NASA CR - 175071

CONTRACT NO. NAS3 - 23894

HIGH FREQUENCY POWER DISTRIBUTION SYSTEM

FINAL REPORT

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by

Induction General, Inc.

GM Industrial Court

3253 Old Frankstown Road

Pittsburgh, PA 15239

Mikund R. Patel

Project Engineer

Prepared for

NASA Lewis Research Center

Space Station Systems Section

Cleveland, OH 44135

David D. Renz

NASA Project Manager

April, 1986

for general release_
NASA CR - 175071

CONTRACT NO. NAS3 - 23894

HIGH FREQUENCY POWER DISTRIBUTION SYSTEM

FINAL REPORT

by

Induction General, Inc.

GM Industrial Court

3253 Old Frankstown Road

Pittsburgh, PA 15239

Prepared for

NASA Lewis Research Center

Space Station Systems Section

Cleveland, OH 44135

Approved

Mikund R. Patel, President

April 1986
ACKNOWLEDGMENT

Induction General, Inc. gratefully acknowledges

contributions from several sources towards the timely

and successful completion of the work documented in this

report: Messrs. D. D. Renz and I. G. Hansen of NASA

Lewis Research Center, who provided continuous guidance

throughout the project; Ms. Sarla Patel, who significantly contributed in the early configuration analysis;

Messrs. R. F. Meserve and A. R. Copp of New England

Electric Wire, who contributed in sharing their valuable

manufacturing experience during the design stage and

also in manufacturing the cable; and Ms. Rosalie T.

Crenelle, who diligently prepared the manuscript.
NASA CR - 175071

HIGH FREQUENCY POWER DISTRIBUTION SYSTEM

ABSTRACT

The objective of this project was to provide the technology of high frequency, high power transmission lines to the

100 kW power range at 20 kHz frequency. In addition to the

necessary design studies, a 150 m long, 600 V, 60 A transmission

line was built, tested and delivered to NASA LeRC for full vacuum tests.

The configuration analysis on five alternative configurations resulted in the final selection of the three parallel Litz

straps configuration, which gave a virtually concentric design

in the electromagnetic sense. Low inductance, low EMI and flexibility in handling are the key features of this configuration.

The final design was made after a parametric study to minimize

the losses, weight and inductance. The construction of the

cable was completed with no major difficulties. The R,L,C parameters measured on the cable agreed well with the calculated

values. The corona tests on insulation samples showed a safety

factor of 3. The EMI test results exceeded the MIL-STD-462

REOl limit due to small non-concentricity introduced by the manufacturing variations. Certain manufacturing improvements are

recommended to reduce EMI.

The project was funded by NASA Lewis Research Center under

Contract No. NAS3-23894, and was monitored by Mr. D. D. Renz.

KEY WORDS: Electrical power, Transmission lines, Distribution, Space station, High voltage, High frequency, Resistance,

Inductance, Capacitance, Analysis, Prototype, Corona, EMI.
TABLE OF CONTENTS

PAGE

EXECUTIVE SUMMARY V

1.0 INTRODUCTION TO PROJECT OBJECTIVE 1.01

2.0 CONFIGURATION ANALYSIS . 2.01

2.1 Initial Alternative Configurations 2.01

2.2 Parameters of Hollow Configurations 2.05

2.3 Parameters of Two Parallel Litz Straps 2.05

2.3.1 Factors Influencing 2-Strap 2.05

Line Parameters

2.3.2 Effect of Stranding and Spiraling 2.05

2.3.3 Skin Effect 2.06

2.3.4 Proximity Effect , 2.07

2.3.5 Resistance Calculation 2.11

2.3.6 Inductance Calculation 2.12

2.3.7 Capacitance Calculation 2.12

2.3.8 Mechanical Force Calculation 2.15

2.3.9 Mass Calculation 2.17

3.0 PARAMETRIC STUDY ON INITIAL CONFIGURATIONS 3.01

3.1 Hollow Configurations 3.01

3.2 Two Parallel Litz Straps 3.01

3.3 Parameter Optimization in 2-Strap Cable 3.04

3.4 Discussion on Initial Configuration 3.04

Designs

4.0 THREE PARALLEL LITZ STRAPS CONFIGURATION 4.01

4.1 3-Strap Configuration Analysis 4.01

4.2 Shielding and End Insulation 4.03

Considerations

4.3 Comparison Between 2-Strap and 3-Strap 4.05

Configurations

4.4 Design Optimization in 3-Strap 4.07

Configuration

ii
TABLE OF CONTENTS

PAGE

4.5 Cable Ratings Revised 4.07

4.6 Line Parameters of the Selected 4.09

3-Strap Design

5.0 MATERIAL SELECTION 5.01

5.1 Conductor Materials 5.01

5.2 Inherent Material Limitations on 5.04

Loss x Mass Product

5.3 Dielectric Materials 5.05

5.3.1 Strand Insulation 5.05

5.3.2 Major Insulation 5.05

5.4 Low Loss Insulating Material 5.06

5.5 Dielectric Loss Calculation 5.13

6.0 CORONA CONSIDERATIONS 6.01

6.1 Corona Inception Voltage Vs. Maximum 6.01

Permissible Void Size

6.2 Dielectric Stress Concentration in 6.05

End-Insulation

6.3 Corona Resistance 6.07

7.0 OTHER CONSIDERATIONS 7.01

7.1 Environmental Considerations 7.01

7.2 Manufacturing and Flexibility 7.01

Considerations

7.3 Joints and Termination Consideration 7.02

8.0 ALL THE ALTERNATIVES COMPARED FOR FINAL 8.01

SELECTION

9.0 THERMAL DESIGN CALCULATIONS 9.01

9.1 Radiation Cooling Alone 9.01

9.2 Conduction Cooling to Continuous 9.03

Base Plate

9.3 Conduction Cooling by Intermittent 9.05

Mounting Supports

9.4 Summary of Thermal Design Calculations 9.09

10.0 MECHANICAL FORCE CALCULATIONS 10.01

ill
TABLE OF CONTENTS

PAGE

11.0 ELECTROMAGNETIC INTERFERENCE 11.01

11.1 EMI Flux Density Estimate 11.01

11.2 EMI Shielding 11.03

12.0 FINAL DESIGN SUMMARIZED 12.01

12.1 Discussion on the Final Design 12.01

12.2 Summary of Basic Formulas for 12.04

3-Strap Design

13.0 FAILURE MODES AND EFFECTS ANALYSIS 13.01

13.1 Exposure to Free Space Temperature 13.01

13.2 Exposure to Ultraviolet Radiation 13.01

13.3 Voids in Insulation 13.02

14.0 SAMPLE CONSTRUCTION AND TESTS 14.01

14.1 One Meter Long Sample 14.01

14.2 3.7 Meter Long Sample 14.01

14.3 Five Meter Long Sample 14.05

15.0 CONSTRUCTION OF FULL CABLE 15.01

16.0 TESTS ON FULL CABLE 16.01

16.1 AC Corona Tests 16.01

16.2 R,L,C Tests 16.04

16.3 EMI Tests with Cable Terminals 16.13

Inside the Chamber

16.4 EMI Tests with Cable Terminals 16.14

Outside the Chamber

APPENDIX A: EDDY LOSSES IN LITZ STRAP LINES A.01

DUE TO TRANSVERSE FLUX

APPENDIX B: PARAMETRIC STUDY ON 2-STRAP B.01

CABLE CONFIGURATION

LIST OF FIGURES L.01

LIST OF TABLES L.05

LIST OF REFERENCES L.07

DISTRIBUTION LIST x L.09

iv
EXECUTIVE SUMMARY

The objective of this project was to provide the technology of high voltage, high power transmission lines to

deliver power at 1000 Vrms and 100 Arms levels, operating at

20 kHz in vacuum. In addition to the necessary design studies,

a 150-meter long, 600 V, 60 A transmission line was built,

tested and delivered to NASA Lewis Research Center for vacuum

testing. The analysis, design, manufacturing and test results

obtained during the course of the project are documented in

this report.

In the initial part of the report, four alternative configurations have been analyzed, namely, the hollow coaxial,

hollow stranded coaxial, hollow parallel and two parallel Litz

straps. A parametric study and the design optimizations on

these configurations, along with the manufacturing considerations, led us to recommend the two parallel Litz straps cable

configuration. Three attractive features of this configuration

are low inductance, good manufacturability and flexibility in

handling. Further improvements were achieved by splitting one

of the two straps in two parts and locating the split halves

on two sides of the full strap. As shown in the report, the

resulting three-strap configuration is a virtually concentric

cable in an electromagnetic sense. For a given cable width,

it reduces the inductance to less than one-half. Furthermore,

since the magnetic flux is now internally contained (see sketch

below), it offers lower EMI. At the same time, it maintains

End-flux

Internal flux containment

in the 3-strap configuration
other attractive features of the 2-strap configuration, including the flexibility in manufacturing and handling. The

3-strap configuration, therefore, was adopted for the final

design.

Alternative conductor materials were evaluated. It is

shown that for a given conductor material, the loss x mass

product of the cable is invariant. Copper meets the set goal

on loss x Mass product at the anticipated operating temperature.

Aluminum would exceed the goal but would pose severe limitations on manufacturing, joints and terminations. Copper was

chosen as the conductor material for its solderability. Alternative insulating materials were also investigated. Polypropylene was selected as the major insulating material for low

dielectric loss.

From corona considerations, the maximum permissible void

size in the insulation has been established. This has a bearing on selecting the manufacturing process and operating practice for the cable. For example, the failure mode and effect

analysis on the cable showed a possibility of creating a delamination void between the middle and the outer straps that may

exceed the permissible void size of 5 mils. Bleeding nitrogen

along the length of the cable may be necessary to avoid corona

under such delamination. Embrittlement and cracking of the

insulation at low temperatures is another possible source of

voids in the insulation.

Two inch wide cable was initially selected as an optimum

design for the originally specified 1000 V, 100 A cable.

However, the manufacturing investigations during the final

design phase revealed that the existing cable extrusion process

would allow only 1.5 inch wide cable. The decreased cable width

required increased cable thickness, resulting in the aspect ratio

of 3.2 versus at least 5 desired for a good flexibility in handling. Upon reviewing this constraint and the near term power

requirement, the cable rating was revised to 600 V, 60 A. The

1.5 inch wide extrusion die limitation, however, is not of a

vi
fundamental nature. Wider extrusion dies and tools can be

developed for higher cable ratings when required. Such cost,

however, was considered unnecessary at this time.

For the 600 V, 60 A cable with 3-strap configuration, the

copper conductor and polypropylene insulation were selected in

the final design. The cable dimensions, design parameters and

design formulas are summarized in Section 12.0. The losses in

the final design are 3.755 watts/m, inductive voltage drop

0.203 v/m, capacitive charging current 0.227 A/m and the mass

1.102 kg/m. If one side of the cable is radiating to free space

at 80°K (liquid nitrogen temperature), the operating cable temperature is expected to be tens of degrees below 0°C. If the

cable is mounted continuously on a 60°C base plate, the operating temperature of the cable would be 61°C. The low temperature rises are due to low losses and large surface area available

for cooling. The mechanical and electrical stresses in the

insulation are adequately low.

Pre-production cable samples were constructed and tested

for R,L,C parameters. The measurements on these samples checked

well with their calculated values. The test results, therefore,

indicated the validity of the design formulas developed for the

cable. The design was then formally reviewed and released for

the construction.

There were several problems encountered during the manufacture of the cable. First of all, maintaining accurate dimensions of Litz straps during and after the manufacturing processes

was difficult. Generally, the strap thicknesses came out 15 to

20% oversize. In the second manufacturing step, applying a 25

to 30 mil polypropylene jacket (major insulation) over the middle strap with no tearing or pinholes was a smooth process and

was completed with no difficulties. In the last step, applying

the outer PVC jacket over the 3-strap assembly, however, proved

difficult because of the fluid pressure which built up at the

corners of the rectangular extrusion die. Because of this, the

vii
outer straps became narrower than the middle strap and the PVC

penetrated excessively inbetween the middle and outer straps.

To eliminate this problem, one-mil teflon tape was half-lapped

around the 3-strap assembly before extruding the outer PVC

jacket. This considerably improved the cable construction.

A detailed Test Plan for testing the cable was prepared

for R,L,C,corona and EMI tests. Several flat samples of polypropylene were tested for AC corona under ASTM standard brass

electrodes. Twenty mils thick samples gave 1550 Vrms corona

inception voltage and 1400 Vrms corona extinction voltage. For

a 440 V operating voltage, this gives a safety factor of 3 for

the solid insulation. The previously mentioned delamination

voids may cause corona under vacuum at a lower voltage.

From a theoretical consideration, the 3-strap cable configuration was expected to give negligible EMI. However, due

to manufacturing variations, the inductances of the two individual outer straps with respect to the middle strap were unequal. This caused unequal currents in the outer straps. The

resulting solenoidal flux radiating from the cable was sufficient to exceed the MIL-STD-462, Section RE01 limit of 27 db

over one picotesla at 20 kHz. Further manufacturing improvements and/or shielding will be necessary in the future to meet

the EMI requirement.

The R,L,C tests were made on full length cable under

20 kHz excitation at full current and full voltage. The resistance agreed well, while the inductance was slightly higher and

the capacitance slightly lower than the calculated values. The

deviations in the L and C are attributed to the oversized dimensions that resulted in the final manufacture. Accounting for

the oversize, the R,L,C test results clearly proved the validity

of the virtually concentric 3-strap cable concept and also verified the formulas developed for calculating the circuit parameters.

viii
1.0 INTRODUCTION TO PROJECT OBJECTIVE

NASA studies have projected the power requirements of

large spacecrafts, power platforms and space stations in the

megawatt level by the year 2000. Analysis on near term power

needs has identified a variety of missions and space station

configurations requiring significant increase in electrical

power up to the hundred kilowatt level. As the need for space

power grows larger, so do the requirements for the power distribution system grow in size, sophistication, and complexity.

Studies by NASA have shown that significant advantages in

weight, cost, efficiency, and design flexibility can be achieved

if power transmission and distribution could be accomplished at

high voltage and high frequency.

A critical element in high voltage, high power distribution system is a high frequency, high power transmission line

which has low loss, low weight and low reactive voltage drop.

This project is to provide this technology for high frequency,

high power transmission lines to the 100 kW power range at

20 kHz, 1000 V and 100 A in a vacuum.

The original specifications and goals for a prototype of

such a line are summarized in Table 1.1. During the course of

the project, the ratings of the first prototype line were

changed to 600 V, 60 A capacity. In addition to the necessary

design studies to meet the set specifications, the goals

included a 150 meter long transmission line to be built, tested

and delivered to NASA Lewis Research Center for vacuum testing.

1.01
TABLE 1.1: SPECIFICATIONS AND GOALS AS SET BY NASA

INITIAL GOALS

A. RATINGS

Voltage - 1000 volts (rms)

Current - 100 amps (rms)

Frequency - 20 kHz

B. DESIGN GOALS

I

2R Loss - .01% per meter (measured at full load current)

Weight - .01 kg per kW per meter

Reactive drop - 1 volt per meter

C. TEMPERATURE

The system shall be designed to reject its heat to a

grounded 60°C base plate or mounting points as required.

D. DESIGN LIFE

The design life shall be ten (10) years of continuous

operation in a vacuum at rated conditions.

E. EMI

The system shall be designed to meet MIL-STD-462 sections

1) RE01 - 30 Hz to 30 kHz

2) RE02 - 14 kHz to 12.4 GHz

3) RE04 - 20 Hz to 50 kHz

F. LENGTH

System shall be 50 meters in length.

REVISED RATINGS

The ratings of the first prototype were later revised

to 600 V, 60 A, with proportionate changes in the loss,

weight and inductance. At the same time, the length

was changed to 150 meters.

1.02
2.0 CONFIGURATION ANALYSIS

Five alternative configurations were investigated in

two phases. The initial phase consisted of four different

configurations, including two parallel Litz straps. A

thorough parametric study was conducted on this group for

minimizing the loss, weight and inductance of the cable. At

this stage, a new three parallel straps configuration was

conceived and was analyzed in the second phase.

2.1 INITIAL ALTERNATIVE CONFIGURATIONS

As for the alternative configurations of the targeted

transmission line, the following four options were initially

studied:

1) Hollow coaxial cable (with solid sheet conductor)

2) Hollow stranded coaxial cable

3) Hollow round parallel lines

4) Two parallel Litz straps

The first three configurations have been thoroughly investigated by Jefferies and Renz in a NASA study

 l

 •" , and are

depicted in Figure 2.1. These configurations with central part

filled with conductor (instead of hollow) offer other three

options, which are theoretically included in the above study

with the inner radius equal to zero. To avoid duplication,

the NASA study is not reproduced here. The interested reader

must read Reference 1 as a complement to this report. However,

the results of that study have been used in the present investigations.

Coaxial cable with hollow center conductor was the configuration selected in Reference 1 for high frequency distribution of space power. Nearly equal radii of the inner and outer

conductor minimized inductance. To provide cable flexibility,

stranded rather than solid conductors were considered. A 5 mm

inside radius stranded coax cable with 0.5 mm conductor thickness was cited as an example to transmit 100 kW at a voltage

of 1000 V AC for a distance of 50 m. This cable would have

2.Q1
Hollow center

Copper conductors -*-''^

Teflon dielectric-''

Hollow coaxial cable (solid sheet conductor)

^^ Copper strands

^ Teflon dielectric

f •*''

' ••_,— Hollow center

Hollow stranded coaxial cable

Hollow center -~^

Copper conductor •—..„

Teflon dielectric

Hollow parallel lines

(a) Configurations investigated.

<u

o

« 6CS

4J

U

3

T3

O

DC formula

\ -AC formula

General formula

012 ?

Conductor thickness, skin depths

(b) Conductance versus conductor thickness.

FIGURE 2.1= Hollow Configurations Investigated

by Previous Workers (Reference 1).

2.02
a power loss of 1900 W, an inductance of 1.45 yH and a capacitance of 0.07 yF. Reference 1 could be used to determine

characteristics of other cable sizes.

There are some disadvantages attached with the hollow

coaxial configurations:

(a) With solid conductor, the cable is rigid and extremely

difficult to handle in practical applications.

(b) With stranded conductor, the inner conductor may buckle

under inward radial electromagnetic force . A support

structure in the center can prevent such collapse, but

would add weight.

(c) The central hollow space adds unnecessary volume.

The third configuration, the hollow round parallel lines,

have all the above disadvantages, plus two more. It has much

higher reactance and occupies almost double the volume compared

to the coaxial configurations.

The fourth configuration, namely the two parallel Litz

straps (Figure 2.2) appears to be a more attractive configuration, and has been thoroughly investigated in this section.

This configuration, having longer leakage flux path, produces

less magnetic flux per ampere, and hence offers a lower reactance design compared to round parallel lines. The strap

configuration has two parameters, a and d, which can be conveniently changed to meet the specified reactance. This configuration also appears to be easy to manufacture and install

on site.

From thermal viewpoint, the flat straps provide more

heat dissipation area. Dielectrically, by rounding the strip

edges, we can achieve more or less a uniform field pattern

between the two straps. Thus, from all major design considerations, flat strap configuration appears to be more favorable

for meeting the objectives of the project. It is therefore

evaluated as a serious condidate to compete with the first

three options.

2.03
h

FIGURE 2.2: Two Parallel Litz Strap Lines.

It offers longer flux-leakage path and hence

lower inductance.

2.04
2.2 PARAMETERS OF HOLLOW CONFIGURATIONS

The design parameters of the first three alternative

configurations (referred to as hollow lines from now on in

this report) are fully documented in Reference 1, which we •

will refer to for further details.

2.3 PARAMETERS OF TWO PARALLEL LITZ STRAPS

Three basic circuit parameters required for the proposed

design are series resistance R to calculate the I2R losses,

series inductance L to calculate the reactive voltage drop

and line-to-line capacitance C to calculate the charging

current. Other constants, such as capacitance and conductance

to ground may also be required for detailed calculations, but

are not considered in the present study.

2.3.1 FACTORS INFLUENCING 2-STRAP LINE PARAMETERS

The major influence factor on the line parameters is,

of course, the geometrical spacings, which will be considered

in detail in latter sections. Before that, the effects of

the following factors are considered below:

a. Effect of stranding and spiraling

b. Skin effect

c. Proximity effect

2.3.2 EFFECT OF STRANDING AND SPIRALING

Because numerous thin and completely transposed strands

are to be used in the proposed line, the current can be

assumed to be uniformly distributed for inductance and resistance calculations, except for small eddy loss corrections as

discussed in the next section. The capacitance of a stranded

line is very nearly equal to that of a solid conductor having

the same periphery, and needs no correction.

As for the effect of spiraling, the resistance would be

higher due to longer strand length required to cover the given

line length. This depends on the transposition pitch and is

taken into account in calculating the resistance. On the

2.05
other hand, inductance depends only on the geometrical arrangement of the conductors. Since the current, on average, is

parallel to the axis of the line, the line length, and not

the strand length, is used in the inductance calculation.

The same argument applies for the capacitance calculations.

The increase in length and hence resistance due to

cabling (stranding and spiraling) is difficult to calculate

from an analytical expression, as it depends on many manufacturing process variables. The easiest way of calculating

this resistance increment factor Kc due to cabling is:

_ Manufacturers catalog weight for given cable length

c ~ Theoretically calculated weight based on straight length

2.3.3 SKIN EFFECT

Alternating currents at high frequencies concentrate

near the outer surface (skin) of the conductor. The skin

depth 6 is given by

where p = resistivity of the conductive material,

]i = permeability of the material, H/m

f = frequency, Hz

•

For copper, 5 is 9 mm at 60 Hz and 0.5 mm at 20 kHz. In order

to keep the eddy losses low, it is necessary to keep the wire

diameter smaller than the skin thickness. Experience in high

frequency design indicates that copper wires for 20 kHz application should be smaller than 0.5 mm (20 mils) in diameter.

Clearly, the high current carrying capacity will require

numerous strands of fine wires in parallel. To keep the eddy

losses low in such bundles of wires, individual strands must

be insulated and continuously transposed to occupy all the

possible locations in the bundle along the full length of

the transmission line. Cables of round shapes, such as Litz

cables used in radio frequency transmission lines, can be

2.06
used. Alternatively, flat braided straps can also be used,

as proposed for the present lines. Calculating the high

frequency AC resistance is complex and involves ber and bei

functions. Practical curves to find the ratio of effective

AC resistance R to DC resistance RQ have been derived by

Grover and published by the U.S. Bureau of Standards

for values of mr up to 100, where r = wire radius and m =

/2/5. As seen in Figure 2.3, the R/RO ratio becomes excessive

beyond mr value of 2. This indicates that the strand's r/<5

ratio should be smaller than /2~ at a given frequency to keep

the skin effect ratio Ks = R/RO less than 1.08.

The skin effect decreases the internal inductance of

the conductor. However, in the stranded flat strap configuration, this decrease is negligible.

2.3.4 PROXIMITY EFFECT

Consider the case of two parallel cylindrical conductors

in close proximity as shown in Figure 2.4. Here, the current

density varies both radially and azimuthally. We may say that

the current strive to occupy the two facing zones of the two.

conductors due to the proximity of each other.

The effect is still more striking in two parallel tubular

conductors. If the tube wall is small, there is no noticeable

radial variation in the current density, and there is, thus, no

simple skin effect. The proximity effect, however, introduces

an asymmetry in the peripheral distribution large enough to

make R and L appreciably different from those calculated by

assuming axisymmetrically uniform current distribution. It

appears that this effect is ignored in calculating R and L of

the hollow round parallel lines in Reference 1. If so, the

actual R and L will be considerably higher than those calculated in that study. For the stranded flat strap configuration, because the strand size would be much smaller than the

skin depth, the.proximity effect on the line parameters will

be small, as discussed below.

2.07
R/RO

1.08

1.03

1.02

1.01

i.oci-^ 0 0.2 0.4 0.6 0.6 !.0 12 1.4 1.6 1.8 2.0

FIGURE 2.3: Ratio of AC to DC Resistance for a Round Wire

in self magnetic field, where r is the wire radius and m = /2/5

2.08
Current crowded at

the darker zone

FIGURE 2.4: Characteristic of Proximity Effect on the peripheral

distribution of current density. Notice that the current is

crowded at the facing zones of the two conductors.

2.09
The equivalent of the above proximity effect in stranded

cables is the eddy losses due to transverse flux perpendicular

to the strand axis. If a single strand of diameter do is

placed in transverse flux density B, the resulting eddy loss,

as derived in Appendix A, is:

f2 B2 j i*

_ ^o watts/meter (2.2)

The leakage (transverse) flux density in the flat strap

configuration varies linearly from maximum at the inner faces

to zero at the outer faces. The maximum flux density BQ is

then:

BQ = y0 H0 = y0 ^ in SI units

If there are N total number of strands, integrating

individual strand losses and converting them into the ratio

of AC to DC resistance, we obtained the following transverse

eddy loss factor in Appendix A:

0.8 f2N

2

d0

6

Kf. = 1 + - in SI units (2.3)

P

2b2El2

The conclusion of this subsection is that the basic R, L

and C parameters of the stranded flat strap configuration can

be calculated assuming uniform current distribution without

significant loss of accuracy, except for a small skin effect

and transverse flux eddy loss corrections in R. Accordingly,

the R, L and C calculations of the stranded flat strap configuration follows.

. 2.10
2.3.5 RESISTANCE CALCULATION

If Kf = overall fill factor of conductor

= conductor cross section

total cross section

a,b,d = geometrical dimensions (Figure 2.2)

p = resistivity of the conductor

Subscript 1 refers to the parameter per meter of

the lines.

Then, the DC resistance per meter loop of the cable

(i.e. one meter lead and one meter return) is given by:

R

dcl =

 = ohros/m loop (2.4)

For copper at 20°C/ p = 1.77 E-8 ohm m. At higher temperature

T°C, it can be calculated by the following IEEE standard

formula:

t

PT _ 234.5 + T . ,~ ,,

P2Q ~ 234.5 + 20

 ( '3)

Now, recalling that Kc = cabling (stranding and spiraling) factor

Ks = skin effect eddy loss factor

Kt = proximity-effect (transverse flux) eddy

loss factor

With all these factors multiplied, the total resistance of a

cable is:

R

acl

 = KsK^c^dcl ohms/m loop (2.6)

DC resistance increases with temperature, while KS and

Kt decreases with increasing temperature. For this reason,

2.11
it would be inaccurate to correct Rac according to Equation

(2.5). For large temperature deviations, Rdc, Ks and Kt

should be separately corrected.

2.3.6 INDUCTANCE CALCULATION

The exact inductance calculation of this configuration

is mathematically involved. The exact formula has been

derived in the Bureau of Standards Bulletin, Vol. 3, No. 1,

Pg. 6, as reproduced in graphical form in Figure 2.5. These

curves are plotted at 60 Hz reactance per foot length one-way.

However, since they neglect skin effect and assume uniform

current distribution, reactance of high frequency stranded

cables can be taken in the proportion of the frequencies as

long as the current distribution is uniform, as in the present

case. If the conductors were solid bus bars, instead of finely

stranded straps, these curves would be invalid at higher

frequencies.

Since width b of the straps is likely to be much larger

than the strap thickness a and the insulation thickness d,

the inductance of this configuration can be simply calculated

by thick infinitely long solenoid (or transformer) type

formula without significant loss of accuracy. Neglecting the

fringing flux at the ends and the magnetomotive force consumed

in the external return flux path, the inductance per meter

length can be shown to be:

l

 = ^F (d + ^f ) H/m loop (2*

7)

This formula agrees well with Figure 2.5 at the lower end.

2.3.7 CAPACITANCE CALCULATIONS

Ignoring the fringing of the electrostatic field (Figure

2.6), the capacitance per meter loop of the line is given by

the following approximate but simple flat electrodes formula:

2.12
(For Upper Curves)

3456 7

13

FIGURE 2.5: Reactance of Rectangular Flat Conductors at 60 Hz

per foot one-way). Source: Bureau of Standards.

2.13
-K SL

* /

T ~v ! *"": P^""~.

FIGURE 2.6: The Strap Lines as a Capacitor

2.14
= ^ = F/meter loop (2.8)

where A = electrode area = b per meter length

e = permittivity of dielectric between electrodes

d = separation between electrodes

The fringing (stray) field introduces significant complexity in the capacitance calculations. Maxwell took this

into account by using the method of conformal transformation

and derived the following accurate formula

[4].

= e I* •^ + 1 + Ln -^ (2.9)

The first term is the familiar expression for the ideal

line without fringing, while the second term gives the edge

correction. The formula is accurate within 0.3% for b=d and

much better for larger b. It can, therefore, be considered

exact for the present application.

The total capacitance for the 50-meter long line is

C = 50 C^ and the total charging current at the receiving end

at voltage V is:

Cc = VCu) amps (2.10)

2.3.8 MECHANICAL FORCE CALCULATION

The force on two round thin parallel straight wires

carrying currents 1^ and ~L-^, 0 degrees out of phase, distance

s meters apart has been shown to be:

21,1. COS0

N/meter loop (2.11)

The force between two straps, each of cross section

a x b, placed opposite each other at distance s, is given by:

PI = KF

where K is the correction factor as plotted in Figure 2.7

(2.12)

[5]

2.15
ORIGINAL! PAGE IS

OE EOOR QUALITY,

0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 20

s-»m *IR SPACE BETWEEN STRAPS

• + b SUM OF SIDES

Repulsion between strap conductors.

F — repulsion between two equal long parallel straps. -

?' — repulsion between two long parallel straight wiraa.

_ 5.40

 /l/l f08

 ? lo-r p0unda per foot of'circuit.

where t is the phase angle between /i and /t which are in amperes, and where «

it the axial spacing in inches. F and f are average force*.

FIGURE 2.7: Force between Rectangular Flat Conductors

for different geometrical spacings

2.16
The above force, in steady state conditions, will have

twice the basic frequency, i.e., 40 kHz. This, over a ten

year period, will amount to 12.6 x 1012 cycles of tensile

stresses. The conventional fatigue studies are done over a

few million cycles. Since the present requirement is much

beyond this conventional range, the tensile stresses in the

major insulation have to be kept exceptionally low. Thermal

aging adds into the complexity in evaluating the electrical

degradation of the insulation. McNutt and Patel have

studied the combined effects of mechanical stresses and thermal

degradation on one class of insulating materials. Such considerations, based on experience and judgement, can be applied

for designing the system to endure the mechanical fatigue over

the functional life of about ten years at 20 kHz.

2.3.9 MASS CALCULATION

If Y is the conductor mass density, Kc and Kf are the

cabling and fill factor, and a x b are the cross sectional

dimensions, the mass per meter loop of the line is simply

given by:

M, = 2KcKfaby kg/m loop (2.13)

2.17
3.0 PARAMETRIC STUDY ON INITIAL CONFIGURATIONS

The effects of various design parameters on the initial

four configurations of the transmission lines are discussed

in this section in order to establish their trade-off values.

3.1 HOLLOW CONFIGURATIONS

The parametric studies on these configurations have been

fl 21

extensively made by Jefferies, Renz and others . To

avoid duplication, this NASA study is not reproduced here.

However, its results are used to derive the optimum designs

for each of the three hollow configurations depicted in

Figure 2.1. These designs are summarized in rows 3 to 7 of

Table 3.1.

3.2 TWO PARALLEL LITZ STRAPS

The equations developed in Section 2.3 are simple enough

for a hand calculator. However, for performing numerous calculations for a parametric study, a computer program was prepared. The details of this program are given in Appendix B.

The AC resistance, inductance, capacitance and weight of the

lines with two parallel Litz straps were calculated for both

copper and aluminum conductors. The insulation mass density

of 1340 kg/m3 (that of PVC) has been assumed in the weight

calculations. Computations were made for major insulation

thickness d ranging from 1 to 3 mm, strap thickness a from

1 to 4 mm, and strap width b from 50 to 100 mm. The results

are summarized in Figures B.I to B.5 of Appendix B. From

these curves, three alternative designs with copper are given

in rows 8 to 10 of Table 3.1. Similar designs with aluminum

conductor are given in rows 11 to 13.

3.01
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3.02
NOTES & STARS FOR TABLE 3.1

NOTES: Losses taken as the most severe constraint. That

is, most designs start with meeting the losses.

All hollow designs are with 0.2 mm (8 mil) copper

conductor and 0.5 mm (20 mil) teflon dielectric.

Center support mass not included in the hollow

designs.

Two parallel Litz straps designs include insulation

mass and use conductor fill factor of 0.4.

Hollow designs used p = 1.77E-8 and parallel Litz

strap designs used p = 2.2E-8 as the copper resistivity. The former is at 40°C temperature and the

latter is at 100°C.

STARS: * 1

* 2

* 3

* 4

* 5

* 6

* 7

* 8

* 9

*10

Goals set by NASA for the present contract work.

Designs from Ref. 1 (shield not included).

Design from Ref. 2 (shield included).

Design from Figure B.I of this report.

Design from Figure B.2 of this report.

Rise across one dielectric layer due to radiation.

Conductors 1 mm apart.

After correcting an apparent misprint in Ref. 2.

See Figures 2.1 and 2.2 for dimensions.

From design considerations alone, this is the best

of all configurations considered in this table,

meeting all the goals with 10% extra margin.

Further improvements with 3-strap configuration

are discussed in Section 4.0.

3.03
3.3 PARAMETER OPTIMIZATION IN 2-STRAP CABLE

As will be discussed in Section 5.2, there is no theoretical scope for optimizing the conductor Loss x Mass product, which remains constant for a given material. If both

loss and mass are evaluated with equal importance by NASA,

one can factor the two in any combination. If one is more

important than the other, than the important attribute can

be favorably factored in the final design.

However, an optimization opportunity exists in the insulation weight. For minimum cable weight, the insulation

should be as thin as possible while keeping the cable coronafree. There is, however, another limitation in making the

insulation thin. The extrusion process of applying the insulation on the straps requires certain minimum thickness to

avoid tearing of the insulation during the manufacturing

process. Experience indicates that a minimum of 25 to 30 mils

insulation thickness is required for a tear-free extrusion.

Once the insulation thickness is fixed, the total conductor and insulation mass can be minimized by selecting the

minimum strap width b while choosing the strap thickness a to

meet the losses and the reactance. Since the loss depends on

the product a x b, for a fixed loss, i.e. for a fixed a x b,

the reactance increases with increasing a and decreasing b,

while the mass of insulation of given thickness decreases

with increasing a, as there is now less insulation thickness

in proportion to the strap thickness. For a given loss and

insulation thickness, the cable weight can, therefore, be

minimized by decreasing b and increasing a until the reactance

increases to its permissible limit. This is done in row 13

of Table 3.1 for aluminum straps, leaving about 6 to 10%

margin in all the parameters, i.e. loss, mass and the reactance,

The same optimization logic also applies for copper straps.

3.4 DISCUSSION ON INITIAL CONFIGURATION DESIGNS

Table 3.1 indicates that in the first group of four

alternative configurations, the hollow copper coaxial lines

3.04
meet the goals on losses and reactance, but not the mass, and

occupy large volume. The hollow parallel line meets only the

loss goal, but misses both the reactance and the mass goals,

and occupies much larger volume.

Two parallel Litz straps with copper conductor can meet

the reactance goal and either loss or mass goal, but not both

exactly. This is also true for the hollow coaxial configurations with copper conductor. The aluminum straps, on the

other hand, meet all the goals comfortably, leaving 6 to 10%

margin for final design adjustments. The parallel straps

occupy the least volume.

Thus, among the four configurations evaluated in this

phase, the parallel Litz straps appear to be the best, the

hollow coaxials rank next best and the hollow parallel ranks

the last.

From only the design point of view/ the parallel straps

configuration with aluminum strands is the most optimum configuration. However, as discussed in Section 7.0, aluminum

poses difficulties in manufacturing and joining. The flat

straps with copper strands offer an optimum configuration

from both the design and manufacturing considerations.

3.05
4.0 THREE PARALLEL LITZ STRAP CONFIGURATION

4.1 3-STRAP CONFIGURATION ANALYSIS

The design and manufacturing considerations studied so

far led us to recommend the two parallel Litz straps configuration. The subsequent study showed that the three parallel straps configuration results in even better design, in

which the middle strap carries full rated current at the rated

voltage, and each of the grounded outer straps carry onehalf of the return current. Such a birfurcated design makes

the cable virtually concentric in the electromagnetic sense,

as evolutionary shown in Figure 4.1. The advantages of the

three-strap over the two-strap design are as follows:

1) The magnetic flux density B in the major insulation

region is now one-half. The reactance is a measure

of the magnetic energy B2/2y stored in the configuration space. Since there are now two major insulation gaps, the net effect is to reduce the cable

reactance to less than

2 x (B/2)2/B2 = 1/2.

2) The magnetic flux in the two major insulation regions

is now in a push-pull mode, thus containing the flux

within the cable volume itself. This significantly

reduces the EMI shield requirement. The main flux

in the two-strap design, on the other hand, returns

external to the cable (like an air-cored solenoid).

3) Due to one-half B, the eddy losses by the transverse

flux reduces to one-quarter. This is only a minor

gain, as the eddy losses are already kept small by

using thin strands. However, this may allow thicker

strands if desired for other reasons.

4) The repulsive force between the central strap and the

outer straps carrying half the current now reduces to

one-fourth. Again, this gain is not much useful, as

the forces are already low.

4.01
(a) Take a Hollow

Coaxial Cable

Flux Lines.

(b) Squash it and cut

top and bottom ends

(c) Remainder, which can be

made with 3 parallel straps

FIGURE 4.1: Three Parallel Straps Configuration.

See how it evolves to be virtually concentric, while

keeping the flexibility in manufacturing and handling,

4.02
5) The flexibility in manufacturing, handling, splicing

and terminating are still maintained.

6) Thermally, the three-strap configuration is not much

different than the two-strap configuration. Due to

large flat surfaces available for cooling, the operating temperatures are anticipated to be much below

the material limits.

Among the above gains, the substantial reduction in the

reactance is viewed most favorably for the overall powersystem management in large space stations.

,4.2 SHIELDING AND END-INSULATION CONSIDERATIONS

Since the main magnetic flux in this configuration is

internally contained, the EMI at some distance away from the

cable would be negligible. However, there exists some fringing

flux at the ends, as shown in Figure 4.1(C). The extent of

this flux is difficult to calculate by hand. To obtain a

<••

quantitative picture of this fringe region, the in-house

literature was searched. A set of three-phase bus bars at the

instant when the middle bus carries 1.0 per unit and the outer

buses -0.5 per unit currents closely resembled the virtually

concentric configuration under present study. A finiteelement magnetostatic flux plot of the three-phase buses is

depicted in Figure 4.2. Notice the internal containment of

the main flux. Also notice that most of the fringing flux

at the ends returns within about one-half the distance between

the outer buses.

Since the EMI shield has to now contain only the fringing

flux, as compared to the main flux in the two-strap configuration, the shield requirement in the three-strap design is

significantly lower.

4.03
y-axis

x-axis

fringing

flux at ends

FIGURE 4.2: Finite-element flux plot of 3-phase bus bars with

central bus carrying peak current. Notice the internal containment of the main flux and that most of the fringing flux at

ends returns within about one-half the distance between the

outer buses. The dotted lines are the bus bars.

4.04
The three-strap design can be further improved by indenting the middle strap (i.e. stepping it down) with respect to the outer straps as shown in Figure 4.3. This has

two advantages:

0

V

VJ

1.0

p.u.

volts

I

r •»

0

V

J

Figure 4.3: Indentation of the

middle strap improves EMI and

end-insulation design.

• The end flux is pulled inward, thus further reducing the shield requirement.

• The additional end-insulation for the middle strap

operating at full line voltage would be adequate

to withstand nonuniform dielectric stress even

without an electrostatic shield over the cable.

4.3 COMPARISON BETWEEN 2-STRAP AND 3-STRAP CONFIGURATIONS

Table 4.1 compares a 50 mm (2 inch) wide 2-strap and

3-strap designs for 1000 V, 100 A ratings. Clearly, the threestrap configuration is superior with respect to the inductance,

tensile stress and EMI shield requirement points of view,

4.05
TABLE 4.1

Comparison between 2-strap and 3-strap 1000 V, 100 A cable designs.

The latter has less than one-half the inductance, double the capacitance and the same resistance, mass and volume.

Configuration

Goals for 50 m

cable

Two Litz

Straps with

a=3, b=50,

d=l mm

Virtually

concentric

Three straps

Indented

a-1.5,3.0,1.5

b=50, d=l mm

R mft

50

47

47

L UH

4

3.7

1.5

C yF

— —

.06

.12

Mass

kg

50

71

71

Volume

m

3

— —

.023

.023

Volts

mil

— —

25

25

Tensile

Stress

Relative

— —

100

25

Shield

Requirement

— —

Yes

Much

reduced

4.06
while comparable with the two-strap design in all other major

considerations. The 3-strap configuration is, therefore,

selected as the final configuration.

4.4 DESIGN OPTIMIZATION IN 3-STRAP CONFIGURATION

Once the 3-strap configuration had been selected as the

most optimum configuration, the design calculations were made

for optimizing the line resistance, inductance and mass of the

3-strap design. The general optimization guidelines discussed

in Section 3.3 for the 2-strap configuration also apply for

the 3-strap configuration. That is, for a given crosssection of the straps (i.e. for a given resistance) and a

given insulation thickness:

9 Wider straps result in lower inductance and higher

capacitance.

e Narrower straps result in higher inductance and

lower insulation mass.

• The loss x mass product of the conductor remains

constant irrespective of the strap width.

Within these guidelines and the 2-strap parametric

calculations reported earlier, an optimum design for the

initially specified 1000 V, 100 A line is depicted in

Figure 4.4. Notice that the aspect ratio of this cable is

2/0.39 = 5.1. For adequate flexibility in handling such a

cable, the aspect ratio of at least 5 is considered desirable.

4.5 CABLE RATINGS REVISED

The process .which is most likely to be used for manufacturing the line is the extrusion process, in which the

conductor passes through a reservoir of insulating material

under heat and pressure. The insulating material melts and

gets deposited on the conductor continuously. This process

requires an extrusion die which is the heart of the process.

4.07
.010 (0.25)

A Extruded Insulation jacket

95°C Class, Flexible.

•Litz Copper Straps of 10 mil

(30 G) copper strands with

1/2 mil Nylese* insulation

per side.

2.0 1.68

(51) U2)

*Tradename of Phelps Dodge Co.

0.390

(10 mm)

FIGURE 4.4: Three-Strap 2.0 inch wide design for

1000 V, 100 A cable, with aspect ratio of 5.1,

All dimensions in inches (mm in the brackets)

4.08
Two leading cable manufacturers, New England Electric Wire

and Cable Company, Lisbon, NH, and Brand Rex Industrial Cable

Corporation, Willimantic, CT, showed interest in manufacturing such cable. However, they are presently limited

in the extrusion width of 1.5 inches. A wider die has to be

specially made and the wider extrusion process has to be

tried out before making the cable. To avoid the excessive

cost of special tooling for the first prototype, the 1000 V,

100 A cable was redesigned for 1.5 inch width. The aspect

ratio of this design came to 3.2, which is much less than the

desired value of at least 5. After reviewing the available

funds, the aspect ratio constraints and the near term power

requirement, the cable ratings were revised to 600 Vrms and

60 Arms. With 0.85 power factor load, this cable would

deliver about 30 kW power.

The cable was redesigned for the new rating, as shown

in Figure 4.5. Notice the aspect ratio of 38/7.1 = 5.4 which

is mid-way between the desired range of 5 to 6.

It should be pointed out that the extrusion die width

limitation is not of a fundamental nature. Wider dies are

within the state of the extrusion art. They can be made and

used when needed. Furthermore, the cable configuration and

the design method are scalable to higher ratings with no

technical barriers. Hence, the selected cable configuration

has no fundamental design or manufacturing barriers for

scaling to higher ratings.

4.6 LINE PARAMETERS OF THE SELECTED 3-STRAP DESIGN

The final design, as it was released for manufacture,

is shown in Figure 4.5 with 20 mil major insulation.

However, as mentioned in Section 15.0, it was necessary to

increase the major insulation jacket thickness to 30 mils to

avoid tearing in the insulation during extrusion. This

increases the inductance by 19% (not 50%, as the copper-build

4.09
INDUCTION GENERAL, INC.

NASA CABLE

36

(1.42)

35

n.38)

7. 1nn

(0.28 inch)

Outer straps 1.3 mm (.050 inch) thick

37 mm (1.46 inch) wide

Middle strap 3.0 mm (.120 inch) thick

35 mm (1.38 inch) wide

Middle insulation jacket, Polypropylene

0.5 mm (.020 inch) thick

Outer insulation jacket, PVC

0.25 mm (.020 inch) thick

No. of strands in the middle

strap=1056 approximately (copper)

No. of strands in each outer

strap=480 approximately (copper)

Strand Insulation: Solderable

Flexible (Aspect Ratio 5.4)

CALCULATED LINE PARAMETERS

Parameter

Resistance (100°C)

Inductance

Capacitance

Mass

Losses

IX drop

Charging current

Dielectric stresses

Unit

mfl/m

llH/m

yF/m

kg/m

Watts/m

Volt/m

Amp/m

Volts/mil

with major insulation

20 mils 30 mils

1.043

.027

.003

1.102

3.755

.203

.227

30

1.043

.032

.002

1.103

3.755

.241

.151

20

FIGURE 4.5: Three-straps 1.5 inch wide design

for 20 kHz, 600 V, 60 A cable. Aspect ratio 5.4.

Prepared By: Mikund R. Patel

Approved By: ^ItftW A. 1 *&L£

ID No. MRP - 850731

Date: July 31. 1985 Rev 2_

4.10
contributes more in the inductance than the major insulation

gap). The table in Figure 4.5 summarizes the line parameters

of the selected design with both 20 mil and 30 mil insulation,

everything else being the same.

4.11
5.0 MATERIAL SELECTION

5.1 CONDUCTOR MATERIALS

Conductor materials that may be used are listed in

Table 5.1. For a conceptual optimization, consider the

following:

Loss = I2R = I2pL/A

and Mass =

where I = Line current

L,A = length and cross sectional area of the line

p,Y = resistivity and mass density of the conductor

material

Loss x Mass = py!2L

2

 (5.1)

Taking the product L x M (Loss x Mass) as an indicative

figure of demerit, it can be minimized for given I and L by

minimizing the product p x y. Table 5.1 shows that the optimum materials in this respect are sodium and lithium. These

materials, however, are difficult to fabricate and require

special storage. The next optimum material is aluminum, which

is widely used as an electrical conductor. However, it is

difficult to join. Moreover, its low ductility makes it

difficult to draw in fine strands. The next optimum common

conductor is copper. Copper is much heavier than aluminum

but is much easier to join and draw in fine strands. Thus,

from these considerations, aluminum and copper emerge as the

potential candidates for the proposed lines. Aluminum offers

low weight, while copper offers better joints. Both aluminum

and copper have been considered in the parametric study in

Section 3.0. Their electrical and mechanical properties are

given in Table 5.2.

5.01
TABLE 5.1: ALTERNATIVE CONDUCTOR MATERIALS

Material

Aluminum

Beryllium

Copper (annealed)

Lithium

Magnesium

Silver

Sodium

Resistivity,

Micro Ohm-cm

(20°C)

2.828

5.9

1.724

8.55

4.6

1.629

4.3

Density

g/cm3

(20°C)

2.7

1.82

8.89

0.54

1.74

10.5

0.97

Resistivity

x Density

(20°C)

7.64

10.73

15.32

4.62

8.00

17.10

4.17

5.02
TABLE 5.2

ELECTRICAL AND MECHANICAL PROPERTIES

OF COMMON CONDUCTORS

Material

Copper

Aluminum

Density

kg/m3

8640

2700

Resistivity

Viflm

(at 100°C)

.022

.0364

Thermal

Conductivity

w/(m°C)

387

219

Young's Tensile

Modulus Strength

psi psi

15 x 10s 30,000

10 x 106 10,000

5.03
5.2 INHERENT MATERIAL LIMITATIONS ON LOSS X MASS PRODUCT

Ignoring the insulation weight for the time being, we

now derive the inherent material limitations on the Loss x

Mass product with copper and aluminum as the conducting

materials. Using Equations (2.2), (2.4) and (2.5), the L X M

product of 50-meter long line carrying 100 A current is:

LxM =50

= 100 E6xKsKtKc

2pv watts. kg (5.2)

Notice that this product is doubly sensitive to the cabling

factor Kc as compared to the skin effect factor Ks and the

transverse flux eddy loss factor Kt. Manufacturers' catalogs

indicate that K., for the present line would be about 1.15 and \*

the calculations indicate Ks and K^ near unity. At 100°C,

py is 190 E-6 for copper and 98 E-6 for aluminum. Substituting these values in.Equation (5.2), we get:

L x M = 25,128 watts.kg for copper, and

L x M = 12,961 watts.kg for aluminum.

These are the minimum theoretical LxM product values that can

be achieved with these materials at 100°C, regardless of the

configuration chosen.

The LxM product goal set for the 100 A line is

500 x 50 = 25,000 watts.kg, which is about the inherent material

limitation of copper. Since the above values do not include

the insulation weight, we can therefore say that the bare

copper conductor's LxM product just nearly meets the goal.

The mass of the most optimum design with any configuration

with copper conductor will therefore exceed by the mass of the

insulation and the shield.

5.04
On the other hand, the limiting L x M product for aluminum is 12,961, which is much below the goal of 25,000. Aluminum can, therefore, exceed the goal.

5.3 DIELECTRIC MATERIALS

Two kinds of insulating materials will be required, the

major insulation to the ground and the minor inter-strand

insulation.

5.3.1 STRAND INSULATION

A variety of film insulations on magnet wires is available

in the market to meet different ANSI and military temperature

requiements. For example, polyimide film is good for continuous service temperature of 220°C. However, one disadvantage

of this wire lies in stripping the strands for joints at termination. This can be overcome by the use of Nylese (registered

tradename of Phelps Dodge Copper Products Corporation) strand

insulation which permits solder pot tinning when done with

care. There are other types of solderable insulation available.

Of these, Nylese has the highest temperature rating of 130°C.

Such an insulation has been used by Hensen , along with

degassed Urethane polymer potting, in manufacturing a 2.3 kW,

20 kHz power transformer for space applications.

5.3.2 MAJOR INSULATION

The major insulation between the straps should be as

thin as possible for compact and light weight design and for

low inductance. On the other hand, thin insulation raises

the operating electrical stresses and the associated corona

activity. Corona considerations are further discussed in

Section 6.0.

Low dielectric loss factor is an important material

property any major insulation in high frequency application

must possess. The next section discusses some of the candidate materials for the major insulation in the cable under

design.

5.05
5.4 LOW LOSS INSULATING MATERIAL

Polyvinyl Chloride (PVC) was initially proposed as

the major insulating material due to its good extrudibility.

With 0.01 loss factor, the dielectric loss calculated from

Equation 5.1 in Section 5.5 was 175 watts in PVC in the

100 kW cable design, where the conductor losses were 470

watts. This dielectric loss being not acceptable, a new

search for a low-loss insulating material was made. Table

5.3 contains the material properties, and Table 5.4 summarizes

manufacturing aspects of the candidate materials. From these

tables, it is seen that Teflon exhibits the least dielectric

loss factor. However, it cannot be extruded over wide straps

with the presently available manufacturing processes. The

PVC and Silicone rubber exhibit high losses. This leaves two

suitable candidates for the present application, namely, the

Polypropylene and Polyethylene. At room temperature and

1 kHz, the latter has lower loss. However, the stability of

this loss at the operating temperature and frequency is yet

another factor which must be considered in the final design.

The frequency dependence of the loss in Figure 5.1 shows that

both the polyethylene and polypropylene are fairly stable

between 1 kHz and 1 MHz frequencies. However, the temperature

dependence of the loss in Figure 5.2 indicates that polypropylene is much stable with respect to temperature changes than

polyethylene. From these points of view, polypropylene has

been chosen fo'r the major insulation over the middle strap.

For each class of polymer, there are usually hundreds

of different grades available from different commercial

suppliers, each exhibiting some differences in critical

material properties. For this reason, a specific grade of

polypropylene which meets the desired low loss requirement

over the operating temperature and frequency range was searched

from different manufacturers of primary cable insulation. The

grade which was initially identified to meet the requirement

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200

Polypropylene

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copolymer

rPolythene D.923

|Polythene D.934

1 2 10

LOG FREQUENCY

10

FIGURE 5.1: Dielectric Loss Tangent Vs. Frequency

at room temperature for some low-loss insulating

materials (Ref. 10). Polypropylene losses are fairly

stable between 1 kHz and 1 MHz.

5.09
7 \-

6 h

5 k

3 I2 b

I JA-paper (density 1-0)

B-Polyethylene-terepthalate (Polyester,

C-Polypropylene Mylar, Melinex)

40 60 80

Temperature, *C

FIGURE 5.2: Dielectric Loss Tangent Vs. Temperature

for some insulating materials (Ref. 10). Polypropylene

losses are low and fairly stable between 0 and 90°C.

5.10
was Pro-Fax SE-023 of Hercules, Inc., Willmington, DE. This

grade, however, was later found to be replaced with an improved version of SE-191. The manufacturer's catalog sheet

on this insulation is depicted in Table 5.5. Compared to

PVC, this material has three advantages:

• Low dielectric loss (the primary reason for

choosing this as the major insulation).

• It has lower dielectric constant, 2.25 versus

2.7 for PVC. This results in lower dielectric

stresses in voids in series with the solid insulation, thus raising the corona inception voltage of

the cable.

• It has lower mass density, 0.904, versus 1.34 g/cm3,

thus resulting in lower insulation mass.

A disadvantage of Pro-Fax polypropylene is its lower

short time dielectric puncture strength, 600 v/mil versus

1000 v/mil for PVC. However, this is of no real concern,

since the operating dielectric stress in the major insulation

is only 600 v/20 mils = 30 v/mil. The more likely insulation

failure mode in the cable is due to corona degradation, as

analyzed in detail in Section 6.0.

Since the outer jacket contributes heavily to the areamoment of inertia in flexural stiffness of the cable, a

flexible material is desired for the outer jacket to keep the

cable stiffness low. Polyvinyl chloride can be made more

flexible than polypropylene by adding required plasticizer in

it. For this reason, the PVC has been selected for the outer

jacket. The outer jacket sees no dielectric stress, hence no

losses, as the outer straps would be grounded in the normal

operation.

5.11
TABLE 5.5

PROPERTIES OF THE SELECTED MAJOR INSULATING MATERIAL,

PRO-FAX SE-191 PROPYLENE COPOLYMER OF HERCULES, INC.

Propylene Copolymer for Wire and Cable Primary Insulation

CHARACTER

Pro-fax SE-191 is an electrical-grade, high-impact-resistant propylene copolymer designed

for maximum toughness at low temperatures, with minimum sacrifice in tensile strength,

heat resistance, hardness, and processibility. It is especially suited for use in filled

telephone cable because its unique stabilizer system resists extraction by petroleum-jellybased filling compounds.

TYPICAL PROPERTIES

PROPERTY

Melt flow (condition L)

Density at 23° C, g/cm3

Tensile strength at yield, psi (MPa)

(0.075-in. specimen— D 1708)

Tensile elongation at break, %

(0.075-in. specimen — D 1708)

Brittleness temperature, °C

Environmental stress-crack resistance, hrs

Thermal stress-crack resistance, 100°C, hrs

Cu/AI dish stability, % change

Dielectric constant, 1 KHz

1 MHz

Dissipation factor, 1 KHz

1 MHz

DC resistivity, volume, ohm-cm

TYPICAL

PRO-FAX SE-191

2.0

0.900

3,350(23.1)

300

-15

>500

No failures

>500

No failures

<100

2.25

2.25

0.0003

0.0003

>1.0X 10's

TEST METHOD

D 1238L

1 2 at 230° C

D792

D638

2 in./min

D638

2 in./min

D746

REA PE-210

REAPE-210

REA PE-210

D1531

D1531

D257

5.12
5.5 DIELECTRIC LOSS CALCULATION

All insulating materials, when subjected to alternating

dielectric stresses, produce internal heat at a rate

P^ = 2irfe0e«E2 watts/m3 or 2irfC1tan5V2 watts/m length (5.1)

where E = dielectric stress Vrms/m; £0 = permittivity of free

space; e» and tan6 = loss factor and loss tangent of the

insulating material; f = frequency, Hz; V = rated rms voltage

and Cj = capacitance per meter length.

For the cable design of Figure 4.5, the dielectric stress

in the major insulation is 600/.0005 = 1.2E6 v/m. For the

chosen grade of polypropylene for this insulation, the loss

factor e» at 20 kHz and 100°C is 0.0003. The dielectric loss

per cubic meter is then:

Pi = 2TT x 20,000 x 8.85E-12 x .0003 x (1.2E6)2

= 480 watts/m3.

The major insulation volume is

2 x .0005 x .036 = 36E-6 m'/meter length of cable.

.*. Dielectric loss = 480 x 36E-6 = 0.017 w/m loop.

For 50-meter long cable, this loss would be .017 x 50 = 0.85

watt, which is now negligible.

5.13
6.0 CORONA CONSIDERATIONS

6.1 CORONA INCEPTION VOLTAGE VS. MAXIMUM PERMISSIBLE VOID SIZE

A critical quantity needed for evaluating an insulation

subject to corona is Ej_, which is the partial discharge inception voltage stress. This can be obtained directly by a partial discharge inception measurement, from extrapolation of

the volt-breakdown time curve, and from careful analysis of .

the voltage stress at a conductor edge or in a void if the

necessary geometric parameters are known. The dielectric stress

is eE in laminar voids and 3eE/(l+2e) for spherical voids, where

the total void thickness is much smaller than the insulation

thickness and E is the stress in solid insulation. Partial

discharges occur in a void when the product of the pressure

in the gas and the void thickness (electrode separation) exceeds the breakdown voltage according to the characteristic

Paschen curve shown in Figure 6.1 ^11]. it shows that the

electrical strength of a gas decreases with pressure up to a

certain value of pressure, below which the strength again

rises as the pressure drops towards vacuum.

Since the laminar voids produce more stress concentration

than the spherical voids, the former is assumed in the following analysis. If a laminar void of size do is developed

between the middle and an outer strap due to any reason, the

breakdown voltage across such a void will depend on the gas

pressure in the void. Since this pressure may vary over a

wide range in and around the spacecraft at different times

of the journey, we shall use the minimum Paschen breakdown

voltage in the calculations, which is about 297 V peak or

210 Vrms between copper electrodes.

There are now two dielectrics in series, a gap of

laminar thickness do and the solid insulation of thickness d.

If the dielectric .stresses are Eo and E in the gap and the

solid, respectively, the voltage across the copper straps

would be V = Eodo + Ed. With the relative dielectric constant

6.01
OXIDIZED

CATHODE

MCKEL CATHODE

PURE ALUMINIUM CATHODE

COMMERCIAL Al CATHODE

0.5 ".0 1.5

pd (Torr cm)

2.5 J.C

FIGURE 6.1: Paschen Curve, Electrical Breakdown Voltage

versus pd for air (Reference 10).

6.02
of 2.25, the polypropylene insulation would take 1/2.25 times

the gap stress, i.e. E = Eo/2.25. With d = 20 mils as an

example, the above voltage then becomes

V = E0d0 2-25

x 20 = E0d0 + 8.89E0 = Eodo

where do is in mils. The product Eodo is the voltage across

gap, which has the minimum Paschen value of 210 Vrms. The

corona would start at this voltage across the gap. The corona

inception voltage of the cable is then

Vi = 210 1 + 8.89(6.1)

Equation (6.1) is plotted in Figure 6.2, which shows

that it takes about 5 mils delamination gap for corona to

start at 600 volts rms. This condition would not normally

lead to an immediate failure, but eventually would under

gradual corona degradation. Should the gap completely fail,

the dielectric stress in the solid insulation would be

600/20 = 30 v/mil, which is two orders of magnitude lower

than the solid insulation's short time breakdown strength.

As seen in Figure 6.2, the corona inception voltage

decreases with increasing gap size and asymptotically approaches the minimum Paschen value of 210 Vrms. This is because the

Paschen minimum is a constant value, although occurring at

different pressures for different gap sizes. As the gap

increases, the corona inception voltage across the gap remains

constant at 210 V, while the voltage contribution of the

decreasing dielectric stress in the solid insulation decreases.

It should be emphasized that the above corona inception

voltage calculations are based on the worst possible combination of the gap pressure and the gap size (i.e. the minimum

Paschen value). It is likely that there is always some pressure in the void due to continuous outgassing of the insulation.

6.03
Vrms

2000 U

1500

1000

500

10 d (mils)

o

FIGURE 6.2: Corona Inceptiqn Voltage Vs. Laminar Void of size dQ

With increasing void size, the voltage asymptotically decreases

to the Paschen minimum, which remains constant. With large voids,

the solid insulation contributes less.

6.04
The location of the gap along the cable length also matters.

If the gap develops at mid-length of the cable, outgassing may

take a long time due to high gas-flow resistance offered by a

long narrow duct.

For increased reliability, it may be desirable to bleed

some gas (such as nitrogen) along the length of the cable

insulation. This would assure the desired gas pressure to

maintain the required dielectric strength of the cable.

Another alternative is to use a suitable adhesive at the

interface of the solid insulation and the conductor straps to

avoid possible delamination. However, the conductor strands

would now have strong bonds with the insulation, and may

break while splicing. For easy preparation and high reliability of joints and terminations, it is desirable to have

the insulation jackets peelable as in the present design.

Furthermore, the planned manufacturing process requires antistick insulation for easy and smooth extrusion. Adding an

adhesive at the conductor-insulation interface may require

special process development before it can be tried on the

present cable. Another reason for having anti-stick peelable

insulation is the cable flexibility. The•inter-woven strands

in the straps bonded with the insulation jacket at the interface would significantly increase the rigidity of the cable.

Based on the above considerations, we proceeded with the

peelable insulation for the first prototype cable.

6.2 DIELECTRIC STRESS CONCENTRATION IN END INSULATION

In the hollow coaxial configurations, the dielectric

stress between the conductors is nearly uniform, and so is in

the parallel strap configuration. However, the stress at the

ends of the parallel straps is not uniform. With respect to

the grounded external shield, the rounded edge of the conductor

makes a cylinder-to-plane configuration in which the dielectric

stress gets concentrated near the conductor edge. If the

6.05
conductor radius is r and end distance e, as shown in the

following figure, then the maximum dielectric stress, from

the published literature lH], is very nearly

max 5

(6.2)

ground

'//y > s * s

A

t

i

s s s s s s

e

i

-, a d

Dielectric Stress

Concentration at Ends

of Middle (Live) Strap

In Litz straps, ends are generally fully radiused, i.e.

r = a/2. In that case, for the same dielectric stresses at

the ends as in the main region, we must have:

I (1.0.25 2|) - j- or e =

ad

a-0.5d

(6.3)

For example, with a = 3 and d = 1 mm, e must be 1.2 mm,

i.e. 20% more than d, the insulation thickness between the

parallel straps. For smaller radius, say r = a/4, e = ad/(a-d)>

1.5, i.e. 50% more insulation at the ends. Recall that in

Section 4.2, we proposed an indentation of the middle strap to

reduce the fringing flux at the ends, which also serves the

purpose of strengthening the end insulation as discussed in

this section.

6.06
6.3 CORONA RESISTANCE

Among all the insulating materials, mica is the best

known on earth for resisting high corona levels for a long

time. Applying a mica tape next to the conductor strap was,

therefore, considered at one time. However, this idea was

dropped since the cable manufacturers considered it likely to

cause difficulty in extruding a void-free insulation over the

straps.

Further, a possibility of filling the inter-strand

space was also considered to reduce corona and outgassing

in vacuum. The filling can be achieved by pre-impregnating

the straps with a suitable insulation. Such a filling,

however, would hinder soldering operation on the straps at

joints and terminations. Moreover, since the dielectric

stresses within the straps are essentially zero, the interstrand space does not need to be filled. The filling,

therefore, was not recommended in the present design.

6.07
7.0 OTHER CONSIDERATIONS

7.1 ENVIRONMENTAL CONSIDERATIONS

The space environment is different than the earth environment in basically two ways, namely, vacuum and no gravity.

The latter has no bearing in the present design considerations.

Vacuum, on the other hand, influences the design in two major

ways. First, the corona inception voltage of the cable, as

determined by Paschen characteristic curve, will be much

different. Second, the cooling must be done only by conduction

and radiation as opposed to convection which is a primary

cooling mode in electric power equipment on earth. These

effects of vacuum on corona and cooling have been taken into

account in the design.

7.2 MANUFACTURING AND FLEXIBILITY CONSIDERATIONS

As for manufacturing, the hollow coaxial cable appears

to be the most difficult configuration. Manufacturing two

concentric conductor tubes separated by an insulation tube is

extremely difficult, as no clearance can be tolerated from the

insulation point of view. Moreover, this configuration will

be extremely rigid and almost impossible to handle during

manufacturing and installation.

The hollow stranded configuration is somewhat less

difficult to manufacture and less rigid. The strands can

perhaps be spirally wound on the central supporting core.

The core and the insulating tube, however, would make the

cable rigid enough to cause handling difficulty.

The parallel Litz straps configuration appears to be

easy to manufacture. The major insulation can be applied by

established extrusion process. No special manufacturing

equipment will be required, except a special set-up, as this

configuration does not fall in the commercial catagory. The

flexibility in handling and installation is a major attraction

of this configuration.

7.01
As for the conductor material, aluminum offers significant weight advantage over copper. However, our discussion

with Litz cable manufacturers indicated that aluminum, not

being as ductile as copper, is extremely difficult to draw in

thin strands and form Litz cables because of excessive breakage

during the manufacturing process. None of the two leading

Litz cable manufacturers make aluminum Litz cables, nor have

shown interest in experimenting with it.

The above manufacturing considerations indicate that the

parallel strap configuration is much easier to manufacture

than the hollow configurations, and the copper strap is much

easier to make than the aluminum strap with existing manufacturing technologies.

For applying the major insulation, two processes can be

considered. First, the extrusion process in which the conductor

passes through a reservoir of insulating material under heat

and pressure. The insulating material melts and gets deposited

on the conductor continuously. Another process is laminating

in which the conductor straps, sandwiched between the insulation straps, pass through rollers under heat and pressure.

Forming a good end insulation, however, appears to be difficult

in this process. The extrusion process, therefore, was selected for manufacturing the present cable.

7.3 JOINTS AND TERMINATION CONSIDERATIONS

The hollow configuration would, in general, be difficult

to join and terminate due to their concentric nature. The

parallel Litz straps, on the other hand, are easy to join and

terminate, as the two conductors can be sufficiently separated

during splice and termination preparation, brought together

again and insulated by applying insulation tape of required

thickness (usually 50 to 100% more than the factory applied

thickness).

For strand joints, aluminum cannot be soldered nor brazed.

It can only be welded. Welding thin aluminum strands is

7.02
normally impractical, as they would quickly melt around the

welding temperature. Copper strands, on the other hand, can

be soldered or brazed, although brazing thin copper strands

is also difficult. The solderability of copper strands is,

therefore, a desirable attribute for joining strands during

field installations and repairs.

The parallel straps can be easily terminated in conductor

blades and bolted to other blades. Again, bolted joints on

aluminum blades are known to loose pressure and cause arcing

due to their higher mechanical creepage rate. Bolted joints

on copper blades, on the other hand, have been successfully

field proven in heavy power equipment over several decades.

For easy joint and terminations, it is desirable that

the strand insulation be solderable without cleaning, or at

least be chemically cleanable. Cleaning the strand insulation

mechanically is impractical in this case with thousands of

strands involved. As discussed earlier, Nylese (of Phelps

Dodge) permits solder tinning.

Based on the above considerations, the copper conductor

rates superior to aluminum and the parallel strap configuration rates superior to the hollow configurations from the

joints and termination points of view.

7.03
8.0 ALL THE ALTERNATIVES COMPARED FOR FINAL SELECTION

Based on the configuration design studies, and the

manufacturing, flexibility, joint and termination considerations presented in the previous sections, an overall comparison

of various alternatives are summarized in Table 8.1. From

Part A of the table, it is clear that the hollow parallel lines

is the least desirable configuration. The hollow coaxial configurations have reasonably low loss, mass and reactance values.

However, they occupy large volume, are difficult to manufacture,

join and terminate and are extremely rigid in handling. Two

parallel Litz straps, on the other hand, have all the desirable

attributes, along with the excellent manufacturability,

flexibility and ease of joints and terminations, while meeting

the loss, mass and reactance goals. The 3-strap configuration

offers less than half the inductance, while keeping all the

desirable attributes of the 2-strap configuration, which make

this configuration the most attractive.

Part B of the table indicates that aluminum strands

excell over copper strands in loss and mass attributes, but

suffer in manufacturability, joints and terminations. From

the latter considerations, copper is the preferred conductor.

As for the major insulation, PVC has a high dielectric

loss factor. Teflon has the lowest tangent 6, but was considered unextrudable over wide straps by the cable manufacturer.

Polyethylene has low dielectric loss over a narrow temperature.

Polypropylene has low loss over a wider temperature range,

while being easily extrudable. The latter, therefore, has been

selected in the final design.

As discussed later in Section 11.2, magnetic shielding

is recommended, as it will be lightweight and less lossy than

conductive shield.

Based on various considerations presented, the three Litz

straps configuration with copper conductor and polypropylene as

the major insulation was selected for the final design. The

line parameters of the 600 V, 60 A, 3-strap design were given in

Figure 4.5. The other design details now follow.

8.01
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8.02
9.0 THERMAL DESIGN

The cable may be cooled by radiation, conduction or both,

and may be mounted in several alternative configurations. The

temperature rises in three possible cooling configurations are

calculated in this section for the 100 A, 1.5 inch wide cable

with 20 mils major insulation and 10 mils outer jacket. The

temperature rise with thicker insulation will be slightly

higher.

9.1 RADIATION COOLING ALONE

The cable may be mounted in such a way that one face of

it would radiate heat to the free space near absolute zero.

In this configuration, the heat rejection would be primarily

by radiation. To calculate the cable temperature rise under

this condition, we assume 80°K reference temperature at the

moment.

Perfect

thermal

insulation

80°K (liquid

Nitrogen temperature)

FIGURE 9.1: Radiation model under study with one face radiating

to 80°K space and other three cable faces thermally insulated.

9.01
The heat radiated from a body at 8l temperature to the

ambient at 9„ is given by

q = 3.66e

1000

y

 ei\ (loo)

watts/in2 (9.1)

where 9 is in °K (°C + 273) and e is the emissivity of the

radiating surface. Calculating the emissivity of a surface in

free space, with simultaneous radiation and absorption of heat

over a spectrum is involved. On earth, the emissivity of most

industrial non-glossy gray-like paints ranges between 0.9 and

0.95. Jacob ' ' shows that the emissivity of many electrical

insulators exceeds 90%, and that almost all electrical nonconductors have the emissivity of at least 75%. For not

knowing exactly the emissivity of the cable surface in free

space, we calculate the cable temperature rise over a range

of emissivity from 0.25 to 0.90.

The heat must be radiated from one side of the cable

surface at the rate of 4.457 watts/meter or 4.457/(39.37 x

1.5) = .076 watts/in2. Then, the calculations made with

Equation (9.1) are summarized in Table 9.2 at the end of this

section, which shows that the cable surface temperature is

likely to be tens of degrees below zero degree centigrade.

In this temperature, of course, the conduction temperature

rise must be added to calculate the conductor temperature in

the outer most strap. Since the insulation thickness and the

heat flow (watts/in2) are relatively small, the conduction

temperature rise across the solid insulation is less than one

degree, as calculated in the next section.

Thus, if the cable is radiating in free space, it will

run quite cool.

With 0.75 emissivity, the cable surface temperature

would be -40°C. At this temperature, the polypropylene and

PVC insulation may develop cracks. This concern has to be addressed in the future if one face of the cable is exposed to

free space in actual use.

9.02
9.2 CONDUCTION COOLING TO CONTINUOUS BASE PLATE

The cable may be mounted continuously on a base plate

maintained at 60°C. In this configuration, the heat rejection

would be by conduction alone.

The conduction cooling is governed by the basic equation

Q = -KAAT/AX where Q = heat flow rate (watts), K = thermal

conductivity (w/m°C), A = area (m2), AT = temperature difference (°C) across distance AX(m). This can be rearranged in

the equivalent circuit form AT = Q.R where R = AX/KA is the

thermal resistance (°C/W). The thermal conductivities of the

cable materials are given in Table 9.1.

The loss in the cable is 4.457 w/m, 25% of which is in

each of the two outer straps, uniformly distributed over the

cable length. If the right hand surface of the cable in

Figure 9.1 is in continuous contact with a 60°C base plate,

then:

(a) The right hand outer PVC jacket conducts 100% of the

heat through its thickness of 10 mils. The resulting

temperature gradient across this thickness is therefore

4.457 = 0.15 x 1 x ,»':L x AT

39.37 (.010/39.37)

.*. AT = 0.20°C

(b) The right hand major insulation of 20 mils polypropylene

conducts 75% of the heat, with the resulting gradient of

0.30°C, and

(c) The left hand major insulation of 20 mils thickness conducts only 25% of the heat, with the temperature gradient

of 0.10°C.

\: The total temperature rise of the cable above the 60°C

i

base plate is the sum of the above temperature gradients,

i.e. 0.20 + 0.30 + 0.10 = 0.60°C. The operating cable temperature would then be less than 61°C.

9.03
TABLE 9.1

THERMAL PROPERTIES OF MATERIALS PROPOSED

IN THE CABLE CONSTRUCTION

Thermal Steady-State

Conductivity Operating

Material win '* c n Limit (°C)

Copper 387 300°C

Strand Insulation 0.20 130°C

Polypropylene 0.15 120°C

Plasticized PVC 0.15 95°C

9.04
The low temperature rise in this mode is attributed to

the low-loss design plus the wide face area available to

dissipate the heat.

9.3 CONDUCTION COOLING BY INTERMITTENT MOUNTING SUPPORTS

The cable can also be mounted on regularly spaced

supports conducting heat to a 60°C base plate. The objective

of the following calculations is to specify the cooling contact

area and their intervals that will keep the cable temperature

below the thermal endurance limits of the materials' proposed

in the construction. The properties of these materials are

summarized in Table 9.1.

The heat conduction paths are shown by arrows in

Figure 9.2, where it is assumed that shaded areas C, distance

£ apart, on both sides of the cable are in contact with the

conduction coolers maintained at 60°C. The equivalent thermal

circuit and the applicable symbols are shown in Figure 9.3.

As calculated earlier, the total losses in the cable

are 4.457 w/m, one-fourth of which is in each of the outer

straps. This gives q1 = 1.11 watts/m. The copper length

AX = 1/2 and the average cross sectional area of half the strap

(assuming 40% fill-factor) is .0014 x .036 x .40 = 20.2E-6m2.

With Kc = 387 w/m°C (from Table ) , the copper thermal resistance is

Rc = 1/2 / (387 x 20.2E-6) = 64£°C/w.

The heat generation is uniformly distributed over the

entire length, rather than lumped at one place. It can be

analytically shown that in such situations, the temperature

rise, AT, can be calculated by assuming that the total heat

generated in length £/2 is conducted across the average

distance £/4 , the thermal resistance across which is Rc/2 .

AT

 = q, x

 y X~T

 = ~^~A — along the strands.

9.05
1a

2

V

i

d $a

/ /

//

y

d Pa

2

S

N

i

T

CONDUCTION

COOLER

CONDUCTION COOLER

FIGURE 9.2: Conduction Cooling Configuration under Analysis.

Shaded area represents cooling contacts C. Arrows indicate

directions of heat flows.

9.06
.

z

§

w.J

CQ

AT (max)

i

1

o

i

CONTACT

AREA

T2

»Pj AT2 P2 <

'q q '

R

c

 Rc

A T l

 q £

n - * * 2

A T 2

• vVvw *

> 1

For the cable shown in

Figure 9.2

qj = 1.11 w/m

Rc = 64 £ °C/w

Rm = .092/c °C/w

RQ - .046/c °C/w .

i

AT

72

 R -

 AX

R

 ~ KA

AT = QR

3

>~~\Vv ' w u v.R.eiereni;t: itnup ou w

3

 R0

qi£ q/ q;z 1

2 2 2

m

= Heat loss (watts) per meter length of outer strap

= Th. Resistance of copper per meter length of outer strap

= Thermal Resistance of major insulationN

= Thermal Resistance of outer jacket J

= Temperature rise across the branch

= Heat flow (watts) through the branch

per one contact area bxc

FIGURE 9.3: Equivalent Thermal Circuit used for Temperature Rise

calculations with conduction alone. Due to symmetry, only half the

cable thickness is considered in the model. The maximum temperature

rise will be at point P

9.07
The thermal resistances of the major insulation of

thickness d per side is:

R -

 d

 = -5 = .092

m K x(bxc) 0.15x36xc

m -^

and for the outer jacket insulation:

R -25 = .046 0 .

o ~ 0.15x36xc c C/w'

The temperature rises shown in Figure 9.3 are:

qz£

AT2 = Q2 (Rjjj) = —j— (Rm) across the major insulation, and

AT3 = Q3 (R0) = qrl (RQ) across the outer jacket.

The maximum temperature rise is then the sum of the above ATs,

i.e.

ATmax = ATi + AT2 + AT3 '

= a.1 I (0.25R + 0.5R + R ) after combining the terms

 c m o

Substituting the values of R , R , R and q calculated c m p i

above, this equation becomes:

AT =1

max

were t and c in meters. The first term on the right represents

the temperature rise along the copper strands at mid-span

between the cooling contacts, while the second term is the rise

across the insulation at the cooling contacts.

For the manufacture of the cable, we have chosen the

operating temperature limit of 95°C for at least 10 years of

life expectancy. Since the reference temperature of the conduction coolers is specified to be 60°C, we have the total

allowable temperature rise, AT , of 35°C. As can be seen ITlciX

9.08
from the above equations, the maximum permissible contact

span is approximately /35/17.76 - 1.4'm when £ is large.

Under this value, we have two degrees of freedom for limiting

the temperature rise, that is by adjusting the contact area (c)

or contact span (£) or both. This choice can be exercised at

any stage in the installation. However, the following relationship between c and £ must be met to limit ATm= to 35°C:

max

a

35 = 17.76£2 + .102- where £ and c are in meters.

A few possible combinations of £ and c which satisfy

this cooling requirement are:

t = 1.25 m, c = 1.6 cm

£ = 1.00 m, c = 0.6 cm

£=0.75m , c=0.3c m

Thus, for conduction cooling in this mode, we need at

least 0.6 cm wide cooling contacts on both sides of the cable

at one meter intervals.

9.4 SUMMARY OF THERMAL DESIGN CALCULATIONS

The calculated operating temperatures of the cable under

the above three possible heat rejection modes are summarized

in Table 9.2, which shows that the temperatures are well within

the permissible limit except at the lower end. Both PVC and

polypropylene at -40°C and below may pose a concern for embrittlement.

9.09
- TABLE 9.2

OPERATING TEMPERATURES UNDER CONDUCTION ALONE AND

RADIATION ALONE FOR 1000 V, 100 A, 1.5 INCH WIDE CABLE DESIGN

HEAT REJECTION MODE

Radiation to freespace (80°K)

_ "* >» _

3.67e / 9i\ / 90\

qr 1000 \iooy vOOy

watts/in2

Conduction alone

Conduction to 60°C base

plate from both sides

of the cable

COOLING

CONDITION

80°K on one side of

the cable with the

cable surface emissivity as follows:

e = 0.90

ae — — n u 7 . / ^ j

e = 0.50

e = 0.25

60°C base plate

continuous on one

side of cable

one cm wide contacts,

one meter apart

OPERATING

TEMPERATURE

Cable surface

temperature*

-54°C

— m/•j n° LIr

-20°C

+28°C

61°C

Conductor

temperature

+95°C

* Due to the internal temperature rise due to conduction, the conductor

temperature will be one degree higher than the cable surface temperature.

9.10
10.0 MECHANICAL FORCE CALCULATIONS

The mechanical force pattern in the three-strap configuration is such that the outer straps experience repulsion from

the center strap and the center strap will experience the equal

and opposite compression on itself. The repulsive force can

be calculated by using 0.5 pu current in Figure 2.7, which

holds for a wide range of the cable geometry. However, for the

chosen geometry of extra-wide, closely-spaced straps, this

force can be accurately calculated from the following analysis.

The flux density Bx at the innerface of the outer strap is:

Bx = yQHx = y0 x °-5 * Il/2" Tesia (peak)

where I is the cable current and b is the cable width (height) .

The flux density at the outer face of the outer strap is zero.

The average flux density in the current carrying region of the

outer strap is, therefore, Bx/2, i.e.

Bavg = ^o1

/

415

 Tesla.

The force per meter length is then:

Fi = Bavg Il/? = 2y0I2/4b = yQI2/2b N/m (peak).

If the straps were totally bonded face-to-face with the

major insulation, the tensile stress in the bond and the

insulation will be:

Fi Ho*2

a, = —r- = - N/mz (peak) .

2b2

For the given cable, this turns out to be:

= 4TTE-7 x 60 = 1>6 N/m2 = 0.0003 psi (peak)

D

 2 x .0372

which is quite small compared with the ultimate tensile

10.01
strength of about 2000 psi for the insulation. Since both

polypropylene and PVC are anti-stick (peelable) materials,

the strap would not be totally bonded face-to-face. In that

case, this force is taken by the end insulation e (see sketch

in Section 6.2), and the tensile stress in the insulation

would be:

a

e

 =

j2

~4be~

N/m2 (peak)

With e = .0015 ra, we have:

°e -

which is very small to be of any concern.

The stresses are very small because the extra-wide straps

chosen to keep the inductance low produces low flux density.

The three-strap configuration is yet another reason for the

low mechanical stresses.

10.02
11.0 ELECTROMAGNETIC INTERFERENCE

11.1 EMI FLUX DENSITY ESTIMATE

Electromagnetic interference radiating from the cable

under development is a major concern. One of the applicable

standards (MIL-STD-462, Section RE01) limits the radiation at

7 cm away from the worst radiating point to be less than

27 db above 1 picotesla. This flux density is difficult to

calculate without a detailed finite-element electromagnetic

study. However, it is shown below why it is expected to be

low.

The "virtually concentric" 3-strap cable configuration

selected for the final design can be viewed as two cable pairs,

put side-by-side, carrying equal and opposite current of onehalf the magnitude. One such pair, shown in the absence of the

other in Figure 11.1, will radiate flux like a solenoid. The

flux densities from the end (radiating point) of one such

solenoid are calculated and tabulated under Figure 11.1 using

the classical thin solenoid theory. It is seen that the flux

density rapidly falls with the distance and becomes . 042yT at

7 cm. If a one-turn search coil, one cable thickness (7.1 mm)

wide and one meter long, is placed at this location, it will

see the induced voltage of 26.7yV at 20 kHz (assuming uniform

flux density within the search coil).

Now, if the other half of the cable is placed next to

the first half considered above, it will produce about -26.7yV.

The net resultant flux density will then be about zero, a subtraction of one "large" number from the other "large" number.

The exact value of this subtraction could not be readily

calculated by any closed form formula.

It should be pointed out that the above flux density estimates are without a shield around the cable. Any shield placed

around the cable will reduce the above calculated flux densities,

11.01
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11.02
11.2 EMI SHIELDING

One consequence of the three-strap configuration is that

most of the magnetic flux is now internally contained, thus

minimizing the electromagnetic shielding requirement. Furthermore, the indentation of the central high voltage strap with the

grounded outer straps reduces also the electrostatic shielding

requirement. Since the EMI field generated by the cable is

mostly due to fringing, it is difficult to establish by analysis

whether or not the shield is required. However, should the EMI

tests indicate a shielding need, it can be applied later on.

There are two alternatives for the shielding material,

the conducting material or the magnetic material. Experience

in such situations indicate that the conducting shield would

produce more losses with no extra benefits. The conductive

shield keeps the flux away by the internally induced eddy

currents, while the magnetic shield guides the flux along the

designed paths. Therefore, if the shield is needed, the magnetic shielding material is expected to be better than the

conductive shielding material.

11.03
12.0 FINAL DESIGN SUMMARIZED

12.1 DISCUSSION ON THE FINAL DESIGN

The final design, as developed in the previous sections,

has been depicted in Figure 12.1. The loss, mass, inductance

and capacitance per meter length of the cable with the revised

ratings are summarized in Table 12.1, where the mechanical

forces and temperature rises under conduction alone and radiation alone are also shown. In actual use, radiation will be

the principle mode of heat rejection. The operating temperature of the cable, therefore, is expected to be tens of degrees

below 0°C, depending on the exact emmissivity of the finally

installed cable surface.

It is observed from the last column of Table 12.1 that

the conductor loss in the revised 60 A cable is now 0.010%

per meter versus 0.01% goal set for the original 100 A cable,

and the mass 0.031 kg/kW/m versus 0.01 kg/kW/m goal for .the

100 A cable. Since the cable current is now 0.6 times the

original 100 A rating, the loss and mass both would now be

reduced to 0.6 each, bringing the loss x mass product down to

0.6 x 0.6 = 0.36. Had we kept the same cable cross section, the

loss would have been reduced to 0.36 (in proportion of I2) and

the mass remained the same (1.0), thus the mass x loss product

again reducing to the same 0.36 ratio.

Had the original goals of .01% loss/m and .01 kg mass/

kg/m were to be uniformly revised, they should now become

1/0.6 times the original values. The new goals should, therefore, be .017% loss/m and .017 kg mass/kW/m, respectively, with

their product being 289 E-6. The last column of Table 12.1

indicates that these numbers for the revised 60 A cable, as

designed, are .010% loss/m and .031 kg mass/kW/m, and their

product 310 E-6. The last number, however, is calculated at

100°C. Since the operating temperature of the cable is expected to be much lower, the actual loss x mass product would be

significantly lower than 310 E-6, thus meeting the goal. The

loss is, however, lower and the mass higher than the uniformly

revised goals. In other words, the design now gives the lowloss attribute a higher importance than the low-mass attribute.

12.01
INDUCTION GENERAL, INC.

NASA CABLE

36

(1.42 )

I 35

T1.38)

7.1mm

(0.28 inch)

Outer straps 1.3 ram (.050 inch) thick

37 mm (1.46 inch) wide

Middle strap 3.0 mm (.120 inch) thick

35 mm (1.38 inch) wide

Middle insulation jacket, Polypropylene

0.5 mm (.020 inch) thick

Outer insulation jacket, PVC

0.25 mm (.020 inch) thick

No. of strands in the middle

strap=1056 approximately (copper)

No. of strands in each outer

strap=480 approximately (copper)

Strand Insulation: Solderable

Flexible (Aspect Ratio 5.4)

CALCULATED LINE PARAMETERS

Parameter Unit

with major insulation

20 mils 30 mils

Resistance (100°C)

Inductance

Capacitance

Mass

Losses

IX drop

Charging current

Dielectric stresses

yF/m

kg An

Watts/m

Volt/m

Amp/m

Volts/mil

1.043

.027

.003

1.102

3.755

.203

.227

30

1.043

.032

.002

1.103

3.755

.241

.151

20

FIGURE 12.1: Three-straps 1.5 inch wide design

for 20 kHz, 600 V, 60 A cable. Aspect ratio 5.4.

Prepared By: Mikund R. Patel

Approved By: /I ' tu/*"** A , / **£&(

ID No. MRP - 850731

Date: July 31, 1985 Rev.

12.02
TABLE 12.1: DESIGN SUMMARY OF 600 V, 60 A, 1.5 INCH

WIDE CABLE, FINAL DESIGN WITH 20 MIL MAJOR INSULATION

DESIGN PARAMETER

PER METER

LOOP

% OF RATING

PER METER *

AC Resistance at 100°C, 20 kHz

Conductor Loss (including 2.4%

eddy loss)

Inductance

IX voltage drop

Capacitance

Charging current

Conductor Mass

Insulation Mass

Force separating the straps

Tensile stress in end-insulation

Dielectric "stresses in major

insulation

Dielectric Loss in Insulation

Conduction cooling contacts

required for 95°C operating

temp.

Operating temp, if one face

radiating to freespace at

80°K

EMI shield requirement

1.043 mfi/m

3.755 watts/m

0.027 yH/m

0.203 v/m

0.003 pF/m

0.227 A/ra

1.102 kg/m

0.057 kg/m

0.065 N/m

20 N/m2

1200 v/mm

0.017 watts/m

.010%

.034%

.378%

.031 kg/kW/m

(.00036 Lbf/inch)

(.003 psi)

(30 v/mil)

1 cm wide, 1 m apart

Tens of degrees below 0°C

Small

* This column should be compared with the proportionately

modified goals for the revised 60 A cable, as discussed

in Section 12.1.

12.03
As seen in the above paragraph, the loss x mass product

of the final design is as close as we can get, since the loss x

mass product is invariant for a given current and a chosen

conductor material. That is, the .loss can be decreased by some

percentage only by increasing the mass by about the same

percentage. Within this fundamental engineering constraint,

these two factors can be adjusted in any desired manner.

The major advantage of the 3-strap design is the significantly lower inductance and lower EMI than in the 2-strap

configuration . Moreover, it offers the same flexibility in

manufacturing, handling, joining and terminating.

12.2 SUMMARY OF BASIC FORMULAS FOR 3-STRAP DESIGN

The formulas for calculating the line parameters R,L,C

and the mass for the parallel Litz straps configurations were

developed during the contract period. These formulas for the

3-strap configuration are summarized in this section in SI

units with the following symbols (see Figure 12.2):

1

e

FIGURE 12.2: Dimension symbols of the 3-strap cable used in

summarizing the .design formulas

12.04
V, I = voltage and current ratings of the cable (rms).

a = thickness of the middle strap,

b = width (height) of the middle strap,

d = major insulating distance between straps,

e = end insulation over the middle strap.

Kc = cabling factor by which the spiraling strand length is

greater than the straight cable length (-1.15).

K_ = skin-effect eddy loss factor. 5

Kj- = transverse flux eddy loss factor.

Kf = fill-factor of the straps (conductor cross-section area/

strap's gross cross-section area).

p,Y = resistivity and mass density of the conductor material,

y ,e= permeability and permittivity of free space.

e,e« = absolute permittivity and loss factor of the major

insulating material between straps.

E = dielectric stress in major insulation (rms).

f = frequency of operation.

do = conductor strand diameter.

N = number of strands in the middle strap.

For shielding and end-insulation considerations discussed

earlier, the outer straps would be slightly wider than the

middle strap width and thinner than one-half of the middle strap

thickness. The dimensions are adjusted such that the middle

strap cross-section equals that of the two outer straps combined.

Then, the line parameters per meter length of the cable

are given by the following closed-form formulas:

Resistance R = K K.K v

 P. ohms/m loop (12.1)

ac i sec i\..pct.D

T , . ,

 Uo /, , a\ H/m loop (12.2)

Inductance i =2Ev3 /

Reactive Drop IXl = 2-irfI^I v/m loop (12.3)

Capacitance C1 = —4^— F/m loop (12.4)

12.05
Charging Current II =

Dielectric loss P.

Force F. =

Conductor Mass M,

27rfC1V A/m loop

2irfC1 tan6V2 watts/m loop

watts/m3 of

insulation

peak N/m loop

2b

= 2K Kfaby kg/m loop

(12.5)

(12.6)

(12.7)

(12.8)

In the above formula for R , the skin-effect eddy loss

cLO 1

factor KS for 10 mil diameter copper strands at 20 kHz is very

nearly unity. For practical Litz straps, the cabling

(spiraling) factor KC is about 1.15 and net copper fill-factor

is about 0.4. The transverse flux eddy loss factor K^ for the

3-strap configuration is one-fourth of that derived in Appendix A of the First Quarterly Report since the flux density in

the major insulating gap is now one-half of that in the twostraps configuration. The factor Kt is then:

Kt = 1 +

0.2f2N

2dQ6

p

2b2E!2

(12.9)

For practical designs, this factor should generally fall

between 1.0 and 1.05, i.e. the eddy losses generally kept

below 5% of the DC I2R losses

12.06
13.0 FAILURE MODES AND EFFECTS ANALYSIS

The reliability is as an important criteria in selecting

the design configuration and the manufacturing process. A

failure mode and effects analysis on the cable was made.

Three possible causes of failures and their effects have been

identified as follows:

13.1 EXPOSURE TO FREE-SPACE TEMPERATURE

As calculated in the thermal design, if one face of the

cable is radiating to free space, the operating temperature

at rated current (60 A) would be tens of degrees below 0°C.

When the cable is under no-load, the cable temperature would

be even lower, approaching the free-space temperature, which

is assumed to be 80°K (-193°C) in the present study. At this

temperature, most insulating materials become brittle and may

crack. When resins are cooled down, they pass through a glass

transition temperature, below which they become glassy (brittle)

The selected insulations, the polypropylene and PVC, both have

glass transition temperatures above 80°K. Both would, therefore, be brittle at 80°K. However, since the cable carries

no mechanical load, other than its own mass, the chances of

brittle cracks may be small.

Should a brittle crack develop across the thickness of

the major insulation, it may lead to electrical failure over

a period of time." A possible suggestion may be to use cryogenic insulations, such as Mylar and Kapton of E. I. DuPont,

which have been successfully used in superconducting power

equipment. These materials,.however, have high dielectric

losses, which are not of a concern in DC cryogenic applications,

but are not practical for the present 20 kHz application.

13.2 EXPOSURE TO ULTRAVIOLET RADIATION

The cable side facing the free space would be exposed to

ultraviolet radiation. Most polymer materials gradually

13.01
degrade under such exposures. Chemical additives are generally

used to stabilize polymers against UV radiations. In the

present application, it is likely that the outer PVC jacket

insulation may not be directly exposed to the radiation. If

a magnetic metal tape is wrapped around the cable for EMI

shielding, the insulation would get adequate protection against

such radiations. In case the EMI .shield is not required and is

not used, a Teflon tape may be wrapped around to protect the

cable. Teflon is much more stable against UV rays.

13.3 VOIDS IN INSULATION

There are two kinds of voids that we should be concerned

about. First, the void within the extruded insulation. The

maximum permissible void size of this kind is calculated to be

5 mils in Section 6.0. The voids left in the extrusion process

under heat and pressure are much less than half a mil. This

failure mode is, therefore, not likely.

The second kind of void that may develop in the cable is

under compressive force on the cable edges as shown in the

Table 13.1 sketch. Under edgewise compression, the flat cable

would tend to deform towards a spherical shape. If this

deformation is under constant volume, no void formation is

possible. The solid insulation would deform under essentially

a constant volume, but the braided Litz straps may not,

depending on how the strands get internally readjusted. Corona

may start under such voids.

13.02
TABLE 13.1: FAILURE MODES AND EFFECTS ANALYSIS

CAUSE EFFECT

FAILURE

MODE

POSSIBLE

REMEDY

Exposure to free

space near absolute zero (especially at

no-load)

Embrittlement of

Insulation,

leading to possible cracks in

major insulation

Electrical

failure

Need suitable insulating

material with low tan 5.

and cryogenic glasstransition temperature.

Exposure to free

space UV

radiation

Insulation instability against

UV radiation

Gradual degradation of

the insulating

material

o Add additives in the insulation to stabilize

against UV

o Wrap metal (EMI shield)

or Teflon tape over to

protect insulation

from UV exposure

Excessive compression between cable

edges in installed condition

Buckling of major

insulation,

leading to gaps

between conductor and insulation (which

does not stick

to the insulation, see Note

and Figure)

Corona discharges,

leading to

electrical

failure

o Make tests on manufactured cable and establish

this possibility and/or

its compression limit.

Include this in Installation QC

o Consider bleeding nitrogen

along the cable insulation

length

o Consider adhesive bond

between insulation and

conductor straps

NOTE: Antistick insulation is desired for smooth extrusion

process, good cable flexibility and easy jointing and

terminations.

Possible delamination (void

formation) under compressive

force on cable edges

13.03
14.0 SAMPLE CONSTRUCTION AND TESTS

14.1 ONE METER LONG SAMPLE

After completing the final design and the subsequent

Design Review at NASA LeRC, the construction of one-meter long

sample was undertaken. The set-up for the sample construction

is shown in Figure 14.1, which also includes major steps of the

process. The actual process chosen for the 150-meter line is

to extrude insulation over the Litz straps (with insulated

strands). However, since Litz straps are not stock items and

making one-meter Litz straps is not economical, the sample

used readily and economically available bare copper straps.

Such straps were vacuum-pressure impregnated with insulating

varnish by the process shown in Figure 14.1. The straps were

then tape-insulated by hand. Two passes of 5 mil, one-inch

wide polypropylene tape were applied on the middle jacket in

half-lap manner, thus building 20 mils jacket per side. The

outer straps were then placed and the outer jacket applied

again by hand-taping. Figure 14.2a, b and c show this cable.

Due to very low voltage required to test this sample,

it was not possible to use the 100 kW, 1000 Vrms power supply

which has been procured for testing the full cable. The

sample was then tested with 1 kHz inductance-capacitance

bridge and a DC milliohm meter. The measured inductance

did not agree with the calculated value because the inductance

of this short sample was much lower compared to both the lead

inductance and the meter resolution.

Therefore, it was decided to make new samples with higher

L so that it can be more reliably measured with the available

bridge. These samples are described in the next section.

14.2 3.7 METER LONG SAMPLE

This sample was constructed with 3 parallel solid copper

bars with dimensions shown in Figure 14.3. They were separated

by 1/8 inch thick 0.75 inch diameter dimensionally stable

14.01
Vacuum/Pressure

Pump

Varnish

Reservoir

MAJOR PROCESS STEPS;

1) Close valves VI, V2, V4 and open V3, V5.

2) Start vacuum. Keep it on for an hour. Stop vacuum pump.

3) Close valves V3 and V5.

4) Open VI. Let varnish enter the sample by suction as much as it can

5) Close VI. Open V4 and V2. Keep V3 closed.

6) Start compression. Keep it on for an hour. Stop compression pump.

7) Close V2 and open V3. Le?ve it this way overnight.

8) Open VI slowly and drain the varnish completely.

9) Open left end of the sample tube and remove the sample.

10) Lay copper straps flat, sandwiched between polypropylene straps.

11) Let the assembly dry overnight under weight.

12) Wrap the assembly with adhesive backed insulating tape.

FIGURE 14.1: Equipment set-up and major process steps for constructing

one meter-long sample. The varnish is vacuum-pressure impregnated on

bare copper straps, which were then tape-insulated by hand.

14.02
ORIGINAL FACE F3

OF POOR QUALITY

(a) Construction of one-meter cable

sample. Outer jacket being taped.

(b) Finished one-meter sample

(c) The sample coiled on a 10 inch

diameter circle, showing its

flexibility (even with cured

varnish on strands).

(d) 1000 V, 100 A, 20/10 kHz

power supply received for

testing full cable at full

current.

FIGURE 14.2: ONE-METER SAMPLE CONSTRUCTION AND 100 KW POWER SUPPLY

14.03
4.8 3.2 4.8 3.2 4.8 mm

X

y

X

X

H Mh

/>

X

Solid Copper Bars

Length = 3.7m

6 , . = 2.2mm

skin

1 kHz

eavg ~

25.4

La

Parameter Calculated Measured

(A) 3-Bars (Concentric) Configuration:

R

JyH

:PF

tafl

DC

.43

910

.78

.4

1005

.8

(B) 2-Bars (Middle-Outer) Configuration:

L H .854 .8

455 509

PF

*mfi

DC

1.04 1.0

Leads

(deducted)

.7

1

0

.7

1

0

FIGURE 14.3: Geometrical Dimensions and Test Results of 3.7 m Sample

14.04
glass-fiber spacers (e = 6) at 5 inch center-to-center spacing,

and then banded together. The "continum average" dielectric

constant of this arrangement was calculated to be 1.75. Since

the inductance measurements were to be carried out at 1 kHz,

the skin depth of 2.2 mm in the straps (shaded area in

Figure 14.3) was used in calculating the inductance at this

frequency. The DC resistance was calculated using the full

strap area.

Two sets of measurements were taken, one in the 3-bar

"virtually concentric" configuration and the other in the 2-bar

configuration. The calculated values, along with the measured

values of the R,L,C parameters for the above two alternative

configurations, are reported under Figure 14.3. It is seen

that the test results compare well with the calculated results.

Notice that the inductance in the 2-bar configuration is

almost twice that in the 3-bar configuration, as predicted.

14.3 FIVE METER LONG SAMPLE

One more sample, with 5-meter long straps of uninsulated

strands and dimensions and separations closer to that adopted

for the final design, was constructed as shown in Figure 14.4.

The straps were separated by 70 mil thick solid (dimensionally

stable) Quinergo Board 1200 (tradename of Quintex Corporation),

which has the dielectric constant of about 7.5 in the raw

(unheat-treated) form. This sample was tested four ways:

(1) in a 3-strap configuration, (2) between middle-to-outer

left strap, (3) between middle-to-outer right strap, and

(4) between outer-to-outer straps. All these test results are

summarized under Figure 14.4. They check reasonably well with

the calculated values except for the inductance in the first

configuration. The reason for this discrepancy could not be

identified. However, it is suspected that large resolution and

large lead inductance (although minimized by further twisting

and shielding) compared to the quantity to be measured may be

partly responsible for this discrepancy.

14.05
II. 3.2 |l.8l.6 mm

25.4

35

Stranded Straps With Solid

Insulation Inbetween

"it"" Length = 5 meters

Straps same as in Sample No. 1,

except separated by 70 mil thick

solid insulation strips (instead

of hand-taped).

e = 7.5

Parameter Calculated Measured

(A) 3-Straps (Concentric) Configuration:

L „ .36 .6

^PF

Rmfi

DC

11,060

5.42

10,130

5.7

(B) 2-Straps (Middle-to-Outer Left) Configuration:

L „ .84 .9

*PF

5530 4910

Lead

(deducted)

.6

1

0

(C) 2-Straps (Middle-to-outer Right) Configuration:

L „ .84 .9

'PF

5530 5040

(D) 2-Straps (Outer-to-Outer) Configuration:

L „ 1.41 1.5

'PF

3230 2510

.6

1

.6

1

FIGURE 14.4: Geometrical Dimensions and Test Results of 5 m Sample

14.06
The general agreement between the calculated and measured

values of the inductances and capacitances in the 3.7 and

5-meter long samples established the validity of the virtually

concentric 3-strap cable concept and the analytical formulas

developed for calculating its line parameters. At this stage,

the 150-meter long full cable construction was released.

14.07
15.0 CONSTRUCTION OF FULL CABLE

The construction of 150-meter long, 600 V, 60 A cable

followed the base line design depicted in Figure 12.1. There

were a few manufacturing problems which needed to be resolved.

These constructional difficulties are discussed here for

future reference.

(1) Manufacturing the Litz braids to an accurately

predictable dimension was difficult. For example, the

finished unextruded strap measurements made by the

writer at New England Electric Wire Corporation indicated

the following sizes in inches:

Middle Strap Outer Strap

Thickness

Width

Actual

.140

1.50

Planned

.120

1.38

Actual

.060

1.60

Planned

.050

1.46

(2)

Thus, the strap sizes came somewhat larger than the

planned. The increased thickness tends to increase the

reactance, while the increased width tends to decrease

the reactance. The thickness and width increasing together, therefore, partly compensate each other in determining the cable reactance.

To reduce the Litz straps to the planned sizes, the

straps were run one more time through compactation rollers

and the extrusion die was slightly opened-up to accomodate

the increased width.

The middle strap was then jacketed with polypropylene

insulation by the extrusion process. Because of some

concerns about tearing the insulation while extruding over

a rough strap surface, a 30 mil thick jacket was applied

instead of the 20 mils planned. The jacketed middle strap

was tested for pinholes by a 4000 Vrms high-potential test

on the cable. No pinholes were detected in this test.

15.01
The jacketed middle strap and the two outer straps

then proceeded to the third step in manufacturing the

finished cable, i.e. extruding the PVC jacket over the

three-strap assembly. In its first attempt, the following

problems were faced:

(a) The outer straps were wider than the middle strap

before going in the extrusion die. However, in the

finished cable on the other end of the die, the outer

straps became somewhat narrower (by about 60 mils)

than the middle strap, and also somewhat asymmetrically placed with respect to the middle strap (see

Figure 15.1). It is suspected that the extrusion

process built up pressure at four corners of a

rectangular die, which in turn pushed the outer

straps inward widthwise. This manufacturing phenomenon tended to increase the cable reactance.

(b) Due to the above pressure build-up, the PVC material

penetrated the corners and also further down inbetween

the middle and outer straps. This affects the cable

design in three ways. it increases the reactance,

increases the overbuild and decreases the flexibility.

(c) The first sample had a significant overbuild (almost

50%), mainly due to the PVC penetrating inbetween

the straps.

As a result of these manufacturing problems, it was

decided to eliminate the PVC penetration between the straps by

wrapping a 1 mil Teflon tape, half-lapped, over the three-strap

assembly before extruding the PVC jacket and by making appropriate changes in the die. After these measures were taken, the

cable production was completed.

15.02
I-040

(a) This is how we

planned

c

about .060'

(b) This is how it came

from extrusion die

in the first attempt

(c) It appears that the fluid pressure

in the die built up at the corners.

This pushes the outer straps widthwise, making them shorter than the

middle strap.

FIGURE 15.1: Manufacturing Problems Encountered

15.03
16.0 TESTS ON FULL CABLE

A complete Test Plan was prepared and submitted to NASA

for approval. It was accepted and followed for the tests

reported in this section.

16.1 AC CORONA TESTS

AC corona tests were made on flat 5 mil thick samples of

polypropylene, the insulating material selected for the major

insulation between the straps. These tests were made at 60 Hz

using ASTM standard brass disk electrodes as shown in Figure 16.1

Two test series were conducted, one on a single layer of

5 mil thick sample and another on sets of 4 layers dielectrically piled in series (i.e. 20 mil total insulation thickness

between the disk electrodes). The test results are summarized

in Table 16.1. • The corona inception voltage was recorded when

the corona discharge exceeded 5 picocoulombs and the corona

extinction voltage was recorded when the discharge fell below

the 5 pc level.

The test results indicate that for 20 mil polypropylene

insulation, the corona starts at 1550 volts (rms) and extincts

at 1400 volts rms. This is comfortable for a cable with 440 V

operating voltage (a safety factor of 3).

16.01
Brass Disk

Electrodes

-Insulation samples

under tests

FIGURE 16.1: Corona Test Set-up

16.02
TABLE 16.1: CORONA TEST RESULTS ON FLAT INSULATION SAMPLES

Report No. 470057 Issued: 9/30/85

Equipment Used

Corona Test Set

Sensitive Research

Electrostatic voltmeter 81247

N065

V028

Cal. Date

6/86

11/85

Test Method

The samples were placed into the Corona Test Set between two A.S.T.M.

brass electrodes. Corona inception and extinction voltage was then

measured on single samples and on sets of four piled dielectrically

in series.

Test Results

Single Sample

1

2

3

4

5

Inception Voltage

890

895

860

890

860

Extinction Voltage

850

845

820

855

810

4 Layer Samples

1

2

3

4 •

5

Inception Voltage

1550

1600

1550

1550

1600

Extinction Voltage

1400

1450

1400

1400

1450

16.03
16.2 R,L,C TESTS

Two -full cable lengths, 65 m and 123 m, were tested.

Both the bridge measurements (Table 16.2) as well as the full

current (short circuit) and full voltage (open circuit) tests

under 20 kHz power excitation were made. The circuit diagram

for the latter tests is shown in Figure 16.2. The results

are tabulated for accuracy in Tables 16.3 to 16.6, and also

plotted in Figures 16.3' to 16.5 to check the trends.

Table 16.7 is a summary comparison of the calculated values

with the test results. The inductance is higher and capacitance is lower than calculated due to the oversize that resulted in the final manufacture.

The capacitance is proportionately more lower, probably

because the dielectric constant of the polypropylene is not

precisely known at 20 kHz. The manufacturer's data gives

e = 2.25 only at two frequencies, 1 kHz and 1 MHz. We have

assumed the same value at 20 kHz in the calculations, although

it could be actually somewhat different.

The AC resistance values at 20 kHz are fairly close to

the calculated values within the manufacturing tolerances.

The test results reported in this section on full cable

once again prove the virtually concentric 3-straps cable

concept and also verify the formulas developed for calculating

the circuit parameters.

16.04
TABLE 16.2: ' BRIDGE MEASUREMENTS ON FULL CABLE LENGTHS

PARAMETER

Resistance

at DC (mfi)

Inductance

at 1000 Hz

(pH)

Capacitance

at 1000 Hz

(yr)

ON 65 m CABLE

Middle

to

Outer 1*

84.08

5.5

.0362

Middle

to

Outer 2

84.36

5.2

.0423

Middle

to

Both Outer

53.63

2.5

.0751

Per

Meter

.825

.0385

.00116

ON 123 m CABLE

Middle

to

Both Outer

102.2

4.3

.168

Per

Meter

.831

.0350

.00137

*This column gives the parameters between the middle to one of the outer straps.

The next column are the parameters of the middle to the other outer strap and

the next one is from the middle to both of the outer straps in parallel. Notice

that the inductances and the capacitances of the two outer straps individually

are not perfectly symmetrical with respect to the middle strap. This is attributed to the manufacturing variations. As discussed in Section 16.3, this

asymmetry increases the EMI.

16.05
I/C2 1/2

S = Power Supply

20 kHz, 100 Arms, 1000 Vrms

CABLE A

-J

Open Circuit

or

Short Circuit

(t

1

I;

==

*n7

y'

X

/

/

V

X

X

/

^

X

/

X

J

=

X

!.

•I*

1

— "v-"J-"~

^_^^

1^*r*

|

'v* ••^T .*

I

t~

^

X

»

»

1

«

^

a

»

x

/

/

/

?/

//

•

^

—

,•>

-

Tnciil at ion

_^ Litz Straps

EMI Shield

(if require

SECTION-AA

FIGURE 16.2: R,L,C Test set-up for 20 kHz

Open-Circuit and Short-Circuit Tests on Full Cable

16.06
TABLE 16.3: SHORT CIRCUIT TESTS ON 65 m CABLE

Current

Amps

20

• 30

40

60

70

80

Voltage

Volts

6.30

9.31

12.59

18.79

21.76

24.76

Power

Watts

22

48

85

191

259

334

Power

Factor

.184

.168

.163

.164

.165

.165

Frequency

kHz

20.07

20.02

20.02

20.00

20.00

19.90

TABLE 16.4: OPEN CIRCUIT TESTS ON 65 m CABLE

Voltage

Volts

200

400

600

800

Current

Amps

1.784

3.556

5.317

7.017

Power*

Watts

4

15

31

53

Power

Factor*

.017

.011

.010

.010

Frequency

kHz

19.23

19.23

19.23

19.09

*These readings, although reported, are to be disregarded.

The nearly perfect capacitive nature of the cable under

open circuit made these power and power factor measurements

extremely inaccurate due to the instrument errors.

16.07
TABLE 16.5: SHORT CIRCUIT TESTS ON 123 m CABLE

Current

Amps

20

40

60

80

Voltage

Volts

10.80

21.39

32.20

42.70

Power

Watts

43

168

390

680

Power

Factor

.199

.191

.196

.199

Frequency

kHz

19.90

19.76

19.80

19.70

TABLE 16.6: OPEN CIRCUIT TESTS ON 123 m CABLE

Voltage

Volts

200

403

602

801

Current

Amps

4.067

7.998

11.960

15.830

Power*

Watts

11

40

100

190

Power

Factor*

.013

.013

.013

.014

Frequency

kHz

18.95

18.75

18.81

.18.73

*These readings, although reported, are to be disregarded.

The nearly perfect capacitive nature of the cable under

open circuit made these power and power factor measurements

extremely inaccurate due to the instrument error.

16.08
Volts

50

123 meter

(Table 16.5)

40 -

30 - 65 meter

(Table 16.3)

20 -

10 -

Amps

FIGURE 16.3: Reactive volts vs. Current during Short-Circuit Tests

on full cable. The dots represent test data, which fall on linear

scales, as expected.

16.09
Watts

1000

500

200

100

50

20

10

.i'-imi:

J.-L-J.-..

... --^ .,

r

:

~l

.-:.:.'.I-

' " r

12.5 20 30 40 50 60- 70 80 Amps

1. Lower curve represents 65 meter long cable (Table 16.3)

2. Upper curve represents 123 meter long cable (Table 16.5)

FIGURE 16.4: Loss vs. Current during Short-Circuit Test

on full cable. The data (dots) fall on linear scale on

log-log paper, as expected.

16.10
Amps

16

15

14

13

12

11

10

123 meter

(Table 16.6

65 meter

(Table 16.4)

200 400 600 800 Volts

FIGURE 16.5: Charging current vs. Voltage during Open Circuit Tests

on full cable. The data (dots) fall on linear scale, as expected.

16.11
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16.12
16.3 EMI TESTS WITH CABLE TERMINALS INSIDE THE CHAMBER

EMI tests in accordance with MIL-STD-462, Sections RE01,

RE02 and RE04 were made at MET Electrical Testing Company in

Baltimore, MD. The tests were made in a shielded chamber with

the cable excited with 5 A, 20 kHz current and extrapolated

to the cable rating of 60 A. The general set-up for these tests

is shown in Figure 16.8. In the first test, the magnetic

radiation at 7 cm away (RE01) was measured on a 5 m long cable

with the cable terminals inside the chamber. The EMI in this

setting significantly exceeded the MIL-STD-461B RE01 limit,

which is 27 db above 1 picotesla at 7 cm from the cable. This

was rather unexpected. Some of the possible reasons of not

meeting the RE01 limit are identified as follows:

(1) The cable terminals were inside the shielded chamber.

The cable-run is virtually concentric, thus minimizing

the EMI to the maximum extent possible with the chosen

configuration. The terminals at both ends of the cable,

however, are not concentric. The terminations used in

the EMI tests are shown in Figure 16.6, which make a

small non-concentric loop at each end. These loops

act as flux-generating solenoids, probably generating

enough flux to exceed the RE01 limit. To eliminate this

termination flux, another 10-meter long cable was tested

with the cable ends outside the shielded chamber, and only

the middle part of the cable running inside the chamber.

The results of these tests are given in the next section.

(2) Another possible reason for exceeding the RE01 limit is

the fact that the inductance of the middle strap with

respect to the two outer straps is not perfectly symmetrical due to manufacturing variations. For example, the

inductance measurements reported in Table 16.2 show that

the inductance from middle to one of the outer straps

was 5.5pH and from middle to the other outer strap was

5.2yH. Thus, the first value is about 6% higher than the

other. The currents in the two outer straps are, there16.13
fore, not equally divided, and have an unbalance of

about 6% of the outer strap current value, or 3% of

the cable's rated current, with unbalance of 1.5%

flowing in each of the two outer straps. This current

unbalance was subsequently confirmed by actual current

measurements in the two outer straps. This means that

the cable, although designed to be virtually concentric,

is only 98.5% concentric and 1.5% non-concentric due to

the manufacturing inexactness that resulted in this first

prototype construction. As shown in Figure 16.7, this

1.570 of the cable, then, acts as a simple solenoid,

generating enough flux to exceed the stringent RE01

limit of 27 db above 1 picotesla.

16.4 EMI TESTS WITH CABLE TERMINALS OUTSIDE THE CHAMBER

Since paragraph (1) above indicated that the non-concentric cable terminals can create solenoidal flux, which does not

represent the main part of the cable, it was decided to make

further tests on a 10-meter long cable with the terminals

placed outside the shielded chamber. The EMI was then measured

under the following three different conditions:

(1) Unshielded cable

(2) Electro-magnetic shield around the cable (EMT)

(3) Conductive shield around the cable (copper tube)

The unshielded cable still exceeded the allowable limits.

With end-terminals outside the chamber, this indicates that

the magnetic flux radiated from the cable is mainly due to the

non-concentric part of the cable as explained in para (2) of

Section 16.3

16.14
f K)

-/

/I

91

\

cable run <

1

Dotted line shows the solenoidal

¥

Lux loop at the cable ends.

Terminal leads connected to 20 kHz power source

at one end and shorted at the other end of the cable.

FIGURE 16.6: Non-concentric terminal-ends

generate flux like a solenoid

outwardly

Internally - - emitti

contained >^ F fl

flux y/

1 1 / \ 1 (

_

^_

IS^

r I -J

Equivalent to

—

I

i

1

1

1 • -f*

w

\

1

1

1 TM

I

1

. ' plus +•

I I I

1 1 1

I I I TM

^i

i

i

»8% 100 51% 50 100 50 +152 -1%

Actual = concentric + solenoid

FIGURE 16.7: Difference in inductances of the middle to two outer straps

makes the actual cable equivalent to a concentric cable plus a solenoid.

The latter acts as a flux-emitter.

16.15
O

-J <

< &

z w

U 2

HH UJ

to o

ca

I-H U

X H

U UJ

GROUND PLANE

CABLE

UNDER TESTS

COPPER

STRAPS

T

7 or

100 cm

1

5cm STAND OFFi

SHIELDED ENCLOSURE

SENSOR

FIGURE 16.8: EMI Test Set-up for

MIL-STD-462, Section RE01, RE02, RE04

16.16
Following is an estimate of the EMI due to-the nonconcentricity introduced by the cable manufacturing variations,

which caused unequal inductances of the two outer straps with

respect to the middle strap.

As shown in Figure 11.1, the magnetic flux at 70 mm

away from half of the cable (equivalent to a 30 A solenoid)

is ..042 pT. A 60 A solenoid would then create .082 yT at

7 cm. The 1.5% solenoid, due to the above-mentioned unbalance

current in the two outer straps, would create .015 x 82000 =

1230 picotesla. This converts to 20 log (1230/1) = 62 db

above 1 pT, which clearly exceeds the RE01 limit of 27 db.

The calculation shows that although the 1.570 non-concentric

part of the cable sounds low, it creates sufficient solenoidal

flux to exceed the specified EMI limit at 20 kHz. Unfortunately,

such EMI calculations on a cable with symmetrical outer straps

could not be made without an involved finite-element study.

However, as mentioned in Section 11.1, the EMI flux produced

at some distance away by the two halves of the cable would

cancel each other, leaving virtually zero net flux. For this

reason, the EMI from a cable manufactured with nearly perfect

symmetry between the straps is expected to have significantly

lower EMI. Finite-element electromagnetic flux plot study

is required to investigate this in further details.

Since the unshielded cable did not meet the EMI limits

(see Figure 16.9), the cable was shielded with 1/2-mil thick,

one inch wide high permeability magnetic tape. This reduced

the EMI only slightly, indicating that the amount of radiated

flux probably saturated the thin tape. The cable was then

shielded with a readily available 1/16-inch thick, 2-inch

diameter mild steel conduit (EMt) and also with copper tubing

of the same size. The reduction in EMI was more with copper

tubing than with EMT. The EMT tube is partly magnetic and

partly conductive, while copper is purely conductive. The

conductive shield reduces the EMI by eddy currents induced in

it which opposes the flux radiated by the cable. Theoretical

16.17
considerations indicate that a purely magnetic shield would

generally perform better than EMT and copper shields at this

frequency.

The full EMI test report has been separately submitted

to NASA. However, the charts summarizing the results are

depicted in Figures 16.9 to 16.19. The following notes apply

to these curves:

1) The measurements were taken after scanning the EMI at

different locations and in different antenna orientations

The reported EMI were at the point of maximum leakage,

which was generally at the edges of the cable. •

2) While measuring the EMI, the shielded chamber had sufficient ambient noise particularly at lower frequencies.

For example, Figure 16.10 is ambient scan from 1 kHz to

50 kHz for the RE04 (and also RE01) setup. At 60 Hz,

the ambient was 85 db, and at 400 Hz it was 40 db.

Since we were primarily interested at 20 kHz and its

higher harmonics, no significant efforts were made to

locate the source of the low frequency ambient.

3) The second harmonic in the 20 kHz power source used for

these tests introduced EMI at 40 kHz. For example, if

there is 37» second harmonc present in the power source

(a normal harmonic content in such equipment), it will

introduce 40 kHz EMI which would be 20 log (I/.03) =

30.46 db lower than the fundamental 20 kHz EMI. This

is about what we see in Figure 16.9.

4) The level of electric field radiation measured in the

RE02 setup (Figures 16.14 to 16.16) was unexpected,

nor can we explain at this time. However, significant

ambient present in the shielded chamber makes this data

doubtful.

16.18
EATON DATA ACQUISITION SYSTEM

E.U.T. : INDUCTION GENERAL/NASA CABLE

TEST PROCEDURE :RE01

COMMENTS :

TEST ENGINEER :B. SMITH

DATE 07 Fi SCAN PARAMETERS TEST TITLE :RE01 NASA CABLE

I Rvr I Start Freq I Stop Freq I Bandwidth I Atten i Detector I Transdu1 7 I 1 0 kHz I 50 kHz I 1 kHz I 0 dB I

GLOBAL PARAMETERS

PEAK: i ALP-10

BROAD BAND CORRECTION: NO

LIMIT LINE : REO"<

SWEEP RATE: MAX

SCANS PER SEGMENT

SEG 1 START TIME: 13:19:21 STOP TIME: 13:19:49

Maximum Scan Data : RE01 NASA CABLE

180-

10.88 kHz FREQUENCY

FIGURE 16.9: EMI RE01 Pre-Scan on Unshielded Cable, showing significant

radiation at fundamental (20 kHz) and its 2nd harmonic. The hash at

lower frequency is ambient noise.

16.19
CATON DATA ACQUISITION SYSTEM

E.U.T. : INDUCTION GENERAL/NASA CABLE

TEST PROCEDURE :RE04 PRESCAN

COMMENTS .-AMBIENT

TEST ENGINEER :B. SMITH

DATE 08 Fh

SCAN PARAMETER:: TEST TITLE- :REQ4 PRESCAN

I Rvr I Start Freq I Stop Freq I Bandwidth I Atten I Detector I Transduc

17 I 1 kHz I 50 kHz I 100 Hz I 0 dB I PEAK I ALP-10

I GLOBAL PARAMETERS

BROAD BAND CORRECTION: MO

LIMIT LINE : RE04

I SWEEP RATE: 1000 Samples/Sec

I SCANS PER SEGMENT : 1

SEG 1 START TIME: 08:46:55 STOP TIME: 08:47:46

Maximum Scan Data : RE04 PRESCAN

JQQ-,

38kHz

"RE.QLCWCY

FIGURE 16.10: EMI RE04 (and RE01) Ambient Scan,

showing higher ambient at lower frequencies.

< H::

16.20
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16.26
ORIGINAL PAG! IS

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 - ";'":———

Since the magnetic radiations, RE01 at 7 cm and RE04

at 1 meter, at 20 kHz are of our primary interest at present,

they are summarized in Table 16.8. The RE04 data in that

table is questionable, as the EMI of EMT shielded cable at

1 meter (RE04) shows higher db than that at 7 cm (RE01) distance. This conflicts with the engineering judgment. At any

rate, the table indicates that the unshielded cable does not

meet the magnetic radiation limits of MIL-STD-462, Section

RE01. The reasons for this, as identified earlier, are as

follows:

1) The pressure buildup in the extrusion die discussed in

Section 15.0 made the middle strap wider than the outer

straps. This caused the end-fringing flux to extend

farther from the cable. If the middle strap was shorter

than the outer straps (as was designed), the end flux

would have been contained close to the cable, as explained in Section 4.2.

2) The manufacturing variations caused inductances of the

two outer straps, with respect to the middle strap, to

be unequal. The resulting unequal currents in the outer

straps caused a small part of the cable to act as a

flux-emitting solenoid, as explained in para (2) of

Section 16.3.

Both of the above sources of the EMI are related with

the manufacturing difficulties experienced in the first prototype construction. Future manufacturing improvements, based

on the experience gained, can eliminate both of the above EMI

sources. For example, Figure 16.20 shows how the end-fringing

flux can be minimized by rolling the outer straps over the

middle strap. The cable, thus made with symmetrical outer

straps, is expected to meet the EMI limits, possibly without a

shield or with a shield of thin, high permeability magnetic tape

(one mil thick of negligible weight) wrapped around it.

16.30
TABLE 16.8

SUMMARY OF EMI MEASUREMENTS ON THE CABLE

PER MIL-STD-462 AND 461B

TEST CONDITION

EMI db at 20 kHz

over 1 picotesla^

RE01

at 7 cm

(limit 27 db)

RE04

at 1 m

(limit 58 db)

Unshielded Cable 73 db 68

Cable shielded with EMT,

electromagnetic (mild

steel) tubing

30 43

(2)

Cable shield with

conductive (copper)

tubing

20 Less.than

meter

sensitivity

(1) These tests were made at 5 A cable current. To

extrapolate them to the rated current of 60 A, add

20 log (60/5) = 22 db in these readings.

(2) This, being at a 1-meter distance, does not make

sense, as it is higher than 30 db measured at 7 cm.

(3) See Figure 16.20 for recommended manufacturing

improvements to reduce EMI to the specified limits

16.31
38

(1.50)

35

(1.38;

-% 7.1 mmK"

(0.28 inches)

Outer straps, 'rolled over the middle

strap to reduce EMI

Middle strap

Major insulation jacket

(polypropylene)

Outer insulation jacket

Teflon tape, 1 mil half-lap, to prevent

the outer insulation penetrating

between the straps.

FIGURE 16.20: RECOMMENDED MANUFACTURING IMPROVEMENTS.

Outer straps rolled over the middle strap will close the main leakage point

at the cable edges, thus further minimizing the EMI. Such a cable made

with improved symmetry in the outer straps is expected to meet the

specified EMI limit.

16.32
APPENDIX A

EDDY LOSSES IN LITZ STRAPS LINES DUE TO TRANSVERSE FLUX

Let us consider a round wire situated in an alternating

magnetic flux, the flux density being distributed linearly as

shown in Figure A.I. We shall neglect the demagnetizing effect

of the eddy currents on the magnetizing flux. This is possible

when the wire diameter is much smaller than the skin depth, as

is true in the present case.

Consider a longitudinal cross section of the wire as

shown in Part (C) of Figure A.I. An alternating e.m.f. will

be induced in the conductor and short circuited eddy currents

will flow as shown. To quantify the phenomenon, let us consider an elemental width dx at distance x from the wire axis.

Let J = current density at distance x Ji

B = flux density at distance x J\

0 = flux linking with the loop at x

X

1 = length of the current path (loop)

t = time

d =- wire diameter or dimension perpendicular to the

flux in rectangular conductor

p = resistivity of the conductor material

As eddy current path is short circuited, the difference in

e.m.f.s is equal to the difference in voltage drops in the

short-circuited paths at x and at x+dx, i.e.

3J \

-§s

 dxy

 d3BX

X Idx = I J__ + -^ dx; d-dx -^rr - Jxd-dx

9Bx „ 3Jx .

-at

 dx = p

 —^

 dx

A.01
T

JL

B.

I

(a) Space varying

magnetic field

(b) Elementary volume

under analysis

1.1 1

«-dx

(c) Eddy current loop

FIGURE A.I: Eddy Currents in Round Conductor

placed in transverse magnetic field

A.02
We neglected the leakage reactance drop because it is small in

comparison with the resistance drop. For sinusoidal time

variations, substituting 3/3t = u , the above equation becomes:

x

3x

From Figure A.I (a) ,

B +B B -B

Bx = ^-±- + —I—i- x = A + Bx, say

where A =

Thus,

 3 J

2

B

2

+B1 and B =

= A —— +. Bx, whose solution is

p

-x = -a- (AX + B -fi)-I- C

The integration constant C can be found from the condition that

the sum of all eddy currents in the conductor cross section is

equal to zero, because the eddy currents are short circuited

ones. Another way of looking at this is that the eddy currents

at +x and -x are equal and opposite. Therefore,

+do/,

-do/.

J dx = 0, from which we get

24p

-'-

 Jx = -f- I AX + T-'

A.03
If we use rms values of the flux densities A and B, then J will

X

be rms. The eddy current losses per unit volume at x-is

J

 2p , and per elementary volume of one meter length, width dx

f±

and height d is J 2pd-dx. The total eddy losses per meter

X

length are then given by integrating this, i.e.

/do/

P eddy = J

-do

2~

d-p J dx (A.I)

We now define the transverse eddy loss factor

„ _ DC Losses + Eddy Losses , so that

t DC Losses

Rac = KtR

dc

The dimension d in round wire varies with x. The eddy

losses in the entire round wire is then found by double integration of the above equation. If there are N strands arranged

as shown in Figure A.2, and the flux density varies from B

at the inner edge to zero at the outer edge of the strap, then

the total eddy losses can be found by the third integration over

all the strands. At the end of these three integrating processes,

we found the f actor. K,.

0.8f 2N

2

d

 6

K = 1 + for round wires,

f c

 P

2

b

2

E12

(A.2)

where d is the dimension perpendicular to the flux direction.

A. 04
(a) Magnetic field linearly

varying from zero to

maximum across the strap

thickness dimension

-ir

oooo

oooo

oooo

oooo

OO QO

ooo o

oooo

oooo

I 1

(b) Strap made of numerous

fully transposed conductor

strands of diameter d~

FIGURE A.2: Rectangular Strap of Round Strands

placed in transverse magnetic field varying from Bo

at one edge and zero at the other edge

A.05
APPENDIX B

PARAMETRIC STUDY ON 2-STRAP CABLE CONFIGURATION

This Appendix contains the following:

(1) Computer Program Listing (BASIC)

(2) List of Computer Symbols (SI Units)

(3) Computer Output Sample (Copper and PVC)

(4) Figures B.I to B.5, depicting Resistance,

Inductance, Capacitance and Mass of 2-strap

cable with copper and aluminum conductors.

Computations were made for major insulation

thickness d ranging from 1 to 3 mm, strap

thickness a from 1 to 4 mm, and strap width

from 50 to 100 mm.

B.01
LIST OF COMPUTER SYMBOLS IN SI UNITS

Dl=rifiSS DENSITY OF CONDUCTOR

D2=nflSS DENSITY OF INSULRTIOH

KF=FILL<SPflCE> FRCTQR OF _STRRP

D0=STRflND PI RMETER

R6=RESISTIVITV OF CONDUCTOR

F=FREOUENCV

E=RELHT1VE PERMIT 1V1TV _OF _lNSjyLLRT_I ON

Ee=RBSOLUT£ PERMIT I VI TV OF FREE_^PRC£

n&=RBSQLUTE PERnEflBlLITV OF FREE SPftCE

J =CRSE < RUT-i .) NUMBER COUNTER

PI =3. 1416

'1>=INSULRTIOH THICKNESS BET»EEN_STRf!PS

H=CQNDUCTOR _SJRf^'..l'^. 1 CKNESS

B=CQNDUCTUR STRBP MIDTHCHEIGHT1

N=NUP1BER UF STRftHPS IN Org STRBP

KS=SKIN EFFECT EPPV LOSS FRCTOR

KC=CRBL- 1 NO <L ST Rftp I NG > _l-JNQTH_FflCJpK

KT=TRRNSVERSE FLL>X EDDY LOSS FRCTOR

RRgRC RESISTRNCE> MILL I OHMS

L= 1 NDUC't HNCE.. M I LL I HENR I ES

C=CfiPRC 1 1 ffrtCE.- M I CROF RRRDS

MC=MEIGHT OF CONDUCTOR

OF IN5ULRTION

HT=TC'TRL HEIGHT WITHOUT SHIELD

VVMn=:> I ELECTRIC STRESS. VOLTS/MM

vvn i L=D i ELECTRIC STRESS^ VOLTS/MIL

L = EXPONENT

B.02
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B.03
COMPUTER OUTPUT SAMPLE (COPPER AND PVC)

Hl-FREQ TK. LINE DESIGN

INPUT DriTR

Dl, D2.. KF= 8648 1348 . 4

D0, R£i, F= 2. 5E-04 2. 2E-08 2008©

c... £0= 2. 7 8. 85E-12

Pi.. ri£i= 3. 1416 1. 25664E-66

CHSE NO. 1

b, fi.. D= . 85 1E-03

N, KT= 407. 435 1. 01872

v/'nn. v/niL= 1000

LINE FRRRPIETERS

RR.. L^ C= 127. 856 2. 83482 . 8597375

WC, WI/ WT= 1 9 . 87 18. 05 29 . 92

CflSE NO. 2

B; ft, D= . 06 lE-^03 1E-83

N, KT= 4£«. 322 1. 01872

WMM,VVniL= 1888 25

LINE PHRRf-iETERS

L^ C= 106. 547 1. 7456S . 671685

WC, Wl, lift- P.3.85 12.86 35.91

B.04
10U

.02

FIGURE B.I: Resistance, Inductance, Capacitance and Mass of two

Parallel Straps with Copper Conductor and 1 mm insulation thickness

ORIGINAL PAGE IS

B<Q5 OE POOR QUALHV
'_!" ' -r-^j I • ••-.-•• • • .

-i •"-- ::1:::-:Lll^-ixin. _:

FIGURE B.2: Resistance, Inductance, Capacitance and Mass of two

Parallel Straps with Aluminum Conductor and 1 mm insulation thickness

ORIGINAL PAGE IS

B.06
bmm • ;-

66: : 70 80 90 100

40

FIGURE B.3: Mass of Parallel Straps with Copper Conductor

and different insulation thicknesses ~.,~,

ORIGINAL PAGE IS

DE POOR QUALITY

B.07
:;; ed ::

;:;7

 : :::BD ;;;::9::

: ID

DO

FIGURE B.4: Mass of Parallel Straps with Aluminum Conductor

and different insulation thicknesses _.

ORIGINAL PAGE fi

OF POOR -- -

B.08
ORIGINAL PAGE IS

OF POOR QUALITY

OJ

•H

)-(

C

QJ

.C

O

Vi

u

I

0)

u

C

(0

-p

O

3

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C

w

0)

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S-l

C

0)

O

i-l

u

-HVd 4

I

0)

O

C

(0

-p

0

3 _;.

FIGURE B.5: Inductance of Parallel Straps with Copper and

Aluminum Conductors with different insulation thicknesses

B.09
LIST OF FIGURES

Figure 2.1:

Figure 2.2:

Figure 2.3:

Figure 2.4:

Figure 2.5:

Figure 2.6:

Figure 2.7:

Figure 4.1:

Figure 4.2;

Figure 4.3;

Figure 4.4;

Figure 4.5;

Figure 5.1;

Figure 5.2:

PAGE

Hollow configurations investigated by 2.02

previous workers

Two parallel Litz strap lines 2.04

Ratio of AC to DC resistance for a 2.08

round wire in self magnetic field

Characteristic of proximity effect on 2.09

the peripheral distribution of current

density

Reactance of rectangular flat conductors 2.13

at 60 Hz

The strap lines as a capacitor 2.14

Force between rectangular flat conductors 2.16

for different geometrical spacings

Three parallel straps (virtually concen- 4.02

trie) configuration, evolution

Finite-element flux plot of 3-phase bus 4.04

bars with central bus carrying peak

current

Indentation of the middle strap improves 4.05

EMI and end-insulation design

Three-straps 2.0 inch wide design for 4.08

1000 V, 100 A cable with aspect ratioof 5.1

Three-straps 1.5 inch wide design for 4.10

600 V, 60 A cable with aspect ratio

of 5.4

Dielectric loss tangent vs. frequency at 5.09

room temp, for some low-loss insulating

materials

Dielectric loss tangent vs. temperature 5.10

for some insulating materials

L.01
Figure 6.1:

Figure 6.2:

Figure 9.1:

Figure 9.2:

Figure 9.3:

Figure 11.1:

Figure 12.1:

Figure 12.2:

Figure 14.1:

Figure 14.2;

Figure 14.3:

Figure 14.4:

Figure 15.1:

Figure 16.1

Figure 16.2

Figure 16.3:

LIST OF FIGURES

PAGE

Paschen curve, electrical breakdown 6.02

stress versus pd for air

Corona inception voltage vs. laminar 6.04

void size

Radiation model under study with one 9.01

face radiating to free space

Conduction cooling configuration under 9.06

analysis

Equivalent thermal circuit used for 9.07

temperature rise calculations with

conduction alone

EMI flux density estimate of one-half of 11.02

the cable, with the other half cancelling

Three-strap 1.5 inch wide design for 12.02

600 V, 60 A cable

Dimension symbols of the 3-straps cable 12.04

as used in summarizing the design

formulas

Equipment set-up and major process steps 14.02

for constructing one-meter long sample

One-meter sample and the 100 kW power 14.03

supply

Geometrical dimensions and test results 14.04

of 3.7 m sample

Geometrical dimensions and test results 14.06

of 5 m sample

Manufacturing problems encountered 15.03

Corona test set-up 16.02

R,L,C test set-up for 20 kHz open-circuit 16.06

and short-circuit tests on full cable

Reactive voltage vs. current during short-16.09

circuit tests on full cable

L.02
LIST OF FIGURES

PAGE

Figure 16.4: Loss vs. current during short-circuit 16.10

tests on full cable

Figure 16.5: Charging current vs. voltage during 16.11

open-circuit tests on full cable

Figure 16.6: Non-concentric terminal-ends generate 16.15

flux like a solenoid

Figure 16.7: Difference in inductances of the 16.15

middle to two outer straps makes the

cable act as a flux-emitter

Figure 16.8: EMI test set-up for MIL-STD 462, 16.16

Sections RE01, RE02 and RE04

Figure 16.9: EMI RE01 pre-scan on unshielded cable 16.19

Figure 16.10: EMI RE04 (and RE01) ambient scan 16.20

Figure 16.11: EMI RE01 test results on unshielded 16.21

cable

Figure 16.12: EMI RE01 test results on cable 16.22

shielded with 2-inch EMT

Figure 16.13: EMI RE01 test results on cable 16.23

shielded with 2-inch copper tube

Figure 16.14: EMI RE02 narrow band test results on 16.24

unshielded cable

Figure 16.15: EMI RE02 narrow band test results on 16.25

cable shielded with 2-inch EMT

Figure 16.16: EMI RE02 narrow band test results on 16.26

cable shielded with 2-inch coppef tube

Figure 16.17: EMI RE04 test results on unshielded 16.27

cable

Figure 16.18: EMI RE04 test results on cable 16.28

shielded with 2-inch EMT

Figure 16.19: EMI RE04 test results on cable 16.29

shielded with 2-inch copper tube

Figure 16.20: Recommended manufacturing improvements 16.32

L.03
LIST OF FIGURES

PAGE

Figure A.I: Eddy currents in round conductor A.02

placed in transverse magnetic field

Figure A.2: Rectangular strap of round strands A.05

placed in space varying transverse

magnetic field

Figure B.I: Resistance, inductance, capacitance B.05

and mass of parallel straps with

copper conductor and 1 mm insulation

thickness

Figure B.2: Resistance, inductance, capacitance B.06

and mass of parallel straps with

aluminum conductor and 1 mm insulation

thickness

Figure B.3: Mass of parallel straps with copper B.07

conductor and different insulation

thicknesses

Figure B.4: Mass of parallel straps with aluminum B.08

conductor and different insulation

thicknesses

Figure B.5: Inductance of parallel straps with B.09

copper and aluminum conductors with

different insulation thicknesses

L.04
Table 1.1;

Table 3.1

Table 4.1

Table 5.1

Table 5.2

Table 5.3

Table 5.4;

Table 5.5;

Table 8.1:

Table 9.1:

Table 9.2

Table 12.1

Table 13.1

Table 16.1:

Table 16.2:

Table 16.3

LIST OF TABLES

PAGE

Initial specifications and goals as set 1.02

by NASA

Comparison of various configurations for 3.02

100 kW, '1000 V, 100 A, 50 m long transmission line

Comparison between 2-strap and 3-strap 4.06

1000 V, 100 A designs

Alternative conductor materials 5.02

Electrical and mechanical properties of 5.03

common conductors

Properties of major electrical insulating 5.07

materials

Candidate insulating materials compared 5.08

from manufacturing viewpoints

Properties of the selected major insulating 5.12

material, Pro-Fax SE-191 propylene copolymer of Hercules, Inc.

Overall comparison between alternatives 8.02

Thermal properties of materials proposed 9.04

in the cable construction

Operating temperatures under conduction and 9.10

radiation for 600 V, 60 A, 1.5 inch wide

cable design

Design summary of 600 V, 60 A, 1.5 inch 12.03

wide cable selected as the final design

Failure modes and effects analysis on 13.01

the cable

Corona test results on flat insulation 16.03

samples

Bridge measurements on full cable lengths 16.05

. Short-circuit tests on 65 m cable 16.07

L.05
Table 16.4:

Table 16.5;

Table 16.6:

Table 16.7

Table 16.8:

LIST OF TABLES

Open-circuit tests on 65 m cable

Short-circuit tests on 123 m cable

Open-circuit tests on 123 m cable

Comparison of calculated and tested

parameters

Summary of EMI Measurements on the

cable per MIL-STD-462 and 461B

PAGE

16.07

16.08

16.08

16.12

16.31

L.06
LIST OF REFERENCES

[ 1] Jefferies, K. S. and Renz, D. D., Parametric Analysis

of Hollow Conductor Parallel and Coaxial Transmission

Lines for High Frequency Space Power Distribution, NASA

Technical Memorandum 83601, Lewis Research Center,

Cleveland, Ohio, March, 1984.

[ 2] Renz, D. D., et. al., Design Considerations for Large

Space Electric Power Systems, NASA Technical Memorandum

83064, Lewis Research Center, Cleveland, Ohio, April, 1983.

[ 3] Grover, F. W., Calculations of Mutual and Self-Inductances,

Dover Publication, New York, 1955.

[ 4] Grivet, P., Physics of Transmission Lines at High and

Very High Frequencies, Vol. I, Academic Press, New York,

1970.

[ 5] Dwight, H. B., Electrical Coils and Conductors, McGrawHill Book Co., New York, 1945.

[ 6] McNutt, W. J. and Patel, M. R., The Combined Effects of

Thermal Aging and Short Circuit Stresses on Transformer

Life, IEEE Transactions, Vol. PAS-95, pp. 1275-1286, 1976.

[ 7] Hansen, I., Description of a 2.3 kW Power Transformer for

Space Application, NASA Technical Memorandum 79138, Lewis

Research Center, Cleveland, Ohio, February, 1979.

[ 8] Dunbar, W. G., High Voltage Design Guide for Spacecrafts,

Air Force Wright Aeronautical Laboratory/Boeing Aerospace

Company Report No. AFWAL-TR-82-2057, Vol. V, January, 1983.

[ 9] Machine Design, Penton/IPC Publication, Cleveland, OH,

April 1985 issue.

[10] Bradwell A., Electrical Insulation, Peter Peregrinus

Ltd. and IEE London (Publishers), 1983.

L.07
LIST OF REFERENCES

[11] Meek and Craggs, Electrical Breakdown of Gases, John

Wiley & Sons, New York, 1973, p. 689.

[12] Jacob M., Heat Transfer, Vol. I and II, John Wiley &

Sons (Publishers), NY, 1949.

L.08
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NASA CR-175071

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L.10