NUREG/CR--3263

SAND83-2622

RV

Printed April 1984

Status Report: Correlation of

Electrical Cable Failure With

Mechanical Degradation

Otmar Stuetzer

$~4~/C1

-Prepared by

Sandia National Laboratories

Albuquerque, New Mexico 87185 and Iivermore, California 94550

for the United States Department of fnprgy

under Contract DE-AC04-76D",OTh9,. i

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NOTICE

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NUREG/CR-3263

SAND83-2622

RV

STATUS REPORT: CORRELATION OF ELECTRICAL

CABLE FAILURE WITH MECHANICAL DEGRADATION

Otmar Stuetzer

Printed: April 1984

Sandia National Laboratories

Albuquerque, NM 87185

Operated by

Sandia Corporation

for the

U.S. Department of Energy

Prepared for

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

Under Memorandum of Understanding DOE 40-550-75

NRC FIN No. A-1051

ABSTRACT

An attempt is being made to assess complete electrical failure

of signal and low-power cables typically used in nuclear power

plant containments and to correlate failure modes with the

mechanical deterioration of the elastomeric cable material.

Work over the past 24 months, although limited to one cable

configuration, has identified creep shortout and insulator

cracking, both aggravated by mechanical stresses, as the phenomena most likely to cause electrical breakdown. Comprehensive tests have been run for six months and are continuing.

Preliminary conclusions can be drawn and are reported.

iii, iv

CONTENTS

Page

EXECUTIVE SUMMARY 1

I. INTRODUCTION 4

II. PURELY ELECTRICAL FAILURES 7

1. Field Breakdown 8

2. Dielectric Loss Failure 11

3. Comment on Circuit Failure 17

4. Other Phenomena 18

5. Summary 20

III. TEST INSTRUMENTATION 21

1. Dedicated Test Facility 21

2. Cable Tensioners 21

3. Special Apparatus 29

IV. CREEP SHORTOUT 31

1. General Comments 31

2. Preliminary High-Strain Experiments 33

a) Shortout Observation 35

b) Deformation Measurements 39

c) Conclusions and Comments 39

3. Systematic Experiments 44

a) Program 44

b) Measurements 46

c) Preliminary Results 54

V. CRACKS IN CABLES 59

1. General Comments 59

2. Preliminary Tests 60

a) Long Cables 60

b) Cracks in Stressed Bends 62

c) Maintenance Accident 62

3. Ongoing Systematic Tests 65

a) Dynamic Cracking 65

b) Static Cracks 66

VI. CONCLUSION 69

REFERENCES 71

APPENDIX A A-1

V
ILLUSTRATIONS

Figure Page

1 Dose-Rate Dependency of Breakdown Stress

(After Asaka, Ref. 13) 9

2 Breakdown Voltage (Vb) and Its Standard

Deviation (a) at 225 0 C 12

3 Dissipation Factor and Volume Resistivity vs

Dose Rate (After Kuriyama, Ref. 16): Dashed

Curve is for Oxygen-Free Atmosphere 13

4 Insulation Resistance in Steam and Chemical

Spray Environment (Flame Retardant EPR)

(After Murata, Ref. 18) 14

5 DC-Leakage Current IL for 1-m Cable Length

(Standard Deviation of 'IL is Orel and Dissipation Factor is D) 15

6 Effect of-Dose Rate on Insulation Resistance

(After Murata et al.. Ref. 18)) 19

7 Overview of Dedicated Test Facility 22

8 Schematic of the Heating Circuits 23

9 Photograph of the Instrument Cart 24

10 Photograph of the Tensioner 25

11 Cross Section of Gooseneck Support and Cable 27

12 Correction for Creep Distance as Function of

the Ratio Between Ellipse Axes (for cases

where angle 0 = 900) 28

13 Cable Bending Tool 30

14 Setup for the Treeing Test 30

15 Room-Temperature Master Curve for HighDensity Polyethylene in Creep (After

Aloisio and Brockway. Ref. 27) 32

16 Compression Creep Modulus for EPR and XLPE

at 700C as a Function of Time (After

St. Onge. Ref. 15) 34

vi
ILLUSTRATIONS (cont.)

Figure Page

17 Compression Creep Modulus for EPR and XLPE

at 10 Hours as a Function of Temperature

(After St. Onge, Ref. 15) 34

18 Relative Sizes and Shapes of Cable and Supports 36

19 Corner Geometry and Force Distribution 36

20 Stress on Support Surface vs Support Radius 37

21 Cut through a New Cable (a) and a Cable Stretched

by 0.9 kg at 1256C for 1 Day (b) 40

22 Creep Measurements at 1250C and 1900C 42

23 Creep Compliance vs Applied Stress for Various

Exposure Temperatures 43

24 DC Breakdown Voltage vs Remaining Polymer Thickness at 2250C 45

25 X-Ray Negative of Tensioner 47

26 Symmetric (a) and Unsymmetric (b) Hang 48

27 Measurement Parameters (38 x) 49

28 Density Distribution 50

29 Average Reduction for 6 Cables 55

30 Thickness Reduction for Worst of 6 Cables 56

31 Worst Thickness Reduction for EPR Layer 57

32 Model for Equation 4 64

33 Crack Experiments 64

34 Total Number of Bending Cracks for 6 Cables 67

vs Time

35 Embrittlement Cracks at 2250C 68

vii
LIST OF TABLES

Table Pa•e

1 60 Hz Breakdown Field and Standard Deviation

for Materials at Different Temperatures 10

2 Creep Shortouts 38

3 Position Dependence of peformation 41

4 Example of Data Evaluation 51

5 Preexperiment Averages 53

6 Crack Statistics After Heating to 2250C

Over 5 Days 61

A-i Cable Test Parameters (27) A-2

viii
ACKNOWLEDGMENTS

The project was initiated and has been monitored by L. L.

Bonzon. who provided essential administrative support and continuous technical guidance. Discussions with K. T. Gillen and

L. D. Bustard were helpful. D. G. Sample assisted in the

design of the X-ray measurement system and assumed responsibility for development and accurate magnification of the pictures. B. D. Haensche investigated suitable image evaluation

methods and provided digitized magnifications with density

traces.

Particular acknowledgment is due to G. J. Simmons, who

designed and built all special instruments and mechanical

devices. designed and installed the test facility, and made

most of the measurements.

ix

EXECUTIVE SUMMARY

The purpose of this investigation is to link the degradation

of the elastomeric cable materials with the actual potential

for circuit failure of safety-related cables in the containment building of a nuclear power plant: failure is defined as

an electrical shortout or as an obvious shortout path.

In the introduction (Section I). the essential difficulties

and limitations of the present work are discussed. A very

large number of independent variables (e.g.. composition.

geometry, additives, aging sequences, rate effects) lead to a

high degree of complexity and force judicious constraints on

experimentation. Further complications are introduced by the

shortcomings of accelerated testing methods. Singleconductor, 480 V, power cables with ethylene propylene rubber

(EPR) insulators and Hypalon jacket were chosen for the

experiments. With this configuration, a number of experiments

large enough to be statistically significant are feasible.

This approach furnishes sufficient data to form a base for

extrapolation by means of scaling models to other geometries

and to other mechanical boundary conditions.

Literature data and several hundred scoping tests (Section II)

are the basis for a discussion of purely electrical failures:

i.e. failures occurring when cables are not mechanically

stressed or handled. It is shown that direct electrical-field

breakdown as well as thermal runaway is very unlikely at rated

voltages even after long-term exposure to high temperatures

(1900C) and to high total doses of gamma radiation (10' rads):

rate effects are negligible or perhaps even beneficial. The

deterioration of insulation resistance due to the presence of

strong continuous radiation (10 Mrad/h). during an accident.

leads to only negligibly small shunt currents; also, high

humidity has little influence in the absence of cracks.

Thus, it is concluded that cables exposed to the environment

in a reactor containment for 40 years, but undisturbed and

unstressed, are highly unlikely to fail. This fact simplifies

the present investigation by limiting it to situations where

the geometry of the cable is affected by mechanical forces.

The remaining failure mechanisms fall into two categories.

First, if the cable is stretched by an applied force (or by

its own weight) over an edge with small curvature, the metal

wires will gradually creep through the soft polymeric insulation resulting in metallic contact between cable wires or a

wire and the cable support, shorting the cable (creep shortout). Second, the polymeric materials embrittled by aging

may crack under mechanical stress. To permit long-term aging,

creep, and cracking measurements, a dedicated facility (Section III) has been established. It consists principally of

seven temperature-controlled heat chambers, in which stressed

and unstressed cable samples are exposed to constant elevated

I
temperatures for long periods. Creep distances can be precision measured by X-raying at regular intervals.

Systematic creep experiments (Section IV) are presently performed on cable samples that are bent in a U-shape over a

4.8-mm diameter support rod with each cable end supporting a

1-lb weight. After 7 months, no situation approaching shortout has been identified. It was found that the most severe

creep deformation occurred at a temperature of approximately

125 0C. Apparently, at higher temperatures, polymer embrittlement proceeds so fast that it effectively inhibits creep.

For the investigation of crack failures (Section V), it is

stipulated that any crack penetrating to the metal wire constitutes cable failure. Three specific situations are being

investigated:

a) Long cables (10 m or more), lying undisturbed in conduits, may develop cracks under thermal cycling. So far,

this phenomenon has only been observed when the cable

temperature exceeded 150°C.

b) Cables that are sharply bent may develop cracks under

mechanical stresses caused by thermal cycling or by

weighting. Severe cracking tends to occur at temperatures above 175°C.

c) Cables embrittled by aging may be stressed and cracked

during maintenance activities. A simple model links the

onset of cracks to the bending radius and the breaking

strain at elongation for the cable materials. Experimental results show good correlation with the model.

During the investigations, phenomena were observed that tend

to decrease the failure probability of cables. For example,

in bent and stressed specimens, the polymeric material becomes

more dense above the support area, which improves the breakdown field strength of the material. In addition there is a

worst position for the lie of the strands in the conductor,

this position leading to maximum creep. One moderating factor,

though, is that this position is mechanically unstable and

tends to eliminate itself.

While the investigation is still proceeding, a number of conclusions (Section VI) may already be drawn. Cables are basically very reliable components with very low failure probability. Failures are mostly due to marginal design or handling practices. Operational integrity for EPR - Hypalon

power cables is predicted under the following conditions:

(1) The cable temperature does not exceed 190°C (thermal

runaway condition), and exposure time at high temperatures is limited, so cracks in unstressed cables do not

occur.

2
(2) Cable overhang over edges with small radius of curvature does not exceed 10 m (creep shortout condition).

(3) Cables for which the breaking strain ratio (elongationat-break ratio) has become less than 0.4 are not handled

or permitted to move mechanically (bending ciack condition).

AlL

" iI -

'V•./

3
1. INTRODUCTION

Improvement in the understanding of the complex aging and

deterioration behavior of elastomeric cable materials exposed

to the environment in the containment of a nuclear power plant

is continuing. However, the correlation of observed material

changes with the electrical failure rate of the cable circuit

has not yet been achieved. This problem is of considerable

complexity due to the large number of variables introduced by

materials combinations, additives, rate effects, circuitry,

aging sequences, environments, mechanical support, and others.

Even a conservative compilation (c.f., Appendix A) shows from

25 to 30 independent parameters leading to hundreds of thousands of different combinations. The situation is further

aggravated by two additional facts:

(1) Cables. per se. are simple and well-developed components.

whose failure rate is very low: many experiments are

ther-efore needed to determine reasonably reliable and

statistically valid failure data. and

(2) Many parameters of possible importance, such as additives

and materials- processing sequences, are proprietary,

often unknown, and/or batch variable; this requires a

multiplicity of tests for statistical evaluation.

Reasonable constraints had to be imposed to keep the investigation within feasible limits and achievable goals. Literature studies and several hundred scoping tests served to bound

and define an acceptable program. The program's aim is to

find and understand the most important operational failure

mechanisms for electrical cables in a nuclear-reactor containment. Extrapolation to 40-yr exposure with the possibility

of a reactor accident at the end of this time is desirable.

Considerable simplification is possible through the realization that for modern qualified cables operated at design voltages or less, the deterioration of most electrical properties

of the cable (such as loss factor or breakdown field strength)

is unlikely to lead to electrical-system failure even if combined (synergistic) effects are taken into account. Two electromechanical failure mechanisms are of concern, however:

(1) creep shortout, occurring when a mechanically stressed

cable is bent over a corner and conductors are pulled towards

each other or towards a grounded electrode, and (2) cracking

of the insulating layers, which in the presence of electrolytic conduction may lead to current diversion.

One constraint imposed on the program is the use of a single

combination of materials (EPR for the insulating material and

Hypalon for the jacket) and a single cable geometry (600 V.

20-A rated low-power cable): this cable type is typically

used in nuclear-power plants. An attempt is being made to

4
generalize the results to other cable types by correlating

the observed cable damage with the rather well-known materials

deterioration. Analytical extrapolation to other geometries

and to a variety of mechanical stresses is accomplished by

correlation to. and then generalization from. simple scaling

models. With this approach, it was feasible to limit the

number of experiments to 84 temperature-time value pairs per

year. The number of tests for each value pair (6 precision

creep measurements and 40 crack tests) was chosen so that it

is large enough to be statistically significant.

A number of additional simplifications are made, that are

generally thought to be on the conservative side (i.e..

increasing the predicted frequency of damage). First, for

crack-caused electrical failure, any crack penetrating through

the jacket to the cable insulator was assumed to be a "failure." Second. it was assumed that radiation exposure mitigates creep shortout by embrittling the polymers; hence, in

the creep experiments, only temperature exposure is used as

an aging mechanism. Third. to extrapolate cracking data, a

working hypothesis was used that links cracking to the end

value of the breaking strain for the materials, disregarding

the combination of radiation and/or time-temperature exposures under which the breaking strain was obtained. Fourth, (and

perhaps not necessarily conservative), creep shortout was

defined as an impedance of less than some fraction of an ohm.

These simplifications decrease the number of necessary experiments and allow a bounding of the problem.

To some degree these simplifications are compensated for by addition of test parameters. Up to now, maximum temperatures

during an accident were generally considered to be below

165°C: for this investigation, maximum test-parameter temperatures of up to 2256C were included. The reasons are threefold: (1) an expectation that predicted containment temperatures may be increased to account for severe accident scenarios, (2) the fact that cables may be overloaded and therefore

hotter than the environment, and (3) the desire to accelerate

the aging process in experiments being conducted.

Further complication is introduced by the shortcomings of

accelerated testing methods. Creep and cracking phenomena

cannot be assumed to exhibit Arrhenius-type behavior. Extrapolations to 40 years may require measurements over many

years, at least for situations where no physical arguments are

readily available to establish the direction of long-term

trends., For this reason, and to permit desirable proof tests,

a dedicated long-term exposure facility has been established.

It consists of seven automatically controlled heat chambers

and three large heat pipes, all designed to operate with minimum attention for long periods. Self-calibrating fixtures

that permit accurate mechanical measurements over long times

have been incorporated.

5
In the following sections the essential details of the work

performed to date will be described. In Section II, literature data and preliminary experiments will be discussed; it

will be shown that purely electrical breakdown is unlikely

for the cable type under consideration. In Section III,

arrangements used to obtain creep and cracking data are

described, together with calibration and correction methods.

Section IV concerns creep measurements and the results of a

first-order (two-dimensional) creep analysis. In Section V.

the cracking data obtained are analyzed. Section VI summarizes the results and discusses possible and probable consequences.

6
II. PURELY ELECTRICAL FAILURES

Electrical cables used in nuclear power stations are very

reliable components. An investigationL of 1400 reactor

trips in 42 pressurized-water reactors between 1961 and 1978

lists only 8 shutdowns caused by cables; 5 of these failures

were due to electrical shorts. An average of 7.2 outages per

plant per year is found by the author; this leads to a 4%

probability of a severe operational cable failure per plant

per year. As there are typically 500-1500 km of cables in a

reactor containment system, 2 the failure probability per

meter of cable per year is very low (about 4 x 10-'). For

example, no accident occurred during the span of the above

investigation, given that no cable failure has been proven

for the Three Mile Island accident.2 To simplify matters,

no attempt is made to separate the type of failures in the

above work.

The goal in this section is to attempt to show that purely

electrical failures (complete field breakdown or thermal runaway melting the insulation) in cables exposed for 40 years

to the environment in the containment of a nuclear power

plant, and then to a design-basis accident, are highly

unlikely. This assumes the cables are not handled or mechanically stressed. Medium voltage (approximately 500 volt) lowpower cables, those normally considered the most likely to

fail, are used, and it is assumed that the cables are operated

within ratings; severe overloads due to circuit breaker failure are not included in the analysis. Insulating materials

of interest are the ones used most often in modern reactor

cabling (e.g.. PE, XLPE, XLPO, and EPR. with Hypalon as a

typical jacket material). The electrical properties of the

various materials are quite similar; average breakdown field

strengths differ by factors of 3, and conductivities by factors of 10 in either direction. For the purpose of this discussion, order of magnitude argumentation is justified; differences between insulator and jacket can then be ignored.

All cables are assumed to be qualified for nuclear power

applications; i.e., test samples have been exposed to a simultaneous or sequential temperature and radiation treatment

designed to simulate a 40-yr exposure to a typical reactorcontainment environment: a design-basis accident simulation

follows, during or after which the electrical quality of the

cable is tested. The tests are not intended to yield and do

not yield statistically meaningful reliability numbers, as

the quantity of cable tested is small. Furthermore, the

Arrhenius approximations used to extrapolate 40 years into

the future are being challenged by some workers.4 However, to avoid 40-yr tests, the qualification procedure

involves radiation overexposure and electrical overtests.

Experience supports the conclusion that the qualification

procedure is meaningful. Few overall failures and fewer.

7
if any, purely electrical failures have occurred even in

cables exposed to reactor-containment environment for long

periods.1,s

More recently, it was confirmed that rate effects 6 and synergistic rate/temperature effects7 can lead to unexpected

mechanical deterioration in cable materials. The resulting

impact on the important electrical properties of cables falls

within the scope of the present investigation. It is also

desirable to extend exposures to higher temperature than presently used in qualification procedures to encompass a greater

temperature range and to reduce acceleration test times.'

It will now be shown that the above effedts do not influence

breakdown properties and have a small and manageable influence

on cable conductivity. It should be noted that the experiments discussed do not intentionally expose cables to mechanical stresses.

1. Field Breakdown

Qualification tests with high dose-rate aging have been

reported by a number of investigators, including Bennet9

(XLPO). Thome'° (XLPE and EPR), and Hosticka et al."1 (XLPE);

all of whom used minor variations of the standard IEEE qualification method. 1 2 The tests revealed no breakdowns except when cables were cracked by rewinding and then immersed in

water (the case excluded in this section).

Of particular interest is a report by GrUb and Langeset.s who

exposed unusually long lengths of EPR cabling to a total of

300 Mrad of radiation (combined gammas, betas, and neutrons)

over 4 years. The authors found no noticeable change in

breakdown field strength over this time and dosage. The average dose rate in the GrUb-Langeset test was about 8 krad/h,

much less than generally used for qualification aging. Doserate effects on the breakdown field strength of EPR, then.

must be very low or negligible. The same conclusion is indicated by Figure 1. which shows dose-rate-effect measurements

of Asaka.1 3

As was the case in Asaka's work, breakdown tests for cables

are generally conducted at room temperature. Occasionally.

tests are also conducted at the highest temperature predicted

during an accident. For EPR, the breakdown field increases

with increasing temperature, as Corbelli and Toniolil4 have

found for temperatures below 100 0 C. One possible explanation

is that breakdown-causing gases diffuse outward more easily

at higher temperatures. Measurements by St. Onge et al.,1$

compiled in Table 1. verify this dependence up to 1700C for

EPR, while measurements for XLPE show a decline of the breakdown field to one-half as temperature increases to 170oC.

Extensive measurements of breakdown field strength above 1700C

were performed during our investigations. An example for EPR

8
70 70 IB I I I

I I I I I

E

E

Iv0

0

W

60-

50sEPR

401-

30 -

201-

o PE INSULATED CABLE

o XLPE INSULATED CABLE

A EPR INSULATED CABLE

. I

FREQUENCY: 50Hz

TEMP.: R.T.

10 I I I I I I I I m

I(

104 5 10 5 5 106

GAMMA RAY EXPOSURE RATE (rad/hr)

Dose-Rate Dependency of Breakdown Stress

(After Asaka, Ref. 13) Figure 1.

9
Table 1

60 Hz Breakdown Field and Standard Deviation for Various Materials

at Different Temperatures (After St. Onge. Ref. 15)

Materials

XLPE. molded

fully cured

XLPE. molded

partially cured

XLPE. cable

Filled XLPE

EPR 1

EPR 2

EPR 3

EPR 4

230C

74(6.7)

70(6.5)

66(7.9)

55(6.8)

41(7.4)

40(2.7)

36(4.6)

46(1.6)

Breakdown

600C

57(2.8)

66(3.0)

49(6.1)

59(1.3)

46(4.0)

44(2.2)

40(5.9)

47(2.7)

field (standard deviation)

(kV/mm)

90 0 C 1100 C 130 0 C

47(3) 37(6.1) 36(6.2)

1500C

36(4.2)

170 0 C

39(6.6)

46(3.8) 39(4.8)

I42(3.2)

49(4.6)

46(1.3)

43(3.6)

38(6.9)

45(5.8)

36(2.2)

49(4.9)

45(3.9)

44(3.7)

44(2.9)

50(1.7)

33(6.8)

40(1.9)

51(5.1)

51(2.4)

40(4.0)

45(2.3)

52(3.2)

37(6.1)

50(4.1)

53(2.8)

45(3.8)

42(4.5)

49(2.8)

30(5.5)

47(3.9)

50(2.7)

44(10.6)

40(2.5)

54(2.3)

26(4.6) 33(5.9)
is shown in Figure 2 for 2250C. It was noted that the average

breakdown voltage increased about 10% over that observed for

the room-temperature voltage, while the standard deviation

stayed about the same.

It is concluded that increasing temperature and radiation

rates do not severely decrease the breakdown field strength

and, by increasing the breakdown strength, are sometimes beneficial. Synergistic effects are unlikely and apparently have

never been reported.

2. Dielectric Loss Failure

A second purely electrical failure mechanism is caused by a

strong increase in dielectric loss or insulator conductivity,

which may lead to circuit starvation or to thermal runaway.

Loss factor (for ac) and insulator resistivity (for dc) are

related parameters used to assess the damage phenomena.

Dose-rate influences on the two parameters were measured by

Kuriyama et al. 1 6 and are presented in Figure 3. There is

only a small change in each parameter as the dose-rate

increases. Seguchi et al."* have exposed EPR and lowdensity PE cables to irradiation with dose rates varying from

5 to 500 krad/h and to a total dose of 100 Mrad. The authors

find there are no dose-rate effects if the materials contain

anti-oxidants: otherwise, the dissipation factor increases by

about 50% with decreasing dose rate, which is in agreement

with Kuriyama's results. It is concluded that dose-rate

effects are small and not important for dielectric losses.

In contrast, the influence of temperature on cable losses are

substantial, although generally temporary (i.e., they last

only as long as the temperature is elevated). In Figure 4.

the insulation resistance as measured by Murata et al.1'

during a LOCA simulation is plotted. Compared to room temperature, temperatures of 174 0 C decrease resistivity by nearly 4

orders of magnitude. The data are only slightly affected by

the aging methods used.

Figure 5 gives an example of higher temperature measurements

(up to 2250C) performed in our investigation. The dc leakage current (1L). averaged over ten 1-m-long cable samples,

is seen to increase by nearly 7 orders of magnitude over the

current measured at room temperature. The ac dissipation

factor (D) first decreases with increasing temperature (as

additives perhaps diffuse out of the polymers) and then rises

to not quite 0.35; dissipation factors were measured at operational ac voltages and frequencies.

The following discussion of thermal runaway is based on a

comparison of the heat generated by losses in the conductor

of the cable and the additional heat caused by the dielectric

losses in the insulator and the jacket. The first quantity

11
BREAK DOWN VOLTAGE

23 - FOR 10 CABLES (EPR)

• T= 225 C

22

U)20 I-

.j • 0 18

-31%

6 > 1 6 

I- .,I 14 -

0

> 12

0 10

8 0

Il.

I I I I I I i, I I I

1 2 3456 7 8910

CABLE NUMBER

Figure 2. Breakdown Voltage (Vb) and Its Standard

Deviation (a) at 2250C

12
0.01.0°

XLPE

uwL

2w.. 1015

olX104 5X104 1X10 5 5x10 5

DOSE RATE (rad/h)

0.54=L

0 | |

EPR

w t

>uJn

1015

0 14

__Ira

0 Ix10 4 5X10 4 j1xO 5x10 0

DOSE RATE (rads/h)

Figure 3. Dissipation Factor and Volume Resistivity vs

Dose Rate (After Kuriyama, Ref. 16); Dashed

Curve is for Oxygen-Free Atmosphere

13
E

C:

wU

0

z

7CD

0

-j

3

103

3

102

3

101

.3

100

. 3

10-1

•200

tv100

I- 0 F

174 0C 1740C. 162cC 16%5020C 1240C

CHVC EMICAL SPRAY -,______6.

! -RATED I !!I VOLTAGE & I CURRENT I S

0.0

0.5

3.5 6.0

5.5

12.0 24.0 120.0

9.0 16.0 72.0 168.0

TIME (hrs) -

Figure 4. Insulation Resistance in Steam and Chemical Spray Environment (Flame Retardant EPR) (After Murata, Ref. 18)
z

w

w

w-j

10-3

(A)

100%

1%

1%

2000 C 150 100 50 25 0C

TEMPERATURE, °C

Figure 5. DC-Leakage Current IL for 1-m Cable Length (Standard Deviation of IL is Orel and Dissipation FacFactor is D)

15
is current dependent, the second depends on the applied voltage. The cable presently under investigation. AWG I14. has a

conductor resistance of 0.00828 ohms/m at room temperature.

At 1000C. this value is increased by-a factor of 1.43. At

the maximum rating. 20 A. the cable current causes heat generation of 4.74 W/m at 1000C, and somewhat more at higher temperatures. The installation has to be able to carry this heat

flux away, and usually is amply dimensioned to do this.

Therefore, dielectric losses may add a comparable amount of

heat generation without undue consequences. To be conservative. we will stipulate that dielectric heat input will not

exceed 20% of the permitted current heat generation. Dielectric losses are therefore limited to 0.95 W/m for the present

cable.

For dc. the dielectric losses are U2 /R where U is the applied voltage and R is the leakage resistance of the cable per unit

length. For an applied voltage of 480 V, it is found that R

must be larger than 0.24 Mohmom to keep the losses under

0.95 W/m. The cable average measurement in Figure 5 (with an

applied voltage of 500 V and a leakage current of 0.4 mA at

an equilibrium temperature of 2250C) yields an average value

of 1.2 Mohmem at the highest experimental temperature. This

cable design can thus be assumed to be safe from thermal runaway up to 2250C.

For ac, the losses are essentially electronic and therefore

generally less than for dc. which contain-an additional ionic

conductivity. For a D-factor of 0.35, the ac losses. U2wCD.

with C being the cable capacitance per unit length, are about

0.015 W/m for the example in Figure 5 and are therefore negligible.

The above deductions assume that the loss resistance and the

D-factor are reasonably.uniformly distributed along the cable.

i.e., that there i-s no "bad spot" solely responsible for the

leakage current. For the measurements in Figure 5. this is

assured by the standard distribution value for the 10 leakage

current measurements, which reaches only a factor of 2 for

the highest temperatures. Unfortunately, many measurements

in literature are made only with one or a very few cable samples.

Not many measurements in literature are made at temperatures

as high as 2250C', If we wish to:determine from such lower

temperature measurements (e.g., Figure 4), the highest safe

operational temperature, we may (for EPR) use the leakage

current curve in Figure 5 for scaling. It is found, that a

cable exhibiting 96 Mohmem at 1750C would show 7.5 Mohmem at

2000C and 0.24 Mohmem, the critical value, at 225 0 C. :The

cables used for obtaining Figure 4 would therefore still be

usable up to 2250C.

16
The two examples above may have involved cables of above

average quality. At Sandia. for some EPR power and control

cables, lower values of R have been measured than for these

examples. A conservative value for a safe temperature is

1900C. Permitting a minimum resistance of 0.24 Mohm-m at

1900C. scaling as described leads to a minimum resistance of

1.3 Mohmom at 1756C. Values in the neighborhood (but below)

this figure have so far been only observed on control cables

of one manufacturer. As control and signal cables work at

much lower voltages than the above discussed power cables.

thermal runaway is not a problem for the former. Nevertheless, addition of a leakage vs temperature test to qualification procedures may be helpful for selection of the best

cables.

Temperature increases up to 1900C (and probably 100 higher)

do not lead to dielectric failure in EPR power and signal

cables. For an XLPE configuration, the conductivity is an

order of magnitude lower than for EPR and the temperature

dependence is about the same (Ref. 15, Figure 4-2). The failure probability for this XLPE cable is therefore lower. Generally, it may be assumed (considering the wide range of

resistivities for EPR discussed above) that XLPE properties

are of comparable magnitude.

3. Comment on Circuit Failure

The present work deals only with primary cable failure, i.e..

complete breakdown of the cable proper. The above discussion

of leakage contains data, however, which permit useful comments on circuit failures.

The first phenomenon of concern is circuit starvation - loss

of so much current through leakage. that the load does not

receive enough power. A 100 m long circuit is considered.

At 480 V, 1900C, and the permitted 0.24 Mohmem resistance, the

circuit would lose 200 mA. which for a power circuit is unimportant. (At 175PC, a circuit using the cable of Figure 5

would only lose 0.6 mA.) For a signal cable operating at

50 V. the losses would be a tenth of the above values or less.

A 10% current loss would cause unacceptable misreadings, if

the information readout is current dependent.

Similarly important is crosstalk, particularly between power

circuits and control circuits. As stated above, a power circuit with a marginal cable may generate a stray current. However, installation regulations demand total spatial separation

of power and signal circuits; where this is the case, no problem exists. Signal-to-signal circuit crosstalk is mitigated

by lower voltages. Instead of the 200 mA of stray current in

the above example. 20 mA would be available. If a small part

of this (e.g., 2 mA) enters a signal line, the accuracy of

indication would suffer.

17
4. Other Phenomena

While radiation is applied, the conductivity of insulating

materials is enhanced due to additional carrier generation.

This phenomenon should be independent of temperature but

could conceivably cause excessively high currents during

accidents. Figure 6. taken from a paper by Murata et al.,*s

shows measurements on EPR cables aged in a variety of ways.

Resistivity varies by less than. a factor of 2. and is not

proportional to the dose rate. Extrapolating to a dose rate

of 10 Mrad/h, dose rates comparable to those expected in an

accident, indicates that a parallel resistance of more than

100 MO would appear across a 1-km length of cabling. At

higher temperatures, the effect would be overwhelmed by

thermal-carrier generation.

The influence of humidity on cable resistivity is observable

but small. Measurements by St. Onge et al. (Ref. 15, Figures

4-3 and 4-5) show an order of magnitude increase in conductivity for XLPE under extreme humidity (soaked cables) compared to dry conditions. A somewhat smaller increase is shown

for EPR. The measurements quoted above (e.g., Figures 4 and

5) include the influence of humidity.

Cables immersed in water, as might be the case after an accident. may exhibit a particular form of field breakdown caused

by water penetrating the cable along micro-fracture structures. A phenomenon named "treeing"''I. 2 0 develops when treelike fractures filled with liquid slowly extend into the insulators. This phenomenon is of importance at high field

strengths. Data are not available for low operational fields.

those smaller than 1 kV/mm for power cables. Treeing becomes

less pronounced at high temperatures, while the influence of

radiation is unknown. An experiment representing the probable

worst case has been in operation for 16 months in our facilities. In this experiment, ten 1-m-long cables have been soaking in water at room temperature under field strengths 5-times

higher than the operational level. No breakdowns have been

recorded.

For a cable with very high insulation resistance, another

breakdown effect has been reported in the literature.20 A

space-charge region may build up under irradiation, causing

high local fields and a discharge breakdown. In reactor circuits. the effect would be strongly mitigated by low circuit

impedance; it can be disregarded at higher temperatures, where

higher conductivity would cause the space charge to leak off.

5. Summary

The above discussion has been brief. A more comprehensive

study of basic electrical breakdown mechanisms and of literature data has been performed and many more preliminary tests

have been evaluated. 2' We believe it has been adequately

18
O%-os-- 0 Mrad NO

-b. 1100 HEAT

3 - -- M TOTAL 200 AGING

-4-- mmm" DOSE 00 0AFE AFTER

__ H00 HEAT

w 10 4 - 200 AGING

z

3 EPR

!1 3

z

10 2 , 0 0.1 0.5 1.0

DOSE RATE (Mrad/hr)

Figure 6. Effect of Dose Rate on Insulation Resistance (After Murata et al., Ref. 18)
shown that a purely electrical failure is highly unlikely up

to temperatures of 1900C. The discussion now turns to electromechanical phenomena.

20
III. TEST INSTRUMENTATION

1. Dedicated Test Facility

To permit long-time aging and creep measurements, a dedicated

facility for the present experiments has been established.

Figure 7 shows an overview, which includes seven temperaturecontrolled heat chambers, equipped with circulation fans and.

on the left, additional calibrated digital thermometers. A

common 480 V ac power supply, under the third chamber but not

clearly visible in Figure 7. permits application of an electrical stress field to any cable: breakdown of this field is

detected by short-circuit indicators. In each of the heat

chambers, six sections of cable are hung, under mechanical

tension, over 4.8-mm diameter cylindrical supports. An additional 40 unstressed samples were put into each chamber at

the beginning of the tests.

The center of the test area is occupied by three 10-m-long

iron pipes, each 50-mm in diameter and surrounded by 25-mmthick insulation (Figure 7). In each of the pipes, six long

cables are heated by passing from 20 to 30 amperes of dc current through the cable conductors. Thermocouples are positioned at 1.7-m intervals on the inside of each pipe: the

common dc power supply, and some of the surveillance instruments, are seen in the foreground of Figure 7. The heating

circuit, with adjustment resistors that keep the temperatures

in the pipes at 950. 1350. and 1650C, respectively, is shown

in Figure 8.

A rail, on which an instrument cart can be moved precisely to

any point on the heat chambers, is visible on the right side

of Figure 7. A theodolite and an X-ray camera are mounted

onto this cart (Figure 9). These instruments are used to

precisely align the cable tensioners inside the heat chambers

and. at regular intervals, to X-ray the cables to detect

deterioration. To simplify this, precise stop points are

marked on the rails, enabling the position of the optical

center of the theodolite or X-ray camera to be accurately

located.

2. Cable Tensioners

A cable-tensioning device, which is designed to permit precise creep measurements, is mounted in each of the heat

chambers (Figure 10). Six cable samples are positioned over

gooseneck-shaped supports and stretched by lead weights (visible at the rear of the mounting struts in Figure 10). For

three of the six cables, the maximum operational electric

field is maintained between cable wire and support by applying

the highest operational ac voltage. 480 V. The three shortindicating circuits are seen located on the right in Figure

10.

21
I

.5-4

.5.4

0

54)

'-f4

0

.5.4

0

ALI

I s,

4 4~ t

4

I

27

OA

I-I

22
+8.64V, 130A - TOTAL

PIPE I

95 0 C

PIPE III

1650C

PIPE II

135°C

1=

Figure 8. Schematic of the Heating Circuits

23
'Ali 4 AIM

I' IEW.

1%,

I

Figure 9. Photograph of the Instrument Cart
---' - -, , ý - Mm P Op 11 1 11 , - 11 1 . " . ý. IMPRIMOWNWAYAWMMO

I

• i•i • ii I iiii • i• i

I ý,

tN

1~

O

O

Figure 10. Photograph of the Tensioner
In the center of the horizontal part of each gooseneck structure (Figure 11), a 1-mm-diameter precision-drilled alignment hole is provided. By means of setscrews, the six holes of

each tensioning device are pointed to the optical center of

first the theodolite and later, automatically, of the X-ray

source. A curved film cassette can be attached to the support frame shown in Figure 10, with the film position indicated on the right in Figure 11. One X-ray picture will furnish a shadowgraph of the six cables on each tensioner; of

interest are the areas showing the gooseneck projection with

holes and the cable wires. An important quantity to be measured is the distance, y. between goosenecks and cable wires

(Figure 11). This "remaining creep distance" is extensively discussed below. To avoid the introduction of sporadic density gradients on the pictures, X-ray exposure, film development, and copy enlargement (a factor of 30) must be done very carefully. A mechanized installation in Sandia's Nondestructive Testing Department provides this service.

The calibration hole in the horizontal part of the gooseneck

support has a second purpose. It is needed not only for original alignment, but also furnishes a correction if, as is

unavoidable, the support distorts during many months of

stress, sometimes at high temperatures. Originally. the illumination source is located on or close to the axis of the

cylindrical hole, as indicated in Figure 11. In this case,

the picture of the hole will be a circle, and the projection

of the creep distance. y. will be correct within a few

microns. As the gooseneck bends down or up, the illumination

source will appear to move in the opposite direction, and the

rear or front end of the gooseneck will shadow off part of

the creep distance. y. The image of the hole boundary then

becomes nearly elliptical with a horizontal major axis. A

sideways motion of the support causes a tilt of the ellipse's major axis by some angle 0. From the ratio of the ellipse's

measured axes, and the measured angle 40, a correction for

the real or apparent decrease in creep distance, y, is found.

The analysis of the correction 2 2 is straightforward although lengthy. Up to now, the corrections obtained were generally less than 20 lim. For the present dimensions of the gooseneck support diameter (4.76 mm). the alignment hole diameter

(1 mm), and the distance to the light source (2 m). the correction for vertical distortion (angle 4 = 900) is shown in

Figure 12. For angles * less than 900, the correction quantity is determined by multiplication with sin 0; pure horizontal motion of the gooseneck does not require a correction

for y. In Figure 12, the expected "uncertainty" in the image

as seen in the X-ray pictures is also indicated. This uncertainty is due to the finite size of the present source (2 mm).

26
CABLE

CON~DUCTOR

lI -.3 D

d

SOURCE

NOT TO SCALE

F

Figure 11. Cross Section of Gooseneck Support and Cable
100 -

Source low

0

Uncertainty

050

10

0.6 0.7 0.8 0.9 1.0

Axis Ratio

Figure 12. Correction for Creep Distance as Function of the

Ratio Between Ellipse Axes (for cases where angle

(v = 90*)

28
3. Special Apparatus

For cracking experiments, cables have to be bent reproducibly.

Figure 13 shows a simple jig developed for this purpose. A

movable lever, connected to a vertical post, wraps the cable

around a cylinder 12-mm in diameter. It usually took about

10 seconds to bend the cable into a hairpin structure: uniformity of motion was attempted.

Located outside the test facility (Figure 7) is a setup to

test the "treeing" mentioned in Section II. Cables are placed

in a container partially filled with a saline solution and

connected to a power supply (Figure 14).

For meaningful loss measurements at operational voltages, a

special high-voltage impedance bridge has been developed.

Additionally, materials-shrinkage experiments were aided by a

simple overflow-volume measuring device, which used a fluid

consisting of an alcohol-water mixture to reproducibly wet

the cable samples investigated.

In summary, a considerable number of devices for holding,

bending, stretching, and cutting cables reproducibly were

developed and tested. particularly during the preliminary

experimentation phase of this investigation. Several of these

devices are described in greater detail in Reference 22.

29
Alt

w

Figure 13. Cable Bending Tool

Figure 14. Setup for the Treeing Test

30
IV. CREEP SHORTOUT

1. General Comments

A cable is stretched by an applied force. F. over a surface

with a radius of curvature. r. At elevated temperatures, the

metal wires will creep through the softened polymeric insulation and cable jacket: after some time, metallic contact

between cable wires or between the wires and the cable support

may occur. The present discussion is limited to the simplest

case where the cable consists of a single wire surrounded by

an EPR insulator and a Hypalon jacket.

The situation of concern involves large compressive strains

with values approaching unity. There is. however, nothing in

the literature concerning such large strains. As a result.

an essentially experimental investigation backed by computer

modeling was undertaken. To introduce basic concepts, an

attempt will be made to first draw conclusions from existing

pertinent data for small-strain situations.

The low-strain behavior of polymeric materials including rubber has been thoroughly investigated. 2 4 Even the long-term or

"aging" phenomena of these materials are subjects treated in

textbooks. 2 9 Unfortunately, most published work is limited

to temperatures below transition temperatures, which means

that only the lower regions of the temperature range of present interest are covered. (For EPR the transition temperature is not well defined, and the transition is not severe

since only a small percentage of the polyethylene incorporated

into the compound shows a crystalline melting point: the

"transition" lies between temperatures of 700 and 900C.26)

At low temperatures and low strain rates, creep is a linear

function of applied stress, a, and a function of a time

dependent "creep compliance." J(t), or its inverse, the creep

modulus, M(t). The creep strain, c, is given by

c =i a J(t) a/M(t) (1)

where a is assumed constant, and J or M is measured at a constant temperature, T.

For the same material, curves representing the time dependence of J. measured at different temperatures, starting

times, and stresses, are similar. Consequently, the curves

can be obtained from one another by translation of the horizontal or vertical or both axes, and by scale changes. Figure

15 shows a Bell Telephone Laboratories "master curve" obtained

by shifting many measurements to obtain overlay with a roomtemperature curve. 2' At room temperature and within the

validity range. Figure 15 predicts a strain of less than 4%

in 40 years for high-density polyethylene (HDPE) under a

31
T _ I I I I

10-1 DIFFERENT SYMBOLS REPRESENT

STRESSES RANGING FROM 0.3-5.5 MPa

AND TEMPERATURES FROM 23-1 10°C

00

a.

w

0.

0

10-2 -

i0~3 i

10-31B.

i fI I

1 DAY

40 YRS

0I

02

7

I I I II r II I I :,

102 106 108 1010 1012 1014

TIME (s)

Figure 15. Room-Temperature Master Curve for High-Density Polyethylene in Creep (After Aloisio and Brockway. Ref. 27)
stress of 600 psi (4.1 MPa). This confirms that creep is

exceedingly slow under most circumstances.

Instead of Equation 1. another method may be used to determine

creep strain. 1 ' An "effective viscosity" is defined as

n = dt/dJ (2)

when the plastic is considered a fluid, and Stokes' equations

are used to determine the movement of the boundaries to which

the stress is applied. The -n values obtained from Equation 2

are usually very high. For the example, the effective viscosities vary from 10o Pa*s at the left side to 10L" Paes on

the right side of Figure 15.

Attempts have been made to extend Equation 1 to strains of

several percent (Ref. 25. p. 157). Linear extrapolation

appears to be justified in a number of cases, but experimental

proof has yet to confirm this. The model using Equation 2.

however, appears to be more suitable for high-strain problems.

It is significant, however, that these models are useful only

up to a critical stress, 2 7 which has to be determined

experimentally.

Higher temperature creep data have been measured by St. Ongels

for crosslinked polyethylene and EPR; the detailed composition

of the materials, however, is proprietary. Figure 16 is the

creep modulus. M. plotted versus time for a number of compositions, with the material kept at a temperature of 700C. The

creep modulus versus temperature, after a 10-hour exposure.

is shown in Figure 17. The data pertain to compressive creep.

of interest here, and permit, with some extrapolation, a crude

order of magnitude estimate of the creep effects expected. At

700C. a representative containment temperature, an EPR sample

may have a creep modulus of 2.8 MPa after 10 hours (the worst

case in Figure 17). As will be discussed below, stresses of

2.8 MPa are quite reasonable if a 1-mm diameter wire is pulled

by a 0.5-kg weight over a surface whose radius of curvature

is a few millimeters. The resulting strain would be unity,

i.e.. the conductor would have cut through the insulation and

jacket. This conclusion is out of the validity range of

small-strain theory, but raises concern.

2. Preliminary Hiqh-Strain Experiments

A number of experiments were performed to obtain data on highstrain, behavior and at the same time to demonstrate actual

creep shortout. 2 2 The experiments will be summarized here.

a) Shortout Observation -- Cable samples were strung

over a supporting cylindrical bar. stretched by weights, and

exposed to various temperatures. Figure 18 shows the relative

33
0.

0

0.

ILIL

ILI

75 *

0

60

A

45 A

30

15 L

0-

10-1 100 101 102 103 104

TIME (hours)

Figure 16. Compression Creep Modulus for EPR and XLPE at

70 0C as a Function of Time (After St. Onge.

Ref. 15)

0.

0. wUIL

Ix)

100 -

' I' I ' I ' I ' I ' I

COMPRESSION CREEP (10 hrs)

* FILLED XLPE -

o XLPE (FULLY

CURED)

0 XLI'! (CABLE)

A EPR 4

A EPR 2

, I I , I i I i I % • I i I

10-8

10 -7

t0.

0.

0

10-

1

0 20 40 60 80 100 120 140

TEMPERATURE (VC)

Figure 17. Compression Creep Modulus for EPR and XLPE at

10 Hours as a Function of Temperature (After

St. Onge. Ref. 15)

34
sizes and shapes of cable and supports. A second supporting

geometry (sharp corner) is shown in Figure 19. For determination of stress from some known stretching weight, an effective

support area must be estimated. For the geometries in Figure

18, the effective area. A, should be proportional to the wire

radius, rw, and the curvature radius of the support, rs.

such that

A = krwrs (3)

where k is the proportionality constant.

For the corner geometry of Figure 19, the support radius rs is

replaced by the cable radius. rc. The proportionality constant, k. should be of the order of unity for both cases.

A detailed but preliminary estimate of stress distribution

and support stress for the above geometries has been developed

by Reaugh. 2 8 He finds that a square-root dependence of

stress on applied force better fits observations than the

generally used linear law. In Figure 20, Reaugh's solution

of stress versus support radius are plotted for two stretching

weights, 0.45 kg and 0.90 kg, and compared with the values

from Equation 3, when k = 1. It is seen that a value of

k = 0.3 would make the experimental approximation straddle

Reaugh's solution. This latter value for k will be used in

subsequent estimates.

For the preliminary tests, the stretching weights were unrealistically high, corresponding to 10-mi to 30-m of cable overhang. The chosen support radii of curvature were lower than

likely in an electrical installation. A considerable number

of creep shortouts occurred: they are listed in Table 2.

All cables tested were of the standard design used throughout

this investigation. The seven-strand copper conductor had an

average radius of 1.15 mm, while the EPR insulation and the

Hypalon jacket had thicknesses of 0.74 and 0.51 mm, respectively. In Table 2. the experimental arrangements are compared. Stresses estimated from Equation 3 with k = 0.3 are

included, along with temperature history and the number of

experiments for each exposure sequence. Electrical stress

was not continuously applied. Creep shortouts were measured

with a low voltage ohmmeter. In a number of cases, the jacket

was still present at the end of the exposure, although somewhat reduced in thickness, while the insulator was totally

squeezed out: these cases are identified in the last column.

In the test listed on the bottom row of Table 2. the heat

spike indicated was 300 0 C applied to the specimen for one hour

before the test. The compilation in Table 2 demonstrates that

at high temperatures (1750C) and high stresses (2 MPa or about

300 psi), shortouts occur readily and quickly. It is difficult to scale the measurements to lower stresses and longer

35
10:1

(9

b) a)

Figure 18. Relative Sizes and Shapes of Cable and Supports

W

CABLE BENT AROUND CORNER

AND STRESSED BY WEIGHT W.

4

Figure 19. Corner Geometry and Force Distribution

36
5

oN

a. 0.5 -

CO b

II

0.4 1.2 2.4

SUPPORT RADIUS (mm)

3.6

Figure 20. Stress on Support Surface vs Support Radius
Table 2

Creep Shortouts

Nominal Exposure Number Number Number

Experiments Stress Temperature Time of of of

(MPa)* Exposure (days) Experiments Shorts Squeezeouts

(a) rs = 1.2 mm 2 1900C. Constant 2.2 2 2

rs - 1.2 mm 0.3 to 1900-1750C, 2.2 10 0

to 3.6 mm 1 Constant

(b) rs = .4 mm 5 1750C. Constant 21 2 0 2

rs = .4 mm 5 1900C, Constant 1 12 0 4

(c) Corner Geometry 2 0-2250C. Rising 5 10 10

0.7 0-225 0 C, Rising 5 20 1

0.7 0-2250C. + heat

spikes 3 10 2

*No electric field

1 psi = 6900 Pa

"It
times, taking into account the aging processes affecting the

cable.

b) Deformation Measurements -- At the end of the exposure, the cable samples were carefully cross-sectioned along

the plane of the most severe deformation. The remaining

thicknesses of insulator and jacket were then measured; Figure

21 shows a cut through a new cable and through a sample

exposed to temperatures of 125 0 C and nominal stresses of 2 MPa

for 1 day. The cuts were made after cooling to room temperature. It is seen that deformation is nonuniform, asymmetric

and, in this example, accompanied by unraveling of the wire

strands. (It is believed that the cutting process itself did

not contribute noticeably to these distortions.)

For this investigation, the minimum thicknesses of remaining

insulator and jacket are of interest. Not surprisingly, considerable scatter was found in the measured values. A main

contributor to the measurement variations is the relative

position of the wire strands with respect to the support

structure. This is demonstrated in Table 3. It is seen that

the average thickness of the composite layer (insulator plus

jacket) may vary by a factor of almost 3, depending on whether

a single strand or two parallel strands lie opposite the support and take up the weight of the cable.

It is evident from Table 3 that the cable insulator yields

much more easily than the jacket. Insulator squeezeouts are

relatively common while squeezeout of the jacket is rare.

This fact complicates the detailed investigation of cable systems.

An illustration of the time dependence of creep is presented

in Figure 22. For two different temperatures (1250C and

190 0 C), the thickness of the two plastic layers is plotted

versus exposure time, with the nominal stress as a parameter;

each point represents an average for two experiments. In the

beginning the thickness is reduced very rapidly, because the

contact area between cable and support is at first very small

and the stress is very high. As exposure time approaches one

day, the contact area approximates the nominal value of Equation 3 and creep rate decreases. At 190 0 C, the bottom diagram

in Figure 22. a shortout of the two cable samples has occurred

after just over two days; the two cables were in the "unfavorable" position indicated in-Table 3 (a).

c) Conclusions and Comments -- From the measurements, it

becomes clear that simple small-strain linear theory does not

hold for the present extremely large-strain phenomena. In

Figure 23. composite creep compliance has been calculated as

a function of nominal stress for a 24-h exposure. The compressive creep compliance decreases strongly with increasing

stress in contrast to the behavior of tensional or torsional

creep compliance. It is not clear at this time what the

39
(a)

(b)

Figure 21. Cut Through a New Cable (a) and a Cable Stretched by 0.9 kg at 125-C for I Day (b)

40
Table 3

Position Dependence of Deformation

77111111771

SUPPORT

F

Thickness (mm) Strain Composite

Insulator Jacket Both Insulator Jacket Strain

Original

Thickness 0.74 0.51 1.25

Remaining 0 0.07 0.07 1 0.86 0.94

Thickness 0 0.09 0.09 1 0.82 0.93

(a) (mm) 0 0.10 0.10 1 0.80 0.92

Avg: 0 0.0867 0.0867 1 0.83 0.93

Remaining 0.01 0.20 0.21 0.98 0.61 0.83

Thickness 0.05 0.21 0.26 0.93 0.98 0.79

(b) (mm) 0.06 0.23 0.29 0.92 0.55 0.76

Avg: 0.04 0.147 0.253 0.95 0.71 0.79

Remaining

Thickness

(c) (mm)

Avg: .0020 0.117 0.170 0.97 0.77 0.86

I-.

/77777777777

SUPPORT Nominal-stress: 14 MPa

Exposure time: 1 day at 1906C
LL

0

cl)

UJ

~6

040

+ 3

z

020

4 1

U)

(fl 0 8 24 48 72 96

EXPOSURE (HOURS)

LL

0

U)I

+

60

50

40

30

20

10

0

0 24

EXPOSURE(HOURS)

4852

Figure 22. Creep Measurements at 1250C and 1900C

42
0.

LUI

()

0.

IL

2

(0

0

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.0

I I I I I I

0 0.45 kg WEIGHT

X 0.90 kg WEIGHT

C] 0.86 kg WEIGHT K 0i.

a0

0

0*

0

6

% %STRAIN =1

40

1900C d.v'%e'

0

see,

24 HOUR EXPOSURE

1250C

b 190 0C

I I I I

I 2 3 6 14

NOMINAL STRESS (MPa)

Figure 23. Creep Compliance vs Applied Stress for Various

Exposure Temperatures

43
exact cause of the compliance decrease is (perhaps strain

hardening or compacting). The phenomenon is beneficial

because the lower compliance delays shortout.

Another probably related, and also helpful, phenomenon concerns breakdown fieldstrength. With the insulator and jacket

thicknesses decreasing, the electrical breakdown voltage

should decrease. This does not appear to be the case. As

Figure 24 shows, breakdown voltages measured for 18 cables at

225 0C do not vary significantly with the thickness of the -

remaining layer. The explanation for this may be that mechanical compression near the support improves the insulating

material (e.g., by removing bubbles, closing microcracks. and

densifying the polymers). Using an experimental facility

developed by K. Gillen,26 a 5% density increase was measured for EPR samples removed from a location just above the

support point. Cable breakdown in the support region therefore becomes less likely with increasing exposure: breakdown

along flaws at other locations remains unaffected.

A third phenomenon further decreasing the breakdown probabil- ity has been observed. Over long times, and probably affected

by vibrations, the "unfavorable" wire strand position (shown

in Table 3. (a)). appears to be unstable. The cable tends to

twist towards the more stable position indicated at the bottom

of Table 3. The wire now sinks in more slowly, and shortout

at the support is delayed.

This work demonstrates the complexity of creep shortout.

Long-term extrapolation from short-time tests is difficult

under the circumstances.

3. Systematic Experiments

As a consequence of the considerations in Sections I and II

and backed by results of the preliminary tests, a long-term

experimental effort has evolved, which is continuing. The

part pertaining to creep phenomena, particularly creep shortout, will be discussed here.

a) Program -- In the main test facility described in

Section II, 6 experiments are being conducted in heat chambers maintained at 750, 1000. 1250, 1500. 1750. and 225 0C. A

seventh experiment is taking place in the open at room temperature. In each test, creep versus time is measured by X-ray

observation of the gradual reduction of insulator thickness

in 6 cable samples. Each cable is hung over a 4.8-mm diameter

steel rod and the cable-ends carry 0.45-kg lead weights. For

three of the six cables in each test, an ac voltage of 480 V.

between cable wire and support, generates an electrical field

tending to pull cable wire and support together. The small

mechanical stress of about 3 Pa (0.4 x 10- psi), which

increases rapidly with decreasing distance, induced by the

44
I I

w

,1

I-. -,J

0

2

0

ILl

QZ

W

0

NO

SDEFORMATION

1 0.q- 0

0 0

0

0

N 0

SLEEVED CABLES

WIREHANGER Ii, 225°C

WIREHANGER III, 225°C

WIREHANGER IV, 2250 C + HEAT SPIKE

2250 C

1.01-

I I

0.1'

0.01 0.1 1.0 10.0

REMAINING THICKNESS (mm)

Figure 24. DC Breakdown Voltage vs Remaining Polymer Thickness at 225 0 C

45
voltage also causes 120 Hz mechanical vibrations. (Figure 10

shows one of the seven tensioning arrangements. Alignment

and correction methods for the measurements have been

described in Section III.)

b) Measurements -- Figure 25 shows an X-ray negative of

the six bent cables on one of the cable tensioners. One of

the bends is magnified five times in Figure 26 (positive) to

show more clearly the calibration hole and the support area.

The scalloped appearance of the wire shadow is due to the

twisting of the seven cable strands, which leads to the effect

tabulated in Table 3. In Figure 26a, the positioning of the

cable on the support is symmetrical, the desired situation.

In Figure 26b. the cable has slipped to the side and the

shadow of the wire at the support point makes an angle, a,

with the horizontal.

All measurements were made from negatives at a magnification

of approximately 40. Figure 27 is an example. The projections of the thicknesses of the EPR and of the Hypalon were

measured on a line drawn perpendicularly to the wire shadow.

indicated by the line with arrow; this line is at an angle,

a. with the vertical as explained above. The accuracy of

the measurements is limited by the fuzziness of the boundaries. With experience, repeated measurements became reproducible. About 5% of the pictures were computer digitized, and

a density trace was made along the measurement line (Figure

28). From these, thicknesses could be more accurately found.

These precision measurements correlated to the standard measurements within 0.5 mm. which is equivalent to 12 Um in the

actual cable samples.

Measurements were tabulated in the form shown in Table 4. The

sample data in Table 4 were taken after 2-mo exposure at

1750C. The first nine rows of the table contain raw data and

the other rows are computer evaluations as follows.

Rows 2. 3. and 4 show the thicknesses in mm for the

EPR and Hypalon layers, and the independently determined total insulation thickness, measured on the

negatives similar to Figure 27.

0 Row 5 lists the tilt symmetry angle, a, introduced

in Figure 26. for possible later use in correlation.

Rows 6. 7. and 8 use the support width. d. defined

in Figure 27, and the major and minor axes of the

elliptical projection of the calibration hole. Support width is known to be 4.76 mm. and the major

hole diameter is very close to 1 mm. The two quantities are therefore useful to obtain scaling factors. In row 6 the scaling factor is printed out

directly.

46
J-\w Jlq\w 2) iJ-\WE 3j-kmc JJW

I. 210501

ý. q-l

Figure 25. X-Ray Negative of Tensioner
ure a)

b)

Figure 26. Symmetric (a) and Unsymmetric (b) Hang

r

48
w WIRE

IYP4 -7

0 /

d

No. 1

Figure 27. Measurement Parameters (38 x)
(31 0

Figure 28. Density Distribution
Table 4

Example of Data Evaluation

U'

Raw Data 1750C. 2-mo exposure

i Cable Number 1 2 3 4 5 6

2 EPR, measured (mm) 14.0 12.0 12.0 14.5 13.0 11.5

3 Hypalon, measured (mm) 9.0 10.0 10.0 10.5 10.5 9.0

4 Total, measured (mm) 23.0 22.0 22.0 25.0 23.5 20.5

5 Alpha. (degrees) 9.0 1.0 1.0 1.0 1.0 0.0

6 Support thickness (mm) 39.5 39.5 39.5 39.5 39.5 39.5

7 Large hole diameter (mm) 38.0 38.0 38.0 38.0 38.0 37.5

8 Small hole diameter (mm) 35.0 37.5 36.0 38.0 37.0 35.0

9 Phi. (degrees) 15.0 35.0 20.0 0.0 50.0 55.0

Computer Evaluations

10 Kappa (Um) 23.7 3.9 15.8 0.0 7.9 20.0

11 K-sin(phi) 6.1 2.3 5.4 0.0 6.0 16.4

12 EPR (Nm) 354.4 303.8 303.8 367.1 329.1 291.1

13 Hypalon (um) 234.0 255.4 258.6 265.8 271.9 244.2

14 Total (um) 588.4 559.2 562.4 632.9 601.0 535.4

EPR Hypalon TOTAL

15 Average (um) 324.9 255.0 579.9

16 Std. dev. (Nm) 28.0 12.7 31.7

17 Deviation (%) 8.6 5.0 5.5

18 Reduced. EPR (%) 44.9 38.7 36.3 39.7 39.4 34.2

19 Reduced. Hypalon (%) 29.7 32.5 30.9 28.8 32.5 28.7

20 Reduced. Total (%) 74.6 71.3 67.2 68.5 71.9 62.8

21 Average. E.H.T (%) 38.9 30.5 69.4

22 Std. dev.. E.H.T (%) 3.3 1.6 3.8

23 Deviation (%) 8.6 5.3 5.4

Group Averages

24 EH.T, 1..3 (%) 67.9 75.3 70.9

25 E.H.T. 4..6 (%) 64.5 72.4 67.7
Row 9 gives the values for the measured tilt angle,

a. versus the vertical, indicated in Figure 27.

* A computer program calculates corrections and averages of interest from the raw data. (The digit

behind the decimal point in rows 10 through 14 in

carried only for program convenience and is not

significant.)

• Row 10 determines the support shadowing correction

(see Figure 12) from the minor to major ellipse axis

ratio. i.e., from the data in rows 6 and 7. The

value is shown in microns.

* Row 11 improves the correction by taking the tilt

angle (row 8) into account.

* Row 12 shows the actual thickness of the EPR layer

in microns. calculated by dividing row 2 by row 6.

* Rows 13 and 14 list the corrected thicknesses for

the Hypalon and for the total polymeric layers. As

the lower part of these layers may be partially

covered by the support, the correction calculated in

row 11 is added.

* Row 15 provides averages for the 6 thickness measurements in rows 2 to 4 in microns. The EPR values

are listed in column 2. Hypalon and total values in

columns 4 and 6, respectively.

• Rows 16 and 17 contain the standard deviations for

row 15, in microns and percent, respectively.

• In rows 18 to 23 are listed the normalized values

for the data presented in rows 12 to 17; for each

cable the total thickness before exposure is defined

to be 100%. The normalization will be discussed

below.

* Finally, in rows 24 and 25. reduced group values are

listed. These help in assessing differences between

the three cables under voltage and the three nonenergized cables. Here the preexposure thickness

for each of the layers is set at 100%.

The reason for introduction of normalized values is that the

original dimensions of the cable samples, although taken

sequentially from the same reel, vary severely. Bending the

cables over support rods increases the variation. In Table

5, some statistical data of interest are shown. The first

three groups of figures show before-exposure averages for all

42 cable samples used in the creep experiments: thicknesses

and standard deviations are measured in microns and the
Table 5

Proexperiment Averages

All EPR

Standard deviation. EPR

Standard deviation (%)

All Hypalon

Standard deviation. Hyp.

Standard deviation (%)

All cable

Standard deviation

Standard deviation

Reduced thickness.

Standard deviation

Standard deviation

Reduced thickness.

Standard deviation

Standard deviation

552.8

55.3

10.0

328.7

35.2

10.7

881.6

69.4

7.9

62.7

3.4

5.5

37.3

3.2

8.6

(%)

EPR. (%)

(%)

(%)

Hyp.. (%)

(%)

(%)
relative standard deviations in percent. For a population of

42, a standard deviation of 10% must be considered large.

Normalization of the total thickness to 100% eliminates the

standard preexposure deviation for this parameter; as the two

bottom groups in Table 5 show, the deviation for the EPR

insulator is halved, and the deviation for the jacket somewhat

improved. The original dimensions are therefore not distorted

proportionally. This points to nonuniformity of composition

or of processing and makes extrapolation more difficult.

c) Preliminary Results -- This investigation attempts to

extrapolate a few years of data to the 40-yr service life of

cables in a reactor containment. A suitable form for discussing such an extrapolation is the log-log plot of remaining

thickness versus times shown in Figure 29. The indicated

power curve is a straight line which conveniently drops uniformly from 100% to 1% in 40 years. A reduction of the total

insulation thickness to 1% means that about 10 jIm of

insulator/jacket material would remain between wire and support. From Figure 24, it is seen that a breakdown at this

thickness is unlikely even for a potential difference of several kilovolts, and much more unlikely at operational voltage.

If the measured thickness stays above this line, we may assume

that creep shortout will not occur. In Figure 29, the remaining cable thickness, averaged over six cable samples, is

plotted for the first few months of data. Some measured

points are connected by curves. We expect thicknesses to

gradually decrease as the EPR creep-modulus curves in Figure

17 indicate. This is usually observed. A slightly positive

slope sometimes occurs as, for example, for the 756C values.

This is probably due to twisting of the cable sample, as the

wire strands seek a stable position. Notice that the greatest

reduction in thickness occurs for the samples exposed at

1250C. Such a worst temperature is expected because the materials first soften with temperature, but enhanced crosslinking at still higher temperature embrittles the polymers

and retards creep.

In Figure 30 (a worst case) the thickness reduction situation

is presented. Instead of the average, the cables with the

smallest remaining thickness (out of the sample of six) have

been selected. Again, the smallest remaining thicknesses

occur for the 125 0 C experiment. The measurements do not tend

to approach the critical curve defined above.

Creep compression of the EPR and Hypalon layers are not proportional; EPR is softer and contributes most of the combined

creep strain. Figure 31 shows worst-case EPR thicknessreduction percentages. The 225 0 C values have the tendency to

approach the critical curve. An explanation is that the

squeezeout of the insulator (see Table 3) may be very rapid

in the beginning before embrittlement slows the process.

54
l100 _ _ _Cables)

L225°C- -- 7 5 oc25°C o 2SOC m A Imý1OO°C • 75 0C

- - 1250C o 1OO°CCO a 1250 C -

(1) A 150°OC

LUI z A 1750C

od 0 225°C--

C10

cc 40

year

1 100 10 1000 10,000

EXPOSURE (DAYS)

Figure 29. Average Reduction for 6 Cables
100

90- o o o

~80-

z 7 0 0

00

M 60 0

a.50 *

0*U 

w 40

z

2 30

z

20 100%x t 0 .4 8

1010 30 60 90 12

EXPOSURE (DAYS)

!0150 210

Figure 30. Thickness Reduction for Worst of 6 Cables
z

Ui

%.0

CO)

CO)

w

z

100-

90-

80-

70-

60

50

40

30 -

2'0

10

10

Lfl

30 60 90 120150

EXPOSURE (DAYS)

Figure 31. Worst Thickness Reduction for EPR Layer
With time. the polymeric materials tend to become more and

more embrittled and the creep modulus increases; the creep

curves in Figures 29 to 31 will flatten out and bend away

from the critical curve. This leads to the expectation that

at the tension load used in these measurements, shortout will

not occur. This will be true if compression creep does not

exhibit the nonlinearities well known for tensional and torsional creep. While this is physically unlikely, experimental

proof is still required.

The above argument pertains only to actual, creep shortout

where all polymeric material has finally been squeezed out

between wire and support. Other phenomena may occur for this

geometric arrangement. It was observed in the 225 0 C test that

the three cable samples under 480-V leads developed temporary

electrical shorts, while severe cracking in the support neighborhood occurred. This situation will be discussed in Section V.

58
V. CRACKS IN CABLES

1. General Comments

Cracks in the insulator-jacket protection of a cable may lead

to electrical shortout. especially during an accident when

the combination of humidity, contaminants, and spray may form

conducting paths to ground or to another crack.

Experiments have been performed during this investigation

that show that the existence of a crack through jacket and

insulator does not always lead to an immediate shortout. even

if the cable is immersed in a conducting liquid. The conditions under which a short does not occur, even though cracks

are present, are very complex and not completely understood.

For the present discussion a criterion will be that cracks

penetrating to the inner conductor may cause a shortout and

are considered as a "failure".

Three specific situations are of primary concern:

a) Long cables lying in conduits or trays may develop cracks

under thermal cycling; it is believed that shrinkage of

the polymers with temperature and time is associated with

the cracking.

b) Cables, sharply bent. (e.g., around conduit corners) may

crack under the mechanical stress produced by temperature

cycling or by overhang weighting.

c) Cables, embrittled by aging. may be mechanically stressed

and cracked by "maintenance" procedures.

Variations and combinations of these characteristic situations

may occur.

The investigation of cracks in cables is made difficult by two

phenomena. First, a crack will originate at a surface flaw.

and the distribution of surface discontinuities is generally

widely variable: once a crack starts enlarging, it relieves

tensional stresses in its neighborhood and prevents further

cracking there. Second. cracks, especially small ones, can

heal again. 2' Measurements of number and size of cracks

occurring for a certain temperature history and stress distribution will show large variation.

Investigations at the Bell Telephone Laboratoriess° have

shown that for LDPE cables the time needed to develop incipient cracking, plotted versus inverse temperature, exhibits

non-Arrhenius behavior. It cannot be assumed that the present

materials combination behaves more simply.

59
2. Preliminary Tests

a) Long Cables -- It is of interest to learn if and when

long cables, lying undisturbed in a conduit and under a minimum of stress, can develop cracks. In a pipe arrangement

similar to the one shown in Figure 7 and discussed above, a

bundle of six, parallel. 10-m-long cables were gradually

heated over a 2-h period by cable current to a temperature

slightly above 300 0 C. The test simulated a severe overload

condition. After cooling it was seen that the cables clamped

at the ends of the long pipe had warped and some had crossed

over the others. Both perpendicular (circumferential) and

longitudinal (axial) cracks were found (see Table 6). The

perpendicular cracks penetrated through both jacket and insulator: the axial cracks occurred in the jacket only. A subsequent test at a much lower temperature of 90 0 C, but otherwise identical, did not produce any cracks in a 10-month

exposure.

Perpendicular cracks are not easy to explain. For a crack to

occur, a tensional stress perpendicular to the direction of

the crack must be present. This stress cannot be generated

by the thermal expansion of the steel pipe. as the expansion

coefficient for both copper wire and for the polymeric insulation are larger than the coefficient for steel. The cooling

cycle cannot be responsible, since it was observed that at

least a part of the cracks occurred before the system cooled

down. A possible explanation may be the shrinkage of the

polymeric materials.

A number of experiments with a specially developed precisionvolume measuring device demonstrated the following behavior

for cables exposed to increased temperature. The polymeric

material's specific volume increased due to the the positive

thermal-expansion coefficient of the polymeric material (see

e.g.. Ref. 15, Figure 3-7). After hours or days in the temperature range of interest, loss of material due to chemical

decomposition decreases the volume. The final volume loss

increases with temperature, approaching a saturation value

with time; a 15% decrease was measured for a several-day

exposure of the cable to 185 0 C.

Assuming a (worst case) three-dimensional uniform shrinkage

of 20%. a strain of less than 0.07 would be generated.

According to Equation 1, using a high-temperature modulus of

2070 kPa (Ref. 15, Figure 5-3). a stress of 145 kPa would

exist, which could gradually open a crack. The present geometry does not have three-dimensional symmetry, however. In

fact, for an Infinitely-long uniform cable no axial tensional

stress would develop. As severe perpendicular cracks do

occur, an axially nonuniform stress distribution must exist.

Bends in the cables do provide asymmetry. A more likely condition is nonuniform "sticking" of the cable insulator to the

60
Table 6

Crack Statistics After Heating to 2256C Over 5 Days

Experiment I II III IV

Number of cables 10 10 10 5

Load per cable (kg) 1.6 0.34 0.34 0.34

Environment Dry Dry Humid Humid

(Heat Spike)

Average number of

cracks per cable 4.6 1.9 2.3 1.6

Standard deviati6n 40% 34% 21% 154%

Average width

of cracks (mm) 0.9 1.1 1.6 0.1

Standard deviation 27% 51% 42%

Integral crack

width (mm) 4.2 2.1 3.7 0.16

(Est.)

Insulator cracks

(percentage) 77% 35% 67% 0

61
cable wire; this phenomenon has often been observed when

cooled-down cables are disassembled.

Cracks parallel to the axis are generated by circumferential

stresses. Measurements have shown that the outer Hypalon.

jacket tends to shrink more than the inner EPR material.

This enhances the circumferential stress and may explain why

parallel cracks almost always occur only in the jacket and do

not penetrate the insulator.

b) Cracks in Stressed Bends -- In this test, 35 cable

samples made of EPR/Hypalon were exposed to a thermal treatment in which the temperature was slowly increased over a

period of 5 days to 2250C in steps of 250C. The cables were

stretched over a sharp edge with weights of either 1.6 kg or

0.34 kg each. Half the experiments were performed in a dry

atmosphere; for the other half, the humidity ranged between

95% and 100%. In the last experiment, a one-hour heat spike

of 2650C was applied before the normal heating sequence.

Listed in Table 6 is the average number of observed cracks.

the average width of the cracks, and the integrated crack

width per cable. Also given is the percentage of cracks which

reached the conductor (through cracks). It was observed that

the formation of cracks increased strongly between 2000C and

225 0 C. The cracks were formed and counted before the system

had time to cool noticeably. Some cracks may have deepened

in the subsequent cooling period. One crack in experiment III

was found to have healed after cooling.

The most severely stretched cables, those in experiment I.

clearly exhibited the largest number and the highest total

width of cracks. It is not clear why the average crack width

was very low in this experiment. The presence of humidity

during the relatively brief aging experiments does not seem

to have noticeable influence. The addition of a heat spike,

simulating a temporary overload, had strong impact. It

resulted in very tiny cracks (or crazing) which did not penetrate to the conductor. This inconsistency (and some of the

observations above) may be related to the adhesion of the

insulator to the center wire. In the first three experiments.

the three cable components (jacket, insulator, wire) stuck

inseparably to each other after cooling. The heat spike in

test IV led to easy separability of the three.

c) Maintenance Accident -- The many data in literature

describing the influence of temperature and radiation exposure

on materials properties have almost always been taken after

cooling to room temperature. The results therefore may not

be descriptive for what happens at high temperature, but

should be applicable for correlation with crack generation at

room temperature, i.e., during "maintenance handling" of aged

cables. A "maintenance accident" was simulated by bending

cable samples around a mandrel. The mandrel diameter was

62
usually approximately equal to that of the cable; this certainly constitutes a severe overtest.

Cracking will occur if the breaking strain or "elongation to

rupture". e. of the deteriorated polymer is exceeded. A simple two-dimensional model for estimating strains on the outer

surface of a bent cable can be developed with the help of

Figure 32. A strip of cable material of width 2p is bent

(without deformation of the cross section) over a cylinder of

radius r. At the outer side of the geometry, where cracks

can start developing, the length of the plastic around a quarter bend is w(r + 2p)/2. At the neutral fiber, the corresponding length is ir(r + p)/2. The resulting bending strain

is then p/(r + p). This strain is augmented by a/M, if an

applied stress, a, exists, where M is the appropriate modulus.

Cracks will develop if the sum of these strains exceeds or

equals the breaking strain, e, that is. if

e + (4)

In experiments performed with cables and not with strips of

material, pure bending without pulling was attempted with a

jig similar to the one in Figure 13. Stress. a, was taken

to be zero. It was assumed that the effects caused by potential sticking of the polymeric material to the center conductor were negligible. Lateral deformation of the geometry

(usually less than 10%), was also neglected. The model cannot

take into account the (ever-present) flaws in the material;

presumably, the value of e accounts for this on a statistical

basis. Considerable scatter in the measured data is to be

expected.

For the cable used, p for the Hypalon jacket has an average value of 2.4 mm, and for the insulator, 1.9 mm. If the

cable is bent over a 4.8-mm-diameter mandrel (r = 2.4),

strains of 0.5 and 0.4, respectively, are created. It is

customary to normalize strain values to the original roomtemperature breaking strain. eo, which is about 3.5 for

these materials. Cracks should therefore become likely when

e/eo equals 0.15 for the jacket (Hypalon) and 0.10 for the

insulator (EPR).

In Figure 33, the time needed for EPR and for Hypalon to reach

the values e/eo = 0.10 and 0.15. respectively., is plotted

versus inverse temperature. The data are extrapolated from

Reference 31 and supported by Sandia measurements on similar

EPR and Hypalon compositions. 2 6 The extrapolation is justified in Reference 22. Five standard cables were aged at 1860C

for 8, 16. 32. and 48 hours and for longer periods. As indicated in Figure 33. the jacket exhibited cracks after 32

hours. The insulator did not crack until after 300 hours.

The experimental results are certainly compatible with the

model developed.

63
NEUTRAL

FIBER

o eI

Figure 32. Model for Equation 4

TEMPERATURE (C)

225 186 147 108

I I - BENDING

CRACKS A

A

100

U

w

0

a. 10

wu

/ JACKET:

O NO CRACKS

I CRACKS

(1 DIAMETER

BENDS)

* CRACKS

(HALFINCH

MANDREL)

INSULATOR:

* NO CRACKS

A CRACKS

2.0 2.1 2.2 2.3 2.4 2.5

INVERSE TEMPERATURE: (1000/[T+2731)

Figure 33. Crack Experiments

64
Some earlier observations for a 225 0 C exposure are also indicated in Figure 33. During the experiments I through III

described in connection with Table 6. straight as well as

prebent cables were exposed to temperature treatment with no

axial stress applied. The cables remained a few hours at

2250C and were then bent (or unbent) after cooling. The bending radius was 6.4 mm in these experiments. According to

Equation 4. cracking should occur for e/eo = 0.08 (Hypalon)

and 0.06 (EPR). The graphs for these values would lie just

above the respective plotted lines in Figure 33. Jacket

cracks were observed as shown. The insulators did not crack

because the exposure did not last long enough.

The results indicate that a basic understanding for cold

cracking of cables exists, at least if embrittlement has been

achieved at high temperatures and therefore in short times.

For extrapolation to long-time behavior, more precise and

more systematic tests are needed.

The model based on Equation 4 may be useful to estimate the

embrittlement ratio e/eo. above which no cold cracks will

occur. For a very sharp bend (kink), the bending radius r

will be very small and the ratio p/(r + p) will tend towards

unity. With e -1. the limiting value e/eo for the materials

used in the experiments described above will be about 0.3. A

very conservative working hypothesis would therefore demand

that e/eo be kept above 0.4 to avoid cold bending cracks.

In the present experiments, no cracking has been observed

above this value.

3. Ongoing Systematic Tests

a) Dynamic Cracking -- The test series presently in operation makes use of available space in the heat chambers used

for the creep tests described in Section IV. Cable samples

are being aged at 750 to 175 0 C in intervals of 25 0 C. and at

2250C. Every month, and semimonthly for the two highest temperatures, the heating cycle is interrupted and the cables

are permitted to cool down so that samples for crack tests

can be withdrawn. The samples consist of 0.2-m-long straight

pieces of cable lying on the bases of the creep tensioners

(Figure 10) and/or on the floor of the heat chambers. In

spite of fan circulation, small heat gradients exist on the

heat chamber floor and the samples are not all at precisely

the same temperature. At the end of the test series, the

straight vertical pieces of the creep test specimens will be

used for crack tests.

After removal the samples are bent in the jig shown in Figure

14, and the cracks are counted. There are always many more

surface cracks observed in the Hypalon layer than cracks

extending through to the wire (through cracks). Also, perpendicular cracks outnumber parallel cracks. Data obtained

up to now are plotted in Figure 34. Groups of 6 samples were

65
taken from the heat chambers at the indicated times, bent and

analyzed. Shown in the graph are the number of through

cracks, with perpendicular cracks in the top graph and parallel cracks at the bottom. Figure 34 shows that, for 2250C

exposure and bending over a 12.7-mm mandrel, through cracks

appear before one month of exposure. At 1750C, through cracks

appear at about a month. The number of cracks grows rapidly

with time. Data for the 1500C exposure are still too few to

exhibit a clear trend. At still lower temperatures no cracks

have yet occurred. As expected, the scatter in the data is

substantial.

The cracking model predicts that an e/eo ratio of 0.04 or

less is needed for cracks to start in the EPR layer. i.e., for

through cracks. In Figure 33, the e/eo = 0.04 line would

lie somewhat above the plotted 0.1 curve. The starting times

for cracking at 2250C and 1750C taken from Figure 34 are compatible with the model plot in Figure 33.

b) Static Cracks -- As reported in Section IV. all 225 0C

creep experiments under voltage shorted out temporarily after

an exposure of 2.5 months. The X-rays (see example in Figure

35) show striations not present in other pictures; in addition. the Hypalon layer has lifted off the support. The striations appear to be associated with cracks.

Inspection of the samples reveals that after the long, hightemperature exposure the polymers are completely embrittled.

Temperature cycling, mechanical vibration naturally present

and caused by the ac field, and asymmetric stress distribution lead to cracking. It is suspected that during the cracking process vapors and fluids in polymer pockets are freed

and permit a surface breakdown discharge3 2 along the crack

walls. The discharge cleans the contamination and the insulator is functional again. Further observations are needed

to verify this explanation. Repetition of the 2250C experiments is planned.

66
70

U-.'

C C 60

00

2.40

-I

cc~ 3

Ow

IL2w

0Q10 Icc

I-t0

2 3 4 6

EXPOSURE (MONTHS)

(0

045

40

M4

rz

<0-

40

70

60

60

40 FII I I I

. 150oC

U *1750C

A 225°C

&U

A

I -- I I I I

301-

20-

101-

0 A 0

1 2 3 4 5

EXPOSURE (MONTHS)

Figure 34. Total Number of Bending Cracks for 6 Cables

vs Time

67
* SL

Figure 35. Embrittlement Cracks at 2250C
VI. CONCLUSION

The work is planned to continue, to generate more accurate

bases for extrapolation and to obtain enough data to develop

some scaling laws. Some preliminary results of the investigation can be given if the following assumptions are made:

a) There are no severe inherent flaws (e.g., metal chips) in

the organic material.

b) No synergistic effects, other than those already observed

in materials aging, will be found.

c) No chemicals other than water or watery electrolytes are

present.

d) No strong mechanical vibrations will occur to shake

embrittled insulation off the cables.

e) Cables are operated within current and voltage specifications.

Under these assumptions, operability approaching 40 years is

to bo expected for EPR - Hypalon cables, with the following

limitations:

f) The cable temperature does not exceed 1900C (above this

temperature thermal runaway may occur).

g) Cable overhang over edges with 3-mm radius of curvature

must not be longer than 10 m and, for 6-mm radius of

curvature, not longer than 30 m (creep shortout critical

dimensions).

h) High temperatures do not last so long that cracks form in

mechanically unstressed cables (long-exposure crack

threshold).

i) Cables for which any combination of thermal and radiation

aging has led to a breaking strain ratio (e/eo) of less

than 0.4, must not be handled or be permitted to move

(bending crack threshold).

The last limitation is quite severe. If the Arrhenius model

for embrittlement holds, the critical value as defined would

be reached in 40 years at a containment temperature of 80C

with no radiation present. There is evidence, however, that

this projection is very optimistic.

A number of installation provisions, which will mitigate or

eliminate certain failureimodes. recommend themselves:

691
(I) Insulating (instead of metallic) caps on conduit corners

will eliminate creep shortout to ground; so will mechanical load relief (tying) for downhanging cables.

(2) Conduits could be completely filled with cables, or the

cables could be tied, to prevent warpage.

(3) Sharp bends, prohibited by many installation specifications, must be avoided.

(4) Materials combinations in support structures, with vastly

different coefficients of thermal expansion which could

tension cables during thermal cycles, must be avoided.

(5) Sealed conduits and connection boxes are useful in preventing conductivity carrying or generating fluids from

reaching cracks.

(6) Reliable circuit breakers will prevent overload damage.

70
REFERENCES

1. H. F. Reisinger, "Component Failures in Pressurized Water

Reactors," Combustion Engineering Inc., Sandia Corporation Contractor Report 13-6442.

2. W. F. Lindsay, Science Applications Inc., personal communication.

3. M. B. Murphy, Sandia Laboratories, personal communication.

4. B. E. Bader. et al., "Proceedings of Workshop on Nuclear

Power Plant Aging," NUREG/CP-0036, November 1983.

5. R. GrUb and B. Langeset. "Voltage Capability of EPR Power

Cables .... " CERN/EF/Beam Tech. 82-1, January 1982.

6. K. T. Gillen and R. L. Clough, "Dose Rate Effects for

Material Aging Studies," SAND80-1146C, presented at Third

International Conference on Radiation Processes, Tokyo,

October 1980.

7. R. L. Clough and K. T. Gillen, "Radiation and Thermal

Degradation of PE and PVC," SAND8O-2149C, presented at

Third International Conference on Radiation Processes.

Tokyo. October 1980.

8. IEEE Standard 382-1980 (see e.g., Figure 7).

9. E. W. Bennet. ITT Suprenant Division. personal communication.

10. F. V. Thome. "Testing to Evaluate Synergistic Effects

from LOCA Environments," SAND78-1718, April 1978.

11. C. Hosticka, et al.. "Qualification of Cables to IEEE

Standards." IEEE Transactions on Nuclear Science, NS-27,

Vol. 1. February 1980.

12. IEEE Standard 383-1974, "Standard for Type Test of Class

IE Electrical Cables, Field Splices, and Connections for

Nuclear Power Generating Stations."

13. M. Asaka, et al.. "Radiation Resistance of Plastic Materials for Cables. Part 2," Fujikura Densen Giho, Vol. 48,

1973.

14. L. Corbelli and F. Tonioli, "Ethylene-Propylene Elastomers as Insulators in Electrical Technology."

Kunststoffe. Vol. 30, Nr. 2/1977. (In German).

71
15. H. St. Onge, et al., "Research to Determine the Acceptable Emergency Operating Temperature of Extruded Dielectric Cable." EPRI Report EL 983. November 1978.

16. I. Kuriyama. et al.. "Effect of Dose Rate on Degradation

Behavior of Insulating Polymer Materials," IEEE Trans.

on Electrical Insulation. EI-14. October 1878.

17. T. Seguchi, et al., "Dose Rate Effects on Chemical and

Mechanical Properties." Takasaki Radiation Chemistry

Research Establishment. JAERI, undated.

18. Y. Murata. et al.. "Electrical Properties Under Radiation

and Steam Environment." Sumitomo Electric Industries.

EIM-79-91, NE-79-13*, 1979.

19. R. W. Sillars, "Electrical Insulating Materials,"

Peregrinus. London, p. 65 ff. (1973).

20. G. C. Derringer. et al., "Basic Study of the Aging Process." (Review Report), Batelle Institute, Columbus,

November 1979.

21. S. Kronenberg, et al., "Gamma Ray Induced Charge Buildup

in Insulators," Army Electronics Command, Fort Monmouth.

AD 786655. September 1974.

22. 0. M. Stuetzer, "Status Report on Reactor Cable Breakdown

Correlation Study," Sandia Laboratories, unpublished,

January 1983.

23. 0. M. Stuetzer, "Test Sample Alignment for Creep Measurements," Sandia Laboratories. March 1983.

24. J. D. Ferry, "Viscoelastic Properties of Polymers."

Wiley. New York (1970).

25. L. C. E. Struik. "Physical Aging in Amorphous Polymers .... " Elsevier, Amsterdam (1978).

26. K. T. Gillen. personal communication.

27. C. J. Aloisio and G. S. Brockway, "Thermomechanical

Reliability of Plastics .... " International Conference

on Plastics in Telecommunications 2. London, September

1978.

28. J. E. Reaugh. "Analysis of Insulator Cracking and Creep

Shortout," Report of Activities under Sandia Contract,

Science Applications. Inc.. November 1982.

29. R. P. Wool and K. M. O'Connor. "A Theory of Crack Healing

in Polymers," J. AppI. Physics, 52, 5923 (1981).

72
30. H. M. Gilroy, "Thermal Oxidative Cracking of Polyethylene

Insulation," Bell Telephone Laboratories Report, 1974.

31. B. E. Klipec and S. 14. Stephan. Samuel Moore Co., Ohio.

1976.

32. 0. M. Stuetzer. "Electrical Insulators in a Reactor Accident Environment." NUREG/CR-1682° January 1981.

73

APPENDIX A

The complexity of failure assessment for cables in reactors

is illustrated by the following parameter count.

A conservative (i.e., low count) estimate of independent parameters p is made. Tests or verification efforts depend on p

and on the number n of samples taken or variations allowed for

each parameter that Influences (noticeably) cable failure.

e.g.. the breakdown voltage VbThe (simplified) test scenario has the cables first aged in a

containment at temperature To. at radiation rate Ro, and

humidity ho for a time to. This time to may be 1 to 40 years

in reality, and many months under accelerated aging. During to

we assume, for simplicity, no cable breakdown. After aging.

accident conditions are assumed with enhanced parameters T.,

Rj. h. for a continuously increasing time t: where t will be

the variable for measurements of Vb(t).

A systematic list of parameters is presented in Table A-I.

The list is self-explanatory. Some not immediately obvious

points are:

A: Materials: "Additives" means both chemicals added before

processing and flame retardants applied afterwards. The

wire material is not only of mechanical importance, e.g.,

during bending, but also produces chemical effects (slow

copper diffusion into the insulator enhances cracking.)

B: Geometry: The number of conductor strands has a firstorder influence on creep effects. "Flaws" means both

lacerations due to handling and geometric inconsistencies

such as metal particulates or bubbles.

C: Circuit: The Load Impedance determines the current in

the cable. "Frequency" alludes mainly to the fact that

breakdown phenomena for alternating current are often

quite different from those for direct current.

D: Aging: For experimentation, to is the preselected time

for accelerated aging. Under real conditions to is the

time at which the accident starts.

E: Accident: o is the conductivity of the immediate cable

environment, e.g., produced by condensation or reactor

spray.

The list contains a number of simplifications (e.g., "flaws"

covers a number of unrelated deficiencies causing quite different effects) and omissions (e.g.. mutual heating of neighboring cables depends on some additional parameters such as

A-i
Table A-i Cable test parameters (27)

Group Parameter Example

A. Materials:

Insulator Composition n = 3

Binders 2

Additives 2

Processing 1

Jacket Composition 2

Additives 1

Processing I

Wire Material 2

B. Geometry and Forces Wire Diameter 1

Insul. Thickness 1

Jacket Thickness I

# of Strands 2

Bending Radius 4

Applied Stress 2

Flaws I

C. Circuit Voltage 1

Load Impedance 2

Frequency 2

Off/On Percentage 2

D. Aging Exposure Temperature To 3

Rad. Rate Ro 3

Humidity ho 2

Exp. Time to 1

E. Accident Parameters Temperature T. 2

Rad. Rate R. 2

Humidity h. I

Conduct. c 2

Measurement Variable Time t into accident

A-2
heat conductivity and separation). The compilation will serve

its present purpose, however.

The number of parameters listed is p = 27. If each parameter

had only one sample value n. we have. of course, defined only

a single experiment. If each parameter is varied once, i.e..

n = 2. the list leads to nP = 227 = 134 x 106 different

test configurations.

In the last column of Table A-i an investigation is characterized that may be called "desirable" (3 materials for insulator. 2 for jacket. and many parameters neglected as indicated

by n = 1). Even this list leads to almost 900.000 variations.

A severe aggravation is introduced by the fact that many of

the above parameters are not well known: materials vary from

batch to batch, some dimensions have wide tolerances. electrolytic conductivities vary with time. etc. To permit a statistical average, each test configuration must be represented

several times - a factor of 3 or more to be applied to the

above numbers.

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BIBLIOGRAPHIC DATA SHEET SAND83-26222

3 TITLIE AND SVST...E .4 RECIPI(NT 6 ACCESSION NIULMSER

Status Report: Correlation of Electrical Cable •5 DATE REP01COMPL,,C.

Failure with Mechanical Degradation 1 VEAR

March. 1984 6 AUT"MOtSI 7 DATE REPORT ISSUED

Otmar Stuetzer April 1984

9 PROJECT TASK"ORN11 UNI 4UMSILER

a PERFORMING OROGANIZATION NAME AND MAILING ADDRESS Iiwar, ZAP coal,

Sandia National Laboratories to FNUMBER

Albuquerque, New Mexico 87185 NRC FIN No. A-1051

11 SIPD%DR,•a% ORGANIZATION %A•E AN• MA'LnNQ ADDRESS lInIcIe Zp Cowoe 12. TYPE OF REPORT

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555 12t PERIOD COVERED,,m,."pl

•3 S.JP.LEMEETAMo NOTES

14 A&STRACT 12X -01 0.N m.

An attempt is being made to assess complete electrical failure of signal and low-power cables typically used in nuclear power plant containments

and to correlate failure modes with the mechanical deterioration of the

elastomeric cable material. Work over the past 24 months, although

limited to one cable configuration,-has identified creep shortout and

insulator cracking, both aggravated by mechanical stresses, as the

phenomena most likely to cause electrical breakdown. Comprehensive tests

have been run for six months and are continuing. Preliminary conclusions

can be drawn and are reported.

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