 0160 and other surge standards. 

2. Computer model predictions of the impact of the 

proposed 100/1300ps surge test on the millions of 

varistors in service shows that these varistors should 

experience a greater failure rate than indicated by 

available information on actual failures. The simplified inductance-free model provides analytical confirmation of this result. This inconsistency raises serious 

questions on the proposed requirement of such a severe test to a wide range of equipment. 

3. The lingering ambiguity on setting a constant opencircuit voltage or ai?justing the voitage whiie the specimen is connected needs to be clarified. A constant 

open-circuit voltage is the generally accepted practice in surge testing. The premises that led to this 

new surge test may justify adjusting the charging voltage after the test specimen has been connected to the 

surge generator; that adjustment, however, results in 

larger amounts of energy being dissipated in surge protective devices, making the apparent incompatibility 

identified above even greater. 

4. The criterion that 80% of the capacitive energy 

must be transferred to the test specimen may be difficult to satisfy and needs clarification. The authors 

have been unable to identify a simple circuit which 

satisfies the criterion while maintaining the required 

rate of decay under open circuit conditions. 

5. While the authors do not question the 

the fuse-blowing scenario, they recommen

review of the statistics of the occurrence of 

ing, of the use of varistors with low clam

age, and of the distribution of actual clam

age within manur'acturing tolerances. The 

model developed in this paper may be a 

in evaluating the effect of these tolerance

thors also urge all users to share informa

observed failure rates, as well as to perfo

ing tests, in order to provide a broader per

these issues. 

REFERENCES ' 

IEC Document 77B(Central Office)4, O

Electromagnetic Compatibility - I~n~

MEISSEN, W., ~bers~annungen i N

nungsnetzen, Elektrotechnische Zeitsdlr

1983. 

VDE 0160 Ausriistung von Starkstrorn

elektronischen Betriebsmitteln, May 1

quotations appearing in this paper from 

this standard are made from the ofic

translation available from the British St

stitution). 

VDE Ol6OAl ~nderun~ 1 zu DIN VDE 

April 1989. 

MEISSEN, W., Private communication

WEISINGER, J., Hybrid-Generator f

lationskoordination, Elektrotechniscl~

vol. 104, 1983. 

ANSIIIEEE C62.45-1987, Guide on 

ing for Equipment Connected to Low-

Power Circuits. 

MARTZLOFF, F.D. and LEEDY, T.E

Varistor Clamping Voltage: Lower is n

Proceedings, 1989 Ziirich EMC Sympo

Transient Voltage Suppression Device

CORP., Somerville, NJ, USA, 1990. 

E. ARGON, I.L. CHANG, G. GUNARA

KAHANER, and M. A. REED, "Mathem

ware: PLOD," IEEE Micro, vol. 8 (4), A

'Certain sofiware is identified in ihis paper in 

equately specify the procedures used to produce t

results. Such identification does not imply thit th

necessarily the best available for the purpose. 

TESTING

VARISTORS

AGAINST

THE VDE 0160

STANDARD
Testing Varistors Against the VDE 0160 Standard

François Martzloff, Life Fellow, IEEE

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

Reprinted from Proceedings, Open Forum on Surge Protection Application, NISTIR4654, August 1991

Significance:

Part 2 Development of standard – Reality checks

Demonstration ad absurdum: 

Accepting the premise of prevalent 100/1300 high-energy surges and subjecting typical metal-oxide varistors to the

stress from a test performed with a prototype generator leads to the conclusion that most of the billions of varistors

in service should fail at alarming rates – but we know they do not. Ergo, the premise is not valid. 

(See also paper “Validating Surge Tests ..." in this Part 2 for a demonstration by numerical modeling.)
Fran~ois D. Martzloff 

National Institute of Standards and Technology 

Abstract - High-energy surge tests have been pe$ormed on metal-oxide varistors of a type in common use

a proposed ZEC standard derived from German Standard WE 01 60. l%e surge generator used for the test w

commercial device developed especially to deliver the I00/1300 ps waveform speczjied by WE Standard 016

on the position of the varistor within its manufacturing tolerance band, failure or degradation can occur, 

concern that this test requirement may be too severe for universal application. 

INTRODUCTION 

Coiicerm hLle occ-urrence of associaid .*kh ciirreiit-:iiiiiiifiije operatio

1983 jijj have id ~e German (TVQEj to specify a iiigii-energy siirge test 

to electronic equipment installed in industrial environments (VDE 0160, 1988 [2]). Essenti

requires discharging into the ac line interface of the equipment under test (EUT) a capacitor of s

that the specified waveform is generated, initially charged at a voltage suitable for producing a peak 

the power-system sine-wave peak (Figure 1). Technical Committee 77 of the IEC has included t

menu of surge immunity tests (TC77BlWG3, 1990 [3]), without limiting the scope of application 

environments intended by the Meissen paper. Thus, this test is likely to become a general requirem

on commercial and consumer equipment, unless its implications are recognized. In the absence 

available surge generator, computer modeling of the test had previously been performed (Fenimore 

1990 141, 1991 [5]). The findings of these simulations have shown that typical varistors, of 

millions have been installed and continue to operate satisfactorily, cannot survive the proposed IE

because excessive energy would be deposited in these varistors during the surge. The recent ava

prototype surge generator made it possible to subject typical varistors to the VDEIIEC surge, as rep

Schaffner*, a manufacturer of surge generators, has now developed a prototype that can produce th

surge; in response to an invitation to try out this prototype, an informal work session was cond

Schaffner facility to subject typical varistors to the VDE 0160 surge. The generator includes 

capacitor, up to 6000 pF, the necessary dc supply to charge the capacitor, a 220-V ac supply (f

environments), and suitable means to decouple the test specimen circuit from the laboratory ac sys

of the circuits are still proprietary, and only the output of the generator is described in thi

chronological recitation of the work session would require first a discussion of the various consid

conditions of the test. Recognizing the natural curiosity of the readers, let it be stated here that 

was destroyed during the test, and the other (barely) survived, consistent with the predictions of t

modeling. Having thus given away the outcome, let us now proceed with the detailed recitat

ronsid_eration; and conditions; 

Y As a poiicy, the idaiiond irisii'iciie of Standards and Teciiiiologii disci'ai~ms aiiy i~xp'ieii edoiseiiieii: of a coiiiiiieicia

identifying such products for the sole purpose of adequately describing the equipment used in the experiment. In this 

the prototype generator used in the tests was the only one known to be available. Furthermore, there is no certainty tha

offer a commercial product based on this prototype. 
in the digital storage oscilloscope did not have the capability of computing the power (i x v) dissipated in the 

varistor and integrating it into total energy deposited. Manual integration of the recorded traces was performed 

after the tests. This computation yields results of sufficient magnitude (that is, large overstress of the varistor) 

to make precise computing unnecessary in evaluating the outcome of the test. 

The test specimens (EUT) were 20-mm diameter varistors, consisting of two 130-V rms rated devices connected 

in series, a good approximation of the practice of applying 250-V rated varistors in the 220-V equipment used 

in Europe (Martzloff and Leedy, 1989 [6]). The nominal voltage, V,,,, of each varistor (voltage measured 

with 0.5 rnA or 1 mA dc injected in the varistor) was determined before the test for each device. One varistor 

pair (referred to as EUT #1) had a nominal voltage of 392.6 V, the other pair (EUT #2), 399.5 V. The 

nominal voltage for a 250-V rms varistor is 390 V, the minimum 354 V, and the maximum 429 V (Harris 

iVianuai, 1996 171). nus, ET~T is siiuaid at 1 % i?"oove the nominal of a 250-xv7 

EUT u2 is at 2.5% above the nominai vaiue. 

To test the varistors under the worst case condition (that is, the varistor at 10% beiow nominai, thus drawing 

energy from the generator for a longer portion of the surge waveform), the test voltage should be raised above 

the voltage specified for nominal test conditions. To place the varistor under conditions equivalent to those 

prevailing for a -10% specimen, a varistor at some tolerance level must be subjected to the same current as that 

occurring for a -10% varistor at the nominal test voltage. With the nominal VDE 0160 test voltage of 2.3 times 

the 220-V peak (714 V), the available EUT varistor specimen can be tested in a manner equivalent to a -10% 

tolerance varistor by raising the test voltage. 

For EUT #I which is 1 % above the V,,, of a 250-V rated varistor, the test voltage should be 10% higher than 

the nominal 714-V peak, plus I%, that is, 792 V. For EUT #2, 2.5% above the V,,,,, the test voltage should 

be 12.5% higher, 803 V. This increased test voltage will place the varistor at the correct value of current on 

its I-V characteristic, but raises the power dissipated in the varistor by the same percentage. Thus, the energy 

deposition in varistors other than -10% tested under the artificially raised test voltage received 11 % or 12.5% 

more energy than what a varistor at -iO% wouid have received. However, considering tine energy ieveis 

observed In the tests reported beiow (about 200% of rated ieveis, this ii-i2.5% excess does not affect t'he 

conclusions. The significant parameter to be observed is the current level, and that correct level was ~ndee

achieved by raising the test voltage. 

The VDE 0160 document states that the specified surge test voltage should be maintained across the terminals 

of the EUT, rather than the usual method of having a preset open-circuit voltage, and then connect the EUT 

without changing the generator setting (the so-called 'let-it-rip' mode [5], and (ANSUIEEE C62.41-1987, 

[8]). Meissen confirmed this interpretation of the document [9], so that the charging voltage of the 

generator capacitor was increased toward obtaining the specified voltage with the EUT connected, using an 

expendable EUT varistor during preliminary tests. However, the prototype generator output voltage, with 

maximum charging voltage and with varistor connected, could only be raised to 774 volts (Figure 2) instead 

of the 792 V or 803 V necessary to place the #1 and #2 varistors in the -10% tolerance situation. Thus, EUT 

#1 was actually tested in a condition corresponding to 7741792 = 98% of the worst case level, and EUT #2 at 

7741803 = 96% of the worst case level. In other words, EUT #1 was tested as if it were at a -8% tolerance 

level, and EUT #2 at a -6% tolerance level with respect to a 0% tolerance on their V,,,. 

The manufacturer's specifications 171 show a 70-5 single-pulse energy rating for the 130-V varistor, or 140 J 

for two in series. Figure 3, from Ref 151, shows the predicted energy deposition as a function of the varistor 

position in its tolerance band, for the test condition where the voltage is maintained across the EUT by 

readjusting the surge generator charging voltage. 
set up by Schaffner included approximately 5 m of leads "suitable for a 16 A load" between th

the output of the generator. Thus, the impedance presented by the test specimen to the applied VD

includes a resistance that will reduce the stress of the varistor; however, this reduction is not readi

by the simple mention in the figure of a 5-m maximum lead length, and the cross-section of the 

not specified. Operators can interpret the test procedure in a way producing maximum stress (a 

large cross section) or a minimum stress (maximum of 5 m of leads with small cross section). 

In accordance with the interpretation of the Figure 4 diagram, the voltage measured and shown 

the total of the voltage developed across the varistor and the lead drop. To evaluate the implic

interpretation, the next test was performed, without changing the generator setting (at its maxim

voltage), with the voltage measurement made at the varistor terminals (Figure 5). Note the 700-V 

test, GI- a 74-V difference (!O%Or-) from th,e value recorded is Fig~re 2. 1: the mode!ing of Refer

[5], tk,e effect ~f t!~is 5-112 test !ead had not been inchibed, so that the c~iiclti~i~iis of the ~iiob

m.I pessimistic than the conseqlieiices of a test coiiditioii wi;4 a lead length included. IIIUS, the vari

iiiider 16% less voltage stress (keep in mind the noniinear relationship between voitage and cur

model prediction, and possibly could survive. 

THE DEATH OF A VARISTOR 

According to a subsequent amendment to the VDE 0160 test specification, the maximum* capac

the duration of the surges may be reduced to 300 ps for equipment installed in circuits protecte

less than 35 A continuous rating. This reduction will provide significant relief to varistors inc

industrial environments. However, the IEC document [3] does not include that reduction. The 

for EUT #1 included two surges with this reduced stress (Figure 6), followed by surges with the 

capacitance and full 1300 ps duration, at the maximum available generator voltage, as shown 

Before and after each surge, the varistor Vnom was recorded to track any shift in characteristics, 

to the maximum shift of 10% allowed in the manufacturer's specifications. 

The test sequence and results for EUT #1 (a specimen in the -8% tolerance p~sition)~ starting 

surges applied: were the following: 

Shot 1: 718 V crest, 400 ps duration, Vnom shift of 1 % (Figure 6) 

Shot 2: 768 V crest, 1100 ps duration, VnOm shift of 1 % 

Shot 3: 774 V crest, 1400 ps duration, Vno, shift of 1 % (Figure 2) 

Shot 4: Repeat, same settings as shot 3 (voltage measured at varistor, Figure 5), 

Varistor (a) of pair punctured 

Varistor (b) of pair externally intact, but VnOm = 0 (short circuit) 

Energy deposited in the varistor; approximately 300 J (215% of rating) 

The same test sequence was then applied to EUT #2, that is, first two shots at reduced stress, 

stress for shot 3 and four additional shots. The V,,, shift grew from 1 % after the first shot to 6% 

shot, as measured after cooling down following the test. By the time the author had returned 

States (20 days iaterj, the snift in Vnom, determined by more systematic measurement at NIST, w

4%. 'The difference between the 6% immediately after the test and the 4% after 20 days may be 

a slow recovery of the material, or a difference in the precision of the measurements, or both. 

* The surge duration is the specified parameter in the VDE 0160 document, therefore the required value of the capac

dependent upon the impedance of the EUT. 
DISCUSSION OF THE RESULTS 

From the simulation predictions, it was expected that the varistors would be destroyed by the test, even though 

the (late) realization of the stress reduction provided by the lead length does somewhat change the situation. 

In other words, the 10% loss of voltage caused by the leads places the varistors used in these tests at 

respectively +2% and +4% in the tolerance band, a condition that the prediction describes as marginal 

survival. The joule rating specified by the manufacturer tends to be conservative, so that it may take more than 

140 J to destroy a varistor. Furthermore, a larger population of test specimens may produce a distribution of 

more failure as well as more survivals as only two test points can only provide an indication, not a certainty. 

However, the conclusion is clear, that varistors of common use in commercial and consumer equipment would 

be in severe jeopardy if the full 10011300 us surge were applied, even with the mitigating effect of the 5-m lead 

--.. length. Discussing the test results with Meissen, we agreed on the following conclusions: 

1. There is no disagreement that the basic phenomenon of fuse blowing can lead to the high-energy surges 

described by Meissen in the heavy industrial environment (circuits with fuses above 35 A). 

2. The prediction of varistor failure through modeling is consistent with the tests; the mitigating effect of the 

allowable EUT lead reduces the forecast of widespread failures, but varistors in the lower tolerance bands are 

still at risk. 

3. The amendments to VDE 0160 providing for reduced maximum capacitance values (see the footnote on page 

3) and reduced duration make the test more realistic. Further evaluation of these reduced stress levels would 

show appropriate limits of application. 

4. However, this stress reduction has not yet been acknowledged by the IEC proposals (Figure 1; showing only 

one value of 1.7 ms is excerpted from the IEC documenti not the amended VDE 0160 where the alternate 

dmitinn nf 0.3 ms is shnwn). This paper is therefnre s~lhmitted tc! the engineering cnmm~nity at !arge as a 

rec~rnrnefidati~fi ~f !kiting the fi~!! duratim ~f a I300 ps surge and its high energy tc the industria! 

environment for which it was first proposed. 

5. The concept of readjusting the surge generator charging voltage to maintain a specified test voltage across 

the specimen is different from the usual practice of maintaining a fixed open-circuit voltage for the generator. 

However, it may be compared to the practice of readjusting the surge generator used for surge arrester tests at 

a specified test current level. As long as the implications of the procedure are recognized, either method may 

be suitable, if uniformly interpreted. 

6. In its present form, the VDE 0160 document leaves open the possibility of different interpretations by 

different operators. Should the principle of a high-energy test be adopted by the IEC, more detailed 

specifications need to be developed and agreed upon by interested parties. 
implications are gratefully acknowledged. 

REFERENCES 

Meissen, W., uberspannungen in Niederspannungsnetzen (Overvoltages in low-voltage 

Elektrotechnische Zeitschrift, Vol. 104, 1983. 

German Standard VDE 0160, Ausriistung von Starkstromanlagen mit electronischen Betreibsmitteln (

with electronic operating controls for use on power systems), May 1988, amended April 1989. 

IEC Draft Standard DIS 77B(CO)lO(c), Electromagnetic Compatibility - Part 4: Testing and 

techniques - Overview on electromagnetic compatibility tests, 1991. 

Fenimore, C. and Martzloff, F.D., Validating Surge Test Standards by Field Experience: High-Energ

Varistor Performance, Conference Record, IEEE/IAS Annual Meeting, October 1990. 

Fenimore, C. and Martzloff, F.D., Incompatibility Between the 10011300 ps Surge Test and Varis

Rates, Proceedings, 1991 Ziirich EMC Symposium. 

Martzloff, F.D. and Leedy, T.F., Selecting Varistor Clamping Voltage: Lower is no

Proceedings, 1989 Ziirich EMC Symposium. 

Transient Voltage Suppression Devices, Harris Corporation, 1990. 

IEEE Guide on Surge Testing-for Equipment Connected to Low-Voltage AC Power Circuits, AN

C62.45-1987. 

Private communications, March 15 and May 15, 1991. 
Figure 1. High-energy waveform specification (From Ref. [3]) 

Vdts I 

Amperes 

800 

400 

kf 

7' 

'-i1r-7- - .... - - .. 2 ms/div 

Figure 2. Voltage across and current through EUT #1 
() 

loo0 i 

-10% 

- - ___--- 0% 

Energy into EUT #1 

for shots #3 2% ii4 

/I ,-- 

' A Maximum allowable energy /I/. // 

Figure 3. 

Energy deposited in varistor as a function of tolerance of device compared to nominal value 

(From Ref. [q) 

Test Measuring instrument 

equipment e.g. oscilloscope 

Sample L1 ..-----c3---- 

Supply system EE 

(L2)N o A 

X2 Supply cable 

(max. 5 m) 

Figure 4. Elementary test circuit diagram (From Ref. 121) 
Volts 

800 

400 

0 

Amperes 

UVV onn

400 

0 

Figure 5. Voltage across and current through varistor only 

Volts 

Amperes 

800 

400 

0 

Figure 6. Voltage across and current through varistor with reduced stress 
VALIDATING

SURGE TEST

STANDARDS

BY FIELD

EXPERIENCE:

HIGH-ENERGY

TESTS AND

VARISTOR

PERFORMANCE
Validating Surge Test Standards by Field Experience:

High-Energy Tests and Varistor Performance

Charles Fenimore, Member, IEEE and François Martzloff, Life Fellow, IEEE

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

© 1992 IEEE

Reprinted, with permission, from

IEEE Transactions on Industry Applications, Vol.28, No.6, November/December 1992

Significance:

Part 2 Development of standards – Reality checks

Demonstration ad absurdum: 

Accepting the premise of proposed IEC 100/1300 high-energy surges being representative of the environment, and

modeling the response of typical metal-oxide varistors, leads to the conclusion that most of the billions of varistors

in service should fail at alarming rates – but we know they do not. Ergo, the premise is not valid and the proposed

high-energy test should not be considered as an across-the-board requirement. 

(See also paper “MOV - VDE" in this Part 2 for an experimental demonstration.)

On the other hand, the tests proposed in IEEE Std C62.41 would not result in systematic failure of commonly used

varistors and consequently appear more realistic
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 28, NO. 6, NOVEMBER / DECEMBER 1992 

Validating Surge Test Standards by Field 

Experience: High-Energy Tests and 

Varistor Performance 

Charles Fenimore, Member, IEEE, and Fran~ois D. Martzloff, Fellow, IEEE 

Abstract-New, high-energy surge tests are emerging in IEEE 

and IEC standards. Field experience offers a valuable criterion 

for validating or invalidating proposed standards. A proposal 

under consideration by the IEC involves so much energy that a 

varistor of the voltage rating commonly used in protecting load 

equipment, if subjected to this test, would almost certainly fail. 

Yet, reported varistor failure rates do not reflect such a situation. Thus, a reexamination of the premises that led to the 

proposed test specifications appears necessary. Proposals for 

high-energy tests as additional waveforms in the new version of 

IEEE C62.41, on the other hand, lead to current and energy 

levels that do not place typical varistors in immediate jeopardy. 

Thus, they appear more consistent with field experience. 

NATURAL approach in defining the surge tests to A, e performed on any equipment is to attempt duplicating the conditions observed in site measurements. 

However, this approach would lead to a situation where 

general conclusions are drawn from limited measurements of specific surge occurrences. It has, in fact, led to a 

multitude of proposals for test standards that may subsequently be applied outside of their original, correct context because no other standard is available at the time. An 

example of this situation may be developing with the 

proposal by Technical Committee 77 of the International 

Electrotechnical Commission (IEC) for a high-energy 

100/1300-ps surge test. 

To evaluate the effects of various proposed or existing 

high-energy stress tests on commonly used varistors, this 

paper presents a simple yet effective model of a surge 

generator. The evaluation proceeds by quantifying the 

current through the varistor and the corresponding energy 

deposited in the varistor. The computed results are compared with the published device ratings to predict the 

likelihood of failure. This likelihood is then compared 

Paper IPCSD 90-40, approved by the Power Systems Protection Committee of the IEEE Industry Applications Society for presentation at the 

1990 Industry Applications Society Annual Meeting, Seattle, WA, October 7-12. This work was supported by the Building Industry Consulting 

Service International (BICSI) and by the US. Army Research, Development, and Engineering Center. Manuscript released for publication 

September 24, 1991. 

The authors are with the Electricity Division, National Institute of 

Standards and Technology, Gaithersburg, MD 20899. 

IEEE Log Number 9203290 

with the available information from field experience on 

failure rates. 

Any immunity test should be conducted with an objective that is more subtle than the goal to "duplicate the 

environment." A test stress is applied to a device not to 

demonstrate that it can survive any of the stresses 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-on, simple, and reproducible 

stress. From surviving the test stress, the inference is 

made, subject to confirmation by field experience, that the 

device does have the ability to survive the infinite variety 

of stresses that it will encounter during its life in the real 

world. In other words, simple test stresses are useful 

because they can be reproduced over a period of time at 

the same facility and between facilities, providing a common language and a standard of comparison that is essential to conduct orderly transactions. Test standards should 

not, however, be misconstrued as representing natural 

phenomena. They are effective only if they discriminate 

between those devices with a potential for long field 

survival and those that are likely to fail. 

The proposed 100/1300-ps IEC test should be reexamined with this philosophy in mind because it appears that 

commonly used varistors would be expected to fail when 

subjected to this test. Anecdotal experience does not 

support the prediction of failure in the field, raising questions as to the general validity of this test. On the other 

hand, high-energy tests derived from new proposals contained in the revised version of IEEE Std. C62.41 do not 

lead to contradiction between field experience and predicted test results. 

PROPOSED IEEE AND IEC HIGH-ENERGY TESTS 

Metal-oxide varistors that suppress surges by absorbing 

energy have proliferated in low-voltage ac power circuits. 

Consequently, new high-energy tests have been proposed 

to assess the ability of these varistors to withstand the 

corresponding stress. In a major revision of the IEEE 

Guide C62.41 [I] (emerging as a Recommended Practice 

[2]), an additional waveform has been proposed to assess 

this ability. The proposal is a 10/1000-ps surge, with 

three "system exposure" levels, which are defined below. 

The IEC Technical Committee TC77 is considering a 

0093-9994/92$03.00 0 1992 IEEE 
FENIMORE AND MARTZLOFF: VALIDATING SURGE TEST STANDARDS BY FIELD EXPERIENCE 

surge test requirement based on the scenario of currentlimiting fuses clearing a fault at the end of a cable, where 

the energy trapped in the system inductance causes a 

large transient at the time the fuse interrupts the current 

[31. That scenario was first described and quantified by 

Meissen [4] and incorporated in German Standard VDE 

0160 [5]. 

The new C62.41 Recommended Practice proposes, 

among other waveforms, a high-energy stress defined by 

an open-circuit voltage and a source impedance at three 

"system exposure" levels. For the "low exposure" level, no 

high-energy stress is proposed; for the "medium exposure" 

level, the surge environment involves a crest of two times 

the system peak voltage with a source impedance of 1 0. 

For the "high exposure" level, the crest is 2.3 times the 

system peak voltage, whereas the source impedance is 

only 0.25 0. 

The IEC proposal appears to be based on the VDE 

0160 standard, which specifies the direct discharge of a 

large capacitor-thousands of microfarads-into the 

equipment under test (EUT). The VDE test procedures 

are not quite clearly outlined at this point but might be 

interpreted as readjusting the capacitor charging voltage 

after connecting the EUT to the surge generator in order 

to maintain the specified test voltage across the EUT. 

That approach would be diametrically opposed to the 

generally accepted practice of performing a surge test 

with a generator having the capability of delivering a 

well-defined open-circuit voltage and short-circuit current 

or an open-circuit voltage associated with a specified 

source impedance (see Fisher and Martzloff [6] and the 

IEEE Guide on Surge Testing [7]). 

Another ambiguity in the VDE 0160 test specification is 

that it might be acceptable to perform a test where the 

voltage waveform is less than the specification, provided 

that 80% of the energy stored in the surge generator 

capacitor is delivered to the EUT. However, there is no 

provision in the test procedure for measuring this energy, 

and it is doubtful that this condition can be achieved with 

a surge generator containing the parallel resistor that is 

necessary to achieve the specified rate of decay (or duration of the tail of the wave) when the EUT offers a high 

impedance. 

Metal-oxide varistors offered by manufacturers include 

ratings of 130 V rms for applications in 120-V systems and 

250 V rms for application in 220-V systems. The motivation for using these varistor ratings is, of course, the 

desire to provide the lowest possible clamping voltage to 

protect sensitive equipment. A paper presented at the 

Ziirich EMC Symposium suggests that premature varistor 

aging may result from this close clamping (see Martzloff 

and Leedy [8]). However, the 130- and 250-V varistor 

ratings are still widely used by equipment manufacturers 

who take the position that they are not afflicted by unacceptable failure rates. Thus, the authors accept that position as reflecting actual field experience and will apply it 

as a criterion for validating or questioning the proposed 

This paper reports the results of modeling the application of a surge test to a family of commonly used varistor 

sizes (14, 20 and 32 mm in diameter). For each varistor 

size, the computations were performed for three levels of 

manufacturing tolerances on the varistor: nominal value, 

- lo%, and + 10%. A varistor with its clamping voltage at 

the maximum acceptable tolerance level (the level shown 

on published I - V curves) will tend to absorb less energy 

than a varistor with a lower clamping voltage because it 

will divert current for a smaller part of the surge. The 

maximum energy deposition in the varistor will occur for a 

varistor having the lowest acceptable clamping voltage 

(typically 20% below the maximum), as indicated by the 

10% tolerance on varistor nominal voltages. Should the 

test generator parameters be at the most severe conditions within its uncertainties (higher peak voltage and 

longer duration than nominal within allowable tolerances), 

the stress on the varistor would be even greater. 

The circuit model used in the computations reported in 

this paper is a simple capacitor-discharge circuit that can 

produce the 10/1000-ps waveform of C62.41 or the 

100/1300-ps waveform of VDE 0160, where each has the 

appropriate selection of the components values. The modeling results, which are discussed in detail below with 

supporting information in the Appendix, indicate that the 

smaller size varistors would not be damaged at the 

"medium exposure" level of C62.41 but would be damaged at the "high exposure" level. The 32-mm varistor 

would easily accept several applications of the "high exposure" level, whereas the 20-mm varistor would have a 

limited life. On the other hand, few varistors will survive 

the VDE 0160 stress. 

Table I presents this information in the form of the 

number of surges that a varistor can survive for the three 

sizes and three tolerance values of varistors and for the 

three type of tests: VDE 0160, C62.41 "high exposure," 

and C62.41 "medium exposure." The results with C62.41 

are in good agreement with anecdotal (unpublished) field 

experience, that is, 14-mm varistors installed at the service 

entrance are often in jeopardy, 20-mm varistors have a 

better chance, and 32-mm varistors are generally successful. Failure rates are not reported formally in the literature, but anecdotal information does circulate. The 

response of industry to the Ziirich paper alerting the 

community to the risk of premature aging caused by 

repeated swells [a] was that 20- and 32-mm varistors do 

not suffer from an unacceptable or alarming failure rate. 

The predicted survival rates of Table I appear to be 

consistent with actual field experience, thus validating the 

stress levels proposed by IEEE C62.41. In contrast, for 

the VDE 0160 stress, the predicted survival rate is so low 

that a conclusion appears inescapable: The VDE 0160 

stress involves an exceptionally high energy level, making 

the application of the test questionable if interpreted as a 

general requirement. The authors do not question the 

scenario leading- to this stress level but do cluestion the - - 

high-energy test standards. -.. - ~ a 
1236 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS. VOL. 28, NO. 6, NOVEMBER / DECEMBER 1992 

TABLE I 

PREDICTED NUMBER OF HIGH-ENERGY SURGES THAT A VARISTOR 

CAN SURVIVE ASA FUNCTION OF SIZE AND CLAMPING 

VOLTAGE TOLERANCE 

Varistor 

Clamping 

Size Voltage VDE 0160 C62.41 C62.41 

mm Tolerance Class 2 High Medium 

none 

none 

none 

none 

none 

1 

none 

1 

5 

none 

1 

8 

1 

3 

20 

8 

80 

800 

80 

3000 

> loh 

500 

8000 

"indefinite" 

20 000 

200 000 

"indefinite" 

IEC proposal to require an across-the-board test at that 

level for all equipment. 

The dramatic effect of the tolerance value on survival 

rate is also apparent. Greater reliability can be achieved if 

users would accept-better yet, request-a slightly higher 

clamping voltage than the lowest clamping voltage offered 

by the manufacturers of varistors and by the manufacturers of packaged suppressors. 

The normal practice in surge testing of low-voltage 

equipment, as described in the IEEE Guide on Surge 

Testing 171, is to specify an open-circuit voltage and a 

short-circuit current to be delivered by the surge generator. With these two parameters specified, the surge generator is considered to be defined for any test involving a 

specimen of high impedance (typically insulation) or low 

impedance (typically a surge diverter). For the unidirectional surges of 10/1000 and 100/1300 ps, a simple 

four-component model circuit can produce these waveforms. An actual surge generator, of course, requires 

careful attention to avoid problems of parasitic 

impedances, but the simple circuit model of Fig. 1 can 

deliver the required waveforms, as shown in Fig. 2 for the 

case of the nominal C62.41 10/1000-ps waveform. 

In the specification of that waveform, the tolerances 

allowed by C62.41 recognize the fact that the open-circuit 

voltage will inescapably have a longer duration but shorter 

rise time than the short-circuit current. Because the 

high-energy aspect of this test makes the current waveform the most significant parameter, the values of the 

components in the model were selected to most closely 

approximate the nominal 10/1000 ws for the short-circuit 

current while allowing the open-circuit voltage to go to 

the longest duration permitted by the tolerances. For the 

VDE 0160 model, the values of the components were 

selected to comply with the 6000-pF requirement while 

producing the specified open-circuit voltage. 

In predicting varistor failure rates, the model can take 

into consideration the possible combinations of manufacturing tolerances on the varistors and the uncertainties of 

the test (which is something that is more difficult to do by 

Fig. 1. Four-component circuit for 10/1000- and 100/1300-ps surge 

modeling. 

0-0 

0 1 2 

TIME (ms) 

Fig. 2. Open-circuit voltage V and short-circuit current I produced by 

circuit model with parameters set for the C62.41 10/1000-ps waveform. 

tests on random samples). In the simple computations 

reported here, three cases have been computed with the 

varistor at the midpoint and the two extremes of its 

manufacturing tolerance. The surge generator parameters 

were set to produce the nominal current waveform in 

order to make a midrange rather than a worst-case prediction. The conclusions on survival rates and validation 

of the proposed tests presented above would not be dramatically affected if the surge generator parameter tolerances were included in the computation. 

The component values of the circuit shown in Fig. 1 

may be selected to generate the desired waveforms of the 

various standards. The selection method is described below. In order to determine the response of the circuit with 

the nonlinear varistor, numerical techniques are used as 

shown in the second step below. 

In the circuit of Fig. 1, the capacitance C is charged to 

an initial voltage q.. The surge generator has a series 

resistance R, and a parallel resistance R,. A small inductance L is tuned to provide the specified rise time. This 

simple LRC circuit is described by a characteristic equation 

where R is defined below. The two decay constants are 

The response of the circuit is a "double exponential" 

waveform [9]. 

Using the allowed tolerances of C62.41 for the model, 

the waveshape of the short-circuit current (in which case 

R is R,) was set at 10/1000 ps. For the open-circuit 

voltage (in which case R is R, + R,), the maximum 
FENIMORE AND MARTZLOFF: VALIDATING SURGE TEST STANDARDS BY FIELD EXPERIENCE 

duration allowed by C62.41 is 2000 ps. The decay times, 

which are expressed as full width at half maximum 

(FWHM) for these two waveforms are designated as t, 

and t,, respectively. The effective source impedance is 

defined in C62.41 as the ratio of the peak open-circuit 

voltage Vp to the peak short-circuit current I,. Its value 

Z = %/Ip has the dimension of an impedance. 

Because the time constants are widely separated, the 

determination of the circuit component values from the 

values of t,, t,, Z, and Vp can be simplified to produce 

approximate values. The characteristic decay values A+ 

and A- are given by 

Applying a logarithm yields 

R tR - R . tR 

L E -- 

log .9 - log .l log 9 ' 

The expressions (1)-(5) uniquely define the characteristics 

of the circuit for given values of the time constants, the 

source impedance, and the peak open-circuit voltage. 

With the parameters of the model test circuit thus 

defined, the solution of the response of the current and 

energy in the varistor is obtained numerically using the 

ordinary differential equation package PLOD [lo]. The 

varistor is presumed to contain an internal series resistance R, and have the I-V relationship 

In particular, for long times t, the short circuit current I,, 

and open-circuit voltage Voc, are given by The first-order system of equations to be solved is given 

by the definition of the capacitor current I and by Ohm's 

law 

At half maximum, one has 

and 

t, = log2. (R, + R,) . C. 

The varistor current I, and I are related by (6) and by 

With a small value of the inductance 

and By exploiting this relationship, a direct numerical solution 

for the varistor current is possible. In addition, the energy 

in the varistor Em is found by integrating 

These relations lead to the four equations: 

The initial charge is given by C . T/,, and the initial current 

and energy in the varistor are zero. The computations 

were performed for the two C62.41 exposure levels and 

for the maximum VDE 0160 stress, as described below. 

MODELING RESULTS 

C62.41-10/1000- ps Stresses 

To evaluate the effects of the test on varistors, a simple 

model of an equivalent circuit of the varistor is connected 

to the terminals of the model generator. The charging 

voltage of the generator is, of course, left unchanged. For 

the range of frequencies involved in these waveforms, the 

only two significant elements of the varistor equivalent 

circuit (Fig. 3) are the pure varistor R,, (I = kV") and 

the series resistance R,,, R, in the model. The parallel 

resistance R,,,, capacitance C, and the series inductance 

The inductance is determined by considering the 10-90% 

rise time tR. The widely separated time constants allow 

the fast component of the current to be estimated by 

at short times t. 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 28, NO. 6, NOVEMBER / DECEMBER 1992 

(LEAD 

INDUCTANCE1 

Fig. 3. Equivalent circuit of a varistor (Source: [14]). 

L of the complete equivalent circuit can be neglected. 

Three diameters of 130-V rated varistors were considered, 

each with its characteristic clamping at - 10,0, and + 10% 

of the nominal value published by one manufacturer. 

Fig. 4 shows the type of plots obtained from the model 

where the current through the varistor and the cumulative 

energy deposited in the varistor are computed as a function of time. Showing the complete set of results for all 

combinations would require excessive space; a summary 

of the results is presented in the Appendix. In the typical 

example of Fig. 5, three curves show the cumulative 

energy for a 14-mm varistor with nominal rating of 130 V 

rms and three tolerance values - 10,0, and + 10% clamping voltage when exposed to the C62.41 "high exposure" 

stress level. 

VDE 0160 loo/ 1300-ps Stress 

Fig. 6 shows the parameters of the 100/1300-ps surge, 

Class 2 described in the most recent amendment to VDE 

0160 [I11 and in the IEC proposal [3] . The voltage level is 

specified as 2.3 times the peak of the ac power system 

voltage. (The amendment also cites a Class 1 category 

with a level of only 2.0 times the peak of the ac power 

system voltage and a shorter duration). Accepting for the 

moment the premises that led to the specification of this 

test, the authors applied the same circuit model used for 

the IEEE waveforms to produce the specified VDE waveform with an energy storage capacitor having the value 

specified in the latest amendment to VDE 0160. (Earlier 

versions of the VDE 0160 standard suggested a 25 000-pF 

capacitor. In the amendment, this value has been scaled 

down to a range of 700 to 6000 pF, perhaps implying that 

the issue is still unsettled, and thus, the IEC proposal is 

still open to feedback from users.) 

In this case, because the VDE places emphasis on 

maintaining the voltage waveform, the model parameters 

were set to obtain an open-circuit voltage close to the 

100/1300-ps values, with the exception of the resulting 

short-circuit current, for which VDE 0160 does not specify a value. Fig. 7 shows the open-circuit voltage and 

short-circuit current computed by the model. 

The computations were performed for the 250-V rms 

rating because the VDE 0160 does not provide specifications for system voltages of less than 220 V rms. Details of 

the results are presented in the Appendix together with 

the corresponding results from the C62.41 stress levels. 

0 1 2 

TlME (rns) 

Fig. 4. Energy deposition E and current I in a 20-mm varistor with 

nominal clamping characteristic (0% tolerance) during the "High Exposure" 10/1000-fis C62.41 surge. 

50 III~IIIII 

- 10% - 

Rating of l4mm varistor _ 

+lo% - 

0 1 2 

TlME (ms) 

Fig. 5. Energy deposition in 14-mm varistors at - 10, 0, and + 10% 

values of clamping characteristics during a "high exposure" 10/1000-ps 

C62.41 surge. 

Fig. 6. Voltage waveform of the 100/1300-ps surge specified by VDE 

0160 and proposed by IEC (Source: [5]). 

OLO 

TIME (ms) 

Fig. 7. Open-circuit voltage V and short-circuit current I produced by 

model with parameters set to approximate the VDE 0160 voltage waveform. 
FENIMORE AND MARTZLOFF: VALIDATING SURGE TEST STANDARDS BY FIELD EXPERIENCE 

Typical manufacturer specifications [I21 include a joules 

rating for maximum single pulses; however, industry standards (Section 6, IEEE Standard on Varistor Test Specijications [131) raise some questions on the application of such 

a simple criterion. 

The cumulative energy levels for the three varistor sizes 

(each at three tolerance levels) were computed with the 

model for the C62.41 and VDE 0160 stress levels. The 

results are shown in Table 11, together with the typical, 

single-pulse joule rating published for these sizes. By 

using this somewhat oversimplified joule criterion (more 

than 10% change in nominal voltage may occur if joule 

rating is exceeded), it would appear that only the 14- and 

20-mm varistor, for the low values of tolerance, might be 

in jeopardy. 

Using the criterion of "pulse rating" proposed by manufacturers [Ill, [14], where the current peak and duration 

are taken into consideration leads to more detailed and 

reliable conclusions, which also agree with field experience (see Martzloff [15]). 

Therefore, the current peak and its duration (FWHM) 

were also computed for the nine combinations of varistor 

parameters and compared with the "pulse rating" corresponding to the duration and peak in each case. The 

detailed results, which are the basis for the summary of 

Table I, are presented in tabular fashion in the Appendix, 

together with a discussion of the finer points of the 

analysis. 

1) Predictions of the impact of the 100/1300-ps surge 

test proposed by the IEC and based on the VDE 0160 

standard show that the millions of varistors in service 

should experience a greater failure rate than that indicated by available information on actual field failures. 

This inconsistency raises serious questions on the proposed requirement of such a severe test to a wide range of 

equipment. 

Furthermore, the lingering ambiguity in the VDE 0160 

standard (and, consequently, in the IEC proposal) on 

whether to set constant open-circuit voltage or to adjust 

the voltage while the specimen is connected needs to be 

clarified. A constant, specified open-circuit voltage combined with a well-defined source impedance is the generally accepted practice in surge testing. 

2) The energy levels and currents resulting from application of a waveform described in the revised IEEE 

C62.41, on the other hand, range from benign for typical 

large varistors to severe for small varistors. Thus, this set 

of stress levels appears to be more consistent with field 

experience, at least as inferred from available anecdotal 

information. 

3) Although the authors do not question the validity of 

the fuse-blowing scenario, which is the basis for the VDE 

0160 and proposed IEC test, they recommend a critical 

TABLE I1 

SINGLE-PULSE RESULTS AND RATINGS (IN JOULES) 

O/C varistor varistor 

14 -10 1 212 43 

Varistor 

Size Tolermm ance 

NOTES: 

1. Five numbers are printed In bold face in the results columns for two values of tolerances in 

the 14-mm and 20-mm varistors. These vdues exceed the rating of the varistor, and thus would 

indicate a high likelihood oi failure at that stress level 

2 The varistor model postulates the same I = kVm relation for the three ratings, with a series re^ 

nstance that decreases as the hameter oi the varistor increases. The lower serles resistance invites 

a greater current diversion into the varistor in the upturn region oi the I V characteristic, where its 

effect is more noticeable, especially for the VDE 0160 and the lower tolerance case for the varistors 

review of the statistics of the occurrence of fuse blowing, 

the use of varistors with low clamping voltage, and the 

distribution of actual clamping voltage within manufacturing tolerances. They also urge all users to share information on the observed failure rates and thus attain a 

broader perspective on these issues. 

VDE 0160 

Class 2 

250V 

This Appendix provides a summary of the 54 separate 

computations made to determine the current in the varistor resulting from the three high-energy tests discussed in 

the paper. Three varistor sizes were considered (14, 20, 

and 32 mm), and their "pulse rating" obtained from [12] 

and [14]. The 32-mm size has been dropped from the 

current product line of [I21 and might appear obsolete. 

However, it was selected because it has been applied in 

the past [14], and thus, more field experience is available 

for that size than for the 40-mm size, which is the present 

offering. 

The computed results are presented in Table A1 for the 

VDE 0160 Class 2 and the C62.41 "high exposure" and 

"medium exposure7' stresses. In each major section of the 

table, the computed current peak and FWHM are tabulated. Next to these computations, the corresponding current peaks are shown from the "pulse ratings" in [12] or 

[14] for the computed duration and for 1, 10, and 100 

applications of that peak of current pulse. 

The usual description of a unidirectional surge is based 

on the FWHM, and therefore, the computations of the 

current in the varistor were aimed at characterizing this 

description of the current waveform. However, the "pulse 

rating" curves in both [I21 and [I41 are based on an 

"impulse duration" defined as the time from virtual origin 

of the wave and the virtual time to half value. In the case 

of the C62.41, with a front time of 10 ,us and a FWHM of 

1000 ps, the difference between the FWHM and the 

"impulse duration" is negligible. In the case of the IEC 

100/1300-ps waveform, the difference is more significant, 

C62.41 

High 

130 V 

C62.41 

Medlurn 1 130 V 
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 28, NO. 6, NOVEMBER / DECEMBER 1992 

VARISTOR 

DIA Tolermm ance 

TABLE A1 

MODELING RESULTS VERSUS DEVICE RATINGS-CURRENT AND DURATION 

VDE 0160 = 10011300 ~s IEEE C62.41 10/1000 ps "High" IEEE C62.41 10/1000 ps 'Medium" 

2.3 x 220 x 1.4; 6000 YF; 2.3 x 120 x 1.4; 0.25 n; 2.0~ 120 x 1.4; 1.0~; 

250-V varistor 130-V varistor 130-V varistor 

of Pulses in Columns of Pulses in Columns 

'Adjustment of approximately half of the rise time made to account for the difference between the computed FWHM and the "virtual duration" used in manufacturers 

specifications. For the short rise time of the C62.41, the difference IS negligible. 

"When allowable peak current for the wrresponding duration and number of pulses exceeds the rated peak current at that duration, the varistor IS deemed in jeopardy; this 

situation is shown by shading the wrresponding area in the rating columns. The unshaded areas represent "survival" of the varistor through the high-energy stress. 

IMPULSE DURATION -In 

Fig. AI. Typical published family of "pulse rating" curves showing 

amplitude, duration, and number of allowable pulses. 

and therefore, the comparisons of Table A1 include a 

40-ps adjustment in the duration (about half of the rise 

time). 

The peak values of the current shown in the table that 

exceed the "pulse rating" have been identified by shading 

the area in the columns. At a glance, it becomes apparent 

that the survival rate to a VDE 0160 exposure can be 

expected to be extremely low; it will be moderate for the 

C62.41 "high exposure7' and will be at its maximum for 

the C62.41 "medium exposure" stresses. 

ACKNOWLEDGMENT 

The authors acknowledge the contribution of T. Key in 

presenting this paper at the 1990 Annual Meeting of the 

Industry Applications Society when, on short notice, they 

were unable to travel to Seattle. 

REFERENCES 

[I] ANSI/IEEE C62.41-1980, Guide on Surge Voltages in Low-Voltage 

ac Power Circuits. 

[2] ANSI/IEEE C62.41-1991, Recommended Practice on Surge Voltages in Low-Voltage ac Power Circuits. 

IEC Document 77~(~entral Office)4, Oueruiew on Electromagnetic 

Compatibility-Immunity Tests. 

W. Meissen, "uberspannungen in Niederspannungsnetzen," Elektrotechnische Zeitschrift, vol. 104, 1983. 

VDE 0160 Ausriistung von Starkstromanlagen mit elektronischen 

Betriebsmitteln, May 1988. 

F. A. Fisher and F. D. Martzloff, "Transient control levels: A 

proposal for insulation coordination in low-voltage systems," IEEE 

Trans. Power App. Syst., vol. PAS-95, 1976. 

ANSI/IEEE C62.45-1987, Guide on Surge Testing for Equipment 

Connected to Low-Voltage ac Power Circuits. 

F. D. Martzloff and T. E. Leedy, "Selecting varistor clamping 

voltage: Lower is not better!," in Proc. 1989 Zurich EMC Symp. 

A. Greenwood, Electrical Transients in Power Systems. New York: 

Wiley-Interscience, 1971. 

E. Argon, I. L. Chang, G. Gunaratra, D. K. Kahaner, and M. A. 

Reed, "Mathematical software: PLOD," IEEE Micro, vol. 8, no. 4 

pp. 56-61, Aug: 1988. 

VDE 0160A1 Anderung 1 zu DIN VDE 0160/05.88, Apr. 1989. 

Transient Voltage Suppression Decices, Harris Corp, 1990. 

ANSI/IEEE C62.33-1982, IEEE Stand. Test Spec. Varistor SurgeProtectiue Decices. 

Transient Suppression Manual, 5th Ed., General Electric Company, 

1986. 

F. D. Martzloff, "Varistor versus environment: Winning the rematch," IEEE Trans. Power Deli~sety, vol. PWRD-1, no. 2, Apr. 

1986. 

Charles Fenimore (M'88) received the B.S. degree from Union College, Schenectady, NY, in 

1966 and the Ph.D. degree from the University 

of California, Berkeley, in 1978. 

He is an applied mathematician in the Electricity Division of the National Institute of Standards and Technology, Gaithersburg, MD. His 

interests include modeling and computing the 

nonlinear characteristics of high-energy/highvoltage electrical systems. He is the author of 

numerous papers on the breakdown of dielectric 

liquids, high-voltage measurements, and the solution of partial differential equations. 

Dr. Fenimore is a member of the Dielectrics and Electrical Insulation 

Society and the Circuits and Systems Society of the IEEE. 
FENIMORE AND MARTZLOFF: VALIDATING SURGE TEST STANDARDS BY FIELD EXPERIENCE 1241 

Fraqois D. Martzloff (F'83) was born in France, bushings development, power electronics, transient measurements, surge 

where he completed his undergraduate training, protection of electronics, applications of varistors, and electromagnetic 

and came to the United States in the 1950's to interference protection. 

continue his graduate studies. He is contributing to the work of several IEEE committees, in 

In 1985, he joined the staff of the National particular, the revision of ANSI/IEEE C62.41, Guide on Surge Voltages, 

Bureau of Standards (NBS, now NIST) to ex- into the Recommended Practice as discussed in the paper, and an 

pand NBS activities in the field of conducted update of C62.45, Guide on Surge Testing. He is serving as vicechair of 

electromagnetic interference with more recent the IAS IEEE Standards Coordinating Committee on Power Quality. In 

emphasis on power quality and surge protection the IEC Technical Committee 77 on Electromagnetic Compatibility, he 

issues. Prior to joining NBS, his long career at is serving as chair of SC77B (industrial systems) and secretary of WG2B 

General Electric included high-voltage fuses and (installation and mitigation). 
APPLYING

REALITY

CHECKS

TO

STANDARDS

ON THE SURGE

ENVIRONMENT
Applying Reality Checks

to Standards on the Surge Environment

Arshad Mansoor Power Electronics Applications Center Knoxville TN 37932 USA Amansoor@epri-peac

François Martzloff National Institute of Standards and Technology Gaithersburg MD 20899 USA f.martzloff@ieee.org

Doni Nastasi Power Electronics Applications Center Knoxville TN 37932 USA DNastasi@epri-peac.com

Reprinted with permission from Proceedings, 23rd International Conference on Lightning Protection, Florence, 1996

Significance:

Part 2 Development of Standards – Reality checks

Three examples of reality checks are given that shed some light on issues raised during standards development

! The apparent reduction in surge voltage activity is explained by the proliferation of surge mitigating devices.

! A proposed high-energy surge would cause failure of ubiquitous SPDs, but they do not fail in the field.

! Allegedly frequent high-level surges would cause frequent failure of light bulbs, but they do not in the field
A DDT VllTP 

RP A T TTV OOWC nA A u A LI A u x\lrnbll I LLIULI\O 

TO STANDARDS ON THE SURGE ENVIRONMENT 

Arshad Mansoor and Doni Nastasi 

Power Eiectronics ~ppiications Center 

Knoxville TN USA 

Abstract - The paper identifies several realities of surge 

environment and equipment survival that are sometimes ignored 

in surge-protection practices. It questions the quest for what 

could be overly conservative requirements for surge immunity or 

surge mitigation by presenting "reality checks" based an field 

experience or laboratory data. A first check focuses on the fact 

that some recent field recordings of surges may be misleadingly 

low in today's surge environment. Additional checks, aimed at 

moderating the overly conservative requirements, include the case 

history of a proposed high-stress 10011300 ps surge test, data on 

failure levels of clock motors and light bulbs that can serve as 

benchmarks for severity levels, and measurements, validated by 

parametric modeling, showing that large currents cannot 

propagate into long cables without causing a flashover of the 

wiring devices at the beginning of the cable, effectively limiting 

the energy-delivery capability of a surge at the end of the cable. 

1. Introduction 

Some proposals for standards on equipment surge 

immunity or on performance of surge-protective devices 

are driven by the commendable quest for conservative 

ratings. However, when this quest produces compounded 

safety factors, the result may not be cost-effective. 

The purpose of our paper is to present facts and 

rationalizations in support of the development of realistic 

standards on the surge environment in low-voltage ac 

power systems. In addition to contributors to the ICLP 

Conferences, two major players in the development of ., ,-+ :-,, ,+,,A..rA- ,., +I., l:A.+,:-r ,..-,, ,,..:,,----r ;ut~11rauu1ualuulucuuo u11 ulc. ug11u~u1g DUI~G GIIYUUIUIGIIL 

are IEC Technical Committee 77 on Electromagnetic 

Compatibility and Technical Committee 81 on Lightning 

Protection. Other contributors addressing switching surges 

a$ well as lightning surges include IEC Subcommittee 28A 

on Insulation Coordination, Subcommittee 37A on SurgePr~!pc!ivp Devices, and FEE Com-mjtke on &geProtective Devices. The standards developed by these 

bodies reflect the collective experience of individual 

contributors as well as the "corporate memory" of these 

groups. In this paper, four items are described that may 

serve as foundations, or reality checks, to be added to this 

corporate memory. The first check can be seen as a 

reminder note to resist the mistake of dismissing surge 

threats because contemporary recordings show benign 

occurrences of surge voitages. Tine three other checks can 

serve as a tempering note, to avoid economically unjustified over-specification of surge protection. 

-- ~ 

* Electriciry Division, Electronics and Elenricd Engineering Laboratory, 

Technology Administration, U.S. Department of Commerce. 

Fran~ois D. Martzloff 

Nationai Institute of Standards and Technology * 

Gaithersburg MD USA 

2. The ninety-five percent rule 

Debates and definitions concerning compatibility 

levels, immunity limits, or emissions limits generally 

acknowledge that requirements drawn from these considerations cannot apply simultaneously to 100% of the cases 

and 100% of the time. The term "high probability" appears 

in many definitions, rather than a hard number. This 

avoidance of hard numbers is the result of the difficulty to 

secure a consensus from individuals who represent 

different communities of end-users with different perceptions of what is an acceptable level of probability. 

The military and aerospace communities often speak 

of zero tolerance or zero defects where economics is 

downplayed. In contrast the commercial, mass-production 

communities speak of realistic requirements where costeffectiveness is a very prominent factor. 

An intermediate position can be found in industry 

where a rational balance is sought between the costs of 

exhaustive mitigation and the consequential costs of limited 

mitigation. 

Our reality checks are offered as a guide to striking 

such a balance, recognizing the diversity of circumstances 

and expectations from a diversity of end-users. 

3. Shrinking surges vs. expanding mitigation 

A reality check can explain an apparently puzzling 

development which has been observed since the beginning 

of studies concerning the characterization of the surge 

environment: starting in the sixties, results published by 

researchers monitoring the occurrence of surges have 

reported a gradual decrease in the relative severity of 

surges (Martzloff & Hahn, 1970 [I]; Allen & Segall, 1974 

:2]; Sd! & Nethercott, 1375 [3]; Gddsiein aid Speranza, 

1982 [4]; Goedbloed, 1987 [5]; Dorr, 1995 [6]). 

This pmz!ing trend can be exp!&ned by reference 

to the reality: the spectacular expansion of surge-protective 

devices (SPDs) in low-voltage ac power circuits now 

makes it almost impossible to make measurements at a site 

where there is not some hidden SPD. Monitors will now 

report the residual voltage of these SPDs, not the me surge 

activity in the environment (Aspnes et al. and discussion. 

1985 [7]). So, it is not that surges have shrunk, but rather 

that mitigation has greatly expanded, masking the persistent 

occurrence of transients which remain a potentiai heat. 

This threat has shifted from observed surge voltages -- 

now mitigated -- to the possibility of excessive surge 

currents into candidate mitigation means with low limiting 

voltage which would be particularly attractive to the surges. 
Future surveys of the occurrence of surges in lowvoltage ac power systems should shift from the recording 

of surge voltages to the recording of surge currents [8]. 

Unless this shift is implemented, the false sense of security 

will be further promoted that potentially damaging surges 

are less frequent than once believed. The 1980 title of a 

seminal IEEE document on the occurrence of surges 

(Guide on Surge Voltages in Low-Voltage AC Power 

Circuits [9]) was appropriate at the time, but ihould now be 

amended to reflect the shift. 

4. The saga of the proposed 100/1300 ps surge 

A proposal was made in the early eighties to require 

a high-stress surge test for industrial equipment. A reality 

check would have revealed the contradiction between the 

failure rates that would result from such a test and the 

observed field failures. 

Starring with observations of the occurrence of long 

duration surges, typically associated wth iuse biowing, the 

proposal was to require an additional test with a surge 

longer than the classic 1,250 - 8/20 ps impulse (Meissen, 

1983 [lo]; VDE 0160, 1989, [ll]; IEC 1000-4-1, 1990 

[12]). The proposed surge would be characterized by a 

10011300 ps waveform, with peaks as high as 1,3 per unit, 

added to the ac sine wave peak. In the original VDE 01 60 

implementation, the test circuit involved the discharge of a 

24000 pF energy-storage capacitor. An amendment to the 

VDE 0160 standard [13] scaled the capacitance value 

down to 6000 pF, still a large stress for the equipment 

under test. 

For instance, such surges would impose large 

stresses to the millions (or perhaps billions) of small varistors now installed in low-voltage ac power systems. One 

would observe a conspicuous failure rate in the field as 

typical 20-mm diameter varistors cannot survive such 

surges, but one does not observe this &Glue rate. An 

objection to this surge specification was first proposed on 

the basis of computer simulation of the event (Fenimore & 

Martzloff, 1990 [14]) because no generator was commercially available to produce that surge. Table 1 shows 

an excerpt from [14] where the current that would be 

caused by the proposed 10011300 ps surge in a 20-mm 

varistor rated 250 V rms was computed and compared to 

the published varistor pulse rating for current-handling 

capability. Only a varistor at the high end of its -+lo% 

tolerance band could survive one application of that surge 

(th.e coxespcnding ~mpkd ccmn! pea! is nn!y 269 A, 

compared to a rating of 300 A). In all the other cases, the 

varistor rating is exceeded and likelihood of failure 

increases for lower tolerances and for more than one pulse. 

Later on, when a prototype generator capable of 

delivering this surge was developed by a surge generator ----.C..-+..--- -m ..At-armA h.r n a nf thm ntathfirr rrnd hv IllalulabLUGl, &am wrulraaru vj vnr v~ YAW UYY~V-Y U..V YJ 

W.Meissen (the originator of the 10011300 ps surge 

proposal), did confirm the fact that 20-mm varistors are 

destroyed by that surge (Martzloff, 1991 [15]). We now 

understand that the IEC proposals and the VDE 0160 

requirements for the 10011300 ps surge have been pulled 

back or are being reconsidered. 

Table 1 

Computed peak currents for a postulated 100/1300 ps surge 

and varistor rated peak current 

5. Failure levels of clock motors 

and incandescent light bulbs 

Tolerance 

on varistor 

nominal 

voltage I rating 

(%) 

- 10 

0 

+10 

Tw~ simp!e devices have been in service ir? --!!inns 

of households and can be seen as surge threshold detectors: 

motor-driven clocks (before the advent of digital clocks) 

and still today, the ubiquitous incandescent light bulb. We 

have some knowledge on the failure rate of some of these 

devices under surge conditions, from which we can derive 

some inferences on approximate limits on the level of 

frequent surges. As we will see in the following two 

subsections, the data are not precise and are subject to 

undetermined statisticai variations. However, the iarge 

number of observations does give some value to the 

inferences, compared to the limited number of observations 

obtained by more precise recording instruments. 

5. I Failures of motor-driven clocks 

Computed results: 

Current peak and 

duration of surge 

in the varistor 

'FWHM: Full width at half maximum 

In the sixties, a report was published citing a 100: 1 

reduction in the failure rate of clock motors when the 

withstand ievei of the motor was raised from 2 icV to 6 i<ii 

[9], thus providing a verification of the "slope" of surge 

occutrrence rat^: vs. peak repafled hy many reezchers [I@, 

Varistor rating: Allowable 

peak A for number of 

pulses of computed 

FWHM 

peak 

(A) 

658 

454 

269 

Figure 1 shows a plot of the frequency of occurrence of surge voltages versus their peak amplitude as 

reported by many researchers, and the two-point line of the 

100:l relationship between 2 kV and 6 kV occurrences. 

The parallel position of all these lines is remarkable. While 

&e p!cs qc&d by tuhl~ rerPE&err %J.P, g~~~_r&v limited J -------- 

to a few thousand hours of observation, at only a few 

locations, the clock data represent the integration of more 

than 10000 clocks over a period of two to three years. 

pvAvtgM* 

(ps) 

625 

515 

400 

dnn IVV 

pulses 

45 

55 

65 

Because the data reflect only the number of clocks 

returned by dissatisfied customers, the exact number of 

clocks involved in "monitoring" the surge voltages, and the 

duration of the observation are not known. Therefore, 

rigorous statistical analysis would be meaningless. The 

point of our reality check, however, is still that a very large 

number of observation points were involved, for a period 

of several years, something that no survey couid do -- but 

for this particular case history, it was all done before the 

environment began to be tamed by the proliferation of the 

new generation of low-voltage SPDs. 

I V 

pulses 

80 

90 

110 

I 

pulse 

210 

250 

300 
Peak Surge 

8-l - Bull & lethercot (co~posite) U - Martzloff Can - Cannova 

6-5 - fioldsteln 4 Speranza UBB - Yernstrom et *I (upper llmlt) 

60e Poedbloed Clock - 2 kV L 6 kV points only 

Figure 1 - Plots of frequency of occurrence versus peak amplitude 

reported in 1970-1987 surveys by independent researchers. The 

line marked "clock" only shows the slope from 2 kV to 6 kV, not 

the actual frequency of occurrence. Note how the lines are 

essentially parallel. (Plot reproduced from 1161.) 

5.2 Failure of incandescent.light bulbs under surges 

TesG recently perfo-a by fie authors 120-V 

incandescent light bulbs show that few bulbs will survive 

surges in excess of 1200 to 1500 V 1., and that a surge of as 

little as 800 V, when it occurs under the most sensitive 

condition, can trigger the failure . Limited tests performed 

on 240 V bulbs yield similar results. Comparison of tests 

performed by a researcher in Austria and by the authors on 

both 120 V and 240 V bulbs might be available at the time 

of presentation of this paper. 

Since we do not hear reports of endemic failure of 

light bulbs beyond what can be expected from their known 

service life, we must conclude that the reality is that there 

is not a high rate of occurrence of surges at levels in excess 

of 1500 V. 

t The measurements reported in this paper have been made with 

instnunentation for which the combined uncertainty should not "...-""A LC r. +OI^ rx.,",, +ha "r,mrr nt-""nl.,;nn ths ...~"r.r..a - 

crcccu u 6" -v I". u'rcr' rrrc yrvs.c.ee vj uyy.,v.rrg ra- rr-wurrment results to the failure levels of light bulbs exposed to 

environments with characteristics that are at best hown within 

an order of magnitude, this level of uncertainry does nor affect 

the practical conclusions. 

This observation merits a brief summary of our 

findings about the mechanism of bulb failure triggered by 

a surge. We emphasize the triggered aspect because surges 

of relatively low amplitude do not cause directly the failure 

of the bulb, but cause a secondary flashover at the power 

frequency that burns out the filament. Surges nf higher 

amplitudes can deposit enough energy into the filament to 

melt it, but our point for the reality check is that even lowamplitude surges can result in failure of the bulb. 

Figures 2, 3, 4, and 5 present oscillograms of the 

voltage across an energized 120-V bulb and of the current 

flowing in the bulb. Figure 2 shows a narrow window 

commensurate with the duration of the surge delivered by 

a 1,2150 ps - 8120 ps surge generator. We observe the 

chopping of the voltage wave, typical of a gap sparkover, 

and the rise of the surge current after the sparkover. The 

scale of the current trace selected to record the surge 

(hundreds of amperes) does not show the normal current 

(1 A) in the bulb. Observations during this test include 

hearing a pinging noise and seeing a bright flash of light, 

followed by darkness as the filament can then be seen 

broken at its points of attachment to the stems. 

I ... i .. i. i. i.... i. .. i. ~-- i.. - ~p I. l.... 1 

Top trace: Voltage across bulb terminals, 200 Vldiv 

Bottom trace: Current in bulb. 100 Ndiv 

Sweep: 10 psldiv 

Figure 2 - Voltage and current in light bulb during application 

of a 1,2450 ps - 8/20 ps surge, resulting in surge sparkover 

Figure 3 was recorded (for a new bulb) with a 

ionger window to dispiay two fuii cycies of the power 

frequency. At that sweep rate, the surge is no longer 

resolved, and its apparent peak on the trace may be lower 

than the actual peak because not enough data samples are 

collected around the peak. However, the timing of the 

surge, and the events following the surge are what is 

:.-&.-* :.. *l.:- c -.-- n- --.-- ------- lllllL I I LC. LLIC UI~C CVCIIL appc~ asa 

voltage spike and a current spike, followed by return to 

practically normal voltage and no visible large current. 

Then, suddenly, a pulse of power-frequency current 

appears, with a large amplitude -- the source of the 

observed flash. We believe that it is this current that causes 

the burn-out of the filament, not the "trigger" surge. 
.................... .................... 

Top trace: Voltage across bulb terminals, 200 Vldiv 

Bottom trace: Current in bulb, 100 Ndiv 

Sweep: 5 msldiv 

Figure 3 - Voltage and current in light bulb during application of 

a 1,U50 ps - 8/20 ps surge at 30°, resulting in surge sparkover, 

followed by power-frequency flashover 

This belief is supported by the recordings shown in 

Figures 4 and 5. Figure 4 is actually the recording of a 

surge application to the bulb of Figure 3, in a test that 

preceded the fatal surge recorded in Figure 3. The applied 

surge was the same (900 V) for both Figure 3 and Figure 4. 

The randomness of the process of igniting the power arc is 

such that in the case of Figure 4, the power arc was not 

ignited, while in the subsequent surge application on the 

same bulb (Figure 3) and in the same conditions, the power 

arc was ignited, resulting in burn-out of the filament. 

Top trace: Voltage across bulb terminals, 200 Vldiv 

Bottom trace: Current in bulb, 100 Ndiv 

Sweep: 5 msidiv 

Figure 4 - Voltage and current in light bulb during application 

of a l,2/50 ps - 8/20 ps surge at 30°, with surge sparkover but 

no power-frequency flashover 

As f~rht er evidence, Fi-me 5 shows the process (in 

a new bulb) when the surge was applied at zero degrees, 

the time at which there is the least power-frequency voltage 

to ignite a power arc. in that test sequence, the iirsi appiied 

surge had an amplitude of 1000 V. Sparkover under the 

impulse did occur (the pinging noise was heard), but the 

power-frequency arc was not ignited and the bulb survived. 

........................+................. ..... 

I i f'f ..L. ..._,.... I...,,....,....I....,.,..,..,.!,...,..., 

Top trace: Voltage across bulb terminals, 200 V/div 

Bottom trace: Current in bulb, 100 Ndiv 

Sweep: 5 msldiv 

Figure 5 - Voltage and current in light bulb during application 

of a 1,2150 ps - 8/20 ps surge at 0°, resulting in surge sparkover 

and delayed power-frequency flashover 

The applied surge was then raised in 100 V steps, 

still with pinging heard but no fatal power-frequency arc. 

Figure 5 was recorded when the applied surge voltage was 

increased to 1300 V. At that level of energy deposited by 

the surge, enough plasma was generated in the path of the 

surge current to eventually ignite the power-frequency arc, 

but it had to wait until the power,frequency voltage had 

rezched i;s pe&. 

To conclude this summary, Table 2 shows the 

relationship between the timing of the surge with respect to 

the sine wave and the amplitude of the surge sufficient to 

trigger ignition of the power-frequency arc. When the 

surge is applied at 90 degrees (the peak of the sine wave, 

making immediate ignition of the power arc easiest), a 

surge of 800 V is sufficient to trigger the power arc. At 

zero degrees, the surge must be raised to 1500 V to 

produce sufficient plasma to result in a subsequent powerfrequency arc. 

The phenomena are of course subject to the 

statistical variations of sparkover. The values shown in 

Table 2 are the averages of tests performed on a total of 20 

bulbs of the same manufacturer, rated 100 W, replicating 

the test at several timing angles. This paper is not the 

medium for reporting in detail our series of experiments 

with other manufacturers and other watt ratings which 

produced similar results; the point is, however, that our 

inferences are not hased on just the 20 bulbs of Table 2. 

Table 2 

Relationship between timing angle of the surge and amplitude 

necessary to produce a fatal power-frequency flashover 

Angle(") 

Peak (V) 

' One specimen produced 'pinging" starting at 1000 V, but no 

power arc. Eventually. the bulb failed at 1600 V without power arc. 

as a direct result of the energy dumped into the filament by the 

surge alone. 

30 

900 

0 

1500 

- - 

15 

1200 

45 

800 

165' 

1100 

135 

850 

90 

800 

150 

850 
6. Limits to pushing surges into branch circuits 

SO~ ppds hwe been =..& tc require SPDs 

intended for installation on indoor circuits to withstand 

surges with relatively high peaks and short rise time. Such 

a requirement would mean that a substantial voltage drop 

would be developed (L. dildt) along the wiring. Added to 

the !ifif r?g vn!tage sf the SPD at the end sf th,e !in=, tk 

voltage necessary to drive such a surge at the origin of the 

line would very likely cause flashover of wiring devices at 

the origin of the line. 

This flashover, occurring during the rising part of 

the surge, would effectively shut off further propagation of 

the surge toward the SPD (except for the energy stored in 

the line during the current rise, which is easily dissipated 

by the SPD). Thus, the requirement of a large surge 

capability for SPDs installed with even a modicum of line 

length would be unrealistic. In support of this statement, 

we present here a summary of measurements on actual 

wiring ZK! EMTP [!7] cnmpter simu!~tim ef a r~qge cf 

parameters. 

Figure 6 shows the experimental circuit with a 

varistor connected at the downstream end of a "branch 

circuit" consisting of two copper conductors of 2-mm2 

cross-section (#I2 AWG), typical of residential wiring. 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 can be set to produce sparkover at some given voltage 

during the test as well as in the model. 

To determine the response of the circuit without the 

clearance limitation and verify that the model produces the 

same result, the gap setting was adjusted for this particular ,--A -- rL-. -- ---- 1.- 

~csr su umi no spaover occurred at the upsueam voirage 

developed for the current delivered by the generator. Once 

the model was validated, parametric variations could be 

performed for any combination of circuit length, applied 

surge, and clearance (gap) sparkover. 

t 9 meters i 

- - /xlm 

uuv A ""U 

Surge 

generator MOV - 

/A 

- t 

I Dig 1 

signal 

analyzer printer 

Figure 6 - Test setup for driving surges into a varistor 

installed at the end of a 9 m branch circuit. 

I. I 

Top trace: Gap voltage, 500 Vldiv 

Center trace: Total current, 500 Ndiv 

Bottom trace: MOV current, 500 Ndiv 

Sweep: 10 psldiv 

Figure 7 - Voltage and currents 

measured in the circuit of Figure 6 

Figure 8 - Plot from model of voltage across the gap 

(compare with voltage trace of Figure 7) and impinging 

current (compare with the two current traces of Figure 7) 

for the circuit of Figure 6 

Inspection of Figures 7 and 8 clearly shows the 

agreement between real-world measurements and model. 

The voltage traces are quite comparable. In Figure 7, the 

two current traces are identical since no current is diverted 

in the arrester. In Figure 7, the current trace is the one 

pstu!ateC! in the mnde!ing. %is c~nespmdence a!!ows 

us to make parametric variations in the model with 

postulated impinging surges of higher values, such as those 

being proposed for consideration in some standards-writing 

groups. Another advantage of the model is that it allows 

postulating a current source -- the consensus choice among 

lightning researchers - rather than using a surge generator 

which is only a charged capacitor with a wave-shaping 

network. Consequently, the surge generator interacts with 

the circuit into which the surge is injected, changing the 

waveform and losing the postulated constant waveform. 

This changing of the waveform makes it more difficu!t to 

perform parametric tests, compared to the ease of 

modeling. Table 3 below shows the results of such 

computations for the current waveform of Figures 7 and 8. 
Computed 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 

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 testlmodel validation. 

Typical wiring devices used in 120-V installations 

have minimum flashover points in the range of 5 to 8 kV 

(this as a result of prevailing clearance requirements rather 

than insulation withstand levels as recommended by IEC 

Publication 664 [18]). Comparing thls flashover level with 

the voltage values of Table 3 shows that, indeed, there is a 

strong possibility that within the range of current levels and 

branch circuit lengths of the table, a flashover would occur 

to throttle further propagation of the surge toward the -A AL- --A I. ---- L A --.. :+ 1. ,.+I.,.- ..,a- A- VkKlbLUI ill UIE; GllU Ul UlG Ul~lLII L1lL.UlL. A11 UUIGl WUIUJ, 

postulating ever-larger surge currents into ever-longer 

branch circuits leads to a contradiction with the physical 

reality of the eventual occurrence of self-limiting flashover. 

This paradox has been discussed in greater details in a 

paper dedicated to that very subject [19]. 

Peak 2w i 3k~ / 5k~ / 7k~ I Length 

7. Conclusions 

10 m 

?n m "" 8,. 

50 m 

1. Reality checks on the surge environment can bring 

a sense of perspective and help developing realistic 

standards for performance and application of surgeprotective devices. This perspective can help avoid 

both over- and under-specification. 

2. The proliferation of surge-protective devices in 

low-voltage ac power systems has made the 

recording of surge vohges practically irrelevant at 

best, and misleading at worst because the recorders 

indicate the let-through voltage of the SPDs, not the 

impinging surge. Surveys of ke surge environment 

should focus on the ability (threat) of a surge to 

deliver a current to those ubiquitous surgeprotective devices. 

23 

C. Q -9- 

9.3 

3. Field failure rates of various devices for which the 

failure level is known can provide some realistic 

and useful information on the rate of occurrence and 

level of surges in the environment. 

4. Modeling the propagation of surges in a circuit, in 

particular computing the resulting voltages and 

comparing them with the known withstand level of 

insulation can provide a restraining influence on 

over-specifiing surge protection requirements. 

3,3 

8,s 

13,7 

Support for the development of this paper was 

provided by the Power Quality research groups of 

Delmarva Power & Light, and Pacific Gas & Electric. 

Support for the testing and modeling was provided by the 

Electric Power Research Institute. 

References 

Martzloff, F.D. and Hahn, G.J., "Surge Voltages in Res~dentlal 

and Industrial Power Circuits," IEEE Transactions PAS-89, No.6, 

JulyIAugust 1970. 

Allen, G.W. and Segall, D., ''Monitoring of Computer Installations 

for Power Line Disturbances," IEEE Winter Power Meeting 

Conference Paper WINPWR C74, 1974. 

Bull and Nethercon 'The Frequency of Occurrence and the 

Magnitude of Short Duration Transients in Low-voltage Supply 

Mdnc," llk.e R& &)?_A E!ec!.rmL Ezgiver. Sept 1964. 

Goldstein, M. and Speranza, P., 'The Quality of U.S. Commercial 

AC Power," IEEE-INTELEC Conference Proceedings, 1992. 

Goedbloed, J.J., 'Transients in Low-Voltage Supply Networks," 

IEEE Transactions EMC-29, No.2, May 1987. 

Dorr, D.S., "Point of Utilization Power Quality Study Results," 

IEEE Transactions IA-31, No.4, July/August 1995. 

Aspnes, J.D., Evans, B.W., and Memtt, R.P., "Rural Alaska 

Electric Power Quality," IEEE Transactions PAS-104, No.3, 

March 1985. 

Martzloff, F.D., "Surge Recordings That Make Sense: Shifting 

Focus from Voltage to Current Measurements," Proceedings. 

EMC '96 Rorna, September 1996. 

IEEE Std 587-180, IEEE Guide on Surge Voltages in AC Power 

Circuits, now updated as ANSyIEEE C62.41- 1991, 

Recommended Practice on Surge Voltages in AC Power Circuits. 

5,2 

!3,9 

22,7 

10. Meissen, W., "iJberspannungen in Niederspannungsneaen 

Elektrotechnishe Zeitschrifi," Vol. 104, 1983. 

1 I. VDE 0160, Ausriistung v6n Starkstromanlagen mit elektronishen 

Betriebsmineln, May 1988. 

12. IEC Publication lOOO+l, Electromagnetic Compatibiliry, Part 4: 

Immunity Tests, Section I: Overview of Immunity tests, 1992. 

13. VDE 0160A1, Anderung 1 zu DIN VDE 016W05.88, 1989. 

14. Fenirnore, C. and Martzloff, F.D., "Validating Surge Test 

Standards by Field Experience: High-Energy Tests and Varistor 

Performance," Conference Record, IEEE-IAS Annual Meeting, 

1990. Aiso in iEEE Transactions 1,428 ~0.6, i992. 

15. Martzloff, F.D., 'Testing Varistors Against the VDE 0160 

Standard," @en Forum on Surge Protection Ap,nlication; 

NISTIR-4654, 199 1. 

16. Martzloff, F.D. and Gruzs, T.M., "Power Quality Site Surveys: 

Facts, Fiction, and Fallacies," IEEE Transactions IA-24, No.6, 

1988. 

17. EPRI, '%Electromagnetic Transient Program (EMTP), Version 2.0; 

Volume 1: Main Program; Volume 2: Auxiliary Routines," EPRI 

Report EL-6421-L, July 1989. 

18. IEC Report 664-1980, Inrulatwn coordination within low-voltage 

-."*^." :..,1..,I:,, ^Irrrr..^". "..,I ^-"^,""" '-"- J)JLCl,&J, YLCLUUll'& CICUIWICC.) UI.U CICC,/UBC UIJLUIlCCJ JYI 

equipment. 

19. Mansoor, A. and Martzloff, F.D., "Driving High Surge Currents 

into Long Cables: More Begets Less," Presented at KEEE-PES 

Summer Meeting, Denver CO, July 19%. to be published in IEEE 

Transactions PWRD at later date. 

72 

!0,4 

31,6 
USING

INCANDESCENT

LAMP 

FAILURE

LEVELS FOR

ASSESSMENT

OF THE SURGE

ENVIRONMENT
Using Incandescent Lamp Failure Levels

for Assessment of the Surge Environment

Hubert Bachl

Cooperative Testing

Institute

Vienna, Austria

François Martzloff

National Institute of Standards

and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

Doni Nastasi

Power Electronics Applications

Center

Knoxville TN 37932 USA

DNastasi@epri-peac.com

Reprinted, with permission, from

Proceedings, 12th International Zürich Symposium on Electromagnetic Compatibility, 1997

Significance:

Part 2 Development of standards – Reality checks

Investigations were conducted in the US as well as in Austria, on 120-v and 240-V incandescent lamps to determine

the levels of surges that can trigger an internal flashover of the hot filament, resulting in filament burnout. 

Repetitive surge application below the threshold do not result in premature failure of the lamp, but above the

threshold, a single application can trigger a fatal flashover. By combining measurement of currents and voltage

during the event with high-speed video recording, the mechanism has been clearly determined.

Depending on the characteristics of the surge (waveform, amplitude, and timing with respect with the powerfrequency sinewave), thresholds of failure range between 800 V and 2000 V. Very few bulbs survive surges above

2200 V. Therefore, the conclusion is inescapable: if such surges were occurring frequently – according to some

SPD advertizing claims – lamps would fail very promptly. We know they do not, ergo the alleged frequency of

occurrence is incorrect.
.
.
.
. - 582 -
.
.
DEVELOPING 

A

CONSUMERORIENTED

GUIDE

ON

SURGE

PROTECTION
Developing a Consumer-Oriented Guide on Surge Protection

Thomas Key, EPRI-PEAC, 942 Corridor Park Blvd, Knoxville TN 37932 TKey@epri-peac.com

François Martzloff, National Institute of Standards and Technology, Gaithersburg MD 20899 f.martzloff@ieee.org

Roger Witt, State Farm Insurance Co., 1 State Farm Plaza, Bloomington IL 61710

Jim May and Stacey Black, Illinois Power Company, 500 South 27th Street, Decatur IL 62525

Reprinted from Proceedings, PQA’97 Conference, 1997

Significance: 

Part 2 Development of Standards – Reality checks

Caught among contradictory stories on the need for surge protection as well as unsupported anecdotes of surgerelated failures, the typical consumer is in a quandary on how to best allocate personal resources to protect the

expensive electronic equipment found in a modem household.

To help provide some answers to this quandary, a team of experts developed and engineering guide on the basics

of surge protection. One of the recommendations addresses the issue of ineffective configuration of utility

connections that are responsible for surge-induced failures. This paper served as a progress report, documenting

the status of the project at the end of 1996. An update on this project was reported under Key et al. 1999, which is

included in Part 6 of this Anthology
1. .-- ^ ^. . ^-. ^. .IN l HUUUL I IUN 

Each year !ightning is estimated to be responsib!e for 250 to 500 ml!!ion do!!ars in propew 

in the United States. This estimate is based on an analysis of insurance claim cempared 

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A iypicai modern residence contains many eiecironic and muiti pori appiiances. in order 

speciiic protection practices, a typicai resiaence was deiinea wiin eiectronic appiiances and 

illustrated in various rooms 8s shown in figure 1. The main systems are home ente

InC!gdlng cab!e TV, home ~fflce Inc!udlng fe!pphQnp, and e!ecfrOnlc klfchen app!rance.. 

cammen vu!nerab!e systems are heme security, intercem, sate!!ite receiver, garage dee

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different experts had different backgrounds and experiences. At the end all agreed on 

principles of how to protect sensitive equipment. However, specific devices and proce

could be applied case by case were more difficult to obtain. Also, there is currently no in

aboiji compaiiGiiiv ieveis and points of -"iiiiierabiiiv for rnaiiji modern 

appliances and systems. Consequently, specific installation methods and precise recomm

on how to protect these electronic systems from surge threats are not yet fully develope

A few appliances seem to have varying susceptibility to surges, The most notable losses, 

to insurance claim records, are telephones and modems, computerized equipment, TV, 

satellite receiver systems. These are generally multi-port appliances-that is, appliances 

to several different systems, such as an entertainment center connected to power, c

telephone or a security system connected to power, sensor, and control. These 

connections, and the likelihood of potential differences between them, are believed to 

factors in surge damage susceptibility of many residential appliances and systems. 

Lightning is not the only threat, but remains a dominant one.. Most of the Continental Uni

experience at least two cloud-to-ground (C-G) flashes per square kilometer per year. Abou

of the area will see three C-G flashes per km2, which is equivalent to about 10 disch

square mile per year. The maximum flash densities are found along the southeastern G

and the Florida peninsula, where the values approach 20/km2 or 50/mi2. Overall about 

C-G flashes strike the United States each year, and lightning is clearly among the nati

severe weather hazards. 

In high lightning areas, perhaps a threshold of three C-G flashes per square km per ye

degree of structural lightning protection is recommended. This threshold may be eco

justified by the growing value, and vulnerability, of residential electronic systems. The

:-L..--.+:-.- A,.-:I.-.C,l.-. n- I-.-.., +rr rrrrnrrm-l;rrC, +C,;- -rn+rr.-.+;n- Dm-;- I;rrh+m;-m -rrr+nrr+;n- -rrr.-

II IIUI III~LIVI I avallault: UI I I IUW LU abbull lp1131 I LI 113 ~IVLC~LIUI I. ua31b II~I ILI III ly ~IVLCLLIUI I pla

well defined in the Lightning Protection Code, NFPA 780. An illustration of the basic el

this protection is provided in figure 2. These protection techniques will not eliminate su

enter the residence either via different wiring systems-including the grounding 

syste~~: cwpled into :vi:ing from nearby flash.es. In additim to the lightning threg!, so

household appliances can act as surge generators, for example a light switch or a furna

Protection against these surges can take several forms. The possibilities are: preventing 

at its origin (impossible for lightning and difficult for surges associated with normal opera

power system), diverting the surge to ground as it impinges on the building, before it e

building (the most effective approach), and finally, clamping by a surge protective device 

the equipment (either by an add-on, plug-in, or built in SPD). Of these three approaches

service entrance and add-on SPDs are options available to the end-user. Service 

protection may be offered by the local utility. If not, a licensed electrical contractor 

installation service. In either case installation of service entrance protection must be d

professional. Various locations for power line or transient-voltage surge suppressors (T

shown in figure 3. 
Legend - 

Air terminal spaced 6 meters (20 ieet) apart aiong ridges and within 0.6 meter (2 iee

ridge ends 

Down conductors (minimum of two) 

Grounding rods; minimum of two, at least 3 meters (1 0 feet) deep Dnnf nrrrinntinnc. en neb nc uragthnr \r~nnc finnnnptarl tn nrnl lnrlinn p\rc+am nuwl ~IUJGCILIUI 13 ~UUI I uo VVWXLI 1-1 vu.11 ~~lr;) VVI II I-VIVU tv ~~VUI IUIV VJVLV~ II 

Air terminals located within 0.6 meter (2 feet) of outside corners of chimney 

Dormers protected with air terminals 

Antenna mast bonded to roof ground conductor 

Gutters and other metal objects bonded to grounding system 

Surae I arrester installed at service Dane1 to ~rotect appliances 

lo. Transient voltage surge suppressors installed at receptacles powering sensitive elec

Figure 2. Typical comprehensive lightning protection of a residence (reprinted with perm

from Underwriters iaboratory inc.j 

\ 

Power Line Suree Protectors 

Service Entrance Arresters @,a

Branch Circuit Suppressor @ 

Receptacle Plug-in Suppressor 

Plug Strip with Suppressor 8 

Figure 3. Options for power-line surge protection in a residence 
A fundamental application issue in surge protection is grounding. Proper grounding is c

the case of ground-seeking surges such as lightning currents. However, even more critical 

port appliances or systems is the difference in ground reference between port connectio

vuinerabiiiiy issue is addressed in the Nationai Eiectric Code (NFPA 70j, which requires 

grounding conductors of all systems entering a facility be bonded together. As shown in 

(left), a split bolt is used to bond different system grounding conductors to the powe

grounding electrodes. Unfortunately this bonding is not always done and a common, but 

and dangerous to personne! 2nd equipme~f, scen=rrio is shwn ir! figwe 4 (right). 

#6 AWG to ground bus 

fin electrical panel 

#6 ~um'per around 

I ,T,L.- .&. , --->..!- -11 

W llCll SLCCI LUllUUlL all 

Driven ends of the conduit 

Ground Rod shall be bonded to the 

grounding wire. 

CATV and 11 teleohone 

Electric 

Meter 

@ service. 

? A single ground 

I not be less than 

rod, may or may Each service has 

ground rod and 

1 25 ohm, as connected togeth

required by code 

Figure 4 - Minimum code-required service grounding (left) and commonly found, but inc

grounding of existing house (right) 

The incorrect grounding of the house, figure 4 right side, was identified by the study grou

primary problem in residential system surge protection. Even with independent surge prot

the power and communications ports, smart electronic appliances are left vulnerable to 

During a surge, an elevation of only one of the port reference potentials relative to the oth

upset or damage appliances. The problem is depicted in figure 5, which shows the conn

a fa~siiiiil~ iiiachiiie (FAX), to bolh the power system and commi;nication syskm. The S

called TVSS in the industry jargon) on the power port of the FAX can be built-in or 

externally by the end user. The arrester at the service entrance, shown in dotted lines, ma

not be present. According to standard practice, the telephone company has installed a 

interface device [NID) at the point of entry. The NID is installed primarily to protect the te

system and human users from hazards of faults and surges that travel from the premise

the telephone system, rather than to protect from surges impinging from the telephone sys

the premises. Nevertheless, should a surge impinge on the telephone system, the NID wi

to ground, in this case the nearest cold water pipe. 

Consider the case of a surge impinging on the entrance of the telephone system. The surg

shown as "Surge i" in figure 5 flows in the long path from the entrance of the telephone s

the grounding point in common with the power system entrance. The inductance of thi

&.....:--ll.. &--- -5 -:-"- I--...-:-- .....A &LA -..,.-- -..""--& --- ":..A 4- --.,-"-I I-..-A--A 

LY~ILCIIIY L~II~ U IIIILIU-I I~IIII~~, CIIIU LI le ~UI ye LUII~I 11 MI I 115e LU aevel~~ I IUI IUI~U allpele

microsecond. The resulting voltage drop along the current path from the telephone port to 
Assuming 10 pH for L and 500 Alps for dildt, the voltage drop that appears between the p

and telephone port of the FAX is Vdifference = 10 pH x 500 A11 ps = 5000 V. This voltag

sufficient to cause a flashover on the printed circuit boards, or a semiconductor failure in 

Most FAX failures can be readily explained by the elevated reference potential associ

independent surge protection of the two ports. 

Power rL;; TVSS System l.4 Arrester 

Entrance 'I:' 

Cold Water Pipe 

- a a - 

Surge i 

Figure 5. Surge voltage potential difference develops between two protected ports of ap

antenna grounding wires. 

I \( aElectric Meter I 

J 

lL MinX14 AWG wire 

I I I I ! I 

.I \ bpional method is to use a UFER 

ground. A 112" dia., 20'-0 long ste

I I rod or #4 bare copper wire encased i

the concrete footing. 

b2O'-O"4 When steel conduit is used to protect 

I grounding wire, all ends of the conduit 

shall be bonded to the grounding wire. 

Figure 6. Recommended service grounding for enhanced surge immunity 
services, power and telephone, must enter at the same point and must be bonded togeth

point. The NEC requires bonding together the two service grounds, but does not require en

same point. As shown in figure 5 system bonding is provided by the cold water pipe

:I -- a- L : n..r r~, ..-I+,,, ,,c,,r:,~ ~:rr,,,,,, ,~:II ,,,. .,, L. ,,,. , 

eleL;LrIC;ally LUIIIIIIUII tu UULI I selvIC;eb. DUL LII~ vu~~ayt: PULGI ILI~ UIIIGIGIIG~: WII UGGUI~ uetiau

separated entry points. Figure 7 shows the details of how an inter-service bonding point 

constructed and installed. 

\TO GROUND BUS IN PANEL 

INTERSYSTEM BONDING POINR 6"x 6 

x 3" weatherproof box with a 6 position 

grounding lug suitable for #14- #6 wire. Labels 

shall be permanently attached. This is for 

connection of telephone, television, and radio ELECTRICAL METER 

antenna grounding wires. 

G 

AWG BARE COPPER 

OUNDING ELECTRODE 

WEATHERTIGHT FITTING 

GROUNDING BUSHING USE1 

rOR iviETAiiiC 

Figure 7. Details of the essential intersystem bonding point 

Beyond grounding, for a surge-protective device (SPD) to be effective it must provide a p

level, or surge clamping, below a level potentially harmful to the equipment. But the lev

be so i()w that device absorbs energ.y i" its iife is or ii faiis pre

under the stresses resulting from a temporary overvoltage. This principle holds for powe

telephone or cable, albeit at different threshold levels. The SPD must also have a surg

handling capability commensurate with the surge currents that might occur at that locatio

The application of an SPD satisfying the criteria cited above must also take into conside

rest of the circuit where the device is to be applied. For instance, operation of the SPD 

cause adverse side effects. In the case of protection of load equipment which is connect

power system as well as to a communication system (telephone, cable or satellite TV) 

attention must be given to this issue. It is possible that protecting the two interface

equipment, each with a separate SPD, might leave the equipment in jeopardy as the 

overvoltages appearing between the two separate systems, see Key, Martzloff, [2,3]. 

Because effective operation of an SPD invoives aiveriing tine surge to ground, the actuai g

system and its connections in a residence must be taken into consideration. The impedan
NEC-is to provide separate grounds for the power system and for the communications 

Even so, many instances are found of such multiple grounding practice in residence

intentionally, or by accident. A typical example is services such as cable TV that used th

_ -1 -..LJ I-. .--I -- :I- ---..--I:-- -----a I:-- TI-:^ ^"..^--^.^-^-. ... ^^ ^- ^^^^-&^I-l^ 

pipe or an uuruuur 1aur;eL as 11s yruur IUII ~y cur lr letiuur I. I r 11s ar rar lyer~ler 11 was ar I al;c;epiaule 

at the time of installation, but can easily be defeated when the water piping inside or out

house is replaced with plastic. So an improper separation of grounds is camouflaged by t

link of steel pipe going through the foundation wall and sealed in the concrete. 

k distance -------.I 

Suppressor 

1 

Arrester 

C62.41 

Location ---1 

Category B 

C62.41 

Location 

Category A 

Figure 8. Cascade arrangement of upstream and downstream SPDs 

OPEN AND UNRESOLVED ~SSUES 

Several issues remain open at this point and will be addressed in the Guide. We hope 

Guide will resolve some of these issues, but others might take more research to resolve. 

cases an iterative process be among uiiiiiies, and eqLIipmeni inan"fac

to optimize the process instead of exchanging blame. 

One important problem that can be resolved by future research is the coordination of 

cacade, (see figure A). !f the se!ectinn nf a sewice entrance SPD is appropriately coordina

the installation of appropriate SPDs within the residence, then the resources will be allocat

effective manner, both technically and economically (see Martzloff, Lai [4,5]). Clea

coordination can only be achieved by cooperation of the utility, SPD manufacturers, and 

occupants and full knowledge of the characteristics of the SPD's and the surge itself. 

Another problem that can be resolved is the effective protection of equipment connecte

different systems. The concept of equalizing ground references has now been su

advocated to motivate SPD manufacturers to offer SPDs with both power and telephone pr

or both power and i'v' protection, in the same enciosure (see figure 9j. The iEEE design

"Surge Reference Equalizer" was published in 1992, [6]. It has not yet found general us
of surge reference equalizers. At this point, the user is left with some uncertainty as to how 

a particular brand may be compared to another. In this intensely competitive market, cl

counter claims need to be sifted via experiment, a role that some utilities might be willing to 

-- &L-& ..---:-.&:--.-. -L.-..lA -&.-I.,.-..-.- UI Ll Id1 LUI 15UI I I'SI Vl ydl llLdllV113 31 IVUIU dUUlG33. 

Figure 9. Surge reference equalizer to protect multi-port appliances 

Last but mt !east, app!icatim of SPD or !ightning protection must consider risk ana!ysis. P

of appliances against surges generated within a building, or impinging at the powe

entrance, can be accomplished with relatively low cost. However additional investmen

required if multiple services (telephone, TV, and power), and multiple grounds are present

protection against a direct lightning strike may also be needed with cost depending on the 

and location. In any case good practice should be sought in the face of an "act of Go

where damage can be minimized by observing appropriate rules of grounding, bond

protection, as the Guide will propose. However, even with the best practices some degre

will always remain. 

More aiieniion to siiige pi~i~~ii~ii practices is warranted bjj the increased iise and iiihei

vulnerability of residential electronic systems and appliances. While there is consensus on 

principles of protection, specifics and installation procedures for cost-effective application

readily available. By applying data and practices for grounding, lightning and surge prote

specific appliance sets; these needed details are beginning to take form. To be succ

Consumer-Oriented Guide on surge protection will have to address complex topics of who

protection including such techniques as multi-port devices and cascade coordinatio

complete, the Guide should provide a comprehensive document that insurance companies 

to educate their policyholders, employees, and agents about how to reduce damage resu

surges caused by lightning. 
L.G. Byerley Ill, et al, Measurement and Use of Lightning Ground Flash Density,

 

Proceedings of the International Aerospace and Ground Conference on Lightning an

Electricity Williamsburg, VA, September 26-28, 1995. 

Key, T.S. and Martzloff, F.D., "Surging the Upside-Down House: Looking into Upsett

Reference Voltages: Proceedings, PQA'94 Amsterdam, October 1994. 

???ar?z!df, F.D., Mansmr, P.., phipps, KO., and Grady, W.M., "Surging the Upside Dn

House: Measurement and Modeling Results," Proceedings, PQA'95 New York, May 

Martzloff, F.D. and Lai, J.S., "Cascading Surge Protective Devices: Coordination ver

IEC 664 Staircase," Proceedings, PQA'91 Paris, October 1991. 

Lai, J.S. and Martzloff, F.D., "Cascading Surge Protective Devices: Options for Effec

Implementations," Proceedings, PQA'92 Atlanta, September 1992. 
DRIVING

HIGH SURGE

CURRENTS

INTO LONG

CABLES:

MORE

BEGETS

LESS
Driving High Surge Currents into Long Cables:

More Begets Less

Arshad Mansoor, Member, IEEE

Power Electronics Applications Center

Knoxville TN 37932 USA

Amansoor@epri-peac.com

François Martzloff, Life Fellow, IEEE

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.

The paper provides quantitative information on this limitation, as a function of wiring length and current rate of rise.
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 

96 SM 399-6 PWRD A paper recommended and approved by the IEEE 

Surge Protectwe Devlces Committee of the IEEE Power Englneering 

Society for presentat~on at the 1996 IEEEIPES Summer Meeting, July 28 

-August 1, 1996, In Denver, Colorado Manuscript submitted December 

28, 1995, made available for pnnting June 19, 1996 

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 

thesis: If a large surge current is postulated as propagating 

downstream (and then taken as a requirement for the downstream 

SPD), the propagation characteristics of this surge current would 

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. 

Surge 

generator 

Digital 

signal 

analyzer 

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 

Top trace: Voltage, 500 Vldiv 

(Center trace: inactive) 

Bottom trace: Current, 500 Ndiv 

Sweep: 10 psldn 

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 * 

CT2 

Laser 

- 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 

,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 ys 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 = 357 1 * 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 ys 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 ys 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 ys. 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, y s 5 10 20 

Actual rise time, ws 4.3 9.5 13.5 

Maximum di/dt, A!ys 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 psldiv 

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. 

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 ail 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: 

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. 

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. 

The importance of looking for blind spots is, once again, 

demonstrated by the parametric computations, a much 

simpler task than exhaustive equipment-exhausting tests. 

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. 

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 

Kermit 

Center. 

measurements reported in this paper were conducted by 

Phipps at the EPRI Power Electronics Applications 

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 

IEEE and IEC colleagues, was provided by Delmarva Power & 

Light and by Pacific Gas & Electric. 

VI. REFERENCES 

[llIEEElANSI 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." 

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 I8I 2-1 81 9. 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- 

125. 

[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 PhD. in electrical engineering 

from 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. 

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 

wh~ch 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-to-cable 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. 
THE EFFECT OF

NEUTRAL

EARTHING

PRACTICES

ON LIGHTNING

CURRENT

DISPERSION IN

A LOW-VOLTAGE

INSTALLATION
The Effect of Neutral Grounding Practices on

Lightning Current Dispersion in a Low-Voltage Installation

Arshad Mansoor, Member, IEEE

Power Electronics Applications Center

Knoxville TN 37932 USA

Amansoor@epri-peac.com

François Martzloff, Life Fellow, IEEE

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 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 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.
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 [6], [lo], [ll]. In agreement with the latter 

three references, we postulated a 100 kA peak, 101350 ys 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 ys 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 ps 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 Simplification 

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 mQ/m and L = 1 yWm, 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 xenario. 

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 

- 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. 
Time (vs) 

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 * I 

I I 

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 

Figure 6 - Waveforms of currents leaving Building 1, as defined in Figure 5, for 

a 100 k.4, 101350 ps surge terminating& the building earthing system 

' Peak occurs early In event 

Figure 7 - Daisy chain TN configuration with building at opposite end of 

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. 

Building 1 Building 2 Building 3 -* - - 20 rn loom - loom * 

10 kA ID kA 9.5 lc4 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. 

Building 1 Buiidlng 2 Building 3 ---+-----+< 20 rn 100 rn 100 m 

IT. 

Q kA 28 W 38 1cA 

100 kA 

- 

- 

Figure 11 - Daisy chain TT 2-wire configuration with building at opposite end 

of distribution transformer struck by a 101350 ps, 100 kA surge 

"Peak occurs late In the event 

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 oole. 10-Q bld~. 10-9ole. 5-Q bldz 

Current in building earth 2 1 3 1 

Current in service 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. CONCLUSIONS 

Modeling several typical TN and TT configurations of 

neutral earthing practices and scenarios of lightning strike point 

provides insights on significant effects, which should lead to 

more effective application of surge-protective devices (SPDs). 

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. 

Martzloff, F.D., Mansoor, A., Phipps, K.O., and Grady, W.M., "Surging 

the Upside-Down House: Measurements and Modeling Results," 

Proceedings, EPRI PQA'95 Conference, 1995. 

Mansoor, A. and Martzloff, F.D., "Driving High Surge Currents into 

Long Cables: More Begets Less," Paper 96 SM399-6 PWRD, IEEE PES 

Summer Meeting, Denver CO, 1996. 

Rakotomalala, A., Auriol, Ph., and Rousseau, A., "Lightning Distribution 

Through Earthing System." Symposium Record, IEEE International 

Symposium on EMC, Chicago IL, USA, August 1994, pp 419-423. 

Lai, J.S. and Martzloff, F.D., "Coordinating Cascaded Surge Protection 

Devices: High-Low versus Low-High," IEEE Transactions IA-29 No.4, 

JulyIAugust 1993, pp 680-687. 

Birkl, J., Hasse, P., and Zahlmann, P., "Systemgerechter Einsatz von 

Ableitern in Niederspannungsnetzen," Elektrotechnische Zeitschrift 

N0.17, 1994, pp 964-971 

Flisowski, Z., and Mazzetti, C., "Efficiency of Lightning Protection 

System as Screening Measure Against LEMP," Proceedings, EMC'96 

ROMA, September 1996. 

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. 

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. 

Dugan, R.C. and Smith, S.D., "Low-Voltage-Side Current-Surge 

Phenomena in a Single-Phase Distribution Transformer System," IEEE 

Transactions PWRD-13, No. 2, April 1988. 

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. 

Transient Voltage Suppression Devices, Harris Corp., 1991. 

ANSIiIEEE C62.11-1991, IEEE Standard for Metal-Oxide Surge 

Arresters for AC Power Circuits. 

ANSIIIEEE C62.34-1996, IEEE Standard for Performance of LowVoltage Surge-Protective Devices (Secondary Arresters). 

ANSIiIEEE C62.4 1 - 1991, IEEE Recommended Practice on Surge 

Voltages in Low-Voltage AC Power Circuits. 

UL Std. 1449, Transient Voltage Surge Suppressors, 1996. 

Berger, K., Anderson, R.B., and Kroninger, H., "Parameters of Lightning 

Flashes," Electra, 1975-80, pp 1548-1565 

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. 

3. 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. 
THE

DILEMMA

OF

SURGE

PROTECTION

VS.

OVERVOLTAGE

SCENARIOS
The Dilemma of Surge Protection vs. Overvoltage Scenarios:

Implications for Low-Voltage Surge-Protective Devices

Arshad Mansoor, Member, IEEE

Power Electronics Applications Center

Knoxville TN 37932 USA

Amansoor@epri-peac.com

François Martzloff, Life Fellow, IEEE

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

© 1998 IEEE

Reprinted, with permission, from

Proceedings, 8th Annual Conference on Harmonics and Quality of Power, Athens, October 1998

Significance:

Part 2 Development of standards

Part 7 Mitigation techniques

The application of surge-protective devices (SPDs) in low-voltage AC power circuits, in particular metal-oxide

varistors (MOVs) has been influenced by the perception that low-limiting voltage is a desirable characteristic. 

Unfortunately, this low limiting voltage – intended for surge protection – makes the devices more susceptible to

fail under conditions of extended temporary overvoltage (TOV).

Like any electronic component, SPDs will fail if overstressed beyond reasonable limits, and this is not a cause for

rejecting their application, but a cause for concern on ensuring that the failure mode – rare as it might be – will be

acceptable. 

This acceptability must also take into consideration the effect of the available fault current that the power system

can deliver at the point of connection of the SPD. This point needs to be more clearly and specifically stated in

emerging standards on SPD applications.
The Dilemma of Surge Protection vs. Overvoltage Scenarios: 

Implications for Low-Voltage Surge-Protective Devices 

Arshad Mansoor, Member, IEEE 

Power Electronics Applications Center 

Knoxville TN 37932 USA 

Abstract - The application of surge-protective devices in lowvoltage systems faces the dilemma of providing effective limiting 

against surges while not attempting to limit the temporary 

overvoltages that do occur in a power system The paper illustrates 

this dilemma with specific scenarios and presents recommendations 

for reconciling these two conditions through adequate design and more explicit standards. 

I. INTRODUCTION 

The concept of b4Whole-House Surge Protection" has become 

a popular subject of discussion and has in fact been 

implemented by several utilities in North America. In this 

approach, the utility will install a surge-protective device (SPD) 

at the service entrance of the customer and provide additional 

plug-in SPDs. These additional SPDs are presumed to be 

wellcoordinated with the service-entrance SPD, and are 

installed within the customer premises, presumably at the point 

of connection of so-called "sensitive appliances" such as home 

entertainment, computers, and sophisticated kitchen appliances. 

The main purpose of these SPDs is to protect sensitive 

equipment against surges, a mission that they can accomplish 

quite well. However, the failure mode of these devices under 

temporary overvoltages (TOVs) that might be expected under 

abnormal but possible conditions of the power system has 

become cause of some concern for utilities and their customers. 

An SPD should not be expected to protect downstream 

equipment in the case of a TOV and then return to normal 

operation, as it does by definition for surge protection. The 

dilemma for SPD designers is whether to select a maximum 

continuous operating voltage (MCOV) high enough to survive 

common TOVs - but at the price of diminished surge 

protection - or to select surge protection with a lower MCOV 

- and then accept failure of the SPD for infrequent but 

possible TOVs. In any case, one should expect that if a TOV 

at any level would cause the SPD to fail, that failure mode 

should be acceptable. 

Fraqois Martzloff, Fellow, IEEE 

National Institute of Standards and Technology* 

Gaithersburg MD 20899 USA 

This dilemma of surge protection versus overvoltage scenarios 

has been created by the industry's obsession with providing 

very low clamping voltages for surge mitigation (Martzloff & 

My, 1989[1]). And now, the need to ensure coordination of 

the "cascade" of the service-entrance SPD and the plug-in 

SPDs has exacerbated this situation. The issue of cascade 

coordination has already been debated at length in the literature 

(Martzloff & Lai, 199 1 [2]); (Stonely & Stringfellow, 199 1 [3]); 

(Hostfet et al., 1992 [4]); (Rousseau & Perche, 1995 151). 

Further debate or exhaustive references to the many papers on 

that subject is not our purpose. It is mentioned here only as a 

contributing factor to the dilemma, but a factor that cannot be 

ignored in a complete assessment. 

11. SURGE PROTECTIVE DEVICES 

FOR LOW-VOLTAGE SYSTEMS 

Ihe introduction of metal-oxide varistors (MOVs) in the 

seventies was a timely innovation, concurrent with the 

increasing use of semiconductors in consumer products. While 

these semiconductors opened new opportunities, their relatively 

low tolerance for surges created a strong demand for better 

surge-protective devices. Unfortunately, market competition 

encouraged a downward "auction" that led to attributing high 

value to low clamping voltages, a situation unwittingly 

encouraged by the listing of "transient suppression levels" 

stipulated in UL Standard 1449 [6]. The list begins at 330 V 

for SPDs intended for 120 V circuits, although there is good 

evidence that most consumer loads do not need such a low level 

of protection (Anderson & Bowes, 1990 171); (Smith & 

Standler, 1992 [8]). 

The generic structure of typical low-voltage residential power 

systems is shown in Figure 1 for the case of a detached home. 

Underground service has similar characteristics. This system 

extends from the outdoor line-side of the service drop all the 

way through the premises wiring, including plug-in type SPDs. 

Figure 1 also shows the various locations where an SPD can be 

installed. 

* Electricity Division, Electronics mad Electrical Engineering Laboratory, 

Technology Administrution, U.S. Del~urtment of Commerce. 

Contributions~m the Nufionul Institute ojstundardr and Technolagy are not 

subject to U.S. copyright. 

Paper accepted for presentation at the 8U International 

Conference on Harmonics and Quality of Power 

ZCHQP '98, jointly organized by ZEEEIPES and NTUA, 

Athens, Greece, October 14-16,1998 

0-7803-5105-3/98/$10.00 0 1998 IEEE 

Typically, there are six locations. The first three: O at the 

outdoor weather-head, Q at the service entrance, and O on the 

line side of the main disconnect are within the scope of IEEE 

Std C62.34 [9]. The next three are within the scope of ongoing 

IEEE project P62.62 [lo]: @I at the load side of the panel 

disconnect, 6 at a permanently wired receptacle, and 8 as a 

plug-in device. Locations O through Q are within the premises 

wiring and therefore under the control of the end-user, while 

locations O and Q are under the control of the utility. 
Branch arcuit 

0 Load side d Service dmp @ Load side d mPh Overevnnt d8W 

@ Meter base adapter @ Penanently cumded maplade 

@Linesirkjdmainovercurrentdevke @Ccmicme&dwplug.ln 

Nat shown: SPDs incorporated in equipment power port 

Figure 1 - Possible SPD locations for a residential building 

An emerging requirement in standards for application of the 

low-voltage SPDs under development at the International 

Electrotechnical Commission (IEC 61643-1 [I 11) as well as at 

the IEEE (P62.62 [lo]) is the provision of a "disconnector" 

intended to discomect a failing SPD. Failure under conditions 

exceeding the SPD capability is recognized as unavoidable, but 

the consequences of such failure are made acceptable thanks to 

the action of the disconnector. 

Unfortunately, some ambiguity has crept in the interpretation 

of this requirement. In some cases is has been interpreted as 

only disconnecting the failed SPD component from the power 

system, but leaving the load energized - and without further 

surge protection (Martzloff, 1998 [12]). If the disconnector is 

of this latter type, the SPDs components of an SPD package 

will fail under TOV conditions, presumably in a safe manner, 

but then allow the TOV to be applied to the downstream load. 

This is undesirable for the typical user who values equipment 

protection above continuity of operation. Finally, there is the 

worst case, as reported in many anecdotal instances, where the 

failure mode of SPDs under TOV conditions has not been 

graceful, to say the least. These instances, while not very 

frequent, have led to new testing requirements for failure 

modes from the Underwriters Laboratories in the updated 

Second Edition of their UL 1449 [6] Standard for low-voltage 

surge-protective devices ("TVSS" in the industry jargon). 

Given this unsettled situation, it will be useful to review the 

scenarios that can lead to failure of an SPD component in an 

SPD package, in particular under TOV conditions. For the 

sake of completeness, we will describe first some failure 

scenarios under surge conditions, as they have some bearing on 

the disconnector design. 

III. FAILURE MODES UNDER SURGE 

CONDITIONS 

For a correctly applied SPD, failure under surge conditions 

should be a very rare occurrence. Nevertheless, one can 

enumerate the following failure scenarios in a field application 

- including misapplications: 

A single, large, and not anticipated impinging surge 

exceeds the capability of the SPD. An example of this 

situation can be the presence of switched capacitor banks; 

A succession of surges, such as multiple lightning strokes, 

exceeds the capability of the SPD. This situation has been 

identified for distribution arresters (Darveniza, 1997 [I 31) 

and might also occur for low-voltage SPDs; 

A thermal runaway is launched in an SPD exposed to high 

ambient temperatures at the time when a surge (within 

specifications for normal ambient) occurs; 

In the questioriable scenario of an alleged "degraded" SPD 

(Stringfellow, 1992 [14]), a thermal runaway is launched 

by the heat generated during a within-specifications surge. 

For all these scenarios, the ultimate failure mode depends on 

the fault current that the power system can deliver at the point 

of connection of the SPD. For maderate fault currents, such as 

on branch circuits, the disconnector can generally provide 

protection. The fuse design can still be a challenge: carry the 

load current, carry the specified surge current, but melt in case 

of a power-frequency fault current resulting from failure of the 

SPD component. For very high available fault currents, such 

as that prevailing at some service entrances close to a large 

distribution transformer, successful clearing may be a greater 

challenge. Furthermore, coordination of overcurrent protection 

is more difficult, compared to branch circuits inside the 

building where the wiring impedance and the rating of circuit 

breakers in the panel can ensure proper coordination. 

IV. FAILURE MODES UNDER TOV CONDITIONS 

Three major types of TOV-induced failures can be identified 

for low-voltage SPDs: 

Moderate TOVs associated with power system faults, such as 

a line-toearth fault in a three-phase system, creating a 1.73 

times nod line voltage in the other phases. Ferroresonance 

can also produce moderate but significant overvoltages. 

Extreme TOVs associated with the commingling scenario 

(accidental fall of conductors of a higher voltage upon 

conductors of a lesser voltage). No conventional, varistor-only 

SPD can be expected to survive such a scenario. 

Double voltage TOVs associated with the loss of neutral in a 

single-phase, threeconductor, earthed center-tap system such 

as the 1201240 V service typical of North American systems. 

Depending on the philosophy of the system designer, in 

particular the utility for the case of a service-entrance SPD. 

survival or expected but acceptable failure can be stipulated for 

the loss-of-neutral scenario. 
a) Moderate TOVs: System Faults 

Among abnormal conditions that can produce temporary 

overvoltages, we give two examples of incidents resulting in 

overvoltages not exceeding twice the normal voltage. Some 

SPDs based on the misconception that a very low clamping 

voltage is desirable might not survive such moderate TOVs. 

Single-phase faults to earth on a three-phase system produce 

a shift in the unfaulted phases. The severity of the voltage rise 

depends on the fault location, the system impedance, and the 

earthing practices. For a 'IT power system in steady-state, the 

neutral will by symmetry be at the same potential as earth, as in 

Figure 2 (a). However, if an earth fault occu on one phase, it 

will cause a shift of potential as the system attempts to maintain 

balance. The neutral will be elevated and a corresponding shift 

will be experienced by the other two phases. In the worst case 

of a completely isolated system with a bolted fault to earth on 

one of the phases, we could have the situation depicted in 

Figure 2 (b). This shows a neutral which has been elevated to 

one per unit, causing the unfaulted phases to drift up to 1.73 per 

unit with respect to earth. 

For three-phase systems with an artificial neutral through an 

earthing transformer or a finite resistance, the earth impedance 

is high. This arrangement limits the fault current during a 

singleline-toearth fault, but allows enough for fault detection 

using overcurrent relays. It also inserts some impedance 

between the system neutral and the actual earth so that the 

voltage on the unfaulted phases will shift toward somewhat less 

than line-line value as shown in Figure 2 (c). 

Figure 2 - Effect of a single phase-to-earth fault 

Series resonances occur in a power system when a series 

circuit consisting of an inductance and a capacitance is excited 

at its natural frequency. As an example of this condition, 

Figure 3 shows a portion of a temporary overvoltage resulting 

from a power system switching incident (restoring power phase 

by phase after interruption [15]), which produced a sustained 

rms voltage exceeding 150% and lasting four seconds. 

Source: EPRl Report [l5] 

Figure 3 - Temporary overvoltage caused by ferroresonance 

b) Extreme TOVs: Commingling 

In this real-life scenario, the low-voltage SPD connected on the 

secondary side of the distribution transformer, and normally 

energized at its rated voltage, is first brought to failure by the 

large overvoltage resulting from commingling conductors. In 

an overbuilt system, a collision of a vehicle with a pole, or 

breaks caused by icing, the conductors of the higher voltage 

distribution system or sub-transmission system can fall on the 

lower voltage distribution system. Such accidental contact 

injects an intruding voltage for a few cycles, until the higher 

voltage breaker clears the fault. 

Figure 4 shows a simplified one-line diagram of the two 

medium-voltage systems (the intruding MV1 and the victim 

MV2) being accidentally commingled. In this figure, the bond 

between the two systems earth connections is shown in dotted 

line to present the generic case of commingling. In an overbuilt 

system with common neutrals, a solid bond exists where the 

dotted tine is shown in the diagram. In the scenario of a simple 

crossing of two systems (not overbuilt along the right of way), 

or delta systems, a solid bond might not be present, and the 

fault current from MV1 will involve the earthing impedances 

shown in the diagram In that case, the intruding voltage Vi 

might be less than the system voltage MV1 but still enough to 

precipitate failure of an SPD on the secondary. 

Figure 4 - Commingling overvoltage applied to the distribution 

transformer primary and reflected on the LV side 
c) Double-voltage TOV (Loss of neutral) 

There are many situations where loss of neutral can occur. 

Some can be of a transient nature, such as a loose connection, 

while some might be permanent until repaired, such as a 

mechanical break or corrosion of the neutral conductor. While 

the latter might not linguistically fall under the label of 

'temporary', the consequences are the same. In a three-wire, 

single-phase system typical of North American practice, this 

condition has been observed many times. 

Figure 5 shows a system where one side of the supply (Ll) is 

lightly loaded, while the other side (L2) is heavily loaded, 

Under normal conditions, the two sides remain at normal 

voltage. Should the neutral connection be lost, then the voltage 

at mid-point is determined by the ratio of impedances on the 

two sides: the L1 side experiences an overvoltage that can 

approach twice nonnal. Any SPD connected to this side will 

then be exposed to the twice-normal voltage, with an available 

current determined by the impedance 22. This current, by the 

very design of the circuit, will not be interrupted by the 

overcurrent protection and is available to generate substantial 

heat in the SPD that has failed as the result of the overvoltage. 

A disconnector designed to clear larger fault currents, such as 

those occumng if the SPD fails while the neutral is connected, 

might not be capable of clearing the limited current. Other 

schemes are then necessary to ensure an acceptable failure 

mode, such as a thermally-activated disconnect. 

1996 National Electrical Code, 230-65, Availabk Short-circuit 

Cumnt 

"Service equipment shall be suitable for the short-circuit 

current available at the supply terminals." 

IEEE C62.341996, Standard for Performance of Low-Voltage 

Surge-Protective Devices (Secondary Arresters), Article 7.6 

"If the manufacturer claims a fault current withstand rating, 

then that rating shall be verified." 

IEC 61643-1 (FDIS November 1997) Surge protective devices 

connected to low-voltage power distribution systems - Part I: 

Performance requirements and testing methodr, Artick 6.2.11 

"The SPD shall be able to carry the paver short-circuit 

current until it is interrupted either by the SPD itself; by an 

internal or external overcurrent disconnector, or by the backup 

overcurrent protection." 

It is noteworthy that none of these documents specify a value 

for the available short circuit, but rather leave it to the 

discretion of the manufacturer, while imposing criteria of 

acceptability after the failure. The NEC tersely requires the 

device to be "suitable," without elaboration. The two standards 

applicable to a service-entrance SPD - where the available 

fault current can be quite high - acknowledge the possibility 

of failure and significance of the available fault current, but do 

not stipulate specific values. 

In an attempt to obtain information on what levels of fault 

currents should be considered, an informal survey was 

conducted among a few utility engineers. It turns out that very 

few utilities limit the available fault current at the service point 

of residential customers but many people are under the 

misconception, as cited below, that all residential service load 

centers and breakers are limited to an available fault current of 

10 kA and therefore they assume that actual available fault 

current must be less than 10 kA. 

Figure 5 - Three-wire, singlephase system 

where a broken or intermittent neutral connection 

creates a voltage imbalance 

V. AVAILABLE FAULT CURRENT 

As described in the preceding paragraphs, several scenarios can 

produce failure of the SPD, ultimately offering a very low (but 

not zero) impedance to the flow of the fault current that the 

power system can deliver. For each scenario, the available fault 

current at the point of connection of the SPD will have a very 

significant effect on the failure mode. 

The significance of available fault current is recognized in 

standards, but the value that should be considered is generally 

left undefined. The following statements can be found in 

published standards. 

Interestingly, none of the individuals consulted on this issue 

could cite a standard or even a document with the status of 

a consensus guide that does stipulate a current value. Two 

perceptions seem to prevail among the individuals consulted, as 

quoted below: 

Based on the observation that breakers for typical 

residential service panels (up to 200 A) generally have an 

interrupting capacity of 10 kA, it would be logical that 'the 

available fault current at the service entrance would be in 

the same order of magnitude. 

Based on the reality that a service entrance connected 

clo,e to a large distribution transformer (such as a garden 

apartment or high-rise) will have available fault currents 

in excess of I0 kA, it would be prudent to review the 

specifics of the situation. 

Furthermore, these standards do not require that the SPD 

application data state a limit of acceptable fault current. 
VI. EMERGING STANDARDS 

a) SPDs installed downstream from the service entrance 

In recognition of the issues raised by failure modes of the SPDs 

within the scope of UL Standard 1449 - downstream of the 

main disconnect - the second edition of this standard [6] now 

requires demonstration of an acceptable behavior in the failure 

mode of SPDs exposed to various overvoltage scenarios. 

Specifically, among the many tests required by UL, three tests 

address the issue: 

Temporary overvoltage with high available current - This test 

is stipulated un&r article 37.2 of UL 1449, calling for extended 

exposure to 125% of normal line voltage with either an 

acceptable temperature equilibrium being attained or until an 

internal disconnect device operates. This test will demonstrate 

capability of sustaining moderate overvoltages, primarily 

providing a margin against high system voltage, but not the 

higher levels of temporary overvoltages covered in the next 

test. The available fault current specified for this test is defined 

as a function of the ampere rating of the service over a range of 

200 A to 25 000 A. 

Fullphase voltage with high available fault current - This test 

is stipulated under article 37.3 of UL 1449, calling for exposure 

to the "full phase voltage" as shown on Figure 2(b) of this 

paper. The same criteria as above apply, namely acceptable 

temperature equilibrium or operation of an internal disconnect. 

The available fault current specified for this test is &fined as 

a function of the ampere rating of the service over a range of 

200 A to 25 000 A. 

Overvoltage with limited current - This test is stipulated under 

article 37.4 of UL 1449, with overvoltage values presumably 

corresponding to a loss of neutral scenario, and the associated 

low values of current supplied by the connected load, as in 

Figure 5. 

It is noteworthy that in the stipulations of these three tests, the 

emotionally charged word "failure" is not used. Instead, a list 

of unacceptable conditions is given, including emission of 

flame, molten metal, flaming particles, charring of adjacent 

material, ignition of enclosure, or creation of openings leaving 

live parts accessible. 

b) SPDs installed upstream from the service entrance 

For SPDs connected upstream from the service panel, the 

recently published IEEE Standard C62.34 does describes a 

loss-of-neutral scenario with limited current, similar to the UL 

37.4 test. However, a demonstration test is not mandated, as 

the consensus development process settled for a weaker 

statement: "if the manufacturer claims a loss of neutral 

withstand capability ... then that capability shall be venjied ..." 

Thus, the inference might be made that if no claim for loss-ofneutral withstand capability is made, no demonstration test is 

required. However, an additional paragraph in the standard 

does mention "...must fail in an acceptable manner. " 

From this brief overview of emerging standards, it appears that 

on the low side of available fault current, both the UL and the 

IEEE standards have recogruzed the issues of acceptable failure 

modes associated with temporary (or quasi-permanent) 

overvoltages. On the other hand, at the high side of available 

fault currents, it seems that insufficient recognition of the issue 

still prevails, as discussed in the preceding section on the 

significance of the level of available fault currents. 

Nevertheless, Figure 6, offered as background information in 

a tutorial addressing considerations on the revenue meter 

environment (Ward, 1980 [16]), shows typical values of fault 

current as a function of distribution transformer size and length 

of service drop. That figure clearly shows values in excess of 

10 kA, but somehow this information has not been fully 

recognized by the SPD community. 

- - 

AVAILABLE FAULT CURRENT 

WITH rOorr SERVICE COR 3 

TRANSFOWER SIZES 

tm AL SERVICE -CABLE LENOW -FEET 

Source: (Ward, 1980 [I 61) 

figure 6 - Fault current amplitude as a function of. 

transformer rating and length and size of service drop 

vn. POSSIBLE SOLUTIONS 

Among possible solutions, two approaches may be 

considered: making the SPD less sensitive to TOVs, and, 

in any case, ensure that if failure is unavoidable under 

extreme stress, it will be in an acceptable mode. 

The obvious way to desensitize SPDs to TOVs is to 

design them with a higher MCOV. However, as the 

higher MCOV in a varistorsnly SPD means a higher 

surge-limiting voltage, there is a limit beyond which such 

an SPD becomes useless (op cit., [2-51). 

A pos:.ble solution may be in reviving the concept of a 

gapped arrester for the upstream SPD of a "whole house" 

scheme (Mansoor et al., 1998 [17]). There, the initial 

let-through associated with the gap volt-time response 

can easily be mitigated by the downstream SPD, while 

the gap prevents the SPD from becoming involved with 

moderate TOVs. 
Of course, for the (rare) commingling scenario, little can 

be done but to ensure a graceful failure. This condition 

should be an implicit requirement, but, as discussed in the 

section on available fault current, the implications of such 

a requirement apparently have not been recognized by all 

interested parties. Even among the community of SPD 

engineers, there has been some reluctance to accept the 

concept that temporary overvoltages should be addressed 

in documents discussing the surge environment. 

VIII. CONCLUSIONS 

The dilemma of providing a suitable surge protection of 

load equipment by means of surge-protective devices, 

while ensuring acceptable response of these surgeprotective devices to unavoidable temporary overvoltages 

raises several application issues that demand attention. 

"Acceptable response" can be interpreted either as survival 

of the SPD (a challenge to the coordination of cascades) or 

as accepting failure, but within well-defined conditions of 

the failure mode (a challenge for applications where high 

available fault currents prevail). 

Coordinating a cascade of surge-protective devices can be 

solved by providing a gapped arrester at the service 

entrance, which will coordinate with the de facto situation 

of low limiting voltage SPDs inside the building. 

The need for a service-entrance arrester to withstand the 

scenario of lost neutral can be satisfied by a gapped 

arrester having sufficient maximum continuous operating 

voltage capability. 

Emerging standards for low-voltage SPDS have given new 

recognition to the importance of taking into consideration 

temporary overvoltages in the design of SPDs. 

Notwithstanding conclusion (3, SPD application standards 

as well as performance and test standards should be more 

explicit in defining how to deal with the issues raised by 

available fault current in case of unavoidable SPD failure. 

IX. REFERENCES 

Martzloff, ED. and Leedy. T.F., "Selecting Varistor Clamping 

Voltage: Lower Is Not Better!" Proceedings, 1989 Zurich 

International EMC Symposium. 

Martzloff, F.D. and Lai, J.S., "Cascading Surge-Protective 

Devices: Coordination versus the IEC 664 Staircase," 

Proceedings. PQA 91 Conference. 

Stringfellow, M.F. and Stonely, B.T., "Coordination of Surge 

Suppressors in Low-Voltage AC Power Circuits," Proceedings, 

Fonun on Surge Protection Application, NISTIR-4657. August 

1991. 

Hostfet, O.T., Hervland, T., Nansen. B. and Huse, J., 

"Coordination of surge-protective devices in power supply 

systems: Needs for secondary protection," Proceedings, 

International Conference on Lightning Protection, Berlin, 

September 1992. 

Rousseau, A. and Perche, T., "Coordination of Surge Arresters 

in the Low-Voltage Field," Proceedings. 7th International 

Telecommunications Conference (IhTELEC), 9SCH35824. 

UL 1449 Standard for Safety for Transient Voltage Surge 

Suppressors, Second Edition, August 1996. 

Anderson. L.M. and Bowes, K.B., 'The Effects of Power-Line 

Disturbances on Consumer Electronic Equipment." IEEE 

Transactions PWRD-5, No.2, April 1990. 

Smith, S.B. and Standler, R.B., 'The Effects of Surges on 

Electronic Appliances," IEEE Transactions PWRD-7, No.3, 

July 1992. 

IEEE Std C62.34, IEEE Standard for Pe$ormance of LowVoltage Surge Protective Devices (Secondary Arresters), 1996. 

IEEE Project P62.62 - Application Guide - Low-Voltage SurgeProtective Devices (ongoing). 

IEC Document 37A163EDIS -Draft IEC 6 1643- 1 : Surgeprotective devices comected to low-voltage distribution systems 

Part 1: Performance requirements and testing methods, November 

1997. 

Martzloff, F.D., "Interpretation and Misinterpretation of TVSS 

Discomector Indications," Scheduled for publication in Power 

Quality Magazine, July/ August 1998. 

Darveniza, M., Turnma, L.R., Richter, B., and Roby, D.A., 

"Multiple Lightning Currents and Metal-Oxide Arresters," IEEE 

Transactions PWRD-12, No.3. July 1997. 

Stringfellow, M.F., "Fire Hazard of Surge Suppressors," 

Proceedings, Fiflh Annual Power Quality Conference, Irvine 

CA, 1992. 

EPRI Report An Assessment of Distribution System Power 

Qualify, Electric Power Research Institute. Palo Alto CA, 1996. 

Ward, D.J., "Secondary Fault Currents at the Service Entrance." 

Proceedings, 55* Annual Electric Meter School and Conference, 

University of Florida, Gainesville FL, 1980. 

Mansoor, A., Martzloff, F.D.. and Phipps, K.O.. "Gapped 

Arresters Revisited: A Solution to Cascade Coordination" Paper 

PE-114-PWRD-0-12-1997, IEEE Winter Power Meeting, 1998. 

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. 

Francois 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 13 

years at the National Institute of 

Standards and Technology. He is 

contributing to several technical committees for the development of standards 

on EMC, surge protection and Power 

Quality in the IEEE and the IEC. 
THE FALLACY

OF

MONITORING

SURGE

VOLTAGES: 

SPDS AND

PCS GALORE!
The Fallacy of Monitoring Surge Voltages: 

SPDs and PCs Galore!

Arshad Mansoor

EPRI PEAC Corp

Knoxville TN 37932

Amansoor@epri-peac.com

François Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899

f.martzloff@ieee.org

Kermit Phipps

EPRI PEAC Corp

Knoxville TN 37932

Kphipps@epri-peac.com

Reprinted from Proceedings. PQA ’99 Conference

Significance:

Part 2 Development of standards

Part 5 Monitoring instruments

Recent projects of monitoring Power Quality in AC systems have focused on voltage surges rather than current

surges. The predictable results of such monitoring – low apparent surge voltages – which in fact only monitors

whatever limiting voltage is allowed by the proliferating SPDs and PCs, do not reflect the surge activity, now confined

to surge currents flowing into the “attractive” paths of the SPDs and the capacitors included in the switch-mode power

supplies of PCs (and other electronic appliances).

At standard-writing times, questions have emerged as to why mention “large” surges when monitoring shows only low

voltages. W ith proper perspective, it becomes apparent that the proliferation of these voltage-limiting, surgeabsorbing SPD s and PCs are the explanation. W hile voltage surges might now

no longer be a threat, the possibility of substantial current surges is indeed a threat to equipm ent.
The Fallacy of Monitoring Surge Voltages: 

SPDs and PCs Galore! 

Arshad Mansoor and Kermit Phipps Fran~ois Martzloff 

EPRI-PEAC National Institute of Standards 

1052 1 Research Dr. and Technology ' 

Knoxville, TN 37932 Gaithersburg MD 20899 8 1 13 

Abstract 

To support the recommendation of shifting transienl rnonilori~ig liwi vvltage surges lo 

surges, the paper presents experimental results as well as numerical modeling results demon

two mechanisms causing an apparent decrease of surge activity in low-voltage ac power c

The first mechanism is the proliferation of surge-protective devices, a situation which is by n

recognized. The second, which should also have been recognized, apparently escaped 

so far: the proliferation of electronic appliances containing a switch-mode powcr supply th

effectively places large surge-absorbing capacitors across the ac power systems. 

1. Introduction 

This paper is unabashedly tutorial, and some of the themes presented here might be quite f

to some of our readers. However, we have observed that these concepts, which could alm

characterized as obvious when given the benefit of hindsight, are still not widely recogniz

Therefore, we will present them to th~s forum, and illustrate their validity by experimental 

measurements and numerical modeling. The three major themes of this presentation are: 

1. There is a logical explanation for the apparent decrease in the level of voltage surges r

in recent power quality surveys: the proliferation of surge-mitigating devices. 

2. The present practice of recording voltage surges can lead to erroneous concepts on sur

protection, which in turn can lead LU equiprr1~111 fililu1t;s. 

3. There is an unfulfilled need to develop and deploy power quality monitors that can 

characterize the energy-delivery capability of a surge event. 

The first theme will be introduced by a historical perspective showing how voltage surges 

became the focal point of monitoring surges, even before the term "power quality" was co

Experimental measurements and numerical modeling will be described to illustrate the eff

the proliferation of surge-protective devices (SPDs) and of new electronic appliances that 

rcspcctivcly as intcndcd and unintcndcd surge-mitigating devices. The next two theme:; w

only briefly discussed because, once the first is accepted, these two follow quite logically. 

I Electricity Division, Electronics and Electrical Engineering Laboratory, Technology Adminrsrrution, 

U.S. Department of Comti~crce. 

Conrriburions from rhe Nariut~al Insrirure of Srunciurd.~ and Tet-hnolngj ore nor ~ubjecr ru U.S. Copyri~lzr. 
2. Historical Perspective 

The proliferation of SPDs in low-voltage ac power circuits has been recognized as nnc: nf t

causes of the apparent decrease of the surge levels recorded in recent power quality survey

change in the occurrence of surges began to be recognized [Dorr, 19951 ' and explanation

offered attributing the phenomenon to the emerging proliferation of SPDs in low-voltage ac 

circuits [Martzloff, 19961. It is noteworthy that for many people, the term "surge" is equ

to "transient overvoltage" to the point that the phenomenon has generally been recorded by 

instruments acting as voltmctcrs, and thc tcrm was without much scrutiny accepted in gcnc

meaning a voltage surge. 

For instance, in the bi-lingual publications of the International Elecrrotechnical Commission 

the French text which parallels the English text uses the term "surtension" (meaning overvo

because there is a mind set, reinforced by the lack of a neutral term in French that could be 

applied to either or both voltage surges and current surges. In the United States, the Under

Laboratories perpetuates this narrow perspective by calling SPDs "Transient Voltage Surge 

Suppressors" (italics ours) [UL Std 1449, 19961, although both recent IEC and IEEE defini

of SPDs introduce and emphasize the concept of current surges as well as voltage surges. 

Now less recognized but significant, a similar cause for the apparent decrease of voltage surg

is becoming important as more and more electronic appliances depend on a switch-mode 

supply with a rectifier-capacitor DC link. Through the rectifies, the DC link capacitor w

typically in the order of 200 pF to 500 pF, offers a low-impedance path to current 

impinging on the power port of these appliances. With many such appliances connec~ed in a

user installation, the effect is that of quite a large capacitor being connected across the ac 

Small wonder then that even largc surge currents (for instance, with the capability of deliv

currents of 3 kA, 8/20 ps) [ANSVIEEE C62.41- 19951 can no longer raise the voltage acros

mains to the high values sometimes reported in earlier surveys of surge voltage occurrence

Another historical mind-set has been to recognize the origin of surges only as a voltage eve

while in fact it can be either a true induced-voltage event, or the end-result of the injection 

surge current somewhere in the power system. Typical induced-voltage surges are associat

with the electromagnetic coupling into the power circuits of the field created by a nearby (b

not direct) lightning flash. Such voltage surges, which can develop substantial voltages in 

impedance circuit loops [Martzloff et al., 19951 can easily be mitigated with relatively sma

SPDs because their energy-delivery capability is relatively small [IEC document 64/1034/C

19981. In contrast. current surges are produced either by the dispersion of the current assoc

with a lightning flash when a direct strike injects current at some point of the power system

close vicinity or at a more remote point of the power distribution system [Mansoor et al., 1

Another source of current surges is switching surges involving the injection of residual ene

into parts of the power system. The energy-delivery capability of these current surges can 

substantial, and be a threat to the survival of improperly sized SPDs. Note in passing the u

thc tcrm "cncrgy-dclivcry capability" and not "cncrgy in thc surgc." Somc of our rcndcrs a

now familiar with that theme [Lindes et al., 19971 - perhaps even tired of seeing it repeate

but the sad truth is that usage of the term "surge energy" is still rampant. 

I Citations appearing in the text as [Author, date] are listed in alphabetical order in Section 7, Bihlio~raphy 
The significance of making the distinction between recording current surges versus recor

voltage surges is very important for equipment designers. A decision to provide only mo

surge withstand capability for an SPD incorporated at the powcr port of thc cquipment m

made because the contemporary surveys reveal few and moderate (voltage) surges. When c

with the misconception that "the lower the clamping voltage, the better" [Martzloff et al

the result can be disastrous. We have in our laboratory 'morgue' two examples of such m

produced devices incorporating an inappropriate SPD that led to early mortality of the pr

Without identifying the culprit - perhaps a harsh word for a designer who was lulled in

position by referring to misleading reports on surge actlvlty - but to illustrate the situati

can name the two products: a compact fluorescent bulb, and a remotely-controlled ceilin

To recite a recurrent theme in our tutorial presentations ("Transients Are Here to Stay"), 

surges might appear to have faded away, but current surges are still here, ready to destroy 

SPD incorporated in a design based on the misperception of fading voltage surges, foster

recording only voltage surges. 

3. Experimental Measurements 

To illustrate the effect of nonlinear SPDs as well as linear capacitors connected across th

the Power Electronics Applications Center (PEAC) "Upside-Down House" [Key et al., 1

was used to inject surges into the service entrance of the I Jpside Down House with variou

combinations of SPDs and/or personal computer (PC) power supplies connected at the en

two branch clrcu~ts, one 9-m long, the other 3h-rn long. In Figure I, currents In the branc

voltages at the nodes are identified respectively as I, I, I, and Vo, V, V,, . The chargin

setting of the Combination Wave surge generator was kept constant to provide a 2 kV, 1

open-circuit voltage (OCV). In a second series of experiments, the setting was increased 

This second series, not reported here in detail because of limited space, confirmed the ex

nonlinear response of varistors and the linear response for capacitors only. 

Figure 1 - Upside-Down House branch circuits 

3.1 Metal-oxide varistors only 

Given the known and predictable behavior of multiple SPDs - often reported in the liter

to the point that bibliographic citations would take several lines to list all of them - this 

the experiment was performed only to provide a baseline. Metal-oxide varistors (MOVs

150 V, 20 mm diameter, were used for this experiment. To record all interesting current

voltages, two shots are necessary when using the 4-channel digital signal analyzer; theref

oscillogram in the figures contains the trace of Vo to serve as a common reference (Figur
Vo 1 kV/div 

1, 550 Ndiv 

1, 500 Aldiv 

I, 5OONdiv /- 

Sweep: 5 pddw 

Vo 5

V, 

V,, 

Currents in branches Voltages at nodes 

790 1000 900 790 410 ' 330 

Figure 2 - Typical recordings1 made during the first experiment - SPDs only. 

In Figure 2, the current traces (left oscillograrn) show the unequal sharing between the tw

MOVs, reflecting the difference in the inductance of the two branch circuits. Because the 

uT 11ic cwu ~I~~IICII-L~ILLI~L ULII~~;I~~S ale nut simultancuus (the cu~-~-cnt inhc longcr branch c

takes longer to build up) their sum seems to exceed the peak of the injected current, I,. 

The voltage traces (right oscillogram) show how the voltage at the service entrance, Vo, is 

mitigated from the 2 kV open-circuit voltage supplied by the generator that would propag

without attenuation in the absence of a surge-mitigating device [Martzloff et al., 19861. T

voltage Vo is the sum of the clamping performed by the varistor at the node V, and the in

voltage drop in the 9-m long connection. Note that this inductive voltage is additive duri

rise of the current I, and subtractive during the fall, hence the apparent "overshoot" in th

of V, compared to the flat-top trace of the varistor at V,. The voltages at V, and V,, are t

typical clamping voltages of the MOVs corresponding to the current they carry. 

In a subsequent experiment with a 4 kV open-circuit voltage setting of the generator (twi

value of Figure 2, but not reported here in detail because of limited space), the voltages a

the varistors, predictably, were nor substantially increased. However, the substantially in

current in the 9-rn long branch circuit (from I000 A to 2800 A, resulting from the nonlin

response of the varistor) produced an increased inductive effect to the point that the volta

the service entrance was raised to 1300 V from the 790 V recorded for the case of Figure 

Table 1, at the end of this section, presents a summary of the peak values recorded in the 

combinations of components, branch circuits, and amplitudes of the injected surge. 

The experimental values shown in Figures 2, 3, and 4, and in Table I were recorded with a digital signa

They have been rounded off to the nearest ten to convey a simpler set of numbers, uncluttered by a precise 

Uncertainty in these measurements is not an issue here as it does not affect the general conclusions. 
In a second set of experiments, less easily predictable would bc thc bchavior of the still 

circuit involving the capacitors of a PC power supply (440 pF each in this experiment) w

receiving a surge originating from the complex RLC wave-shaping network of a Combin

Wave surge generator (Figure 3). 

V, 1 kV1div 

1, 550 Ndiv 

1, 500 Ndiv 

I, 5OONdiv 

Sweep: 5 psldiv 

Currents in branches 

- -1- 

Sweep: 5 psldiv 

v " 

(peak V) 

Voltages at nodes 

Tir? 

Figure 3 - Typical recordings made during the second experiment - Capacitors on

Is 

(peak A) 

In Figure 3, the voltage at the service entrance, V,, is mitigated from the 2 kV open-circu

voltage supplied by the generator. However, because of the interaction between the RLC 

components in the wave-shaping network of the Combination Wave generator on the one 

and the capacitances of the PC power supplies and inductances of the branch circuits on 

hand, this voltage rings around an average voltage level of 1300 V, reaching a peak of 17

To reflect this situation, the tabulation of the voltages in the figure shows two lines, peak 

average. Thus, the mitigation effect is degraded by the ringing. Nevertheless, one can ex

that a< more PC% wni~ld he added, the ringing freqi~ency would hecome lower and the vo

peaks lower. 

3.3 Capacitor and MO V 

'Y 

(peak A) 

In a third set of experiments, a capacitor was connected at the end of the Y-m branch circ

and an MOV was connected at the end of the 36-m branch circuit. Interest in this particu

configuration was motivated by the desire to show how an MOV would mitigate the ring

that was observed in the preceding experiment at the end of the 36-m line (V,, in Figure 

'db 

(peak A) 
V, 1 kVldiv 

1, 550 Ndiv 

l9 500 Aldiv 

I,, 500Ndiv 

Sweep: 5 psldiv Sweep: 5 pldiv 

Currents in branches Voltages at nodes 

(peak V) 

1600 769 470 360 330 

Figure 4 - Typical recordings made during the third experiment - capacitor and MO

Experiments 2 and 3 (Figures 3 and 4) do show decreasing levels of voltage surges at var

points of the Upside Down House, compared to the open-circuit conditions. However, (l

interactions between the capacitors of the generalor RLC wave-shaping network on the on

and the PC capacitors on the other hand, make detailed interpretation of the waveforms te

and beyond the scope of our illustrative examples. Numerical modeling with a current so

as reported in the following section, avoids this interaction and provides further evidence 

"PCs Galut~" effc~t. FUI 1t;adel~s interested in thc dctails, Tablc 1 documents thc rcsults 

concerning peak voltages noted from the oscillograms kept on file but not reproduced her

Readers interested primarily in the big picture may skip a close examination of this table. 

Table 1 - Summary of node voltages for component combinations 

Dctailcd cxpcrimcnt Device 

description number Combination 

9 m 36 m 

1-1 3.1 - MOVs onlv MOV 

Capacitor 

3.3 - Capacitor and MOV + Capacitor F MOV 

I 3.2 - Ca~acitors onlv I r&Ec I 

Ringing at the open-ended line is the cause of this voltage being greater than Vo 
Our previous experience with modeling cascaded SPDs [Lai et al., 19931 and the behavio

Upside Down House with installed SPDs [Martzloff et al., 19951 can be readily applied t

prediction of the behavior of the three circuits subjected to the experimental measuremen

an imposed current source, free from the unavoidable interactions that occurred in the pre

experimental measurements. This approach yields the best of the two methods: a compu

that has been well demonstrated as suitable for modeling nonlinear SPDs, and the freedo

Impose any fixed waveform - what reality Imposes on a res~dence is not the surge from 

impedance-limited surge generator - combined with the possibility to model many bran

circuits and many combinations of SPDs andlor capacitors. 

4.1 Modeling combinations of loads and branch circuit lengths 

In a series of modeling runs similar to the combinations of the experimental measuremen

model used the circuit of Figure 5. A current source feeds a fixed current surgc via a com

service drop to the panel bus of the service entrance, where three branch circuits made of 

diameter conductors (#I 2 AWG) take off, with length of respectively 5 m, 10 m, and 20 

choice of these lengths was based on curiosity about the effect of the long distance conne

(and thus an effect that might bc delayed 01 degraded) to the irnmediate occurrencc of an 

overvoltage at the end oS shorter branch circuits. The currents in the circuit are identified 

Figu~e 5 respectively as I,, I, I,, I,,, and I,,, wth the node voltages at the servicc bu\ and h

circuit ends respectively as V, V, V,, and V,,. 

FlgUre 5 - Service drop and branch circuits for modeling combinations 

As in our previous modeling applications, the EMTP program [EPRI, 19891 was used. T

computational artifacts, a finite 10 kQ resistance was always postulated at the ends of all 

circuits, whether these were left in "open" or "loaded" condition. The loads that were mo

included combinations of SPDs, capacitors, and 100 Q resistors. The SPDs were all 20-m

diameter, 130-V rated MOVs. The capacitors included a capacitance of 440 pF and a ser

resistance ("ESR") of 0.25 0 to represent an electrolytic capacitor. Initial conditions for 

models stipulated an initial charge on the capacitor to represent the normal condition of th

link. The 100 Q resistor was selected as a typical value for a 150-W power-consuming ap

connected at the end of the branch circuit. 
g

Compared with the 100 0 load at the end of the branch circuits, the series impedance ha

small effect and the impinging surge current I, divides almost equally (3000 A I4 = 750 A

nearly identical four branches I,, I,, I,,, and I,, (I, being the current in the service entranc

'branch', with a length of 0). Thus, the voltage developed by this 750 A current across 

would simply be expected to be 75 000 V (seventy five thousand volts). We say "would 

expected" because, of course, the insulation level of a real-world low-voltage insulation 

withstand such a voltage. The result of this theoretical case is another illustration of the 

"More Begets Less," according to which a high-amplitude, steep-front surge cannot prop

in branch circuits because a flashover will occur at the origin [Mansoor et al., 19981. 

Currents in branches: Time in milliseconds, peaks in amperes Voltages at nodes: Tlme in milliseconds, peak

Figure 6 - Currents and voltages for the baseline case of 100 R loads without SPD

Figure 7 shows the example of two PCs, onc at thc cnd of thc 5-m branch circuit, the oth

the end of the 10-m branch circuit: the surge current is shared (unequally) between the tw

capacitors, with a corresponding decrease of the voltage level at each DC link, and also a 

reduction of the voltage at the open end (unprotected) of the 20-m branch circuit. This e

shows the beginning of the "PCs Galore !" effect. Several intermediate and further steps 

modeled by adding combinations of surge-mitigating devices to build our case file and co

the expected effects, but we will spare the reader from a tedious recitation. 

The results shown in Figure 6 call for several comments: 

The unequal division of the currents I, and I,, reflects the effect of the larger impedan

10-m branch circuit, compared to the 5-m branch circuit. 

Substantial mitigation is obtained at the nodes V, and V,,, as a result of the filter-like 

of the line inductance and the capacitance of the PC power supply. 

Thc voltngc surgc dcvclopcd at thc nodc Vo propagates, unabatcd, to the end of thc op

circuited 20-m branch circuit, a reminder that the notion of voltage surges being atten

as they propagate in building wiring [IEC Report 664, 19801 was incorrect. 
Currents in branches: Time in milliseconds, peaks in amperes Voltages at nodes: Time in milliseconds, pea

Figure 7 - Currents and voltages for the case of two PCs without SPDs 

The voltage results are especially worth noting, in the context of what a power quality m

would report in such an installation, depending upon its point of connection. 

From the (impossible) high voltages of the 100 Q baseline (no SPDs), the voltage th

be recorded by a monitor installed at the service entrance is now reduced to "only" 1

A voltage-only monitor installed at the point of use of the power (typical selection o

~nstallat~on In many surveys), namely the two reccptacles at 5 m anti 1 U m fcedlng th

would report respectively 5 10 V and 290 V, creating the illusion Cfallacy) that there 

significant surge activity at these points. 

In reality a current surge I, of nearly 2000 A is carried by the PC rectifier into the cap

via a line fuse. Such a high current - undetected by a voltage-only monitor - coul

wcll bc fatal for thc rcctificr or thc input fusc, or for thc capacitor of thc PC powcr p

failures were found in post-mortems of equipment recently performed at PEAC. Th

scenario is also what occurred in the experimental test of Section 3 for a 4 kV open-

voltage and capacitor-MOV combination (shown in next-to-last row of Table 1). 

4.2 Modeling PCs Galore 

As a grand finale for illustrating our major theme, the case of an increasing number of P

modeled, with each of the PCs connected at the end of a dedicated branch circuit, with a

circuits having the same length of 20 m, still with the 10-m long service drop driving a 3 

current surge into the installation. 

Table 2 shows the resulting voltages at the service entrance and at the point of connectio

PC>, as well as 111e i~npusecl irr~yir~gi~~g culreIil ad Ll~e resullii~g cu1-rerrls in the individua

ports of the PCs. Indeed, the effect is linear with increasing numbers of PCs, and the re

decrease in the voltages expected to be reported by a power quality monitor installed at 

of use is quite apparent. 
Number Impinging current 

of I ('1 

Inspection of thc tablc shows that thc cffcct is practically proportional to thc numbcr of P

use in the installation. While we have used the short acronym of PC in the title and prec

discussions, the proliferation that we observe is not limited to PCs, but includes many el

appliances, such as home entertainment, heating and air conditioning with adjustable spe

that use a DC link with large energy-storage capacitor. 

Current at PC port 

(A) 

5. Action Items 

The customary closure of a paper is to list conclusions. However, in this case we sugges

items for industry, rather than academic conclusions: 

Voltage at service entrance 

('4 

Thc cxamplcs givcn hcrc clcarly show that thc fallacy of littlc surge activity can bc c

by limiting power quality measurements to voltage surges. With undetected current 

activity, users of electronic appliances will be puzzled by unexplained failures in the 

reporls of little voltage surge activity. 

We offer the explanation that these unexplained failures are likely to be associated w

the (heretofore not characterized) surge current delivery capability of the environmen

Such failures could have been avoided, had designers been better informed. 

With our leading theme now solidly established, questions arise about what to do ab

This matter requires the dedicated attention of both manufacturers and users of powe

monitoring instruments. 

Unfortunately, more than three years after beginning the crusade to overcome the fal

limiting surge measurements to voltage surges [Martzloff, 19951, there is no commer

available power quality monitor capable of characterizing the energy-delivery capabi

a surge event. 

Worse yet, among the international delegates to an IEC working group chartered to d

standard on measurement of power quality parameters (which is likely to dominate th

of future monitoring instruments), there are some claims being made that "a current 

not a power quality parameter" and therefore it should not be included in the emergin

standard prescribing measurements methods. This misconception needs to be correc

Therefore, the crusade must be pursued with perseverance, and it is the goal and hop

authors that sufficient recognition of the fallacy will eventually create a market dema

appropriate instruments, which forward-looking manufacturers may have developed 

be in the process of developing in anticipation of such recognition. 

Voltage at PC po

(V) 
ANSI/IEEE C62.4 1- 1995, IhEh Kecommended Practice on Surge Voltages in Low- Voltage AC 

Circuits. 

Dorr, D., "Point of Utilization Power Quality Results," IEEE Transactions IA-31, No.4, JulyIAu

1995. 

EPRI, "Electromagnetic Transient Program (EMTP)," Version 2.0; Volume 1 : Main Program: 

Volume 2: Auxiliary Routines, EPRI Report EL-6421-L, July 1989. 

IEC Report 664-1980 (sic) Insulation coordination within low-voltage systems including cleara

creepage distances for equipment, 1980. 

IEC 64/1034/CD, General basic information regarding surge overvoltages and surge protection 

voltage a.c. power circuits, Committee Draft, 1998. 

Key, T.S. and Martzloff, F.D., "Surging the Upside-Down House: Looking into Upsetting Refer

Voltages," Proceedings, PQAf94 Conference, 1994. 

Lai, J.S. and Martzloff, F.D., "Coordinating Cascaded Surge Protection Devices: High-Low ver

High,"IEEE Transacrions, [AS-29 No.4, July/August 1993. 

Lindes, G., Mansoor, A., Martzloff, F.D., and Vannoy, D., "Surge Recordings that Make Sense 

Deposition: Yes ! - 'Joule Content': Never !" Proceedings, PQA'47 Conference, 1997 

Mansoor, A., Martzloff, F.D., and Nastasi, D, "Applying Reality Checks to Standards on the Su

Environment," Proceerlmgs, 23rd lnternatiot~al Cmlference on L~glzttlitzg Protection, I;lorence, 

Septemh~r 1996 

Mansoor, A. and Martzloff, F.D., "Driving High Surge Currents into Long Cables: More Begets 

IEEE Transactions PWRD-12, July 1997. 

Mansoor, A. and Martzloff, F.D., "The Effect of Neutral Earthing Practices on Lightning Curren

Dispersion," IEEE Transactions PWRD-13, July 1998. 

Martzloff. F.D. and Gauper. H.A.. "Surge and High-Frequency Propagation in Tndrlstrial Power 

IEEE Transactions IA-22, JulyIAugust 1986. 

Martzloff, F.D. and Leedy. T.F., "Selecting Varistor Clamping Voltage: Lower Is Not Better!" 

Proceedings, 1989 Zurich International EMC Symposium, 1989. 

Martzloff, F.D. and Lai, J.S., "Cascading Surge-Protective Devices: Coordination versus the IEC 

Staircase," Proceedings, PQA 91 Conference, 1991. 

Martzloff, F.D., "Keeping up with the Reality of Today's Surge Environment," Conference Proc

Power Quality Solutions, Intertech Pub., Ventura CA, USA, September 1995. 

Martzloff, F.D., Mansoor, A., Phipps, K., and Grady, W.M., "Surging thc Upside-Down IIouse: 

Measurements and Modeling Results," Proceedings, PQA'95 Conference, EPRI, 1995. 

Martzloff, F.D., "Surge Recordings that Make Sense: Shifting Focus from Voltage to Current 

Measurements," Proceedings EMC'96 ROMA Symposium, Rome, September 1996. 

UL Std. 1449, Transient Voltage Surge Suppressors, Second Edition, 1996. 
ON THE

DISPERSION

OF

LIGHTNING

CURRENT

AFTER A

DIRECT FLASH

TO A BUILDING
On the Dispersion of Lightning Current

After a Direct Flash to a Building

François Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

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 of

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
25th International Conference on Lightning Protection 5.10 

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 possible scenarios 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 (power 

or communications), electrically-conducting mechanical 

elements, metallic structures, etc., has a de facto lightning 

'protection' system of intended or unintended air terminals 

and down-conductors - except that their connection to the 

earthing system might have unpredictable and unwanted 

side effects. 

As an input toward developing consensus, this paper 

reviews in a first part two examples of simulations that have 

been performed by others, and in a second part relates a 

real-world anecdote of a corroborated case of a direct flash 

to a residence. 

2. SIMULATING DISPERSION 

2.1 Examples of simulations 

A Joint Working Group of the International Electrotechnical Commission (IEC) has recently developed a 

Technical Repad on surge pmntp,ctian[2]. This &-year 

effort involved the participation of five ZEC Technical 

Committees interested in the subject. The data base 

considered by the group included, among many sources, 

two published papers, identified in the Bibliography of the 

report, authored independently by members of the group. 

In both studies, a 10f350 ps waveform was postulated, 

and a time-invariant earthing resistance and inductance 

were postulated. Currents in the available paths to earth 

and voltages at selected points of the systems were 

computed. For the purpose of this paper, three figures only 

are reproduced here for a qualitative glimpse on the results. 

Figure 1, simplified from Ref [2], shows the nature of 

circuit components and configuration: two buildings and the 

distribution transformer linked by a cable in a linear 

arrangement. The point of strike is Building 1. The detailed 

numerical values, which are given in the referenced paper, 

are not significant for this comparison of the two studies. 

* Electricity Division, Electronics and Electrical Engineering Laboratory, Technology Administration, U.S. Department of Commerce 

Contributionsfrom the National Institute of Standards and Technology are not subject to U.S. Copyright. 

ICLP zooo 41 9 
Figure 1 - Circuit components and configuration used in the Hasse simulation (Ref [Z]) 

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 resistance and 

an inductance, not drawn in the figure, but modeled in the 

computation according to the 20-m length of each radial 

drop. The point of strike of the flash is the earthing system 

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 the building (IMa!nns) 

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 resulting 

currents, according to Mansoor (Ref [3]) 

The significance of citing these two independent 

studies, separated by an ocean, by two different languages 

used by the authors, and by two different simulation 

programs, is that quite compatible conclusions were 

reached after exchanging information, first across the table 

during IEC working group meetings, then later on, via 

intensive e-mail messages, as a working relationship 

blossomed in spite of logistics barriers. This important 

point will be elaborated further in Section 7 of this paper. 
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: 

.... 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. 

This engineer had bought a house from the previous 

owner who had lived many happy years there without any 

problem, so that our engineer made the (unwarranted) 

assumption that the house and its electrical wiring were in 

good order. The house was surrounded by several tall, 

mature trees so he thought that the cone of protection from 

the trees would benefit the house. Alas ! All-knowing Zeus 

recognized that this engineer needed to be taught a lesson 

on reality and thus sent a downward stepped leader toward 

the general area of the engineer's house ... 

By now, dear readers, you have guessed that our 

mythical engineer is none but the author of this paper ... who will now offer this true story for your edificution. 

First, the "where" : Figure 4 shows a simplified (here we 

go again ...) topology of the house, a two-story woodji-ame 

with basement and attic. Utilities (power, telephone, and 

cable TV), all entered, via overhead service drops, at the 

rear of the house, while water and sewer underground 

pipes were at the front of the house. The telephone system 

was not involved in the incident and therefore is not shown 

in the figure. The power installation included the usual 

revenue-meter (outside) and service panel (inside) with 

circuit breakers controlling a multitude of branch circuits. 

Only three are shown in the figure: lightingfixture in the 

attic, TV on the secondfloor, and a counter-top receptacle 

(via ground-fault interrupter) for the kettle (@in Figure 4) 

sitting on the enamelled cast-iron kitchen sink. (The 

significance of this detail will sulface shortly.) 

-Firm bonds - Dubious bonds 

Figure 4 - Simplified configuration anecdotal 
Water service and indoor piping were all copper, with 

a bond between the ground bus of the service panel and the 

nearest cold-water pipe. There was no visible grounding 

conductor to a (non-existent) made-electrode, but the 

perception existed that, given the vintage of the house 

(circa 1920), the water piping was su.cient, in addition to 

the multiple-grounded neutral of the power company. 

Upon moving in the house, I had installed in the 

service panel a surge arrester (circa 1965 vintagejconsisting of a silicon-carbide disk in series with a 

stamped-metalair gap. The cable TV service, as origlnalJylnalJy 

installed by the utility, only had a 50-cm long picket 

allegedly serving as "ground". Having been exposed to 

the concept of bonding, I had installed a bond between the 

picket and the nearby outdoor waterfaucet. A very passive 

- but soon to become active -part of the installation was 

the typical sewer system made of lead-bonded cast-iron 

pipe extending a to vent through the roof and connected to 

the street sewer, still with cast-iron pipes and thus offering 

the topology of a well-grounded air terminal, albeit below 

the peak of the roo$ But I am getting ahead of myselj as 

I had never considered this vent as a lightning air terminal, 

since the house was surrounded by taller trees and thus 

"obviously" within their cone of protection. 

Now for the "When " and "How" : On the day when 

the tale unfolded, my wi$e (the corroborating eye-witness) 

and I were standing in the kitchen, listening to the 

approaching thunder and watching the big drops of rain 

just beginning to splash on the window. Then, a bright 

flash outside, with an immediate, deafening thunderclap, 

and also we both saw a smallflash under the kettle. "That 

was a close one " we both said, whereupon I proceeded to 

check all appliances in the house. Several were 

inoperative, but a check of their branch circuit breakers 

revealed that they had tripped, and resetting them restored 

order. The only one that did not work was the old TV set, 

although there was no evidence of severe damage or 

burned smell, and we considered ourselves lucky - until 

a smell from the attic attracted my attention: the ceiling of 

the attic (which was covered by cellulose-base panels) was 

smoldering ! 

Fortunately - and not by accident - a handy fire 

extinguisher allowed me to quench the smoldering, while 

my wife called the Fre department. To their credit, they 

were infront of the house within minutes. I told them that 

I believed that the $re extinguisher had done the job; 

nevertheless, one fireman proceeded to climb on a ladder 

to the am'c window and hacked it away to Let the smoke 

out, while another entered the house, pushing me aside. 

with a high power water hose in tow - which fortunately 

he did not turn on. Afrer ripping several of the ceiling 

panels to verify that the fire indeed was out, the firemen 

lejf, with our emotional- thanks and^ the applause of the 

neighbours gathered in front of the house. 

4.3 The homeowner's epilogue 

One obsolescent TV receiver, which was not repaired, 

but catalysed the purchase of a new and upgraded set 

(missing the opportunity to do an extensive postmortem as in the "Case of the Cozy Cabin" [4]). 

Several hundred dollars expended to repair the window 

destroyed by the firemen, install a splice on one attic 

rafter weakened by charring, and- replacing the ripped 

panels. 

After raovering- fhmthe shock, a- realization of how 

lucky to have been in tbe house at th~ time of the 

incident, and glad for the foresight of having a fire 

extinguisher on every floor of the house ! 

4.4 The engineer's epilogue 

Such a traumatic experience called for an investigation 

of the incident. The first observation was that the previous 

owner had installed insulation between the attic rafters, 

stapling the aluminum foil of the bats to the rafters, but not 

overlapping them across the edges of the rafters. This 

arrangement, concealed by the panels, created several gaps 

along the 5-m distance separating the sewer vent pipe from 

the light fixture at the apex of the attic, but reducing the 

total gap to a few centimetres - an easy side-flash 

scenario, resulting in the ignition of the dust and surface 

fizz ofthe rough-fiom-sawrnill~rafters. 

The second observation, a few days after the incident, 

was to notice a small rust spot on the kitchen sink where the 

kettle usually sat: there was a small hole in the otherwise 

good-condition glaze, exposing the underlying cast iron: 

The flash seen under the kettle (0) was the cause of the 

enamel puncture; several kilovolts must have been required 

to break down the series-connected insulation of the heating 

element inside the kettle, and the porcelain glaze of the 

sink. The electronic ground-fault circuit breaker controlling 

the receptacle had to be reset, but it was not damaged, and 

subsequent use of the kettle did not cause it to trip, so we 

concluded that the brief breakdown of the insulation of the 

heating element was not a massive event. 

The immediate action was to install a bond between the 

sewer vent pipe and all extraneous metal in the attic. The 

long-term effect on the engineer was a consciousnessraising on the issue of surge protection of multiple-port 

appliances, even though a bond had been provided between 

the incoming cable TV service and the power system [5]. At 

that time, the concept of the surge-reference equalizer [6] 

had not yet surfaced, and no commercial device was 

available to provide that function. In fact, the proliferation 

of plug-in surge-protective devices launched by the 

introduction of metal-oxide varistors had not yet occurred. 

A casual inspection of the fist-sized surge arrester at 

the service panel showed no distress, an indication of 

adequate design for the rare scenario of a direct strike to a 

building [7]. 
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 frrJm 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 current 

dispersion illusPated in Section 2 have met with some 

skepticism among the North-American surge-protection 

community and perhaps others. Part of this skepticism is 

also based on the relatively rare occurrence of massive 

failures for secondary arresters (distribution transformer 

secondary terminals and residential service entrance) 

designed to withstand the "classical" 8/20 ps or 4/10 ps 

surges, at crest levels of a few to a few tens of kiloamperes. 

Furthermore, the two simulations cited in Section 2 were 

based on the assumption that earthing electrodes have a 

constant resistance during the flow of the lightning current, 

an assumption that is questioned on the basis of preliminary 

results of measurements made in Florida in connection with 

triggered lightning experiments [8]. 

In contrast with these simplified scenarios, the realword anecdote would be a challenge for any numerical 

modelling but demonstrates evidence of substantial 

overvoltages developed in the* installation (insulation 

puncture at the kettle) during the flow of this undetermined 

lightning current dispersion among the complex available 

paths to earth. The anecdote also offers an example of a 

surge arrester with modest current-handling capability 

surviving the scenario of a direct strike to a building. 

A symptom of the incomplete consensus is the 

noticeable lack of a discussion of risk analysis in the report 

developed by the IEC Joint Working Group [2]. This topic 

was initially included in the document outline, raising 

high expectations, but, confronted with incompatible 

proposals, the group gave up on that initiative. The 

proposed methodologies ranged from elaborate and detailed 

mathematical formulae - which turned out to be using 

somewhat arbitrary postulates - to common-sense, almost 

intuitive considerations. 

7. A PROPOSAL FOR THE DISPERSION 

OF LIGHTNING INFORMATION 

In a 1963 freedom-seeking speech that still resonates 

today, the mantra "l'hve a dream" was coined. On a much 

more modest scale, the author has a dream of unfettered 

information-sharing on lightning. Having cited the 

preceding examples of developing, but still incomplete. 

consensus on the dispersion of lightning current, here is the 

proposal (or is it a challenge ?): Hopefully helpful timely 

participation, on a world-wide basis via electronic mail 

could supplement - not compete with - the established 

routes for information sharing, at a much accelerated pace. 

We are still mostly in a mode of developing standards - 

a notoriously slow process - by volunteers or delegates 

often hampered by travel budgets, or of publishing peerreviewed papers - unquestionably a wise process, but 

entailing long delays between generation and ultimate 

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 

General 

Many publishing organizations, in their instructions for 

the peer-iehew proc&s,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, citatiom - in the form 

of a "Bibliography" with or without annotations. 

For this paper, "References" are limited to the strict 

minimum necessary to support a particular point being 

made. To illustrate where extensive listing of "references" 

might lead, a likrature search was conducted with 

"lightning" as a leading key word, and next with one 

additional word. The results are listed below, showing the 

number of "hits" found for the period of just 1969-1999 - the accessible on-line data base did not include Benjamin 

Franklin's seminal letters to the Royal Society on lightning 

protection of houses and the Purfleet munitions storage [9]. 

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 unrealistic for a 

single researcher to examine in detail the contents of fifteen 

thousand papers. Injecting the concerted filtering and 

sharing action of today's active researchers into a readily 

accessible data base - the author's dream - would be a 

great improvement. 

Specific references 

Anderson, R.B. and Eriksson, A.J.. "Lightning parameters 

for engineering applications." ELECTRA, No.69, March 

1980. 

IEC 62066: 2000, General basic information regarding 

surge overvoltages and surge protection in low-voltage 

ax. power circuits. 

Mansoor, A. and Martzloff, ED., 'The Effect of Neutral 

Earthing Practices on Lightning Current Dispersion in a 

Low-Voltage Installation," IEEE Transactions PWRD13; 

July 1998. 

Key, T.S., Nastasi, D., Phipps, K.O., and Martzloff, F.D., 

"Some enlightening case histories on lightning damage." 

Submitted for presentation at ICLP 2000. September 2000. 

Key, T.S., and Martzloff. F.D., "Surging the Upside-Down 

House: Looking into Upsetting Reference Voltages," 

Proceedings, PQA '94 Conference, Amsterdam, October 

1995. 

IEEE Std I 100-1999 ("Emerald Book") IEEE 

Recommended Practice for Powering and Grounding 

Electronic Equipment. IEEE, 1999. 

IEEE PC62.41.1-2000 - Drafi IEEE Guide on the Surge 

Environment in Low-Voltage (I000 Vand less) AC Power 

Circuits. 

EPRI "Investiga,tion of lightning entry into a secondary 

service using rocket-triggered lightning," EPRI Report 

TR-110418, April 1998. 

Cavendish. Watson. Franklin and Robertson, "A report 

of the Committee appointed by the Royal Society 

to consider a Method jor securing the Powder 

Magazines at Purjleet," Philosophical Transactions 

of the Royal Society, London, 1772. 
ROLE AND

STRESS OF

SURGEPROTECTIVE

DEVICES

IN SHARING

LIGHTNING

CURRENT
The role and stress of surge-protective devices

 in sharing lightning current

François D. Martzloff

National Institute of Standards and Technology

Gaithersburg MD 20899

f.martzloff@ieee.org

Arshad Mansoor

EPRI PEAC Corp

942 Corridor Park Blvd

Knoxville TN 37932

Amansoor@epri-peac.com

Preprint of paper to be presented at EMC Europe 2002, September 2002

Significance

Part 2 Development of standards – Reality checks

Part 4 Propagation and coupling – Numerical simulations

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 of a building and their involvement in that part of the lightning current that exits

the building wia the 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 occupant; the results show that a significant part of the exiting lightning current

can involve those downstream SPDs with some 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 occurence of reported failures.
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 considering 

surge stresses on equipment higher than those currently 

described in the IEC document 61000-4-5 on immunity 

testing [8]. 

The purpose of the paper is to examine in detail the 

sharing of lightning current, not just by the SPDs at the 

service entrance, but also by all SPDs that might be 

involved in the exit path of the lightning current. Such 

sharing might well involve SPDs incorporated in the 

equipment located close to the service entrance, but not 

always included in the numerical simulations that have 

been performed by many researchers (Altmaier et al., 

1992) [1]; (Standler, 1992) [23]; (Rakotomalala, 1994) 

[20]; (Birkl et al., 1996) [3]; (Mansoor and Martzloff, 

1998) [15]; (Mata et al., 2002) [19]. In its recent 

development of a Guide and a Recommended Practice 

on surges in low-voltage ac power circuits [13] the 

IEEE has refrained from identifying SPDs as being 

those that may be connected at the service entrance. 

Instead, it refers to "SPDs involved in the exit path" 

without reference to their point of installation. 

Given the tendency of equipment manufacturers to 

include an SPD at the equipment power input port, the 

issue of "cascade coordination" arises. Several previous 

papers (Martzloff, 1980) [17]; (Goedde et al., 1990) 

[5]; (Lai and Martzloff, 1991) [14]; (Standler, 1991) 

[22]; (Hostfet et al., 1992) [7]; (Hasse et al., 1994) [6] 

have explored the concept of cascade coordination 

involving two or more SPDs connected on the same 

power supply but at some distance from each other. 

The legitimate wish of the energy service providers to 

specify robust SPDs at the service entrance results in 

SPDs having a relatively high Maximum Continuous 

Operating Voltage (MCOV). On the other hand, some 

equipment manufacturers tend to select SPDs with a 

low MCOV under the misconception that lower is better 

(Martzloff and Leedy, 1989) [18]. This dichotomy can 

result in a situation where the low-MCOV SPDs 

included in equipment might well become involved in 

the "exit path" and thus become overstressed in the case 

of a direct flash to the building. This situation is made 

more complicated by the fact that commercial SPDs 

packages are assembled from typical distributors' supplies that can have an allowable tolerance band of 

±10% on the voltage-limiting rating. 

To explore the possibility and implications of a 

questionable coordination, numerical simulations were 

performed on a simplified model of a building featuring 

SPDs installed at the service entrance and SPDs that 

may be incorporated in equipment connected inside the 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 and 

Martzloff, 1998) [15] have validated the intuitive 

expectation that the tail of the 10/350 μs waveform 

often postulated for simulations will be shared among 

the available paths simply according to the relative 

values of resistance in the paths leading to the earthing 

electrodes. This fact is apparent in the results of Figure 

2, for example at the 350 μs time: when inductive 

effects have dwindled, the current IH in the 10-Ω

earthing resistance of the building is ten times smaller 

than the total current exiting the building [IN+IL1+IL2] 

toward the power distribution system in which multiple 

earthing electrodes offer an effective earthing resistance 

of only 1 Ω. It is also worthy to note that this sharing is 

controlled by the relative values of the resistances, so 

that any earth conductivity differences associated with 

local conditions will wash out. 

The combination of the service-entrance 150-V MOV 

on Line 2 and the 130-V MOV incorporated at the 

power port of the equipment constitutes a so-called 

"cascade". When two such cascaded SPDs are to be 

coordinated, a decoupling impedance must be provided 

between the two SPDs so that the voltage drop caused 

by the current flowing in the decoupling impedance – in 

this example the impedance of the 2,5 mm2

 diameter 

wires – and added to the limiting voltage of the 130-V 

MOV, will cause enough of the current to flow through 

the 150-V MOV to reduce stress on the 130-V MOV. 

The simulation was performed for three values of the 

impedance (length) of the connection, i.e., 0,1 m, 1 m, 

and 10 m to assess the effect of this impedance for 

practical situations. Figure 3 shows the results for 

these three cases and Table 1 shows the resulting energy 

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 the 

effectiveness of a cascade for the case of direct flash to 

the building. In an actual installation, there would be 

more than one piece of equipment, presumably each 

with a 130-V built-in MOV at the power port. One 

might expect that some sharing among these multiple 

SPDs would reduce the energy stress imposed on these 

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 in 

a situation where the difference between the two SPDs 

of the cascade is not large, because of the de facto 

situation of low values of MCOV that the industry has 

unfortunately adopted. Of course, if tolerances were 

also taken into consideration for the service entrance 

MOV, the extremes of distributions for both MOV 

would make an effective coordination between a 

nominal 150-V MOV and a nominal 130-V MOV even 

more problematic. 

 

II. 3 Nonlinearity of circuit elements 

 

Most of the reported simulations, as cited above, have 

been performed with a conservative postulate of a 100 

kA 10/350 lightning discharge. The median of the 

current peaks compiled in the seminal Berger et al. 

paper [2] is only 20 kA. Occasional reservations have 

been voiced on the validity of these data collected with 

technology dating back to the 1970’s. A recent (July 

2000) actual case history was communicated to the 

authors by a colleague for two major lightning storms 

recorded in the area of Tampa in Florida by means of 

the Lightning Detection System [24], during which over 

30 000 flashes were detected in a period of less than 12 

hours, with only one at the 150 kA level, and a median 

of 20 kA, confirming the Berger at al. data. 

One could expect that the dispersion of the lightning 

current that results from the combined action of linear 

elements (resistance and inductance) with nonlinear 

components (MOVs) might produce a different sharing 

of the current as the decoupling element is linear but the 

SPDs are nonlinear. To explore this hypothesis, the 

computations for the case of Figure 4 and Table 2 were 

repeated, for peak currents of 100 kA (the original value 

of the computation), 50 kA, and 25 kA (about the 

median of the statistics). Table 4 shows the results of 

these computations. It is interesting to note that as the 

applied stroke is decreased 4 to 1 (from 100 to 25), the 

total energy deposited in the varistors is decreased by a 

factor of 3200/610 = 5.2. This relative greater decrease 

is caused by the larger portion of the current exiting via 

the linear-path neutral, further relief for all the SPDs 

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 reference 

parameter of a direct lightning flash to a building should 

be a 10/350 μs current with a peak of 100 kA, the 

numerical simulations performed for a simplified 

system with one surge- protective device installed at the 

service entrance, and one or more built-in SPD in 

downstream equipment indicate that the downstream 

SPD is very likely to be overstressed and fail, most 

likely catastrophically. 

2. There are several possible explanations for the 

apparent contradiction between a prediction of downstream equipment failures based on this postulated 

lightning parameters, and equipment field experience 

that does not report such frequent failures, although of 

course anecdotes abound. 

• The occurrence of a direct flash to a building can 

cause such extensive damage that a post-mortem for 

investigating the specifics of a prevailing ineffective 

coordination is not performed at that time and the 

issue is ignored. 

• Enough uncontrolled clearance flashovers occur in 

the installation to provide significant relief for any atrisk SPDs incorporated in downstream equipment. 

• In an installation where many built-in or plug-in 

SPDs are present, the sharing illustrated by Figure 4, 

combined with a low probability of a 100 kA stroke, 

might reduce the stress on downstream devices to a 

value within their capability. In particular, many 

commercial plug-in SPDs advertise capabilities of 

hundreds of joules, unlike the 20 joules of a single 

MOV, which might be provided at the input port of 

electronic equipment. 

• Insufficient field failure data have been obtained, 

compiled, shared, and published to enable realistic 

assessment of frequency and severity of occurrences 

involving an unsuccessful cascade coordination. 

 

3. It is impractical at this point to mandate high energy 

handling capability for built-in SPDs. Such a move 

might meet with strong objections from manufacturers 

whose products have satisfactory field experience, and a 

risk analysis might show it to be not cost-effective. 

4. Economic and political realities related to the type 

and mission of the installations to be protected should 

be kept in mind. Clearly, mass-market applications 

such as cost-conscious consumers, in a framework of 

regulated or unregulated installations, are different from 

bottom-line-conscious industrial applications, and even 

more so in the case of national assets – be they cultural 

or military. 

5. Another approach for manufacturers might be to 

avoid placing low MCOV varistors at the input port of 

their equipment. Rather, they should select an SPD 

with an MCOV and resulting surge-protective level as 

high as their equipment can inherently stand. This is a 

“selfish” approach which is mentioned here halfseriously, half-facetiously: there are enough low MCOV 

SPDs installed by users or included in other equipment 

in a typical system that those unfortunate low-MCOV 

devices will take up the stress, leaving unscathed the 

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 PWRD13, 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,” IEEE 

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. 
François Martzloff 

END OF FILE “ALL PAPERS PART 2” 

April 2003