NISTIR 90-3935

AlllOfc, 3“isasi

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Final Report

January 1988-December 1989

L.F. Goodrich, R.B. Goldfarb, S.L. Bray

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Institute of Standards and Technology

Boulder, Colorado 80303-3328

JANUARY 1990

QC

100

.U56

NO. 90-3935

1990

NISTIR 90-3935

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Final Report

January 1988-December 1989

L.F. Goodrich, R.B. Goldfarb, S.L. Bray

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Institute of Standards and Technology

Boulder, Colorado 80303-3328

JANUARY 1990

Prepared for

Office of Fusion Energy and Division of High Energy Physics

Department of Energy

Washington, DC 20545

^

united states

DEPARTMENT OF COMMERCE, Robert A. Mosbacher, Secretary

I NATIONAL INSTITUTE OF STANDARDS ^PVHIVl AND TECHNOLOGY, Raymond G. Kammer, Acting Director

FOREWORD

The National Institute of Standards and Technology (formerly the National

Bureau of Standards) program to develop standards for conventional superconductors has been one of our most successful and visible projects in support of

DoE superconductor applications. Appendix A is a list of publications from

this research program. We have been successful in standardizing measurement

techniques in practical superconductivity. This has resulted in savings of

time and money as well as better communication among researchers, producers,

and users of superconductors. The work described here is a further effort in

this area. We are shifting our emphasis to large conductor standards and

topics such as ac losses which have become of increasing concern with the

development of fine filament conductors. Small conductor work continues as

well with emphasis on round robin methods to measure the degree to which our

suggested techniques have increased measurement uniformity and accuracy.

in

CONTENTS

Page

Foreword m

ABSTRACT vii

I. CRITICAL CURRENT 1 0

II. AC LOSSES 10

III. ADDITIONAL ACTIVITIES 17

IV. ACKNOWLEDGEMENTS 23

V. REFERENCES 23

Appendix A - PUBLICATIONS OF RESEARCH PROGRAM 25

Appendix B - INTEGRITY TESTS FOR HIGH-TC AND CONVENTIONAL CRITICALCURRENT MEASUREMENTS SYSTEMS

L. F. Goodrich and S. L. Bray 31

Appendix C - BATTERY-POWERED CURRENT SUPPLY FOR SUPERCONDUCTOR

MEASUREMENTS

S. L. Bray, L. F. Goodrich, and W. P. Dube ... 41

Appendix D - NB 3 SN CRITICAL- CURRENT MEASUREMENTS USING TUBULAR

FIBERGLASS -EPOXY MANDRELS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer . 47

Appendix E - CRITICAL CURRENT MEASUREMENTS OF NB 3 SN SUPERCONDUCTORS:

NBS CONTRIBUTION TO THE VAMAS INTERLABORATORY COMPARISON

L. F. Goodrich and S. L. Bray 53

Appendix F - THERMAL CONTRACTION OF FIBERGLASS - EPOXY SAMPLE MANDRELS

AND ITS EFFECT ON CRITICAL- CURRENT MEASUREMENTS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer . 65

v
Appendix G

Appendix H

Appendix I

Appendix J

Appendix K

Appendix L

Appendix M

Appendix N

THERMAL CONTRACTION OF FIBERGLASS - EPOXY SAMPLE

HOLDERS USED FOR NB 3 SN CRITICAL- CURRENT MEASUREMENTS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer . 69

VAMAS INTERCOMPARISON OF CRITICAL CURRENT MEASUREMENT IN

NB 3 SN WIRES

K. Tachikawa, K. Itoh, H. Wada, D. Gould, H. Jones

C. R. Walters, L. F. Goodrich, J. W. Ekin, and

S. L. Bray 79

PROPOSED STUDY ON THE EFFECT OF SAMPLE BONDING TECHNIQUES

ON THE MEASURED CRITICAL CURRENT OF NB 3 SN SUPERCONDUCTORS

L. F. Goodrich and S. L. Bray 87

CURRENT RIPPLE EFFECT ON SUPERCONDUCTIVE D.C. CRITICAL

CURRENT MEASUREMENTS

L. F. Goodrich and S. L. Bray 99

CALORIMETRIC MEASUREMENTS OF AC LOSS IN NB 3 SN

SUPERCONDUCTORS

M. Takayasu, C. Y. Gung, M. M. Steeves, B. Oliver,

D. Reisner, and M. 0. Hoenig 107

GUIDELINE FOR NB 3 SN CRITICAL CURRENT MEASUREMENTS USING

FIBERGLASS -EPOXY COMPOSITE SAMPLE MANDRELS (SUMMARY)

L. F. Goodrich 115

REPORT ON SAMPLE MOUNTING

L. F. Goodrich 123

GUIDELINE FOR NB 3 SN CRITICAL CURRENT MEASUREMENTS USING

FIBERGLASS -EPOXY COMPOSITE SAMPLE MANDRELS

L. F. Goodrich 129

vi
DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Final Report

January 1988 -December 1989

L. F. Goodrich, R. B. Goldfarb, and S. L. Bray

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Institute of Standards and Technology

Boulder, Colorado 80303

A cooperative program with the Department of Energy, the National Institute of

Standards and Technology, other national laboratories, and private industry is

in progress to develop standard measurement practices for use in large-scale

applications of superconductivity. This report describes research for the

period January 1988 through December 1989. It contains the results of

critical current studies on a large conductor Reference Material (RM) , the

effect of power-supply current ripple, the measurements of large conductors,

and an interlaboratory comparison (round robin) of Nb 3 Sn wires. Short-range

variations in magnetic hysteresis loss in multifilamentary Nb 3 Sn were studied.

The results of participation in several interlaboratory comparisons are

described.

Key words: ac losses; circuit diagrams; critical current; current ripple;

hysteresis; magnetization; niobium-tin; niobium- titanium; power supply; round

robin; standards; superconductors; susceptibility.

vii

I . CRITICAL CURRENT

A. Large Conductor Reference Material

Critical current measurement requires complex techniques and apparatus.

The quality of the measurement is strongly dependent on the integrity of the

overall system. There are numerous measurement variables that, if not

properly addressed, can cause erroneous results. The NbTi Superconducting

Wire Standard Reference Material (SRM 1457) [1] was developed to provide a

holistic method for testing I c measurement systems. The SRM has been

productively utilized from both a scientific and commercial standpoint;

however, its usefulness is limited by its maximum current capacity. At a

magnetic field of 6 T, the I c of the SRM is only 125 A, but it has been used

to test 25,000 A I c measurement systems [2]. This situation suggested the

need for a well characterized superconductor having a significantly higher I c

.

It was decided that a RM (reference material) rather than an SRM was the

best solution to the need for higher current capacity in a well-characterized

conductor. The characterization of a RM is not as stringent as a SRM and,

thus, its properties are not certified at the level of a SRM. There are two

basic reasons, one practical and one technical, why this is appropriate for

the large conductor. Technically, there are measurement problems unique to

large conductors that, at present, limit our confidence in measurement

accuracy. A matter of practicality is the limited distribution anticipated

for the large conductor RM as compared to the SRM. This limited distribution

diminishes the benefit of the additional time and difficulty required to

develop a SRM.

A preliminary screening of candidate conductors for the RM resulted in

the selection of a NbTi rectangular monolith. A U.S. wire manufacturer has

agreed to supply NIST with the required length of this conductor (""500 m) at

no charge. The cross-sectional dimensions of the conductor are 1.8 mm x 3.6

mm and its I c at 6 T is approximately 2700 A, or about 22 times that of the

SRM. A 12 T, 12.7 cm-diameter (5") bore magnet, that is well suited to large

conductor I c measurements, has been acquired, tested, and calibrated. A

sample test fixture that incorporates 3000 A vapor cooled current leads and a

magnet- to- test- fixture coupling that is intended to support large Lorentz

forces has been designed, constructed, and tested. The test fixture will

accommodate multiturn coiled samples having diameters up to 11.4 cm.

Calibrations have been made on current, voltage, magnetic field, and

temperature

The wire was received as a continuous length from which 130 sample units

were made. A systematic sampling pattern was used to select 30 units from

these 130. The remaining 100 units will be available for sale. The 30 sample

units are used to characterize all units. It was decided that about 10

critical current specimens would be chosen from these 30 units. Each sample

unit, or spool, has approximately 3 meters of wire, which is more than enough

length for an I c specimen. The extra length on these units and the remainder

of the 30 units are to be used for photographs, other measurements, and

contingencies. The filament twist pitch, copper- to-superconductor ratio,

1
average filament diameter, and superconductor area are examples of other

measurements now being made. An anticipated contingency is the possibility of

testing a suspect measurement with an adjacent specimen.

The critical current or the electric field (or resistivity) versus

current is being characterized over a range of variables. The approximate

range of electric fields is from 0.02 to 0.5 /iV/cm. The approximate range of

resistivity is from 10" 12 to 10" 11 ft- cm (except at the higher magnetic

fields). The range of applied magnetic fields is 5 to 9 T. The sample

temperature ranges from 4.02 to 4.22 K. The highest current is about 3850 A

and the lowest is about 670 A. Measurements are made with both current

directions with the applied magnetic field parallel to the wide face of the

conductor

At present data acquisition on the test samples is nearly completed and

statistical modeling is being developed. The target date for the RM to be

available is in 1990. Several additional related publications are planned. A

paper on a simple test (integrity test) of the critical-current measurement

system is included as Appendix B to this report. A final paper on our high

current supply is included as Appendix C.

B. Self-Field Effect

A variable that can significantly affect short sample I c measurements in

high current conductors is the magnetic field introduced by the current

carried in the test sample and its current leads. This "self field" is

superimposed on the applied background field and can either increase or

decrease the magnitude of the total magnetic field, depending on its

orientation with respect to the background field. The effect of this variable

on the measured I c depends on the magnitude and the spatial geometry of the

sample current including the current return path. For aspected conductors

(those that have unequal cross-sectional dimensions, a rectangular conductor

for instance) , the I c depends on the orientation of the magnetic field in the

plane of its cross-sectional axes. Consequently, the measured I c will fall in

a range of possible values depending on the field orientation. The purpose of

this study is to measure that range and to determine techniques that can

reduce the measurement uncertainty introduced by this effect.

The NbTi monolith that was selected for the large conductor reference

material was also used in this study. The knowledge obtained in this study

regarding its behavior with respect to the self-field effect should be

beneficial in characterizing it for use as a RM. A sample holder was designed

and constructed that allowed measurement of the self-field effect for several

different sample current geometries. The initial self- field measurements have

been completed and the resulting data have been analyzed. This has proven to

be a formidable task because of the complexity of accurately determining the

nature of the magnetic field perturbations caused by the self- field of the

various sample and current- lead geometries. The agreement between the

measured values and the calculated values was very good; for example, a 9.1%

change in the measured I c was predicted to be 10.6%. A publication of these

2
results and the results of other self- field experiments is planned in the next

year

There is also concern about variation in magnetization among wires that

are processed in different ways. This problem presents another possible

application for our self-field calculations; to determine the actual shape of

I C (H) from 0 to 8 T without the self- field effect. These data are relevant to

the magnetization of a conductor. At this time it is not known how changes in

I c in the region from 5 to 8 T will reflect changes in I c at low magnetic

fields. Thus, variations in I C (H) need to be understood and controlled in

order to control variation in magnetization. We have demonstrated our

capability of measuring I C (H) from 0 to 8 T on an Superconducting Super

Collider prototype conductor (see Section I-E). The I c was about 1700 A in

zero applied magnetic field. A systematic study of the correlation between

variation at low fields and variation in the region of 5 to 8 T may allow

estimates of low field variation using the variation measured in the region of

5 to 8 T, which can be determined during more routine measurements.

C . Nb3Sn Round Robins

A significant research activity has been the continuing participation in

the VAMAS (Versailles Agreement on Advanced Materials and Standards) and the

US-Japan interlaboratory comparative measurements, or round robins, on the

critical current (I c ) of Nb 3 Sn. Twenty- four laboratories from Europe, Japan,

and the United States participated in the round robin. The purpose of this

international collaborative effort is to determine the sources of variation in

the measured I c of Nb 3 Sn superconductors and, ultimately, to arrive at an

accepted measurement technique that is general enough to be practical and yet

specific enough to yield consistent and accurate results at different

laboratories. In response to discrepancies in the results of these

comparative measurements, the NIST study was expanded to address specific

measurement variables (thickness of G-10 sample holder and thermal

contraction) that were thought to be partially responsible for these

discrepancies. Four papers relating to the expanded NIST study are included

as Appendices D, E, F, and G to this report. The first paper on the joint

results of the VAMAS round robin is included as Appendix H. The second joint

results paper was published in the last year [3]. The NIST work has

identified and suggested a solution to a problem that was a significant source

of the measurement discrepancies.

Preliminary comparisons of the data showed variations about the average

measured I c of +50% at a magnetic field of 12 T. The largest variations in

the measured I c were thought to be the result of sample damage incurred during

handling. However, even with this outlying data omitted, the variation for

some of the samples is still +25%. The NIST systematic study of the effect of

measurement holder (tubular sample holder made from G-10 fiberglass-epoxy

composite) geometry revealed that a seemingly small change in geometry can

result in a 40% change in the measured I c at a magnetic field of 12 T.

Specifically, the radial thermal contraction of the measurement mandrel

depends on its wall thickness and, thus, so do the conductor prestrain

3
(at 4 K) and, ultimately, the measured critical current. This study revealed

that axial sample strain was a principal variable in the measurement.

Three different conductors were used in these measurements; one was

manufactured in Europe, one in Japan, and one in the United States. These

conductors varied in overall diameter, filament diameter, manufacturing

process (internal tin or bronze process), and copper to superconductor ratio.

Owing to these variations, some samples were more sensitive to certain

measurement variables than others. Although such sensitivities probably

resulted in a greater overall data scatter, they were beneficial in the sense

that they emphasized the importance of the specific measurement variables.

The large discrepancies in the measurements suggest that there is more to

be learned about short-sample Nb 3 Sn I c measurements. Several topics that

require further investigation are sample holder materials, reaction variables,

bonding techniques, and sample homogeneity. In the category of sample holder

materials, two measurement variables are of particular concern: One is the

reaction holder and the other is the measurement holder; both can cause large

variations in the measured I c

. In the case of the reaction holder, usually an

oxidized stainless steel tube, the sample can be damaged when it is removed

from the holder following the reaction process, if diffusion bonding between

the sample and holder has occurred. This problem might be alleviated by

either an improved oxidization process for the stainless steel holders or by

the use of a different reaction holder material that is less susceptible to

bonding. In the case of the measurement holder, differential thermal

contraction between the sample and holder can result in various states of

sample strain depending on the sample holder material and its geometry and the

bonding method. To address this problem, measurement holder materials and

geometries that do not result in adverse mechanical interaction between the

sample and holder need to be studied further.

The existing American Society for Testing and Materials (ASTM) "Standard

Test Method for D-C Critical Current of Composite Superconductors" (B714-82)

addresses the problem of differential thermal contraction between the sample

and holder. The ASTM standard states that "The tensile strain, induced by the

differential thermal contraction of the specimen and holder, shall not exceed

0.05% for Nb 3 Sn." It was assumed that the sample holder materials supplied to

a subset of the round robin participants, National Electrical Manufacturers'

Association (NEMA) G-10 and G-ll tubes, met this specification. However, the

effect of the holders wall thickness on its circumferential thermal

contraction was not considered and, consequently, the ASTM maximum strain

specification was exceeded in some measurements. In retrospect, the

measurement discrepancies introduced by the use of sample holders having

different wall thicknesses and, thus, different thermal contractions might

have been anticipated and avoided.

The purpose of the round robin measurements is to document and better

understand the sources of inconsistencies in the measured critical current of

Nb 3 Sn conductors. This understanding is ultimately necessary for determining

a consistent and practical measurement method. In theory, a method that would

give consistent results is easily achieved by imposing extensive constraints

on all aspects of the measurement; however, this approach is impractical

4

f
considering the variety of I c measurement systems that are presently in use.

A practical measurement method should impose the smallest number of

constraints but yield acceptable results.

More challenging than the issue of consistent measurements is that of

accurate measurements. The basic problem that these round robin tests reveal

is that the correct and/or best method for critical-current measurements of

Nb 3 Sn is presently unclear. This difficulty is a result of the strain

sensitivity of Nb 3 Sn conductors. The question is, at what strain level should

the sample's I c be measured and how can this selected strain level be

consistently reproduced (with practicality in mind) at different laboratories?

Theoretically, this problem could be solved by simply requiring that all

laboratories measure the I c of Nb 3 Sn as a function of tensile strain; however,

there are several reasons why this is not a practical solution. The

measurement of I c as a function of strain requires specialized equipment that

is only available in a few laboratories. Also, among the laboratories that do

have this measurement capability, there are often considerable discrepancies.

These may be partially because that I c

- strain measurements are usually made

using straight, rather than coiled, samples. The straight samples are

normally shorter than coiled samples; consequently, the increased current

transfer voltage and decreased voltage tap separation associated with the

shorter samples can reduce measurement accuracy significantly. In general,

the I c- strain measurement is a complex one that introduces a number of new

measurement variables and, thus, it is unlikely that this approach would

enhance measurement consistency.

NIST was asked to prepare a document on a "Proposed Study of the Effect

of Sample Bonding Techniques on the Measured Critical Current of Nb 3 Sn

Superconductors" (Appendix I) . This is a detailed test plan to determine if

grease can be used as a sample bonding material instead of solder. This

document was published in the Proceedings of the 6^ Japan-US Workshop on

High-Field Superconducting Materials and Standards Procedures for High-Field

Superconducting Materials Testing.

D. Large Conductor I c Test Method

The work that has already been completed in the area of large-conductor

measurements and the planning for future studies has resulted in some special

concerns regarding the measurement of large -conductor critical currents.

These concerns fall into three related categories: increased electromagnetic

forces (Lorentz force), power supply effects, and large -conductor

characteristics

The high current levels required for large-conductor measurements result

in increased Lorentz forces on the sample that must be adequately supported by

the test fixture. As an example of the magnitude of these forces, aim

coiled sample carrying 3000 A in a magnetic field of 6 T experiences a total

Lorentz force of 18,000 N (4,037 lbs) or 180 N/cm (103 lbs/in) along its

length. Failure to constrain the sample under the influence of these large

forces can cause several different measurement problems, where the severity of

5
the problem depends on the amount of sample motion. The circuit loop formed

between an instrumented segment of the sample and its voltage tap leads in the

presence of a strong magnetic field acts as a potent electrical generator.

Under these conditions, even small sample motions can greatly increase the

voltage signal noise. Moreover, if the sample motion is sufficient, heat

generated by the motion can result in premature sample quenches. In the

extreme case of inadequate sample constraint, the conductor can be

irreversibly damaged by the Lorentz force. This situation is further

complicated by the need to maintain good thermal contact between the sample

and the liquid helium bath. In other words, overzealous attempts to provide

adequate mechanical support (e.g. thick epoxy ove r layments ) can result in

inadequate thermal contact and premature sample quenches.

Also, the generation of high currents requires the use of large power

supplies that often introduce a new set of measurement problems. Many of

these power supplies produce significant current ripple (periodic departure

from dc output level) , and this varying component of the sample current can

cause errors in the I c measurement [4] . (Another paper that calculates the

effect of current ripple is included as Appendix J.) This ac component of the

sample current produces a corresponding ac component of the sample voltage

that is then applied to the input of the dc voltmeter. The ac voltage is

present even at low current levels, i.e. those below the flux flow portion of

the voltage -current curve, because of the inductive loop formed between the

sample and voltage tap leads. The reaction of the dc voltmeter depends on the

amplitude of the ac voltage. At low amplitudes, for the two meters that have

been tested, the output response to an ac input signal was approximately the

time average of that signal. This characteristic of the voltmeter in

conjunction with the presence of current ripple can cause a reduction in the

measured I c

. At low ac amplitudes this behavior has been modeled and tested

to allow a quantitative estimation of its effect. However, at some ac

threshold level, which depends on the particular meter being used, the meter's

response to the ac voltage becomes erratic. At this ac level, the current

ripple has disabled the measurement system.

Ground loops are a problem that can exist in I c measurement systems

regardless of the size of the system's current supply; however, the effect of

ground loops is often amplified in a high current system [5,6]. The required

differential voltage output from a current supply used for I c measurements is

typically from 3-6 V. This low differential voltage is allowed by the

extremely low load resistance of the superconductor and its current leads.

Although the differential voltage at the current supply's output is low, the

voltage of the output with respect to earth ground (system voltage) can be

quite high (e.g. 40 V RMS). If there is only one point in the measurement

system's electrical circuitry that is coupled to ground, this high voltage

does not present a problem. On the other hand, if a second point in the

system is coupled to ground (a ground loop) , the high system voltage provides

a strong driving force for currents to flow from the system to ground. In

particular, if this ground current flows through the system's voltmeters,

which are often capacitively coupled to ground, the resulting spurious

voltages can lead to inaccurate or even meaningless measurement results. The

high system voltage can be reduced by connecting one point of the current

supply's output circuit to Earth ground, but this may also introduce a ground

6
loop. This conflict seems inherent in many high current I c measurement

systems

Other complexities of large conductor I c measurements result from the

conductor's increased cross-sectional dimensions and its electromechanical

characteristics. The large cross-sectional area of these conductors has

several ramifications for the I c measurement. First, the higher current

capacity that is allowed by the larger conductors necessitates longer current

contacts in order to avoid excessive resistive heating of the sample. Also,

the current transfer length (the minimum sample length that allows uniform

distribution [7] over the conductor's cross section, of the asymmetrically

injected sample current) is greater for the larger conductors. This situation

is complicated by the fact that the background magnetic field is an effective

impetus for current transfer. Consequently, in order to insure uniform

current distribution within a sample of a reasonable length, the majority of

the sample should reside in the high field region of the background magnet,

with the instrumented segment of the sample centered within a relatively small

portion of this region. Unfortunately, the cost of the background magnet

increases exponentially with the volume of its high field region. This

presents a practical limit on the size of the high field volume and, thus, on

the length of the instrumented segment of the sample. Moreover, this length

determines the voltage signal strength and, consequently, affects the

precision and accuracy of the I c measurement.

The mechanical construction of large conductors, whether monoliths or

cables, can introduce spatial variations in their electromagnetic properties.

Large monolithic conductors usually have a rectangular cross section to allow

easier winding. The unequal cross-sectional deformation that is necessary to

form a rectangular conductor can cause longitudinal variations in its

electromagnetic properties. A rectangular conductor also exhibits

electromagnetic asymmetries associated with the direction of the applied

magnetic field about its longitudinal axes. Both of these phenomena

are due to the aspect ratio effect [8]. In the case of cables, the degree and

type of mechanical deformation varies periodically along the length of each

strand. This mechanical variation results in a corresponding electromagnetic

variation. Specifically, these variations are caused by localized damage,

angular effects, and aspect ratio effects [9,10,11,12]. In both cases,

monoliths and cables, these spatial variations complicate the I c measurement.

The measurement of large conductor critical currents involves all of the

complexities associated with small conductors, many of which are amplified in

the case of large conductors, as well as some problems that are unique to the

large conductors. A large -conductor I c test method must be specific enough to

address all of these problems while maintaining enough generality to be of

practical benefit. A large conductor test method that is specific

to SSC (Superconducting Super Collider) cables has been proposed by Brookhaven

National Laboratory [13].

7
E. Critical-Current Measurements on Sample

Used in AC Loss Round Robin

The VAMAS Technical Working Party requested that we make critical -current

measurements on the U.S. wire used in the ac loss round robin. These critical

currents will be useful for analyzing and comparing the ac loss data. These

I c

's are reported at various applied magnetic fields, at 4.2 K, and with a

critical-current criterion of 0.1 juV/cm (see Table 1-1 and Figure 1-1). The

measurements were made on a straight sample in a radial-access, split-pair

magnet. The estimated uncertainty of the I c values is ±2%.

Table 1-1. Critical Currents at Various Applied Magnetic Fields.

/z 0 H, T 0.0 0.5 1 2 3 4 5 6 7 8 9

I c , A 1675 1344 1101 827 675 566 473 384 293 197 101

8
Figure 1-1. Critical current versus applied magnetic field at 4.2 K and a

critical-current criterion of 0.1 /xV/cm.

9
II. AC LOSSES

A. Short-Range Variations in Magnetic Hysteresis Loss

of Multifilamentary Nb3 Sn

We have experimentally investigated short-range variations in magnetic

hysteresis loss of multifilamentary Nb 3 Sn. The work was a collaboration among

R. B. Goldfarb (National Institute of Standards and Technology), D. B.

Smathers (Teledyne Wah Chang Albany), and M. M. Steeves (Massachusetts

Institute of Technology). The study was undertaken for two reasons. The

first is that measurements of hysteresis loss often use small, individually

reacted samples. We want to be confident that a small sample of individually

reacted wire is representative of a larger length of wire. By measuring the

losses for adjacent samples, we were able to quantify variations attributed to

metallurgy. The second reason is that we wished to determine the best way to

react samples that are metallurgically uniform. We compared samples prepared

with long "cold tails" and samples reacted with short "hot tails." In

addition, we estimated precision as a function of several measurement

variables

Hysteresis is the major loss mechanism in superconductors when fields are

cycled. It arises from irreversibility, owing to flux vortex pinning, of the

penetration of supercurrents into the conductor. As the field is cycled, the

penetrated supercurrents do not recede. Instead, oppositely circulating

supercurrents enter from the surface as new flux vortices are created. In

addition to energy dissipation, there is more information that can be

extracted from these measurements of magnetization versus field. The width of

the hysteresis loop is related to the critical current density for a given

filament diameter. The remanent magnetization at zero field gives an

indication of the residual field that could be expected in a superconducting

magnet constructed with the wire. Time dependence of the magnetization, on a

scale of seconds, gives information on flux creep.

Multifilamentary Nb 3 Sn is a metallurgically complicated system. The

intermetallic compound is very brittle and must be formed in situ. The

critical current of the reacted wire is very sensitive to mechanical strain.

Tin-core Nb 3 Sn wire was reacted on 8 mm (5/16—18) and 5 mm (10—24) threaded

stainless steel rods. Long "cold tails" extended out of the reaction furnace.

The coiled wire was unscrewed from the rods after reaction. Four samples,

each about 30 cm in length, were cut from each coil. The ends of each sample

coil were cut with a slitting saw to avoid crushing the filaments together.

The samples were soaked in phosphoric acid rust remover ("Naval Jelly") to

clean the surfaces without etching away the copper matrix. The samples'

density and volume were determined by Archimedes' principle. By etching some

samples in dilute nitric acid, we determined that, even with the more tightly

wound samples, the Nb 3 Sn filaments were not broken.

We used a vibrating- sample magnetometer to measure the hysteresis losses.

A feature of this method is that it requires only small samples compared to

magnetometers based on flux integration. For measurements in transverse

field, the composite wire is threaded onto a plastic screw sample holder, and

10
the field is applied along the axis of the sample coil. The temperature of

measurement was 4 K. We calibrated the magnetometer with Ni wire wound in the

same geometry as the superconductor samples. Maximum applied fields were

typically 2.4 MA/m (3 T in free space). The field was stepped.

Magnetization was measured after any matrix eddy currents were allowed to

decay. Magnetization was calculated as magnetic moment per unit volume of

wire (not per unit volume of Nb 3 Sn) . The area of the hysteresis loop

corresponds to the loss. The signal from the empty sample holder and rod is

diamagnetic, about one tenth the signal when the sample is measured. However,

the sample holder is nonhysteretic . Therefore the hysteresis loss and width

of the loop are not affected.

We measured the variations in hysteresis losses for contiguous samples of

multifilamentary Sn-core Nb 3 Sn reacted as a single coil. The characteristics

of the wire are given in Table II- 1.

Table II-l. Characteristics of Nb 3 Sn wire

Nb alloy

Diffusion barrier

Tin diffusion method

Wire diameter

Fraction Cu stabilizer

Local Cu/Nb ratio

Filament diameter

Approximate number of filaments

Number of filament subelements

1 wt.% Ti

vanadium

Sn core, modified jelly roll

0.78 mm

0.54

1.3 cm

3—3 . 5 /an

4000

18

The variation in hysteresis loss attributed to short-range metallurgical

variations (over about one meter) in the wire is less than 2.5%. In addition

to metallurgical variations in the samples, the other sources of variation,

listed in Table II-2, were examined and quantified.

One sample coil was stretched about 25% and then compressed. The winding

radius was increased slightly too. The hysteresis loss increased by 0.2%.

Unlike critical current, hysteresis loss appears to not be very sensitive to

strain.

Table II-2. Limits to precision

Metallurgical variations 2.5%

Volume determination 0.2%

Sample centering sensitivity 0.6%

Day-to-day repeatability 0.3%

Sample ends sawed vs. cut 0.5%

Thread size and pitch 4.0%

11
Two wire samples were reacted with short welded ends instead of long cold

tails. The purpose of welded tails or long tails is to prevent loss of Sn

from the cut ends of the wires during reaction. The loss for one sample with

apparently good welds, the "reference" sample, was as expected. The loss for

the second sample was 30% less. The ends of this sample were trimmed and the

loss was remeasured. This time the loss was 80% greater than that of the

reference sample. It is possible that improperly welded ends allowed Sn to be

lost from the sample. It is important to weld ends carefully when using this

method to prepare samples.

This work was presented at the Applied Superconductivity Conference,

21—25 August 1988, San Francisco, paper MC-6.

B. VAMAS AC Loss Interlaboratory Comparison

R. B. Goldfarb was the NIST participant in the VAMAS round- robin on ac

losses in multifilamentary Nb-Ti. Magnetic hysteresis losses were measured

for samples of the "U.S. wire" and the "Japan A wire." The characteristics of

the wires are given in Table II-3.

Samples of the wires were wound on the threads of a 5 -mm diameter nylon

screw. Because the Japan A wire had a small diameter, three strands were

wound together. The magnetic field was applied along the axis of the sample

coil, approximately transverse to the wire axis. Magnetization was measured

at 4 K with a vibrating- sample magnetometer. The magnetometer was calibrated

with pure nickel wire wound in the same geometry as the samples. The fields

were stepped rather than ramped. The area of the loops of magnetization

versus field corresponds to the magnetic hysteresis loss. The results are

exhibited in Table II -4. The applied field for each loop ranged from +Bm

to —

m , where Bm is the maximum value indicated in the table. The estimated

precision of each measurement is given.

Table II-4 gives the hysteresis loss per total sample volume. The loss

per volume of Nb-Ti is higher. For maximum fields of 3 T, the superconductor

loss is 107.4 kJ/m^ for the U.S. wire, and 95.8 kJ/m^ for the Japan A wire.

Table II-3. Characteristics of VAMAS wires

U.S. Wire Japan A Wire

Outer diameter (mm) 0.742 0.35

Filament diameter (/xm) 4.6 6.3

Number of filaments 10,980 760

NbTi/Matrix 1/1.38 21.5/78.5

Twist pitch (mm) 13 6

I c at 3 T (A) 653 50

J c per NbTi area (A/m^) 3.6 x 10 9 2.1 x 10 9

Vol

NbTi (%) 42.0 21.5

12
Table II -4. Hysteresis loss measurements

Wire Spool

Name Number

Sample

Vol. (cm^)

Maximum

Field (T)

Hysteresis

Loss (kJ/m^)

Precision

(%)

U.S. 2 0.1709 3 46.9 3

8 0.1733 3 45.1 1 14 0.0758 3 45.1 1

3 45.3 1

1 26.9 1 0.5 18.9 1

Japan A 0.1924 3 20.4 5

0.0962 3 20.6 3

1 12.4 3

0.5 7.4 1

The results of the VAMAS interlaboratory comparison on ac losses in Nb-Ti

were discussed at two meetings in 1989: the Japan—U.S. Workshop on High- Field

Superconducting Materials in February, and the International Cryogenic

Materials Conference in July. The sample wires, denoted U.S. (or "D" ) and

Japan-A (or "E"), were measured for positive field cycles of 1 to 0 to 1 tesla

with no transport current. The participants were instructed to use their

standard technique; there were not many restrictions placed on the

measurements. The laboratories measured magnetic hysteresis using flux

integration, vibrating- sample magnetometry , or calorimetry.

We analyzed the results reported at the February meeting. The mean loss

for the U.S. sample was 5.08 in units of millijoules per meter length of wire.

(Ordinarily, we prefer to express loss as energy per unit volume.) The

standard deviation was 0.59 mJ/m. The mean loss for the Japan-A sample was

0.505 mJ/m. The standard deviation was 0.059 mJ/m. The variations were

small, despite the variety of measurement methods. We compare in Figure II -1

the loss measurements for those laboratories reporting on both samples; each

point represents data from one laboratory. The linear fit reveals that the

variations are a consequence of systematic differences at the participating

laboratories. We believe that the primary cause of variation in reported

losses is inaccurate magnetic field calibration. Magnetization calibration is

usually easier to achieve and is done routinely.

As part of the VAMAS program, we measured ac losses on a coil of Nb-Ti

wire designated Japan-B (or "F2"

)

using a calorimeter on loan from

Laboratoires de Marcoussis, France. Calorimetry is difficult and requires

long lengths of wire. Its advantage is that all sources of dissipation are

measured. An ac transport current was used, but there was no applied magnetic

field. Our data (solid circles), scaled from 60 Hz to 50 Hz, are shown in

Figure II-2, together with the results of other participants reported at the

International Cryogenic Materials Conference.

13
JAPAN-A

Figure II-l. Each point represents losses measured at one laboratory on two

superconductor wires, "US" and "Japan-A."

Figure II-2. Calorimetric measurement of ac power loss, scaled to 50 Hz.

The lines are fits to the data of three laboratories. The

points are NIST data.

14
C. Hysteresis Loss Measurements in Support of the

U.S. Demonstration Poloidal Coil Project

We have provided measurement services to U.S. wire manufacturer Teledyne

Wah Chang Albany (TWCA) to characterize their tin-core Nb 3 Sn conductors for ac

losses. Figure II-3, from the TWCA data base, shows the hysteresis loss

(measured at NIST) and the critical current density for three series of wires.

Their conductor was used by the Massachusetts Institute of Technology

(MIT) in the U.S. Demonstration Poloidal Coil Project. We have measured halfcycle hysteresis losses for MIT and found that half-cycle losses were 42.6% of

full-cycle losses. Appendix K is a paper by the MIT group on calorimetric

loss measurements which incorporates our data.

We have also measured hysteresis losses in several ISH^Sn wires made by

IGC Advanced Superconductors. Some of the data appear in their paper

published in IEEE Transactions on Magnetics 25 . 2204 (1989).

D. Participation in Workshop on AC Loss in

Superconducting Magnets for ITER

At the invitation of Lawrence Livermore Laboratory, we participated in a

workshop on ac losses for the International Thermonuclear Experimental Reactor

(ITER) in July 1989. AC losses are expected to contribute significantly to

the refrigeration load of this machine. We reviewed the major sources of ac

losses for ITER: (1) magnetic hysteresis loss in the superconductor, which is

controlled by filament diameter; (2) self-field loss from the pulsed transport

current, which is reduced by cabling; (3) eddy-current filament coupling loss,

which is reduced by twisting the filament bundle; (4) eddy currents in the

superconductor matrix and conduit; and (5) magnetic hysteresis loss in the

conduit

We discussed and gave example data for three methods for characterizing a

conductor for losses: (1) calorimetry, (2) magnetization as a function of

field, and (3) ac susceptibility vs. temperature. While most magnetization

measurements are done with fields transverse to the conductor, data are needed

on longitudinal magnetization for the ITER magnets.

E. Disclaimer

Certain commercial products are identified to specify the experimental

study adequately. In no case does such identification imply recommendation or

endorsement, nor does it imply that the products are necessarily the best

available for the purpose.

15
±3T

HYSTERESIS

LOSS

(

NON

-COPPER

mJ/cm^ )

Figure II-3. Hysteresis loss and critical current density for Nb 3 Sn wires

made by Teledyne Wah Chang Albany. Each curve is composed of

many points representing different billet designs.

16
III. ADDITIONAL ACTIVITIES

There were discussions of standards and round robin measurements at a

number of meetings over the last year that are pertinent to the DoE standards

program. A summary and comments from the floor are presented.

A. VAMAS Technical Working Party Meeting

July 8-16, 1988, Southampton, England

There were 20 people that attended this meeting, 14 from European

countries, 5 from Japan, and 1, Loren Goodrich, from the USA. Dr. David Gould

(EEC) is the VAMAS coordinator. Prof. K. Tachikawa (JPN) is the chairman of

both subgroups, Superconducting Materials and Cryogenic Materials. Dr.

Hitoshi Wada (JPN-NRIM) is the secretary of the TWP Office.

The topics of discussion included the analysis of the Nb 3 Sn criticalcurrent round robin test, the NbTi ac loss round robin test that is now in

progress, and future critical-current round robins. NIST participation in

these discussions was very important since we were the sole representative for

all the USA participants and our input at the meeting was fundamental in the

analysis of the Nb 3 Sn round robin.

Prof. Tachikawa presented the first draft of the Japanese report on the

analysis of all the USA, JPN, and EEC data on all three Nb 3 Sn wire samples.

This draft had been circulated to all 24 participating laboratories well

before the meeting. The results are presented in numerous plots of I c vs.

magnetic field and anonymity of each participant's data is insured by

identifying the individual data sets solely by arbitrarily selected letters

rather than laboratory names. Standard deviations (after excluding

extraordinarily outlying results) of the measured I c at 10 and 15 T for each

sample are given. NIST had sent several corrections to the USA data,

suggested changes to the sections that related to our contribution, and a

suggested re-write of the summary and recommendations section. Other than the

wording, we at NIST felt that the report was a reasonable description of the

results. The statistical analysis was somewhat misleading because the errors

were not random, but our new summary emphasized the systematic error evident

in these data. The NIST study of sample holder effects provided further

evidence that strong systematic effects can be present in these critical

current measurements

Dr. Colin Walters prepared another analysis of the round robin data.

This analysis was intended to supplement rather than to be an alternative to

Prof. Tachikawa 's report. The difference in the two approaches was that Dr.

Walters fitted I c vs. magnetic field for each specimen of each participant

using the approximate Kramer flux pinning law. This has some difficulties

over wide field ranges, but does smooth the data and allow for interpolation

or extrapolation. The results are compared in this report at a magnetic field

of 11 T. At the time of this meeting, the results of this analysis were far

less complete than the Japan draft report and copies of this report had not

been received by laboratories outside Europe before this meeting.

17
VAMAS has its own publication medium and it is intended that both

reports, Prof. Tachikawa's and Dr. Walter's, will be published therein.

Dr. Gould will attempt to create a somewhat shorter summary report that will

use elements from both of the larger reports. This report will be 6 to 10

pages long and is intended to be published in Cryogenics . Dr. Gould thought

it would be best if we held the paper on NIST results until the first joint

report is made. It may be possible to publish Dr. Gould's report and the NIST

report in the same issue of Cryogenics .

Four NbTi conductors have been selected for the ac-loss round robin test,

two from JPN (sample A: 6.3 gm filaments and CuNi matrix, sample B: 0.5 gm

filaments and CuNi matrix), one from USA (4.6 gm filaments and Cu matrix), and

one from EEC (0.17 gm filaments and CuNi matrix). The two larger filament

samples (JPN A and USA) have been distributed and the deadline for results was

set for the end of October (subsequently changed to the end of December 1988)

We were asked to collect these results from the USA participants and send them

to Dr. Hitoshi Wada (secretary of the VAMAS TWP office) and to

Prof. L. van de Klundert (one of the European participants)

Both of the small filament samples were to be supplied as 20 cm long

coils; however, few laboratories can measure a sample this large. Prof. L.

van de Klundert suggested that everyone measure the sample using his method, a

resonance technique using a transformer. However, it was stated that this

would undermine the purpose of these tests, to compare the individual methods

normally used by each participant. We need to contact the U.S. specialist to

find out if a free wire sample or another coil shape could be measured.

The four selected samples will provide a range of ac losses, but the two

smaller filament samples will be very difficult to measure. A 500 m length of

the 0.5 gm filament diameter sample is expected to have a loss of 1 . 5 W at 50

Hz and 1200 V. The 0.17 gm filament diameter sample will have a much lower

loss and it is not insulated; consequently, this sample will require

insulation prior to sample winding. We were asked to measure the I c of the

USA wire over a wide range of magnetic fields, including 0 T. The zero field

I c will be well over 1000 A and will have self-field effects.

B. American Society for Testing and Materials (ASTM) Meeting

August 22, 1988 at ASC in San Francisco, California

An ASTM B1.08 meeting was held in conjunction with the Applied Superconductivity Conference (ASC) in San Francisco on August 22, 1988. Although

there were over 900 registrants for the conference, a notice of the meeting

was posted, and an announcement of the meeting was made at the initial plenary

session, only ten people attended the ASTM meeting. The meeting was held late

in the evening. Unavoidably, the meeting conflicted with one poster session

and a social hour; however, this was the most open time period during the

conference. The decision to hold the meeting in conjunction with the ASC

conference was based on the already crowded conference schedule for

superconductivity. A large attendance was anticipated based on the superconductivity community's expressed realization that there are new problems

with measurements and definitions, especially for the high-Tc superconductors.

18
In retrospect, the meeting announcement should have been specifically directed

at a wider audience than the present ASTM members. In the future, it might be

more productive to hold an ASTM meeting of this type at a smaller conference

or appropriate workshop as a scheduled part of the agenda.

The meeting closely followed the detailed agenda with open discussion on

each item. Measurement concerns relating to both conventional (NbTi and

Nb 3 Sn) and to high-Tc superconductors were discussed. There was an overall

balance in the discussions relating to conventional materials, to high-Tc materials, and items common to both material types.

One topic of discussion relating to conventional materials was the extent

to which existing standards (B 713-82 and B 714-82) are being used. There was

a consensus among the group that these standards are completely relevant in

regard to the new high-Tc materials, but additional guidelines are needed. It

was generally agreed upon that there are some measurement considerations

unique to high-Tc conductors at their present level of development. The

existing ASTM standards could be appended to address these considerations

without sacrificing their present integrity and commercial adequacy, but this

may be premature

There were no objections to reapproving the existing standards. Some

minor changes and additions to these standards, prior to initiating the

reapproval process, are being considered. Other specific comments are listed

below.

• Representatives from France indicated that the ASTM standard test method was

used for the basis for the development of a similar French standard. Their

test method is used beyond the 600 A limit of the ASTM standard.

• There is a problem with normalizing the results of ac loss measurements:

expressing the loss in terms of W/m^ does not take into account the I c of the

conductor

• A controversy exists over the minimum copper- to -superconductor ratio needed

for magnet stability.

• It should be kept in mind that, regardless of their unique properties, the

high-Tc materials are still superconductors and the existing standards are

valid for them.

• Given the present critical-current densities for the high-Tc materials, it

is difficult to test these conductors at effective resistivities below that of

copper

• The measured properties of the high-Tc conductors are strongly dependent

upon the criteria used.

• Standards for high-Tc conductors are premature at this time based on the

rapidly changing technology and the limited commercial application at their

present level of development. Should an ASTM standard be developed for highTc prior to the existence of a significant commercial market?

19
• Standards for high-Tc materials would need to accommodate unpredictable

changes in their properties; thus, this type of standard would have limited

utility.

• The superconductor industry could be impaired by overspecification of their

products

• Accommodation of high-Tc conductors by the measurement standard may be

premature at this time, but a Guide or a Guideline may be appropriate.

• It should not be the role of ASTM to set safety standards, but general

safety concerns should be mentioned and appropriate references given. There

are concerns over the toxicity and combustibility of the high-Tc materials.

Combustibility of fine filament conventional materials is also of concern.

• Are the present standards being used? If not, how can they be changed to be

more useful in the commerce of superconductors?

• The ASTM Standard Test Method was used in specifying the conductor for the

Superconducting Super Collider, but extensive additional specifications were

also necessary. Could the Standard be made more complete or would this result

in an overspecification for most conductors?

• The ASTM Standard should be general so as to accommodate all applications.

• Items suggested as candidates for addition to the Standard Definitions:

filament sausaging, n-value, Meissner effect, and ac losses.

• The n-value (shape of the voltage-current relationship) is an important

parameter but the error in its measurement can be large.

\ • The level of ac loss for different conductors varies greatly depending

on the application for which they were designed.

C. 9^ Workshop on Niobium Titanium Superconductors

January 22-25, 1989, Pacific Grove, California

This workshop was held in Pacific Grove, California from January 22-25,

1989. L. F. Goodrich gave an audience participation talk on ASTM (American

Society for Testing and Materials) Standards for superconductors with a focus

on NbTi conductors. There was an overview of the existing standards and

possible additions were described. One addition that was considered was a

test method for large conductors, more than 600 A. A general question that

was presented was: how can the ASTM Standards be made more useful? The

general consensus was that the definitions could be expanded, and we are still

feeling our way in large conductor testing. Other specific comments are

listed below.

• We need to wait and see if the present SSC specification is effective

before considering it as an appendix to the present ASTM standard. The spring

back test and other tests are new and substantially untested. Given the

20
public trust in ASTM, we shouldn't base any additions on the present

specifications since we are still feeling our way along.

• There used to be concern about I c measurements and that's why there was a

lot of interest in ASTM standards in 1980 to 1982. Now it isn't clear that

there is a problem.

• A wire manufacturer said his customers need self-field information.

D. 6th Japan-US Workshop on High Field Superconductors

February 22-24, 1989, Boulder, Colorado

This workshop was held in Boulder, Colorado from February 22-24, 1989.

L. F. Goodrich gave an audience participation talk on ASTM Standards for

superconductors with a focus on Nb 3 Sn conductors. There was an overview of

the existing standards and possible additions were described. A general

question that was presented was: how can the ASTM Standards be improved?

There was no consensus on a number of key topics, such as reaction mandrel

material, measurement mandrel material, and bonding method. The implication

of this is that a future round robin will have to prove which method gives the

best or most consistent results. Other specific comments are listed below.

• The sample holder material should not be restricted at this time; several

different materials should be tested.

• Choose sample holder material and mounting technique that is easily

reproducible

• As is , the standard is not specific enough with respect to sample holder

material, geometry, etc.

• The answer you get does not matter because it is a comparative measure - give customer an absolute comparison.

• Suggested holder-bonding: thin walled G-10 tube, varnish, measure with

force in.

• Is all varnish the same; is there variation in the G-10 material; are there

ASTM standards for these materials?

• There Is ambiguity in the ASTM test method with regard to the allowed

differential contraction between the sample and holder.

• Practicality is a consideration, holder should be reusable.

• The problem with soldering to stainless steel is an unknown shunt current

and an artificially lower n-value.

• Stainless steel holders can result in excessive precompression.

21
• Reaction holders made from different materials (graphite, niobium,

stainless steel, and copper) result in different I c

's even when measurements

are made on the same measurement holder.

• There will be a problem in conductor procurement unless a consensus is

reached on a measurement mandrel material to be used for acceptance testing.

At present both G- 10 and stainless steel are being used.

E. Versailles Project on Advanced Materials and Standards (VAMAS)

Technical Working Party (TWP) Meeting

July 24, 1989, Los Angeles, California

There was a VAMAS TWP meeting on superconductivity at the International

Cryogenic Materials Conference in Los Angeles, California , on July 24, 1989.

The topics of discussion included the analysis of the Nb 3 Sn critical-current

round robin test, the NbTi ac loss round robin test, and future round robins.

H. Wada, National Research Institute for Metals (Japan), is Head of the Office

for the TWP, and he has printed minutes of this meeting (TWP Office ReportA(7)). NIST made two presentations to aid in the next VAMAS round robin

measurement on Nb 3 Sn critical current. These are included as Appendices L and

M of this report.

A detailed description of a test method for Nb 3 Sn critical current was

prepared for the VAMAS TWP Office and is included as Appendix N. This

measurement guideline is intended to be one of a few different measurement

techniques in a second Nb 3 Sn I c round robin that is being planned to determine

the preferred measurement method for M^Sn.

K. Tachikawa, the chairman of the VAMAS TWP, has set the schedule of

future round robins in the TWP Office Report-A(7) . A Nb 3 Sn I c round robin is

scheduled to start in October 1989 and end in July 1990. Also a large

conductor I c round robin will start in February 1990. Our NbTi Reference

Material might be a good candidate for these measurements. A I c round robin

on new conductors is scheduled to start in February 1990. The first

intercomparison of ac loss measurements (NbTi conductors) will be completed by

December 1989. A comprehensive ac loss measurement round robin will start in

April or May of 1990.

F. American Society for Testing and Materials (ASTM) Meeting

October 25-26, 1989, Denver, Colorado

There was an ASTM meeting in Denver, Colorado, on October 25-26, 1989. L.

F. Goodrich, now the chairman of Subcommittee B01.08 on superconductors,

presented a status report on superconductivity to Committee B01 on electric

conductors. The existing ASTM Standards, B713 and B714, are to be submitted

for reapproval.

22
IV . ACKNOWLEDGMENTS

D. F. Vecchia and J. D. Splett are developing the statistical model and

analysis for the large-conductor, critical-current Reference Material.

T. C. Stauffer and R. M. Folsom prepared samples for critical current

measurements. A. N. Srivastava and L. L. Dulcie assisted with computer

programming. R. L. Spomer, R. J. Loughran, and R. R. Buell assisted with the

magnetization measurements. R. J. Loughran made the calorimeter measurements

with assistance from R. W. Cross. J. R. Cave graciously arranged for the

loan of the calorimeter. J. W. Ekin participated in discussions.

V. REFERENCES

[1] SRM 1457 can be obtained from: Office of Standard Reference Materials,

Room B311, Chemistry Building, National Institute of Standards and

Technology, Washington, DC 20234, phone (301) 975-6776.

[2] E. M. W. Leung, H. G. Arrendale , R. E. Bailey, and P. H. Michels, "Short

Sample Critical Current Measurements Using a Superconducting

Transformer," Adv. Cry . Eng

. Vol. 33 . Plenum Press, New York (1988), pp

219-226.

[3] K. Tachikawa, "The VAMAS Intercomparison of Critical Current Measurements

in Superconducting Nl> 3 Sn Wires," Cryogenics 29_, pp . 710-715 (1989).

[4] L. F. Goodrich and S. L. Bray, "Current-Ripple Effect on Superconductive

d.c. Critical Current Measurements," Cryogenics 28., pp . 737-743 (1988).

[5] L. F. Goodrich, S. L. Bray, W. P. Dube, E. S. Pittman, and A. F. Clark,

"Development of Standards for Superconductors, Interim Report Jan. -Dec.

1985," NBSIR 87-3066, National Bureau of Standards, Boulder, Colorado

(April 1987).

[6] Ralph Morrison, Grounding and Shielding Techniques in Instrumentation . John Wiley & Sons, New York (1977).

[7] J. W. Ekin, "Current Transfer in Multifilamentary Superconductors I.

Theory," J. Appl. Phys . , 49, 3406-3409 (1978).

[8] L. F. Goodrich, W. P. Dube, E. S. Pittman, and A. F. Clark, "The Effect

of Aspect Ratio on Critical Current in Multifilamentary Superconductors,"

Adv. Cryo . Eng. - Materials . Vol . 32 . Plenum Press, New York (1986).

[9] J. W. Ekin, L. F. Goodrich, J. Moreland, E. S. Pittman, and A. F. Clark,

"Electromechanical Properties of Superconductors for High-Energy Physics

Applications," NBSIR 86-3061, National Bureau of Standards, Boulder,

Colorado (December 1986) .

[10] L. F. Goodrich, E. S. Pittman, J. W. Ekin, and R. M. Scanlan, "Studies of

NbTi Strands Extracted from Coreless Rutherford Cables," IEEE Trans . on

Magnetics . MAG 23, 1642, 1987.

23
[11] L. F. Goodrich and S. L. Bray, "Current Capacity Degradation in

Superconducting Cable Strands," IEEE Trans, on Magn. 25(2) 1949-1952

(March 1989) .

[12] J. W. Ekin, L. F. Goodrich, S. L. Bray, N. F. Bergren, and R. B.

Goldfarb, "Electromechanical Properties of Superconductors for HighEnergy Physics Applications," NISTIR 89-3912, National Institute of

Standards and Technology, Boulder, Colorado (November 1989).

[13] M. Garber and W. B. Sampson, "Test Methods for Cable Critical Currents

and Normal State Resistance," SSC-N-488, Brookhaven National Laboratory,

Upton, New York (March 1988)

24
APPENDIX A

PUBLICATIONS OF RESEARCH PROGRAM

1. A. F. Clark and J. W. Ekin, "Defining Critical Current," IEEE Trans. Mag.

MAG-13 . 38-40 (1977).

2. R. L. Powell and A. F. Clark, "Definitions of Terms for Practical

Superconductors, 1. Fundamental States and Flux Phenomena," Cryogenics

17, 697-701 (1977).

3. R. L. Powell and A. F. Clark, "Definitions of Terms for Practical

Superconductors, 2. Critical Parameters," Cryogenics IJS, 137-141 (1978).

4. J. W. Ekin, "Current Transfer in Multifilamentary Superconductors. I.

Theory," J. Appl. Phys . 49, 3406-3409 (1978).

5. J. W. Ekin, A. F. Clark, and J. C. Ho, "Current Transfer in

Multifilamentary Superconductors. II. Experimental Results," J. Appl.

Phys. 49, 3410-3412 (1978).

6. A. F. Clark, J. W. Ekin, R. Radebaugh, and D. T. Read, "The Development

of Standards for Practical Superconductors," IEEE Trans. Mag. MAG -15 . 224-227 (1979).

7. F. R. Fickett and A. F. Clark, "Development of Standards for

Superconductors, Annual Report FY 79," NBSIR 80-1629, National Bureau of

Standards, Boulder, Colorado (Dec. 1979).

8. F. R. Fickett and A. F. Clark, "Standards for Superconductors," Prof.

1979 Mech. and Mag. Energy Storage Contractors' Review Meeting, Conf790954, U.S. Dept, of Energy, Washington, DC (Dec. 1979).

9. D. T. Read, J. W. Ekin, R. L. Powell, and A. F. Clark, "Definitions of

Terms for Practical Superconductors: 3. Fabrication, Stabilization, and

Transient Losses," Cryogenics .19, 327-332 (1979).

10. F. R. Fickett and A. F. Clark, "Development of Standards for

Superconductors," Prof. 8th Int'l. Cryo . Eng. Conf. (ICEC8), IPC Science

and Technology Press (1980) pp . 494-498.

11. F. R. Fickett, S. B. Kaplan, R. L. Powell, R. Radebaugh, and A. F. Clark,

"Definitions of Terms for Practical Superconductors: 4. Josephson

Phenomena," Cryogenics 20, 319-325 (1980).

12. G. Fujii, J. W. Ekin, R. Radebaugh, and A. F. Clark, "Effect of Thermal

Contraction of Sample Holder Material on Critical Current," Adv. Cryo.

Eng. Vol. 26 . Plenum Press, New York (1980) p. 589-598.

25
13. G. Fujii, "Present Practices in Japan for the Measurement and Definition

of Various Superconducting Parameters," Tech. Report ISSP Univ. of Tokyo,

Ser. A, No. 1063 (July 1980).

14. G. Fujii, J. W. Ekin, R. Radebaugh, and A. F. Clark, "Effect of Thermal

Contraction of Sample-Holder Material on Critical Current," Tech. Report

ISSP Univ. of Tokyo, Ser. A, No. 1074 (Aug. 1980).

15. A. F. Clark, "Development of Standards for Practical Superconductors,"

Proc. Superconductivity Technical Exchange Meeting, IAPG , PIC-ELE-SC

209/1 (1980), p. 103 (presentation only).

16. R. Radebaugh, G. Fujii, D. T. Read, and A. F. Clark, "A Standards Program

for ac Losses in Superconductors," Proc. IVth Int'l. Congress on

Refrigeration (1980), Comm. A h-10 (1980), pp . 1-4.

17. F. R. Fickett, L. F. Goodrich, and A. F. Clark, "Development of Standards

for Superconductors, Annual Report FY80," NBSIR 80-1642, National Bureau

of Standards, Boulder, Colorado (December 1980).

18. G. Fujii, "Present Practices in Japan for the Measurement and Definition

of Various Superconducting Parameters," Cryogenics 2X, 21-38 (1981).

19. H. R. Segal, Z. J. J. Stekley, and T. A. DeWinter, "Development of

Critical Current Measurement Standards," Proc. 1980 Applied

Superconductivity Conference, IEEE Trans. Mag. MAG -17 . 73-76 (1981).

20. G. Fujii, M. A. Ranney, and A. F. Clark, "Thermal Expansion of Nb 3 Sn and

V3 Ga Multifilamentary Superconducting Cables, Fiberglass - Epoxy and

Cotton-Phenolic Composite Materials," Jap. J. Appl. Phys . 20, 1267-1270

(1981) .

21. A. F. Clark, G. Fujii, and M. A. Ranney, "The Thermal Expansion of

Several Materials for Superconducting Magnets," IEEE Trans. Mag. MAG -17 . 2316-2319 (1981).

22. L. F. Goodrich and J. W. Ekin, "Lap Joint Resistance and Intrinsic

Critical Current Measurements on a NbTi Superconducting Wire," IEEE

Trans. Mag. MAG-17 69-72 (1981).

23. F. R. Fickett and L. F. Goodrich, "NBS Superconductor Standardization

Program," Proc. 7th Int. Conf. on MHD Electrical Power Generation . 87-89

(October 1981 ) .

24. L. F. Goodrich, J. W. Ekin, and F. R. Fickett, "Effect of Twist Pitch on

Short-Sample V-I Characteristics of Multifilamentary Superconductors,"

Adv. Cryo. Eng. 28, 571-580 (1982).

25. L. F. Goodrich and F. R. Fickett, "Critical Current Measurements: A

Compendium of Experimental Results," Cryogenics 22, 225-241 (1982).

26
26. A. F. Clark, L. F. Goodrich, F. R. Fickett, and J. V. Minervini,

"Development of Standards for Superconductors, Interim Report October

1980- January 1982," NBSIR 82-1678, National Bureau of Standards, Boulder,

Colorado (July 1982)

27. Standard Definitions of Terms Relating to Superconductors, Annual Book of

ASTM Standards, ASTM B713-82, Part 2.03, pp . 591-594, American Society

for Testing and Materials, Philadelphia, PA (1983).

28. Standard Test Method for D-C Critical Current of Composite

Superconductors, Annual Book of ASTM Standards, ASTM B714-82, Part 2.03,

pp. 595-598, American Society for Testing and Materials, Philadelphia, PA

(1983)

29. L. F. Goodrich, "The Effect of Field Orientation on Current Transfer in

Multifilamentary Superconductors," IEEE Trans. Magn. MAG -19 . 244-247

(1983)

30. L.^ F. Goodrich, D. F. Vecchia, E. S. Pittman, and A. F. Clark, "Critical

Current Measurements on a NbTi Superconducting Wire Standard Reference

Material," Adv. Cryo . Eng. 30, 953-960 (1984).

31. A. F. Clark, L. F. Goodrich, and F. R. Fickett, "Experience in

Standardizing Superconductor Measurements," Journal de Physique, Colloque

Cl, supplement n°l, pp . Cl - 379 -Cl - 382 (Tome 45, janvier 1984).

32. L. F. Goodrich, D. F. Vecchia, E. S. Pittman, J. W. Ekin, and A. F.

Clark, "Critical Current Measurements on an NbTi Superconducting Wire

Standard Reference Material," NBS Special Publication 260-91, National

Bureau of Standards, Boulder, Colorado (September 1984).

33. A. F. Clark and L. F. Goodrich, "Characterization of a Standard Reference

Superconductor for Critical Current and a Summary of Other Standard

Research at NBS," Proc . Tenth International Cryogenic Engineering

Conference (1984), Butterworth & Co., Ltd., pp . 433-437 (1984).

34. L. F. Goodrich, J. V. Minervini, A. F. Clark, F. R. Fickett, J. W. Ekin,

and E. S. Pittman, "Development of Standards for Superconductors, Interim

Report January 1982 -December 1983," NBSIR 85-3027, National Bureau of

Standards, Boulder, Colorado (January 1985).

35. F. R. Fickett, "Standards for Measurement of the Critical Fields of

Superconductors," NBS J. Res., 90(2), 95-113 (March-April 1985).

36. L. F. Goodrich, W. P. Dube, E. S. Pittman, and A. F. Clark, "The Effect

of Aspect Ratio on Critical Current in Multifilamentary Superconductors,"

Adv. Cryo. Eng. - Materials 32, Plenum Press, New York (1986).

37. W. P. Dube and L. F. Goodrich, "Quench Detector Circuit for

Superconducting Testing," Rev. Sci. Instrum. 57., 680-682 (1986).

27
38 . L. F. Goodrich and A. F. Clark, "Joint Resistivity in Superconductors,"

Proc. Tenth International Cryogenic Materials Conference , Published in

the U.S.S.R., pp. 256-258 (1986).

39. R. B. Goldfarb and J. W. Ekin, "Hysteresis Losses in Fine Filament

Internal-Tin Superconductors," Cryogenics 26, 478-481 (August 1986).

40. L. F. Goodrich, S. L. Bray, W. P. Dube, E. S. Pittman, and A. F. Clark,

"Development of Standards for Superconductors, Interim Report January

December 1985," NBSIR 87-3066, National Bureau of Standards, Boulder,

Colorado (April 1987)

41. L. F. Goodrich, "Development of Standards for Superconductors, Interim

Report January 1986-December 1987," NBSIR 88-3088, National Bureau of

Standards, Boulder, Colorado (February 1988).

42. L. F. Goodrich and S. L. Bray, "Current Ripple Effect on Superconductive

d.c. Critical Current Measurements," Cryogenics 28, 737-743 (November

1988)

43. L. F. Goodrich, S. L. Bray, and A. F. Clark, "Current-Ripple Effect on

Superconductive dc Critical Current Measurements," Adv. Cryo . Eng. - Materials 34, 1019-1026, Edited by A. F. Clark and R. P. Reed (Plenum

Publishing Corp., 1988).

44. L. F. Goodrich, S. L. Bray, and T. C. Stauffer, "Nb 3 Sn Critical-Current

Measurements Using Tubular Fiberglass-Epoxy Mandrels," IEEE Trans, on

Magn. 25(2) 2375-2378 (March 1989).

45. K. Tachikawa, K. Itoh, H. Wada, D. Gould, H. Jones, C. R. Walters, L. F.

Goodrich, J. W. Ekin, and S. L. Bray, "VAMAS Intercomparison of Critical

Current Measurement in Nb 3 Sn Wires," IEEE Trans, on Magn. 25.(2) 2368-2374

(March 1989)

46. R. B. Goldfarb, D. L. Reid, T. S. Kreilick, and E. Gregory, "Magnetic

Evaluation of Cu-Mn Matrix Materials for Fine-Filament Nb-Ti

Superconductors," IEEE Trans, on Magn. 2.5(2) 1953-1955 (March 1989).

47. S. L. Bray, L. F. Goodrich, and W. P. Dube, "Battery Powered Current

Supply for Superconductor Measurements," Rev. Sci. Instrum. 60, 261

(1989) .

48. L. F. Goodrich and S. L. Bray, "Critical-Current Measurements of Nb 3 Sn

Superconductors: NBS Contribution to the VAMAS Inter laboratory

Comparison," Cryogenics 29., 699-709 (July 1989).

49. L. F. Goodrich, S. L. Bray, and T. C. Stauffer, "Thermal Contraction of

Fiberglass-Epoxy Sample Mandrels and Its Effect on Critical - Current

Measurements , " Proceedings of the 6th Japan-U.S. Workshop on High-Field

Superconducting Materials and Standard Procedures for High-Field

Superconducting Materials Testing . 91-93, Edited by K. Tachikawa, K.

Yamafuj i , H. Wada, J. W. Ekin, and M. Suenaga (Japan 1989).

28
50 . L. F. Goodrich and S. L. Bray, "Proposed Study on the Effect of Sample

Bonding Techniques on the measured Critical Current of Nb 3 Sn

Superconductors , " Proceedings of the 6th Janan-U.S. Workshop on HighField Superconducting Materials and Standard Procedures for High-Field

Superconducting Materials Testing . 99-108, Edited by K. Tachikawa, K.

Yamafuj i , H. Wada, J. W. Ekin, and M. Suenaga (Japan 1989).

51. L. F. Goodrich, S. L. Bray, and T. C. Stauffer, "Thermal Contraction of

Fiberglass-Epoxy Sample Holders Used for Nb 3 Sn Critical-Current

Measurements," to be published in: Adv. Cryo . Eng . -Materials 36.

52. L. F. Goodrich and S. L. Bray, "Integrity Tests for High-Tc and

Conventional Critical-Current Measurement Systems," to be published in:

Adv. Cryo. Eng . -Materials 36 .

53. L. F. Goodrich, R. B. Goldfarb, and S. L. Bray, "Development of Standards

for Superconductors, Final Report January 1988 -December 1989," to be

published as NISTIR 90-xxxx, National Institute of Standards and

Technology, Boulder, Colorado (1990).

29

APPENDIX B

INTEGRITY TESTS FOR HIGH-TC AND CONVENTIONAL

CRITICAL- CURRENT MEASUREMENTS SYSTEMS

L. F. Goodrich and S. L. BrayPresented at the International Cryogenic Materials Conference

Los Angeles, California

24-28 July 1989

to be published in: Advances in Cryogenic Engineering

Materials, Volume 36

31
INTEGRITY TESTS FOR HIGH-TC AND CONVENTIONAL

CRITICAL- CURRENT MEASUREMENT SYSTEMS*

L. F. Goodrich and S. L. Bray

Electromagnetic Technology Division

National Institute of Standards and Technology

Boulder, Colorado 80303

ABSTRACT

Critical-current measurement systems must be extremely sensitive to the

small differential voltage that is present across the test specimen as it

changes from the zero resistance state to the flux- flow resistance state.

Consequently, these measurement systems are also sensitive to interfering

voltages. Such voltages can be caused by ground loops and by common mode

voltages. Specific methods for testing the sensitivity of critical-current

measurement systems and for detecting the. presence of interfering voltages

are discussed. These include a simple procedure that simulates the zero

resistance state and the use of an electronic circuit that simulates the

flux-flow resistance state.

INTRODUCTION

The determination of a superconductor's critical current (I c ) requires

the measurement of extremely low voltages, ^ on the order of 1 /iV.

Consequently, the I c measurement system must be quite sensitive to the

resistive voltage that appears as the test specimen changes from the

superconducting to the normal state, and be insensitive to other sources of

voltage that might otherwise corrupt the measurement. The I c measurement

system is susceptible to sources of interfering voltage that might be

negligible for many other measurements. Ground loop and common mode

voltages^ are prime examples of these sources of interference.

Because of the nebulous character of these voltages, it is often

difficult to predict, based simply on the design of the measurement system,

whether or not the system is prone to these problems. However, some

practical test methods that are useful for checking the sensitivity and

accuracy of the measurement system and for detecting the presence of

interfering voltages have been developed. The test methods do not directly

indicate the sources of problems; they simply indicate their presence.

Consequently, the solution of these measurement problems depends on

knowledge of their likely sources. As a diagnostic tool, these tests are

best used for evaluating the success of specific changes that are intended

to alleviate these problems.

Z"Contribution of NIST, not subject to copyright.

32
There are two general test methods, the zero resistance test and the

finite resistance test. The zero resistance test is used to detect the

presence of interfering voltages, and the finite resistance test is used to

evaluate the sensitivity and accuracy of the measurement system. Both tests

can be conducted at either room or cryogenic temperature. The room

temperature tests are usually more convenient, but less definitive because

they do not depict the actual measurement conditions as precisely as the

cryogenic tests.

Finally, the development of an I c measurement system is complicated by

the expensive and volatile nature of the liquid cryogen. This is

particularly true when the cryogen is helium and the measurement system is

computer controlled. During the first I c measurements for a new or modified

measurement system, large quantities of liquid helium are often expended

during the inevitable debugging process. To address this problem, a simple

electronic circuit that simulates the voltage-current (V-I) characteristic

of a superconductor has been developed and tested. The simulator is an

effective tool for developing the measurement system to a high level, before

expending liquid helium. Also, for complex I c measurement systems, the

simulator has proven useful as part of a pre-operation check. In this way,

problems can be detected and corrected prior to transferring liquid cryogen

from the storage Dewar to the measurement system cryostat.

TEST METHODS

Finite Resistance Test

The finite resistance test is simply a four-wire resistance measurement

where the superconductor specimen is replaced in the measurement system by

an appropriate copper conductor. The idea is to measure a specimen that has

a known resistance to assess the accuracy and sensitivity of the measurement

system. In order to closely approximate the actual measurement conditions,

it is important for the copper specimen's resistance, over the length that

is spanned by the voltage taps, to be similar to that of the superconductor

at its critical current. This allows testing of the measurement system at

an appropriate voltage and current.

Zero Resistance Test

Interfering voltages are often difficult to detect because they are not

always easily distinguished from actual specimen voltages. For example, the

interfering voltages can have the character of a current- transfer voltage-^

or even a flux- flow voltage. This is particularly true of the hightransition- temperature (high-Tc ) superconductors because their V-I

characteristics are not as well understood as those of the conventional, or

low-Tc , materials. The zero resistance test effectively simulates an ideal

superconductor where the V-I characteristic is V(I) - 0.

Ideally, the test is carried out with the measurement system configured

exactly as it would be for an I c measurement, with one exception: the

voltage tap leads are not both connected to the superconductor specimen.

Rather, one of the leads is connected to the specimen and the second lead is

connected to the first lead close to, but not in direct contact with, the

specimen. This forms a null voltage tap pair (see Fig. 1). There are

situations where the null voltage tap pair should have enhanced inductance

to simulate the inductance of the differential voltage tap pair. 1^ In this

configuration the measured voltage should be equal to zero regardless of the

current carried by the specimen. Any voltage that is detected, as the

current is cycled, is an interfering voltage. The important point is that

all of the electrical current paths that are present for an I c measurement

33
Fig. 1. Illustration of sample and lead

configuration for zero and finite

resistance tests.

are preserved in this configuration and the common mode voltage applied to

the input terminals of the voltmeter is also the same. The only difference

is that the differential voltage has been eliminated to allow easy

recognition of any interfering voltages.

If the current is increased from zero to some maximum value and then

decreased back to zero with a linear current ramp, equal and opposite

inductive voltages will be generated for increasing and decreasing current.

These voltages are easily recognized because they are of constant magnitude.

In the case of a continuous data acquisition system, where the specimen

voltage is recorded while the current is being continuously increased, the

inductive voltages can be reduced by tightly twisting the voltage leads, but

they cannot be eliminated. However, the inductive voltage is just a dc

offset and, thus, it does not affect the determination of the I c

. For

discrete measurement systems, where the voltage data are acquired with the

current held constant at selected set points, inductive voltages are not a factor. Like inductive voltages, thermal electric voltages are always

present and, if they are held constant during the data acquisition cycle, do

not affect the I c measurement.

The interfering voltages that present a problem for I c measurements are

those that change with changing current. These are the ground loop and

common mode voltages. The zero resistance test is effective for detecting

the presence of these voltages.

A more convenient, but less definitive, zero resistance test can be

made at room temperature using a copper conductor, as in the finite

resistance test. Once again, the voltage tap leads must be connected

together and then connected to a single tap to eliminate the differential

voltage. It is still important to retain as much of the actual I c measurement system as possible. For example, all instruments that are used

34
during I c measurements (chart recorders, computers, and so on) should be

connected for the zero resistance test. These peripheral devices often have

ground connections and are sometimes the source of interfering voltages.

For some measurement systems, the test fixture and specimen are not

electrically isolated from ground. When this is the case, the room

temperature test should be made with the copper conductor connected to

ground through an appropriate resistance.

Superconductor Simulator

A simple electronic circuit was designed to simulate the intrinsic V-I

characteristic of a superconductor. This circuit was used to characterize

the response of a nanovoltmeter when subjected to a highly asymmetric

periodic voltage that results from passing a dc-biased ac current through a superconductor.^ A more general application of this circuit has been to aid

in the development and testing of critical-current data acquisition systems.

Another important parameter in the determination of I c is the

measurement of the n value. The parameter n is defined by the approximate

intrinsic voltage -current (V-I) relationship,

V - V0 (I/I 0 ) n

,

where I 0 is a reference I c at a voltage criterion V0 , V is the sample

voltage, I is the sample current, and n reflects the shape of the curve with

typical values from 20 to 60. A higher number means a sharper transition.

In the measurement of the V-I characteristic, the sensitivity of the

voltmeter is the key factor. Voltage accuracy is less significant in the

determination of the I c for a sample with a high n value. For example, with

n - 30, a voltage error of 10% translates into a 0.3% current error.

The details of the circuit design are given in Ref. 4. The input to

the circuit comes from a shunt resistor connected in series with the current

supply. The output current from the simulator passes through two shunt

resistors, a "high output" and "low output." Typically, the nanovoltmeter

being tested is connected to the low output resistor, which generates a signal in the microvolt range, and a recording instrument is connected to

the high output resistor, which generates a signal that is 10^ times as large as the low output signal. Another channel of the recording instrument

is connected to the analog output of the nanovoltmeter. These two channels

are then compared and the measurement system may be thus characterized under

conditions similar to an I c measurement. The simulated values of the I c and

n can be adjusted.

The simulator does not reproduce all of the possible elements of an

actual superconductor's V-I characteristic. Current- transfer voltages and

the complex voltage patterns associated with flux dynamics, for example, are

not produced by the circuit. However, the simulator does produce the

predominant element of a superconductor's V-I characteristic, its abrupt

increase in voltage with increasing current. It also produces a very low

and well defined output voltage. These two capabilities make the simulator

very useful for the development and trouble shooting of I c measurement

systems

EXAMPLES OF TEST RESULTS

Measurement System Details

The copper specimen used for this study was cylindrical and measured

14.6 cm in length and 8.9 cm in diameter, and it had a voltage tap

35
separation of about 0.1 cm. All of the tests made using this specimen were

conducted at room temperature. Two different I c measurement systems were

used for these tests. The details of these measurement systems are given

elsewhere. 5 For both the finite resistance test and the zero resistance

test, the current was steadily increased from zero to a maximum level and

then steadily reduced to zero while the voltage was recorded with a digital

processing oscilloscope.

Finite Resistance Measurements

Figure 2 shows the results of a finite resistance test. The hysteresis

in the data is due to inductive voltage. The lower portion of the

hysteresis loop is for increasing current and the upper portion is for

decreasing current. The measured resistance is approximately 1.4 nQ. Based

on the diameter of the copper conductor (8.9 cm) and the voltage tap

separation (0.1 cm), this implies a copper resistivity of 0.9 ^H-cra.

The actual resistivity of the copper is probably closer to 1.7 ^Q-cm.

The discrepancy is probably due, primarily, to the lack of precision in the

measurement of the voltage tap separation. To allow space for soldering,

the voltage taps are staggered around the circumference of the conductor.

Given the relatively small tap separation and the finite size of the taps,

an accurate measurement of the longitudinal separation is difficult.

-Another source of the discrepancy is nonuniform current distribution within

Fig. 2. Finite resistance test, voltage

versus current for a copper sample.

36
the conductor. For an accurate resistivity measurement, the separation

between either current contact and its adjacent voltage tap should be at

least five times the diameter of the conductor. This allows uniform current

distribution in the region of the voltage taps. For the conductor used in

these tests, the separation between the current leads and voltage tap leads

is less than the diameter of the conductor. Ideally, a longer conductor

with a greater tap separation would be used for this test. Nonetheless,

this test demonstrates the sensitivity of the measurement and, to the extent

that the measured and actual values of the copper's resistivity are in the

same range, the accuracy of the measurement is demonstrated.

Zero Resistance Measurements

The results of a zero resistance measurement are shown in Fig. 3. This

test was made using the same measurement system that was used for the finite

resistance test. The voltage scale for this plot is nanovolts. Again the

hysteresis is caused by inductive voltage. The continuous curve (upper

portion of the loop) is for increasing current and the discrete data are for

decreasing current. The important point is that, if the inductive voltage

is subtracted from the data, the measured dc voltage is essentially equal to

zero. This measurement system has a voltage noise of about ±5 nV and a voltage measurement uncertainty of about ±2 nV ±2% of the signal.

Current, A

Fig. 3. Successful zero resistance test,

voltage versus current for a copper

sample

37
In contrast, Fig. 4 shows the results of a zero resistance test for

another measurement system. In this case, the voltage scale is in

microvolts and, even if the inductive voltage is subtracted from the data,

the measured voltage is not equal to zero. This is an example of an

interfering voltage. In an actual I c measurement, the interfering voltage

would be detected along with any differential voltage, thus altering the I c measurement. In fact, the abrupt increase in voltage that occurs at

approximately 900 A might be mistaken for a superconducting transition. The

features of the increasing current (continuous curve) are reproduced for the

decreasing current (discrete points).

DISCUSSION

Other examples of integrity tests can be found in Ref. 5, where various

combinations of voltmeters, power supplies, and load grounding conditions

are given. These combinations can change the results significantly; a voltmeter that works well with one current supply may not work with another.

In general, if the load can be grounded near the test sample, the level of

interfering voltages can be reduced. The resistance of the voltage tap

leads can also be a factor; the higher the lead resistance is, the larger

the interfering voltage.

Current, A

Fig. 4. Unsuccessful zero resistance test,

voltage versus current for a copper

sample

38
The required cross sectional area of the copper test specimen used in

the finite resistance test depends on the I c of the superconductor, the

selected I c criterion, and the test temperature (room or cryogenic) . For

high-current systems the required size of the copper specimen can become

impractical. However, low-current systems may require only a copper test

specimen that is comparable in cross sectional area to that of the

superconductor. The iength of the copper specimen is also important to

ensure an accurate measurement of its resistivity. It should be long

enough, in comparison with its cross sectional area, to ensure uniform

current distribution in the area of the voltage taps.

CONCLUSIONS

A set of simple procedures that will test the integrity of measurement

systems used for critical-current determinations on high-Tc and conventional

superconductors has been developed. These tests include a finite

resistance, a zero resistance, and a superconductor voltage-current

simulator. In the measurement of the critical current, voltage sensitivity

is a key factor. The zero resistance test is the most effective test to

detect the presence of interfering voltages such as ground loop or commonmode voltages and will determines the voltage sensitivity limit of a measurement system.

ACKNOWLEDGMENTS

The authors extend their thanks to W. P. Dube' for the idea of a zero

resistance test, and to R. L. Spomer for drafting.

This work was supported by the Department of Energy, Office of Fusion

Energy and Division of High Energy Physics.

REFERENCES

1. L. F. Goodrich and F. R. Fickett, Critical current measurements: a compendium of experimental results Cryogenics 22:225 (1982).

2. Ralph Morrison, "Grounding and Shielding Techniques in Instrumentation,"

John Wiley & Sons, New York (1977).

3. J. W. Ekin, Current transfer in multifilamentary superconductors. I.

theory J^_ Annl . Phvs

49:3406 (1978).

4. L. F. Goodrich and S. L. Bray, Current ripple effect on superconductive

d.c. critical current measurements Cryogenics 28:737 (1988).

5. L. F. Goodrich, S. L. Bray, W. P. Dube', E. S. Pittman, and A. F. Clark,

"Development of Standards for Superconductors, Interim Report

January- December 1985," NBSIR 87-3066, National Bureau of Standards,

Boulder, Colorado (1987).

39
gBBmagi
APPENDIX C

BATTERY-POWERED CURRENT SUPPLY FOR SUPERCONDUCTOR MEASUREMENTS

S. L. Bray, L. F. Goodrich, and W. P. Dube

published in: Review of Scientific Instruments

Volume 60(2), pp. 261-264, February 1989

41
Battery-powered current supply for superconductor measurements

S. L. Bray, L.F. Goodrich, and W. P. Dube

Centerfor Electronics and Electrical Engineering, National Bureau ofStandards, Boulder, Colorado 80303

(Received 16May 1988; accepted for publication 20 October 1988)

To measure the critical current of superconductors, a high output current supply is required. In

addition to high current capability, the supply should be designed to reduce ground loop

problems, respond linearly to an input control signal, and minimize output noise. A current supply with these qualifications has been constructed and tested. Although the supply was

originally designed for testing conventional superconductors at high current levels, it has also

been successfully used in measurements on the high-critical-temperature ceramic

superconductors where the maximum current output was less than 1 A. The supply can produce

1000 A output current with a noise level of approximately 0.05 A peak-to-peak. The specifics of

the current supply’s design and performance are given.

INTRODUCTION

This current supply is an experimental design that has been

gradually developed through continuous modification and

improvement for an explicit application, critical-current

measurements. These are extremely low-voltage measure- ments and, as such, they present special measurement problems. Specifically, the critical-current measurement requires

a determination of the current level where the resistance of a superconducting sample becomes nonzero. In other words, a

transition from zero voltage to finite voltage must be measured. 1 Consequently, small spurious voltages (such as those

resulting from ground loops2 ' 3 ), that might in some measurements be negligible, can become dominant in critical- current measurements. Also, current noise can result in a systematic error in the measured critical current or in completely random measurement results, depending on the noise

level.

4 Finally, the inherent thermoelectric instability of a superconductor necessitates precise and stable current control. This instability also necessitates a sample protection

device when operating at high current densities.

5 These concerns are reflected in the design of this current supply; however, specifics of the design should be considered quite flexi- ble depending on the particular application. Although the

only use for which this supply has been tested is superconductor measurement, the general elements of its design may

be useful in many other applications.

I. CIRCUIT DESIGN AND PERFORMANCE

A. Input, control, and power stages

The basic circuit is a voltage-controlled current source

(see Fig. 1) consisting of three main stages: input, control,

and power. The principal power source for the supply is a 4- V battery consisting of two 1 500-A-h, 2-V wet cells. This 4-V

battery supplies the output current through the power stage

of the supply. Also, two 6-V wet cell batteries are used to power the control stage, and a 1 1 5-V ac-powered/ + 1 5-V

dc-output supply is used at the input stage of the supply

circuit. The power stage of the circuit uses six Darlington

transistor pairs in parallel, with each pair consisting of a 25-

A (T3) and a 300-A (T4) power transistor. Each of the 300-

A transistor emitters is connected to the positive bus bar

261 Rev. Sci. Instrum. 60 (2), February 1989

through a separate 1-mfl resistor. The resistors provide the

required current-proportional feedback signals for the control stage. The collectors of these transistors are all connected to the 4-V battery through a single l-mfl current limiting

resistor. The collectors of the 25-A transistors are all connected directly to the 4-V battery and the bases are each

connected to a series combination of a 0.3-mH inductor and

a 1-fl resistor. These inductors and resistors are beneficial in

stabilizing the circuit with respect to parasitic oscillation.

The opposite ends of the inductors are connected together,

forming a branch point. The emitter of another 25-A (T2)

transistor is connected to this branch point and provides the

current to drive the Darlington pairs. This transistor amplifies the output signal from the control circuit. A precision

0. 1-mfl resistor connected between the negative battery ter- minal and the negative bus bar provides a current-proportional voltage signal for measuring the total current output

of the supply.

The control stage of the circuit is based on an operational amplifier (OA2) that compares the signal from the input

stage with the feedback signal from the Darlington emitter

resistors and applies the resulting signal to the power stage. In the feedback path, another operational amplifier (OA3)

is used as a summing amplifier in order to combine and amplify the six signals from the feedback resistors. A trim potentiometer is used on this operational amplifier to allow

adjustment of the control circuits bias point. Without this capability, one of two undesirable conditions may exist, depending on the offset voltages of the operational amplifiers

in the control loop. These offset voltages may result in an

offset current at zero input voltage, or they may result in excessive deadband (where a significant threshold input vol- tage is required to produce an output current ) , depending on

the size and polarity of the offset voltages. The trim potentiometer can be used to correct either of these conditions. The

potentiometer is adjusted so that a small deadband region

exists at turn on. This adjustment ensures that there will be no offset current and that the deadband region is not excessive. In addition to providing a small margin of safety with

respect to offset current, the residual deadband also masks

any initial turn-on nonlinearity from the optically coupled

input stage. In other words, the deadband prevents the sup261

42
Isolation Boundary

Fig. 1. Circuit diagram of the batterypowered current supply.

ply from responding to the input signal until the optical iso- lator is in its linear region. Since the summing amplifier in- verts the feedback signal, another inverting amplifier ( OA4

)

is required in the feedback path to restore the polarity of the

signal. The resulting signal is then sent to the inverting input

of the comparator. The comparator uses a feedback capacitor to reduce the ac gain and, thus, stabilize the circuit. The main function of the input stage is to provide electrical isolation between the input signal generator and the

supply circuit in order to avoid ground loop problems. 3 This

isolation is achieved through the use of an optical isolator, or optocoupler (OC1 ). The input stage is linearized in a manner similar to that of the control stage. Again, a comparator

(OA1) is used to compare the reference and feedback sig- nals. The output signal from the comparator is amplified by

a single transistor ( T 1 ) . The feedback signal is supplied by a second optocoupler (OC2), which is in series with the first. This is necessary to provide electrical isolation. Moreover,

the second optocoupler imitates the first and, thus, generates an appropriate feedback signal without electrically coupling

the input and control stages. It is important to note that

separate power supplies are used on each side of the isolation

boundary (vertical dashed line on circuit diagram). Again,

this avoids electrical coupling of the stages. The feed-for- ward optocoupler ( OC 1 ) was found to be a significant noise source because of an unused base terminal which was picking up 60-Hz noise. This problem was greatly reduced by

262 Rev. Sci. Instrum., Vol. 60, No. 2, February 1989

connecting the base terminal to the emitter terminal with a

0. l-/iF capacitor. This treatment of the feedback optocoupler (OC2) resulted in oscillation; consequently, the ca- pacitor was omitted on this optocoupler.

B. Supply protection

The battery powered current supply was not designed

for continuous operation at its maximum current level. Its duty cycle is limited by the maximum power dissipation of

the 300-A output transistors and the 1-mft collector resistor

of Fig. 1. Consequently, a method of overcurrent protection

that uses thermal cutoff switches on both the collector resis- tor and the output transistors has been designed. For redundancy, two of the output transistors are monitored with cutoff switches, and two cutoff switches are used on the

collector resistor. The four cutoff switches are placed in series with the coil of a 12-V relay and with the two 6-V batteries. Under normal operating conditions, the thermal cutoff

switches are in the closed position. This, in turn, holds the

12-V relay in a position that connects the Darlington output

stage to the control signal amplifier. T2. If the temperature

of any of the four thermal cutoffs exceeds its set point, the

relay current is interrupted and the relay switches the Darlington input point from T2 to common, which results in an

interruption of the output current. After sufficient cooling,

the thermal cutoff switches may be manually reset. The max- imum current output of the supply depends on many factors

Superconductor measurements 262

43
including load resistance, duty cycle, and the available cooling for the output transistors. Consequently, the output capability of the supply and the specifics of the protection sys- tem will vary from one application to another.

C. Performance

The observed performance of the circuit with respect to output capability, linearity, current drift, and noise are described. The maximum current output of the supply in this particular application is approximately 1000 A. The linearity of the circuits’ response was measured by comparing the current supply output with the control signal input. In this case, the input signal was a linear ramp. The results of this measurement are shown in Fig. 2 where the difference between the actual output current and the ideal linear current (DELTA I) is plotted versus the dc output current (CURRENT). This plot was constructed by generating a linear fit equation from the curve of input-voltage versus output-current in its most linear region. The ideal linear current was

calculated from this equation and compared with the measured current over the entire input-voltage range. This plot

shows that the difference between the ideal and actual currents is less than 1 A for current output levels between 50-

400 A. At the maximum* output of 1000 A the difference is approximately 30 A. Linearity measurements were made

under a number of different conditions and these data are

typical. Although this plot has been scaled to accentuate the

deviation from linearity, the maximum nonlinearity represents only 3% of the measured output current. For this application of the current supply, this level of

nonlinearity is acceptable. However, for an application

where a higher level of linearity is required, several elements

of the present design should be considered. First, linearity of

the input stage depends on a constant ratio of the large-signal current gains of the two optocouplers (OC1 and OC2).

Careful matching of these two optocouplers can enhance the

l ie;. 2. Plot of the difference between the ;ict ual output current and the ideal linear current vs the dc output current.

263 Rev. Sci. Irstrum., Vcl. 60, No. 2, February 1989

linearity of the supply’s response. Also, the comparator of

the control circuit (OA2) results in a nonlinearity due to its closed-loop configuration. Moreover, the comparator is configured as an inverting amplifier in order to limit the gain of

the offset-trim signal; however, this results in an impedance

reduction at the inverting input. As a result, loading of the

feedback signal by the 10-kfl input resistor can reduce the

supply’s linearity. A reduction of this resistance can reduce

this nonlinearity but an increase in the offset-trim level will also result. Finally, because the feedback signal is a sum of

the signals generated by six different resistors and Darlington pairs, linearity depends on perfect matching of the resis- tors or consistent current sharing among the transistors. The

former is more easily achieved; thus, both the 1-mfl emitter

resistors and the 6.2-kfl summing resistors should be closely

matched to enhance linearity.

The current drift of the supply was measured at two

nominal current levels, 50 and 200 A. In both cases, the

input signal was increased until the nominal current output was reached and then held constant until the end of the test. The input signal and the output current were continuously

recorded during the test, which had a total duration of 30

min. For the 50-A test the current output increased by

0.30% after 1 min and by 2.1% after 30 min. At 200 A, the

current increased by 0.18% after 1 min and by 0.78% after 30 min. Several tests were conducted w’ith different load re- sistances, and these results are typical of all the tests. Also,

drift measurements were made under varying initial temperature conditions. These tests suggest that the current

drift is thermally induced. For example, when the supply is at ambient temperature at the beginning of the test, the current drifts upward with time regardless of the nominal current output. However, if a 200-A test is conducted immediately before a 50-A test, the current drifts downward during

the 50-A test as the supply temperature decreases. As evi- denced by these data, a disproportionate amount of the total current drift occurs in the first minute of the test. Drift measurements that isolated the input stage’s contribution to the

total current drift were made. These tests showed that the

initial current drift was due primarily to the input stage. This

condition might be improved by closer matching of the optocouplers, and it could certainly be improved by omitting the

optically coupled input stage from the design in applications

where ground loops are not a problem.

The supply’s performance, with respect to output noise, was tested with two load resistances. In the first case, a lowinductance, 50-mfl load (twisted wire) was used. This relatively large load resistance limited the output current to 20

A, but it also resulted in a sufficiently large current-proportional voltage signal for accurate measurement of the current ripple. The peak-to-peak current ripple was 0.01 A at 20-A output current. In the second case, a load resistance on

the order of 1 mfl was used to allow the supply to be operated at high currents. The voltage signal from a 0. 1-mfl resis- tor was used to measure the current ripple. For this case, the

measured peak-to-peak current ripple was about 0.05 A for output currents from 10 to 900 A. The actual current ripple

may be less than this, but, because of residual voltage amplifier noise, this was the limit of this measurement.

Superconductor measurements 263

44
II. DISCUSSION

Although this design has proven very successful in this particular application, many of its elements can be changed

or even deleted in order to better suit a different application.

For example, in some critical-current measurement systems

ground loops are not a problem and, consequently, the optically coupled input stage might be omitted in order to in- crease the circuits linearity while decreasing its complexity.

Reference 2 gives specific test methods for evaluating criti- cal-current measurement systems. For critical-current measurements the resistance of the current supply’s load ( a su- perconductor and connecting cables) is extremely low, thus

allowing the relatively low voltage of the main battery. This

low battery voltage would limit the current output for a re- sistive load. In this case, a higher voltage battery would be

required in order to maintain the supply’s output capability.

A disadvantage of a battery-powered current supply is that the batteries require periodic recharging. Furthermore,

if the system is used on a daily basis at high output levels,

regular overnight recharging is required. These conditions

lead to the danger of overcharging. Consequently, the power

supply that is used for battery charging is outfitted with two

adjustable l-to-10-h timers. Because the supply is usually

unattended during recharging, two timers are used to provide redundancy. In this system both timers must malfunction in order for the batteries to continue charging beyond

the selected time. Originally, lead-antimony batteries were

used in this supply, but they have since been replaced with

lead-calcium batteries. The calcium batteries do not generate as much hydrogen and oxygen gas when charging and, as

a consequence, they do not require the maintenance or pres- ent the fire hazard that conventional batteries do. However,

when overcharged, the lead-calcium batteries exhibit elec- trolyte swelling, which presents the danger of acid overflow.

Generally speaking, low-level measurement systems re- quire low-noise power supplies. This is the main benefit of a battery-powered supply as compared to conventional high

current supplies. Also, a battery-powered supply offers very

high ground isolation, which is crucial in some systems. For

any application, the supply’s relatively simple design and

versatility are important benefits. In many cases, the benefits

of a battery-powered supply outweigh its minor disadvantages.

264 Rev. Sci. Instrum., Vol. 60, No. 2, February 1989

ACKNOWLEDGMENTS

The authors extend their thanks to E. S. Pittman for

assistance with the linearized optocouplers, to R. M. Folsom

for creating circuit diagrams, and to H. E. Spilman for cir- cuit testing. This research was supported in part by the U.S.

Department of Energy through the Office of Fusion Energy

and the Division of High Energy Physics.

APPENDIX: COMPONENT DESCRIPTIONS

An effort was made to avoid the identification of commercial products by the manufacturer’s name or number,

but in some cases these products might be indirectly identified. In no instance does this identification imply endorsement by the National Bureau of Standards, nor does it imply

that the particular products are necessarily the best available

for that purpose.

All of the resistors used in this circuit are 1/8 W, 1%,

metal-film resistors, except the 1 and 0.1 mfl resistors,

which are high current and low-temperature-coefficient resistors. All the capacitors are ceramic. The inductors are made from No. 18 insulated wire, wound on ferrite cores. The 4 V battery is two 2 V, 1500 A-h, lead-calcium wet cells wired in series. The 6 V batteries are 8 A-h rechargeable wet

cells. The 15 V supply is a 115 V ac input, 25 mA output,

commercial supply. OA1 is a number 3140 operational amplifier; OA2 and OA4 are number 741 operational amplifiers; and OA3 is a number OP5 operational amplifier. The

optical isolators, OC1 and OC2, are both number 4N25. The

transistor T1 is a 2N3904; T2 and T3 are TIP35 transistors.

The output transistors, T4, are 500 A/300 A (peak collector

current/dc collector current), 625 W/400 W

(25 °C/100 6C) power transistors.

'L. F. Goodrich and F. R. Fickett, Cryogenics 22. 225 (1982).

2L. F. Goodrich et al., National Bureau of Standards Internal Report 87- 3066, 1987. ’Ralph Morrison, Grounding and Shielding Techniques in Instrumentation

(Wiley, New York, 1977).

4 L. F. Goodrich, S. L. Bray, and A. F. Clark, Adv. Cryog. Eng. 34, 1019

( 1988).

5 W. P. Dube and L. F. Goodrich, Rev. Sci. Instrum. 57, 680 ( 1986).

Superconductor measurements 264

45

APPENDIX D

Nb 3 Sn CRITICAL- CURRENT MEASUREMENTS USING TUBULAR FIBERGLASS -EPOXY MANDRELS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

IEEE Trans, on Magnetics 25.(2) 2375-2378 (March 1989)

47
IEEE TRANSACTIONS ON MAGNETICS, VOL. 25, NO. 2, MARCH 1989 2375

Nb 3 Sn CRITICAL- CURRENT MEASUREMENTS USING TUBULAR FIBERGLASS - EPOXY MANDRELS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

National Bureau of Standards

Boulder, Colorado 80303

Abstract

A systematic study of the effect of sample mounting

techniques on the superconducting critical-current

measurement was made in conjunction with the VAMAS

(Versailles Agreement on Advanced Materials and

Standards) interlaboratory comparison (round robin)

measurements. A seemingly small change in mandrel

geometry can result in a 40% change in the measured

critical current of a Nb 3 Sn sample at 12 T. This is a result of a change in the conductor pre-strain at 4 K

caused by variation in thermal contraction between

thick- and thin-walled fiberglass-epoxy composite

(G-10) tubes. An approximate measure of the variations

in thermal contraction (from room to liquid nitrogen

temperature) indicate a 0.2% greater contraction for

the thick-walled tube. This difference, combined with

strain sensitivity measurements, is consistent with the

observed decrease in critical current. Previous

publications on the thermal contraction of G-10 have

addressed the plate geometry, but not the tube

geometry. The contraction of a G-10 plate is highly

anisotropic. The radial contraction of a tube is

different than the contraction of a plate, however,

because the circumferential fiberglass is put into hoop

compression by the epoxy, and the resulting contraction

is a competition between the two structural components.

This appears to be the source of the variation in

thermal contraction with tube wall thickness.

Introduction

The measurements reported here are part of a collaborative effort involving twenty- four laboratories

from the European Economic Community, Japan, and the

USA under the Versailles Agreement on Advanced

Materials and Standards (VAMAS). The purpose is to develop critical-current (I c ) measurement techniques

for Nb 3 Sn superconductors that will yield consistent

results at different laboratories. Comparisons of

preliminary results showed a large variation in the

measured I c

. In response, the NBS study was expanded

to address specific measurement variables that might

have caused these discrepancies.

The predominant variables that affect I c measurements of Nb 3 Sn wires^’^ fall into four categories:

sample reaction conditions, measurement mandrel (sample

holder) material and geometry, the method used for

bonding the sample to its mandrel, and damage incurred

during shipping or in transferring the sample from the

reaction mandrel (tubular stainless steel sample

holder) to the measurement mandrel. In the last three

categories, the sensitivity of Nb 3 Sn to mechanical

strain results in variations in the measured I c

.^ For

these measurements, similar samples were reacted at a central location under (presumably) the same conditions; consequently, the reaction variables were

not suspected of being a significant factor. The

measurement mandrel material and geometry can work in

conjunction with the sample bonding technique to create

a dominant measurement variable: the axial strain of

the conductor. Under a condition where (1) the circumferential thermal contraction of the measurement

mandrel is greater than the axial thermal contraction

of the conductor and (2) the bonding method creates a strong coupling between the sample and the measurement

mandrel, significant axial strain of the conductor at

4 K may result. This study discovered that the

circumferential thermal contraction of the G-10

fiberglass -epoxy tubes used for the measurement

mandrels depends on the tube's geometry, owing to its

composite structure. In particular, the ratio of the

tube wall thickness to its radius strongly affects its

thermal contraction. Based on these observations, a systematic study was made of the effect of G-10

measurement mandrel geometry on the measured I c

.

For the measurement mandrel study, a glass -filled

-epoxy adhesive was used for bonding the sample to the

mandrel. This technique insured the strong bond and

rigid coupling between the sample and mandrel that is

necessary for positive stress transmission.

Experimental Details

The Nb 3 Sn wire measured in this study had a diameter

of 0.68 mm with 37 sub-bundles, each contained 150 Nb

filaments. The conductor fabrication method was an internal- tin diffusion process. A single diffusion

barrier of Ta separated the filament region from the

outer layer of Cu. The sample reaction temperature was

- 700°C for 48 h.

The measurement mandrels were constructed from G-10

tubes that had a cylindrical copper current contact

rigidly attached to each end. Each contact had a superconductive bus bar (Cu and Nb 3 Sn tape) connected

to it. The sample was transferred from the reaction

mandrel onto the measurement mandrel by carefully

unthreading the sample from the reaction mandrel and

slipping it onto the ungrooved measurement mandrel.

The one current contact that was removed for this

transfer was then reattached to the measurement mandrel

and the sample was tightened onto the mandrel. Each

end of the sample was then soldered onto its current

contact. To avoid persistent currents, an effort was

made to not have a complete superconductive loop formed

by the sample and bus bar around the current contact.

Three pairs of adjacent voltage taps were placed along

the center of the sample. Each pair had a separation

of about 10 cm and there was a 1 to 2 cm gap between

adjacent pairs. The sample was bonded to the mandrel

with glass -filled epoxy and allowed to cure. The epoxy

coat was kept thin to enhance the heat transfer from

the sample to the helium bath. This unit was then

mounted in the test fixture where the electrical

connections were completed.

The I c measurements^ were made with forward

(antiparallel magnetic fields, Lorentz force in) and

reverse (parallel magnetic fields, Lorentz force out)

current directions to determine the effect of selffield and sample- to -mandrel bonding (strain effect).

Unless otherwise stated, the I c values given are averages over all taps, observations, and current

directions at an electric field strength criterion of

10 ^V/m. The precision of the I c measurement was about

±0.2%. A statement regarding the accuracy of these

measurements is presently inappropriate because of the

strong systematic effects of sample mandrels.

Another parameter that was determined during the I c measurement is n. The parameter n is defined by the

approximate voltage -current relationship,

Manuscript received August 22, 1988. Contribution of NBS, not subject to copyright.

U.S. Government work not protected by U.S. copyright.

48
2376

v - Vq ( I / Iq ) n

, where I 0 is a reference I c at a voltage criterion VQ , V

is the sample voltage, I is the sample current, and n

reflects the shape of the curve with typical values

from 20 to 60. A higher number means a sharper

transition. A lower value of n can be an indication of

sample damage or strain.

Results

Critical Current

A summary of the I c data is given in Table I and in

Fig. 1. The information contained in Table I is the I c data, sample mandrel dimensions, and intercomparisons

of I c data. The I c data are given at magnetic fields

from 6 to 12 T. The effects of sample mandrel geometry

and thermal cycling were studied. The column headings

of Table I contain key words that relate to these

variables. Each column, which is labeled with a letter

from A to D, is a data set taken for a single thermal

cycle. Specimens are numbered 1 and 2. Different data

columns that are for the same specimens (columns A and

B for example) indicate a thermal cycle where the

specimen was measured at 4 K, warmed to room

temperature, again cooled to 4 K, and remeasured. In

some cases, the mounting configuration of the specimen

was changed between these measurements (columns B and C for example) . The sequence of the thermal cycles is

from left to right in the table.

The key words heading the columns of Table I ( thick . thin , and bored) are used to indicate the measurement

variables. Thick and thin refer to the ratio of wall

thickness to outer radius of the sample mandrel. Under

the Sample Mandrel Dimensions section of the table,

these specific values are listed for each data set. Bored refers to a specimen that was originally mounted

and measured on a thick walled mandrel. The specimen

was then brought to room temperature (thermally cycled)

and the mandrel inner diameter was bored to a larger

dimension, resulting in a reduced wall thickness. The

specimen's I c was then remeasured.

A comparison of the I c data in columns A and B is a measure of the I c variation due to thermal cycling.

The maximum variation was 0.4% at 10 T (the highest

magnetic field for column A) as shown in the intercomparison portion of the table ([A-BJ/B). This small

variation upon thermal cycling was representative of

all observations. Column C was a special case of a thin walled mandrel in that it was bored (machined with

the specimen mounted on the mandrel) . Because this was

the same specimen as A and B, the comparison is a systematic one. The intent of this procedure was to determine the effect of mandrel wall thickness on the

measured I c of a particular specimen. The intercomparison data ([B-C]/C) shows a strongly field-dependent

variation of I c (indicative of a strain effect) with a maximum difference of 40% at 12 T. This is thought to result from a change in the conductor pre- strain (at

4 K) due to variation in thermal contraction between

thick- and thin-walled G-10 tubes. However, machining

may have changed the mandrel's outer diameter at room

temperature, due to stress relief, and, thus, the prestrain state of the conductor. More details regarding

thermal contraction of G-10 measurement mandrels and

the resulting I c degradation due to axial strain are presented below. The comparison between C and D showed

only a small difference, 3 to 4%, with very little

field dependence ([C-D]/D). This indicates that the

effect of boring was about equivalent to that of simply

mounting a specimen on a thin walled tube.

Figure 1 is a semilogarithmic plot of I c versus magnetic field. This type of plot was used because

Table I. I c Data and Mandrel Dimensions

Critical current (amperes) at 10 pV/m

Specimen 1 1 1 2 thick thick thin

bored

thin

4 0 H, T (A) (B) (C) (D)

6 343.3 343.2 397.3 386.3

7 268.4 267.6 318.4 308.9

8 207.0 207.4 253.8 246.0

9 156.5 156.6 200.1 193.5

10 113.6 114.0 154.7 149.3

11 78.5 116.0 111.9

12 50.2 83.5 80.4

Sample Mandrel Dimensions

(A) (B) (C) (D)

Wall ,mm 9.3 9.3 2.5 2.0

Rad. ,mm 15.6 15.6 15.6 15.6

Wall/Rad 60% 60% 16% 13%

Intercomparison of Critical Current Data

m 0 h , T (A-B)/B (B-C)/C (C-D)/D

6 0.0% -13.6% 2.8%

7 0.3% -16.0% 3.1%

8 -0.2% -18.3% 3.2%

9 -0.1% -21.7% 3.4%

10 -0.4% -26.3% 3.6%

11 -32.3% 3.7%

12 -39.9% 3.9%

it illustrates the percentage differences between the curves. The thick curve is significantly lower than

the others. Figure 2 is a semilogarithmic plot of n versus magnetic field. The significant element of this

plot is the lower values of n for the thick curve at the higher magnetic fields. This is consistent with a strain degradation.

Thermal Contraction

Thermal contraction measurements on NEMA (National

Electrical Manufacturers' Association) G-10 glass-epoxy

composites have been published for the plate

geometry^’ ^ but not for the tube geometry. The

contraction of a G-10 plate is highly anisotropic. In

the nomenclature of composites, the two directions in

the plane of the fiberglass fabric are warp and fill.

The direction perpendicular to the plane of the fabric

is referred to as the normal direction. The thermal

contraction from 293 K to 4 K is about 0.24% for the

warp direction and 0.71% for the normal direction. The

contraction in the fill direction is expected to be a little more than that of the warp direction. The

contraction in the warp direction is dominated by the

fiberglass fabric and the normal direction is dominated

by the epoxy. The G-10 tubes used for measurement

mandrels were rolled spirals of fiberglass fabric

embedded in an epoxy matrix, rather than tubes machined

from plate material. The radial direction for a rolled

tube is normal to the fabric. The radial contraction

of a tube is different than the contraction of a plate,

however, because the circumferential fiberglass is put

into hoop compression by the epoxy, and the resulting

contraction is a competition between the two structural

components. The dependence of the radial contraction

on wall thickness is expected to be caused by this

competition.

49
n

Value

2377

Magnetic Field, T

Figure 1. A semi logarithmic plot of I c at an electric

field criterion of 10 /^V/m versus magnetic field.

Magnetic Field, T

Figure 2. A semilogarithmic plot of n value versus

magnetic field.

A thick-walled G-10 tube was originally selected as

a common sample mandrel material to allow adaptation of

one size tube to the various test fixtures used by the

different laboratories. However, a thick-walled tube

is difficult to manufacture without de laminations that can cause irregular contraction and voids when they are machined. Also, the outer diameter can change when the

inner diameter is bored. Another observation was that

the radial contraction can be slightly asymmetric; this

will result in an approximately elliptical rather than

circular cross section at low temperatures. In order

to make an approximate correction for this, the thermal

contraction measurements are the average of two approximately orthogonal measurements of the tube's

diameter

As briefly mentioned above, thermal contraction

measurements were made on several G-10 tubes in order

to estimate the resulting conductor strain. In the

’interest of expediency, these measurements were limited

both in accuracy and precision; nonetheless, the

measurements give an indication of the variation in

thermal contraction between the tubes. These tests

consisted of measuring the outside diameter of the G-10

tube at room temperature, submerging the tube in liquid

nitrogen, allowing it to reach thermal equilibrium, and

then quickly removing it and remeasuring the diameter.

These measurements were made with a precision micrometer, but they were limited by the practical

difficulties of the measurement.

The measurements showed a substantial difference

between the radial thermal contraction of G-10 tubes

with different wall thicknesses. The results of the

thermal contraction measurements are shown in Table II. As indicated by the column headings in the right hand

portion of the table, four different samples were

measured. Of the four samples, the geometry of sample

1 was most like that of the actual measurement

mandrels. This sample had a stepped cross-sectional

geometry characterized by a longitudinal variation in

its inner and outer diameters. The outside diameter of

the sample was reduced at its ends and its inside

diameter was enlarged at its center. This stepped

geometry was necessitated by the cylindrical current

contacts that were flush mounted over the ends of the

mandrel. Reducing the outside diameter allowed the

contacts to slip over the mandrel's ends; enlarging

only the central portion of the mandrel's inside

diameter allowed a reduction in the mandrel's wall

thickness only in the area of the sample, while

maintaining the required support in the area of the

current contacts. The other three samples had constant

inside and outside diameters along their lengths.

Table II. Thermal Contraction of G-10 Tubes

Mandrel Dimensions Outer Diameter

Thermal Contraction

Ratio Wall Sample 1 Sample 2 Wall/Radius Thickness Ad/d Ad/d

59% 9.27 mm 0.39% 0.41%

49% 7.70 mm 0.38% 0.36%

39% 6.10 mm 0.33% 0.32%

28% 4.32 mm 0.30% 0.28%

16% 2.54 mm 0.25% 0.23%

13% 2.03 mm 0.25% 0.24%

Ratio Wall Sample 3 Sample 4 Wall/Radius Thickness Ad/d Ad/d

20% 3.18 mm 0.25%

30% 4.76 mm 0.28%

50
2378

Samples 1 and 2 were progressively bored to larger

inside diameters while samples 3 and 4 had factory

diameters (not machined) and were measured to test the

equivalence between factory tubes and bored tubes with

respect to thermal contraction.

For each sample, the table shows the tube's wall

thickness, the ratio of the tube's wall thickness to

its radius in percent, and the thermal contraction of

the tube between room and liquid nitrogen temperatures

(Ad/d). The data indicate that the thicker walled

tubes contract more than the thinner ones. In

addition, the good agreement between the data for

samples 1 and 2 indicates that the contraction of the

actual measurement mandrels is similar to that of

simple tubes. The agreement of the data for samples 3 and 4 with those of 1 and 2 indicates that the bored

tubes behave in a manner similar to that of factory

tubes. Samples 1 and 2 with wall/radius ratios of 59%,

16%, and 13% have respectively similar dimensions to the thick , thin bored , and thin measurement mandrels of

Table I.

A bias in this study exists as a result of the

thermally transient measurement conditions (with a systematic difference in thermal mass) and the fact

that it was conducted at liquid nitrogen, rather than

liquid helium, temperature. Both of these factors

result in an underestimate of the actual thermal

contraction between room and liquid helium temperatures. Based on the temperature dependence of the

thermal contraction of G-10 plate ,

1 the thermal

contraction to 4 K was estimated as -0.28% for the thin

walled mandrel and -0.48% for the thick walled mandrel.

The thermal contraction of the Nb 3 Sn conductor was only

-0.21%. Consequently, the G-10 contraction may have

introduced additional amounts of conductor pre- strain.

The amount of compressive pre-strain might depend on several factors including the mandrel's wall thickness

and the strength of the mandrel- to-conductor bond.

Considerable I c tensile-strain-sensitivity data were

available for this sample; however, no explicit data

were available for compressive strain sensitivity.

Consequently, the strain scaling law,^ in conjunction

with the available tensile strain data, was used to estimate the expected I c degradation for compressive

strain. The results of these calculations predict a -33% reduction in the I c at 12 T for the thick-walled

specimens as compared to the thin-walled ones; the

observed reduction was -40%. These measurements and

calculations are only approximate, but the trend is

apparent. In all cases, the thick walled samples had

reduced I c

's and this reduction was in approximate

agreement with that predicted by the strain scaling

law.

Conclusions

The bonding method and the mandrel material and

geometry have a significant effect on the critical

current measurement of Nb 3 Sn. Thick-walled G-10 tubes

contract about 0.2% more than thin-walled G-10 from room to liquid helium temperature. Thus, the amount of

differential thermal contraction between the mandrel

and the sample depends on the wall thickness of the

G-10. The effect of strain on the I c of a particular

sample depends on its upper critical field. The sample

tested here has an upper critical field of -19 T and

the difference between critical current measurements on

thick- and thin-walled G-10 was about 40% (predicted

33%) at 12 T. Another sample with an upper critical

field of -24 T was also measured and, in this case, the

difference between critical current measurements on

thick- and thin-walled G-10 was about 17% (predicted

24%) at 12 T.

G-10 tubes that have diameters similar to that of

those tested, and wall thicknesses about 20% or less of

the tube's radius, are well matched in thermal contraction to Nb 3 Sn wires and use of this type of tube for

measurement mandrels should result in a minimum of

conductor strain. The scaling of this effect for tubes

of various diameters has not been demonstrated, but

scaling is probably accurate for a limited range of

tube diameters

Acknowledgments

The authors extend their thanks to J. W. Ekin for

discussions on these results; to R. M. Folsom for

sample preparation, data reduction, and plotting; to

D. L. Rule for help with data plotting; to W. E. Look

for assistance with the magnetic field calibration; and

to the other VAMAS participants.

This work was supported by the Department of Energy,

Office of Fusion Energy and Division of High Energy

Physics

An effort was made to avoid the identification of

commercial products by the manufacturer's name or

label, but in some cases these products might be

indirectly identified by their properties. In no instance does this identification imply endorsement by

the National Bureau of Standards, nor does it imply

that the products are necessarily the best available

for that purpose.

References

[1] L. F. Goodrich and F. R. Fickett, "Critical

Current Measurements: A Compendium of

Experimental Results," Cryogenics . vol. 22, pp.

225-241, May 1982.

[2] Standard Test Method for D-C Critical Current of

Composite Superconductors, Annual Book of ASTM

Standards, ASTM B714-82, Part 02.03, pp . 595-98,

American Society for Testing and Materials,

Philadelphia, PA (1983).

[3] J. W. Ekin, "Strain Scaling Law for Flux Pinning

in Practical Superconductors. Part 1: Basic

Relationship and Application to Nb 3 Sn

Conductors," Cryogenics . vol. 20, pp . 611-624,

Nov. 1980.

[4] L. F. Goodrich, "Development of Standards for

Superconductors, Interim Report January 1986 - December 1987," NBSIR 88-3088, National Bureau

of Standards, Boulder, CO (February 1988).

[5] G. Fujii, J. W. Ekin, R. Radebaugh, and A. F. Clark, "Effect of Thermal Contraction of Sample

Holder Material on Critical Current," Adv.

Crvog. Eng . -Materials . vol. 26, pp. 589-598,

Plenum Press, New York (1980).

[6] A. F. Clark, G. Fujii, and M. A. Ranney, "The

Thermal Expansion of Several Materials for

Superconducting Magnets," IEEE Trans . Mae .

MAG17, 2316, 1981.

51
.

fit
APPENDIX E

CRITICAL CURRENT MEASUREMENTS OF Nb 3 Sn SUPERCONDUCTORS:

NBS CONTRIBUTION TO THE VAMAS INTERLABORATORY COMPARISON

L. F. Goodrich and S. L. Bray

Published in Cryogenics

Volume 29, pp . 699-709, July 1989

53
Critical current measurements of NbaSn

superconductors: NBS contribution to the

VAMAS Interlaboratory Comparison*

L.F. Goodrich and S.L. Bray

Center for Electronics and Electrical Engineering, National Bureau of Standards, Boulder,

Colorado 80303, USA

Received 7 July 1988

Critical current measurements on several Nb3 Sn superconductors were made as part of an

interlaboratory comparison (round robin). These measurements were made in conjunction

with twenty-four laboratories from the European Economic Community, Japan and the USA

as part of the Versailles Agreement on Advanced Materials and Standards (VAMAS). The

results of the NBS measurements, including the effect of sample mounting techniques on

the measured critical current, are given. A systematic study of the effect of measurement

mandrel (tubular sample-holder made from G10 fibreglass-epoxy composite) geometry

revealed that a seemingly small change in that geometry can result in a 40% change in the

measured critical current at a magnetic field of 12T. Specifically, the radial thermal

contraction of the measurement mandrel depends on its wall thickness and, thus, so does

the conductor prestrain (at 4 K) and, ultimately, the measured critical current. Techniques

for reducing variation in the measured critical current are suggested.

Keywords: critical currents; superconductors; Nb3 Sn; VAMAS; NBS

The measurements reported here represent the NBS

contribution to a collaborative effort involving twentyfour laboratories from the European Economic

Community, Japan and the USA under the Versailles

Agreement on Advanced Materials and Standards

(VAMAS). The purpose of this effort is to develop critical current (7C ) measurement techniques for Nb3 Sn superconductors that yield consistent results at different

laboratories. These were part of an interlaboratory

comparison (round robin) where each of the participating

laboratories was supplied with samples of three different

conductors that had been reacted in various coil geometries to match the 7C measuring instruments of the

various laboratories. Each laboratory then measured the

7C of the three samples at several applied magnetic field strengths and two orientations (parallel and antiparallel

to the sample coils’ self field). Preliminary comparisons

of the laboratories’ results showed a large variation in the measured critical current. In response, the NBS study was expanded to address specific measurement variables

that might have resulted in these discrepancies. The

analytical benefit of round robin collaborations is evi- denced by these results where the interlaboratory

comparison of data demonstrated that a subtle measure- ment variable can result in a large difference in the

measured 7C

.

•Publication of the National Bureau of Standards, not subject to copyright

General considerations for accurate critical current

measurements have been given 1,2. The predominant

variables that affect critical current measurements of

Nb3 Sn wires fall into four categories: sample reaction conditions, measurement mandrel (sample holder)

material and geometry, the method used for bonding the

sample to its mandrel and damage incurred during

shipping or in transferring the sample from the reaction

mandrel (tubular stainless steel sample holder) to the

measurement mandrel. In the latter three categories, it is the sensitivity of Nb3 Sn to mechanical strain that results

in variations in the measured 7^. For these measurements, like samples were reacted at a central location

under presumably the same conditions; consequently, the

reaction variables were not suspected of being a significant factor. However, in order to address the question of

shipping damage, specimens of each sample were reacted

and measured at NBS; thus, some measure of the

conductor’s sensitivity to reaction conditions was obtained. The measurement mandrel material and geometry

can work in conjunction with the sample bonding technique to create a dominant measurement variable, the

axial strain of the conductor. Under a condition where 1, the circumferential thermal contraction of the measure- ment mandrel is greater than the axial thermal contraction of the conductor, and 2 the bonding method creates a strong coupling between the sample and the measurement

mandrel, significant axial strain of the conductor at 4 K

may result. It was discovered in this study that the

Cryogenics 1989 Vol 29 July 699

54
Critical current measurements of Nb2 Sn superconductors: L.F. Goodrich and S.L. Bray

circumferential thermal contraction of the G10 fibre- glass-epoxy tubes used for the measurement mandrels in

this study depends on the tube’s geometry, owing to its composite structure. In particular, the ratio of the tube’s

wall thickness to its radius strongly affects its thermal

contraction. Based on these observations, a systematic

study of the effect of G 1 0 measurement mandrel geometry on the measured critical current was made.

For the measurement mandrel study, a filled epoxy

adhesive was used for bonding the sample to the mandrel.

This technique ensured the strong bond and rigid coupling between the sample and mandrel that is necessary

for positive stress transmission. In one case, varnish was

used as the sample mounting adhesive to study the effect of the Lorentz force on the critical current of a weakly

constrained sample. Another variable that depends on

the bonding technique is the thermal coupling between

the sample and the liquid helium bath. Although a complete epoxy coating of the sample results in uniform

mechanical constraint, it also presents a thermal impedance that in some cases may cause a lower measured

critical current. The purpose of the measurements presented here is to document and better understand the sources of inconsistencies in the measured critical current of Nb3 Sn

conductors. This understanding is ultimately necessary

for determining a consistent and practical unified

measurement method. In theory, a measurement method

that would give consistent results is easily achieved by

imposing extensive constraints on all aspects of the

measurement; however, this approach is impractical

considering the variety of Ic measurement systems that

are presently in use. Consequently, a practical measure- ment method should impose the least number of

constraints that yield acceptable results. More challenging than the issue of consistent measure- ments is that of correct measurements. The basic problem

that these round robin tests reveal is that the correct and/or best method for critical current measurements of

Nb3 Sn is unclear. Thus, a statement regarding the

accuracy of these measurements is currently inappropriate. A future joint publication on the results from all participants is planned.

Samples

In order to avoid the identification of commercial

products by the manufacturer’s name, the samples

measured in this round robin will be identified by letter only. Some physical data are provided to indicate the

general type of conductor used in these tests. Sample X had a wire diameter of 0.68 mm and 37

sub-bundles each containing 150 Nb filaments. The

conductor fabrication method was an internal tin dif- fusion process. A single diffusion barrier of Ta separates

the filament region from the outer layer of Cu. The sample

reaction temperature was 700 °C for 48 h. Sample Y had a wire diameter of 1.0 mm and 7 sub-bundles each containing 721 Nb filaments. The

conductor fabrication method was a bronze diffusion

process. Each sub-bundle had a Nb diffusion barrier and

the matrix was a CuSnTi alloy. The sample reaction

temperature was 670 °C for 200 h. Sample Z had a wire diameter of 0.80 mm and 114

sub-bundles each containing 54 NbTa filaments in a bronze matrix. The conductor fabrication method was a bronze diffusion process. The Cu stabilizer was located

at the centre of the wire separated by a Ta barrier. The

sample reaction temperature was 700 °C for 96 h. Specimens of each sample were heat treated (reacted)

at their central laboratory (central reaction). Other

specimens of each sample were to be reacted at each

individual measurement laboratory (self reaction). A

comparison of results from these two groups gives an

indication of reaction and shipping variables. Each

specimen was reacted on a stainless steel tube that had

a spiral groove on it to define the coil pitch. The surface

of the stainless steel had been oxidized to reduce the

chance of diffusion bonding between the sample and the

reaction mandrel. When bonding does occur, the sample

can be damaged when it is removed for mounting onto

the measurement mandrel. The self reactions performed

at NBS were done in a three-zone, vacuum tube furnace.

The temperature of the furnace was controlled during the

reaction to a precision of + 5 °C. The furnace temperature

calibration was checked and was accurate to ± 10°C.

Experimental details

The measurement mandrels were constructed from a G10

tube that had a cylindrical copper current contact rigidly

attached to each end. Each contact had a superconductive

bus bar (Cu and Nb3 Sn tape) connected to it. The sample was transferred from the reaction mandrel onto the

measurement mandrel by carefully unthreading the

sample from the reaction mandrel and slipping it on to

the measurement mandrel (not grooved). The one current contact that was removed for this transfer was then

reattached to the measurement mandrel and the sample was tightened onto the mandrel. Each end of the sample was then soldered onto its current contact. An effort was made to avoid a complete superconductive loop formed

by the sample and bus bar around the current contact. This was done in order to avoid persistent currents. Three

pairs of adjacent voltage taps were placed along the centre

of the sample. Each pair had a separation of about 10 cm

and there was a 1-2 cm gap between adjacent pairs. The

sample was bonded to the mandrel with filled epoxy

(or varnish in one case) and allowed to cure. The epoxy

coat was kept thin to enhance the thermal conductivity

from the sample to the helium bath. This unit was then

mounted onto the test fixture where the electrical connections were completed.

The measurements were made at about 4.2 K in a 12 T,

12.7 cm diameter superconductive solenoidal magnet,

which has a homogeneity of +0.2% over the region

spanned by the voltage taps. Since the ambient atmospheric pressure at this test site is low, resulting in an

equilibrium liquid helium temperature of 4.0 K, the dewar

pressure was elevated. The pressure in the helium dewar

was controlled with a diaphragm-type manostat and the

temperature was deduced from the measured pressure. A

heater on the bottom of the helium dewar was used to remove the natural stratification and bring the bath into

equilibrium at the elevated pressure. The voltage-current curves for all three voltage tap

pairs were recorded simultaneously with a digital processing oscilloscope. An analogue nanovoltmeter with an

700 Cryogenics 1989 Vol 29 July

55
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

amplified output was used to measure the sample voltage

of each tap pair. A battery powered current supply was

used for the sample current

4

. The sample voltage was

measured at numerous sample currents. These points

along the voltage-current curve were analysed to determine 7C

. The measurements were made with forward

(antiparallel magnetic fields, Lorentz force in) and reverse

(parallel magnetic fields, Lorentz force out) current directions to determine the effect of self-field and sample-tomandrel bonding (strain effect). Unless otherwise stated,

the 7C values given are averages over all taps, observations

and current directions at an electric field strength criterion

of 10 pV m

_1

. The precision of the 7C measurement was

about +0.2%.

Another parameter that was determined during the 7C measurement is n. The parameter n is defined by the

approximate voltage-current relationship,

V= F0 (7/7J

where 70is a reference critical current at a voltage

criterion V0 , V is the sample voltage, 7 is the sample

current, and n reflects the shape of the curve with typical

values from 20 to 60 (a higher number means a sharper

transition). A lower value of n can be an indication of

sample damage or strain. Some specimens were difficult to measure because they

would quench (thermal runaway into the normal state) at relatively low sample voltages (electric field strength

about 10 /iV m

_1

). This resulted in some variation in the

7C values due to the small voltage range that could be

analysed. In some cases this low quench level was traced

to damaged segments (weak links) adjacent to a current

contact. These weak links may have been due to sample

damage introduced in removing the sample from the

reaction mandrel, transferring the sample to the measure- ment mandrel, or by strain in the transition length near

the current contact. The transition between the contact

and the G10 has two intrinsic problems. One problem

arises from the differential thermal contraction between

the contact and the G 10, which causes a stress concentration. The other problem is the differential contraction

between the copper contact and the sample, which

introduces additional sample pre-stain. Difficulties in

measuring the Z sample may have been due to its high

current density (535 A mm -2

at 10T) and bronze surface

layer (larger contact resistance) causing sample heating.

Results

Sample X

A summary of the data on sample X is given in Table 1 and Figures 1-4. The information contained in Table 1

is the critical current data, sample mandrel dimensions

and intercomparisons of critical current data. The critical current data are given at magnetic fields from 6 to 12T.

The effect of sample mandrel geometry, bonding method,

thermal cycle, and central or self reaction variables were

studied. Table 1 contains key words that relate to these

variables. Each column, which is labelled with a letter from A to G, is a data set taken for a single thermal cycle.

Specimens are numbered from 1 to 4. Different data

columns that are for the same specimens (columns A and

B for example) indicate a thermal cycle where the

specimen was measured at 4 K, warmed to room temperature, cooled again to 4 K and remeasured. In some

Table 1 Data on sample X; critical current for sample X

Specimen 1 1 1 2 3 3 4

/+H(T) Central Central Central Central Self Self Self thick thick thin thin thin thin thin bored spiral varnish

(A) (B) (C) (D) (E) (F) (G)

6 343.3 343.2 397.3 386.3 388.9 386.6 374.5

7 268.4 267.6 318.4 308.9 312.4 310.3 297.9

8 207.0 207.4 253.8 246.0 249.7 247.6 235.9

9 156.5 156.6 200.1 193.5 197.4 195.6 184.8 10 113.6 114.0 154.7 149.3 153.2 151.6 142.1

11 78.5 116.0 111.9 114.6 112.8 106.6 12 50.2 83.5 80.4 82.5 81.1 74.6

Sample mandrel dimensions

(A) (B) (C) (D) (E) (F) (G)

Wall (mm) 9.3 9.3 2.5 2.0 1.5 1.5 2.0 Rad. (mm) 15.6 15.6 15.6 15.6 15.6 15.6 15.6 w/r 60% 60% 16% 13% 10% 10% 13%

Intercomparison of critical current data

^oH(T) Delta Delta Delta Delta Delta Delta (A-B)/B ( B—C) / C (C-D)/D (E-D)/D (F-D)/D (G-D)/D

6 0.0% — 1 3.6% 2.8% 0.7% 0.1% -3.1%

7 0.3% - 16.0% 3.1% 1 .1 % 0.5% - 3.6%

8 - 0.2% - 1 8.3% 3.2% 1 .5% 0.7% -4.1%

9 -0.1% - 21 .7% 3.4% 2.0% 1.1% - 4.5%

10 - 0.4% - 26.3% 3.6% 2.6% 1 .5% -4.8%

11 - 32.3% 3.7% 2.4% 0.8% - 4.7%

12 - 39.9% 3.9% 2.6% 0.9% - 7.2%

Cryogenics 1989 Vol 29 July 701

56
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

Magnetic Field, T

Figure 1 Percentage variation of /c about the average /c versus magnetic field for specimen G (self, thin, varnish) of sample X. Tap

symbols: O. 1 ; , 2; A- 3

Magnetic Field, T

Figure 2 A semilogarithmic plot of /

c at 10/rV m

_1 versus magnetic field for specimens of sample X. fc = 10/iV m

_1

. O O.

Central thick; A

' ' A. central bored; , central thin; #-•-#, self thin spiral; A , self thin; , self thin varnish

cases, the mounting configuration of the specimen was

changed between these measurements (columns B and

C for example). The sequence of the thermal cycles was

from left to right in the table. The key words heading the columns of Table 1 (central,

thick, thin, etc.) are used to indicate the measurement

variables. Central indicates reaction of the specimen at

702 Cryogenics 1 989 Vol 29 July

Magnetic Field, T

Figure 3 A semilogarithmic plot of n value versus magnetic field for specimens of sample X. O O, Central thick; A • A, central bored; , central thin; #-•-#, self thin spiral; A A.

self thin; , self thin varnish

Magnetic Field, T

Figure 4 Percentage difference between /c measured with reverse and forward current directions versus magnetic field for specimens of sample X. Ec = 10/iV m

-1

. O O. Central thick; A A. central bored; , central thin; #-•-#. self thin spiral; A A.

self thin; , self thin varnish

a central laboratory and self indicates reaction at the

individual measurement laboratory. Thick and thin refer to the ratio of wall thickness to outer radius of the sample

mandrel. Under the Sample Mandrel Dimensions section

of the table, these specific values are listed for each data

set. Bored refers to a specimen that was originally

mounted and measured on a thick-walled mandrel. The

57
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

specimen was then brought to room temperature

(thermally cycled) and the mandrel inner diameter was

bored to a larger dimension, resulting in a reduced wall

thickness. The specimen’s 7C was then remeasured. The

typical method for bonding samples to measurement

mandrels was to coat the sample and the entire surface

of the mandrel with epoxy. Two variations of this technique are indicated by spiral and varnish. Spiral

indicates that the sample was bonded to the mandrel with

a thin stripe of epoxy along its length. Varnish indicates

that the sample and the mandrel were completely coated

with varnish rather than epoxy.

The centrally reacted specimens were numbers 1 and

2. A comparison of the 7C data in columns A and B is a measure of the 7C variation due to thermal cycling. The

maximum variation was 0.4% at 10 T (the highest magnetic field for column A) as shown in the intercomparison

portion of the table ([A-BJ/B). This small variation upon

thermal cycling was representative of all observations.

Column C was a special case of a thin-walled mandrel

in that it was bored (machined with the specimen

mounted on the mandrel). Because this was the same

specimen as A and B, the comparison is a systematic one. The intent of this procedure was to determine the effect of mandrel wall thickness on the measured Ic of a

particular specimen. The intercomparison data ([BC]/C) shows a strongly field dependent variation of 7C (indicative of a strain effect) with a maximum difference

of 39% at 12 T. This is believed to result from a change

in the conductor pre-strain (at 4 K) due to variation in thermal contraction between thick and thin walled G-10

tubes. An approximate measure of the thermal contraction (from room to liquid nitrogen temperature) variations indicated a 0.2% greater contraction for the thickwalled tube. Strain sensitivity measurements on this sample indicate that this additional pre-strain would

result in an Ic degradation in the order of 33%. The

composite nature of G-10 plate

5-7

in conjunction with

the compound geometry of a tube may result in a

variation in radial thermal contraction between tubes of

different wall thickness. Machining may have changed

the mandrel’s outer diameter at room temperature due

to stress relief and, thus, the pre-strain state of the

conductor. More details regarding thermal contraction

of G-10 measurement mandrels and the resulting critical current degradation due to axial strain are presented

below. The rest of the intercomparisons in Table 1 were made relative to D (central, thin) since it was thought to be a more ideal baseline for comparisons. The comparison

between C and D showed only a small difference, 3-4%,

with very little field dependence ([C-D]/D). This indicates

that the effect of boring was about equivalent to that of

simply mounting a specimen on a thin-walled tube.

The self reacted specimens on thin wall tubes were numbers 3 (E and F) and 4 (G). For data set E the

specimen was bonded to the mandrel with a thin stripe

of epoxy along its length (spiral). This spiral mounting

technique was employed to examine the effect of a less rigid containment structure. The comparison between D

and E was within 2.6%. Furthermore, the comparison of

F (E with a continuous epoxy coat) with D was indicative

of the small difference, 1.5%, between central and self reacted wires. The systematic difference between E and

F was, thus, 1.7% or less. This demonstrates that a continuous epoxy coat was not very detrimental in this

case. Specimen 4 (G) was bonded to the mandrel with a coat of varnish in order to investigate the bonding

properties of a weaker and less permanent adhesive. The

7C values in column G were averages of data taken with

forward current only (Lorentz force into the mandrel).

Some measurements were made with reverse current

(Lorentz force away from the mandrel) at 6 and 12 T but

these results were not included in this table. For this data

set, the 12 T (magnetic field of lowest Lorentz force)

measurements were made first to reduce the possibility

of irreversible strain damage. However, even at this relatively low Lorentz force, there was an irreversible shift in the forward 7C after the reverse 7C measurements were

made.

Figure 7 is a plot of the percentage difference between

the observed forward 7c s and the 7C averaged over all taps

and observations at each magnetic field. For the data

between 6 and 1 1 T, the variation of measured 7c s among

the pairs of voltage taps was typical of specimen homogeneity. The indicated precision of the repeat determinations (indicated by multiple data points for the same tap

at the same field) was also typical of these measurements.

The lower group of data points at 12 T were the first measurements of this specimen and were made with

forward current (Lorentz force in). The upper group of

data points at 12 T were taken after reverse 7C (Lorentz

force out) measurements were made (the data for reverse

7C is not shown on the plot). These 12 T data indicate an

irreversible enhancement in the forward 7C which can also

be seen in the (G-D)/D column of Table 1 where a

discontinuity exists betweeen 1 1 and 12 T. The limitations

of varnish as a bonding agent are demonstrated by these

results. First, the limited tensile strength of the varnish is incapable of constraining the wire on the mandrel even

under moderate, outward Lorentz force. Second, even under compressive loads, the varnish seems incapable of

reproducing the results obtained with epoxy. Sample

inhomogeneity could also explain this difference.

Figure 2 is a semi-logarithmic plot of 7C versus magnetic

field. A semi-logarithmic plot was selected because it illustrates the percentage differences between the curves. The central, thick curve is significantly lower than the

others and the thin, varnish curve is slightly lower. Figure

3 is a semi-logarithmic plot of n versus magnetic field. The significant element of this plot is the lower values of

n for the central, thick curve at the higher magnetic fields. This is consistent with a strain degradation. The magnitude of the differences among the other curves is insignifi- cant for this measurement variable. Figure 4 is a plot of

the percentage difference between reverse and forward 7C versus magnetic field. The thin, varnish curve illustrates the large strain effect, 13% at 12T, when the sample is not adequately constrained to the mandrel. The smaller

difference at 6 T was probably due to the decreased strain

sensitivity at the lower magnetic field. The other curves

illustrate the competition between the self-field effect,

which is significant at the lower magnetic fields, and the

strain effect, which is dominant at the higher magnetic

fields. The maximum difference for the specimens bonded

with epoxy was only 2% at 12 T which indicates that this bonding method is very positive.

Sample Y

A summary of the data on sample Y is given in Table 2

and Figures 5-9. The centrally reacted specimens were

Cryogenics 1 989 Vol 29 July 703

58
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

Table 2 Data on sample Y

Critical current (amperes) at 10 /rV m 1

Specimen 1 2 3 3 4 4

Central Central Central Central Self Self

thick thick thick thin bored

thin thin

^oH(T) (A) (B) (C) (D) (E) (F)

6 398.1 372.5 397.1 447.8 429.1 7 332.7 311.0 333.1 377.6 359.2

8 279.4 261.2 280.0 320.5 302.4

9 235.0 220.0 235.3 272.9 255.7

10 184.9 198.9 232.7 216.1

11 154.8 166.9 198.3 182.1 182.9

12 128.6 139.1 168.1 153.4

Sample mandrel dimensions

(A) (B) (C) (D) (E) (F)

Wall (mm) 9.3 9.3 9.3 2.5 2.0 2.0 Rad. (mm) 15.6 15.6 15.6 15.6 15.6 15.6

ratio 60% 60% 60% 16% 13% 13%

Intercomparison of critical current data ^oH(T) (A-C)/C (B-C)/C (C-D)/D (E-D)/D (F-DJ/D

6 0.3% - 6.2% - 1 1 .3% -4.2%

7 -0.1% - 6.6% - 1 1 .8% - 4.9%

8 -0.2% -6.7% -12.6% - 5.6%

9 -0.1% -6.5% - 1 3.8% - 6.3%

10 -7.0% - 1 4.5% -7.1%

11 -7.2% - 1 5.8% -8.2% -7.8%

12 -7.5% -17.3% -8.7%

Magnetic Field, T

Figure 5 Semilogarithmic plot of l

c at 1 0 /A/ m

-1

versus magnetic

field for specimens of sample Y. Fc = 1 0 /iV m

-1

. O

• • O- Central

thick C; A A. central bored; self thin; • •, central thick A; A. A. central thick B

numbers 1, 2 and 3. With respect to mandrel geometry,

mounting technique and the reaction site, these specimens were the same. Data sets A and C are in good agreement

(0.3% maximum variation); however, set B is in contradiction with both A and C (7.5% maximum difference).

There are several possible explanations for this discrepancy including shipping damage, mounting damage,

sample inhomogeneity and variation in mandrel thermal

contraction. These data are insufficient to determine the

most likely source of this effect. Data set D was acquired on specimen 3 after boring and again is an indication of

the pre-strain effect. The magnitude of this 7C enhancement was 17% at 12 T and is in approximate agreement

with a 0.2% change in the pre-strain state of the wire as explained for sample X above. Sample Y’s higher upper

critical field would explain a lower strain sensitivity than

that of sample X (39% at 12T).

Data sets E and F, self reacted specimen number 4, are compared to data set D and again there is a similar 7C discrepancy (8.7% at 12 T) between two comparable

specimens. The additional possible explanation for this

is the fact that these specimens were reacted at different

sites. A comparison of data sets E and F was another

check on the effect of thermal cycling. These abbreviated

data sets and resulting thermal cycles were the result of one of several magnet quenches and, as in the previous

cases, the effect of thermal cycling on 7C appears to be

about 0.4%.

The most significant result for this sample is, again,

the dependence of 7C on mandrel wall thickness.

Figure 5 is a semi-logarithmic plot of 7C versus magnetic

field. These curves show two data bands: an upper band,

associated with the thin-walled specimens, and a lower

band, associated with thick-walled specimens. While

there is considerable variation in the curves, the dominant

variation is between the thick and thin-walled mandrels.

Figure 6 is a semi-logarithmic plot of n versus magnetic

field. The larger data scatter for the central thick C curve may be due to the tap-to-tap 7C variation observed on

704 Cryogenics 1989 Vol 29 July

59
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

Magnetic Field, T

Figure 6 A semilogarithmic plot of n value versus magnetic field for specimens of sample Y. O

' ‘

‘ O. Central thick C; A A.

central bored; self thin; • #, central thick A; A A. central thick B

Magnetic Field, T

Figure 8 Percentage variation of /c about the average /c versus magnetic field for specimen C (central, thick) of sample Y. Tap

symbols; O. 1 : 2; A. 3

o

Magnetic Field, T

Figure 7 Percentage difference between /c measured with reverse and forward current directions versus magnetic field for specimens

of sample Y. fc = 1 0/iV m

-1

. O

• •

• O. Central thick C; A A.

central bored; self thin; # • . central thick A; A A. central thick B

Magnetic Field, T

Figure 9 Percentage variation of /

c about the average /c versus magnetic field for specimen D (central, thick, bored) of sample Y. Tap

symbols: O, 1 ; 2; A. 3

this specimen. The magnitude of the differences among

the other curves is insignificant for this measurement

variable. Figure 7 is a plot of the percentage difference

between forward and reverse 7C versus magnetic field. In

this case the maximum difference was only about 1%.

These curves are similar to those for sample X, Figure 4,

and indicate that the specimen was well bonded to the

mandrel.

Figures 8 and 9 are plots of the percentage difference

between the observed forward Ic s and the 7C averaged

over all taps and observations at each magnetic field. Both of these plots contain data for specimen 3, the

Cryogenics 1 989 Vol 29 July 705

60
Critical current measurements of Nb2 Sn superconductors: L.F. Goodrich and S.L. Bray

difference being that Figure 8 contains central thick C

data and Figure 9 contains central thick-bored D data.

The significant difference between the two plots is that the

tap-to-tap 7C variation, observed in Figure 8, was not

present after boring as shown in Figure 9. One possible

explanation is a reduced longitudinal variation in G10

thermal contraction after boring. The variation of

measured /c s for the other specimens was similar to that

of Figure 9.

Sample Z

A summary of the data on sample Z is given in Table 3

and Figures 10-12. The centrally reacted specimens were numbers 1 and 2 with specimen 3 being self reacted. Data

sets A, B, and C are in good agreement, with a 1.6%

variation in 7C at 12T. The centrally reacted specimens were measured on mandrels that had an outer diameter

of about 19 mm rather than the typical 31 mm of all the

other specimens. Prior to this study, this diameter was

convenient for the anticipated measurements and not

expected to be a significant variable. Given these limited

data and this extra variable, no additional conclusion can be made regarding mandrel geometry and reaction

site.Figure 10 is a semi-logarithmic plot of 7C versus magnetic field. Figure 77 is a semi-logarithmic plot of n

versus magnetic field. The significant information contained in this plot is the relatively high n value for this conductor. Figure 12 is a plot of the percentage difference

between forward and reverse 7C versus magnetic field. In

this case the maximum difference was only about 4%.

With the possible exception of the self reacted specimen,

these curves are similar to those for samples X and Y,

Figures 4 and 7, and indicate that the specimen was well

bonded to the mandrel. Within the limits of the measureTable 3 Data on sample Z

Critical current (amperes) at 10 /A/ m

1

Specimen 1 2 3

Central Central Self ^oH(T) (A) (B) (C)

6 544.1 7 454.5 443.1 8 381.4 372.2 9 320.6 311.5 10 270.6 268.9 263.5

11 226.4 224.3 221.1 12 187.6 185.6 184.6

Sample mandrel dimensions

(A) (B) (C)

Wall (mm) 3.0 3.0 2.0 Rad. (mm) 9.3 9.3 15.6

ratio 32% 32% 13%

Intercomparison of critical current data /^H(T) (A-B)/B (C-B)/B

6

7 - 2.5%

8 - 2.4%

9 - 2.8%

10 0.6% -2.0%

11 0.9% - 1 .4%

12 1 .1 % - 0.5%

Magnetic Field, T

Figure 10 A semilogarithmic plot of /c at 10 /iV m

-1 versus magnetic field for specimens of sample Z. fc = 10//Vm _1

. A A.

Central A; O

• •

• O. central B; , self

Magnetic Field, T

Figure 11 A semilogarithmic plot of n value versus magnetic field for specimens of sample Z. A A, Central A; O

' ‘

‘ O. central B; , self

ment precision for this sample, homogeneity between

taps was similar to the other samples.

Thermal contraction measurements and

strain calculations

Thermal contraction measurements on NEMA-type

(National Electrical Manufacturers’ Association) G10

glass-epoxy composites have been published for the plate

706 Cryogenics 1989 Vol 29 July

61
f Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

Figure 12 Percentage difference between /c measured with re- verse and forward current directions versus magnetic field for specimens of sample Z. £c = 10/jV m

-1

. A A. Central A; O

' •

• O. central B ; , self

geometry 5 “ 7 but not for the tube geometry. The contraction of a G10 plate is highly anisotropic. In composite

nomenclature, the two directions in the plane of the

fibreglass fabric are warp and fill. The direction perpendicular to the plane of the fabric is referred to as the

normal direction. The thermal contraction from 293 K

to 4 K is about 0.24% for the warp direction and 0.71 %

for the normal direction. The contraction in the fill direction is expected to be a little more than that of the

warp direction. The contraction in the warp direction is dominated by the fibreglass fabric and the normal direction is dominated by the epoxy. The G10 tubes used for measurement mandrels were rolled spirals of fibreglass

fabric embedded in an epoxy matrix, rather than tubes

machined from plate material. The radial direction for a

rolled tube is normal to the fabric. The radial contraction

of a tube is different than the contraction of a plate,

however, because the circumferential fibreglass is put in

hoop compression by the epoxy and the resulting contraction is a competition between the two structural

components. The dependence of the radial contraction on wall thickness is expected to be caused by this competition.

A subset of the round robin participants used

measurement mandrels from a single batch of G10 tubes

to reduce the number of measurement variables among

these laboratories. A thick-walled G 1 0 tube was originally

selected as a common sample mandrel material to allow

adaptation of one size tube to the various test fixtures. However, a thick-walled tube is difficult to manufacture

without delaminations that can cause irregular contraction and voids when they are machined. Also, the outer

diameter can change when the inner diameter is bored.

Another observation was that the radial contraction can

be slightly asymmetric; this will result in an approximately elliptical rather than circular cross section at low

temperatures. In order to make an approximate correction for this, the thermal contraction measurements are

the average of two approximately orthogonal measurements of the tube’s diameter.

As briefly mentioned above, thermal contraction

measurements were made on several G10 tubes in order

to estimate the resulting conductor strain. In the interest

of expediency, these measurements were limited both in accuracy and precision; nonetheless, the measurements

give an indication of the variation in thermal contraction

between the tubes. These tests consisted of measuring the

outside diameter of the G 10 tube at room temperature,

submerging the tube in liquid nitrogen, allowing it to reach thermal equilibrium and then quickly removing it and remeasuring the diameter. These measurements were made with a precision micrometer but they were limited

by the practical difficulties of the measurement. Two tube

geometries were measured; both had 31.2 mm nominal

outside diameters, but one tube had a 12.7 mm inside

diameter while the other had a 26.2 mm inside diameter.

These two geometries had 9.3 mm and 2.5 mm walls,

respectively, and are representative of the thick- and

thin-walled mandrels. These measurements showed a

substantial difference between the radial thermal contraction of G10 tubes having different wall thicknesses.

Specifically, the thin-walled tube’s radial thermal

contraction was ~0.23% while the thick-walled tube’s

contraction was ~ 0.41 %. A bias in this study exists as a

result of the thermally transient measurement conditions

(with a systematic difference in thermal mass) and the

fact that it was conducted at liquid nitrogen, rather than

liquid helium, temperature. Both of these factors result

in an underestimation of the actual thermal contraction

between room and liquid helium temperatures. Based on

the temperature dependence of the thermal contraction

of G10 plate 5,6

, the thermal contraction to 4K was

estimated as ~ 0.28% for the thin-walled tube and

~0.48% for the thick-walled tube. The thermal contraction of the Nb3 Sn conductor was only ~0.21%; consequently, the G10 contraction may have introduced

additional amounts of conductor pre-strain. The amount

of compressive pre-strain might depend on several factors

including the mandrel’s wall thickness and the strength

of the mandrel-to-conductor bond.

For samples X and Y, considerable 7C tensile-strainsensitivity data were available; however, no explicit data

were available for compressive strain sensitivity. Consequently, the strain scaling law3 in conjunction with the

available tensile strain data was used to estimate the

expected 7C degradation for compressive strain. For

sample X the results of these calculations predict a ~ 33%

reduction in the 7C at 12 T for the thick-walled specimens

as compared to the thin-walled ones. As mentioned above,

the observed reduction was ~ 39%. Sample Y had a lower

strain sensitivity with a predicted 7C degradation of

~24% at 12T. The actual reduction in 7C for the thick

mandrel as compared to the thin was ~ 18%. As stated

above, these measurements and calculations are only

approximate but the trend seems to be obvious. In all cases, the thick-walled samples had reduced 7c s and this reduction was in approximate agreement with that predicted by the strain scaling law.

Discussion

Sample damage is a major concern because it can go

undetected and have a significant effect on the accuracy

Cryogenics 1989 Vol 29 July 707

62
Critical current measurements of Nb3 Sn superconductors: L.F. Goodrich and S.L. Bray

of the Ic measurement. Two potential sources of damage

are shipping, when the packaging of reacted samples is inadequate, and sample transfer from the reaction mandrel to the measurement mandrel, particularly when

diffusion bonding between the specimen and the stainless

steel reaction mandrel has occurred.

Diffusion bonding has been observed even in cases where the surface of the stainless steel was oxidized prior

to sample reaction. Consequently, a more complete and

reliable technique for oxidizing the reaction mandrel

needs to be developed. If specimen damage during

transfer is to be avoided, dexterity and patience are

required. Unthreading the specimen from a grooved

mandrel seems to be better than attempting to flex the

coil to a larger diameter in order to slip it over the grooves.

Long coils tend to tighten as they are unthreaded unless

the torque is evenly distributed among the coil turns during the process. The other potential source of sample damage, shipping,

is more easily avoided. Careful packaging of the samples

can eliminate concerns over shipping damage. The reaction mandrel alone cannot protect a sample from shipping

damage. The packaging needs to protect the sample from

collisions with external objects as well as from inertial

collisions with other specimens within the container or with the container itself. Within this study it was not possible to separate the

effect of reaction variables or sample inhomogeneity from

that of shipping damage; however, considering the poor

packaging that was used for some of the samples, it is suspected that sample damage may have played a role

in these measurements. The damage or loss of samples

during shipping and the delay of sample shipments by

customs are characteristic problems of international

round robin measurements and, hopefully, with ex- perience these problems will be reduced.

The study concerning the effect of boring and thickversus thin- walled G-10 tubes was not originally intended

but, rather, the results of initial measurements led

naturally to it. Thus, the number of specimens and the

schedule limited the completeness and organization of

these results. A more selective data presentation was not

used because it would imply that there were fewer

difficulties and that less was learned in these studies than was actually the case. A round robin is a unique measurement setting where

the interlaboratory comparison of results is essential to the advancement of measurement accuracy. Although

only the NBS results have been presented herein, it was

the preliminary comparison between these results and

results from the other laboratories that lead to the

expansion of the NBS study and, ultimately, to the source

of the measurement discrepancies.

Conclusions

Two major conclusions are that the bonding method and

the mandrel material and geometry have a significant

effect on the critical current measurement of Nb3 Sn. Thick wall G10 tubes contract about 0.2% more than

thin wall G10 from room to liquid helium temperature.

Thus, the amount of differential thermal contraction

between the mandrel and the sample depends on the wall

thickness of the G10. The method of bonding the sample

to the mandrel also determines how much of the dif- ferential thermal contraction is transmitted to the sample.

A positive bond will change the strain state of the sample more than a weak bond; however, a weak bond may allow

the sample to be strained by the Lorentz force. The

difference between critical current measurements on thick

and thin wall G10 was about 20% at 12 T for a Nb3 Sn

conductor with an upper critical field of 24 T and about

35% at 12 T for one with an upper critical field of 19 T. The effect of self versus central sample reaction was not

indicated as a major source of error in the resulting critical current measurements. The most direct comparison of

this effect can be made on sample X, where the effect was

less than 3% at 12 T. Furthermore, sample X should have

been more sensitive to reaction conditions because of its relatively short reaction time, which may not have fully

reacted the Nb filaments.

Some preliminary recommendations based on these

results can be made. G10 tubes with wall thickness less than about 20% of the tube radius are well matched in thermal contraction to Nb3 Sn wires and use of this type

of tube for measurement mandrels should result in a minimum of conductor strain. The boring of thick-walled G10

tubes for use as measurement mandrels may, due to structural voids, result in irregular thermal contraction.

A well oxidized stainless steel reaction mandrel should

be used to reduce diffusion bonding of the sample to the

reaction mandrel and the possibility of damage when the

sample is transferred to the measurement mandrel. A thin,

continuous coat of epoxy will provide a good bond

between the sample and the mandrel; however, in cases where end cooling of the sample is inadequate, this method may result in a lower measured 7C

. Also, from a

practical standpoint other bonding materials, such as varnish or grease, that are easily removed from the

measurement mandrel may be preferable to epoxy for routine measurements. These bonding materials may be

adequate in cases where the measurements are conducted

with the Lorentz force into the mandrel.

The results presented here are preliminary in the sense

that they have yet to be correlated with measurements

from all 24 participating laboratories. This correlation is a necessary element in a study of this type and may clarify

the results presented here. The effect of thick- versus thin-walled measurement mandrels discovered in this study indicates that in order for various laboratories to

arrive at the same 7C results, a more detailed unification

of the measurement techniques would be required.

However, this would not ensure the correctness of the

results, only a consistency of results. A unified definition

of the desired measurement, aimed at achieving the

maximum usefulness and practicality of the resulting

data, is required in order to make a correct measurement.

Hopefully, a future joint publication will address some

of these issues and result in a simple measurement method

that has only the restrictions necessary to achieve this end.

Acknowledgements

The authors extend their thanks to J.W. Ekin for dis- cussions on these results, to R.M. Folsom and T.C.

Stauffer for sample preparation, data reduction and

plotting, to D.L. Rule for help with data plotting, to W.E.

Look for assistance with the magnetic field calibration

and to the other VAMAS participants.

708 Cryogenics 1989 Vol 29 July

63
Critical current measurements of Nb2 Sn superconductors: L.F. Goodrich and S.L. Bray

This work was supported by the Department of Energy,

Office of Fusion Energy and Division of High Energy

Physics.

An effort was made to avoid the identification of

commercial products by the manufacturer’s name or

label, but in some cases these products might be indirectly

identified by their particular properties. In no instance

does this identification imply endorsement by the

National Bureau of Standards, nor does it imply that the

particular products are necessarily the best available for that purpose.

References

1 Goodrich, L.F. and Fickett, F.R. Critical current measurements:

a compendium of experimental results Cryogenics (1982) 22 225

2 Standard Test Method for D-C Critical Current of Composite

Superconductors Annual Book of ASTM Standards ASTM B714-82,

Part 02.03, American Society for Testing and Materials,

Philadelphia, PA, USA (1983) 595-598

3 Ekin, J.W. Strain scaling law for flux pinning in practical

superconductors. Part 1 : Basic relationship and application to Nb3 Sn conductors Cryogenics (1980) 20 611 4 Bray, S.L., Goodrich, L.F. and Dube, W.P. Battery powered current supply for superconductor measurements Rev Sci Inst (1989) 60 261

5 Fujii, G., Ekin, J.W., Radebaugh, R. and Clark, A.F. Effect of thermal contraction of sample holder material on critical current Adv Cryog Eng Mater (1980) 26 589

6 Fujii, G., Ranney, M.A. and Clark, A.F. Thermal expansion of Nb3 Sn and V3 Ga multifilamentary superconducting cables, fiber- glass-epoxy and cotton-phenolic composite materials Jap J Appl

Phys (1981) 20 L267

7 Clark, A.F., Fujii, G. and Ranney, M.A. The thermal expansion of

several materials for superconducting magnets IEEE Trans Magn

(1981) MAG-17 2316

Cryogenics 1989 Vol 29 July 709
APPENDIX F

THERMAL CONTRACTION OF FIBERGLASS -EPOXY SAMPLE MANDRELS

AND ITS EFFECT ON CRITICAL- CURRENT MEASUREMENTS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

published in:

Proceedings of the 6 t*1 Japan-US Workshop

on

High-Field Superconducting Materials

and

Standards Procedures for High-Field

Superconducting Materials Testing.

Boulder, Colorado

February 22-24, 1989

Editors: K. Tachikawa, K. Yamafuj i and H. Wada

and

J.W. Ekin and M. Suenaga

pp . 91-93, Japan (1989)

65
Thermal Contraction of Fiberglass-Epoxy Sample Mandrels

and Its Effect on Critical-Current Measurements*

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

National Institute of Standards and Technology

Boulder, Colorado 80303

A systematic study was made of the

effect of sample mandrel thermal

contraction on the superconducting

critical-current (I c ) measurement. This

study was conducted in conjunction with

the VAMAS (Versailles Agreement on

Advanced Materials and Standards)

interlaboratory comparison (round robin)

measurements. The Nb 3 Sn wire measured

in this study had a diameter of 0.68 mm.

The sample coils were mounted using a filled-epoxy adhesive on mandrels that

were made from fiberglass-epoxy (G-10)

tubes of several different wall

thicknesses

.

Figure 1 is a semilogarithmic plot of

I c versus magnetic field for three

different sample mandrels, designated

thick, bored , and thin . The variable

among the three sample mandrels is the

ratio of their radius to their wall

thickness. The relative dimensions of

the three mandrels are given in Table I.

In all cases, the mandrels were

constructed from G-10 tubes. The

precision of the I c measurement was

±0.2%. The uncertainty of the I c measurements may be as large as ±5%

because of the strong systematic effects

of sample mandrels.

The bored designation refers to a specimen that was first mounted and

measured on the thick walled mandrel.

The specimen was then brought to room

temperature and the mandrel's inner

diameter was bored to a larger size,

Table I. Mandrel Dimensions

thick bored thin

Wall, mm 9.3 2.5 2.0

Rad. , mm 15.6 15.6 15.6

Wall/Rad 60% 16% 13%

* Contribution of NIST, not subject to

copyright

.

resulting in a reduced wall thickness.

The specimen's I c was then remeasured.

On the other hand, the thin designation

refers to a specimen that was simply

mounted on a mandrel that had been

already machined to a reduced wall

thickness. The reason for the two

different approaches was to demonstrate

that neither sample variability nor

changes in sample strain introduced

during boring were principally

responsible for variations in the

measured I c

. A comparison between the thin and

bored specimens shows only a small

difference, 3 to 4%, in I c with very

little field dependence. This indicates

that the effect of boring was about

equivalent to that of simply mounting a

Figure 1. A semigarithmic plot of I c for an electric field criterion of 10

fiV/m versus magnetic field.

66
specimen on a thin-walled tube. In

comparison to the thin and bored

specimens, the thick specimen shows a significant reduction in I c ; the

difference between the curves increases

with increasing magnetic field (40% at

12 T) , which is indicative of a strain

effect.

The thick- and the thin-walled sample

holders are made from the same materials

and constructed using the same

techniques. The difference in their

radial thermal contraction is a result

of their highly anisotropic structure.

In the case of plates that are made of

this material, the thermal contraction

normal (perpendicular) to the plate is

considerably greater than in the plane

of the plate. A competition between

these two contractions occurs in the

tube geometry where the direction of

greater contraction is radial and the

direction of lesser contraction is

circumferential. Consequently, the wall

thickness of the tube affects its

overall radial thermal contraction.

Thermal contraction measurements were

made on several G-10 tubes in order to

estimate the resulting conductor strain.

These measurements showed a substantial

difference among the radial thermal

contraction of G-10 tubes that have

different wall thicknesses. The results

of the thermal contraction measurements

are shown in Table II. The thermal

contraction was determined by measuring

the outside diameter of the G-10 tube at

room temperature, submerging the tube in

liquid nitrogen, allowing it to reach

thermal equilibrium, and then quickly

removing it and remeasuring the

diameter. These measurements were made

with a precision micrometer, but their

accuracy was limited by the practical

difficulties of the measurement. The

estimated uncertainty in the value of

Ad/d was ±10%. As indicated by the

column headings in the right hand

portion of the table, four different

samples were measured.

For each sample, the table shows the

tube's wall thickness, the ratio of the

tube's wall thickness to its radius in

percent, and the thermal contraction of

the tube between room and liquid

nitrogen temperatures (Ad/d) . Samples 1 and 2 were progressively bored to larger

inside diameters while samples 3 and 4

had factory diameters (not machined) and

were measured to test the equivalence

between factory tubes and bored tubes

with respect to thermal contraction.

The data indicate that the thickerwalled tubes contract more than the

thinner ones. The agreement of the data

for samples 3 and 4 with those of 1 and

2 indicates that the bored tubes behave

in a manner similar to that of factory

tubes. Samples 1 and 2 with wall

thickness/radius ratios of 59%, 16%, and

13% have respectively similar dimensions

to the thick , bored, and thin

measurement mandrels.

A bias in this thermal-contraction

study exists as a result of the fact

that it was conducted near liquid

nitrogen, rather than liquid helium,

temperature. This factor results in an

underestimate of the actual thermal

contraction between room and liquid

helium temperatures. Based on the

temperature dependence of the thermal

contraction of G-10 plate, ^ the thermal

contraction to 4 K was estimated as ~0.28% for the thin-walled mandrel and

Table II. Thermal Contraction of G-10

Tubes Between Room and Liquid Nitrogen

Temperatures

Tube Dimensions Outer Diameter

Thermal Contraction

Ratio

Wall

to Wall Sample 1 Sample 2

Radius Thickness Ad/d Ad/d

59% 9.27 mm 0.39% 0.41%

49% 7.70 mm 0.38% 0.36%

39% 6.10 mm 0.33% 0.32%

28% 4.32 mm 0.30% 0.28%

16% 2.54 mm 0.25% 0.23%

13% 2.03 mm 0.25% 0.24%

Ratio

Wall

to Wall Sample 3 Sample 4

Radius Thickness Ad/d Ad/d

20% 3.18 mm 0.25%

30% 4.76 mm 0.28%

67
-0.48% for the thick-walled mandrel.

The thermal contraction of the Nb 3 Sn

conductor is only -0.21%. Furthermore,

the Nb 3 Sn conductor was rigidly bonded

to the mandrel with a filled- epoxy

adhesive. Consequently, the G-10

contraction may have introduced

compressive strain to the conductor.

Considerable data on the sensitivity

of I c to tensile strain were available

for this sample; however, no explicit

data were available for compressive

strain sensitivity. Consequently, the

strain scaling law,^ in conjunction with

the available tensile strain data, was

used to estimate the expected I c degradation for compressive strain. The

results of these calculations predict a -33% reduction in I c at 12 T for the

thick-walled specimens as compared to

the thin-walled ones; the observed

reduction was -40%. These measurements

and calculations are only approximate,

but the trend is apparent. In all

cases, the thick-walled samples had

reduced critical current and this

reduction was in approximate agreement

with that predicted by the strain

scaling law.

0-10 tubes that have diameters similar

to those tested and wall thicknesses

about 20% or less of the tube's radius

are well matched in thermal contraction

to Nb 3 $n wires. Use of this type of tube

for measurement mandrels should result

in a minimum of conductor strain. The

scaling of this effect for tubes of

various diameters has not been

demonstrated, but scaling is probably

accurate for a limited range of tube

diameters

.

This, work was supported by the

Department of Energy, Office of Fusion

Energy and Division of High Energy

Physics

.

An effort was made to avoid the

identification of commercial products by

the manufacturer's name or label, but in

some cases these products might be

indirectly identified by their

properties. In no instance does this

identification imply endorsement by the

National Institute of Standards and

Technology, nor does it imply that the

products are necessarily the best

available for that purpose.

References

1. G. Fujii, J. W. Ekin,

R. Radebaugh, and A. F. Clark,

"Effect of Thermal Contraction of

Sample Holder Material on Critical

Current , " Adv. Crvog. Eng . -Materials , vol. 26, pp. 589-598, Plenum Press,

New York (1980)

.

2. J. W. Ekin, "Strain Scaling Law for

Flux Pinning in Practical

Superconductors. Part 1: Basic

Relationship and Application to

Nb 3 Sn Conductors," Cryogenics . vol.

20, pp. 611-624, Nov. 1980.

68
APPENDIX G

THERMAL CONTRACTION OF FIBERGLASS - EPOXY SAMPLE

HOLDERS USED FOR Nb 3 Sn CRITICAL- CURRENT MEASUREMENTS

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

Presented at the International Cryogenic Materials Conference

Los Angeles, California

24-28 July 1989

to be published in: Advances in Cryogenic Engineering

Materials, Volume 36

69
THERMAL CONTRACTION OF FIBERGLASS - EPOXY SAMPLE

HOLDERS USED FOR Nb 3 Sn CRITICAL- CURRENT MEASUREMENTS*

L. F. Goodrich, S. L. Bray, and T. C. Stauffer

Electromagnetic Technology Division

National Institute of Standards and Technology

Boulder, Colorado 80303

ABSTRACT

It is typical for Nb 3 Sn-Cu superconductor specimens to be wound into

coils on tubular specimen holders for critical-current measurements. If the

thermal contraction of the holder is different than that of the specimen,

axial strain may be applied to the specimen upon cooling from room to

liquid-helium temperature. This strain can affect the measured critical

current. The thermal contraction was measured for three different Nb 3 Sn-Cu

superconductors. Also, the thermal contraction was measured for several

different specimen holders, all of which were made from fiberglass - epoxy

composites. The specimen holder measurements were made using an electricalresistance strain- gage technique, and they were confirmed by direct

mechanical measurements. The tubes varied in diameter, wall thickness, and

fabrication technique. Some of the tubes were made directly from tube

stock, and others were machined from plate stock. The results of these

measurements show that the thermal contraction of tube stock is strongly

dependent on the ratio of its wall thickness to its radius, while the

contraction of tubes machined from plate stock is relatively independent of

these dimensions. Critical-current measurements of Nb 3 Sn-Cu specimens

mounted on these various holders show that the presence of differential

thermal contraction between the specimen and its holder can significantly

affect the measured critical current.

INTRODUCTION

Critical-current (I c ) measurements of Nb 3 Sn-Cu superconductors can be

affected by the relative thermal contraction of the material on which the

test specimen is mounted. This measurement variable became apparent in the

recent VAMAS (Versailles Project on Advanced Materials and Standards)

interlaboratory comparative measurements (round robin) of the critical

current of Nb 3 Sn.^ Differential thermal contraction between the specimen

and its holder in cooling from room to liquid-helium temperature can cause

either a tensile or compressive strain of the specimen, either of which can

affect the measured I c

. For coil specimens that are mounted on the

surface of cylindrical holders, the strain is predominantly along the axis

of the specimen. The amount of strain depends on the magnitude of the

^ 1 7 Contribution of NIST, not subject to copyright.

70
differential contraction, the relative strengths of the specimen and holder,

and the mechanical coupling between the specimen and its holder.

The literature contains considerable data on the compressive pre- strain

of Nb 3 Sn filaments caused by differential thermal contraction between the

filaments and the matrix material. However, very little data on the overall

thermal contraction of Nh^Sn-Cu wires are presently available. The thermal

contraction of a Nb 3 Sn-Cu cable is given in Ref. 4; however, this conductor

has a tungsten core that reduces its thermal contraction. Consequently, the

thermal contraction of three different Nb 3 Sn wires was measured as a part of

this study.

Fiberglass-epoxy composites are commonly used sample holder materials.

These materials are anisotropic in three mutually perpendicular directions.

The three directions are associated with characteristics of the fiberglass

fabric, and they are designated as the warp, fill, and normal directions.

The fabric orientation for plate stock is shown in Fig. 1. The normal

direction is perpendicular to the fabric planes, while the warp and fill

directions are parallel to the fabric planes. The density of the fabric is

not the same in both directions of the weave. The warp and fill directions

are determined by the fabric's thread count. The number of threads per unit

length of fabric is lower in the warp direction than in the fill direction.

This structural anisotropy causes a three-dimensional variation in thermal

contraction. The contraction in the fill direction is slightly greater than

in the warp direction, but the contraction in the normal direction is

considerably larger.^’

Figure 1 also shows the typical fabric orientation for a rolled tube.

The radial thermal contraction of the tube results from a competition

between the larger contraction in the normal direction and the smaller

contraction in the warp direction. Consequently, it depends on the ratio of

the tube's wall thickness to its outside radius (wall- to-radius ratio). For

thin-walled tubes the radial thermal contraction approaches that of a plate

in the warp direction, and for thick walled tubes it approaches that of a plate in the normal direction.

A holder whose thermal contraction is relatively independent of its

wall- to-radius ratio can be made by machining a cylindrical tube from thick

fiberglass-epoxy plate stock with the axis of the tube perpendicular to the

surface of the plate (plate tube). For this orientation (Fig. 1) the radial

contraction is based on the contraction in the warp and fill directions,

Plate Plate Rolled

Tube Tube

Fig. 1. Geometries of fiberglass -epoxy composites.

71
which are both similar to that of a NT^Sn-Cu wire. In addition to thermal

contraction measurements of both rolled tubes and plate tubes, I c measurements were made using both types of specimen holder to confirm the

relationship between thermal contraction and the I c measurement.

EXPERIMENTAL DETAILS

Fiberglass-Epoxv Composites

Two different types of fiberglass-epoxy composite were used in this

study, NEMA (National Electrical Manufacturers' Association) G-10 and G-ll.

All of the rolled tubes were G-10 and the plate tubes were G-ll. This was

not a matter of choice but, rather, one of material availability. The

thermal contraction of the G-ll is slightly less than that of G-10. Neither

of these materials was cryogenic-radiation (CR) grade, which is designated

as G-10CR or G-11CR. The manufacturing specifications for the CR grade

materials are more stringent and their performance at cryogenic temperatures

is more predictable.^

Thermal Contraction

All of the thermal contraction measurements were conducted between room

temperature and liquid-nitrogen (LN2 ) temperature. We assume that, if a material's contraction is well matched to a Nb 3 Sn-Cu wire from room to LN2

temperature, then the additional differential thermal contraction that

occurs between LN2 and liquid-helium temperature will be insignificant.

The Nb 3 $n-Cu thermal contraction measurements were made using quartz

reference tubes and reacted Nb 3 Sn-Cu specimens that measure approximately

35 cm in length. The specimen is placed in a quartz tube and attached to

one of the tube's ends. At room temperature, the quartz tubes are

approximately 1 mm longer than the specimens. This differential length is

measured with a micrometer, the tube and specimen are cooled to LN2

temperature, and the differential length is remeasured. The thermal

contraction of the specimen is deduced from the measured change in the

differential length and the known contraction of the quartz tube. The

uncertainty of these measurements is estimated to be ±5%.

Two different methods were used to measure the thermal contraction of

the fiberglass tubes. The first was a mechanical method, where the tube was

submerged in liquid nitrogen, allowed to reach thermal equilibrium, removed

from the nitrogen, and the diameter was then quickly measured with a precision micrometer. To address the thermally transient nature of these

measurements, the tube's diameter was measured as a function of time while

it was warming toward room temperature. These measurements allowed an

extrapolation of the data to liquid-nitrogen temperature. In the case of

the plate tubes, two orthogonal measurements were made at each time, one in

the warp direction and one in the fill direction. The uncertainty of the

mechanical measurements of thermal contraction was estimated to be ±10% of

the measured value (for example, 0.20% ±0.02%).

The second type of thermal contraction measurement used electricalresistance strain gages. ^ For this measurement, two well matched 350 ft strain gages were used in a half-bridge configuration. One of the gages was

bonded to the test specimen and the other was bonded to a reference

material, a quartz tube, whose thermal contraction was known. All of the

strain-gage measurements reported here were taken with the strain gages

mounted on the circumference of the test specimen since this represents the

contraction which is relevant for the Nb 3 Sn-Cu coil. Both the test specimen

72
and reference material were cooled to liquid nitrogen temperature and the

resulting output from the strain bridge was measured.

Two factors will cause a change in the resistance of the strain gages

when cooled from room to LN2 temperature. First, the resistivity of the

strain gages' grid alloy will change with temperature. Since this change in

resistance will be nearly equal for matched strain gages, there will be

little effect on the output of the strain bridge. The second source of

change in resistance is caused by thermally induced strain. This strain

results from differential thermal contraction between the strain gage and

the test specimen. If the thermal contraction of the test specimen material

is different than that of the reference material, the strain bridge will be

unbalanced and its output will be indicative of the difference in thermal

contraction between these two materials. Since the thermal contraction of

the reference material is known, the thermal contraction of the test

material can be deduced. The uncertainty of the strain- gage measurements of

thermal contraction was estimated to be ±5% of the measured value (for

example, 0.20% ±0.01%).

Critical Current

For the I c measurements, a Nb 3 Sn-Cu wire with a diameter of 0.68 mm was

used. This conductor was made by an internal-tin diffusion process, and it

has 37 sub-bundles of 150 Nb filaments. A single Ta diffusion barrier

separates the filament region from the outer Cu layer. The specimen was

wound onto a stainless steel tube and then vacuum heat treated at 700 °C for

48 h. The stainless steel tube has a helical groove machined into its

surface to retain the specimen and define its geometry. Following heat

treatment, the specimen was removed from the stainless steel reaction holder

and transferred to the fiberglass measurement holder. The outside diameter

of the fiberglass holder is 3.12 cm and its surface is not grooved. A thin

continuous layer of filled epoxy adhesive was painted over the surface of

the specimen and holder. The typical specimen length was approximately

80 cm. Three pairs of adjacent voltage taps were placed along the center of

the specimen. Each pair had a separation of about 10 cm, and there was a 1 to 2 cm gap between adjacent pairs. An electric field criterion of 10 ,uV/m was used for determining the critical current. The uncertainty of the I c measurements may be as large as ±5% because of the strong systematic effects

of the specimen holders.

RESULTS

Thermal Contraction Measurements

The results of the thermal contraction measurements are shown in Fig. 2

where the thermal contraction of the tubes' diameters (Ad/d) is plotted as a function of the tubes' wall- to-radius ratios. The thermal contraction of

three different superconductors (the VAMAS I c round robin conductors 1

) was

also measured. The dashed horizontal line labeled "Nb 3 Sn-Cu" indicates the

measured thermal contraction (0.28%) of the conductor whose I c was measured

in this study. The measured thermal contraction of the other two conductors

is 0.28% and 0.26%. The two types of data symbols that are labeled

"Mechanical" and "Strain Gage" show the thermal contraction of the G-10

rolled tube specimens as measured by the two different techniques. With the

exception of the "Mechanical" data, all of the data shown in this figure

were obtained using the strain-gage technique.

The data show a strong dependence of the thermal contraction on the

wall- to-radius ratio. The approximate thermal contraction of G-10CR plate

stock in cooling from room to liquid-nitrogen temperature is 0.224%, 0.264%,

73
and 0.687% in the warp, fill, and normal directions respectively, and it is

0.202%, 0.227%, and 0.585% for G-11CR. 4,5 As expected from the structural

geometry of the rolled tubes, their thermal contraction fall between that of

the warp and normal directions of plate stock. The measured thermal

contraction of the G-ll plate tubes is shown for three wall- to-radius ratios

and in two structural directions, warp and fill. In both directions, the

thermal contraction is relatively independent of the wall- to-radius ratio

and it is comparable to that of plate stock. The error bars associated with

the high wall- to-radius ratio data points represent the range of values that

were measured on three different specimens with three repeat determinations

for each specimen.

Additional measurements were made using strain gages that were mounted

on the circumference of the plate tube halfway between the warp and fill

axes. The thermal contraction at these positions was within the

experimental uncertainty of the value measured in the warp direction. This

indicates that the effective circumferential thermal contraction of the

plate tube is closer to the value measured for the warp direction.

Mechanical thermal contraction measurements were also made on the plate

tubes and were within the experimental uncertainty of the strain- gage

measurements. Additional thermal contraction measurements were made on

tubes with different outer diameters to test the scaling of thermal

contraction with the wall- to-radius ratio. All of the data of Fig. 2 were

taken on tubes having outer diameters of about 3.18 cm and various inner

diameters. However, two rolled tubes with outer diameters of 11.4 and

Fig. 2. Plot of room to liquid-nitrogen temperature thermal

contraction of G-10 rolled tubes and G-ll plate tubes

as a function of the tubes' wall- to-radius ratio.

74
12.4 cm, and vail- to-radius ratios of 14 and 8%, respectively, were also

measured. The thermal contraction of these two tubes was within the

experimental uncertainty of the rolled tube data of Fig. 2. Finally, an

11.4 cm outer diameter (94% wall- to-radius ratio) plate tube was measured

and was within the experimental uncertainty of the plate tube data of

Fig. 2.

Critical-Current Measurements

The results of the I c measurements are shown in Fig. 3, where the I c is

plotted as a function of applied magnetic field for several different

conductor specimens and specimen holders. There are three variables, in

addition to the applied magnetic field, for these I c measurements: the

location where the specimen was reacted, "central" or "self"; the specimen

holder's wall- to-radius ratio, "thick" or "thin"; and the type of tube used

for the specimen holder, "rolled" or "plate". The self reacted samples were

reacted at NIST, whereas the central reacted samples were reacted at another

site and shipped to NIST for I c measurements. The thick and thin

designations do not indicate specific wall- to-radius ratios; instead, they

indicate two general categories of tube geometry. All of the thin tubes

have wall- to-radius ratios that are no greater than 13%

Fig. 3. A semilogarithmic plot of I criterion of 10 /iV/cm as a field

c at an electric field

function of magnetic

75
(self 10% and central 13%) and the thick tubes have wall- to-radius ratios

that are at least 60% (plate 80% and rolled 60%). The measured I c

's are

nearly the same, within experimental uncertainty, for all of the specimens

except the one that was mounted on the thick, rolled tube. The measured I c is significantly lower for this specimen at all magnetic fields. The

difference is about 14% at 6 T and 40% at 12 T. The fact that the I c degradation increases with increasing magnetic field is consistent with a strain effect. The magnitude of the change in I c and the measured thermal

contraction are in good agreement with the strain effect.-* At 12 T, a 33%

reduction was calculated from strain-effect measurements^ that were made at

14 T, and the measured I c reduction was 40%.

DISCUSSION

A practical reality of round robin measurements is that the dimensions

of the specimen holders vary between different laboratories. For consistent

measurements, the specimen holders should be designed so that the I c measurement is insensitive to this variable. The thermal contraction of a tubular specimen holder that is made from an anisotropic material can vary

with its geometry. This presents the potential for variations in the strain

state of the specimen and, thus, variations in the measured I c

. An apparent

solution to this problem is to use an isotropic material for the specimen

holders and to use a bonding technique that rigidly couples the specimen to

its holder. This will ensure that the strain transmitted to the specimen,

due to thermal contraction, is independent of the holder's geometry and,

thus, equivalent from laboratory to laboratory. This approach addresses the

issue of measurement consistency, but it does not address accuracy.

The I c measurement should, arguably, be made with a minimum of

externally applied strain on the superconductor. This requires a strong

bond between the specimen and holder to avoid specimen strain under the

influence of the Lorentz force, and it requires that the thermal contraction

of the sample holder be well matched to that of the superconductor. Also,

the holder should, ideally, be made from an electrically insulating material

to prevent current sharing with the test specimen. Unfortunately, an

isotropic and insulating material with a thermal contraction similar to that

of Nb 3 Sn-Cu is not readily available. Fiberglass-epoxy plate tubes are a practical alternative to the ideal isotropic specimen holder. The thermal

contraction of a plate tube is slightly anisotropic; however, it is

relatively independent of the tube's dimensions. Furthermore, the thermal

contraction in the radial direction (the pertinent direction for a coil- type

specimen) is similar to that of Nb 3 Sn-Cu. Based on the thermal contraction

of G-10CR and G-11CR plate, the thermal contraction of G-10 plate tubes

may be slightly closer to that of a Nb 3 Sn-Cu wire than the G-ll plate tubes

that were measured here. It is possible that the difference in the

circumferential thermal contraction between the warp and fill directions

will result in a spatial variation of the strain state of the Nb 3 Sn-Cu coil

sample, but the I c data indicate that this effect is not significant.

The plate tubes present some practical disadvantages. First of all,

machining a specimen holder from plate stock is considerably more difficult

than from tube stock. Also, the length of a plate tube is limited by the

thickness of the available plate stock. This, in turn, limits the length of

the superconductor specimen for a given coil diameter and pitch.

Furthermore, short specimen holders are often incompatible with existing I c test fixtures. Satisfactory specimen holders of greater length could

perhaps be constructed by bonding a series of short plate tubes together.

This technique might require an alignment of the warp and fill fibers

between individual tube sections because of the anisotropic radial thermal

contraction of plate tubes.

76
CONCLUSIONS

For I c measurements, fiberglass -epoxy composites are suitable specimen

holder materials. However, the design of the specimen holder should take

into account the anisotropic nature of the material and the resulting

variability in thermal contraction. These characteristics of the material

can result in large variations in the measured I c for specimens mounted on

holders of different designs. A cylinder or tube can be machined from a thick fiberglass -epoxy plate with the axis of the cylinder perpendicular to

the surface of the plate. This type of tube (plate tube) has a thermal

contraction that is relatively independent of its dimensions and that is

similar to that of a Nb 3 Sn-Cu specimen. Alternatively, the specimen holder

can be made from tube stock and machined to a wall- to-radius ratio that

results in a thermal contraction that closely matches that of the Nb 3 Sn-Cu

specimen.

ACKNOWLEDGEMENTS

The authors acknowledge the contribution of R. L. Spomer, S. Bird, and

T. Harris who made some strain- gage measurements on the G-10 rolled tubes as part of a Colorado School of Mines class project. The authors extend their

thanks to M. Thoner (Vacuumschelze) for a discussion of the plate tube

geometry, to R. M. Folsom for sample preparation and assistance with the

measurements, to J . D. McColskey for instructions on mounting strain gages,

and to R. Gerrans for drafting.

This work was supported by the Department of Energy, Office of Fusion

Energy and Division of High Energy Physics.

An effort was made to avoid the identification of commercial products

by the manufacturer's name or label, but in some cases these products might

be indirectly identified by their particular properties. In no instance

does this identification imply endorsement by the National Institute of

Standards and Technology, nor does it imply that the particular products are

necessarily the best available for that purpose.

REFERENCES

1. K. Tachikawa, K. Itoh, H. Wada, D. Gould, H. Jones, C. R. Walters,

L. F. Goodrich, J. W. Ekin, and S. L. Bray, VAMAS intercomparison

of critical current measurement in Nb 3 Sn wires IEEE Trans . Magn

25-2:2368 (1989).

2. G. Fujii, J. W. Ekin, R. Radebaugh, and A. F. Clark, Effect of thermal

contraction of sample holder material on critical current

Adv. Crvog . Eng

26:589 (1980).

3. J. W. Ekin, Strain scaling law for flux pinning in practical

superconductors. Part 1: basic relationship and application to Nb 3 Sn

conductors Cryogenics 20:611 (1980).

4. A. F. Clark, G. Fujji, and M. A. Ranney, The thermal expansion of

several materials for superconducting magnets IEEE Trans . Mag

MAG -17:2316 (1981).

5. M. 3. Kasen, G. R. MacDonald, D. H. Beekman, Jr., and R. E. Schramm,

Mechanical, electrical, and thermal characterization of G-10CR and

G-11CR glass-cloth/epoxy laminates between room temperature and 4 K

Adv . Crvog . Eng

26:235 (1980).

6. Measurements Group, Inc., Tech Note-513, Measurement of thermal

expansion coefficient using strain gages (1986).

77

APPENDIX H

VAMAS INTERCOMPARISON OF CRITICAL CURRENT MEASUREMENT IN Nb 3 Sn WIRES

K. Tachikawa, K. Itoh, H. Wada, D. Gould, H. Jones

C. R. Walters, L. F. Goodrich, J. W. Ekin, and S. L. Bray

IEEE Trans, on Magnetics 25.(2) 2368-2374 (March 1989)

79
IEEE TRANSACTIONS ON MAGNETICS, VOL. 25, NO. 2. MARCH 1989

VAKAS INTERCOMPARISON OF CRITICAL CURRENT MEASUREMENT IN Nb^Sn WIRES

K . Tachikawa* , K. Itoh**, H.

+ Vada**, D. Could***, H. Jones"^

C.R. Walters , L.F. Goodrich , J.V. Ekin and S.L. Bray

*Tokai University, Hiratsuka, Kanagawa 259-12, Japan

‘‘National research Institute for Metals, Tsukuba, Ibaraki 305, Japan

‘Commission of the European Communities, Brussels, B-1049, Belgium

Clarendon Laboratory, Oxford, 0X1 3PU, United Kingdom

Rutherford Appleton Laboratory, Didcot, 0X11 OQX, United Kingdom

National Bureau of Standards, Boulder, Co 80303, USA

Abstract

The VAMAS technical working party in the area of su- perconducting and cryogenic structural materials has

recently carried out the first world-wide intercomparison of critical current, I e

, measurement on multifilamentary Nb,Sn wires. Three sample wires were supplied from each of EC (European Communities), Japan and

USA. The total number of participant labs were 24 (EC

11, Japan 8 and USA 5). There were few restrictions

for the I measurement at participant labs. The standard deviations of the I values reported from these

labs varied among test samples, and were 6-21Z of

averaged I^’s at 12 Tesla.

Introduction

The superconductivity technology may have tremendous

impact on important areas of science and technology and

should be developed under the concept of the long term

project whereby international cooperation would play an essential role. This is the underlying idea in the

VAMAS which stands for the Versailles Projects among

summit participant countries on Advanced Materials and

Standards. The VAMAS Technical Working Party (TWP) on superconducting and cryogenic structural materials consists of representatives of participant countries. The

TWP has carried out an intercomparison on the critical

current, I , measurement in Nb,Sn multifilamentary

wires: I is the most important superconducting

parameter from the practical point of view.

The purpose of the present intercomparison (round

robin) test on I is to identify parameters affecting

the 1 by accumulating and evaluating the measurement

results. The eventual goal of the research is to provide recotmnendations for the performance of short

sample critical-current measurements of Nb^Sn superconductors.

Participants, Samples and Test Procedure

11 EC, 8 Japanese and 5 US laboratories listed in

Table I have participated in the round robin test on the I . The distribution of test samples and accumulation of resulting I data in EC, Japan and USA were performed by the respective central labs;

BCMN(Belgium), NRIM, and KBS

.

Multifilamentary Nb.Sn wires with relatively small

current carrying capacity (<500 A at 8 Tesla) were chosen as the test samples which could easily be tested

at any participant laboratory. Three sample wires were supplied, one from each of EC, Japan and USA; these

samples are labeled disorderly Just as sample A, B and

C In this paper.

Manuscript received August 22, 1988.

Sample A fabricated by a bronze method has a wire

diameter of 0.8 mm and 114 sub-bundles each containing

90 Nb-Ta filaments in a bronze matrix. The Cu stabilizer is located at the center of the wire and

separated from the filament region by a Ta barrier.

The volume fraction of Cu stabilizer in the wire is

much smaller compared to those in other 2 samples.

Sample B has a wire diameter of 1,0 cm and 7 subbundles each separated by a Nb barrier from the Cu stabilizer outside. Each sub-bundle contains 721 Nb filaments in a Cu-Sn-Ti alloy matrix. This sample was also

fabricated by a bronze process.

Sample C prepared by an internal-Sn diffusion method

has a wire diameter of 0.68 nut and 37 sub-bundles each

containing 150 Nb filaments. The filament region is

separated from the outer layer of Cu by a single Ta barrier.

Specifications and the cross sectional views of

these samples are given in Fig 1. Upper critical

fields, H

£ 2

> for sample A and sample B are enhanced by

Table 1. Participant laboratories in I round robin

test.

C

Europe (11)

Atominstitut der Oesterr eichischen Univ. (Austria)

Inst. Experimental Physik, Oester. Univ. (Austria)

S.C.K./C.E.N (Belgium)

S.N.C.I., C.N.R.S. (France)

Kernforschungszentrum Karlsruhe (FRG)

Siemens AG (FRG)

Vacuumschmelze GmbH (FRG)

E.N.E.A, Centro di Frascati (Italy)

High Field Magnet Lab., U. Nijmegen (Netherlands)

Clarendon Laboratory (UK)

Rutherford Appleton Laboratory (UK)

Japan (8)

Electrotechnical Laboratory

Furukawa Electric Co., Ltd.

Hitachi Ltd.

I.M.R., Tohoku University

I. S. I. R., Osaka University

Japan Atomic Energy Research Institute

Kobe Steel, Ltd.

National Research Institute for Metals

USA (5)

Brookhaven National Laboratory

F.B.N.M.L., Massachusetts Institute of Technology

Lawrence Livermore National Laboratory

National Bureau of Standards

University of Wisconsin

(KMS-9-^'4 ^9 '

1 u 89 IEEE

80
additions of Ta and II, respec lively.

Specimen vires from these samples, each several

meters long, were distributed to participant labs

through relevant central labs. Such vires vere cut into parts as specimens at each participant lab and one

specimen, vound on a heat treatment holder of drum

shape, vas collected from participant labs and heat

treated at the relevant central labs (central

reaction). The reacted specimens were then returned to

the participant labs for measurements. In some cases,

the participant labs also performed their own heat

treatments on additional specimens of the samples (self

reaction). The self reaction of samples vas not possible at all labs.

Specifications of heat treatment holders vere different from lab to lab. At most of the labs stainless

steel tubes with a spiral groove on the outer surface

vere used to have a definite coil pitch. The surfaces

of such tubes vere usually coated vith ceramics or oxidized prior to specimen mounting, in order to avoid

reaction of the specimen vith the holder. At some labs

the same holders vere used for both heat treatment and

measurement, thereby reducing the possibility to damage

the specimen by handling.

Central reaction vas carried out in the following

manner. For samples A and B, all specimens vere heat

treated in one vacuum furnace at once. The temperature

of the furnaces vas veil controlled within ± 5 *C with

time and in space. For sample C, specimens vere individually heat treated in a dynamic vacuum. Both ends

of a specimen of internal-Sn processed sample C vere

extended to a position where the temperature vas kept

below the melting point of Sn in order to avoid the

outflow of molten Sn from the wire. The heat treatment

conditions for samples A, B and C were 700 *C for S6hr,

670 *C for 200hr and 700 *C for 48hr, respectively.

More than half of self reactions vere done in vacuum. At other labs specimens were encapsulated and

heat treated in an argon or a hydrogen atmosphere. The

sample A sample B sample C

Fabrication Method Bronze Bronze Internal Sn

Wire Diameter (mm) 0.8 1.0 0.68

Structure NbTa/CuSn Nb/CuSnTi Nb/Cu/Sn

Cu/non-Cu 0.22 1.68 0.83

Bronze/ Cores 2.8 2.5 3.1

Filament Diam.(ym) 3.6 4.5 2.7

No. Filaments 10,000 5,047 5,550

Heat Treatment 700 C 670 C 700 C 96h 200h 48h

Fig.l Specifications and cross sectional views of

round robin test samples.

:?69

heat treatment temperature was also veil regulated in

the self reaction.

Apparatuses and Measurement Conditions

Informations about measurement apparatuses and ex- perimental conditions at each laboratory vere described

in formatted sheets prepared and distributed prior to the measurements.

At roost of participant labs superconducting solenoid

magnets vere used to generate fields of 8-16.5 Tesla.

At some labs copper solenoid magnets or hybrid type

magnets vere used which can generate fields over 20

Tesla. In any of these magnets rms ripple fields which

may cause heat generation in the specimen, were less

than 0.12 of the field generated. Field determination

vas carried out by means of NMR, Hall probe or rotating

coil magnetometer, with estimated accuracies of 0.12 to 12 .

At most of the labs transistor type power sources

vere used to supply electric currents to the specimen,

rms ripple noises were less than 0.12 of the full

power. The accuracy of current determination vas vithin 0.52 at most of the labs. In order to amplify

the voltage signals generated in the specimen, nano- voltmeters vere used at most of the labs. At some labs

the signals were directly led to high sensitive XY re- corders. The typical level of noises observed varied

from 0.01 uV to 2 wV, depending on the apparatuses

used

The total length of a specimen, the distance between

voltage taps and the distance between a current tap and

its nearest voltage tap were much varied among labs.

The shortest specimen used in a lab had a total length

of about 215 mm and 1.5 turns of winding on the

measurement holder. The longest specimen used vas 2000 cm long and had 30 turns on the holder.

The variety of materials, e.g., fiberglass reinforced plastic (FRP), stainless steel, alumina, hastelloy and brass vere used as the measurement holder. The

Table II. Homogeneity study on sample A at SCK/CEN.

Averages and standard deviations of I 's for fields of 7-10 Tesla.

Field (Tesla) 7 8 9 10

Average (A) 424.5 354.8 296.6 246.4

Std. Dev. (A) 6.0 5.8 5.0 5.8

Std .Dev/Ave(2) 1.5 1.7 1.7 2.4

Table III. Homogeneity study on sample B at Clarendon

Lab. Averages and standard deviations of

I 's for fields of 8-14 Tesla,

c

Field (Tesla) 8 10 12 14

Average (A) 303.8 214.1 150.9 102.3

Std. Dev. (A) 2.5 2.8 1.6 1.0

Std.Dev/Ave(Z) 0.8 1.3 1.0 1.0

81
:ru

holders had ring or bar shaped current terminals of copper. Most of holders had a spiral groove on them.

The specimen was mounted in a groove on the holder vith

both ends soldered to current terminals, and fixed

using a bond such as grease, varnish, epoxy, and

solder. In some cases no bond was used.

The Lorentz force caused by the interaction of the

transport current with the applied field has an essential effect on the I . This force acts as a tensile or compressive stress to specimen when the spiral specimen

generates a central field parallel or antiparallel to the direction of the applied field. At most labs, the

direction of both fields were antiparallel. The effect

of the field direction on the I £ was examined at several labs.

The I was defined at a current where a certain

voltage gradient or a certain resistivity appeared

along the superconducting specimen. Values of I at 5,

10, and 100 iiV/m and at integer numbers of magnetic

fields were requested to be report^ in the_^esent

round robin test; values at 10 * and 10 ft m were

optional

.

The relationship between the voltage V, and the

transport current, I is empirically expressed as V I where the exponent n is nearly constant in the small

voltage region. A larger n corresponds to a 'sharper

transition in the specimen, n values were also

requested to be reported. In the ease where n values were -not xeported by the participant, they were es- timated by using the following relation,

n-l/log I c (100uV/m)/I

c (10uV/m).

Homogeneity Study

As the purpose of this test was the intercoraparison

of results obtained at different labs, it was absolutely important that all the test specimens supplied

should have identical superconducting properties. It is difficult, however, to fabricate a sample with

homogeneous properties along the whole length of a wire

because of the complicated structure of Nb^Sn multifilamentary composite. The homogeneity in superconducting properties was examined on sample A at SCK/CEN

and on sample B at Clarendon Lab.

The homogeneity study at SCK/CEN was performed on 21

specimens of sample A. Specimens were taken from the

various parts of a test wire, each wound on a holder

and heat treated at Rutherford Lab together with those

of central reaction. The I

£ measurement was carried

out at 4.3 K and at 7-10 Tesla.

115 I g data defined at a voltage of 10 uV/m for 21

specimens were analyzed at SCK/CEN. For voltages

larger than lOuV/m specimens showed a tendency to quenching. 15 specimens quenched under a voltage of

lOuV/m. The quench currents were usually lower com- pared to the true I c

’s expected, and were not taken

into account in the statistical treatment of the I data. For each magnetic field averages and standard

deviations of I c ’s were calculated according to the

lOuV/m voltage and listed in Table II. Standard deviations are 2.4? of average of I 's at 10 Tesla and

decrease to 1.52 at 7 Tesla. C

The homogeneity in sample B wa3 examined at Clarendon Lab. 7 specimens were taken from the test wire,

each mounted onto a stainless steel holder, and heat

treated at Rutherford Lab altogether in a furnace. The

I

£ measurements were carried out at 4.2 K and at 7-15

Tesla. Averaged I^'s at a voltage of 10 uV/m and

standard deviations are summarized in Table III.

B (tesla)

Fig. 2 * c

’ s at 10 juV/m as a function of field for

sample A of central reaction; results of participant labs.

Fig. 3 I c

's *t 10 wV/m as a function of field for

sample B of central reaction ; results of participant labs.

82
Crilical

Current

(A)

Fig. 4 ^ c

' s at 10 JuV/m as a function of field for

sample C of central reaction ; results of participant labs.

2371

Standard deviations are within 1.32 at all fields.

Based on these homogeneity studies, it may be concluded that the present test samples A and B are rather

homogeneous in superconducting properties over the

whole length of wire and adequate for the round robin

test. The homogeneity study has not yet been carried

out on sample C.

Results of the Round Robin Test

In Figs 2, 3 and 4 I 's at 10 uV/m are shown as a function of applied field for samples A, B and C,

respectively. The name of labs is coded like a, b, c in no special order. These results were obtained in

the antiparallel self field except for those of labs q and s.

The semilogarithmic plots show that for samples A

and B the logs of the I decreases almost linearly with

increasing field up to about 15 Tesla, and then drops

off more rapidly above this field. I ’s in sample C decrease in a similar way but with a steeper slope, and

start to drop off above about 12 Tesla. The averaged n values generally decrease with increasing magnetic

field and are much more scattered than the I ’s, especially for sample A. c

In Tables IV and V are shown averages and standard

deviations of I £ data and n values obtained at participant labs, respectively. Four extraordinarily outlying data sets were excluded in the calculations because these data sets were obtained in cases where the

specimens stuck to the reaction holders and were likely

damaged in removal or for other similar reasons. There

was no single magnetic field where all labs reported

data on any sample. A future publication comparing all

data sets at a single magnetic field, by means of a Kramer extrapolation method, is planned. Coefficients

of variation of I ’s at 10 juV/m and 12 Tesla are about

7, 6 and 21 2 of the averages for samples A, B and C of

central reaction, respectively, and become larger at higher magnetic fields. These values are appreciably

larger than those obtained in the homogeneity studies

which are 2.4 and 1.3 2 for samples A and B, respectively.

Table IV. Averages and standard deviations of critical currents at 10 aV/m for

samples A, B and C of central reaction.

Field

(Tesla)

sample A sample S sample C

No. Ave.

(A)

cr cr (A)/Ave

(2)

No. Ave

(A) a <j (A)

/Ave

(2)

No. Ave

(A)

cr a

(A)

/Ave

(2)

6 1 544.1 0.0 0.0 4 447.4 18.7 4.2 3 393.8 25.7 6.5

7 4 447.9 26.2 5.9 7 372.2 15.4 4.1 7 309.3 24.3 7.9

8 5 388.7 21.8 5.6 8 315.2 13.0 4.1 6 248 .

20.8 8.4

9 7 327.4 23.3 7.1 10 265.3 11.1 4.2 9 197.3 20.6 10.4

10 9 277.1 13.0 6.5 15 229.1 13.1 5.7 11 150.0 18.5 12.3

11 7 234.4 14.3 6.1 13 195.3 12.6 6.5 13 106.9 16.2 15.1

12 11 195.7 13.6 6.9 17 163.6 10.4 6.3 16 72.5 15.2 20.9

13 10 158.0 14.5 9.2 14 134.6 7.2 5.3 13 45.4 10.7 23.6

14 10 130.5 15.5 11.9 16 114.3 7.8 6.8 13 30.0 8.4 28.0

15 10 106.7 12.9 12.1 12 92.2 5.0 5.4 10 16.9 5.0 29.4

16 6 91.9 10.8 11.8 10 75.8 4.8 6.4 7 8.5 2.6 30.5

17 3 71.8 6.3 8.8 4 57.5 2.5 4.4 2 3.4 0.2 6.4

13 4 57.4 9.0 15.7 7 43.7 2.6 8.4 2 2.1 0.7 33.7

19 3 36.8 5.3 14.4 4 30.6 2.6 8.4 1 1.0 0.0 0.0

20 2 26.4 4.4 16.7 5 19.4 2.7 13.9

No:number of data, <r :standard deviation, a /Ave :coeff icient of variation

83
2372

In Fig. 5 I 's at 10 uV/m and at 12 Tesla re- ported from labs are compared. In cases no I £ values were reported at 12 X, they were estimated

by extrapolations not exceeding 2 T. In the figure,

AI /o' denotes the difference in I between lab and

average normalized to the standard deviation, <r . Labs are arranged in Incremental order of mean laboratory value of Al^/cr . The arrows attached

to some of the data symbols indicate that these data

actually lie either above or below the vertical limits

of the plot. The maximum to minimum spread in I at each lab is almost within 2 times of cr , rather small

compared to the total maximum to minimum spread (6

times of O’ ). These results clearly indicate systematic differences among the labs in the measured I of each sample. These differences may be due to theC

Table V. Averages and standard deviations of n values for samples A, B and C of central reaction.

Field

(Tesla)

sample A sample B sample C

No. Ave

.

a a /Ave

(Z) No. Ave. O* 0" /Ave

(Z) No. Ave cr cr /Ave

(Z)

6 4 37.2 2.4 6.5 3 32.2 2.8 8.5

7 6 39.3 5.0 12.8 7 32.8 7.6 23.3

8 4 78.3 33.9 43.4 8 36.8 5.6 15.2 6 32.5 4.9 14.9

9 5 87.8 38.5 43.8 10 34.2 4.5 13.1 9 28.6 3.5 12.1

10 7 58.0 21.4 36.9 13 32.6 5.5 16.8 11 26.7 5.2 19.3

11 6 50.0 23.3 46.7 13 32.3 8.2 25.3 13 26.8 8.8 32.8

12 9 52.7 18.7 35.5 17 30.3 6.6 21.8 16 21.9 7.7 35.0

13 8 43.9 16.5 37.7 15 28.3 7.5 26.7 13 18.7 7.4 39.7

14 7 35.1 14.8 42.3 16 27.8 • 6.2 22.2 13 14.5 5.1 34.9

15 6 37.3 13.1 35.1 12 28.1 5.1 18.0 10 12.4 3.5 28.5

16 3 33.0 3.6 10.8 10 25.9 4.0 15.4 7 10.5 2.7 25.5

17 4 24.2 5.3 21.9 2 8.8 1.6 17.7

18 1 22 0.0 0.0 7 20.6 4.3 20.9 3 6.3 0.7 10.8

19 4 19.0 3.5 18.2 1 5.2 0.0 0.0

20 5 12.8 2.3 17.8

No:number of data, o’ :standard deviation c /Ave:coef f icient of variation

Fig. 5 Comparison of I *s among participant labs at 12 T. Symbols (O, • ), (A, A ) and ( 0, ) refer to the samples A, B and C (central

reaction, self reaction), respectively.

84
2373

different sample holders, sample handling, the measurement method, aid variations in instrument calibration.

More detailed papers addressing these possible sources

of systematic differences are planned for future publication.

The ratios of I £ (ff) to I ( M) ate shovn in Fig. 6 for sample B. I c (ft) and I (

|> denote the I c

's for

which the self and the applied fields are in parallel

and antiparallel directions, respectively. If the

specimen is tightly fixed on the holder, there should

be no difference between I ( t t) and I (tl), the ratio

of I (ff) to I (tl) being unity. The tightness of the

fixing dependson the bonding material and its amount.

Epoxy resins such as ’stycast' seem to give better

fixing than grease, as is apparently shown in Fig. 6. Some of the curves of the ratio I (t!)/I ( t

drop

below unity at higher fields in Fig. 6. If a specimen

is uniformly deformed by concentric hoop stress, the

ratio will approach unity as the field increases.

Therefore, nonuniform deformation of specimen may

result in the ratio less than unity.

In order to see if the ’self reaction' carried out

at each lab was the exact copy of the 'central

reaction', data of I 's for two reactions were compared. This comparison between self and central reactions did not include all of the available data, but

rather, it only included data from labs that measured

both of their specimens on similar holders. Averaged

I 's of self reacted sample B are nearly the same as tnose of centrally reacted one. However, averaged I 's of self reacted sample A are always smaller than those

of centrally reacted one, while those of self reacted

sample C are smaller at lower fields and larger at higher fields than those of centrally reacted one. These results indicate that I c

's of samples A and C are

more sensitive to the heat treatment condition than

those of sample B. The reaction time for sample C is

the shortest of the three samples and, consequently,

the sensitivity of sample C to both reaction time and

temperature may be greater than for the other samples.

Furthermore, sample C's apparent sensitivity to the

reaction conditions may be due to an inadequate confinement of the internal tin at the specimen ends

during the heat treatment. The coefficient of variation of I of self reacted samples, which may reflect

the difference in reaction conditions among labs, correspond to 4-8 Z of I c almost independent of magnetic

field and sample. The possibility of sample damage incurred during shipping is also a variable in the comparison between the self and central reactions.

Strain Effect

The tensile strain effect on I 's were examined at NBS, Osaka Univ. and Rutherford Lib using the round

robin test samples. Although the measurement details

were different, there was good agreement among three

sets of data. As examples, results for 14 Tesla at KBS

and those for 15 Tesla at Osaka Univ. are compared in

Fig. 7 where I 's have been normalized to the maximum

values. The I c of Nb-Sn conductor increases with increasing tensife strain up to 0.2-0.3Z where the precompressive strain imposed on the Nb-Sn compound is

mostly released. As can be seen in this figure, the

tensile stain sensitivity of I is largest for sample C

and relatively small for sample B. This may be related

partly to the H - of these samples; sample C has a relatively low 6^ of ~ 19 Tesla, while it is ~ 24

Tesla for samples A and B. A more complete comparison

of the strain effect data will be the subject of a future paper.

Effects of the thickness of the FRP holder on the I £ were examined at NBS using specimens of samples B and

B (tesla)

Fig. 6 Ratio, I c (|f)/I CtO vs. field curves for sample

Strain ( 7.

)

Fig. 7 I vs. strain curves for samples A, B and C obtained at NBS and Osaka Univ. I is normalized

to its maximum value.

85
4

C. Specimens were first mounted on thick walled FRP holders. After the I measurement the holders were

bored so as to have tnin walls. It was found that 1^ s for these bored holders with thin walls were slightly

enhanced. The thermal contractions of thick and thin

wall holders are different due to the textured structure of FRP, and estimated to be about 0.42 for thick

wall and about 0.252 for thin wall on cooling from room

temperature to 4.2K. It was concluded that the com- pressive strain in the specimen resulting from the

thermal differential contraction between the FRP holder

and the specimen was responsible for the discrepancy in

I in those specimens mounted on thick and thin wall

holders.

The sample holder material and geometry, in conjunction with the specimen bonding method, can have an effect on the strain state of the specimen and this will

have a significant effect on the results of the I £ measurements. Thus, this could be a major source of

the observed systematic variation in the measured I c 's.

References

[1] H. Jones, "First Meeting of the VAMAS Technical

Working Party on Superconductors and Cryogenic

Structural Materials", Cryogenics , vol. 26, pp. 488-489, August (1986).

[2] L.F. Goodrich and S.L. Bray, "Critical-Current

Measurements of Nb,Sn Superconductors: N3S Contribution to the VAMAS Interlaboratory Comparison",

submitted for publication in Cryogenics

.

Conclusions

Results of the present round robin test may be sum- marized as follows.

a) Homogeneity test on the I showed the present

samples were enough homogeneous along the whole

length of wire to be used in the round robin test.

b) Coefficients of variation of I *s were 72, 62 and

212 at 12 Tesla for samples A, B and C, respectively. They were largest for sample C which was

most sensitive to strain and heat treatment conditions. Coefficient of variation increased at higher

fields for all of samples.

c) n values shoved larger coefficient of variation than

I . n values were rather small in case the specimen

was soldered on a metallic holder.

d) Strain effects on the I of three wires were examined. It was pointed out that the strain in the

specimen was a major origin for the scatter in I£

.

e) Materials of the specimen holder and tightness of

fixing have significant effects on the 1^ through

the strain effect.

Several parameters influencing the I value of Nb^Sn

vires have been extracted and analyzed through the

VAMAS round robin test. However, the present round

robin test was performed with few restrictions on the

measurement method, and more strict determination of

measurement conditions will be required to minimize the

scatter in I and to furnish a really effective standard method of the I £ measurement.

A more complete report on the present round robin

test will be submitted to the steering committee of

VAMAS from the Technical Working Party. A number of

papers addressing specific aspects of this round robin

test are planned for future publication.

Trade names are used in this document to specify the

measurement details. In no instance does this identification imply endorsement by the authors or their

Institutions, nor does it imply that the particular

products are necessarily the best available.

The authors wish to thank all the scientists and

representatives who have been involved in the present

round robin test for their cooperation.

86
APPENDIX I

PROPOSED STUDY ON THE EFFECT OF SAMPLE BONDING TECHNIQUES ON

THE MEASURED CRITICAL CURRENT OF Nb 3 Sn SUPERCONDUCTORS

L. F. Goodrich and S. L. Braypublished in:

Proceedings of the 6^ Japan-US Workshop

on

High-Field Superconducting Materials

and

Standards Procedures for High-Field

Superconducting Materials Testing.

Boulder, Colorado

February 22-24, 1989

Editors: K. Tachikawa, K. Yamafuj i and H. Wada

and

J.W. Ekin and M. Suenaga

pp . 99-108, Japan (1989)

87
Proposed Study on the Effect of Sample Bonding Techniques on

the Measured Critical Current of Nb 3 Sn Superconductors

L. F. Goodrich and S. L. Bray

National Institute of Standards and Technology

Boulder, CO 80303

PREFACE

This proposed study was prompted, in part, by presentations that were

given at the 6th Japan-U.S. Workshop on High Field Superconductors (February

22-24, 1989 Boulder, Colorado) and subsequent discussions. The results of the

recent VAMAS Nb 3 Sn critical-current round robin indicate that increased

consistency in inter laboratory measurements might be achieved by a more

detailed specification of the critical-current measurement technique.

However, a competition exists between the benefit of a rigidly specified

measurement technique, which insures measurement reproducibility, and a less

restrictive technique that would be practical for a greater number of

laboratories. Ideally, the measurement technique should have the least number

of restrictions that is consistent with measurement reproducibility. This

proposal addresses only one aspect of the critical - current measurement

technique, the sample- to-measurement-holder bonding method. The measurement

study proposed herein is typical of research that is routinely performed at

NIST; however, the existing commitment of our available resources prevents us

from undertaking this work now. Also, from a technical perspective, the

critical-current measurement technique proposed here is not one that is

normally used at NIST. Consequently, another laboratory, which normally uses

a measurement technique that is similar to this one, may be better suited to

this research. Dr. Hitoshi Wada of Japan's National Research Institute for

Metals has expressed an interest in conducting a study of this type and this

document was prepared at his request.

88
INTRODUCTION

The results of the VAMAS critical-current (I c ) round robin demonstrated

the existing level of measurement discrepancy among a large number of

laboratories for several different Nb 3 Sn samples. The discrepancy was

significant. A major source of this discrepancy was determined to be varying

degrees of axial sample strain caused by variations in the thermal contraction

of different sample holders.

The sample bonding technique can play an important role in this effect

because the strength of the bonding determines the amount of the sample

holder's thermal contraction that is transmitted to the sample. A relatively

weak bond can inhibit this transmission of thermal contraction but it may also

allow sample motion under the influence of the Lorentz force, which, in turn,

can cause variations in the measured I c

. A high strengch bonding material

used in conjunction with a sample holder whose thermal contraction is

comparable to that of the Nb 3 Sn sample alleviates both of these mechanical

problems but it introduces a practical problem. The commonly used sample

holder material that has satisfactory thermal contraction is a fiberglass - epoxy composite (G-10). The high strength bonding material that is typically

used for G-10 sample holders is a filled epoxy adhesive. The problem with

this method is that the epoxy forms a fairly permanent bond which makes

repeated use of the sample holder difficult.

What is clearly needed is a bonding material that provides adequate

sample support but that is easily removable from the sample holder. The

results of the VAMAS measurements indicate that grease may be adequate for

this purpose as long as all of the I c measurements are made with the Lorentz

force directed inward, toward the sample holder. Prior to accepting grease as

the preferred sample bonding material, a more definitive study of its

suitability should be conducted. A general systematic method for conducting

this study is proposed here.

A second Nb 3 Sn I c round robin where several different measurement

techniques are tested by each participating laboratory in order to determine

the preferred measurement method for Nb 3 Sn is envisioned. This study would

test a number of measurement variables including the sample holder material.

Consequently, a prior determination of the preferred bonding material would

help simplify this subsequent round robin. The number of laboratories

participating in this second round robin might be considerably reduced, from

the first round robin, because of the difficulty involved in complying with

the necessarily restrictive specifications for the various measurement

techniques. However, in the event that the results of the second round robin

resulted in a general consensus regarding which measurement variables must be

specified in order to assure consistent measurements among different

laboratories , a third round robin might be useful for determining the

effectiveness of the measurement method for a larger group of laboratories.

89
GENERAL PROPOSAL

Once again, the objective of this study is to determine the effectiveness

of grease as a sample bonding material in making reproducible I c measurements.

For this proposed study on sample bonding, stainless steel was selected

as the sample holder material, because it represents a worst case condition

with respect to differential thermal contraction between the sample and its

holder. In other words, if grease proves to be adequate as a bonding material

on stainless steel, it should be adequate for other sample holder materials as

well. Although the thermal contraction of G-10 is better matched to that of

the Nb 3 Sn samples, there is no guarantee that G-10 will ultimately be selected

as the preferred sample holder material.

Stainless steel sample holders present some measurement problems that

other materials do not: The stainless steel sample holder provides a shunt

path in which a portion of the metered test current may flow and, thus, it may

artificially increase the measured I c

. The amount of current flowing in this

shunt path depends on several variables including the joint resistance between

the sample and the current contacts, the resistance between the current

contacts and the sample holder, the resistance of the sample holder (which

varies with the sample holder's diameter, wall thickness, and length), and the

voltage across the sample. Based on the resistivity of stainless steel, an

estimate of the current -contact joint resistance, the sample length, and the

electric field criterion, a simple calculation can be made to estimate the

maximum value of the shunt current. This calculation indicates that for most

measurement conditions the shunt current is negligible in comparison with the

sample current; however, in cases the applied magnetic field is high and the

sample current is very low the, shunt current can be significant. Also, the

stainless steel's relatively high thermal capacitance and thermal conductivity

may artificially enhance the sample's stability, particularly in the case of a

soldered sample. Nonetheless, stainless steel is a very commonly used sample

holder material and since it seems to present the most severe test of the

sample - to-holder bond, it was selected for this proposed study.

In general, the proposed method for determining the effectiveness of

grease as a bonding material is to measure the I c of a number of Nb 3 Sn

samples, where some of the samples are mounted to their holders using grease

and some are soldered to their holders, and then compare the variation in the

measured I c for the two sample mounting techniques. For all of the samples,

the reaction and measurement variables should be held constant in order to

isolate the variable of interest, sample bonding.

There are two phases in the study. The first is a necessary but

insufficient test of grease's suitability as a bonding material. In this

phase, all of the mounting variables should be held constant in order to

insure that discrepancies in the I c measurements that are actually due to

variations in the sample mounting technique are not attributed to inadequate

sample bonding. This consideration also requires good homogeneity among the

test samples. In some cases it may be advantageous to measure the same sample

several times while changing the mounting technique to avoid measurement

variations caused by sample inhomogeneity. The mounting technique that is

90
employed for this phase of the study should be designed to optimize the

grease's bonding capability. This phase of the study will thus demonstrate

whether grease is an adequate bonding material under ideal conditions. If,

under these ideal conditions, the grease proves inadequate in producing

consistent measurement results, then it would undoubtedly be inadequate in

actual practice where many different laboratories would be mounting and

measuring the samples. Of course, this assumes that the mounting technique

selected for this study does in fact optimize the grease's bonding capability.

In a sense, the first phase of the study will test the validity of this

assumption

Assuming that the first phase of the study demonstrates that, at least

under tightly controlled conditions, grease is an adequate bonding material,

the second phase of the study should then be conducted. The second phase of

the study addresses the fact that, in actual practice, mounting variables will

exist between different measurement laboratories; to some degree there will

result in variations in the I c measurements. The purpose of the second phase

is to determine which mounting variables have a pronounced effect on the Ic

measurement. The mounting variables that are purposely introduced during the

second phase should be those that are believed to occur most frequently in

practice. In other words, the second phase of the study should simulate the

situation that occurs in an inter laboratory comparative measurement. Once

these variables have been isolated, specifications for sample mounting with

grease can be proposed that address these dominant variables.

In both phases of the study the soldered samples will act as a standard,

of sorts, with which the grease mounted samples will be compared.

Because the solder bond is a very positive one, if the measurement variability

for the grease mounted samples is comparable to that of the soldered samples,

it will indicate that the strength of the grease bond is adequate. Like the

grease mounted samples of the first phase, the soldered samples should be

mounted with as few variables as possible to allow a legitimate comparison

between the two mounting methods. For the same reason, the soldered samples

of phase two should be mounted with the deliberate introduction of mounting

variables, thus simulating the variability that is typical from one laboratory

to another for soldered samples. The results of the phase two study will

demonstrate whether the inclusion of realistic mounting variables causes

greater data scatter for the grease mounted samples than for the soldered

samples. The justification for conducting this study in two phases, rather

than simply conducting phase two of the study, is that the second phase is

considerably more difficult than the first and it need not be conducted if the

first phase is unsuccessful. Equally important is the fact that phase one

directly addresses the effectiveness of grease as a bonding material without

the introduction of other variables.

PROCEDURE

General Philosophy

Because this study will be conducted at another laboratory, it is

impractical to specify all of the details of the sample reaction, mounting,

and measurement procedures. For example, a detailed specification might

91
require that the sample holders have an outside diameter that is slightly

different than that normally used at the testing laboratory, and compliance

with this somewhat arbitrary specification would unnecessarily increase the

required time and cost for this study. Consequently, the philosophy of this

proposal is to state general guidelines for the study while allowing the

measurement laboratory to substantially adhere to its normal practices.

Sample Reaction

The samples should be reacted on well oxidized stainless steel tubes in

order to avoid diffusion bonding between the sample and the tube. When

diffusion bonding occurs, the probability of damaging the sample when

transferring it to the measurement holder is greatly increased. Even in the a

case where diffusion bonding does not occur, extreme care must be exercised in

transferring the sample to avoid damage due to bending strain. The reaction

holders should have a helical groove for defining the sample coil's pitch.

The sample should be well seated in this groove prior to reaction to insure

reproducibility of the coil diameter from sample to sample.

Sample Holders

The stainless steel measurement holders should have the same diameter,

pitch length, and groove geometry as the reaction holders to insure that the

reacted sample can be fitted to the measurement holder with a minimum of

sample deformation. For phase one of the study, the samples should be well

seated in the measurement holder's groove to insure consistent coil diameters

and, thus, consistent bending strain. This is also important for minimizing

the thickness of the bonding material between the sample and holder in the

case of grease mounted samples. Moreover, a thick layer of grease is more

likely to yield under the influence of the Lorentz force than a thin layer,

where the sample is substantially in contact with the stainless steel holder.

A good choice for the holder's groove geometry is a 60° "V" with a fairly

sharp point. The 60° groove provides good sample support in both the radial

and transverse directions. The sharp point of the groove insures that the

conductor contacts the groove at the points of tangency rather than at the

bottom of the groove. In the case of the measurement holder, this should

minimize the thickness of the bonding-material layer between the sample and

the holder. In the case of the reaction holder, it may reduce diffusion

bonding. The groove should be deep enough that the sample contacts the holder

within the groove rather than at the interface formed between the groove and

the holder's surface. This is important for transverse stability of the

sample. One detrimental consequence that results from making the groove too

deep is that the sample will require greater deformation during the transfer

process in order to clear the holder.

Sample Mounting

For the grease mounted samples, the measurement holder should be prepared

by filling the sample groove with grease, avoiding any excess grease outside

the groove. This procedure insures that equal amounts of grease are applied

to all of the samples and that there is not an excess of grease,

92
that might inhibit cooling. Since all of the measurements will be made with

the Lorentz force directed inward, there is limited benefit in having excess

grease on top of the sample.

Before attempting to transfer the sample from the reaction holder, it

should be checked for diffusion bonding to the holder. If there is a

significant amount of bonding, that might result in sample damage, the sample

should be discarded. There are two methods that can be used for removing the

sample from the reaction holder. The first method requires a very smooth

interface between the sample and holder but, when it can be employed, it

allows sample removal with minimal deformation. In this method the sample is

unthreaded from the holder by lightly gripping the sample coil's surface while

turning the holder. Alternatively, the sample coil's diameter can be expanded

by carefully twisting the coil until the holder can be extracted.

Although this second technique results in some sample deformation, it is

the same deformation that must be introduced when the sample is mounted on the

measurement holder. In the case of the grease mounted samples, the sample

cannot be threaded on to the measurement holder because the grease must be

applied to the sample groove prior to mounting the sample. Although the

soldered samples could be threaded onto the measurement holder prior to

soldering them in position, this would introduce an undesirable variable

between the soldered samples and the grease mounted samples. Consequently,

both the soldered and the grease mounted samples should be transferred using

the same technique.

Before soldering the grease mounted samples to both of their current

contacts, the slack in the sample coil should be removed by firmly seating it

in the sample groove. This is easily achieved by first soldering one end of

the sample coil to its current contact and then, starting at that contact,

pressing the sample into its groove in a continuous motion that results in all

of the coil's slack being transferred to the opposite cunent contact. The

sample can then be fixed in this taut position by soldering it to the current

contact. The grease that is pressed out of the groove along the sample's

sides during this operation increases its transverse stability and should not

be removed.

In the case of the solder mounted samples, slack should be removed prior

to soldering and then mechanical retainers should be used at both ends of the

sample to hold it in the taught position during the soldering operation. To

insure reproducibility, the ends of the sample should not be allowed to move

during soldering. The sample holders should be pre- tinned prior to mounting

the sample to simplify the soldering operation and to, thus, reduce the

probability of sample damage.

An attempt to consistently regulate the shunt current that flows in the

stainless steel holder should be made. If the current contacts are

electrically isolated from the sample holder (with the exception of the sample

current path) , the shunt current should be fairly consistent from sample to

sample. However, if the contacts are not isolated, then the shunt current

will depend on, among other factors, the resistance of the solder joints at

the current contacts and the resistance of the interface between

93
the contacts and the holder. Consequently, the length of the contact joint

should be the same for all of the samples and the resistance of the contactto-holder interface should be constant.

Current Reversal

Grease will undoubtedly present a limitation in the I c measurements that

other more permanent bonding materials do not. Its relatively low tensile

strength will probably disallow its effective use for I c measurements where

the Lorentz force on the sample coil is directed outward. This, in turn,

disallows sample current reversal for testing the effect of the sample coils

self -generated magnetic field on the measured I c

.

Statistics, Field History, and Repeat Determinations

The number of each sample type, grease mounted or soldered, that need be

measured for this study cannot be specified beforehand. It is important that

the number of measurements is large enough to be statistically significant,

but the number of measurements required to meet this criterion depends on the

observed measurement variability for each of the sample types. Because the

Lorentz force may play an important role in the I c measurements of the grease

mounted samples, the magnetic field history of the test sample may effect the

measurements. For example, if the first I c measurement is made at the field

where the Lorentz force on the sample is a maximum, the subsequent

measurements might be different than for another field sequence.

Consequently, the measurement sequence should be carefully considered and

recorded. Also, if the Lorentz force causes irreversible sample motion,

subsequent measurements at the same magnetic field may have different results.

Repeat determinations of the I c should be used to detect mechanical

deterioration of the sample bond.

Grease Mounting Variables

The first variation on the phase-one measurements that should be studied

is the effect of eliminating the bonding material altogether the effect of the

bonding material can be isolated. By measuring the I c of several samples that

are tightly mounted on grooved holders without any bonding material and

comparing the resultant measurement variability with that of bonded samples.

In the second phase of the study, common mounting variables should be

introduced to the measurements. For the grease mounted samples, a common

variable might be the relative tightness of the sample coil on the measurement

holder. This variable is important for two reasons. First, the level of

bending strain in the sample will depend on the coil diameter and, thus, it

will depend on how closely the sample is fit to the holder. Also, the

thickness of the grease between the sample and the holder will vary with the

sample tightness. This might lead to variations in the rigidity of the sample

under the influence of the Lorentz force.

The difficulty in studying the effect of this variable is in measuring

the sample tightness. One method for quantifying the sample tightness is to

place the sample on the holder, without using any grease, mark the position of

94
the sample ends on the holder, remove the sample from the holder, apply grease

to the holder, remount the sample in a slack position, and measure the

distance from the ends of the sample to the marks on the holder. This

technique should result in a reproducible level of slack in the sample coil.

It is important not to tighten the sample beyond the position in which it will

ultimately be tested because some of the grease may be irreversibly extruded

from the groove and the bond strength might suffer. By measuring the I c as a

function of sample tightness the sensitivity of these measurements to the

grease layer thickness and the sample bending strain can be studied.

Another mounting variable that exists at different laboratories is the

geometry of the sample holder groove. This geometry may effect the mechanical

stability of the grease mounted samples when under the influence of the

Lorentz force. It is impractical at this point to measure the effect of a

wide range of different groove geometries on the measured I c ; however, there

is a limiting case that should be measured. A holder having no groove at all

should represent the lower limit of transverse sample stability. By studying

this extreme case the importance of providing transverse stability, by means

of a sample groove, can be determined. In preparing a smooth sample holder

for this study it is important to make its outside diameter such that the

resulting diameter of the mounted sample coil is equal to that of the phase - one samples that are mounted on grooved holders. This will assure that the

bending strain is not an inadvertent additional variable in this study.

In the phase -one measurements, it is specified that all the mounting

grease be applied to the sample groove before transferring the sample to the

measurement holder, thus controlling the quantity and position of the grease.

In practice, the top of the sample is often coated with grease after it is

transferred to the measurement holder. This procedure may effect the measured

I c in several ways. The additional grease may increase the mechanical

coupling between the sample and holder. Consequently, more of the

differential thermal contraction between the sample and holder may be

transmitted to the sample, thus increasing the level of compressive strain.

The increased sample- to -holder coupling may also make the sample more stable

under the influence of the Lorentz force. Finally, if a heavy overlayment of

grease is applied to the sample it may inhibit cooling and reduce the measured

I c

. To test these effects, samples should be measured with several levels of

grease overlayment.

Different laboratories most likely cool their samples at significantly

different rates; in the limiting cases, some may slowly cool the sample in

helium vapor while others may insert the sample directly in a liquid helium

bath (or perhaps liquid nitrogen) . This variable may affect the measured I c of grease mounted samples. The grease bonding material has virtually no

mechanical strength until it is cooled below its freezing temperature.

Consequently, the strength of the coupling between the sample and its holder

is changing during the cooling process while differential thermal contraction

between the sample and holder is occurring. This fact may result in different

mechanical states for samples cooled at different rates. If this variable

does have an effect on the measured I c , it is probably greatest when a thick

overlayment of grease is applied to the sample. Aside from providing a

greater potential for a strong coupling between the sample and holder , a thick

95
layer of grease also creates a thermal insulation layer between the holder and

the helium bath that might increase the difference in the cooling rates for

the various components of the sample-holder assembly. To test the effect of

the sample cooling rate, at least two samples should be measured; both should

have a thick grease overlayment and one should be slowly cooled while the

other should be inserted directly in a liquid helium bath.

Grease is a generic term that applies to a very broad range of heavy

lubricants that certainly have significantly different physical properties.

Two such materials that have been used for sample bonding materials are

petroleum jelly and vacuum grease. Undoubtedly, other materials have been, or

might be, used for this purpose. The variation in their physical properties

might be a significant interlaboratory measurement variable. To test the

effect of this variable on the I c measurement, a representative group of

greases that are used for sample mounting should be assembled and comparative

I c measurements should be made using these various greases.

Solder Mounting Variables

The significant mounting variables associated with soldered samples are

presumably fewer than those for grease mounted samples. This is one of the

benefits of using soldered samples as the comparative standard for this study.

However, to legitimately compare soldered and grease mounted samples in the

second phase of this study, it is necessary to simulate the typical

variability found in soldered samples that are prepared at different

laboratories. Assuming that the soldered samples are carefully handled to

avoid damage, the dominant mounting variable is probably the strain state of

the conductor, which may vary depending on the soldering technique that is

employed.

The technique that has been described for testing the effect of sampleto-holder tightness for grease mounted samples is also applicable to soldered

samples. Since the thickness of the solder layer between the sample and

holder is probably of no consequence, this test would simply measure the

effect of varying the sample's bending strain. Another soldering technique

that may be used and that might result in a different strain state for the

sample is to simply start at one end of the sample and solder it along its

length without first anchoring its ends to the current contacts. Several

samples should be prepared and measured using different soldering techniques

in order to test the variation of solder mounted samples.

Concluding Remarks

This document provides general guidelines for studying the suitability of

grease as a bonding material in Nb3Sn I c measurements. The details of the

study will undoubtedly require adjustments based on intermediate results. For

example, if the initial phase-one measurements are not adequately consistent,

the sample mounting technique will require adjustment and further testing

before the second phase of the study can begin. In the event that these

measurements indicate that grease is an adequate bonding material for the

particular conductor and sample holder that is tested, this does not mean that

it will be adequate for all conductors or sample holders. Larger conductors

96
will be subjected to greater Lorentz forces and the grease may be incapable of

constraining these forces. Also, the grease's adhesion to other sample holder

materials may be different than it is for stainless steel. Finally, the

second phase of this study is important in learning about the effect of

various mounting variables but it cannot be fully conclusive because it will

not address all of the possible interlaboratory variables. However, presuming

that the phase one measurements are satisfactory, the mounting technique could

be tightly specified to insure measurement consistency.

97
fMHHHHMHUHBNHni
APPENDIX J

CURRENT RIPPLE EFFECT ON SUPERCONDUCTIVE d.c. CRITICAL CURRENT MEASUREMENTS

L. F. Goodrich and S. L. Bray

Cryogenics 28, pp. 737-743 (1988)

99
Current ripple effect on superconductive d.c.

critical current measurements*

L. F. Goodrich and S. L Bray

Electromagnetic Technology Division, National Bureau of Standards, Boulder, CO 80303,

USA

Received 70 May 7988

The effect of current ripple or noise on d.c. critical current measurements was systematically

studied. Measurements were made on multifilamentary Nb-Ti superconductor. A low-noise,

battery-powered current supply was required in this study in order to make the pure d.c.

critical current measurements. Also, an electronic circuit that stimulates a superconductor's

general current-voltage characteristic was developed and used as an analysis tool. In order

to make critical current measurements in which current ripple was present, the battery supply

was modified to allow the introduction of controlled amounts of a.c. ripple. In general, ripple

in a current supply becomes more significant in current supplies rated above 500 A because

effective filtering is difficult. The effect of current ripple is a reduction in the measured d.c.

critical current; however, ripple of sufficient amplitude can result in arbitrary measurement

results. The results of this work are general and quantitatively applicable to the evaluation

of critical current data and measurement systems. A theoretical model was developed to

further support and explain the ripple effect. An unexpected benefit of this work was a more

precise method for general critical current data acquisition. Problems common to all large

conductor critical current measurements are discussed.

Keywords: critical currents; superconductors; Nb-Ti

The presence of current ripple (any periodic departure

from d.c. output level) reduces the measured d.c. critical current, 7 C (see Reference 1). In addition to quantifying

this effect, an attempt was made to determine its source. Several possibilities are discussed. A recent literature

search indicates that this is the first documented study

of this type. The I c measurements reported here were made using a straight sample geometry and a radial- access magnet.

Experimental apparatus and procedure

Battery powered current supply

A battery current supply was modified to allow the

creation of current ripple with variable frequency and

amplitude2. The maximum d.c. output of this supply is 1000 A. The maximum tested frequency, a.c. amplitude,

and d.c. bias for this supply are 10 kHz, 170 A, and 400 A,

respectively (these values were not achieved simultaneously). The performance of the supply as a source of a.c. current at several frequencies was evaluated by measuring

the peak amplitude and RMS values of the output current. To judge the shapes of the waveform, the ratio of the

peak amplitude to the RMS value was compared to the

theoretical value for the given waveform (1.73 for a

‘Official contribution of the National Bureau of Standards, not

subject to copyright

0011 2275/83/110737 07 S03 00

< 1988 Eutterworth & Co (Publishers) Lid

triangular wave and 1.41 for a sinusoid). For both

waveforms and all frequencies (60 Hz triangle; 60, 360,

and 1000 Hz sinusoids) the variation between the theoretical and measured ratios increased with increasing a.c. amplitude. For I c measurements at 8 T, this variation was nearly independent of frequency and had a maximum

value of less than 12% at the maximum tested a.c. amplitude of 70 A. At amplitudes below 40 A, the max- imum variation was less than 5%.

Sample and instrumentation

The sample measured in this experiment was a commercial multifilamentary Nb-Ti superconductor with a diameter of 0.65 mm. The filament twist pitch was 0.8 twist cm

-1

. There were 6006, 5 pm diameter filaments

with a Nb diffusion barrier around each of the filaments.

The sample was contacted with several sets of twisted-pair

voltage taps. This allowed simultaneous acquisition of

data with several voltmeters to determine any variation

in the observed effect, any difference in voltmeters, and

any effect due to voltage-tap geometry. One voltage-tap

pair was connected to the sample in such a way that the

mutual inductance between the voltage-tap pair and

sample was enhanced. It spanned the same segment of

the sample as another lower mutual-inductance pair. Two

other voltage-tap pairs had their wires connected near

the sample, and only one of the wires was connected to the sample (referred to as null tap pairs). This provided

100

Cryogenics 1988 Vol 28 November 737
Current ripple effect on critical current measurements: L.F. Goodrich and S.L. Bray

pairs that should not have any non-inductive voltage.

One of these pairs had low and the other had high mutual

inductance. This allowed for a determination of how the

voltmeters measure a purely inductive voltage.

The sample voltage has several components. For this work, the intrinsic d.c. transport V-I characteristic of the

superconductor was assumed to have the voltage proportional to the current raised to the power n, where n

for the tested sample was about 30 (see Reference 1). There might be a slight current transfer voltage but

otherwise the non-inductive component of the voltage

should be zero, within measurement limits, until the

current approaches the 7 C

. Here, the highly current

dependent flux-flow voltage occurs. The unavoidable

loop made by the pair of voltage taps and the segment

of the sample between them, will give an inductive voltage

proportional to the time derivative of the sample current. There will be a highly asymmetric voltage signal due to periodic entry of the peak current into the flux-flow

region of the superconductor at high enough bias currents.

a bias voltage and then amplified by OA2. The adjustable

bias voltage produced by OA 1 simulates the 7 C

. No output

occurs until the input voltage has risen above the bias

voltage. The gain of the amplifier may be varied and

simulates the n of a superconductor. The voltage output

of this amplifier stage drives the base of transistor T 1. A

non-linear current results at the emitter of the transistor.

This non-linear current results from the combination of

the diode (Dl) and output transistor (Tl). The current

passes through two shunt resistors that provide the

circuit’s output signals. Typically, the nanovoltmeter to be tested is connected to the 0.01 Q resistor and a recording instrument is connected to the 100Q resistor. Another channel of the recording instrument is connected

to the analog output of the nanovoltmeter. These two

channels are then compared and the nanovoltmeter may

be thus characterized under conditions similar to an I c measurement. The 0.01 Q series resistor is externally

connected to the remainder of the simulator circuit to allow thermal stabilization of the nanovoltmeter to

resistor connection.

Superconductor simulator

A simple electronic circuit was designed to simulate the

intrinsic V-I characteristic of a superconductor. This

circuit was used to characterize the response of a nanovoltmeter when the input voltage is a highly asymmetric

periodic voltage which results from passing a d.c.-biased

a.c. current through a superconductor. A more general

application of this circuit had been to aid in the

development and testing of critical current data acquisition systems. This circuit was developed based on an idea

by Ekin and Brauch of NBS. The original idea used a diode in series with a current sensing resistor. The

combination was placed in parallel with a high current

shunt resistor. This passive circuit gives a non-linear V-I

characteristic. Unfortunately, the curve is not very sharp

(n % 10) and it is not adjustable. An active simulator

circuit is shown in Figure 1. The input of the simulator

circuit comes from a large shunt connected in series with

the current supply. This input voltage is compared with

Figure 1 Schematic diagram of an electronic circuit that simulates

the intrinsic voltage current characteristic of a superconductor

738 Cryogenics 1988 Vol 28 November

Data acquisition

Two types of voltmeter u'ere used to measure the

sample voltage. One was an analog nanovoltmeter and

the other was a digital nanovoltmeter. Both were direct

current meters with an amplified analog output, which was fed to a digital processing oscilloscope. The voltage

signal of a current sensing resistor was split, amplified,

and fed into tw'o channels of the digital processing

oscilloscope. One of these current signals was filtered to give the d.c. component and the other was unfiltered

(passed wide band). A total of eight channels (two current

and six voltage) were used to acquire data simultaneously.

The sample voltage and current w’ere measured at relatively fixed current ripple and many d.c. bias currents

(point method). These points along the voltage-current

(V-I) curve were analysed to determine the 7 C

. This

procedure is different from the more typical procedure

where the current is ramped continuously with time while

the voltage and current are recorded. Ramping the

current was not an option in this experiment because the

a.c. and d.c. components of the current had to be

accurately separated and the steady state voltage

measured. The point method has some general advantages over the ramp method. The point method allows

for a longer time averaged voltage to be measured (if voltage noise is a problem) without greatly increasing the

time required for taking a complete V-I curve. This is achieved by selecting a distribution of current bias points

that results in a more precise measurement. This requires

taking a few widely spaced points in the low current portion of the curve to establish the base line and then

taking a progressively closer-spaced point distribution as the current increases toward 7 C

. The point method greatly

reduces the effect of inductive voltage and sample motion.

The only disadvantages are that it takes longer because

the voltage has to settle for each data point and the

general shape of the curve has to be known a priori. The

7 C was defined as the d.c. current at which the measured

electric field strength rose to 0.2 /A7 cm' 1

. The precision

of the 7 C measurement was about ± 1 %.

101
Current ripple effect on critical current measurements: L.F. Goodrich and S.L. Bray

Results

Voltmeter tests

In order to develop a quantitative understanding of the

nanovoltmeter's input-output relationship for a.c. voltage inputs and thus determine the effect on the measured

Je, two a.c. tests were conducted on both the digital and

analog nanovoltmeters. The first test (symmetric test)

consisted of applying sinusoidal voltage signals of various

amplitudes and frequencies to the nanovoltmeter input

and measuring the resulting output signal. A.c. signals of

appropriate amplitudes were obtained by dividing the

output voltage of an a.c. signal generator. The objective

of the second test (asymmetric test) was to determine the

response of the nanovoltmeters to the highly asymmetric

a.c. voltage that is characteristic of a superconductor

carrying a sinusoidal current. The superconductor simulator circuit was employed to generate the required

asymmetric nanovoltmeter input signals.

The results of the symmetric tests were considerably

different for the digital and analog meters. In both cases, the d.c. output of the meters increased with increasing

a.c. input amplitude. However, the output signal from

the analog meter was approximately an order of magnitude greater than that of the digital meter for a given a.c. input. Defining the a.c. rejection (ACR) of the nanovoltmeter as 1 minus the ratio of the d.c. output voltage

to the peak a.c. input voltage provides a useful measure

of a meter’s relative tolerance of a.c. input voltages. This

is not the common mode rejection ratio (CMRR)3

. For

the analog meter, a 60 Hz, 500 pV peak a.c. signal resulted

in a d.c. output of 5.2 pV, or an ACR of 99.0%. The same

input signal applied to the digital meter resulted in a d.c. output of 0.66 /iV, or an ACR of 99.9%. This input signal was on the high end of those encountered in this ripple

experiment, indicating that a detectable artificial voltage

may be observed for the analog meter. For a voltage-tap

pair with a 1 cm separation, a typical mutual inductance was about 2 nH (53 pV peak for 60 Hz and 70 A amplitude; 880 pV peak for 1000 Hz and 70 A amplitude). The

mutual inductance of the high inductance taps was about

6 nH.

The dependence of the d.c. output on the input signal’s

frequency (60, 360. and 1000 Hz) was also investigated.

Again, there was a significant difference between the

response of the analog and digital meters. Although the

d.c output of both meters did not show monotonic

dependence on the input signal’s frequency, the analog

meter showed a considerably stronger frequency dependence. For the analog meter, a 250 pV peak a.c. signal

at 60, 360, and 1000 Hz resulted in a d.c output of 0.8,

3.5, and 0.5 pV, respectively. The same input signals

applied to the digital meter resulted in a d.c. output

of 0.14, 0.16, and 0.10/iV, respectively. In the case of the asymmetric test, the behaviour of the

two meters was essentially the same. A broad range of

d.c. biased sinusoidal signals was applied to the input of

the simulator circuit resulting in a large variety of typical

asymmetric nanovoltmeter input signals. As before, three

a.c. frequencies (60, 360, and 1000 Hz) were tested. In order to determine the nanovoltmeters’ responses to these

input signals, the signals were observed and analyses on

a digital processing oscilloscope. The peak-to-pcak, timeaveraged, and RMS values of the input signals were thus

determined and compared with the output of the nanovoltmeters. The results were that, regardless of the meter

being tested or the character of the input signal (fre- quency, d.c. bias level, and a.c. amplitude), the nanovolt

meter’s output signal was always within 5% of the

time-averaged (d.c.) value of its input signal. In other

words, the nanovoltmeter behaves like a very low fre- quency, low-pass filter, which results in an output from

the nanovoltmeter that is a time average of its input

signal. This behaviour is consistent with the symmetric

test where the meters were shown to measure the time

average of their input signals.

Theoretical results

The motivation for modelling this effect is to help

understand what is being measured and how to generalize

these results to different samples, ripple, and voltmeters.

It will become clear that the shape of the V-l curve and

the amplitude of the ripple will change the magnitude of

this effect. To extend the d.c. I c criterion to a general

current waveform in a way consistent with the design

limits of a superconductor application, the measured

voltage criterion should reflect the ohmic power loss within the superconductor. Differentiation between voltage and electric field strength criteria will be dropped in

this section.

It is instructive to consider the power dissipated in a normal conductor when the conductor is carrying d.c.- biased a.c. current. A normal conductor is similar to a superconductor with an n of 1 and, thus, would have a

finite critical current, based on a voltage criterion. The

magnitude of the effect of a.c. ripple on the measurement

of d.c. 7 C can be calculated for this case. For the purpose

of this argument, start with the definition of 7 C as the

RMS value of the total current (a.c. and d.c.) at which

the RMS of the total voltage equals some criterion. This

would be correct for a normal, ohmic sample because it is analogous to the derivation of a.c. power. For sinusoidal waveforms, the time average of the instantaneous

power is the product of the RMS of the current and the

RMS of the resistive voltage. However, measurements of

this voltage would require a complex voltmeter that could

separate the resistive from the inductive voltages and be

capable of measuring the RMS of the total (a.c. and d.c.) resistive voltage. Based on this definition of 7 C and value

of n equal to 1, the measured d.c. 7 C would be equal to the square root of: the 7C without ripple squared minus

the RMS of the ripple squared. For example, the power

of 99.9 A d.c. plus 5 A RMS a.c. is equal to the power of

100 A d.c. [99.9 = ((100

2 -(5)2 )*]. The two components of

current add in quadrature, thus for a low percentage

ripple the effect on the d.c value is very small, for n = 1. A d.c. voltmeter will measure the time average of the

voltage, which is close to the correct quantity to measure

for a superconductor with finite n. Consider the case where the d.c. bias current is well below the flux-flow

region and the a.c. amplitude is just large enough to put

the total current into the flux-flow region. The instantaneous voltage signal will be zero except at the peak of

the current when it will look like a periodic delta function.

Considering the a.c. power analogy, the current is fairly

constant for the period of non-zero voltage and the time

average power becomes the product of the time average

of the two. Thus the d.c. voltmeter is the proper instruCryogenics 1988 Vol 28 November 739

102
Current ripple effect on critical current measurements: L.F. Goodrich and S.L. Bray

ment to use based on the definition of 7 C that relates to the average power. It can be shown that for n = 2 and a

sinusoidal waveform, the time average of the voltage gives

the same d.c. value of I c as the time-averaged power for the case where n = 1. In general, the difference between

the time-averaged voltage and the time average power is a difference of 1 in the exponent. For an n of infinity,

the d.c. voltmeter will give a reduction in the measured

7C that is equal to the a.c. amplitude. This sets the limits

for the effect and these are consistent with the experimental data. Current ripple should be considered as a percentage of 7C and this in combination with n can be

used to estimate the effect. Calculating the time average of the instantaneous,

resistive voltage would model the measurement of d.c. critical current with ripple. This can be done by starting

with the approximate voltage-current relationship.

v= Kd/ijr

where: 70

is a reference critical current at a voltage

criterion of V0: V is the instantaneous voltage: 7 is the

instantaneous current: and n reflects the shape of the curve with typical values from 20 to 60 (a higher number

is a sharper transition). The time average of this voltage,

Vr is ta* °

IKKi.)"]

* i

J 0 r dr

where the period was set to 1 s(co = 2n s~ *) in general. For

a sinusoidal and d.c. current, the instantaneous current can be expressed as

7" = [i + a sin(2:rr)]

n = i"[l + (a/i) sin(27rr)]"

where i is the d.c. component of the current and a is the

amplitude of the sinusoidal component. Thus, the timeaveraged voltage becomes

Ka = KUJIoT [1 + (a/i) sin(27rr)]

n dr

AC Amplitude (%)

Figure 2 Model curves of the percent /c reduction versus the percent a.c. amplitude for several values of n

n approaches infinity, and these plots are consistent with

this limit. Consider the physical meaning of a point on

this limit line where the ripple amplitude is 100%; at this point, the 7C reduction is 100%. The amplitude of the

ripple is equal to the measured d.c. critical current with

ripple present (

= i c) and 70 = 2/c

.

Comparison on model and experimental data

The 7 C for increasing d.c. bias current was measured as a function of a.c. amplitude and frequency using the two

orK = KU/I.YG

where G is defined as the value of the integral and = V0 when 7" = i"G. Also, if a d.c. instrument is used to measure the sample current, i will always be the measured

quantity regardless of the a.c. and the d.c. current levels. Based on these definitions, 70 is the d.c. critical current without ripple and the value of i, where the sample voltage

equals the 7e criterion, is the measured d.c. critical current with ripple present, ic

. In order to keep the equations

based on measured quantities instead of theoretical

quantities, such as 7 0, the percent difference, D, is defined

as the difference between 70 and i c normalized to ic D = 100 (70 - ic

)/i c = 100 (G11 n] - 1) This definition is consistent with a/ic ratio in the integral

(the ripple is relative to the measured quantity i c). The integral G was determined by numerical integration for a number of values of n and a/ic ratios. These

results are shown for two different ranges of current ripple

in Figures 2 and 3. The x axis is percent ripple amplitude

(percent of a/ij and the y

axis is the percent reduction

In measured critical current. D. The line of 7 C reduction

equal to the a.c. amplitude was predicted as the limit as

Figure 3 Model curves of the percent /c reduction versus the percent a.c. amplitude for several values of n

740 Cryogenics 1988 Vol 28 November

103
Current ripple effect on critical current measurements: L.F. Goodrich and S.L. Bray

different types of voltmeters and various pairs of voltage

taps. The reduction in the measured d.c. I c versus a.c. amplitude is shown in Figures 4, 5, 6 and 7. These data were taken with a magnetic field of 8 T where the I c without ripple was 140 A. The symbols indicate actual

data points and the solid lines are the model curves. The

a.c. amplitude is the peak value of the a.c. component of

the current waveform (not peak-to-peak or peak of total current). A reference line (n = oc) was drawn for an / c reduction equal to the a.c. amplitude. For all frequencies

and waveforms, these data are just below this reference

line indicating that the effect is to reduce the measured

Ic by slightly less than the a.c. amplitude. The — 60 Hz

frequency was labelled negative to indicate that it was a

triangular waveform at 60 Hz rather than a sinusoidal

waveform. At all of the other frequencies, the waveform

was sinusoidal with very little distortion. The zero

frequency data had only the intrinsic ripple of the supply,

which was about 0.3 A in this operation. The triangular

waveform was tested to determine whether the detailed

shape of the ripple and the relationship of RMS to peak

changed the measured values. For cases where the

comparison of two different current waveforms is based on their having equal amplitudes rather than equal RMS

values, these data suggest that the detailed shape does

not significantly effect the measured 7 C

. The a.c. amplitude

is the relevant quantity for scaling the effect on the

measured d.c. I c

. The data shown in Figures 4 and 5 were taken using

the analog nanovoltmeter while the data shown in Figures

6 and 7 were taken using the digital meter. The 60 Hz

data of Figures 6 and 7 were measured on the identical

pair of voltage taps as the 60 Hz data of Figures 4 and

AC Amplitude (%)

Figure 4 Percentage /c reduction versus the percentage a.c. amplitude. Model curves are shown for several values of n. Experi- mental data taken with an analog voltmeter are shown at several

frequencies. The —60 Hz frequency was labelled negative to indi- cate that it was a triangular waveform at 60 Hz rather than a sinusoidal waveform. Frequency (Hz): O. 0; D, 60; A. —60; V,

360; <\ 1000

Figure 5 Percentage /c reduction versus the percentage a.c. amplitude. Model curves are shown for several values of n. Experi- mental d.jta taken with an analog voltmeter are shown at several frequencies. For key to symbols, see Figure 4 caption

5. The digital meter data, at frequencies other than 60 Hz, were measured simultaneously with the corresponding

analog meter data but with another pair of voltage taps

that spanned the same segment of the sample. This other

pair of voltage taps had a higher mutual inductance but

the results for both pairs of taps were within the experimental limits. The results measured by the two voltmeters were very close. This indicates that the measured d.c.

AC Amplitude (%)

Figure 6 Percentage lc reduction versus the percentage a c. amplitude. Model curves are shown for several values of n. Experi- mental data taken with a digital voltmeter are shown at several frequencies. For key to symbols, see Figure 4 caption

Cryogenics 1988 Vol 28 November 741

104
Current ripple effect on critical current measurements: L.F. Goodrich and S.L. Bray

Figure 7 Percentage /c reduction versus the percentage a.c. amplitude. Model curves are shown for several values of n. Experimental data taken with a digital voltmeter are shown at several frequencies. For key to symbols, see Figure 4 caption

value of Ic does not depend on the voltmeter used. There

were, however, other differences between these two voltmeters that will be discussed later. The effect of current

ripple was measured for a number of sample segments

and each segment gave similar results. The experimental data fall between the n = 20 and the

n=100 model curves, within experimental error. The

measured value of n was about 30. A model curve with

n = 30 would be a lower limit to the effect. These

experimental values indicate that the effective n may be

higher because of the range of electric fields that occur during the peak current (n is a weak function of £) or because sample heating increases the value of n. The

model and data are in close agreement, within ±2%

for ripple below 20% and within ± 10% for ripple up

to 80% using n = 30.

The model curves can be used to estimate the effect of

current ripple on the measured d.c. critical current. For

approximating the effect of commercial power-supply

ripple on critical current measurements, the curves of

Figures 4 and 6 may be the most useful because of their

lower range. Both ripple and the reduction in critical current can be generalized to a percentage effect. Heating

through a.c. losses may increase the size of this effect at higher amplitudes depending upon the sample geometry.

Voltmeter effects will also become more significant at higher amplitudes.

Discussion

Artificial current transfer voltage

The analysed data from both the digital and the analog

voltmeters gave the same values of / c

. However, some

/ c data taken with the analog nanovoltmeter indicated

the presence of a current transfer voltage 1

. This was not

742 Cryogenics 1988 Vol 28 November

true of the data with the digital meter. It is believed that

this behaviour is simply an artifact of the measurement

system and not the result of a true current transfer voltage.

Two characteristics of the measurement system may

explain this effect. One concerns the sample current supply and the other concerns the nanovoltmeter. First, at high a.c. amplitudes the sample current supply’s a.c. amplitude increases with increasing d.c. bias level which

results in an increasing inductive voltage at the sample

voltage taps. Second, as was determined in the symmetric

voltage test, the nanovoltmeters have a small d.c. voltage

output when subjected to an a.c. input, and this d.c. voltage increases with increasing a.c. amplitude. Consequently, during a cycle of data acquisition, the nanovoltmeter has an increasing d.c. voltage output which is interpreted by the data analysis program as current

transfer voltage. The digital meter does not indicate this

artificial current transfer voltage because of its considerably greater rejection of symmetric input signals.

Finally, good agreement between the data taken by the

two meters is achieved by simply subtracting the artificial voltage from the analog meter data. When this current

transfer correction is large, significant errors in the

determination of the 7 C can result. This is probably the reason that the high amplitude analog voltmeter data at 360 Hz (worst case ACR) was out of line with the other

frequency data.

A.c. voltage threshold

Although both analog and digital nanovoltmeters have

demonstrated predictable input-output behaviour with

respect to low amplitude a.c. input signals, there is an

apparent a.c. amplitude threshold above which their

behaviour becomes erratic. This erratic behaviour varied

from a random output to a very low frequency (less than

1 Hz) and high amplitude output. These outputs could

be positive or negative. Both meters display this erratic

behaviour but the a.c. amplitude threshold for the digital

meter is approximately 10 mV, whereas the threshold for

the analog meter is approximately 1 mV. These results are similar to the symmetric test where the digital meter was shown to have a considerably greater tolerance for

a.c. input signals.

Null taps and passive filtering

Data taken on the null tap pairs (described above) were

used to determine the voltmeter’s response to the symmetric voltage signals during the experiment. These data were consistent with the symmetric test, the artificial current transfer voltage, and the erratic behaviour, all of

which were presented above. This is further evidence for

the assertion that these effects are caused by the a.c. inductive voltage. A passive RC filter was inserted at the

input of the nanovoltmeter to reduce the a.c. amplitude

of the voltage signal. This did reduce the artificial current

transfer voltage, and it increased the threshold of the

erratic behaviour. However, the passive filter introduced more voltage noise. A passive filter may be an option for some experiments. Another possible method of reducing

the inductive voltage is to balance the inductance by

opening an opposite area within the twisted pair. This

105
Current ripple effect on critical current measurements: L.F Goodrich and S.L. Bray

area could be adjusted with a high frequency signal during

sample preparation.

Transient losses

The self-field loss is the primary type of transient loss

present in the case of sample current ripple for this sample

geometry4

. The only effect of transient loss on the 7 C measurement is sample heating, which may raise the

temperature of the sample depending upon cooling

conditions. The power dissipation is a strong function of

frequency, amplitude, and current density. Computed

self-field loss (for the 30 cm sample) at 8 T and the highest

a.c. current amplitude was 3 mW at 60 Hz (50 mW at 1000 Hz). The effect of current ripple was also measured

at 5T with a.c. amplitudes up to about 170 A. The

difference in the power between 5 and 8 T was a factor of

6 for comparable amplitudes and frequencies. There was

a slight frequency and amplitude dependence for the 5 T

data (higher frequency or amplitude reduced I c by more),

but the results were still at or below the reference line,

reduction less than or equal to the a.c. amplitude. The

lack of a strong frequency, amplitude, and current density

dependence of the measured reduction of Ic indicates that

sample heating due to transient losses was not significant

for this experiment.

Other sample geometries might have larger transient

losses due to current ripple. A change in the sample

geometry was made to examine this possibility. The

sample was centred in a holder, and four conductors

(symmetrically located on a 2 cm diameter circle) were

all used for the return current path. This arrangement

reduced the effect of the magnetic field from the return

current path. 7 C data was taken at 5 T with four returns and with one return for various frequencies and a.c. amplitudes. The results with a single return had a slight

frequency and amplitude dependence. At 60 Hz (sinusoidal and triangular), these data were within 1 % (2 A)

of the four-return results. At 360 Hz and 100 A a.c. amplitude, the 7 C was reduced by 3% (8 A) as compared

to the four-return results. This further reduction is attributed to the additional sample heating due to the

transient loss created by the magnetic field of the return

current path. For other geometries, such as a coiled

sample, this effect could be much larger.

High frequency data

Additional 7 C data were taken at higher frequencies

(3 and 10 kHz), but experimental difficulties made the

results at these frequencies less reliable. Transient losses,

inductive voltage, waveform distortion, and reduction in

available amplitude all became more significant at higher

frequencies. The motivation for investigating higher

frequencies was to estimate the effect of frequencies near

the range of spikes from silicon-controlled rectifiers (SCR)

found in most large power supplies. These data indicated

the same trend of reduction in 7 C but with more experimental uncertainty. Perhaps a better model of an SCR

spike would be a spike rather than a continuous dutycycle waveform. An extrapolation of the work presented

here would indicate that the effect of the spikes would

only be to add a finite amount to the time-average of the

resistive voltage.

Conclusions

The effect of ripple scales with the percentage of 7C, and the effect depends upon the shape of the V-I curve

(n). For an n of 30 or more and large ripple, current ripple

reduces the measured d.c. critical current by slightly less than the a.c. amplitude. The effect of ripple is nearly

independent of frequency for the range 60-1000 Hz. A

d.c. voltmeter measures a relevant quantity, the time

average of the voltage. A simple electronic circuit was

used to test a d.c. voltmeter’s response to an asymmetric

waveform that simulates a superconductor’s V-I characteristic. Reducing the mutual inductance of the voltage

taps is important when testing with power supply ripple

present.

A large conductor test method has to address the reality

of power supply current ripple. Power line frequencies

and harmonics (up to three phase, fully rectified) are the

dominant frequencies. Current ripple of 1-3% at the

maximum current is typical for power supplies in the

range 600-6000 A. This ripple is generally a function of

output current, but the percentage ripple can be higher

for intermediate currents. This can be a significant source

of measurement error. The direct effect on the sample

voltmeter can also introduce measurement error. Sample

heating due to transient losses may also be significant for some sample geometries.

Acknowledgements

The authors extend their thanks to W. E. Look and

H. E. Spilman for their assistance with sample preparation and computer programing, and to R. B. Goldfarb

for discussions on transient losses. Research supported

by the US Department of Energy through the Office of

Fusion Energy and the Division of High Energy Physics.

An effort was made to avoid the identification of

commercial products by the manufacturers name or

label, but in some cases these products might be indirectly

identified by their properties. In no instance does this

identification imply endorsement by the National

Bureau of Standards, nor does it imply that the particular

products are necessarily the best available for that

purpose.

References

1 Goodrich, L.F. and Fickett, F.R. Critical current measurements:

a compendium of experimental results Cryogenics (1982) 22 225

2 Bray, S.L. and Goodrich, L.F. Battery powered current supply for superconductor measurements National Bureau of Standards NBS1R 88-3088 ( 1988)

3 Goodrich L.F. et at. Development of standards for superconductors. interim report Jan. to Dec. 1985 National Bureau oj Standards NBSIR 87-3066 (1987) 4 Ekin, J.W. Materials at Low Temperature American Society for Metals, USA (1983) Ch 13

Cryogenics 1988 Vol 28 November 743

106
APPENDIX K

CALORIMETRIC MEASUREMENTS OF AC LOSS IN Nb 3 Sn SUPERCONDUCTORS

M. Takayasu, C. Y. Gung, M. M. Steeves, B. Oliver,

D. Reisner, and M. 0. Hoenig*

published in:

Eleventh International Conference on Magnetic Technology (MT-11)

Tsukuba, Japan, 1989

* None of these authors are NIST employees.

107
Eleventh International Conference on Magnet Technology (MT-11)

Tsukuba, Japan, 1989

CALORIMETRIC MEASUREMENT OF AC LOSS IN N^Sn SUPERCONDUCTORS*

M. Takayasu, C. Y. Gung, M. M. Steeves, B. Oliver, D. Reisner and M. O. Hoenig

Massachusetts Institute of Technology

Plasma Fusion Center

Cambridge, Massachusetts 02139, USA

ABSTRACT

The A.C. loss of twisted internal-tin, modified-jelly-roll Nb3 Sn wire intended for the US-DPC coil was

measured in triangular wave fields at ramp rates up to 1.3 T/s by an isothermal calorimetric method. To

minimize the time delay of measurement, a constant-flow control method was developed. Measurement with

ripple and triangular fields allowed hysteresis loss and coupling loss to be separated. Hysteresis loss obtained

by the present calorimetric apparatus agreed quite well with magnetization measurements obtained at NIST.

Conclusions drawn from the preliminary tests indicate the major component of A.C. loss to be hysteresis

below 2 T/s for the observed low values of the residual resistivity ratio, RRR (10 - 30). The bundle coupling

loss will become comparable or even dominant at the expected maximum operational ramp rate of 9.5 T/s,

especially if the RRR values were to improve significantly. Estimates of A.C. loss for the US-DPC coil

operating at JAERI are made.

INTRODUCTION

The United States Demonstration Poloidal

Coil (US-DPC) is being built in order to test

advanced, high-current-density, pulsed Nb3 Sn

cable-in-conduit superconductors for ohmic heating

coils for tokamak fusion devices [1]. AC loss is

one of the critical factors for the US-DPC Coil,

which uses commercial Nb3 Sn wire. The

superconducting wire was manufactured by

Teledyne Wah Chang Albany (TWCA). It is

titanium- alloyed internal-tin, jelly-roll Nb3 Sn wire

of the diameter 0.78 mm with an internal Cu/Nb

ratio of 1.821 with a vanadium barrier, designed to

maximize current density while minimizing A.C.

loss [2]. The nominal non-copper current density is

700 A/mm2 at 10T [3]. A.C. loss measurements of

the 225 strand, 30 kA US-DPC conductor will be

carried out at ramp rates of up to 9.5 T/s in early

1990 at the Demonstration Poloidal Coil Facility

(DPCF) at JAERI, Japan.

In order to perform laboratory

measurements of A.C. loss, an isothermal

calorimeter was built in which loss is measured as a

Supported by US DOE Contract DE-AC02-78ET51013.

function of helium boil-off. The calorimeter

permitted the loss measurement of 12 m long single

strand samples (5 cc) in a continuous triangular

wave field of a 12T Bitter magnet at the Francis

Bitter National Magnet Laboratory, MIT. To

minimize measurement time delay, a constant flow

control method was developed. Tests of single

strand samples of TWCA US-DPC wire were

performed with triangular wave fields at ramp rates

of up to 1.3 T/s. Ripple field loss test were also

performed. Together, die triangular field and ripple

field tests allowed hysteresis and coupling loss to be

mathematically separated. The coupling loss is

mainly bundle coupling loss, that is, an eddy

current loss between adjacent filament bundles.

Based on these measurements, an estimate of the

A.C. loss of US-DPC coil under test conditions is

given.

EXPERIMENTAL

APPARATUS

A calorimetric apparatus [4] was developed

to measure the A.C. loss of single strands, as well

as cabled superconductors, with and without

transport current. Fig. 1 shows a schematic of the

108
Fig. 1 Schematic diagram of calorimetric A.C. loss measurement apparatus.

apparatus. A 12 m long single strand (MIT

designation 87-17) was first wound non-inductively

in a single layer on a 100 mm diameter, 75 mm long

machinable ceramic, and was subsequently heat

treated. The test sample was enclosed by an open

bottom bell jar in the 127 mm inner diameter tail

section of a metal cryostat. The helium boil-off was

measured at room temperature by a mass flow

meter. The apparatus was calibrated by a D. C.

resistive heater.

A constant flow control system was used in

order to minimize the time delay required to

establish an equilibrium state of the flowing helium

gas. Constant mass flow was maintained by an

automatic feedback system which controlled a background heater. The sensitivity was about 5

mW, and the resolution was less than 1 mW.

Typical A.C. loss measurements ranged from 20 - 40 mW.

TEST RESULTS

Measurements of AC loss were made at

4.2K,

A) by means of high frequency ripple fields at

steady DC fields and.

Fig. 2 Experimental data of total A.C. loss at 4.2K

due to a ripple field as a function of steady D.C.

field with the superimposed ripple field of peak-topeak amplitude Bp=0.086T and frequency f=7.5

Hz. The solid line shows theoretical loss using

measured RRR of 1 1.9.

109
Fig. 3 Measured total A.C. loss at 4.2K due to a ramp field as a function of the maximum field Bm. The lower solid line shows theoretical bundle

coupling loss with the ramp time Tm=6 s.

B) by means of triangular wave ramped fields of up

to 9 teslas with maximum ramp rates of 1.3 T/s.

A. RIPPLE FIELD TESTS

Fig. 2 shows A.C. loss due to a ripple field as a function of D.C. field. The frequency and the

amplitude (peak-to-peak) of the ripple field were

0.086T and 7.5 Hz (1.3 T/s), respectively. The

solid line is calculated from the existing theoretical

equation [5], taking RRR to be 11.9, a measured

value of the tested wire. The negative slope of the

solid line is due to the magnetic effect on the

electrical resistivity of the copper matrix. The

agreement between the experimental data and the

calculated values is fairly good, except that the

experimental points show a peak around 3T of the

D.C. field.

B. RAMP FIELD TESTS

Fig. 3 shows measured total loss due to a ramped field as a function of maximum field with

various ramp times. The lower solid line calculated

using existing theory shows the bundle coupling

loss [5]. The difference between the total loss and

the calculated bundle coupling loss appears to be

hysteresis loss, plotted in Fig.4.

The four solid circles in Fig. 4 show the

hysteresis loss measured by magnetization [4].

Original data were full-cycle hysteresis losses.

They were reduced to half-cycle loss using

Fig. 4 Hysteresis loss obtained by the subtraction

of the calculated bundle coupling loss from the

measured total loss shown in Fig. 3. The solid line

is calculated from Eqs. (1) and (2) with a=7.5

microns.

Fig. 5 Relationship between the half cycle

hysteresis and full cycle hysteresis losses, which

were measured by the magnetization method at

NIST. The maximum field values are 3T () and

7T (). The half cycle hysteresis loss is seen to be

42.6% of the full cycle one.

magnetization data obtained experimentally for

similar TWCA US-DPC wire. The comparison is

shown in Fig.5. The half cycle hysteresis loss was

42.6% of full cycle loss.

110
If a superconductor has no transport current,

the hysteresis loss per unit volume per cycle may be

given by Eq. (1) [5].

Qh = i| JcadB, (1)

where Jc is the critical current as a function of the

field B, Bm is the maximum field and a is the

filament radius.

An approximation of the critical current Jc

with the field B is given by the following equation

obtained experimentally by tests using similar USDPC wires at magnetic fields between 8T and 15T.

Jc - 6.117 x 1()3 exp(-0.219 B) [A/mm2] (2)

The solid line in Fig. 4 is based on Eq. (1)

with Eq. (2), taking the filament diameter to be 15

microns. While the filament size is larger than

expected, an effective filament size would be larger

than the design value (3 microns) or optically

observed filament size (about 6 microns).

Agreement between the experimental data and the

calculation is qualitatively significant

MEASUREMENT ACCURACY

The accuracy of A. C. loss measurement

using our calorimetric technique depends on the

following:

(1) Extraneous boil-off caused by current leads,

cryostat, and other structure materials.

(2) Heat transfer through the bell jar.

It is caused by temperature difference of the

helium boiling point due to the pressure

difference between helium inside and outside of

the bell jar.

(3) Acoustic resonance in the vent tube.

It is high frequency oscillation noise of helium

vapor flow.

(4) Leakage of helium gas.

(5) Mass flow change due to the time-varying

magnetic field.

One disadvantage of the calorimetric method

is that the time required to reach equilibrium

(usually takes more than 40 min) [4]. By using the

constant flow control system, the time required can

be reduced to less than 10 min. In this way, the

operating pressure and helium bath temperatures can

be kept constant during measurement.

When pressure outside the bell jar in the

cryostat changes, undesirable heat transfer occurs

through the vent tube wall. In order to prevent this,

the vent tube was originally vacuum shielded. The

high frequency acoustic noise (problem (3), above)

was however enhanced. The noise became

negligible when a single layer vent tube was used

without a vacuum shield. By adopting the constant

flow operation, the pressure fluctuations in the

cryostat during measurement could be minimized.

Leakage of boil-off vapor from calorimeter

in liquid helium was increased with both D.C. and

A.C. magnetic fields (problem (4), above). This

problem was solved using Stycast 2850FT to seal

the G10 canister.

Problem (5), not yet fully understood,

causes the most serious background noise for an

accurate loss measurement. When a triangular

waveform is applied, the mass flow change per unit

time due to this noise depends on the amplitude of

magnetic field, but not ramp time. This means that

the noise effect on the measurement loss per cycle

proportionally increases with ramp time. The data

obtained with more than 30 sec of the ramp time had

a large error due to this noise. The data shown in

Figs. 3 and 4 were obtained with a ramp time Tm of

12 sec or less. The resultant error was less than

10%.

Table 1. Measured and design values of the half cycle hysteresis

loss per cycle in unit of mJ/cc.

Maximum Field Present Calorimetric Magnetization

Bm (T) Method Method (NIST)

3 86 83*

7 140

* 42.6% of the full cycle hysteresis loss measured.

Ill
DISCUSSION AND CONCLUSIONS

The A.C. loss test results obtained by the

calorimetric method have an error of less than 10%.

The hysteresis loss per cycle of US-DPC wires with

a triangular wave, Eq. (3), was obtained using

Eqs.(l) and (2) with a computer fitting to the

experimental data shown in Fig. 4.

Qh = 178.2 { 1 - exp(-0.219 B^)} [mJ/cc] (3)

Table 1 shows both the half cycle hysteresis

losses obtained from Eq. (3) and using the

magnetization method. As discussed above, the

values obtained by the magnetization method are reduced to the half cycle hysteresis loss by taking

42.6% of the full cycle values. The half cycle

hysteresis loss per cycle of US-DPC wire was 86

mJ/cc for a 0 - 3T - 0 cycle and 140 mJ/cc for a 0 - 7T - 0 cycle. These values agreed quite well with

values obtained by the magnetization method at NIST.

The other major A.C. loss component is the

bundle coupling loss. As shown in Fig. 3,

however, bundle coupling loss was minor in our

experimental range of frequency. The coupling loss

increases with the ramp rate and copper matrix

resistivity. The effect of magnetic field on copper

resistivity or RRR must be taken into account.

Even if RRR is as large as 1000 without magnetic

field, the value is decreased to less than 40 at a magnetic field of 10T. If the copper resistivity is a

linear function of magnetic field B, given as p(B,4.2K) = po(B=0, 4.2K) + a B [Ohm m], the

coupling loss per cycle due to the triangular wave

can be given using existing theory [5] by

In Eq. (4), lpis the twist pitch and p

is the effective

transverse resistivity. Since lp

of the wire is 9.8

mm, and if po=2.0xl0’8/RRR [Ohm m] and a=4.8

x 10-H [Ohm m/T], the coupling loss per cycle is

given as follows,

Q. = 50.68^ ln(l+2.4xlO'3 RRR BJ [mJ/cc] (5)

Fig. 6 shows the bundle coupling loss

calculated using Eq. (5), as a function of the ramp

rate of a triangular wave field. The maximum field

value Bm is 10T. The hysteresis loss per cycle is

also shown in the figure. It is independent of ramp

Fig. 6 The expected bundle coupling loss of USDPC wire for various RRR values as a function of

the ramp rate of a triangular wave field at the

maximum field of Bm=10T. The hysteresis loss is

also shown by a dotted line.

rate. As seen in Fig. 6, the bundle coupling loss

will be a few times larger than the hysteresis loss at

the expected maximum US-DPC operating ramp

rate of 9.5 T/s.

Conclusions drawn so far from the tests

with triangular and ripple waveforms indicate the

major component of A.C. loss for a twisted TWCA

US-DPC wire to be hysteresis below 2 T/s for the

observed low values of the residual resistivity ratio,

RRR (10 - 30). However, the bundle coupling loss

(eddy current loss between adjacent filament

bundles) will become comparable or even dominant

at the expected maximum operating ramp rate of 9.5

T/s, especially if the RRR value were to improve

significantly. It is noted that the RRR value of A.C.

superconductor must be optimized to provide

maximum protection at minimum A.C. loss.

The US-DPC is composed of three double

pancakes. The length of each pancake is 75 m. The

coil will be operated in series with DPC-U1 and U2

of JAERI using a continuous triangular wave

current, ramped between 0 and 30 kA (maximum)

in a few seconds. The maximum field, which the

innermost turn will be exposed to, is 10T. By

using the calculated field distribution of the USDPC conductor, both the hysteresis loss and the

bundle coupling loss were calculated from Eqs. (3)

and (5), respectively. Fig. 7 shows the results

obtained with RRR=30 under the most severe

conditions (maximum current of 30 kA, and ramp

112
MO

oo

E

UJa

>

o

ccUJa.

(/) COO

d

<

LENGTH FROM H« INLET (m)

Fig. 7 Calculated A.C. loss per cycle per unit

volume along the conductor of the center pancake of

US-DPC.

RAMP TIME Tm (3)

Fig. 8 Calculated A.C. loss power per second of

the whole three double pancakes of US-DPC.

time Tm of Is). Here the A.C. loss per cycle is

plotted as a function of position along the conductor

of the center pancake from the point of helium

entry. As seen in this figure the bundle coupling

loss is dominant at the inner turns.

The A.C. loss power, that is loss per unit

time, can be given by dividing Eqs. (3) and (5) by

the cycle time 2Tm. Calculated total A.C. loss

power per second of the whole US-DPC coil with

the maximum current of 30 kA and RRR of 30 is

shown in Fig.8 as a function of ramp time. The

coupling loss becomes comparable or dominant

with Tm of Is (the expected maximum operating

ramp rate). The maximum total A.C. loss expected

under these operating conditions is about 5 kW.

ACKNOWLEDGMENT

The authors wish to thank T. Ichihara on

leave from Mitsubishi Electric Corp. and S. H.

Autler of University of Massachusetts for their

initiation of the calorimeter and helpful discussions,

and are also grateful to Y. Iwasa for helpful

suggestions, and J. S. Schultz for magnetic field

analysis of US-DPC coil. The authors are indebted

to R. B. Goldfarb and R. C. Spomer of NIST for

hysteresis measurements by magnetization method.

REFERENCES

1. Steeves, M.M. et al. The US-DPC, a poloidal

coil test insert for the Japanese demonstration

poloidal coil test facility. IEEE Trans. Mag.,

1988, 24, 1307-1310.

2. Hoenig, M.O., Steeves, M.M. and Gibson,

C.R., The selection of 30kA ohmic heating coil

conductor. EEEE Trans. Mag., 1988, 24, 1452-

1454.

3. Smathers, D.B. et al. Production of tin core

modified jelly roll cable for the M.I.T.

multipurpose coil. IEEE Trans. Mag., 1988,

24, 1131-1133.

4. Goldfarb, R.B., Ekin, J.W., Hysteresis losses

in fme filament internal-tin superconductors.

Cryogenics, 1986, 26, 478-481.

5. For example, Wilson, M.N., Superconducting,

Clarendon press,Oxford, 1983, pp. 162- 182.

113

APPENDIX L

GUIDELINE FOR Nb 3 Sn CRITICAL CURRENT MEASUREMENTS USING

FIBERGLASS -EPOXY COMPOSITE SAMPLE MANDRELS (SUMMARY)

L. F. Goodrich

Presented at VAMAS TWP meeting

International Cryogenic Materials Conference

Los Angeles, California

23 July 1989

115
Guideline for Nb3Sn Critical Current Measurements

Using Fiberglass-Epoxy Composite Sample Mandrels

(Summary)

L. F. Goodrich

Superconductor and Magnetic Measurements Group

National Institute of Standards and Technology

Boulder, Colorado

Presented at VAMAS TWP meeting

International Cryogenic Materials Conference

Los Angeles, California, 23 July 1989

The materials used for this technique are stainless steel tubes for reaction

mandrels, G-10CR cylinders cut from plate stock (plate tubes, see Fig.1)

for measurement mandrels, and silicone-based vacuum grease as the

bonding material.

Plate

Fabric

Planes

Plate

Tube

Rolled

Tube

Normal

i k

Fabric

Planes

Warp

Fig. 1. Illustration of fiberglass-epoxy geometries.

116
Preface

This guideline and others were requested as an attempt to achieve

consistent interlaboratory measurements.

A competition exists between the benefit of a rigidly specified

measurement technique, which ensures measurement reproducibility,

and a less restrictive technique that would be practical for a

greater number of laboratories.

This measurement guideline is intended to be one of a few different

measurement techniques in a second Nb3Sn l c round robin that is

envisioned to determine the preferred measurement method for Nb3Sn.

At this time, there is no consensus regarding the materials used

for the reaction mandrel, measurement mandrel, or sample bonding.

There are advantages and disadvantages for all known materials.

This guideline was written for one combination of materials.

It may not be the best, but it represents one feasible option.

117
Sample Reaction Mandrel

Helical grooved (not too deep) and oxidized tube.

Well seated, but not tight sample.

Bend wire through holes on each end to retain coil.

Internal-tin processed wire ends outside furnace.

Sample Measurement Mandrel

G-10 "plate tube" (see Fig. 1).

Same pitch length and groove geometry as reaction.

Dia. approximately 0.5% larger than reaction mandrel.

Thin layer of grease in groove.

Sample well seated in groove.

Sample Mounting

Check for diffusion bonding.

If bonded, the sample should be discarded.

Two methods of removing sample from reaction mandrel, unthreading or

slightly expanding coil.

If threading onto measurement mandrel, then apply grease to groove with

sample up between grooves.

118
Sample Mounting (cont.)

Solder one current contact, then remove the coil’s slack, hold the sample

taut and solder the other current contact.

Neither contact should have more than 0.9 turns otherwise a slowly

decaying current can be induced.

Attach voltage taps to wire after it is seated.

Sample Cooling Rate

Grease has no mechanical strength until it freezes

May result in different mechanical states for samples

cooled at different rates.

Cool sample over a time period of at least 5 minutes.

Current Reversal

Grease has a relatively low tensile strength.

Lorentz force on the sample coil directed inward only.

Field History and Repeat Determinations

First measurement at a magnetic field closest to

where the Lorentz force is maximum.

Repeat determination of l c

at this first magnetic field.

119
Discussion

Diffusion bonding is a significant problem.

Baking in a sprayed graphite lubricant may help.

Thermal expansion of reaction mandrel may be a problem because the wire

may shorten upon anneal.

Does this lead to alumina or a stainless steel with an alumina coat?

A low wall-to-radius ratio G-10 or G-1 1 rolled tube may be another option

for a measurement mandrel that is well matched in thermal contraction to

a Nb3Sn conductor.

Grease may not be necessary and it may have an adverse effect.

The angle of the groove may have an effect; 90° may be better than 60°.

Cautions

This was not written as a step-by-step procedure.

This is a short version of this guideline; see the text version for more

complete details.

120
Standardization of Coil Dimensions

Serious consequences to the selection of a coil

diameter

If the coil diameter is selected to allow measurements in the smallest

magnet bore, then the probability of sample damage due to handling

increases.

If a large coil diameter is selected, a number of laboratories may not be

able to participate.

Significant benefits to standardizing coil diameter

This may be the most effective way to achieve consistent l c measurements.

Batch fabrication of reaction and measurement mandrels with common

materials could be cost effective and allow more interchangeable

combinations of reaction and measurement mandrels.

This could facilitate the exchange of samples among laboratories.

121

APPENDIX M

REPORT ON SAMPLE MOUNTING

L. F. Goodrich

Presented at VAMAS TWP meeting

International Cryogenic Materials Conference

Los Angeles, California

23 July 1989

123
Report on Sample Mounting

L.F. Goodrich

Superconductor and Magnetic Measurements Group

National Institute of Standards and Technology

Boulder, Colorado

Presented at VAMAS TWP meeting

International Cryogenic Materials Conference

Los Angeles, California, 23 July 1989

124
Reaction Mandrels

Requirements

* Need grooves and retainers.

* Want to avoid diffusion bonding.

* Thermal expansion may be important.

Options

* Oxidized stainless steel.

- Easy to machine.

- Diffusion bonding can be a problem.

- High thermal expansion.

* Stainless steel with ceramic coat.

- Cost and adherence of ceramic?

- Thermal expansion is still high.

* Alumina

- Cost of machining is high.

- Durability?

- No diffusion bonding.

- Low thermal expansion.

125
Measurement Mandrels

Requirements

* Reversible sample bonding.

* Thermal expansion is important.

Options

* G-10 rolled tubes with low wall/radius and grease bonding.

- Easy to machine.

- Thermal contraction can be matched or lower, (tighten coil)

- Thermal contraction may vary.

- Thin layer of grease or bare.

* G-10 plate tubes with grease bonding

- Hard to machine.

- Thermal contraction lower than Nb3Sn wire, (tighten coil)

- Thermal contraction not variable.

- Thin layer of grease or bare.

126
Measurement Mandrels (cont.)

* Stainless steel with grease bonding.

- Easy to machine.

- Thermal expansion slightly more than Nb3Sn wire?

- Grease may support.

- Augmented sample stability may be a problem.

- Shunting current may not be a problem.

- Could use the same mandrel for measurement

and reaction, (differential expansion?)

* Stainless steel with solder bonding.

- Easy to machine.

- Thermal contraction more than Nb3Sn wire?

- Augmented sample stability is a problem.

- Shunt Current may be a problem.

- Can not use same mandrel for measurement and reaction.

* Alumina with grease bonding.

- Hard to machine

- Thermal contraction less than Nb3Sn wire.

- Could use same mandrel for measurement and

reaction, (differential expansion?)

127

APPENDIX N

GUIDELINE FOR Nb 3 Sn CRITICAL CURRENT MEASUREMENTS

USING FIBERGLASS -EPOXY COMPOSITE SAMPLE MANDRELS

L. F. Goodrich

Prepared for: Office of the VAMAS Technical Working Area 6

Superconducting and Cryogenic Structural Materials

129
Guideline for Nb 3 Sn Critical Current Measurements

Using Fiberglass-Epoxy Composite Sample Mandrels

L. F. Goodrich

Superconductor and Magnetic Measurements Group

National Institute of Standards and Technology

Boulder, CO 80303

PREFACE

This guideline was prompted, in part, by presentations given at the 6th

Japan-U.S. Workshop on High Field Superconductors (February 22-24, 1989

Boulder, Colorado) and subsequent discussions. The results of the recent

Versailles Project on Advanced Materials and Standards (VAMAS) Nb 3 Sn criticalcurrent round robin indicate that increased consistency in interlaboratory

measurements might be achieved by a more detailed specification of the

critical-current (I c ) measurement technique. However, a competition exists

between the benefit of a rigidly specified measurement technique, which

ensures measurement reproducibility, and a less restrictive technique that

would be practical for a greater number of laboratories. Ideally, the

measurement technique should have the least number of restrictions that is

consistent with measurement reproducibility.

This measurement guideline is intended to be one of a few different

measurement techniques in a second Nb 3 Sn I c round robin that is envisioned to

determine the preferred measurement method for Nb 3 Sn. Each participating

laboratory, in this second round robin, will make measurements using each

technique. The number of laboratories participating in this second round

robin might be considerably reduced, as compared with the first round robin,

because of the difficulty involved in complying with the necessarily

restrictive specifications for the various measurement techniques. However,

in the event that the second round robin results in a general consensus

regarding which measurement variables must be specified in order to assure

consistent measurements among different laboratories, a third round robin

might be useful for determining the effectiveness of the measurement method

for a larger group of laboratories.

At this time, there is no consensus regarding the materials used for the

reaction mandrel, measurement mandrel, or sample bonding. There are

advantages and disadvantages for all known materials. This guideline was

written for one combination of materials. It may not be the best, but it

represents one feasible option. The materials used in this technique are

stainless steel tubes for reaction mandrels, G-10CR cylinders cut from plate

stock for measurement mandrels, and silicone based vacuum grease as the

bonding material.

This measurement guideline assumes that the reader has the basic

knowledge and experience with I c measurements of NbTi and Nb 3 Sn conductors.

If not, the basic information on these measurements can be found elsewhere [1,

2 , and 3 ]

.

130
INTRODUCTION

The results of the VAMAS I c round robin demonstrated the existing

measurement discrepancy among a large number of laboratories for several

different Nb 3 Sn samples [4]. The discrepancy was significant. A major source

of this discrepancy was determined to be varying degrees of axial sample

strain caused by variations in the thermal contraction of different sample

mandrels [5].

The sample bonding technique can play an important role in this effect

because the strength of the bonding determines the amount of the sample

mandrel's thermal contraction that is transmitted to the sample. A relatively

weak bond can inhibit this transmission of thermal contraction but it may also

allow sample motion under the influence of the Lorentz force, which, in turn,

can cause variations in the measured I c

. A high strength bonding material

used in conjunction with a sample mandrel whose thermal contraction is

comparable to that of the Nb 3 Sn sample alleviates both of these mechanical

problems but it introduces a practical problem. The commonly used sample

mandrel material that has satisfactory thermal contraction is a fiberglass - epoxy composite (G-10). The high strength bonding material that is typically

used for G-10 sample mandrels is a filled epoxy adhesive. The problem with

this method is that the epoxy forms a fairly permanent bond which makes

repeated use of the sample mandrel difficult. Clearly what is needed is a

bonding material that provides adequate sample support but that is easily

removable from the sample mandrel. The results of the VAMAS measurements

indicate that grease may be adequate for this purpose as long as all of the I c measurements are made with the Lorentz force directed inward, toward the

central axis of the sample mandrel.

Grease is a generic term that applies to a very broad range of heavy

lubricants that have significantly different physical properties. Two such

materials that have been used for sample bonding materials are petroleum jelly

and silicone -based vacuum grease. In this document, grease means siliconebased vacuum grease .

Recent research at NIST indicates that a G-10 cylinder appropriately

machined from plate stock has a thermal contraction that does not vary

significantly with its geometry and is close to that of a Nb 3 Sn wire sample.

Earlier research indicated that the thermal contraction of rolled fiberglass

epoxy composite tubes varies significantly with the wall thickness of the tube

[5]. Because these composites are anisotropic, the orientation of the

material's internal structure, relative to the tube, is important. Figure 1 illustrates several fiberglass -epoxy composite geometries. A plate tube can

be made by machining a cylindrical tube from thick fiberglass -epoxy plate

stock with the axis of the tube perpendicular to the surface of the plate.

There is a commercially available rod that is machined from plate stock, but

the axis of this rod is in the plane of the plate. A rod of this type will

have a very anisotropic radial thermal contraction which is unacceptable for

this application.

A plate tube can be machined to make a sample measurement mandrel. The

inner and outer diameters of the plate tube can be machined to receive current

131
contacts and to be mounted on the test fixture without significantly affecting

the thermal contraction of the G-10. There is an advantage to having extra

wall thickness however; it reduces the chance of delamination. The thickness

of the plate stock will determine the maximum overall length of the plate

tube. It is better to start with a single plate than to build up the

thickness by gluing multiple plates together. The plate tube can be machined

from a plate using a hole saw or by cutting out a parallelepiped then turning

it on a lathe. A hole saw is a tool that can be used to drill large diameter

holes. It can also be used to extract a cylindrical section from a plate,

which can be turned on a lathe to the final diameter. Water is commonly used

as a machining lubricant for G-10; it also reduces the dust and keeps the

temperature of the G-10 low.

MATERIALS

The materials used in this technique are stainless steel tubes for

reaction mandrels, G-10 cylinders cut from plate stock for measurement

mandrels ( plate tubes ) , and silicone-based vacuum grease as the bonding

material. The stainless steel tubes and silicone-based vacuum grease are

readily available and have consistent properties for this application. The

G-10 plate, especially plate of the necessary thickness, is not as available

and the consistency of its properties is not known. Thus, a single lot of G10 should be acquired. Dr. Hitoshi Wada of Japan's National Research

Institute for Metals has expressed a willingness to purchase a common lot of

this material, thus, minimizing variations in properties.

A G-10, NEMA (National Electrical Manufacturers' Association)

designation, measurement mandrel material was selected because it can be

oriented to have a thermal contraction that does not vary significantly for

various geometries and it is close to that of a Nb 3 Sn wire sample. Cryogenicradiation (CR) grade material should be purchased, if possible, since the

manufacturing specifications for CR grade materials are more stringent and

their performance at cryogenic temperatures is more predictable. The use of

CR grade material would also allow for better reproducibility in the future.

Thick G-10CR plate is available from a number of sources in the U.S. and

a sheet or two may be sufficient to supply the participants with material. A

full sheet is about 91 by 122 cm (3 by 4 feet) and a 5 cm thickness would cost

about $2,000 plus freight (110 kg). A full sheet may be enough material for

400 sample mandrels. The key parameter is the thickness of the plate since

this will determine the maximum length of the measurement coil. The

measurement mandrel needs to be machined from the plate stock with the

cylinder axis normal to the plate surface in order to better match the thermal

contraction of the Nb 3 Sn coiled sample. The thickness of the plate should be

at least 5 cm. The plate stock could be cut into strips that are slightly

wider than the measurement mandrel diameter for distribution to participants

of this round robin. This would facilitate machining the cylinders, which is

desirable since G-10CR is somewhat difficult to machine. Each participant

should be supplied with enough material to construct several (5 to 10)

mandrels to allow for different wire diameters and mistakes during machining.

132
PROCEDURE

General Philosophy

Because this measurement technique will be conducted in many

laboratories, it is impractical to specify all of the details of the sample

reaction, mounting, and measurement procedures. For example, a detailed

specification might require that the sample mandrels have an outside diameter

that is slightly different than that normally used at the testing laboratory,

and compliance with this somewhat arbitrary specification would unnecessarily

increase the required time and cost for this study. Consequently, the

philosophy of this guideline is to be as general as is practical and, thus,

allow the measurement laboratories to adhere to their normal practices as much

as possible.

Please read through all of these procedures before implementing this

guideline. It was not written as a step-by-step procedure. Some elements are

described out-of -sequence because of the inter-relationship between steps and

various options that are presented.

Sample Reaction Mandrel

The samples should be reacted on well oxidized stainless steel tubes in

order to avoid diffusion bonding between the sample and the tube. When

diffusion bonding occurs, the probability of damaging the sample when

transferring it to the measurement mandrel is greatly increased. Even when

diffusion bonding does not occur, extreme care must be exercised in

transferring the sample to avoid damage due to bending strain. The reaction

mandrel should have a helical groove for defining the sample coil's pitch.

The sample should be well seated in this groove, but not tight, prior to

reaction to insure reproducibility of the coil diameter and pitch. The wire

can be retained on the reaction mandrel by bending the wire through a hole at

each end of the mandrel. Internal -tin processed conductors need to have all

wire ends well outside the reaction furnace, unless the operator is specially

trained to weld the wire ends.

Sample Measurement Mandrel

The G-10 measurement mandrel should have the same pitch length, groove

geometry, and a diameter that is approximately 0.5% larger than the stainless

steel reaction mandrel to insure that the reacted sample can be fit to the

measurement mandrel with a minimum of sample deformation. The samples should

be well seated in the measurement mandrel's groove to ensure consistent coil

diameter and, thus, minimize bending strain. This is also important for

minimizing the thickness of the bonding material (grease) between the sample

and mandrel. Moreover, a thick layer of grease is more likely to yield under

the influence of the Lorentz force than a thin layer, where the sample is

substantially in contact with the G-10 mandrel.

A good choice for the mandrel's groove geometry is a 60° "V" with a

fairly sharp point. The 60° groove provides good sample support in both the

radial and transverse directions . The sharp point of the groove insures that

133
the conductor contacts the groove at the points of tangency rather than at the

bottom of the groove. In the case of the measurement mandrel, this should

minimize the thickness of the bonding-material layer between the sample and

the mandrel. In the case of the reaction mandrel, it may reduce diffusion

bonding. The groove should be deep enough that the sample contacts the

mandrel within the groove rather than at the interface formed between the

groove and the mandrel's surface. This is important for transverse stability

of the sample. One detrimental consequence that results from making the

groove too deep is that the sample will require greater deformation during the

transfer process in order to clear the mandrel.

Sample Mounting

Before attempting to transfer the sample from the reaction mandrel, check

it for diffusion bonding to the mandrel. If there Is significant bonding that

might result in sample damage, the sample should be discarded. Two methods

can be used for removing the sample from the reaction mandrel: The first

requires a very smooth interface between the sample and mandrel but, when it

can be employed, allows sample removal with minimal deformation. In this

method the sample is unthreaded from the mandrel by lightly gripping the

sample coil's surface while turning the mandrel. Alternatively, the sample

coil's diameter can be expanded by carefully untwisting the coil slightly

until the mandrel can be extracted. Although this second technique could

result in some sample deformation, it may be the better method depending on

groove depth, the number of turns, and the operator.

It is necessary to have a complete, but thin, coat of grease under the

sample on the grooved G-10 mandrel. This is difficult or impossible to

maintain if the sample is threaded onto the measurement mandrel. However, the

sample could be threaded completely onto the measurement mandrel, then

expanded slightly and moved just out of the groove to a position between the

grooves on the mandrel. With the sample in this position, the groove could be

filled with grease and then the sample could be slipped back into the groove.

In the other case, where the coil is expanded to clear the mandrel, the grease

should be applied to the sample groove prior to sample transfer.

Voltage tap wires can be attached to the sample after it is seated in the

groove. There is a chance that the grease will interfere with the soldering

operation or that the elevated temperature will affect the properties of the

grease, but these are not expected to be significant. If the voltage tap

wires are attached to the sample before the sample is set into the greased

groove, there are two possible options depending upon how the sample is

transferred from reaction to measurement mandrel. If the sample Is threaded

onto the measurement mandrel, the voltage tap wires could be attached to the

sample after it has been transferred to the measurement mandrel but before It

has been set in its groove. In the case where the sample is transferred

directly to a greased measurement mandrel, the voltage tap wires could be

soldered to the sample while it is on the reaction mandrel. In all cases, the

single voltage tap wires can be counter wound onto the coil and pairs of wires

twisted together after the sample is attached to the current contacts.

134
The amount of grease that is necessary will vary depending upon the

details of the groove geometry and sample diameter. In most cases it is

expected that a thin coat on the surfaces of the groove would be a sufficient

amount of grease. A little extra grease could be applied to reduce the chance

of an incomplete coat. Too much extra grease in the groove will make it more

difficult to keep the grease layer thin, since the excess grease will act like

a hydraulic fluid and require excess pressure to extrude past the sample. A

thick layer of grease might lead to a less rigid support for the sample under

the influence of the Lorentz force. Any excess grease outside the groove

should be removed after the sample is in place so that sample cooling is not

inhibited. In cases where sample motion due to mechanical vibration is

expected, grease could be added to fill the sides of the groove after the

sample is seated. Since all of the measurements will be made with the Lorentz

force directed inward, there is limited benefit in having excess grease on top

of the sample.

Before soldering the sample to both of its current contacts, remove the

slack in the sample coil by firmly seating it in the sample groove. This is

easily achieved by first soldering one end of the sample coil to its current

contact and then, starting at that contact, pressing the sample into its

groove in a continuous motion that results in all of the coil's slack being

transferred to the opposite current contact. The sample can then be fixed in

this taut position by soldering it to the current contact. Neither current

contact should have more than 0.9 of a turn soldered to the current contact.

If more than one turn is soldered to the current contact there is a chance

that this loop will carry a slowly decaying current (induced by a change in

the background field), which will affect the magnetic field of the sample.

Sample Cooling Rate

Different laboratories probably cool their samples at significantly

different rates; in the limiting cases, some may slowly cool the sample in

helium vapor while others may insert the sample directly in a liquid-helium

bath (or perhaps liquid nitrogen). This variable may effect the measured I c

. The grease bonding material has virtually no mechanical strength until it is

cooled below its freezing temperature. Consequently, the strength of the

coupling between the sample and its mandrel is changing during the cooling

process while differential thermal contraction between the sample and mandrel

is occurring. This fact may result in different mechanical states for samples

cooled at different rates. This guideline suggest that the sample should be

cooled from room temperature to liquid helium (or liquid nitrogen) temperature

over a time period of at least 5 minutes.

Current Reversal

Grease will undoubtedly present a limitation in the I c measurements that

other more permanent bonding materials do not. Its relatively low tensile

strength will probably disallow its effective use for I c measurements where

the Lorentz force on the sample coil is directed outward. This, in turn,

disallows sample -current reversal for testing the effect of the sample coil's

self -generated magnetic field on the measured I c

.

135
Field History and Repeat Determinations

Because the Lorentz force may play an important role in the I c measurements of grease-mounted samples, the magnetic field history of the test

sample may affect the measurements. For example, if the first I c measurement

is made at the field where the Lorentz force on the sample is a maximum, the

subsequent measurements might be different than for another field sequence.

Consequently, the measurement sequence should be carefully considered and

recorded. It is recommended that the first measurement be at the field

closest to where the Lorentz force is maximum (the force peaks at about 5 T

for Nb 3 Sn) . Also, if the Lorentz force causes irreversible sample motion,

subsequent measurements at the same magnetic field may have different results.

Repeat determinations of the I c should be used to detect mechanical

deterioration of the sample bond. Repeat determinations are suggested for

this guideline, especially for the first magnetic field at the peak Lorentz

force

.

DISCUSSION

Diffusion bonding between the sample and the reaction mandrel has been a

significant problem. In addition to oxidization of the surface, baking (prior

to winding a Nb 3 Sn sample) a sprayed graphite lubricant into stainless steel

reaction mandrels has been suggested and we found it beneficial. With use,

the surface of the stainless steel becomes more passive. Other reaction

mandrel materials such as graphite and alumina have been suggested. A related

issue that has been mentioned, is that the thermal expansion of the reaction

mandrel material will affect the ultimate I c of the wire that is reacted on

it. The postulate is that the wire irreversibly shortens during the initial

anneal of the reaction heat treatment and then the wire is stretched more by

mandrels that expand more, resulting in a lower I c. This would suggest that

reaction mandrel materials such as graphite or alumina would be better than

stainless steel.

There are some unresolved issues regarding sample motion and support on

the measurement mandrel. The grease bonding material may not be necessary

and, in fact, it may cause some adverse effects. Some possible effects were

mentioned above, such as; grease may interfere with the soldering operation,

it may inhibit seating the sample into the groove, and it might inhibit sample

cooling. If the thermal contraction of the mandrel is less than the sample

from room to liquid-helium temperature, then the sample will tighten onto the

mandrel upon cooling and a bonding material may not be necessary except in

cases were mechanical vibration is expected. It is also possible that grease

will act like a lubricant and enhance conductor motion. Another related

consideration is the angle of the groove. It is possible that a 90° groove

angle would be better than 60° because the sample will seat more firmly. A

60° angle was selected for this guideline because it is more commonly used to

machine grooves.

There have been discussions about the feasibility of a standard coil

diameter and pitch. There are serious consequences to the selection of a coil

diameter. If the coil diameter is selected to allow measurements in the

smallest magnet bore that is used for I c measurements, then probability of

136
sample damage due to handling increases. Alternatively, if a larger coil

diameter is selected, a number of laboratories may not be able to participate.

There are also significant benefits to standardizing coil diameter and pitch.

This may be the most effective way to achieve consistent I c measurements.

Batch fabrication of reaction and measurement mandrels with common materials

could be cost effective and allow more interchangeable combinations of

reaction and measurement mandrels. This could also facilitate the exchange of

samples among laboratories.

With standardization of coil diameter and pitch, there are a number of

possibilities that might become cost effective. Alumina reaction mandrels is

a good example. Another option is reaction mandrels made by machining a

groove into a stainless steel tube, then spraying a coat of alumina (or some

other ceramic) onto the tube. This would be a lower cost option than

machining an alumina- tube reaction mandrel, and it would not present the

problem of diffusion bonding that an oxidized stainless steel mandrel does.

Another option that standardization creates is the use of rolled G-10 tubes

for measurement mandrels. A low wall- to -radius ratio could be selected to

give a thermal contraction that is well matched to the thermal contraction of

Nb 3 Sn. Rolled tubes would be much easier to fabricate than plate tubes. A

lot of material could be purchased and, thus, the variability of the thermal

contraction should be negligible. A rolled- tube measurement mandrel might be

appropriate for this test, since we are only trying to find a consensus method

that gives consistent results.

At this time, stainless steel seems to be the more common reaction

mandrel material, so this guideline is based on it. It is expected that the

measurement mandrel material and bonding method will be the more dominant

factors and will be tested regardless of reaction mandrel material.

REFERENCES

1. L. F. Goodrich and F. R. Fickett, "Critical current measurements: a

compendium of experimental results," Cryogenics 22., 225-241 (1982).

2. Standard Definitions of Terms Relating to Superconductors, Annual Book of

ASTM Standards, ASTM B713-82, Part 2.03, pp . 591-594, American Society

for Testing and Materials, Philadelphia, PA (1983).

3. Standard Test Method for D-C Critical Current of Composite

Superconductors, Annual Book of ASTM Standards, ASTM B714-82, Part 2.03,

pp . 595-598, American Society for Testing and Materials, Philadelphia, PA

(1983)

4. K. Tachikawa, K. Itoh, H. Wada, D. Gould, H. Jones, C. R. Walters, L. F.

Goodrich, J. W. Ekin, and S. L. Bray, "VAMAS intercomparison of critical

current measurement in Nb 3 Sn wires," IEEE Trans, on Magn. .25.(2) 2368-2374

(March 1989)

5. L. F. Goodrich, S. L. Bray, and T. C. Stauffer, "Nb 3 Sn critical-current

measurements using tubular fiberglass -epoxy mandrels," IEEE Trans, on

Magn. 25(2) 2375-2378 (March 1989).

137
138

Fig.

1.

Illustration

of

fiberglass

-epoxy

composite

geometries.
NIST-1 14A

(REV. 3-89)

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

1. PUBLICATION OR REPORT NUMBER

NISTIR 90-3935

BIBLIOGRAPHIC DATA SHEET

2. PERFORMING ORGANIZATION REPORT NUMBER

3. PUBLICATION DATE

January 1990

4. TITLE AND SUBTITLE

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Final Report

January 1988-December 1989

5. AUTHOR(S)

L.F. Goodrich, R.B . Goldfarb, and S.L. Bray

6. PERFORMING ORGANIZATION (IF JOINT OR OTHER THAN NIST, SEE INSTRUCTIONS) IT CONTRACT/GRANT NUMBER

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY —

SAmaEBSaMMKMB^2Qga9 Boulder, Colorado 80303-3328 8- type of report and period covered

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (STREET, CITY, STATE, ZIP)

Office of Fusion Energy and Division of High Energy Physics

Department of Energy

Washington, DC 20545

10. SUPPLEMENTARY NOTES

DOCUMENT DESCRIBES A COMPUTER PROGRAM; SF-185, FIPS SOFTWARE SUMMARY, IS ATTACHED.

11.

ABSTRACT (A 200-WORD OR LESS FACTUAL SUMMARY OF MOST SIGNIFICANT INFORMATION. IF DOCUMENT INCLUDES A SIGNIFICANT BIBLIOGRAPHY OR

LITERATURE SURVEY, MENTION IT HERE.)

A cooperative program with the Department of Energy, the National Institute of Standards

and Technology, other national laboratories, and private industry is in progress to

develop standard measurement practices for use in large-scale applications of superconductivity. This report describes research for the period January 1988 through December

1989. It contains the results of critical current studies on a large conductor Reference

Material (RM) , the effect of power-supply current ripple, the measurements of large conductors, and an interlaboratory comparison (round robin) of Nb^Sn wires. Short-range variations in magnetic hysteresis loss in multifilamentary Nb„Sn were studied. The results of

participation in several interlaboratory comparison are described.

12.

KEY WORDS (6 TO 12 ENTRIES; ALPHABETICAL ORDER; CAPITALIZE ONLY PROPER NAMES; AND SEPARATE KEY WORDS BY SEMICOLONS)

ac losses; circuit diagrams; critical current; current ripple; hysteresis; magnetization;

niobium-tin; niobium-titanium; power supply; round robin; standards; superconductors;

susceptibility

13. AVAILABILITY 14. NUMBER OF PRINTED PAGES

X UNLIMITED

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