AlllDE

NAn INST OF STANDARDS & TECH R.I.C.

A1 11 02788858

/Development of standards lor supercondu QC100 U5S NO.88-3088 1988 V19 C.1 NBS-P

I'tCKJ PUBLICATIONS

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Interim Report

January 1986-December 1987

LF. Goodrich, Editor

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Bureau of Standards

Boulder, Colorado 80303

FEBRUARY 1988

Stimulating America s Progress

1913-1988

QC

100

-U56

#88-3088

1988

C.2

Work performed under Department of Energy contract DE-AI01-76PR06010
r
Research Information Center

A'ationaJ Bureau of Standanh;

Gaithersburg, Maryland 2Utb y

NBSIR 88-3088

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Interim Report

January 1986-December 1987

L.F. Goodrich, Editor

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Bureau of Standards

Boulder, Colorado 80303

FEBRUARY 1988

Work performed under Department of Energy contract DE-AI01-76PR06010

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Annbler, Director

FOREWORD

The broad objective of this research program is the development of standards

for the definition and measurement of superconductor parameters. The utility

of these standards in commerce depends upon their wide acceptance and use. To

this end, an underlying objective of the program must be the maintenance of

technical credibility throughout the industrial and research communities. A

consistent demonstration of accuracy, precision, and general attention to

detail greater than that commonly achieved by the users is obviously requisite

to this goal.

A less prominent, but equally important, objective of this program is the

expansion of fundamental understanding, measurement capabilities, and the

technological base. These particular goals provide the necessary foundation

for achieving the broader program objectives and, thus, are of primary

importance

.

Finally, it is necessary to maintain flexibility in direction and emphasis in

order to effectively exploit the unforeseeable products of research v/ithin the

context of the basic program objectives.

The work described here strongly reflects this philosophy.

iii

Considerable progress has been made xn i_he expansion and rennemenr of

measurement capabilities for critical currents in practical superconductors.

Most of these improvements were motivated by specific measurement applications; however, they have resulted in a general increase in measurement

versatility. Both the expanded capaoilittes and tne resulting m.easurem;ents

and investigations are reported.

Measurem.ent problems associated with the determination of critica_ current tor

large conductors and ac loss measurements are the major components of the

work. Many of the results have already been presented at workshops and

published in the literature. Those publications not previously included in our annual reports are collected in the appendixes

Effect of Current Pyinnle on Critical Current Measurem.ents

The effect of current ripple or noise on dc critical current m;easurem.ents was systematically studied. A recent literature search indicates that this is the

first documented study of this tj'pe. A low noise, battery powered current

supply was required in this stud}’’ in order to make the pure dc critical

current measurements. Also, an electronic circuit that sim.ulates 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 rhe

introduction of controlled amounts of ac ripple. The effect of current ripple

is a reduction in the measured dc critical current; however, ripple of

sufficient amplitude can result in arbitrar]/ measurement results. The results

of this w^ork are general and quantitatively applicable to the evaluation of

critical -current data and measurement s^'s terns . A theoretical model was developed to further support and explain the ripple effect. .-.n unexpected

benefit of this work was that it led to a more precise method for general

critical-current data acquisition.

Critical Current Measurements on Cable Strands

A capabiliry to measure localized variations in critical currents was developed in this standards program. This capability allows the measurement

of the critical current profile inherent in cable strands and v;as used in

measuring strands extracted from iroTi Rutherford cables as part of another

Department of Energy program. A large variation in critical current •'up to

35%) was obsein.'ed along the length of the cable strand. Also, the voltage

distribution along the strand was shovm to be dependent upon r.agnetic field

orientation and current direction. Many voltage taps v.-ere required ir. order

to accurately construct the strand's voltage profile. Further, the

deformation geometry of the strand dictated a close spacing of voltare * ip

ojhich, in turn, resulted in a reduced voltage signal. In order to meri.sure

these low voltage signals, a low noise current supply was required tn rai:u,ii: an acceptable signal- to-noise ratio. The battery powered current supplv w.is

ideallx* suited to this application. The results of these measurements,

including suggestions for specific measurement techniques, and an effec::vtcritical current definition applicable to cabled conductors, are reported a:,d.
wiii be incorporated into future recommendations on standard measurement

methods for cable conductors.

NbTi Round Robin

The critical current of a large NbTi conductor was measured as part of a round

robin collaboration with several other laboratories. This measurement

presented an opportunity to establish and test a large conductor measurement

capability. Specifically, a high-output low-noise current supply was developed and calibrated to allow critical-current measurements of the quality

achieved previously using the low-current system. Also, superior magnetic

field calibration equipment and techniques were developed and applied to the

10 T split pair magnet for use in these measurements. The effect of current

supply ripple on these high current measurements was found to be consistent

with the low current studies. Preliminary results indicate that one of the

participants' measurements was about 10% lower than the consensus value.

These low values were attributed to a magnet calibration error and to power

supply ripple and spikes.

Nb 3 Sn Round Robin: NBS Critical Current Measurements

A 12 T, 12.7 cm diameter solenoidal magnet system was acquired (with NBS

funds) , tested, and calibrated. This system represents an increase in high

field capability and, owing to its large bore diameter, an increase in large

conductor critical-current measurement capability. The magnet calibration

techniques developed for the NbTi round robin were employed in the calibration

of this system. The system was used in the most recent Nb 3 Sn round robin

measurements. The large number of samples that were to be measured in this

round robin dictated a test fixture and related equipment that would allow

rapid mounting. Instrumentation, and measurement of the various samples with a minimum of system down time. Equipment was designed and constructed to meet

these needs. The results of these measurements indicate that sample handling

and mounting techniques continue to be a dominant variable in Nb 3 Sn critical

current measurements. In fact, a systematic study indicated that a seemingly

small change in the mounting technique can result in a 40% change in the

measured critical current at 12 T.

Critical Current Measurement System

As the design of the battery powered current supply has evolved and improved,

demands for its use by several different programs and in two adjacent

laboratories have greatly increased. Moreover, the dependence upon this

current supply has reached the point where measurements are frequently

postponed until this supply is available rather than use the commercial

supplies that are at hand. The battery supply's advantage over these

particular commercial supplies is twofold. First, it has much lower noise

and, second, it is less susceptible to ground loop problems. The latter

quality is due to its superior electrical isolation and is accomplished

through the use of an optically coupled input stage. This type of electrical

isolation is in many ways superior to previous methods and, consequently, its use is becoming commonplace in sensitive measurement devices (see Reference

[3] of Chapter V). As a result of the increasing dependence on, and use of.

VI
this supply, circuitry for its protection against electrical damage due to

misuse has been designed and implemented. Details of the overall supply

design and applications are reported.

Magnetic Evaluation of Cu-Mn Matrix Material for Fine-Filament Nb-Ti

Superconductors

The addition of Mn to the Cu matrix of f ine - filament Nb-Ti superconductors has been recently shown to reduce proximity- effect coupling between

closely spaced filaments. The reduction in coupling is thought to be caused

by magnetic impurity scattering of electron pairs by the Mn moments rather

than simply the increased electrical resistivity of the matrix. Coppermanganese, the proposed matrix alloy, is a classical spin glass. Spin-glass

effects could result in hysteresis loss and remanent magnetization in

composite superconductors, features that could counteract the reduction in

loss due to the Mn additions. We have examined some of the magnetic

properties of Cu-Mn alloys with Mn content in the range 0.16 to 3.75 wt . % to

determine whether spin-glass effects could impair the performance of finefilament Nb-Ti conductors. It appears that, for Mn content in this range, the

adverse magnetic properties of the matrix are not significant.

Coupline Losses Measured by Vibrating- Sample Magnetometry

Coupling losses in multifilamentary superconductors may be measured with

a vibrating- sample magnetometer by varying the waiting period before recording

the magnetization signal after a field change. A Nb-Ti multifilamentary wire

and a multistrand cable with different twist pitches were measured for

different waiting periods. For wire with long twist lengths, coupling losses

may be up to two times the hysteresis loss.

Vll

CONTENTS

Page

FOREWORD iii

SUMMARY V

ABSTRACT xi

I. EFFECT OF CURRENT RIPPLE ON CRITICAL CURRENT MEASUREMENTS

L. F. Goodrich 1

II. CRITICAL CURRENT MEASUREMENTS ON CABLE STRANDS

L. F. Goodrich 13

III. NbTi ROUND ROBIN

L. F. Goodrich 15

IV. Nb 3 Sn ROUND ROBINS; NBS Critical Current Measurements

L. F. Goodrich and S. L. Bray 18

V. GRITICAL CURRENT MEASUREMENT SYSTEM

S. L. Bray and L. F. Goodrich 44

VI. MAGNETIC EVALUATION OF Cu-Mn MATRIX MATERIAL FOR

FINE-FILAMENT Nb-Ti SUPERCONDUCTORS

R. B. Goldfarb, D. L. Rled, T. S. Kreilick, and E. Gregory . . 55

VII. COUPLING LOSSES MEASURED BY VIBRATING- SAMPLE MAGNETOMETRY

R. B. Goldfarb 61

Appendix A - Gurrent-Ripple Effect on Superconductive dc Critical

Current Measurements

L. F. Goodrich, S. L. Bray, and A. F. Glark ... 64

Appendix B - Hysteresis Losses in Fine-Filament Internal-Tin

Superconductors

R. B. Goldfarb and J. W. Ekin ?

IX

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

Interim Report

January 1986 -December 1987

L. F. Goodrich, Editor

Electromagnetic Technology Division

Center for Electronics and Electrical Engineering

National Engineering Laboratory

National Bureau of Standards

Boulder, Colorado 80303

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

Standards, 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 1986

through December 1987. It contains the results of critical current studies on

the effect of power- supply current ripple, measurements on cable strands, an interlaboratory comparison (round robin) on a large NbTi monolithic conductor,

and a Nb 3 Sn round robin. Several useful current supply circuits have been

developed. The reduction in coupling losses in multifilamentary NbTi

conductors has been addressed by a study of the magnetic properties of matrix

material consisting of dilute alloys of Mn in Cu. In addition, vibratingsample magnetometry is shown to be adaptable to the measurement of coupling

losses, in addition to hysteresis losses, in muitifilamentary conductors.

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

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

robin; standards; superconductors; susceptibility.

xi

I EFFECT OF CURRENT RIPPLE ON CRITICAL CURRENT MEASUREMENTS

L. F. Goodrich

The effect of sample -current power-supply ripple on the measurement of dc

critical current was reported in reference [1], which is included as Appendix

A in this report. A battery current supply was modified to allow the creation

of current ripple with variable frequency and amplitude (see Sec. V). This

was used to systematically study the effect of current ripple [1]. The

presence of current ripple (any periodic departure from a dc output level)

reduces the measured dc critical current (I^)

Iri the present report, further

calculations and comments are made on this effect that will make the results

more complete and general.

Data Acquisition

The sample voltage and current were measured at relatively fixed current

ripple and numerous dc bias currents (point method) . These points along the

voltage - current (V-I) curve were analyzed to determine I^,. This procedure is

different from the more typical procedure where the current is ramped

continuously with time whii- the voltage and current are recorded. Ramping

the current was not an option xn this experiment because the ac and dc

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 average voltage to be

measured (if voltage noise is a problem) without significantly 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

Iq. 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

Theoretical Model

The motivation for modeling the effect of current ripple on measurement of dc critical current is to help understand what is

measured and how to generalize the results to different samples,

voltmeters. As described in reference [1], calculating the time

the instantaneous, resistive voltage would model the measurement

critical current with ripple. This can be done by starting with

approximate voltage - current relationship.

the

being

ripple

average

of dc

the

and

of

V = Vo ( I / lo )^,

1
where Iq is a reference critical current at a voltage criteria of Vq

instantaneous voltage, I is the instantaneous current, and n reflects

shape of the curve with typical values from 20 to 60 (a higher number

sharper transition). The time average of this voltage (V^-^) is

V is the

the

is a

1 Vta = ( Vo / [ lo 1’^

) J dt,

0

1 where the period was set to 1 s (w=27t s’-^) in general. For a sinusoidal and

dc current, the instantaneous current can be expressed as

= [ i + a sin (27rt) = i’^ [ 1 + {a/i} sin (27rt)

where i is the dc component of the current and a is the amplitude of the

sinusoidal component. Thus, the time average voltage becomes

1 Vta = Vq [ i / Iq ]^ / [1 + (a/i) sin (27rt) dt

0

or

Vta = Vo [ i / lo G,

where G is defined as the value of the integral. Notice that = Vo when

lo^ = i^ G. With these definitions, lo is the dc critical current without

ripple and i is the measured dc critical current with ripple present. In

order to keep the equations based on measured quantities instead of , theoretical quantities, such as lo, the percent difference (D) is defined as the difference between lo and i normalized to i

D = 100 (lo - i)/i = 100 -1 ).

2
This definition is consistent with a/i ratio in the integral (the ripple is

relative to the measured quantity i)

The integral G was determined by numerical integration for a number of

values of n and a/i ratios. These results are shown for three different

magnitudes of current ripple on Figures 1, 2, and 3. The x-axis is percent

ripple amplitude (percent of a/i) and the y-axis is the percent reduction in

measured critical current (D) . The line of reduction equal to the ac amplitude was predicted [1] as the limit as n goes to 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 reduction is 100%. The amplitude of the ripple is equal to the

measured dc critical current with ripple present (a=i) and Iq = 2i.

Comparison of Model and Experimental Data

The model curves are compared to the experimental data from reference [

in Figures 4, 5, 6, and 7 for two different ranges and two different voltmeters. A few model curves are used for clarity and the experimental points

are the symbols. The two types of voltmeters used to measure the sample

voltage are described in [1]. The -60 Hz frequency was labeled negative to

indicate that the ripple current was a triangular waveform at 60 Hz rather

than a sinusoidal waveform. The experimental data fall on the model curve with an n of 40 within experimental error. The measured value of n was about

30. A model curve with an n of 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 weak

function of E) 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

±5% for ripple up to 100% using an n of 30.

The model curves can be used to estimate the effect of current ripple on

the measured dc critical current. For approximating the effect of commercial

power supply ripple on critical current measurements, the curves of Fig. 3 ma 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 ac losses may increase the size of this effect at higher

amplitudes depending upon the sample geometry [1]. Voltmeter effects will

also become more significant at higher amplitudes [1].

Superconductor Simulator

A simple electronic circuit was designed to simulate the intrinsic Vcharacteris tic of a superconductor. This circuit was used during currerr ripple studies (Appendix A) in order to characterize the response of

nanovoltmeter when subjected to a highly asymmetric periodic voltHt?--- -.s results from passing a dc biased ac current through a superconduc^or . A

general application of this circuit has been to aid in the developir' :

' testing of critical-current data acquisition systems. This circu: ? developed based on an idea by J . W. Ekin and J. C. Branch of thi.s ’ The original idea used a diode in series with a current sensing f This combination was placed in parallel with a high-current sh';n' . • '

3
Fig. 1. Model curves of the percent reduction versus the percent ac amplitude for several values of n that are listed on the right.

4
Fig. 2. Model curves of the percent reduction versus the percent ac amplitude for several values of n that are listed on the right.

5
Ic

REDUCTION

(^)

Fig. 3. Model curves of the percent reduction versus the percent ac amplitude for several values of n that are listed on the right.

6
Ic

REDUCTION

(Z)

100 N

Fig. 4. The percent reduction versus the percent ac amplitude. Model

curves are shown for several values of n and experimental data

several frequencies using an analog voltmeter.

7
Ic

REDUCTION

Fig. 5. The percent reduction versus the percent ac amplitude. Model

curves are shown for several values of n and experimental data at

several frequencies using a digital voltmeter.

8
Ic

REDUCTION

(^)

20

15 -

10 -

5 -

0 -

Fig. 6

RC RMPLITUDE

The percent reduction versus the percent ac amplitude. .Model

curves for several values of n and experimental data at several

frequencies using an analog voltmeter.

9
Ic

REDUCTION

(Z)

0 5 10 15 20

nc RMPLITUDE ('/.)

Fig. 7. The percent reduction versus the percent ac amplitude. Model

curves for several values of n and experimental data at several

frequencies using a digital voltmeter.

10
This passive circuit gives a nonlinear V-I characteristic. Unfortunately, the curve is not very sharp (n about 10) and it is not adjustable. A active

simulator circuit is shown in Fig. 8. The input of the simulator circuit

comes from a large shunt connected in series with the current supply. This

input voltage is compared with a bias voltage and then amplified. The

adjustable bias voltage serves to simulate 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 a transistor. A non-linear current

results on the emitter of the transistor. This non-linear current passes

through a blocking diode and two shunt resistors. The output of the circuit

is taken from one or the other of these shunt resistors. Typically, the

nonovoltmeter to be tested is connected to the 0.01 resistor and a recording

instrument is connected to the 100 resistor. Another channel of the

recording instrument is connected to the analog output of the nonovoltmeter.

These two channels are then compared and the nanovoltmeter may be thus

characterized under conditions similar to an 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

Reference

[1] L. F. Goodrich, S. L. Bray, and A. F. Clark, "Current-Ripple effect on Superconductive DC Critical Current Measurements," Adv. Cryo . Eng.-

Materials , Vol. 34, pp . 1019-1026, Plenum Press, New York (1988).

11
-t-iSV GAIN Cn>

Fig. 8. Schematic diagram of an electronic circuit that simulates the

intrinsic voltage - current characteristic of a superconductor.

12
II. CRITICAL CURRENT MEASUREMENTS ON CABLE STRANDS

L. F. Goodrich

Detailed critical current measurements were made on the spatial and

angular dependency of the local critical current, I^,, of strands extracted

from Rutherford cables [1,2] as part of the electromechanical DoE/HEP program.

Previous standards work on cable critical current was published as [3]. In

the present report, the implications of those results for a recommended cable

test method are discussed after a summary of the results.

Samples

A study was made of the effects of strand location and magnetic field

angle on current degradation that occurs in cables that have been compacted

into a keystone shape. The samples came from prototype cables for high energy

physics applications.

To perform this study a number of voltage taps were placed along the

length of the strand to determine its local electrical properties after it had

been deformed by the cabling process. Three regions of the cable were

studied: the thin edge . the face (top or bottom), and the thick edge . Each of

these regions was divided with voltage taps into three approximately equal

length, adjacent segments. The three segments of each cable edge were denoted

as the central flat edge and the two corners (bent portion)

Results

Results on a number of strands indicate that cabling can lead to very

localized reduction in I^, within a given strand. A weak link that represents

only 5% of the total length can contribute 95% of the total voltage.

The local critical current varies by as much as 35% along the length of a given strand at a given magnetic field orientation. This profile with

position can also change with the magnetic field orientation by as much as

15%. The Iq of the thin edge of the cable can be as much as 20% lower than

the thick edge.

The effective critical current was defined as the critical current

determined by adding the voltage contributions of each strand segment. This

total voltage was then converted to an effective electric field by dividing b” one pitch length. This was useful because it related more directly to

measurements made on the whole cable where several pitch lengths were incl- '’'d between the voltage taps.

Conclusions

These results have several implications for short-sample critical

: t testing. First, both magnetic- field orientations need to be tested to

determine the limiting case for critical current. Second, the differen- • between thick and thin edge degradation indicates that changing the dir- ’

of the test current (for each field orientation) will affect the criticurrent. This is so because the location of the peak magnetic field <h c p13
with current direction. Thus, the peak field location relative to the thin

edge should be recorded. Third, to determine an effective critical current

the voltage taps should be placed an integral number of cable pitch lengths

apart, preferably on the same strand to reduce transverse voltages. Fourth,

the current contacts should be an integer number of cable twist pitch lengths

long in order to give more uniform current injection and reduce transverse

voltages. Fifth, the legitimacy of the definition of effective critical

current being defined as a spatial average of the local needs to be

demonstrated by actual practice. Dependence upon sample stability and the

extent of local variations may need to be considered.

These results would be incorporated into any future recommendations on standard measurement methods for cabled conductors.

References

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

NbTi strands Extracted from Rutherford Cables," IEEE Trans, on Magnetics,

Vol. MAG-23, No. 2, pp . 1642-1645 (March 1987).

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

"Electro-Mechanical Properties of Superconductors for High Energy Physics

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

Colorado

[3] 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. Cryog. Eng . -Materials , Vol. 32, 833, Plenum Press, New York (1986).

14
III. NbTi ROUND ROBIN

L. F. Goodrich

NBS was invited by Intermagnetics General (IGC) to participate in a round

robin measurement of critical current on a large NbTi conductor. Other

participants were Brookhaven National Laboratory (BNL) , General Electric

Medical System Group, and IGC. The experiment was intended to measure the

uniformity of critical current test standards among several independent

laboratories. The cumulative results of the test were to be reported by IGC.

Sample and Measurement Specifications

The conductor was a large NbTi monolith with a 1.8 x 3.6 mm cross section

and a copper- to-superconductor ratio of 2.5 to 1. It had 774 filaments, each

about 52 lira in diameter. The primary measurement conditions were at 4.2 K, 6 to 7 T, and the external magnetic field oriented transverse to the conductor

axis and parallel 1119 to the wide face of the conductor. Data at two resistivitv

criteria (1 x lO'-^-^ and 1 x lO'-*-^ Q-cm) were requested. The minimum critical

current specification was 2500 A at 6 T and 1900 A at 7 T using the 1 x 10'^^

Q-cm criterion.

Several determinations of critical current were made at a number of

magnetic fields (6.0, 6.2, 6.5, 6.8, 7.0, 7.2 T) in the range of interest.

Therefore, the effect of random uncertainty should be negligible compared to

the systematic uncertainty. The total systematic uncertainty was estimated

using a straight summation of the individual uncertainties (worst case) in

magnetic field, current, temperature, voltage, strain, and current transfer.

No correction was made for the small (-0.3% amplitude) current ripple;

however, it was included in the systematic error. The aspect ratio effect was about 13.0% at 6 T and 12.6% at 7 T (higher I^, with the magnetic field

parallel to the wide face of the conductor) [1].

Measurement System

At the time of these measurements, our only magnet system suitable for

large conductor tests was a split pair magnet with a straight sample geome“r ’. This resulted in current transfer that was significant for this sample

compared to the lower criterion of 1 x 10'^^ fi-cm. Thus, there was a larger

uncertainty for the lower criterion. The total sample length was 42 cm, vi’“’ 4 cm long current contacts. The voltage tap separation was 1.95 cm, centerin a 6.35 cm diameter magnet bore. The sample was bonded with epoxy to a

’ stainless steel mandrel. A greater length of the sample in full,

perpendicular magnetic field would be necessary to reduce the curren" rra*

voltage on the taps. As of August 1987, we have a 12.2 T, 12.7 cm dia.

solenoidal magnet system purchased by NBS which will be useful for f-;r*' - large conductor tests.

System Calibration

A number of calibrations were done on the high curren^ sys'- ; * ^ round robin measurements. A voltage source and precision resis'i.i

15
were used to check the calibration of the sample voltage and current amplifier

channels. The current sensing resistor was calibrated against a lower-current

resistor that had a previous calibration. The most involved calibration was

the split pair magnet field calibration. In order to base this calibration on nuclear magnetic resonance (NMR) , an NMR calibration of another magnet

(normal, room temperature magnet) was required because the access of the

superconducting magnet was not large enough for the necessary room temperature

operation of the NMR probes. Once the normal magnet was calibrated, this

calibration was transferred to a rotating coil gaussmeter. Then the rotating

coil gaussmeter, inserted into a room temperature reentrant Dewar, was used to

calibrate the superconducting magnet. Previously, magnetic field calibrations

were done using the rotating coil gaussmeter calibrated at one point in a standard permanent magnet.

Effect of Current Ripple

A new technique for reducing the current ripple was used in this

experiment (see the Battery Powered Current Supply section of this report for

the detailed circuit). Basically it uses a low-noise, fast- response current

supply to actively filter the current from a noisy, 3000 A, commercial power

supply. Critical current measurements were made with and without ripple

reduction to allow a precise comparison to determine the effect of ripple on

the measured critical current. These measurements were made at three magnetic

fields, 6, 7, and 9 T. A summary of the comparison is given in Table I. The

amplitude of the ripple with active filtering was 0.3% for the two lower

fields and 0.4% for 9 T, the lower current case. The effect of this level of

ripple is assumed to be negligible. Without filtering, the amplitude of the

ripple was 2.4, 2.2, and 2.7% respectively. The predictions came from the

model curve of the current ripple section of this report. The measurements

and model curves (uncertainty in the value of n) are only precise to ±0.1%,

but the extra digits were carried along for clarity. These results show

further agreement between the model and experimental data to within the precision of the measurements. This also indicates that the effect of current

ripple can be accurately modeled as a percentage effect [2].

Preliminary Results

Some preliminary results were communicated to NBS by IGC. There was good

agreement between measurements made by BNL and NBS. IGC's measurements were

about 10% lower than these. The low measurements at IGC were traced to a magnet calibration error and power supply current ripple and spikes. The

magnet was recalibrated with a Hail probe. The current ripple was on the

order of 2% rms . Some of the spikes originated from near-by heavy electrical

equipment. To reduce the power supply noise and thus its effect on I^, these

measurements are now conducted during periods when the majority of the heavy

electrical equipment is not operating. As a further check of the measurement

system, the NbTi Standard Reference Material (SRM-1457) was measured and the

results confirmed the system calibration.

16
AcknowledEments

The author extend his thanks to W. E. Look for sample preparation, data

reduction, and assistance with current shunt and magnetic field calibrations.

Table 1. A comparison of measured and calculated effect of current ripple on

the measured critical current in a large conductor. Extra digits provided are

for clarity.

Magnetic

field

Ic

0.2 yuV/cm

Percent

ripple

Approximate

value of n Measureddifference

Model

difference

6 T 2983 A 2.4 40 0.55% 0.56%

7 2245 2.2 35 0.59 0.43

9 886 2.7 30 0.51 0.53

References

[1] 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. Cryog. Eng . -Materials , Vol. 32, 833, Plenum Press, New York (1986).

[2] L. F. Goodrich, S. L. Bray, and A. F. Clark, "Current-Ripple effect on Superconductive DC Critical Current Measurements," Adv. Cryo . Eng .

Materials, Vol. 34, pp . 1019-1026, Plenum Press, New York (1988).

17
IV. Nb 3 Sn ROUND ROBINS: NBS Critical Current Measurements

L. F. Goodrich and S. L. Bray

NBS has participated in three Ic round robins on Nb 3 Sn conductors during

the past three years. Two of these round robins developed from a Japan-USA

cooperation on high- field superconducting materials for fusion energy, which

started in 1980. The latest round robin developed under the Versailles

Agreement on Advanced Materials and Standards (VAMAS) , which involved

laboratories from the European Economic Community (EEC) , Japan, and the USA.

In each case, a reference wire from a country or collection of countries was supplied and individual specimens of the wire were made and distributed to all

laboratories

The first Japan-USA round robin involved four laboratories from Japan

(Electrotechnical Lab., Japan Atomic Energy Research Institute, National

Research Institute for Metals, and Research Institute for Iron, Steel, and

Other metals) and four laboratories from the USA (Brookhaven National Lab.,

Massachusetts Institute of Technology, NBS, and University of Wisconsin). The

exchange of reference wires took place in November 1985. A paper by D . C. Larbalestier and K. Tachikawa on the results of this round robin was presented

at the 1986 Applied Superconductivity Conference (Sept. 28-Oct. 3, 1986), but

it was not published.

The second Japan-USA round robin involved seven laboratories from Japan

(Electrotechnical Lab., Furukawa Electric, Hitachi, Japan Atomic Energy

Research Institute, Kobe Steel, National Research Institute for Metals, and

Research Institute for Iron, Steel, and Other Metals) and five laboratories

from the USA (Brookhaven National Lab., Lawrence Livermore National Lab.,

Massachusetts Institute of Technology, National Bureau of Standards, and

University of Wisconsin) . This round robin was set up with significant

constraints on the sample mounting and testing procedures that were designed

to determine if a reduction could be made in the measured Ic differences that were observed at high magnetic fields (H > 10 T) in the first Japan-USA round

robin test. Ultimately, the intent of this collaboration and these

constraints was to improve the measurement procedures for Nb 3 Sn conductors.

The main constraint was that- all laboratories use measurement mandrels

constructed from a supplied batch of G-10 tubes. Other suggestions were that

all samples be bonded to the G-10 mandrel with an epoxy adhesive and that the

Iq be measured for both current directions (parallel and anti -parallel

magnetic fields)

The VAMAS round robin was designed with few restrictions, to involve as much of the superconductor community as would like to participate. This

involved an additional ten laboratories from EEC; Atominstitut der

Oestereiden Universitaten (Austria), Clarendon Lab. (UK), Service National der

Champs Intenses (Grenoble, France), INFN-Sez. (Genova, Italy), University of

Nijmegen (Netherlands) , Kernforschungszentrum Karlsruhe (FRG) , Rutherford

Appleton Lab. (UK), Siemens (FRG), SCK/CEN (Belgium), Vacuumschmelze GmbH

(FRG) . With the timing of the second Japan-USA and this VAMAS round robin,

the second Japan-USA round robin became a subset of the larger VAMAS round

robin. Thus, the distinction between these two may be lost with time.

18
This section of the report is a preliminary presentation of the

measurements made at N3S Boulder for these last tvo round robins. Other N3S

measurements of critical current as a function of strain vill be presented

elsewhere. It is noted that meaningful results of these measurements depend

upon the collaboration among the laboratories; the knowledge of the results of

other participants helps the interpretation of one's own results. To this

end, all 22 laboratories that participated in the VA2-iA.S round robin are acknowledged. A round robin is a unique setting where the intercom.parison of

results is essential to the advancement of measurement accuracy. The basic

problem, that these round robin tests reveal is that the correct and/or best

method of critical current measurement for lTb 3 Sn is currently unclear. Thus,

a statement regarding the accuracy of these measurements is presently

inappropriate. A future joint publication on the results from all

participants is planned.

Sam.o les

The sam.ples in this round robin will be identified by letter only in

order to avoid the identification of commercial products by the manufacturer's

nam.e . Some physical data is provided to indicate the general type of

conductor used in these tests. In some cases these products m.ight be

indirectly identified by their particular properties. In no instance does

this imply endorsement by the National Bureau of Standards, nor does it imiply

that the particular product was the best or worst available for this purpose.

Sample X had a wire diameter of 0.68 mm and 37 sub-bundles each

containing 150 Nb filaments. The conductor fabrication m.ethod was an internal - tin, diffusion process. A single diffusion barrier of Ta separates

the filam,ent region from the outer layer of Cu. The sam;ple reaction

temperature was 700 "C for 48 h.

Sample Y had a wire diameter of 1.0 mm and 7 sub -bundles each contain i ’'g 721 Nb filaments. The conductor fabrication method was a bronze, diffusion

process. Each sub-bundle had a Nb diffusion barrier and the m.atrix was a CuSnTi alloy. The sample reaction temperature was 670 °C for 200 h.

Sample Z had a wire diameter of 0.8 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 t’'- center of the wire separated by a Ta barrier. The sample reaction tempera" •' 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

reacted at each individual measurement laboratory ( self reaction . A

comparison of results from these two groups would give an indic«i"i ‘V. f reaction and shipping variables. Each specimen was reacted on a s"a:rT

steel tube that had a spiral groove on it to define the coil pitch. I • surface of the stainless steel had been oxidized to reduce the cl t--- i diffusion bonding the sample to the reaction mandrel. If bondirg did

the sample could be dam.aged when it was removed for mounting or.t" •* measurement mandrel. The self reactions performed at NBS were d ;•

19
three -zone, vacuum tube furnace. The temperature of the furnace was controlled during the reaction to within ±5 °C. The furnace temperature

calibrations checked and was within ±10 °C.

Apparatus

The measurement mandrel was constructed from a G-10 tube 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 on to 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 reattached and the sample was tightened onto the

mandrel. Each end of the sample was soldered onto its current contact. An

effort was made to avoid a complete loop made by the sample and bus bar on the

contact. This was done in order to circumvent slowly decaying persistent

currents. 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. At this point 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 so that the thermal conductivity from the sample to

the helium bath was not as limited. This unit could then be mounted onto the

sample cryostat 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. Since the ambient atmospheric pressure at

this test site is low, resulting in a 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 deduced using

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 taken

simultaneously on a digital processing oscilloscope. An analog nanovoltmeter,

with an amplified output, was used to measure the sample voltage of each tap

pair. A battery powered current supply was used for the sample current. The

sample voltage was measured at numerous bias currents. These points along the

voltage-current curve were analyzed to determine I^. These measurements were made with forward (anti-parallel magnetic fields, Lorentz force in) and reverse current directions in order to determine the effect of self- field and

sample - to -mandrel bonding (strain effect). Unless otherwise stated, the

values given in this text are averages of all taps, observations, and current

directions at 10 /iV/m. The precision of the measurement was generally

about ±0.2%.

Another parameter that was determined during the 1^, measurement was n. The parameter n is defined in the approximate voltage - current relationship.

V = Vo ( I / lo )^^,

20
where Iq is a reference critical current at a voltage criteria of 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 is a sharper transition)

The 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 yi/V/m) . This resulted in more variation in

the values due to the small voltage range that could be analyzed. 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 measurement mandrel, or making the transition

from the current contact to the G-10 body of the mandrel. The transition

between the contact and the G-10 has two intrinsic problems. One problem

arises from the differential 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 more

sample pre-strain. Difficulties in measuring the Z sample may have been due

to its high current density (535 A/mm^ at 10 T) and bronze surface layer

(larger contact resistance).

Preliminary Results

A summary of these data on sample X is given in Table 1 and Figs. 1, 2,

and 3. 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 labeled with a letter from A to G , is a data set taken for a single thermal cycle of a specimen indicated by a number from 1 to 4 . Some specimens, such as cycles A,

B, and G for specimen 1 and also cycles E and F for specimen 3, were thermally

cycled (between room and liquid helium temperatures) and measured again. The

sequence of the thermal cycles was form left to right in the table. Specimens

are numbered 1 through 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 12 T.

The key words heading the columns of Table 1 (central, thick, thin, etc. are used to indicate the measurement variables. Gentral indicates reaction o! the specimen at a central laboratory and self indicates reaction at the

individual measurement laboratory. Thick and thin refers to the ratio of wa’

thickness to outer radius of the sample mandrel. Under the Sample Mandrel

Dimensions section of the table, these specific values are listed for ea 1. data set. Bored refers to a specimen that was originally mounted and neci > > on a thick walled mandrel. The specimen was then brought to room temper "• (thermally cycled) and the inner diameter was bored to a larger dimensior,

resulting in a reduced wall thickness. The specimen's was then rerr- The typical method for bonding samples to measurement mandrels was to • • sample and the entire surface of the mandrel with epoxy. Two variati

this technique are indicated by spiral and varnish . Spiral indi-atfs * \ - n

sample was bonded to the mandrel with a thin stripe of epoxy along i* \

21
Table 1 . Data on Sample X

Critical current for Sample X

Specimen 1 1 1 2 3 3 4

MoH. 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

T Delta

(A-B)/B

Delta

(B-O/C

Delta

(C-D)/D

Delta

(E-D)/D

Delta

(F-D)/D

Delta

(G-D)/D

6 0.0% -11.8% 2.8% 0.7% 0.1% -3.1%

7 0.3% -14.3% 3.1% 1.1% 0.5% -3.6%

8 -0.2% -16.9% 3.2% 1.5% 0.7% -4.1%

9 -0.1% -20.7% 3.4% 2.0% 1.1% -4.5%

10 -0.4% -25.6% 3.6% 2.6% 1.5% -4.8%

11 -31.5% 3.7% 2.4% 0.8% -4.7%

12 -39.2% 3.9% 2.6% 0.9% -7.2%

22
(V)

Fig. 1. Ic at 10 /iV/m versus magnetic field for specimens of sample X.

23
Fig. 2. n value versus magnetic field for specimens of sample X.

24
[(I+)-(I-)]/Iav0

+0. 15

+ 0 . 10 -

Ec=10 ^\J^-

CENT THICK

CENT BORE

CENT THIN

SE THN' SPL

SE THN

SE THN VAR

+0. 05 -

-0. 05 h

- 0. 10 -

•0. 15

7 8 10 11

Field (T)

Fig. 3. Percentage difference between measured with reverse and forward

current directions versus magnetic field for specimens of sample X

25
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 of data in columns A and B is a measure of 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-B]/B).

This small variation upon thermal cycling was representative of all observations. Column C was a special case of a thin wailed 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 of a particular specimen. The intercomparison data ([B-C]/C)

shows a strongly field dependent variation of (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 thick walled tube.

Strain sensitivity measurements on this sample indicate that this additional

pre-strain would result in an degradation on the order of 33%. The

composite nature of G-10 [1, 2, and 3] in conjunction with the compound

geometry of a tube may result in a variation in radial thermal contraction

between tubes of different wall thicknesses. 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. The rest of the intercomparisons in

this table were made relative to D ( central . thin ) since it was thought to be

a more ideal baseline for intercomparisons. 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.

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 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 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 less permanent adhesive. The 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 after the reverse measurements were made. Figure 4

is a plot of the percentage difference of the observed forward from the

average at each magnetic field. For these data between 6 and 11 T, the

26
5

TRP SYMBOL

1 O

o o

A A

8 10

MRGNETIC FIELD (T)

Percentage difference of observed versus magnetic field for

specimen G ( Self , thin , varnish ) of sample X.

27

'

^
variation of measured among the pairs of voltage taps was typical of

specimen homogeneity. The indicated precision of the repeat determinations

was also typical of these measurements. The first observations at 12 T are

the lower group of data points. The upper group of data points were taken

after the reverse measurements. This irreversible enhancement can also be seen in the discontinuity of the intercomparison of G and D in Table 1. 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.

Figures 1, 2, and 3 provide a graphic summary of these data on sample X. Figure 1 is a semi - logarithmic plot of versus magnetic field. A semilogarithmic plot was selected because it better illustrates the percentage

differences between the curves . The four upper curves are in approximate

agreement with data taken at other laboratories. The central . thick curve is

significantly lower than the others and the thin, varnish curve was slightly

lower. Figure 2 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

insignificant for this measurement variable. Figure 3 is plot of the

percentage difference between reverse and forward versus magnetic field.

The thin . varnish curve illustrates the large strain effect, 13% at 12 T, 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.

A summary of these data on sample Y is given in Table 2 and Figs. 5, 6,

and 7. The centrally reacted specimens were 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 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 12 T)

Data sets E and F, self reacted specimen number 4, are compared to data

set D and again there is a similar discrepancy (8.7% at 12 T) between two

comparable specimens. The additional possible explanation for this is the

28
Table 2. Data on Sample Y

Critical Current for Sample Y

Specimen 1 2 3 3 4 4

Central Central Central Central Self Self

0 thick thick thick thin thin thin

bored

(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

w/r 60% 60% 60% 16% 13% 13%

Intercomparison of Critical Current !Data

Delta

(A-C)/C

Delta

(B-O/C

Delta

(C-D)/D

Delta

(F-D)/D

Delta

(E-D)/D

6 0.3% -6.2% -11.3% -4.2%

7 -0.1% -6.6% -11.8% -4.9%

8 -0.2% -6.7% -12.6% -5.6%

9 -0.1% -6.5% -13.8% -6.3%

10 -7.0% -14.5% -7.1%

11 -7.2% -15.8% -7.8% -8.2%

12 -7.5% -17.3% -8.7%

29
Fig. 5. Ic at 10 /iV/m versus magnetic field for specimens of sample Y.

30
Field (T)

Fig. 6, n value versus magnetic field for specimens of sample Y.

31

f\j
[(I+)-(I-)]/Iave

+ 0. 15

Ec=10 uV/m

+ 0. 10 -

CENT THK C

CENT BORE

SELF THIN

CENT THK A

CENT THK B

+ 0 . 05 ^

+ 0 . 00 -

-

0 . 10 -

- 0 . 15

Fig. 7. Percentage difference between I^. measured with reverse and forward

current directions versus magnetic field for specimens of sample Y.

32
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 cycle were induced by one of

several magnet quenches. And, as in the previous cases, the effect of thermal

cycling on appears to be about 0.4%.

The most significant result for this sample is, again, the dependence of on mandrel wall thickness.

Figures 5, 6, and 7 provide a graphic summary of these data on Sample Y.

Figure 5 is a semi - logarithmic plot of 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 scatter 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 variation observed on 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 versus magnetic field. In this case the maximum

difference was only about 1%. These curves are similar to those for sample X,

Fig. 3, and indicate that the specimen was well bonded to the mandrel.

Figures 8 and 9 are plots of the percentage difference of the observed

forward from the average at each magnetic field. Both of these plots

contain data for specimen 3, the difference being that Fig. 8 contains

central - thick - C data and Fig. 9 contains central - thick - bored -D data. The

significant difference between the two plots is that the tap-to-tap

variation, observed in Fig. 8, was not present after boring as shown in Fig.

9. One possible explanation is a reduced longitudinal variation in G-10

thermal contraction after boring. The variation of measured for the

other specimens was similar to that of Fig. 9.

A summary of these data on sample Z is given in Table 3 and Figs. 10, 11,

and 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 at 12 T. The centrally reacted specimens were measured on mandrels that had an outer diameter of about 19 mm rather than the typical ~1 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

Figures 10, 11, and 12 provide a graphic summary of these data on Sample Z. Figure 10 is a semi - logarithmic plot of versus magnetic fi.

Figure 11 is a semi - logarithmic plot of n versus magnetic field. Tie

significant information contained in this plot is the relatively hi,*. ri for this conductor. Figure 12 is a plot of the percentage diffe:»rr

forward and reverse versus magnetic field. In this case the ii: < - : difference was only about 4%. With the possible exception of the -<• If

specimen, these curves are similar to those for samples X and Y. hi ’’

33
Fig. 8, Percentage difference of observed Ij, versus magnetic field for

specimen C ( Central . thick ) of sample Y.

34
MRGNETIC FIELD (T)

Fig. 9. Percentage difference of observed I^. versus magnetic field for

specimen D ( Central . thick , bored ') of sample Y.

35
Table 3. Data on Sample Z

Critical Current for Sample Z

Specimen 1 2 3

/i H, T Central Central Self

0 (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

w/r 32% 32% 13%

Intercomparison of Critical

Current Data

MoH, T Delta

(A-B)/B

Delta

(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%

36
600 T

O

180

6

Ec-10 uV/m

CENT A

CENT B

SELF -

_L

7

J 1 I L. 8 9 10 11

Field (T)

Fig. 10. Ij, at 10 /iV/m versus magnetic field for specimens of sample Z.

37
Field (T)

Fig. 11. n value versus magnetic field for specimens of sample Z.

38
[(I+)-(I-)]/Iave

+ 0 . 15

+0. 10 h

^ 0. 05 !-

+ 0. 00 f*

- 0 . 05

- 0 . 10 ^

-

0 . 15

Ec=10 uV/m

CENT A -

CENT B

SELF —

J ^ ^ ^

J

7 8 9 10 11 12

Field (T)

Fig. 12, Percentage difference between measured with reverse and forward

current directions versus magnetic field for specimens of sample Z

39
7, and indicate that the specimen was well bonded to the mandrel. Within the

limits of the measurement 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 G-10 glass-epoxy composites

have been published for the plate geometry [1, 2, and 3] but, not for the tube

geometry. The contraction of a G-10 plate is highly anisotropic. In

composite nomenclature, 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 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 thick walled G-10 tube was selected as a common sample mandrel material

in order to allow one size tube to be used by all the participants of the

Japan-USA round robin. A thick walled tube is difficult to manufacture

without getting 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 measurements below were the average of

two approximately orthogonal diameter measurements.

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 gave a preliminary 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 have 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 G-10 tubes

having different wall thicknesses. Specifically, the thin walled tube's

40
radial thermal contraction was -0.23% while the thick walled tubes 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. Using the temperature dependence of the thermal

contraction of G-10 plate [2 or 3], the thermal contraction to 4 K was estimated as -0.28% for the thin walled tube and -0.48% for the thick walled

tube. 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.

For samples X and Y, considerable tensile - strain- sensitivity data was available; however, no explicit data were available for compressive strain

sensitivity. Consequently, the strain scaling law [4] in conjunction with the

available tensile strain data was used to estimate the expected degradation

for compressive strain. For Sam.ple X the results of these calculations

predict a -33% reduction in the 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

degradation of -24% at 12 T. The actual reduction in for the thick mandrel

as compared to the thin was -18%. As stated, these measurements and

calculations are only approximate but the trend seems to be obvious. In all

cases, the thick walled samples had reduced I^'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 of the measurement. Two potential

sources of damage are specimen shipment, where the packaging of reacted

samples is inadequate, and specimen 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. Damage

introduced by specimen transfer is easier to detect and evaluate than the

other source but, if specimen damage during transfer is to be avoided,

dexterity and patience are required. Unthreading the specimen from a gro

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 ; they are unthreaded unless some torque is applied to each turn with 'h' p of the hand. Some observations on packaging methods that were use*’; f

" shipping reacted specimens are listed here. It is not possible in <;• - to separate the effect of shipping damage from the effect of central v- rs self reaction or from sample homogeneity. The reaction mandrel alone a: r protect a sample from shipping damage. The packaging needs to pro'f f- * , sample from collisions with external objects as well as from in- rfia!

41
collisions with other specimens within the container or with the container.

The damage or loss of samples during shipping and the delay of sample

shipments by customs are characteristic problems of international round robin

measurement and, hopefully, with experience these problems will be reduced.

The study concerning the effect of boring and thick- versus - thin walled G10 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. Any future publication will most likely contain a more condensed version of these results to give proper weight to the significant

and the definitive.

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 Nb 3 Sn. Thick wall G-10 tubes contract about 0.2% more than

thin wail 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 method of bonding the sample to the

mandrel also determines how much of the differential 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 G-10 was about 20% at 12 T for a Nb 3 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.

G-10 tubes with wall thickness less than about 20% of the tube radius could be

used for the measurement mandrel and they would be well matched in thermal

contraction to the sample. Do not use thick walled G-10 tubes with voids and

machine them to make a measurement mandrel because the voids will cause irregular thermal contraction. A well oxidized stainless steel reaction

mandrel should be used to reduce diffusion bonding of the sample to the

mandrel and the resulting 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.

The results presented here are preliminary in the sense that they have

yet to be correlated with measurements from all 22 participating laboratories.

This correlation is a necessary element in a study of this type and may

42
clarify the results presented here. This second round robin indicates why

studies of this type need to be iterative. For example, it was anticipated

that by simply using the same mandrel material, different laboratories would

arrive at the same, correct results. However, due to differences in

bonding technique, the results were different. This indicates that in

order for various laboratories to arrive at the same Ij, results, a more detailed unification of the measurement techniques would be required.

However, this would not insure the correctness of the results. If, for

example, all of the participants of this study would have used a continuous

coat of epoxy to bond their samples, the results may have been in close

agreement but the results would have been lowered by the increased

conductor pre-strain. A unified definition of the desired measurement 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 for accuracy.

Acknowledgments

The authors extend their thanks to J . W. Ekin for discussions 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 all 22

laboratories that participated in these round robin measurements.

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. Cryog.

Eng . -Materials , Vol. 2A> 589, Plenum Press, New York (1980).

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

V3 Ca Multifilamentary Superconducting Cables, Fiberglass - Epoxy and

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

L267-L270 (1981).

[3] A. F. Clark, G. Fujji, and M. A. Ranney, "The Thermal Expansion of

Several Materials for Superconducting Magnets," IEEE Trans. Mag.

MAG-17 . 2316-2319 (1981).

[4] 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).

43
V. CRITICAL CURRENT MEASUREMENT SYSTEM

S. L. Bray and L. F. Goodrich

Power supply modifications to reduce the amount of current ripple and

thus improve the overall accuracy and precision of critical current measure- ments were made. Also, additional modifications were made to systematically

study the effect of power-supply current ripples. These circuit designs and

innovations will be useful for and applicable to critical current measurements

of all kinds.

In order 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

having these qualifications has been constructed and tested [1]. In addition

to its normal operation as a constant-current source, two other unique modes

of operation have been developed and tested. In one mode, the supply is used

as a source of dc biased ac currents for testing the effect of dc -power- supply

ripple on critical current measurements [2]. In the other mode, the supply is

used in conjunction with a larger and noisier current source to form a hybrid

supply. In the hybrid system, the smaller supply is used to increase the

current output while actively filtering the noise output of the larger supply.

Finally, systems for rapid quench (thermal runaway of the sample into the

normal state) detection and supply reset have been incorporated in the design.

The purpose of a quench detector is to prevent damage to the sample due to

resistive heating. The specifics of all designs, modes of operation, and

performance results are given.

Low-Ripple Current Supply

This is an experimental design that has been gradually developed through

continuous modification and improvement for an explicit application. Consequently, the specifics of the design should be considered quite flexible

depending on the particular application. The performance of the circuit under

different load conditions and with various input signals is being evaluated.

Although the only application for which this supply has been tested is superconductor measurement, many other possible uses for a current supply of this

type may exist.

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 1350 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 115 Vac/15 Vdc supply is used at the input stage of the

supply circuit. The power stage of the. circuit utilizes 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 through a separate 1 mQ 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

44
Fig. 1. Circuit diagram of the battery powered current supply in the low

noise mode.

45
single 1 mfi 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 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 0.1 mfi resistor connected between the negative battery

terminal 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 upon an operational amplifier

(0A2) 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 (0A3) is

used as a summing amplifier in order to combine and amplify the six signals

from the feedback resistors. In order to avoid an offset current from the

supply when the input signal is at zero volts, a trim potentiometer is used on

this op-amp. This potentiometer is adjusted such that the normal output of

the op-amp is slightly negative. This, in turn, results in a negative output

signal from the comparator (0A2) and, thus, no offset current from the supply.

Since the summing amplifier inverts the feedback signal, another inverting

amplifier (0A4) 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 opticoupler (OCl). The input stage is linearized in a manner similar to that of the control stage (see page 66 of [3]). Again, a comparator (OAl) is used to compare the reference and feedback signals. The

output signal from the comparator is amplified by an single transistor (Tl).

The feedback signal is supplied by a second opticoupler (0C2) which is in

series with the first. This is necessary in order to provide electrical

isolation. Moreover, the second opticoupler 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 feedforward opticoupler (OCl) 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 connecting this unused base terminal to the

emitter terminal with a 0.1 ixt capacitor.

The performance in this low current ripple mode of operation was tested

with two load resistances. In the first case a low- inductance , 50-mQ 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 currentproportional voltage signal for accurate measurement of the current ripple.

46
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 mfi was used to allow the

supply to be operated to high currents. The voltage signal from a 0.1 mQ

resistor 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.

Quench Detector with Rapid Reset

The control stage of the current supply contains several reactive elements

that enhance stability and reduce noise. These elements, however, result in a time lag between the input signal and the supply's response. Traditionally,

sample quench protection has been achieved by amplifying the sample voltage

[4] and using this signal to reset the control signal generator in the event

of a quench. However, this time lag between control signal reset and current

interruption is great enough to jeopardize the sample. Consequently, a method

for rapid current interruption was devised and implemented. Figure 2 is a schematic diagram of the quench protection circuit. This circuit is a combination of an adjustable amplifier stage, an absolute value stage, an inverter, an electronic timer, a single transistor current amplifier, and an optically coupled output stage. The amplified sample voltage is input to the

absolute value stage which accommodates bipolar sample currents. The inverter

provides the required high-to-low transition for the electronic timer. The

timer holds the protection circuit output in the high state long enough to

insure that the control signal generator will be reset. The current amplifier

drives the output stage and the output stage provides electrical isolation

between the quench protection circuit, the control signal generator, and the

current supply. The signals sent to the control generator and the current

supply by the quench protection circuit are high-to-low resistance

transitions. In the case of the current supply, this output is connected

between the base of transistor T2 (Fig. 1) and ground. The transition from

high to low resistance, thus, interrupts the supply current by driving T2 into

cutoff

This circuit is capable of interrupting the supply current in approximately 50 /is , which is more than adequate for insuring sample protection.

ac and dc Current Supply

In order to test the effect of dc-power-supply ripple on critical current

measurements [2], the basic current supply was modified to allow the generation of dc biased ac currents. There were two basic areas of modification.

First, the circuit had to be altered to accommodate simultaneous cont; ’ by two separate input signals, a linear ramp and a sinusoid. The ramp .sign-t'-

provided the dc offset current while the sinusoid provided the ac ripple

current. Figure 3 is a schematic diagram of the circuit used to gene.-_jt

composite control signal. The two signal generators, ramp and ac . a-^e isolated from one another by unity gain operational amplifiers, OAl ai The two input signals are then combined by the inverting summer. QA . Another inverting summer, 0A4, is used to restore the polarity of rb47
Fig. 2. Quench detector circuit with rapid shut-down of the current

supply.

Fig. 3. Schematic diagram of the ac and dc current supply.

48
This signal is then applied to the input of the current supply's control

circuit

The second modification involved the ac response of the control circuit.

The basic control circuit described above contains several filtering elements

that are useful for noise reduction and stability; however, these elements

greatly attenuate the desired ac output current in this mode of operation.

Consequently, these filtering elements were altered to enhance the ac response

of the supply. The resulting increase in current noise was insignificant in

comparison with the deliberately generated ac current and, thus, was not

considered detrimental in this application. Specifically, the 0.01 /iF

feedback capacitor connected between the inverting input and the output of OAl

(see Fig. 1) was reduced to 82 pF. Also, the 0.01 fj,F capacitor connected

between the base and emitter of OCl was removed. The 3 /jF capacitor

connecting the inverting input of 0A2 to common was reduced to 0.02 fiF

Finally, the 1 /iF feedback capacitor connected between the inverting input and

the output of 0A2 was reduced to 0.1 nF

These changes gave the desired

increase in ac response while maintaining stability. The intrinsic (no input

ac voltage control signal) current ripple in this mode was about 0.5 A peakto-peak for dc output currents from 0 to 400 A.

In this mode of operation, the maximum tested frequency was 10 kHz, ac amplitude was 275 A peak- to -peak , and dc bias was 400 A (not all at the same

time). Moreover, the maximum output voltage of the current supply is limited

to less than 4 V by its battery voltage, the Impedance of the current path

between the battery and the load, the collector-emitter saturation voltage of

the output transistors, the dc and ac levels of the output current, and by the

frequency of the output current. If the Impedance of the load and the ac and

dc output current levels are such that the load voltage reaches the maximum

available output voltage, saturation results. In these particular tests

measures were taken to reduce the load impedance and, thus, extend the

operating limits of the supply.

Hybrid Supply

Although the low noise character of the battery powered current supply is

very desirable for critical current measurements, its most notable limitation,

particularly for large conductor measurements, is its maximum output capability of approximately 950 A. Unfortunately, high output commercially available current supplies often have high noise levels that limit their suitability for critical current measurements. In light of this, a method for using

the battery powered supply in conjunction with a larger current supply to form

a high output and low noise hybrid supply has been developed and tested.

The high-current supply that was used in these tests is a 3000 A, thr-^e

phase, SCR switching supply. Like the battery supply, this supply is

controlled by an external input signal and uses feedback to maintain a line.i.

input-output relationship. Unlike the battery supply, the 3000 A supply h ^ substantial output noise level. This noise is predominantly harmonics of

60 Hz and has a peak-to-peak amplitude of 144 A (36 A rms) at a dc outpui f 3000 A. The relationship between the peak-to-peak and rms values will not.

49
the usual one because the noise is not sinusoidal. This noise level is

sufficiently large to cause errors in dc critical current measurements [2].

Figure 4 is a block diagram of the hybrid current supply. As shown, the

control circuits of both the battery supply and the 3000 A supply receive two

input signals, one from the ramp generator and one from the 3000 A shunt

resistor. The ramp signal controls the dc current from each supply and the

shunt resistor provides feedback signals. In the case of the 3000 A supply,

this feedback signal is simply used to maintain a linear input-output

relationship. The feedback signal sent to the battery control circuit

encounters a dc blocking capacitor that allows only the ac , or noise, portion

of the signal to pass. This noise signal is inverted and combined with the

ramp signal to control the output of the battery supply. As indicated in the

diagram, the two supplies are connected in parallel and, therefore, the total

current output is the sum of the individual supply outputs. Since the two

supplies put out similar noise signals that are of opposite polarity, this

summation tends to cancel the total output noise from the hybrid supply.

In order for the battery supply to simultaneously decrease the noise

output and increase the dc output of the 3000 A supply a different control

circuit than those previously described was required. Ideally, in this

application, the battery's behavior should be as follows. As the ramp input

signal becomes nonzero, the output current from the battery supply abruptly

increases from zero to a relatively low dc bias level in order to allow ac currents without clipping (the battery supply is unipolar and can only produce

positive currents). As the output current from the 3000 A supply increases,

its noise level also increases and the resulting negative feedback signal to

the battery supply causes its output to reduce this noise. Finally, as the

3000 A supply nears its peak output, the battery supply's dc output ramps

quickly to its peak value such that both supplies reach their peak output

levels simultaneously. This complex behavior of the battery supply is only

necessary in order to limit its duty cycle. Otherwise, the battery supply

could be held at peak output (less the required ac peak output) throughout the

majority of the 3000 A supply's ramp cycle. The battery control circuit was designed to minimize the supply's output at all times and, thus, prolong

discharge of the battery cells. It is important to note that this nonlinear

behavior of the battery supply does not result in a nonlinear output from the

hybrid supply. This is due to the fact that the battery supply is contained

within the 3000 A supply's control loop. In other words, the 3000 A supply

compensates for nonlinear variations in the battery supply's contribution to

the total output current.

Figure 5 is a diagram of the control circuit used to achieve this

nonlinear behavior. Beginning at the output side of the control circuit, the

original power circuit of Fig. 1 is driven by the output of an optical

isolator (nonlinear) . The isolator is driven by an inverting summer which

combines three input signals. The first input signal is the 3000 A shunt

resistor feedback signal as described above. The second signal is a constant

voltage obtained from a Zener diode. The gain of the amplifier that drives

the Zener diode is high in order to insure early zener breakdown. This signal

provides the current-bias. The potentiometer that is connected to the Zener

diode allows adjustment of the bias -current level. The input to the bias

50
Fig. 4. Block diagram of the hybrid current supply.

DC

Fig. 5. Schematic diagram of the control circuit used to filter and boost

the current.

51
portion of the control circuit comes from the ramp generator through an optical isolator. The third input signal to the summer is the current boost

signal. In this case, the output potentiometer provides adjustment of the

current gain or, in other words, determines the ramp rate of the battery

supply output. The signal to the potentiometer is provided by an inverting

amplifier that has an adjustable threshold voltage. This threshold voltage

determines the point at which the boost current begins its ramp cycle. The

diode between the current boost potentiometer and the amplifier blocks the

positive polarity signal that would result from the threshold voltage being

greater than the input voltage. Again, the input voltage is provided by the

ramp generator. The threshold and current gain adjustments must be made in a coordinated manner such that the battery supply peaks in unison with the

3000 A supply. When the circuit is adjusted correctly, the result is a low

noise linear ramp current that peaks near 4000 A. The peak-to-peak noise

level of the composite supply has been measured at 17.6 A (3.5 A rms) for an output of 3000 A. This is a 88% reduction in the peak-to-peak noise and a 90%

reduction in the rms noise.

Battery 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 mfi collector resistor of Fig. 1. Consequently, a method of overcurrent

protection that uses thermal cutoff switches on both the collector resistor

and the output transistors has been designed. Referring to Fig. 6, under

normal operating conditions all four thermal cutoff switches (TC1-TC4) are in

the closed position. This 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 one of the four thermal cutoffs exceeds its set point, the

relay current is interrupted and the relay switches to the other position. In

this position, input to the Darlington output stage is connected to common and

the output current is interrupted. The diodes in series with the relay

current are used to prevent undesired relay energization in the unlikely case where two shorts between the 4 V battery and the thermal cutoff switches occur

thus preventing the current interruption from occurring. These modifications

are now under way and will be tested upon completion.

Acknowledgments

The authors extend their thanks to W. P. Dube for numerous discussions

and being a constant source of advice, to R. M. Folsom for creating circuit

diagrams, and to H. E. Spilman for circuit testing.

52
Fig. 6 Schematic diagram of the protection for the current supply.

53
References

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

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

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

(April 1987) .

[2] L. F. Goodrich, S. L. Bray, and A. F. Clark, "Current-Ripple effect on Superconductive DC Critical Current Measurements," Adv. Cryo. Eng.-

Materials, Vol. 34, pp . 1019-1026, Plenum Press, New York (1988).

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

[4] W. P. Dube and L. F. Goodrich, "Quench detector Circuit for

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

54
VI. MAGNETIC EVALUATION OF Cu-Mn MATRIX MATERIAL FOR FINE-FILAMENT Nb-Ti

SUPERCONDUCTORS

R. B. Goidfarb and D. L. Ried

National Bureau of Standards, Boulder, Colorado 80303

T. S. Kreilick and E. Gregory

Supercon, Inc., Shrewsbury, Massachusetts 01545

Abstract

Copper-manganese alloys have been proposed as matrix material for the

reduction of coupling losses in fine-filament Nb-Ti superconductor wires.

Magnetization and susceptibility measurements show that adverse magnetic

effects arising from the spin-glass properties of this matrix are minimal for

concentrations of Mn up to at least 4 percent.

Introduction

The addition of Mn to the Cu matrix of f ine - filament Nb-Ti

superconductors has been recently shown to reduce proximity- effect coupling

between closely spaced filaments [1,2]. Small interfilament separation

promotes the formation of filaments with uniform diameter [3,4], thus avoiding

the deleterious effects of "sausaging" [5]. The result is a large total Nb-Ti

cross section (and therefore a large critical current for the composite wire)

with small filament diameters (to minimize hysteresis losses)In the absence

of Mn, coupling losses can offset the reductions in hysteresis losses that are achieved by using fine filaments. The reduction in coupling is thought to be

caused by magnetic scattering of electron pairs by the Mn moments [1,2, 6, 7]

rather than simply the increased electrical resistivity of the matrix [1,8].

Copper-manganese, the proposed matrix alloy, is a classical spin glass,

characterized by magnetic viscosity and anisotropy that develop below a "spinfreezing" temperature. Several review articles on spin glasses have been

published [9-17]. Spin-glass effects could result in hysteresis loss and

remanent magnetization in composite superconductors, features that could

counteract the reduction in loss due to the Mn additions

We have examined some of the magnetic properties of Cu-Mn alloys with .Mn content in the range 0.16 to 3.75 mass percent (wt.%) to determine whether

spin-glass effects could impair the performance of fine- filament Nb-Ti

conductors. It appears that, for Mn content in this range, the adverse

magnetic properties of the matrix are not significant.

Spin-Glass Properties

A spin glass exhibits a peak in magnetic susceptibility as a function

temperature which is associated with a "freezing" of magnetic moments.

Freezing is a relative concept, depending on magnetic field strength and

in addition to temperature [18]. Above the freezing temperature, Tf, t!. alloy behaves paramagnetically [19], sometimes with large moments. v -r

55
below Tf, magnetic relaxation effects appear [20], In this paper, we do not

discuss the time (or frequency) dependence of the magnetic properties.

Applying a magnetic field to a sample at 4 K results in an isothermal

remanent magnetization, IRM, measured after reducing the field to zero. The

IRM is a function of the maximum applied field. Cooling the samples from

above Tf to 4 K in a field, and then reducing the field to zero, gives a thermoremanent magnetization, TRM. The TRM is also a function of field. For

a large enough experimental field, the IRM will equal the TRM [21]. Fine

details of the hysteresis loops of magnetization vs. field near the origin for

dilute Cu-Mn alloys are discussed in Refs. [22] and [23].

Spin glasses have three properties that might affect their suitability

for use as matrix material for multifilamentary superconductors. The

magnitude of each of these effects would be a function of the volume of the

matrix that contained the Mn, usually just the interfilamentary space.

First, large magnetic moments could produce a significant magnetization

at high fields. The field profile of a superconducting magnet could be

distorted if it were made with wire containing such matrix material.

Second, if the operating temperature of the wire is below Tf, there could

be a large IRM in the matrix material after the superconducting magnet is

cycled to high fields. (The direction of the IRM would be the same as that of

the trapped flux in the superconductor.) Third, if the magnet were intended

to be cycled, hysteresis in the matrix material could be a problem. These

last two possibilities could partially defeat the purpose of using wire with

fine filaments.

(The field-cooled unidirectional anisotropy at 4 K in Cu-Mn is not fixed

very well compared to other spin glasses; it can be switched by moderate

fields. As a consequence, there is a symmetrical hysteresis loop after

exposure to moderate fields. It is a "false" hysteresis due to the switching

of the unidirectional anisotropy as the field is cycled. The reader

interested in this point is referred to Refs. [21] and [24].)

Proximity Effect

Proximity- effect coupling arises when interfilament separation is on the

order of the coherence length, in the matrix (approximately equal to the

electron-pair decay length in the "dirty limit"). Manganese additions to the

matrix cause magnetic scattering of the electron pairs and a reduction in

In the spin-glass state, however, there appear to be two opposing tendencies

[6]. On the one hand, the freezing of the Mn spins is thought to inhibit

electron scattering, tending to increase On the other hand, the exchange

interactions in the spin-glass state could enhance electron-pair breaking,

thereby reducing The latter effect appears to dominate and, generally, £

is inversely proportional to Mn content [6].

56
Measurements

Measurements were made on Cu-Mn alloys containing 0.16, 0.72, 2.00 and

3.75 wt.% Mn (0.19, 0.83, 2.31, and 4.31 at . % Mn) . Manganese content was checked by dc plasma emission spectroscopy by an independent laboratory, which

reported 0.21, 0.80, 2.15, and 3.68 wt.% Mn. We will use the manufacturer's

chemical compositions. The alloys were melted in a vacuum induction furnace

in an Ar atmosphere. The ingots were machined to a diameter of 6 cm,

preheated to 816 °C, hot-extruded to 1.9 cm diameter, cold drawn to 3.2 mm

diameter, and annealed in Ar for 1 h at 650°C. Samples were cut into

cylinders, about 1.5 cm in length, 3.2 mm in diameter.

Susceptibility vs. Temperature at Low Field

The ac susceptometer was a two -position , mutual - inductance type [25].

Susceptibility was measured along a sample's axis as a function of increasing

temperature in an 80 A/m (1 Oe) rms , 10 Hz field. Eddy-current contributions

to the susceptibility signal were negligible for the thin samples at this low

frequency. No demagnetization factor corrections were necessary. A typical

susceptibility curve for one of the samples is shown in Fig. 1. Peaks in the

susceptibility curves gave the spin- freezing temperatures, Tf, and maximum

volume susceptibility values at Tf, in Table I.

For comparison, we refer to the careful susceptibility measurements on Cu-Mn single crystals by Nagata et al. [26] They measured Tf values of 9.90

and 14.67 K, and effective magnetic moments of Mn of 4.80 and 4.86 /ig, for

1.08 and 2.02 at . % Mn, respectively. (Similar effective moments are given in

Refs. [27] and [28] for other concentrations of Mn.

)

The susceptibilities in Table I are not insignificant. They exceed, for

example, the susceptibilities of many austenitic stainless steels [29-31].

Even at these levels, however, the susceptibility of the Cu-Mn matrix should

not adversely affect superconductor performance. However, since Tf is above

the boiling temperature of helium for all except the 0.16 wt.% Mn alloy, the

low- temperature magnetization-vs .- field behavior of these samples is examined.

Magnetization vs. Field at 4- K

Magnetization measurements of the 0.72, 2.00 and 3.75 wt.% Mn alloys wer made as a function of field at 4 K with a vibrating- sample magnetometer. We

report measurements of magnetization at 5.6 MA/m (70 kOe , referred to below

"7 T" , the magnetic flux density in free space), hysteresis loss for cycled

fields, and remanent magnetizations at zero field. The results are summarizin Table II.

At the maximum applied field of 7 T, the 3.75 wt.% Mn sample had a magnetization of 44 kA/m, greater than that of many austenitic steels. Th‘

suggests that there could be some field distortion due to the matrix mat-

•' if it were used in the manufacture of superconducting magnets. The

magnetizations of the 2.00 and 0.72 wt.% Mn samples at 7 T were 26 kA r

- ' ’ kA/m, respectively, with much less potential for problems. Characte: i sr •: spin glasses, the magnetization curves were far from saturation at

57
The hysteresis losses per 7-T field cycle are small. Compared to typical

hysteresis in multifilamentary conductors, about 100 kJ/m^ (10® erg/cm®), the

losses are negligible. For the 3.75 wt . % Mn sample, the loss is approximately

10 kJ/m® . For the 2.00 and 0.72 wt . % Mn alloys, the losses are about 4 and

0.4 kJ/m®, respectively. In applications where the field direction is not

reversed, there will be virtually no hysteresis.

Reducing the field to zero with the samples at 4 K allows the IRM to be

measured. For the 3.75 wt . % Mn sample, the IRM gets no larger than 1.8 kA/m

once the maximum field per cycle exceeds 2.4 MA/m. For the 2.00 and 0.72 wt .

Mn samples, the values of IRM get no larger than 600 and 80 A/m, respectively,

for smaller maximum fields. Similar remanent magnetizations at 4 K have been

reported by others [27,32,33]. Even for the 3.75 wt .Mn alloy, the IRM is

small

Cooling the samples from above Tf to 4 K in the field, and then reducing

the field to zero, gives the TRM. For a 7-T cooling field, the TRM values

approximately equaled the IRM values noted above.

Conclusion

Spin-glass properties of Cu-Mn alloys have the potential to impair

superconductor performance when they are used as the matrix material for finefilament Nb-Ti wires. For the case of dilute alloys, with less than 4% Mn,

the effects are minimal.

The main concern is possible distortion of the superconducting-magnet

field profiles when matrix material containing 3.75 wt .Mn (or more) is used.

A lesser concern is that there will be a small remanent magnetization owing to

the matrix material when the wire is cycled to high field at low temperatures.

Hysteresis losses are insignificant.

Acknowledgments

R. L. Spomer assisted with some of the measurements. E. W. Collings made

several helpful comments and suggestions on the manuscript. This work was supported, at NBS , by the Department of Energy, Office of Fusion Energy and

Division of High Energy Physics, and at Supercon, by the Department of Energy,

Small Business Innovation Research Program.

58
References

[1] E. W. Ceilings, Adv. Cryo . Engr . (Materials) 34, 867 (1988).

[2] T. S. Kreilick, E. Gregory, J. Wong, R. M. Scanlan, A. K. Ghosh, W. B. Sampson, and E. W. Ceilings, Adv. Cryo. Engr. (Materials) 34, 895

(1988) . [3] E. Gregory, T. S. Kreilick, J. Wong, A. K. Ghosh, and W. B. Sampson,

Cryogenics 27, 178 (1987).

[4] T. S. Kreilick and E. Gregory, Cryogenics 27, 401 (1987).

[5] J. W. Ekin, Cryogenics 27, 603 (1987).

[6] J. Niemeyer and G. von Minnigerode, Z. Physik B 36, 57 (1979).

[7] D. K. Finnemore, H. C. Yang, V. G. Kogan, and S. L. Miller, Adv. Cryo.

Engr. (Materials) 30, 909 (1984).

[8] T. S. Kreilick, E. Gregory, J. Wong, J. Less Common Metals (1988), in

press

[9] G. Heber, Appl. Phys . 10, 101 (1976).

[10] J. A. Mydosh, Am. Inst. Phys. Conf. Proc . 24, 131 (1975).

[11] J. A. Mydosh, J. Magn. Magn. Mater. 7, 237 (1978).

[12] P. A. Beck, Prog. Mater. Sci. 23, 1 (1978).

[13] K. H. Fischer, Phys. Stat. Sol. B 116, 357 (1983).

[14] K. H. Fischer, Phys. Stat. Sol. B 130, 13 (1985).

[15] C. Y. Huang, J. Magn. Magn. Mater. 51, 1 (1985).

[16] K. Moorjani and J. M. D. Goey, Magnetic Glasses, Amsterdam: Elsevier,

1984, Ch. 7.

[17] D. Chowdhury, Spin Glasses and Other Frustrated Systems, New Jersey:

Princeton University, 1986.

[18] R. B. Goldfarb, K. V. Rao , and H. S. Chen, Solid State Comm. 54, 799

(1985)

[19] A. F. J. Morgownik and J. A. Mydosh, Phys. Rev. B 24, 5277 (1981).

[20] C. N. Guy, J. Appl. Phys. 50, 7308 (1979).

[21] R. W. Knitter, J. S. Kouvel, and H. Claus, J. Magn. Magn. Mater. 5, 356

(1977).

[22] C. N. Guy, J. Phys. F: Met. Phys. 12, 1453 (1982).

[23] P. Monod, J. J. Prejean, and B. Tissier, J. Appl. Phys. 50, 7324 (1979).

[24] P. A. Beck, Phys. Rev. B 33, 3585 (1986).

[25] R. B. Goldfarb and J. V. Minervini, Rev. Sci. Instrum. 55, 761 (1984).

[26] S. Nagata, P. H. Keesom, and H. R. Harrison, Phys. Rev. B 19, 1633

(1979) . [27] J. Owen, M. E. Browne, V. Arp, and A. F. Kip, J. Phys. Chem. Solids 2,

85 (1957).

[28] J. M. Franz and D. J. Sellmyer, Phys. Rev. B 8, 2083 (1973).

[29] E. W. Collings and S. C. Hart, Cryogenics 19, 521 (1979).

[30] E. W. Collings and R. L. Cappelletti, Cryogenics 25, 713 (1985).

[31] K. Pieterman, A. Ketting, and J. Ch. Geerse, J. Physique (Colloq.) 45,

Cl-625 (1984).

[32] J. A. Careaga and R. Tournier, cited in J. Souletie and R. Tournier, U

Low Temp. Phys. 1, 95 (1969).

[33] The conversion of magnetization from /ig per Mn atom to A/m for dilute

Gu-Mn is: [/ig/Mn] x [at.% Mn] x 7.854 x 10^ = [A/m]. For example,

0.0123 ^g/Mn for 0.83 at.% Mn corresponds to a magnetization of 80 A

59
Figure 1

peak

Volume ac susceptibility vs. temperature for Cu-3.75 wt . % Mn.

at 25.2 K corresponds to the spin- freezing temperature, Tf.

The

Table I. Spin- freezing temperatures (Tf) and maximum volume susceptibilities

(Xmax) dilute Cu-Mn alloys.

Wt.% Mn At.% Mn Tf (K) Xmax ^max (^8^)

0.16 0.19 < 4.0 2.1 X 10-3 1.7 X 10 -“ 0.72 0.83 9.2 4.5 X 10-3 3.6 X 10 -“ 2.00 2.31 18.0 1.0 X 10-2 8.0 X 10-“ 3.75 4.31 25.2 1.8 X 10-2 1 CO X O

Table II. High-field magnetization (M) , hysteresis loss, and isothermal

remanent magnetization (IRM) , for dilute Cu-Mn alloys at 4 K.

Wt.% Mn M at 7 T Hysteresis IRM

(kA/m) (G) (kJ/m^) (erg/cm^ ) (kA/m) (G)

0 ..72 12 150 0 ,.4 4 X 103 0 ,.08 1 2 ..00 26 325 4 ..0 4 X 10 “ 0 ,.6 8

3 ,.75 44 550 10 ,,0 1 X 10 ® 1 ,.8 23

60
VII. COUPLING LOSSES MEASURED BY VIBRATING- SAMPLE MAGNETOMETRY

R. B. Goldfarb

Abstract

Coupling losses in multifilamentary superconductors may be measured with

a vibrating- sample magnetometer by varying the waiting period before recording

the magnetization signal after a field change. A Nb-Ti multifilamentary wire

and a multistrand cable with different twist pitches were measured for

different waiting periods. For wire with long twist lengths, coupling losses

may be up to two times the hysteresis loss.

Introduction

Coupling in multifilamentary superconductors refers to the tunneling of

Cooper pairs from one superconducting filament to another through a normal

metal matrix, a phenomenon known as the proximity effect. These electron

currents are Induced by changes in applied field in an attempt to shield the

composite conductor. They give rise to magnetic moments which may be detected

by magnetometers based on vibrating a sample or on integrating flux [1].

Magnetic moments from coupling currents add to those from superconductor

shielding and trapped flux to increase the area of a magnetization-vs .- field

hysteresis loop. Unlike superconductor shielding currents, coupling currents

decay with time. Thus the measured area of a hysteresis loop depends on the

sweep rate of the field.

Measurements made by the integrated flux method [2] require a swept

field. With increasing cycle time, the hysteresis loss remains constant, but

the coupling loss decreases. The hysteresis loss may be isolated by

extrapolation.

A vibrating- sample magnetometer may be used with either swept or stepped

fields. The former method requires perfectly balanced pick-up coils to

eliminate induced voltages from the changing applied field. Since balanced

coils are difficult to realize, I use the latter quasistatic method.

Sufficient waiting time is allowed between field steps for the superconducting

field magnet to come to equilibrium and for coupling currents in the

superconductor sample to decay. Thirty to 40 seconds are usually adequate fo'' the lock-in amplifier to stabilize.

Typical methods for reducing coupling losses are to twist the multifilar

conductors [3], increase the resistivity of the matrix, or add magnetic

scatterers to the matrix [4]. In this note, I report on some measurements o' magnetization as a function of field for a multifilamentary Nb-Ti conductn:- • several geometries to illustrate how coupling losses can be measured with ; vibrating- sample magnetometer.

61
Experiment

Samples were in the form of wires or fiat woven cable. Some of the wire

characteristics are given in Table I [5]. A thin Nb diffusion barrier was on

the surface of each Nb-Ti filament. Wire samples were coiled in two layers

around the threads of a size 10 nylon screw sample holder, about 4.5 mm in

diameter. The magnetic field was applied approximately transverse to the

wire, along the axis of the coil.

The cable was made with 30 strands of wire. A cable sample consisted of

two carefully cut 4-mm segments with the field applied in the cable plane.

Because of the transposition of the strands in the cable, the field was only

approximately transverse to the individual strands. One wire sample was carefully cut into 4-mm segments, for comparison with the cable sample.

The wire stock had a twist length of 2.5 cm. In the cabling operation

the twist length of each strand was increased somewhat. However, because the

cable sample consisted of 4-mm segments, there was an "effective" twist length

of 0.4 cm. For purposes of comparison, I further twisted some of the wire to

a final twist length of 0.45 cm.

Measurements of magnetization as a function of field were made at 4 K,

with maximum applied fields of ±800 kA/m (±10 kOe , corresponding to a flux

density in free space of ±1 T) . Each complete hysteresis loop was traversed

in 60 equal steps per cycle. The waiting time between reaching a measurement

field and recording the magnetization was either 40 or 90 seconds. The area enclosed by the loops equaled the total losses under these conditions.

Table II gives the total loss per sample volume (Nb-Ti plus matrix) for

wires and cable, for different twist lengths and waiting times. It may be

concluded from these data that the intrinsic hysteresis loss per cycle for

this wire is 31 kJ/m®. Additional losses, due to coupling, depend on time and

twist-pitch. These coupling losses could amount to twice the hysteresis loss.

Conclusion

Interfilamentary coupling losses may contribute to the losses measured by

vibrating- sample magnetometry . They can be eliminated by sufficiently long

waiting times between field steps or by cutting the sample in short segments.

Short twist lengths are effective in reducing coupling losses in the actual

wires

Acknowledgments

Wire and cable stock, and wire characteristics based on delivered

material, were provided by R. M. Scanlan, Lawrence Berkeley Laboratory. The

work was supported by the Department of Energy, Office of Fusion Energy and

Division of High Energy Physics.

62
References

1] A comparison of the advantages of each type of magnetometry is given in:

R. B. Goldfarb and A. F. Clark, Adv. Cryo . Engr. (Materials) 32, 779

(1986)

[2] W. A. Fietz, Rev. Sci. Instrum

36, 1621 (1965).

[3] c. H. Morgan, J. Appl. Phys . 41

3673 (1970).

[4] E. W. Collings, Adv. Cryo. Engr. (Materials) 34, 867 (1988).

[5] This wire is described

design and fabrication

in: T. S of strand

. Kreilick, E. Gregory, and J.

for SSC magnet applications,"

Klipping and I. Klipping, eds . , Butterworth, UK (1986) pp . 725-729. The

billet number is Supercon 351E, alternately designated "LBL2". Cable

identification is SC322.

Table I. Characteristics of Nb-Ti wire.

Number of filaments 4326

Filament diameter 6 . 3 /im

Edge-to-edge filament spacing 0.8 /im

Wire diameter 0.65 mm

Twist length 2.5 cm

Table II. Total loss for different twist lengths and waiting times.

Run

number

Sample

geometry

Twist

length (cm)

Waiting

time (s)

Loss per

cycle (kJ/m^)

1 Wire coil 2.5 40 91

2 Wire coil 2.5 90 58

3 Wire coil 0.45 40 31

4 Wire segments 0.4 * 40 31

5 Cable segment 0.4 * 40 32

6 Cable segment 0.4 * 90 32

* "Effective" twist length equal to segment length.

63
Appendix A ^VANCES IN CRYOGENIC ENGINEERING MATERIALS, VoL 34

Edited by A.F. Clark and R.P, Reed

(Plenum Publishing Corporation, 1988)

CURRENT -RIPPLE EFFECT ON SUPERCONDUCTIVE DC CRITICAL CURRENT MEASUREMENTS*

L. F. Goodrich, S. L. Bray, and A. F. Clark

Electromagnetic Technology Division

National Bureau of Standards

Boulder, Colorado 80303

ABSTRACT

The effect of sample - current power-supply ripple on the measurement of

dc critical current is reported. Measurements were made on multifilamentary

NbTi superconductor. In general, ripple in a current supply becomes more significant above 500 A because effective filtering is difficult. The

presence of current ripple reduces the measured dc critical current. Ripple

also causes noise at the input to the voltmeter used for the measurements.

The quantitative effect of current ripple was studied using a battery

current supply modified to allow the creation of ripple current with

variable frequency and amplitude. Problems common to all large -conductor

critical-current measurements are discussed.

INTRODUCTION

The presence of current ripple (any periodic departure from a dc output

level) reduces the measured dc critical current (Ic) • addition to quantifying this effect, an attempt was made to determine its source. Several possibilities are discussed.

The Ij, measurements reported here were made using a straight sample

geometry and a radial-access magnet. The 1^ was defined as the dc current

at which the measured electric - field strength rose to 0.2 /iV/cm. A battery

current supply was modified to allow the creation of current ripple with

variable frequency and amplitude. Two types of voltmeters were 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 input to a digital processing oscilloscope.

The voltage signal of a current-sensing resistor was split, amplified, and

fed into two channels of a digital processing oscilloscope. One of these

current signals was filtered to give the dc 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

was measured at relatively fixed current ripple and numerous dc bias

currents. These points along the voltage-current (V-I) curve were analyzed

to determine I^,. The precision of the I(, measurement was about ±1%.

* Contribution of NBS , not subject to copyright.

1019

64
The performance of the battery power supply as a source of ac 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 ideal value for

the given waveform (1.73 for a 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 ideal and measured ratios increased

with increasing ac amplitude. This variation was essentially independent of

frequency and had a maximum value of less than 12% at the maximum tested ac amplitude of 70 A. At amplitudes below 40 A, the maximum variation was less

than 5%.

The sample voltage has several components. For this work, the

intrinsic dc 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.^ There might be a slight currenttransfer voltage but otherwise the real component of the voltage should be zero, within measurement limits, until the current approaches the 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 simple electronic circuit was designed to simulate the intrinsic V-I

characteristic of a superconductor. The circuit was used to determine the

nanovoltmeter's response to a highly asymmetric voltage waveform. It was also useful for testing and developing data acquisition systems. This

circuit was developed based on an idea by J . W. Ekin and J. C. Branch of

this institution. The original idea used a diode in series with a current

sensing resistor and this combination in parallel with a high-current shunt

resistor. This passive circuit gives a nonlinear V-I characteristic but the curve is not very sharp (n about 10) and it is not adjustable. For the

adjustable active circuit, a voltage signal proportional to the sample

current is the input and a highly nonlinear current is the output of the

circuit. The circuit consists of a variable gain (variable n) operational

amplifier with the current-proportional signal as the input and a variable

voltage bias point (variable I(,) . The amplifier output drives the base of a transistor, which has a diode and current sensing resistors in series with

the emitter current. These current sensing resistors provided various

voltage level waveforms. A 0.01 II series resistor is used to produce a low

voltage (-1 /iV) waveform for the nanovoltmeter to monitor. A 100 n series

resistor provides a high voltage signal (-10 mV) that allows the indirect

determination of the input signal to the nanovoltmeter using a wide band

amplifier to measure the total voltage waveform.

The sample measured in this experiment was a commercial

multifilamentary NbTi superconductor with a diameter of 0.65 mm. The

filament twist pitch was 0.8 twist/cm. There were 6006, 5-^m-diameter

filaments with a Nb diffusion barrier around each of the filaments. The

sample was contacted with several sets of twisted pairs of voltage taps.

For this experiment, this allowed simultaneous acquisition of data with

several voltmeters to determine 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 larger. It spanned

the same segment of the sample as another lower-mutual- inductance pair. Two

other voltage-tap pairs had their wires connected together near the sample

and only one of the wires was connected to the sample (referred to as null

tap pairs). This provided pairs that should not have any real (resistive.

1020

65
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.

RESULTS

The for increasing dc bias current was measured as a function of ac amplitude and frequency using two different types of voltmeters and various

pairs of voltage taps. The reduction in the measured dc versus ac amplitude is shown in Fig. 1. The ac amplitude is the peak value of the ac component of the current waveform (not peak-to-peak or peak of total

current). These data were taken using an analog nanovoltmeter. The range

of ac amplitude covered 0 to 50% of the dc bias-current (I) which was about

140 A at 8 T. A reference line was drawn for an Ij, reduction equal to the

ac amplitude. For all frequencies and waveforms, these data are just below

this reference line indicating that the effect is to reduce the measured Ij. by slightly less than the ac amplitude. The -60 Hz frequency was labeled

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 operating

mode. All of the non-zero frequency data fall on a straight line; all of

the curvature takes place between zero ripple and the first point. The

triangular waveform was added to determine if the detailed shape of the

ripple and the relationship of rms to peak, changed the measured values.

These data suggest that the detailed shape is not significant. The ac amplitude, however, does have a significant effect on the measured dc I^.

Data taken using a digital nanovoltmeter is shown in Fig. 2. The 60 Hz data of Fig. 2 were measured on the identical pair of voltage taps as the 60 Hz data of Fig. 1. The data at other frequencies were measured simultaneously with the corresponding data on Fig. 1, 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 dc

does not depend on the voltmeter used. There were, however, other differ- ences between these two voltmeters that will be discussed later. The effect

of current ripple was repeated on a number of sample segments and each

segment gave similiar results.

In order to develop a quantitative understanding of the nanovoltmeter

input-output relationship for ac voltage inputs, and thus determine the

effect on the measured I^,, two ac 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. AC

signals of appropriate amplitudes were obtained by dividing the output

voltage of an ac signal generator. The objective of the second test

(asymmetric test) was to determine the response of the nanovoltmeters to the

highly asymmetric ac 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 dc output of the meters

increased with increasing ac 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 ac input. Defining the ac rejection

(ACR) of the nanovoltmeter as one minus the ratio of the dc output voltage

1021

/ 66
Fig. 1. reduction versus ac amplitude at several

frequencies using an analog nanovoltmeter.

Fig. 2. reduction versus ac amplitude at several

frequencies using a digital nanovoltmeter.

1022 67
to the peak ac input voltage provides a useful measure of a meter's relative

tolerance for ac input voltages. This is not the common mode rejection

ratio (CMRR).^ For the analog meter, a 60 Hz, 500 /iV peak ac signal

resulted in a dc output of 5.2 /iV, or an ACR of 99.0%. The same input

signal applied to the digital meter resulted in a dc output of 0.66 fj.V

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 peak for 60 Hz and 70 A amplitude; 880 /iV peak for 1000 Hz and 70 A amplitude). The mutual

inductance of the high inductance taps was about 6 nH.

The dependence of the dc 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 dc output of both meters showed a non-monotonic dependence on the input

signal's frequency, the analog meter showed a considerably stronger

frequency dependence. For the analog meter, a 250 fiV peak ac signal at 60,

360, and 1000 Hz resulted in a dc output of 0.8, 3.5, and 0.5 /xV respectively. The same input signals applied to the digital meter resulted

in a dc output of 0.14, 0.16, and 0.10 /jV respectively.

In the case of the asymmetric test, the behavior of the two meters was essentially the same. A broad range of dc 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 ac frequencies (60, 360, and 1000 Hz) were tested. In order to determine the

nanovoltmeters ' response to these input signals, the signals were observed

and analyzed on a digital processing oscilloscope. The peak to peak, time

average, 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

(frequency, dc bias level, and ac amplitude), the output signal was always

within 5% of the time-average (dc) value of the input signal. In other

words, the nanovoltraeter behaves like a very low-frequency, low-pass filter,

passing the dc component and rejecting the ac component of the input signal.

This behavior is consistent with the symmetric test where the meters were shown to ignore sinusoidal input signals.

MODEL

The motivation for modeling 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-I curve and the shape of ripple will change the magnitude of this effect. To extend the dc I^, criteria to a general current waveform in a way consistent

with the design limits of a superconductor application, the measured voltage

criteria should reflect the ohmic power loss within the superconductor.

Differentiation between voltage and electric field strength criteria will be

dropped in this section.

The magnitude of the effect of ac ripple on the measurement of dc I^, can be calculated for the case of a normal, ohmic sample which would have an

n of 1 . For the purpose of this argument, start with the definition of I^, as the rms value of the total current (ac and dc) 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 ac power. For

sinusoidal waveforms, the time average of the instantaneous power is the

68 1023
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 (ac and dc) resistive voltage. Based on this

definition of Ij, and an n of 1, the measured dc I(, would be equal to the

square root of, without ripple squared minus the rms of the ripple

squared. For example, the power of 99.9 A dc plus 5 A rms ac is equal to the power of 100 A dc (99.9 - ((100)^- (5)^)^). The two components of

current add in quadrature, thus for a low percentage ripple the effect on

the dc value is very small, for n of 1

A dc 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 dc bias current is well below the flux-flow

region and the ac amplitude is just large enought 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 ac 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 dc voltmeter is the proper

instrument to use based on the definition of that relates to the average

power. It can be shown that for an n of 2 and a sinusoidal waveform, the

time average of the voltage gives the same dc as the time average power

for the case of n equals 1. In general, the difference between the time

average voltage and the time average power is a difference of 1 in the

exponent. For relatively low amplitudes, the effect can be approximated

using the first few terms of the binomial series for the instantaneous

voltage. For an n of 30, this gives 98.3 A dc plus 5 A rms ac is equal to 100 A dc (i.e. the measured dc 1^ would be reduced by 1.7 A). For an n of

infinity, it is apparent that the dc voltmeter will give a reduction in the

measured I^, that is equal to the ac 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 I^ and this in combination with n can be used to estimate the effect.

DISCUSSION

The analyzed data from both the digital and the analog voltmeters gave

the same I^ . However, some I^ data taken with the analog nanovoltmeter

indicated the presence of a current- transfer voltage.^ This was not true of

the data taken with the digital meter. It is believed that this behavior is

simply an artifact of the measurement system and not the result of a true

current- transfer voltage. There are two characteristics of the measurement

system that may explain this effect. One concerns the sample current supply

and the other concerns the nanovoltmeter. First, at high ac amplitudes the

sample current supply's ac amplitude increases with increasing dc 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 dc voltage output when subjected to an ac input, and this dc

voltage increases with increasing ac amplitude. Consequently, during a cycle of data acquisition, the nanovoltmeter has an increasing dc voltage

output which is interpreted by the data analysis program as current- transfer

voltage. The digital meter does not indicate this artificial currenttransfer voltage because of its considerably greater rejection of symmetric

input signals. Finally, 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 I^, 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.

1024

69
Although both analog and digital nanovoltmeters have demonstrated

predictable input-output behavior with respect to low amplitude input

signals, there is an apparent ac amplitude threshold above which their

behavior becomes erratic. This erratic behavior 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

behavior but the ac 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 ac input signals.

Data taken on the null tap pairs (described in the introduction) 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 behavior, all of

which were presented above. This is further evidence for the assertion that

these effects are caused by the ac inductive voltage. A passive RC filter

was inserted at the input of the nanovoltmeter to reduce the ac amplitude of

the voltage signal. This did reduce the artificial current transfer

voltage, and it increased the threshold of the erratic behavior. 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 area can be adjusted with a high frequency

signal during sample preparation.

The self-field loss is the primary type of transient loss present in

the case of sample current ripple for this sample geometry.^ The only

effect of transient loss on the 1^ 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 ac current amplitude was 3 mW at 60 Hz (50 mW at 1000 Hz). The

effect of current ripple was also measured at 5 T with ac amplitudes up to about 50% of the 340-A dc I^,. 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 by more) , but the results were still at or below the

reference line, reduction less than or equal to the ac amplitude. The lack

of a strong frequency, amplitude, and current density dependence of the

measured reduction of I,, 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 determine this.

The sample was centered 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 and reduced the net torque on the sample cryostat. I^. data was taken at 5 T with four returns and with one return for various

frequencies and ac amplitudes. The results with a single return had a slight frequency and amplitude dependence. At 60 Hz (sinusoidal and

triangular), these data were within 2 A of the four-return results. At 360

Hz and 100 A ac amplitude, was reduced by another 8 A. 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

Additional data were taken at higher frequencies (3 and 10 kHz), but

experimental difficulties made the results at these frequencies less

70

1025
reliable. Transient losses, inductive voltage, va%*efom distortion, and

reduction in available amplitude all becane note significant at higner

frequencies . The notivation for investigating higner rrequencies «'es to estimate the effect of frequencies near the range of spikes fron

silicon-controlled rectifiers (SCR) found in nost large pover supplies.

These data indicated the sane general trend of reduction in 1^ but vith tore

experinental uncertainty. Perhaps a better nodel of an SCR spike would be a spike rather than a continuous dut%^-cycle wevefom. An extrapolation of the

work presented here would indicate that the effect of the spikes would only

be to add a finite anount to the tine average of the resistive voltage.

CONCLUSION'S

“nis report concludes t’nat the effect of ripple should scale with t'ne percentage of Ij, and t'ne effect will depend upon the shape of the V-I cur’.'e (n) . For an n of 30 or nore and large ripple, t'ne effect of current ripple

on the neasurenent of dc critical current is to reduce it by slightly less

than t'ne ac anplitude. A dc voltneter neasures a relevant quantity, the

tine average of the voltage. A sinple electronic circuit was used to test a dc voltneter’ s response to an as\*nnetric wavefom t'nat sinulates a superconductor's V-I characteristic. Reducing t'ne nutual inductance of t'ne voltage taps is inportant w'nen testing wit'n power supply ripple present.

A large -conductor test nethod has to address t'ne reality of power

supply current ripple. Power line frequencies and 'namonics (up to 3-p'nase.

fully rectified) are t'ne doninant frequencies. Current ripple of 1 to 3% at

t'ne naxinun current is t^.’pical for power supplies in t'ne range of 600 to 6000 .A. Tnis ripple is generally a f-unction of output current 'out t'ne percentage ripple can 'oe 'nig'ner for intemediate currents. This can be a significant source of neasurenent error. The direct effect on t'ne sanple

voltneter can also introduce neasurenent error. Sanple heating due to transient losses nay also be significant for sone sanple geonetries.

Tne aut'nors extend t'neir t'nanics to E. Loo'x and K. E. Spilnan for

t'neir assistance wit'n sanple preparation and conputer prograning; and to

R. B. Goldfar'D for discussions on transient losses. Research supported by

the 'l'.S. Departnent of Energy." t'nroug'n t'ne Office of Fusion Energy and t'ne D ivis i on of Hi g'n Ene r

g:>- P'nys i c s .

An effort was nade to avoid the identification of connercial products

'ey t'ne nanufacturer ' s nane or label, 'out in sone cases t'nese products night

'oe indirectly identified 'oy their particular properties. In no instance

does t'nis identification inply endorsenent by t'ne National Bureau of

Standards, nor does it inply that t'ne particular products are necessarily

wns DeSw n o zt cnsc p'^zrposs.

1. L. F. Goodric'n and F. R. Fic’Kett, "Critical Current .heasurenents : A

Conpendiun of Experinental Results," Cryogenics 22., 255-241 (1962).

2. L. F. Goodrich, S. L. Bray, ?. Dune. E. S. Pittnan, and A. F. Clar'x, "Developnent of Standards for Superconductors. Interin Report

Jan. to Dec. 1985," N3SIR 87-3066, National Bureau of Standards,

Boulder, Colorado (.April 1987).

3. J. «'. E'-cin, "Superconductors," C'napt. 13 in .Haterials at Lew

Tenterature . edited by R. ?. Reed and A. F. Clark, Anerican Soc for

.Metals, p. 489-494 (1983).

1026

71
Appendix B

Hysteresis losses in fine filament

internal-tin superconductors

R.B. Goldfarb and J.W. Ekin

Electromagnetic Technology Division, National Bureau of Standards,

Boulder, CO 80303, USA

Received 20 June 1986

Hysteresis losses were measured on a series of fine filament NbsSn superconductors made by the

internal-tin process. Hysteresis was measured as a function of filament diameter and interfilament separation using a vibrating sample magnetometer in transverse magnetic field. Losses were

greater than expected from the critical state model that expresses loss as a function of filament diameter. Micrographs of the reacted wire cross-sections showed some interfilament bridging for

all wires. This gave rise to effective filament diameters that were greater than actual diameters. The critical interfilament separation, above which the losses would be expected to follow the

critical state model, was determined.

Keywords: Nb3Sn; superconductors: fine filament internal-tin superconductors;

a.c. losses; hysteresis: magnetization

Fine i'llanicnt superconductors are of interest in a.c. loss studies because, according to the critical state model,

hysteresis loss is proportional to filament diameter. Thus, one would e.xpect that wires composed of fine lllaments would exhibit small losses. In this study we used a \ibrating sample magnetometer in transverse magnetic

field to measure the d.c. hysteresis loss for a series of NbjSn superconductors made b\ the internal-tin process. There was no transport current. Filament

diameter and edge to edge interfilament separation were

systematically varied to determine whether filament coupling caused larger than expected hysteresis losses. Such magnified losses may be expressed in terms of an

effectise lilament diameter that is greater than the actual filament diameter. The effective diameter is deduced from

the hysteresis loss using the critical state model. There has been some precious work on hysteresis

losses in fine filament superconductors. Dubots ei a!.' compared measured and calculated hysteresis in Nb-Ti

conductors with filament diameters ranging from 10 to

t). I /xm. They found that, for diameters < I jam. effective filament diameters remained at ^1 jam. Information on

interlllament separation was not given. Carr and V\’agneH

attributed the asymmetrical hysteresis curve they observed in Nb-Ti with l.b jam filament diameter to

relatively large surface currents. Ghosh and Sampson’

studied fine filament Nb-Ti with filament diameters in the range 1-5 jam. They note that, for diameters < 3 jam. the magnetization land therefore hysteresis loss) is larger than expected from the critical state model. They allribute

this to degradation of the magnitudes of critical current owing to filament necking or damage.

In this Paper we show that certain NbjSn internaltin conductors with filament diameters < 5 have

effective filament diameters greater than expected from

the critical slate model. The ratio of effective to actual NbjSn filament diameters appears to be a linear function 001 1-2275/86/080478-04 $03.00

® 1986 Butterworth & Co (Publishers) Ltd 478 Cryogenics 1 986 Vol 26 August/September

of interfilament separation. The slope of the line is a function of the local area ralioof matrix material to Nb for each w ire. The greater effective filament diameters appear

to be a consequence of bridging between filaments.

Sample preparation and measurement

The multifilamentary NbjSn wires used in this study were composed of subelements containing 150 filaments each. The array of subelements was surrounded by a tantalum

diffusion barrier. Unreacted NbjSn wires were wrapped

on the threads of size 10-24 stainless steel screws. The

thread pitch was such that the windings were not in contact with each other. The screws had been previously oxidized to prevent the wires from slicking to the screws. The internal-tin process wires were reacted in a vacuum,

with the cold ends out of the furnace, according to the following schedule: one day at 340'^C. four days at 5X0 C

and four days at 700°C. This schedule is adequate to fully react the wires for the filament diameters in this study'*. The residual tin content of each wire after reaction was

estimated by the manufacturer to be cS wt''u. After reaction, the wire coils were unscrewed from

the stainless steel screws and the long ends cut off. Each

coil was = 1.3 cm in height. The wire volumes were measured using Archimedes' principle. The w ire diameter was the same for each sample. O.bS I mm. Thus, the wire

length for each coil was obtained. The average Nb filament diameters before reaction and numberof filaments in each sample were prov ided by the manufacturer. The reacted filament diameters and

total v olume of NbjSn in each coil were calculated based on a presumed increase of 3S"o in filament cross-sectional area after reaction'. Local area ratios of matrix to Nb were determined by the manufacturer from the arrangement of

the Nb rods in the billets before drawing: the Nb may be thought of us enclosed by adjacent hexagons and the

72
Hysteresis losses in superconductors: R.B. Goldfarb and J.W. Ekin

Table 1 Parameters lor six Nb3 Sn muittfilamentarv superconductors

Sample Nb

diameter (^m)

Nb3Sn

diameter

(/an)

Local area

ratio

Nb

separation

(/an)

Nb3Sn

separation

(/an)

Numoer of filaments Length

icm) Volume NCjSn

imm^i

3 4 2 4.9 1.5 2.1 1.4 2850 17 8 9.71 6 40 4.7 1.8 2.4 1.7 2850 17 3 8 55

9 3.0 3.5 1 5 1.5 1.0 5550 17 1 9 24

12 2 8 3 3 1.8 1.7 1.2 5550 17.0 8 01

15 2.3 2.7 1.5 1.2 0.8 9150 17.0 8 89

18 2.2 2.6 1.8 1.3 0.9 9150 16 8 8 04

uToss-scaional areas of Nh and matrix material measured. The average centre to centre filament distance,

i. ma> be determined from the local area ratio./., and Nb

filament diameter, d. b\ the equation

Z. = Cx 3 -) - 1 (1)

The edge to edge Nb separation is simply s — d. Table I gives relevant parameters for the six samples. Note that the Nb diameter plus Nb separation equals the NbjSn diameter plus NbjSn separation. Both sums are equivalent to the centre to centre filament distance (not shown). The matrix to Nb local area ratios of 1.5 and 1.8 correspond to area ratios for reacted NbjSn of 1.0 and 1.2. respectively. The magnetic moment was measured at 4.2 ± 0.2 K

with a vibrating sample magnetometer in the axis of the

coil (i.e. field transverse to the wire if we ignore thread

pitch of =5.5"). The maximum applied field was

1.6 M.^ m“' (20 kOe). There was no transport current and

the coils were open circuited. .Magnetization was computed as magnetic moment per unit volume of Nb3 Sn (i.e. not including the volume of the matrix material). Magnetization was plotted Venus e.xternal (applied) magnetic

field. A typical magnetic hysteresis loop for one of the samples (sample 18) is shown in Figure 1

Results

Hysteresis loss was computed by numerically integrating

the loop area, ignoring the initial branch. .According to the

critical state model, for transverse field geometry*;

External Field (kOe)

-20 -10 0 10 20

External Field (MA/m)

01

5; S

Figure 1 Typical hysteresis loop for a multifilamentary NbjSn wire made by the internal-tin process (sample 18). Magnetization was

computed as magnetic moment per unit volume of Nb3 Sn

Figure 2 ''ypical scanning electron micrograph showing some

interfilament bridging (sample 18 centre sube'ementi

H' = 8 r/eiT ^ //o 3~ tSl units) |2:

Thus, an effective filament diameter. mav be defined

that is a function of the hysteresis loss. H . for a maximum

applied field.//,. Since we did not have data for the cniical current densitycy^. at 1.6 NtA m“' (20 kOe). we deduced

these based on the critical current of each wire measured at 6.4 .M.A m“' (80 kOe) together with the total NbjSn cross-sectional area and the measured value of the

full penetration field. Hp. at 1.6 M.A m”‘ (20 kOc J^. //- and filament diameter, d. are related according to the

critical state model*: Hp = dJjTT (SI units)

Hp may be obtained as one half of the field required ii reverse the magnetization' From the hysteresis f H.

at 1.6 M.A m"' •:2<i kOei is =:'5 k.A m“' (*^4-'- Oe' ‘ ' .i.:he loops. This suggests a factor of 10 increase compared

to the measured value of at 64 ,MA m"' kO.

.Applying this factor of 10 to each conductor g;-.- th. values ofy^ at 1.6 M.A m“' t20 kOe) shown ;r. Taru 2 'Ae then compute the effectiv e filament diameters of the w rrv and. additionally. the ratiooftheeffeciiveioaciujl N* Sn

filament diameters. Optical and scanning electron muri-.w^ p

. re.-

that some of the filaments m each wire w .-'.- >' .- i e . isolated but were connected by bridges' An i

• ‘ this is shown inF;xun-2 Such filaments had ; . their original hexagonal spacing and j:;g“ I in rows in the process of wire drawing*

"Evidence fo' inierfiijmeni b'lOging n tn*;^ ^ a ' pfesented by A.K Ghosh and W B ^mpson*

Cryogenics 1986 Vol 26 August/'Septernbcf 479

73
Hysteresis losses in superconductors: R.B. Goldfarb and J.W. Ekin

Table 2 Hysteresis loss and effective filament diameters for Nb3Sn

Sample Estimated

value of Jc

at 1 .6 MA m"'

(20 kOe) (GA m"^)

W

(MJ m'^)

1

Effective NbjSn

filament diameter ipm)

Effective to actual NbjSi

diameter

ratio

3 49.4 0.870 10 46 2.12

6 50.6 0.714 8 37 1.78 9 50,6 1.369 16,05 4,55

12 54.3 0 885 9.67 2,94

15 53,8 1419 1 5.66 5.80 18 51.9 0.805 9.20 3.56

Discussion

As shown in Figure 3. the relationship between actual and

effective NhjSn diameters is not monotonic. However, some insight is obtained by plotting the ratio of effective to actual NbjSn diameters versus NbjSn interfilament separation (see Figure 4). It appears that effective filament diameter is a function of both actual filament diameter and interfilament separation. Thus, loss per volume of NbjSn superconductor is a function of both parameters. Further refinement is possible by distinguishing

among those conductors w ith the same matrix to Nb local area ratios. These are identified in Figures 3 and 4 for the rwo ratios. We see in Figure 3 that there is less of an

increase in effective filament diameter for the conductors

with the larger local area ratio. However, we note that in Figure 4 the abscissa (NbjSn interfilament separation) is not independent of local area ratio [see Equation ( 1 )|. The

apparently linear dependence of the effective to actual diameter ratio on interfilament separation has no

physical significance. A critical interfilament separation can be defined It corresponds to the separation above w hich the hysteresis

losses would depend on filament diameter as expected from the critical state model W'e determined the critical separation for reacted Nb,Sn. as shown in Figure 4. by

Actual Nb3Sn Filament Diameter (/xm)

Figure 3 Effective filament diameter as a function of actual NbjSn filament diameter for two local area ratios of matrix to Nb.

Matrix to Nb area ratio: O, 1.5; •, 1.8

NbgSn Interfilament Separation {pm)

Figure 4 Ratio of effective to actual NbjSn filament diameter as a function of NbjSn interfilament separation for two local area ratios of matrix to Nb. Matrix to Nb area ratio; O, 1 .5; •, 1 .8

linearly extrapolating the data to a unity value of effective

to actual diameter ratio. The critical separation appears to be a function of the local area ratio. Critical v alues for the separation between reacted NbjSn filaments were = 1 .6 ;um for an area ratio of 1.5 and 2.0 pm for a ratio of

1.8. The corresponding critical values for the separation between unreaaed Nb filaments were 2.4 and 1.5 /am.

respectively.

Conclusions

Certain fine filament superconductors made by the

internal-tin process have effective filament diameters that are greater than actual filament diameters. The dif- ferences between effective and actual diameters increase

as actual diameters decrease, the effect being more

pronounced in those wires with the smaller local area

ratio of matrix material to Nb. W'hen the ratio of effective

to actual diameters is plotted as a function of edge to edge

interfilament separation, a more refined picture emerges. The diameter ratio increases linearly with decreasing

interfilament separation for each local area ratio. The

larger than expected effective filament diameters are

likely to be due to interfilament bridging in these wires. For the practical purpose of conductor design, a

critical interfilament separation can be determined. It corresponds to the spacing above w hich the hy steresis loss would depend on filament diameter as expected from the

critical state model. The critical NbjSn interfilament separation for the conductors used in this study was 1.6 and 2.0 pm for matrix to Nb area ratios of 1.5 and 1.8. respectively. The critical separation is offered simply as a practical design parameter related to the critical state model and for guidance in the manufacture of fine filament internal-tin superconductors such as those described in this Paper.

Acknowledgements

M. Suenaga made the high field critical current measurements and generously shared samples of the

480 Cryogenics 1986 Vol 26 August/September

74
Hysteresis losses in superconductors: R.B. Goldfarb and J.W. Ekin

unreacied wires. .A.K. Ghosh astuieK pointed out the importance of local area ratios. G.M. Ozersansky

pro\ided \aluahle information on the phssical

characteristics of the wires. E.S Pittman made the optical and scanning electron micrographs. The wires were manufactured b> Intermagnetics General Corporation*.

The experimental work was sponsored by the Air Force

Office of Scientific Research. USA. Most of the results of

this study were presented at the Fourth US-Japan

VK'orkshop on High Field Superconducting Materials for Fusion. 1986. This workshop and the microanalysis were

sponsored b> the US Department of Energy. Office of Fusion Energy.

‘Certain commercial materials are identified to adequately specify

the experimental study. In no case does such identification imply recommendation or endorsement by the National Bureau of Standards, nor does it imply that the material is the best available for the purpose

References

1 Dubots. P.. Fevrier. A.. Renard. J.C., Tivergnier. J.P., Coyer. J. and Hoang Gia K> NhTi wire> with uliru-llnc filament- lor 5U-60 Hz use; Inlluence of the filament diameter upon los-c- lEEE Tram Magn ll'fX5) M.AG-2I 177

2 Carr. W.J. Jr and Wagner. G.R. Hysteresis in a fine filament NhTi composite Ad\ Cnog Eng ( 19X4) 30 92.' 3 Ghosh. A.K. and Sampson. W.B. Magnetization and cntical currents of NhTi wires with fine filamenis.4r/i Cnog Eng | |9sh) 32 X09

4 Ghosh. A.K. personal communication (WSh)

5 Ozeryansk). G.M. personal communication |l9sp)

6 Carr. W .J. Jr.JC Loss and Macroscopic Thcon- ofSupcrconducwr^ Gordon and Breach. New York. US.A (1983) P’

Goldfarb. R.B. and Clark. A.F. Hysteretic losses in Nh-Ti

superconductors y.4p/i/ Pfna (I98.>) 57 3X09

8 Ghosh. A.K. and Sampson. W.B. Magnetization siudie- of

multilllamentary Nh.Sn wire-, paper presented at the Fininh

L'S-Japan Workshop on High Field SuperconductiMiy Matenals

for Fusion ( I9XP)

s

Cryogenics 1986 Vol 26 Augusl/Sepiember 481

75
U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBSIR 88-3088

2. Performing Organ. Report No. 3. Publication Date

February 1988

4. TITLE AND SUBTITLE

HBS-1 14A (REV. 2-8

DEVELOPMENT OF STANDARDS FOR SUPERCONDUCTORS

5. AUTHOR(S)

L.F. Goodrich

6. PERFORMING ORGANIZATION (If joint or other than NBS, see in struction s) national bureau of standards

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & 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

I I 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 Bureau of

Standards, 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

1986 through December 1987. It contains the results of critical current

studies on the effect of power-supply current ripple, measurements on cable

strands, an interlaboratory comparison (round robin) on a large NbTi monolithic

conductor, and a Nb^Sn round robin. Several useful current supply circuits

have been developed. The reduction in coupling losses in multifilamentary

NbTi conductors has been addressed by a study of the magnetic properties of

matrix material consisting of dilute alloys of Mn in Cu. In addition, vlbratingsample magnetometry is shown to be adaptable to the measurement of coupling

losses, in addition to hysteresis losses, in multifilamentary conductors.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize oniy proper names; and separate key words by semicoions) ac losses; circuit diagrams; critical current; current ripple; hysteresis;

magnetization; niobium-tin; niobium-titanium; power supply; round robin; standards;

superconductors; susceptibility

13. AVAILABILITY

1_XJ Unl imited

I I For Official Distribution. Do Not Release to NTIS

I I Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402.

Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

88

15. Price

USCOMM-DC 6043-P80 * U.S. GOVERNMENT PRINTING OFFICE: 1988-573-180/85,120 REGION NO. 8