François D. Martzloff 

General Electric Company 

Schenectady NY 

f.martzloff@ieee.org 

Reprinted, with permission, from IEEE Transactions on Power Apparatus and Systems, PAS-102, May 1983 

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

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

Paper Award, 1985, Surge-Protective Devices Committee

Significance: 

Part 4 – Propagation and coupling of surges 

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

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

the surge of wire runs is pointed out and illustrated. 

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

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

than the front time of the impulse. 

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

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

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

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

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

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

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

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

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

THE PROPAGATION AND ATTENUATION 

OF SURGE VOLTAGES AND SURGE CURRENTS 

IN LOW-VOLTAGE AC CIRCUITS 

F.D. Martzloff, Senior Member 

Corporate Research and Development 

General Electric Company Schenectady, 

Schenectady, New York 

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

difference between surge impedance (characteristic impedance) of 

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

pointed out and illustrated. A comparison is made between the 

propagation of a surge through an isolating transformer and a 

ferro-resonant line conditioner. The effect of connection options 

are shown for one or several surge protective devices connected at 

the end of a 3-wire line. 

INTRODUCTION 

Considerable progress has been made during the last decade 

toward recognizing the occurrence of surge voltages in low-voltage 

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

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

practical information is still scarce on the propagation of these 

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

through transformers. 

This paper provides concrete examples on the propagation of 

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

from which some conclusions can be drawn and sound practices 

recommended. 

The tests reported here have been performed with the voltage 

and current waveshapes recommended by the recently published 

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

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

an idealized transmission line, how complex systems transform 

pure standard waves, and how some connection practices for surge 

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

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

IEEE 587 ring wave. 

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

examples reported are given in order to alert the reader against 

some pitfalls resulting from occasional misconceptions observed by 

the author during discussions and reviews of many surge-related 

problems. 

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

IEEE Surge Protective Devices Comittee of the IEEE 

Power Engineering Society for presentation at the IEEE 

PES 1982 Summer Meeting, San Francisco, California, 

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

made available for printing May 17, 1982. 

TEST CIRCUITS AND TEST GENERATORS 

The test circuits selected to represent typical low-voltage ac circuits include runs of nonmetallic jacketed 600 V wire such as that 

found in residential wiring and runs of rigid conduit with wires 

pulled in the conduit, such as that found in industrial installations. 

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

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

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

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

impulse, and the 8/20 ps current impulse. The 0.5 ps - 100 kHz wave is produced by a KeyTek 424 surge generator, capable of 

superimposing the voltage pulse at a controllable time of the 60 Hz 

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

is produced by a Haefely P6R generator, capable of supplying up 

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

KeyTek 424 with a different plug-in unit (Figure 3) or by a laboratory circuit using storage capacitors and an ignitron switch. The 

initial test involving transmission line propagation is made with a 

narrow pulse obtained from a Velonex 350 pulse generator. 

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

loaded, produced by KeyTek 424 generator and 

PN281LSC plug-in 

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

Haefely P6R pulse generator 

IIIIIIII 

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

KeyTek 424 pulse generator with PN247 plug-in 

0018-95 10/83/0500-1163$01.00 @ 1983 IEEE 
The oscillograms were recorded with a Tektronix 7633 oscilloscope. Voltage measurements were made with two 1000:l 

Tektronix P6015 probes in differential configuration, so that the 

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

Current measurements were made with a 0.010 T&M Research 

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

100 AIV. 

The oscillograms are shown together with a schematic of the 

circuit configuration. The three wires of the various lines are 

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

convention on color. The impulse was applied between the point 

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

connection of the differential probes and the resultant recording 

are connected by the arrows. 

PROPAGATION OF VOLTAGE SURGES 

Transmission line behavior 

To establish the baseline of the propagation characteristics, a 

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

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

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

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

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

computed. As a side experiment, the classical nonreflection 

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

surge impedance* is also observed for a terminating resistance of 

100 a. The slight mismatch indicated by the small reflection 

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

attributable to the connection of the line with closely spaced wires 

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

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

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

matching impedance is between these two values. 

NON-METALLIC SHEATHED CABLE 

XI2 AWG WITH GROUND. 

Fignre 4. Transmission line behavior of non-metallic, 

plastic jacket wire 

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

but too often not recognized in qualitative discussions of 

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

classical transmission line concepts to wiring systems if the front 

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

For a 1.2150 ps impulse, this means that the line must be at least 

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

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

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

voltages at the crest values of the impulses. 

Short lines behavior 

The response of lines shorter than the 200 m limit identified 

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

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

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

A difference between surge impedance (also known as "characteristic 

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

tIIllllllI 

i 

Figure 5. Response to a 1.2/50 ps impulse of wiring in nonmetallic jacket 

(A) Output of unloaded generator and with 

line connected 

(B) Sending end and receiving end, 

open receiving end 

(C) Expanded trace of oscillogram (B) 

so- "n 

Figure 6. Response to a 1.2/50 FS impulse of wiring 

in metal conduit 

(A) Effect of grounding options at the sending end 

(B) Sending end and receiving end voltages 

In both Figures 5 and 6, the front of the wave (SD side) is. 

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

impedance loading the impulse generator. It is even more noticeable in Figure 6(A), which shows the voltages at the sending end 

for the connection with white, green, and conduit tied together. 

The greater capacitance of this configuration, compared to white 

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

the multiple reflections occurring at the mismatches produced by a 

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

jagged appearance of the oscillogram. 

The front of the impulse is further expanded in Figure S(C), 

showing the difference between the sending and receiving ends of 

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

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

reflections occurring during the rise time of the impulse. Likewise, the slight differences between the sending end and the 
receiving end occurring early in the rise do not affect the final 

voltage at the receiving end of the conduit line. 

These tests show that the propagation of voltage surges in 

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

not procwe appreciable attenuation of the surges, nor does it 

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

significant for the case of an appliance or industrial equipment with 

a control circuit drawing very little load while the power circuit 

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

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

of the line impedance and load impedance, but that beneficial 

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

standby condition of the control circuit, the very circuit containing 

the most sensitive electronic components of the appliance or 

equipment. 

Branched circuits 

Departing further from the simple and sanitary behavior of a 

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

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

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

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

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

open receiving ends goes through oscillations that only vaguely 

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

unidirectional impulse has vanished. 

Figure 7. Response to a 1.2150 ps impulse 

of a branched circuit 

This simple branch circuit behavior demonstrates why it would 

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

simulated with simple test waves.* 

This dichotomy between simulating nature and performing standard tests has 

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

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

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

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

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

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

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

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

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

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

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

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

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

those devices that do not survive the test wave. 

Does an isolating transformer help? 

The author has witnessed and engaged in many discussions on 

the merits of isolating power transformers, sparked by the misconception indicated by statements such as "spikes are attenuated by 

transformers" or "spikes do not pass through transformers." Figures 8 through 12 are offered to support the position that these 

quotations are misconceptions. When properly applied, isolating 

power transformers are useful to break ground loops, but they do 

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

normal mode. 

Figure 8 shows the propagation - or worse, the enhancement 

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

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

side of this "isolating" transformer. 

0 

W 

G 

1 kVA GENERAL PURPOSE 1201120 

TRANSFORMER 

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

through an isolating transformer 

Figure 9 shows similar behavior in a transformer offered as a 

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

isolation and low effective capacitance between primary and secondary windings, but here again, the author has observed that users 

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

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

the simple two-winding transformer of Figure 8, but we also note 

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

second half-cycle. Hardly an improvement. 

1 kVA LOW CAPACITANCE -!j "LINE ISOLATOR" TRANSFORMER 

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

through a "line isolator" transformer 

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

secondary, which represents the extreme case of a low-power electronic control in the standby mode. Figure 10 shows the primary 

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

and a 100 W (150 a) load on the secondary side, at the same generator setting as Figure 8. With the 10 W load that might be typical of an electronic control in standby mode, the combined series 

reactance of the transformer and shunt resistance of the load produce the output shown in Figure 10(A), still slightly higher than 

the input. 

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

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

produce a greater attenuation than resistive loads for the inductive 

series impedance of the transformer, at the frequency spectrum of 

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

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

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

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

electronics connected on the secondary side of the transformer. 

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

suppression. Figure 11 shows a 6 kV impinging ring wave 

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

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

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

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

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

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

side of the line conditioner. This oscillogram is, in fact, a 

photograph of two successive measurements, one with no load on 

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

are exactly superimposed. Compare this with the regulation of the 

generator output voltage noticeable in Figure 1, where a 100 

resistor is connected directly at the terminals of the generator. 

Figure 11. Decoupling of a 

by a ferro-reson 

0.5 

ant 

fis - 100 kHz ring wave 

line conditioner 

This decoupling reflects the nonlinear behavior of the ferroresonant line conditioner, which is significant in this case, compared to the linear behavior of transformers: for surge sources of 

lower impedance than the generator used in these tests, or for frequencies lower those than contained in the 0.5 ps - 100 kHz ring 

wave, the transformer attenuation would become lower, in direct 

proportion to the corresponding impedance change, while the 

ferro-resonant transformer would keep the decoupling unchanged. 

The two oscillograms of the output were recorded with the 

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

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

voltage is indicated by the gray band recorded on the oscillograms 

by photographically superimposing repetitive traces of the line 

voltage. For timings other than the peak, the small voltage oscillation on the output voltage would be completely contained within 

the normal peak-to-peak band of the 60 Hz line voltage. 

While these measurements were being taken, an additional 

observation was made. Figure 12 shows the response of the line 

conditioner to surges occurring at different times in the 60 Hz 

cycle, as indicated by the different vertical position of the traces at 

their beginning. The ferro-resonant mechanism is responsible for 

this different response. In itself, this is not very important in the 

present context, but it does provide another example of the 

importance of performing surge testing at different angles along 

the power-frequency cycle (as recommended in the discussions 

presented in IEEE Std 587) because the outcome of the test may 

be influenced by the timing of the surge. 

Figure 12. Effect 

u 

of timing of the surge with respect to the 

power frequency voltage 

PROPAGATION OF CURRENT SURGES 

Line impedance: surge impedance or impedance to the surge? 

As mentioned in the transmission line behavior, a distinction 

has to be made between two concepts that unfortunately can be 

confused because of the language: surge impedance of the 

transmission line, and impedance of the line to the surge. The 

first is the classical transmission line parameter, also called 

"characteristic impedance", Z, = a, and applies for long 

lines and short pulses. It is independent of the line length and frequency. The second, impedance to the surge, is indeed dependent 

on the line length, and is the impedance of the complex (real and 

imaginary) network of distributed parameters, R,L,C, of the wiring configuration. This impedance is also dependent on the frequency, so that rigorous analysis would involve computation over 

the frequency spectrum of the impulse of interest. For practical 

applications, it would be more convenient, although not rigorous, 

to define the impedance of a line to the surge as the ratio of voltage to current, stating the current wave form. 

Thus, inspection of Figure 13 shows a current crest of 400 A 

flowing in the line with shorted end and a voltage crest of 1700 V 

at the sending end, with a current wave form of 25/70 ps It is 

noteworthy that the short-circuit impulse of 8/20 ps produced by 

the generator has been stretched out by the effect of the line 

impedance. This impedance is mostly inductive, as shown by the 

fact that the crest of the voltage occurs during the initial current 

rise where di/dt is large, with a resistance detectable by a finite 

voltage at the time di/dt is zero - that is, at the crest of the 

current, not counting the capacitance. Thus, one might define the 

impedance to the surge of this 75 m line as being 1700 V/400 A 

for a 25/70ps wave, or 4.25 'ohms', a far cry from the 100 Cl 

characteristic impedance determined by the first measurement 

reported in this paper. This impedance is essentially proportional 

to the line length, in contrast to the constant value of the characteristic impedance. 
Figure 13. Current and voltage in 75 m line, 

unidirectional impulse 

The complex nature of the line configuration is also evident in 

the voltage observed between the shorting jumper at the receiving 

end of the line and the green ground wire: the voltage is not just 

half of the sending voltage but, rather, the superposition of that 

half-voltage and higher frequency components which are not seen 

in the sending-end voltage. 

Likewise, Figure 14 shows a first current crest of 48 A in the 

shorted 75 m line with a sending-end voltage of 5000 V when the 

0.5 ps - 100 kHz generator is driving the line. This corresponds 

to an impedance to the surge of 100 'ohms', not very different 

from the characteristic impedance, For the second crest, however, 

the current crest is 30 A with a voltage crest of 1300 V, or about 

45 'ohms' for the significant frequency of the second crest. 

Figure 14. Current and voltage in 75 m line, 

oscillatory wave 

The addition of the grounding wire (Figure 15) to the circuit of 

Figure 14 does not considerably change the crests of voltage and 

current but introduces the added complexity of a secondary oscillation superimposed on the driving oscillation. 

Figure 15. Effect of added grounding wire for 75 m line, 

oscillatory wave 

Figure 16, similar to Figure 13, shows the propagation of a 

current surge in conduit-enclosed lines. For the same generator 

short-circuit wave form of 8/20 ps, the resultant current and voltages are shown for a 25 m conduit run. The ratio of 

voltagelcurrent yields a value of 1500 V/850 A for the 20150 ps 

current wave form, or an impedance of 1.8 'ohms'. 

Figure 16. Current and voltage in 25 m conduit run, 

unidirectional impulse 

Figure 17 shows the impedance of the 75 m line as a function 

of frequency, as measured by an impedance vector meter. The 

values of impedance defined as approximations for impulses are 

also shown on this graph. 

Therefore, as a first approximation, a more useful view of the 

relative impedance values in a wiring system can be derived from 

this concept of "impedance to the surge" than from the use of 

characteristic impedance, provided that the user of this approximate concept does not lose sight of the approximations implied in 

the concept. 
Figure 17. Impedance versus frequency of 75 m line 

For instance, consider the case of a wiring system where a 

decreasing "staircase" of voltage surges is expected as the wiring 

progresses within the building, starting from the servlce entrance. 

Such a staircase is described in the IEC recommendations on Insulation Coordination [2]. The staircase is obtained by using the 

interface effect of the series and shunt impedai~ces of the wiring, 

including line impedance, transformer (if any) series impedance, 

and shunt impedance of connected loads. While the series 

impedances are likely to constant for a given system, the shunt 

impedances will vary with the loads. Alternatively, the decreasing 

voltages can be the result of installing surge protective devices at 

interfaces between sections of the wiring where the voltage 

decrease is to take place Such a coordinated wiring system will 

require careful consideration of the line impedances for the various ranges of surge-effective frequency to be expected, so that 

proper coordination can be ensured between the successive protective devices installed at the interfaces [81. 

The pitfalls of unsanitary wave forms 

While an impulse generator is essentially an energy storage element (capacitor, line or inductance) discharged into the test specimen through some wave-shaping network, producing a clean wave 

shape as described in standards specifications is not a trivial undertaking. Unless precautions are observed, the stray inductance of 

capacitors or the stray capacitance of inductors as well as the wiring impedances can introduce unwanted oscillations - an "unsanitary" wave form. 

Figure 18 gives an example of the problems that an unsanitary 

wave form can introduce. An attempt was made to apply a laboratory surge generator (built for energy deposition testing) to force a 

8/20 ps current into the 75 m line, since the 8/20 ~s short-circuit 

wave becomes stretched as discussed previously. The reasonably 

clean current wave form of Figure 18(A) would be quite acceptable as a test wave where total current, crest current, or energy are 

Figure 18. Effect of unsanitary current wave on line voltage 

the significant parameter. However, when applied for the purpose 

of evaluating line impedance by measuring and computing V/l, the 

small ripple occurring on the current rise produces the L(di/dr) 

oscillations seen in Figure 18(B). Thus, such an unsanitary wave 

form is totally useless for that purpose. 

The effect on residual surges 

of connections options for suppressors 

A noticeable lack of agreement has been observed among various application information sources on the most effective transient 

suppression configuration to be applied. Taking, as an example, 

the task of specifying the protection of an appliance or equipment 

connected at the end of a line with no opportunity to divert the 

transient closer to the source (for instance, at the service 

entrance), the options would be to connect one, two, or three 

surge suppressors between the three wires (black, white, and 

green) at the end of the line. However, more needs to be known: 

Will the impinging surge be in the normal mode (black to white) 

or in the common mode ([black-and-whitel-to-green)? Where in 

the equipment is the most sensitive component: line-to-line (most 

likely) or line (black OR white)-to-green? Clearly, the situation is 

confusing, and there will not be a single, simple answer applicable 

indiscriminately to all cases. The National Electrical Code [91 

specifically allows the connection of surge arresters (Article 280-22) 

if the interconnection occurs only by operation of the surge arrester during the surge. Since the standby current of a varistor or the 

leakage current of an avalanche diode suppressor is very low, the 

intent of this requirement can be met. Furthermore, there will 

not be any interference with the operation of Ground Fault Circuit 

Interrupters if the total number of suppressors does not result in a 

large current. 

The set of measurements recorded in Figure 19 shows an 

example of these many options with increasing protection, albeit at 

increasing cost, from a single suppressor to three suppressors. 

The selection would depend on the vulnerability level and location 

of the equipment to be protected. The impinging surge is 

assumed to be black-to-[white and green], since white and green 

are tied together at the service entrance. The line is the 75 m line 

previously investigated, and the surge is that available from the 

generator set for a 2000 A 8/20 ~s short-circuit impulse. Rather 

than attempt to modify the setting of the generator for each case 

in order to maintain constant current crest for the various 

configurations (an impossible task if wave form is also to be maintained), the generator was left unchanged, to discharge a constant 

total energy in the system - not a bad hypothesis for the real 

world. This test was performed with 20 mm diameter varistors 

rated for 130 V rms line voltage, as an example. Similar results 

would be obtained with other types of clamping suppressors. The 

point is not so much the clamping voltage measured, as it is the 

relative differences for the various options shown. The current 

crests are all in the range of 300 to 380 A, which is not a 

significant change for comparing clamping voltages. 

If only one suppressor is allocated to protect the equipment, 

the black-to-white suppressor connection affords maximum protection for the electronics which are also likely to be connected blackto-white. However, the voltages between either black or white 

and green are large; this is the stress that will be applied to the 

clearances of the equipment. This example shows a current surge, 

which might seem relevant only to surge suppressor applications, 

becomes a voltage surge issue, which is relevant to insulation 

coordination of clearances. 

The configuration with suppressor black-to-green does not 

afford very good protection for components connected black-towhite; therefore, it should be used only if there is a special need 

to clamp black-to-green at a low voltage. 

An improved protection is obtained with a suppressor black-towhite complemented by a second suppressor white-to-green. 

Another option, not investigated here but often used in applications of three-electrode gas tubes, would be the connection two 
CIRCUIT TOTAL CURRENT I V-BLACK-TO-WHITE I V-BLACK-TO-GREEN I V-WHITE-TO.GREEN 

tv 10us mv 8@QT lo# mv nKIdl 10~8 

380 A 360 V max 380 V max 100 V max 

/\ damaL *mL 4 ?GF_yLV?h'c#v, I 7 dk 8 zQ!!E T ~T~o&& 1 / 6 &Y- 1 FA b 

Figure 19. The effect on residual voltage surges of connections options for one, two, or three suppressors 

suppressors, one between a line and the grounding conductor, the 

other between the second line and the grounding conductor. The 

ultimate protection is, of course, one suppressor in every position, 

but this should be required only for exceptionally sensitive loads. 

CONCLUSIONS 

The examples of surge propagation described provide the basis 

for several practical conclusions that should provide guidance in 

designing or evaluating surge protection schemes. 

1. Surge propagation in wiring systems should be considered as a 

case of classical transmission lines only if the lines are long 

enough to contain the surge front. 

2. For typical voltage or current surges produced by lightning or 

switching, the surge impedance (characteristic impedance) is 

not the significant parameter. Rigorous analysis requires considering the frequency spectrum of the impulse and the line 

impedance at the significant frequencies of that spectrum. 

Approximations can be made for specific current surge wave 

forms. 

3. Isolating power transformers are intended to serve as ground 

isolators, or ground-loop breaks. They do not provide appreciable attenuation of line-to-line transients unless they are 

operating with their series reactance combined with a welldefined shunt load on the secondary. 

4. Ferro-resonant line conditioners can provide attenuation of 

fast line-to-line transients with ratios of 100:l or higher. 

Adding a small fixed load on the output side can raise this 

attenuation to 150:1, or more. 

5. The connection options for surge suppressors must be 

matched to the protection requirements for optimum protection at minimum cost. Universally applicable solutions always 

tend to be more expensive. 

6. Careful design is required for impulse generators. Improvisation can lead to meaningless results and wasted time. 

7. In testing for surge protection evaluation, the timing of the 

surge with respect to the power line frequency can be 

significant. 

8. The pure and sanitary test waves specified by test standards 

are intended to obtain reproducible results rather than to 

duplicate surges occurring in reality. Complex wiring system 

(within a building or within equipment) will promptly 

transform the pure wave form into a distorted form, but that 

does not prevent consistent results, since an agreement exists 

on the initial test wave. 

ACKNOWLEDGMENTS 

Discussion of surge protection techniques among members of 

the IEEE Working Group on Surge Characterization (for instance, 

the effect of connection options) has provided valuable insights 

into the problems and has motivated the examples presented in 

this paper. The contributions made by reviewers of the paper provided further insight on the subject and are gratefully acknowledged. 
REFERENCES 

IEEE Std 587-1980, "IEEE Guide for Surge Voltages in 

Low-Voltage AC Power Circuits," The Institute of Electrical 

and Electronics Engineers, Inc., New York. 

IEC Publication 664, "Insulation co-ordination within lowvoltage systems including clearances and creepage distances 

for equipment," International Electrotechnical Commission, 

Geneva, Switzerland, 1980. (Available in the United States 

from ANSI.) 

IEEE Std 465.1-1977, "IEEE Standard Test Specifications for 

Gas Tube Surge Protective Devices," The Institute of Electrical and Electronics Engineers, Inc., New York. 

IEEE Std C62.32-1981, "IEEE Standard Test Specifications 

for Low-Voltage Air Gap Surge-Protective Devices," The 

Institute of Electrical and Electronics Engineers, Inc., 

New York. 

IEEE Std C62.33-1982, "IEEE Standard Test Specifications 

for Varistor Surge-Protective Devices," The Institute of 

Electrical and Electronics Engineers, Inc., New York. 

F.A. Fisher and F.D. Martzloff, "Transient Control Levels," 

IEEE Transactions on Power Apparatus and Systems 95 (11, 

JanuaryIFebruary, 1976. 

F.D. Martzloff and F.A. Fisher, "Transient Control Level 

Philosophy and Implementation," paper presented at the 2nd 

EMC Symposium on Electromagnetic Compatibility, Montreux, June 28-30, 1977. 

F.D. Martzloff "Coordination of Surge Protectors in LowVoltage AC Power Circuits," IEEE Transactions on Power 

Appararus and Systems 99 (I), JanuaryIFebruary, 1980. 

National Electrical Code 1981, National Fire Protection Association, Boston, Massachusetts. 

Franqois D. Martzloff (M'56) was born 

in France, and received his undergraduate degree at the Ecole SpBciale De 

Micanique et D'Electricitk in 1951; he 

received the MSEE degree from Georgia Tech in 1952 and the MSIA degree 

from Union College in 1971. 

Since 1956 he has been with the 

General Electric Company, where he 

gained experience in the Transformer 

and Switchgear Divisions. Upon joining General Electric Corporate Research and Development in 1961, he became involved 

in power semiconductor circuits and overvoltage protection. He 

has participated in the introduction and application of metal 

oxide varistors since 1971. 

In IEEE, Mr. Martzloff is active on the Surge Protective 

Devices Committee. He is chairman of the Working Group on 

Surge Characterlzation in Low-Voltage Circuits. He is also a 

member of the Ad Hoc Advisory Subcommittee of the USA 

Advisory Committee on IEC SIC 28A and ANSI C.62 Subcommittee on Low-Voltage Surge Protective Devices. He has been 

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

transient protection. 

Discussion 

A. C. Liew (National University of Singapore, Kent Ridge, Singapore): 

The author has once again presented us with an interesting and practically useful paper. 

Regarding the application of classical transmission line concepts to 

wiring systems and the concept of surge impedance, it is well known 

that surge impedance or characteristic impedance is applicable directly 

only until the time of arrival of the first reflection. After that, a lumped 

equivalent circuit is usually used or great effort in keeping track of the 

reflected and transmitted waves must be taken. This is evident in figure 

14 of the paper. 

The author's comments of the following observation made by us are 

appreciated. 

We have found that for nearby lightning strokes, the induced 

voltages on the wiring system of a building (even when supplied by an 

underground cable at the service entrance) does not have to be very 

large to cause operation of sensitive Ground Fault Circuit Interrupters 

or Earth Leakage Circuit Breakers (current-operated type). Even with 

the installation of low voltage lightning arresters (500 V or 380 V type 

for a 415 V system) before the Ground Fault Circuit Interrupter, sufficient unbalance surge currents to ground can flow to cause its operation. This was traced to the distributed capacitances to ground of the 

wiring system. With the liberal connection of suppressers to the input 

terminals of sensitive equipment, the situation is likely to be further aggravated. Thus, while no damage occurs as with successful surge suppression, this nuisance tripping may be intolerable in certain cases. 

Manuscript received August 9, 1982. 

F. D. Martdoff: Indeed, as pointed out by Liew and myself, there are 

limitations to the application of classical transmission line concepts to 

wiring systems. However, while Liew states that these are well known, 

my experience in discussing the topic has shown me that in many instances, in the heat of a discussion, or under the pressure of a postmortem, some erroneous or misapplied concepts can surface. 

To avoid these situations, some repetition of known facts, presented 

with concrete examples, may be helpful and provide useful guidance. 

Thus, the purpose of the paper is not to report discoveries, but to make 

better known the limitations and pitfalls cited in the paper, in the context of concerns on surge propagation and attenuation. 

E. K. Howell: While Liew reports operation of Ground Fault Circuit 

Interrupters coincident with nearby lightning strokes and attributes this 

operation to surge currents resulting from distributed capacitance of the 

wiring system, the information provided is not sufficient to warrant any 

specific conclusions, explanations, or recommend&ions. 

Most Ground Fault Circuit Interrupters (GFCI) today use electronic 

signal processing and provide limiting and integration of the fault current signal, which tends to prevent operation by the fast surge currents. 

However, a flash-over in the wiring system may initiate a sub-cycle 

follow-through current, at power system frequency, having sufficient 

magnitude and duration to require operation of the interrupter. Furthermore, spark-gap types of low voltage ligntning arresters have, inherently, a follow-through current which is limited by a varistor but 

may be large enough to properly cause GFCI operation if that current is 

allowed to pass through the ground fault current sensor. 

There is also the possibility that the electronic circuit was susceptible 

to the surge voltage, rather than the current, as the result of insulation 

breakdown or some parasitic high-frequency coupling within the GFCI 

device. The present Underwriters Laboratory GFCI Standard No. 943 

requires immunity (no tripping) to 3 kV crest of the 0.5 ps-100 kHz 

voltage surge waveform. No current surge response requirement exists 

today. If tripping in response to a surge of voltage or current is sufficiently intolerable, then specifications defining acceptable performance 

should be considered, for either general use or special-purpose devices. 

We both thank Professor Liew for the opportunity to clarify this subject. 

Manuscript received October 12, 1982.