Power Electronics Applications Center

Knoxville TN 37932 USA

Amansoor@epri-peac.com

ç

National Institute of Standards and Technology

Gaithersburg MD 20899 USA

f.martzloff@ieee.org

© 1996 IEEE

Reprinted, with permission, from IEEE Transactions on Power Delivery, Vol.12, No.3, July 1997

Significance:

Part 2 Development of Standards – Reality checks

Part 4 Coupling and propagation of surges

In the propagation of a surge current injected at the service entrance of a building, two significant factors can

prevent the propagation of a postulated “large” surge current to the end of the branch circuits of the facility. 

1. The combination of the inherent inductance of the wiring and the high rate of current change for such a

current to begin flowing into the branch circuit results in a high voltage at the driving end (V = L x di /

dt).

2. In the absence of a surge=protective device at the service entrance, the withstand voltage of the wiring

devices at the driving end – the service entrance – is very likely to be exceeded by the voltage that this

rising current will develop along the branch circuit. 

The resulting flashover will abort further propagation of the surge current toward the far end, thus

establishing a limit to what is physically possible. If there is a surge-protective device at the service

entrance, the scenario becomes a matter of cascade coordination.

Provides quantitative information on this limitation, as a function of wiring length and current rate of rise.

Filename: More begets less
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 

I is the same as the current delivered by the surge generator. 

Another effect of the added inductance is the increase in the 

time from origin to the first current zero, 33 ps in Figure 5, 

compared to 25 ys 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'"l (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 ps in that case). The 

result is that the simulation of Figure 6 took 43 seconds on a 

486-based PC, compared to 263 seconds for Figure 7. 

Therefore, the lumped-parameter model is perfectly adequate to 

represent reality, and performing a transmission-line analysis [5] 

is an unnecessary consumption of computing time and resources. 

Figure 6 - Impinging current and resulting upstream 

voltage as computed with lumped-parameters model 

Figure 7 - Impinging current and resulting upstream 

voltage as computed with transmission-line model 

In both Figures 6 and 7, the effect of the branch circuit 

inductance on the resulting voltage is apparent as the peak voltage 

occurs at the beginning of the rise (as soon as the "gentle toe" 

effect ceases), not at the peak of the current. The step change in 

the voltage trace corresponds to the reversal of the current in the 

varistor, showing the relative contributions of the varistor effect 

and of the inductive effect as seen from the upstream end. 

Table 1 below shows the results of such computations for the 

waveform of Figures 5, 6 and 7. As mentioned above, the 

insertion of an inductance in the load connected to the surge 

generator increased the rise time beyond the standard 8 ps. In 

making the parametric computations, we chose to stay with this 

10 ys value to maintain continuity with the test/model validation. 

TABLE 1 

Upstream voltage (in kV) necessary to drive a current of the peak 

value shown (columns) and rise time of 10 ps into a branch circuit 

of length as shown (rows), terminated with a 130-V rated varistor 

Length\Peak 2kA 3kA 5kA 7kA 10kA 
Figure 8 - Three surge current waveforms with different rise 

times used to compute the values of Table 2 

Figure 8 shows three waveforms of same amplitude, with 

nominal rise time of 5 ps, 10 ps, and 20 p, obtained by taking 

half or double of the frequency used in Eq. (2). The actual rise 

time [1.25 x (time from 10% to 90%)], as opposed to the nominal 

rise time used to describe the waveforms, was computed as well 

as the maximum rate of rise for each wave. The maximum rate of 

rise (which is obtained when the second derivative of the current 

is equal to zero) occurs initially, once the gentle toe is over, and 

determines the maximum resulting voltage produced by the 

inductive effect. Table 2 shows the corresponding values of the 

rise time, maximum rate of rise, and resulting voltage for a 

branch circuit length of 10 m and amplitude of 5 kA. Note that 

for a 1-to-4 increase in nominal rise time, the maximum di/dt 

decreases only by one half, with the same decrease appearing in 

the resulting voltage, showing once again that initial rate of rise 

is more important than rise time and amplitude. 

TABLE 2 

Effect of the rate of rise of the postulated current on the 

resulting voltage at the upstream end of the branch circuit 

Nominal rise time, ps 5 10 20 

Actual rise time, ws 4.3 9.5 13.5 

Maximum di/dt, A!ps 1250 850 630 

Resulting voltage, kV 7.0 5.2 3.6 

In the scenario tested and modeled so far, no flashover 

possibility was considered. Nevertheless, the values shown in 

Table 1 clearly indicate that some real-world circuit lengths and 

surge parameters postulated in some SPD application standards 

under development can produce high upstream voltages that will 

cause a flashover of the upstream wiring devices. 

C. The paradox of "more begets less" 

Common-sense intuition might lead the unwary to expect that 

higher surge currents would impose a greater stress on the circuit 

components, including the downstream varistor. Also, a longer 

branch circuit, with its corresponding higher inductance, could 

be expected to have the capability of storing more energy during 

build-up of the surge current toward the downstream varistor, 

into which that stored energy ultimately has to be dissipated. 

Cascade coordination studies [4], [6], [a], have shown that in 

some cases, the downstream varistor continues to carry current 

long after the impinging surge current has gone past its peak. 

To explore the validity of such expectations, we performed 

tests and modeling, with an actual gap in the test circuit, and a 

switch in the model circuit, to bypass the current at the upstream 

end when sparkover voltage is attained. By measuring the 

current that flows in the branch circuit toward the downstream 

varistor and the voltage across the varistor, the energy deposited 

in the varistor during the total surge event can be determined. 

Likewise, the modeling can determine the current in the varistor, 

hence the voltage across it, and allow computation of the energy. 

In [4], agreement was reported between, on the one hand, 

computing the deposited energy through actual measurement of 

the current and voltage, followed by computation of the energy 

by means of the digital signal analyzer used for measurements 

and, on the other hand, the model computations. Therefore, in 

the tests reported here, we were satisfied to verify waveform 

agreement between the actual varistor current measurement and 

the computed varistor current, and let the model alone compute 

the energy deposited in the downstream varistor. 

Figure 9 shows the real-world recording of the situation that 

develops for a "clearance" sparkover of 2 kV. This relatively 

low value, compared to the 6 kV to 10 kV level that we might 

expect from typical low-voltage wiring devices, is made necessary for the test case where only 9 meters of branch circuit were 

considered, and the setting of the surge generator was maintained 

at the same nominal 3 kA short-circuit current. The object, of 

course, is to demonstrate that the clearances are likely to flash 

over, as indicated by progressively higher values of the necessary 

upstream driving (or resulting) voltage shown in Table 1. 

Under the conditions of Figure 9, sparkover of the gap 

occurred at approximately 1 ps. After sparkover, the current 

delivered by the surge generator is the sum of the currents in the 

gap and in the branch circuit. Its peak (3.2 kA) is greater than 

those of Figures 2 and 5 because the generator does not need to 

overcome the varistor that reduced the voltage available for 

driving the current, nor the impedance of the 9 meters of cable. 

Top trace: Resulting voltage, 500 Vldiv 

Center trace: MOV current, 500 Ndiv 

Bottom trace: Total current, 500 Ndiv 

. , 

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 all the energy has to go to the 

downstream varistor. At the higher threat level of 7 kA, the 

voltage produced in the inductance of 10 meters of line, added to 

the varistor voltage, is sufficient to sparkover the 6 kV gap, 

relieving the varistor from further involvement beyond that of 

discharging the energy stored in the line. In the case of the 30-m 

long line, this transition occurs between 2 kA and 3 kA. 

Turning now to the first observation, that higher current or 

greater inductance result in less stress, this apparent paradox is 

caused by the fact that with the higher values of di/dt and L, the 

voltage at the clearance rises more quickly to the flashover point. 

Consequently, the build-up of energy in the line inductance is 

shut-off earlier so that the current level in the line reached at that 

point is lower and, in spite of the greater inductance, the stored 

energy ?h L i2 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 measurements reported in this paper were conducted by 

Kermit Phipps at the EPRI Power Electronics Applications 

Center. Support for the modeling was provided by the Electric 

Power Research Institute. Support for the development of this 

reality check, motivated by the discussions and contributions of 

IEEE and IEC colleagues, was provided by Delmarva Power & 

Light and by Pacific Gas & Electric. 

VI. REFERENCES 

[l]IEEE/ANSI 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, January/ 

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 1812-1819. Also in IEEE 

Transactions, IAS-29 No.4, July/August 1993, pp 680-687. 

[5] Standler, R.B., "Coordination of Surge Arresters and Suppressors 

for Use on Low-Voltage Mains." In Proceedings, 1991 Ziirich EMC 

Symposium, pp 517-524. 

[6] Hostfet, O.T., Hervland, T., Nansen, B. and Huse, I., "Coordination 

of of surge-protective devices in power supply systems: Needs for 

secondary protection," In Proceedings, 20th International Conference 

on Lightning Protection, 1992. 

173 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- 1-C 1LJ. 

[9] Martzloff, F.D., "Coupling, Propagation, and Side Effects of Surges 

in an Industrial Building Wiring System," IEEE Trans-actions IA-26, 

No.2, MarchIApril 1990, pp 193-203. 

[lo] EPRI, "Electromagnetic Transients Program (EMTP), Version 2.0; 

Volume 1; Main Program; Volume 2: Auxiliary Routines" EPRI Report 

EL-6421-L, July 1989. 

[ll] Standler, R.B., 'Transmission Line Models for Coordination of 

Surge-Protective Devices," In Proceedings, IEEE Inter-national 

Symposium on Electromagnetic Compatibility, 1993, pp 358-362. 

[12]Martzloff, F.D., Mansoor, A., Phipps, K.0, and Grady, M. 

"Surging the Upside-Down House: Measurements and Modeling 

Results," In Proceedings, PQA'95 Conference, EPRI, 1995. 

[13] IEEEYANSI C62.45-1992 "Guide on Surge Testing for Equipment 

Connected to Low-Voltage AC Power Circuits." 

Arshad Mansoor (M' 1995) Is Electrical Systems Engineer at the EPRI 

Power Electronics Applications 

Center (PEAC). He received his MS 

and Ph.D. in electrical engineering 

hom the University of Texas, Austin, 

in 1992 and 1994 respectively. His 

areas of interest include power 

quality, power systems transients, 

harmonics, surge protection, and 

EMTP model development. 

FranqoisMartzloff (M'1956, F'1983) 

Born and educated in France, with 

additional MS degrees from Georgia 

Tech and from Union College, 

worked at General Electric for 29 

years and now ten years at the 

National Institute of Standards and 

Technology. He is contributing to 

several committees for development 

of standards on EMC and surge 

protection in the IEEE and the IEC. 
Discussion 

M. Darveniza (University of Queensland, Australia 4072): 

The authors are to be congratulated for drawing attention to an important feature in surge 

protection. Namely, the magnitude of the voltage up-line from a surge protective device 

will exceed the protective level of the SPD, the extent of the over-voltage depending on 

distance and on waveshape of the incident surge. Because of this, flashover at an 

upstream device (for example, another SPD) will limit the severity of the surge stressing 

the downstream SPD. 

Two examples are offered which support the author's statement "more begets less" and 

which can he rephrased by saying that in some cases, a less onerous surge may impose 

more severe overstress than a more onerous incident surge. The two examples are: 

A cable-entry substation protected by an upstream SPD connected at the 

overhead line-tocable junction. The most severe stress at the substation occurs 

when the surge incident from the line onto the cable is just not large enough for 

operation of the SPD at the line-cable junction. 

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.