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NBSIR 77-853

CONS-3800-1

HELIUM RESEARCH IN SUPPORT OF

SUPERCONDUCTING POWER TRANSMISSION

INNMLREPORT BECEIVED

CRYOGENJCS^imaSiQIli

INSTITUTE FOR BASIC STANDARDS

NATIONAL BUREAU OF STANDARDS

BOULDER, COLORADO 803Q2,

PREPARED IDE

E U.S. ENERGY RESEARCH AND DEVELOPMENT ADMINISTRATION

OFFICE OF CONSERVATION

DIVISION OF ELECTRIC ENERGY SYSTEMS

UNDER CONTRACT NO. E(49-l)-3800

(July 1975 - September 1976)

ebruary 1977

NBSIR 77-853

CONS-3800-1

HELIUM RESEARCH IN SUPPORT OF

SUPERCONDUCTING POWER TRANSMISSION

ANNUAL REPORT

(July 1975 - September 1976)

M.C. JONES, V.D. ARP, W.R. PARRISH, . D.E. DANEY, P.R. LUDTKE, N.V. FREDERICK

AND B.A. HANDS*

CRYOGENICS DIVISION

INSTITUTE FOR BASIC STANDARDS

NATIONAL BUREAU OF STANDARDS

BOULDER, COLORADO 80302

(*Guest Worker from Department of Engineering Sciences,

THE U.S. ENERGY RESEARCH AND DEVELOPMENT ADMINISTRATION

OFFICE OF CONSERVATION

DIVISION OF ELECTRIC ENERGY SYSTEMS

UNDER CONTRACT NO. E(49-l)-3800

February 1977 .

-

University of Oxford)

PREPARED FOR

U.S. DEPARTMENT OF COMMERCE, Juanita M. Kreps, Secretary

Dr. Betsy Ancker-Johnson , Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Acting Director
FOREWORD

In order to bring superconducting electric power transmission for ulilities

application from its conceptual stage to a practical realization, developments

were needed in several fields apart from the conductor itself. These included

electrical insulations, economical of space, which would function at the low

temperatures of the conductor; cable designs to withstand the thermal stress

accompanying cooldown; refrigeration machinery of heretofore unheard of reliability; and low-cost, reliable cryogenic envelopes to house and thermally

insulate the low-temperature cables. But, in addition to these, it was recognized that the circulation of helium at temperatures and pressures required

to refrigerate superconductors in channels kilometers long presented a unique

thermal-hydraulic situation; no such flow system had been operated prior to the

inception of this cable development work.

The combination of cable refrigerant channels, whose walls have low thermal

capacitance at liquid helium temperatures, and the strongly temperature dependent density of the helium at the thermodynamic states encountered in the refrigeration of the superconductor could be shown on theoretical grounds to lead to

flow instabilities in some circumstances. This situation was compounded by the

fact that available instrumentation for basic helium measurements were either

unavailable or primitive and some special problems resulted from the high voltages present in the cable.

These are the general problems addressed by the current program of re- search at the Cryogenics Division of the National Bureau of Standards. The

goals are to identify the conditions under which stable flow and temperature

of the refrigerant can be expected in order to maintain the conductor at its

proper working temperature, notwithstanding the perturbing effect of network

fault currents; further, to conduct such research and development on measurement

techniques as to ensure that the performance of model and prototype cables and

their refrigeration systems can be intelligently monitored.

ii
CONTENTS

Page

1.0. CURRENT LEADS COOLED WITH SUPERCRITICAL HELIUM FOR SUPERCONDUCTING POWER TRANSMISSION LINE TERMINATIONS 1 1.1. Introduction , 1

1.2. Similarity Considerations 4 1.3. Experimental Details 6 1.4. Thermal Runaway Studies 13

1.4.1. Computer Program for Lead Performance 22

1.4.2. Experimental Results 24

1.4.3. Discussion of Results and Comparison with Theory ... 32

1.5. Oscillatory Behavior 35

1.5.1. General Description of Observed Oscillation 35

1.5.2. Discussion of Oscillations 39

1.6. Conclusions and Recommendations '^^ 1.7. References . . .

2.0. HELIUM MEASUREMENTS 48

2.1. Data Transmission from the Superconducting Power Line

to Ground Instrumentation 4 8 2.1.1. Power Supply 48

2.1.2. Equipment Specifications 48

2.1.3. Transmission to Ground 50

2.1.3.1. Capacitive Coupling 50

2.1.3.2. Fiber-optics 50

2.1.3.3. Radio Transmission 50

2.2. Thermometry 51

2.3. Pressure Determination from Density and Temperature

Measurements 51

2.4. Microwave Cavity Pressure Transducer 52

2.4.1. General Circuit Concept 5 3 2.4.2. Material 53

2.4.3. Cavity Geometry 54

2.4.4. Diaphragm Thickness and Transducer Sensitivity .... 59

2.4.5. Electrical Circuit 64

2.4.6. Transducer Tests 68

2.4.7. The Coaxial Connector 77

2.4.8. Summary and Conclusions 77

2.5. Helium Impurity Study 77

2.5.1. Assistance to Brookhaven National Laboratory 77

2.5.2. Instrvimentation for Air Impurities in Helium 77

2.6. References 81

3.0. HELIUM PROPERTIES

8 2 3.1. Improved Helium Property Computer Codes

3.2. Thermal Conductivity of Helium ^-^ 3.3 References 8 3

4.0. FLOW FACILITY FOR MODELLING HELIUM FLOW PASSAGES OF HIGH

ASPECT RATIO (L/D) ("LONG LINE" APPARATUS) 84

4.1. Apparatus 84

4.2. Instrumentation 84

4.3. Status 9 0

iii
LIST OF FIGURES Page

Figure 1.1. Sketch showing model lead construction 3 Figure 1.2. Flow schematic of lead test apparatus 7 Figure 1.3. Installation of lead no. 1 8

Figure 1.4. Installation of lead no. 2 . . 9 Figure 1.5. Installation of lead no. 3 10

Figure 1.6. Installation of lead no. 4 11

Figure 1.7. Alternative installation of lead no. 3 12

Figure 1.8. Temperature profiles for lead no. 1 14

Figure 1.9. Temperature profiles for lead no. 3. Helium flow rate

0.0117 g/s, pressure 1.0 MPa, helium inlet temperature

4.5-6 K 15

Figure 1.10. Temperature profiles for lead no. 3. Helium flow rate

0. 0233 g/s, other conditions as for figure 1.9 16

Figure 1.11. Temperature profiles for lead no. 3. Helium flow rate

0. 0457 g/s, other conditions as for figure 1.9 17

Figure 1.12. Temperature profiles for lead no. 4. Helium flow rate

0. 0233 g/s, other conditions as for figure 1.9 18

Figure 1.13. Temperature profiles for lead no. 4. Helium flow rate

0.070 g/s, other conditions as for figure 1.9 19

Figure 1.14. Comparison of temperature profiles for lead no. 3.

Connection A as for figure 1.5; connection B as for

figure 1.7. Helium flow rate 0.0168 g/s, other

conditions as for figure 1.9 20

Figure 1.15. Comparison of temperature profiles for lead no. 3.

Helium flow rate 0.0308 g/s, other conditions as for

figure 1.14 21

Figure 1.16. Comparison of computed thermal conductivity of copper

with experimental data 25

Figure 1.17. Total Carnot work per ampere of lead current vs m/I. . . 26

Figure 1.18. Total Carnot work per ampere of lead current vs IL/A. . . 27

Figure 1.19. Thermal runaway tests for lead no. 2. Helium flow rate

0.005 g/s, pressure 1.0 MPa, helium inlet temperature

4.5 K 29

Figure 1.20. Thermal runaway tests for lead no. 2. Helium flow rate

0.010 g/s, other conditions as for figure 1.19 30

Figure 1.21. Thermal runaway tests for lead no. 2. Helium flow rate

0.0149 g/s, other conditions as for figure 1.19 31

Figure 1.22. Generalized plot of thermal runaway results 34

iv
List of Figures (continued) Page

Figure 1.23. Comparison of experimental steady state temperature

profile with calculations for various hP 36

Figure 1.24. Oscillation amplitude vs. lead current for lead no. 1. . 37

Figure 1.25. Current-flow rate conditions necessary to eliminate

oscillations in lead no. 1 for a helium inlet temperature

4.3-4.5 K. The family of lines shifts to the left for

higher inlet temperatures 38

Figure 1.26. Frequency of oscillations as a function of inlet helium

density 40

Figure 1.27. P/ T regions in which oscillations were observed for

lead no. 1 41

Figure 1.28. p,T regions in which oscillations were observed for

lead no. 2 42

Figure 1.29. Stability boundaries from the calculations of Rott com- pared with observations of this work. The dashed lines

represent experimental conditions where oscillations

were not observed, the solid lines where oscillations

were observed 44

Figure 2.1. Schematic outline of the transducer signal path from

high voltage to ground potential 49

Figure 2.2. 440 MHz resonant cavity pressure sensor 55

Figure 2.3. Geometrical parameters of the microwave cavity 56

Figure 2.4. Assembly of the microwave cavity pressure transducer. . . 57

Figure 2.5. Design detail on the two pieces which are soldered

together to form the cavity. 58

Figure 2.6. Schematic outline of the diaphragm distortion under

pressure, defining relevant parameters (not to scale).. . 60

Figure 2.7. Normalized pressure sensitivity of the resonant frequency as a function of the diameter ratio d/b.

Parameters are defined in the text ^3

Figure 2.8. Design of the capacitive probe to the cavity 65

Figure 2.9. Block diagram of the electrical circuit 66

Figure 2.10. Output frequency of the voltage controlled oscillator

as a function of applied voltage 67

Figure 2.11. Approximate representation of the reflected electromagnetic power sensed at the diode detector as a function of frequency. The top curve representes the

ideal case, for three different values of the coupling

strength. The middle curve represents the curve between the cavity and the directional coupler. The

bottom curve portrays the additional perturbations as the coaxial line length is increased. The dashed line

represents the discriminator voltage cutoff as discussed in the text. 69

V
List of Figures (continued) Page

Figure 2.12. Detailed schematic diagram of the detection and

demodulation circuits

Figure 2.13. Schematic outline of the power supply connections

and the analog output voltage circuit . . . Figure 2.14. Calibration of cavity #2 at 76 K, showing some residual hysteresis , .'.

Figure 2.15. Calibration of cavity #2 at helium temperature,

showing substantial hysteresis

Figure 2.16. Uncorrected calibration data for transducer #3, for

three different temperatures. The apparent irreversibility is opposite in direction to that expected from

hysteresis, and is interpreted as drift in the electronics. The discontinuity in the 76 K data was observed during an experimental interruption

Figure 2.17. Schematic of helium purity calibration apparatus. . . Figure 4.1. Flow schematic of Long Line apparatus

Figure 4.2. General view of assembled apparatus

Figure 4.3. View from below of test section mounted inside

experimental vessel

Figure 4.4. Positions of test section instrumentation

Figure 4.5. Thermometer well-pressure tap detail

70

71

74

75

76

79

85

86

87

88

89

vx
ABSTRACT

This report is the second annual report of research on helium related

problems in support of superconducting power transmission development. The

report is in four sections.

In the first section; results are presented from experimental and com- puter modeling of the performance of current leads cooled with supercritical

helium gas. Performance characteristics studied are burn-out conditions and

existence of oscillation in the helium gas.

The second section, on helium measurements, reports some conclusions on the feasibility of data transmission from high-voltage regions to grounded

read-out instrumentation, on thermometry and on helium impurity measurements.

A major part of this section is a detailed description with test results of

microwave cavity pressure transducers for use at helium temperatures.

A third section, on helium properties, reports some improvements in com- puter codes for helium properties and discusses some recent data on the

thermal conductivity of helium.

In the final section, a description is given of a recently completed

flow facility which has been built for research on flow and heat transfer

dynamics of supercritical helium in channels of high aspect ratio modeling

superconducting power transmission line channels.

Key Words: Burn-out; helium-cooled current leads; helium flow facility;

helium impurities; microwave pressure transducer; thermal-acoustic oscillations .

vii

1.0. CURRENT LEADS COOLED WITH SUPERCRITICAL HELIUM FOR

SUPERCONDUCTING POWER TRANSMISSION LINE TERMINATIONS

1.1. Introduction

Much has appeared in the literature on the thermodynamic optimization of

current leads, but it becomes increasingly clear for electric power applications that considerations of efficiency must give way to those of reliability,

safety and operability. This has been the emphasis of the NBS program. High

voltage aspects have not concerned us, however, since these are under study

elsewhere and the thermal performance appears to be a separable problem.

Our concern at the outset was that, although helium-cooled current leads were common everyday accessories in many laboratories, almost all of them

were of the self-sustained type cooled by boil-off helium from a dewar at a little over atmospheric pressure. SPTL leads would require the use of high

pressure helium and the gas flow would be set independently of current. This

had implications as far as thermal runaway, or burn-out, is concerned, and we also had some concerns about the possibility of thermal acoustic oscillations

in the helium, a commonly observed phenomenon in tube penetrations into

liquid helium dewars. Our approach, therefore, was to build small models of

candidate designs and to run these at current densities and helium conditions

typical of proposed SPTL leads.

Unfortunately, at the time we began this work, only one such lead had

been designed for this application and this was of the conventional Efferson

braid type adapted to a.c. conditions by having transposed braids. We therefore chose basic model geometries whose operational characteristics, if not

detailed construction, could be scaled up to desired capacity.

Four models of different design were studied; the essential features of

each are given in table 1 and figure 1. The first two leads were designed to

encompass a wide variation in flow and heat transfer characteristics. Thus

in no. 1 - a simple tubular conductor - the gas friction factor and heat

transfer coefficient for a given Reynolds number should be the lowest limit

achievable, as should also the heat transfer area per unit length in relation

to the conductor cross section. In lead no. 2 - a braided lead in an annular

space - these factors were expected to be about as high as can be achieved.

The heat transfer aspects of a lead, as shown below in section 1.2, are

2 characterized by the dimensionless group H = hPL /kA, where h is the heat

transfer coefficient, P the wetted perimeter (or heat transfer surface area

per unit length) , L the lead length, k the conductor thermal conductivity at

reference conditions and A the cross sectional area of the conductor. This

factor based on an unenhanced laminar flow heat transfer coefficient is also

listed in table 1.1.

Lead no. 3 is a special tapered lead design of larger capacity constructed by the Los Alamos Scientific Laboratory. A large heat transfer area is

obtained in this case by soldering an externally threaded copper tube in the

center of a hollow conductor generating three parallel spiral channels

against the bore, the central passage being plugged.

While we had tested nos . 1 and 2 with a return lead which was similar in

construction to no. 2, we needed a new return lead of greater length and

capacity for use with no. 3. We chose a length of copper heat exchanger

tubing consisting of a coiled copper spiral ribbon insert soldered between a central core and an outer tube used as the conductor. We also tested this

lead as no. 4. This choice had one advantage that heat transfer and pressure

drop characteristics were well documented [1] as well as its being naturally

an efficient heat transfer device. Tests on leads 3 and 4 were restricted to one pressure (1 MPa) and an inlet helium temperature range of 4.5-6 K while

those for nos. 1 and 2 covered the range from 0.25 to 1.5 MPa and 4.5 to 12 K.
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In section 1.2 we discuss the similarity laws for current leads. After

a brief description of the experimental details of our test apparatus in

section 1.3, we discuss our observations of thermal runaway and computations

of this phenomenon in 1.4. In section 1.5 oscillatory behavior observed in

testing the leads is discussed. Finally, in section 1.6 we list some recom- mendations for lead design resulting from this study.

1.2 Similarity Considerations

It is useful to consider under what conditions a model and full scale

current lead can have identical temperature profiles. This can be done by a dimensional analysis of the governing differential equations and boundary

conditions. We assume that the temperature boundary conditions at the lead

ends (both conductor and gas) are the same and that the same material is used

for the conductor. This analysis does not depend upon the often used simplifications that the Wiedemann-Franz law is obeyed, or that the thermal con- ductivity is a constant, or that the heat transfer coefficient and specific

heat of the gas is a constant. Furthermore, a variable conductor cross sectional area may be admitted.

For a given construction and conductor material with thermal conductivity k(T), resistivity p (T) , cross sectional area A(x), operating between

the lower and upper temperatures T and T for the conductor and 6 and 9 '^'^ o u o u for the gas, the governing differential equations for one dimensional heat

flow in the conductor and gas are

d

(j^A

dT dQ ^ pl^ ^ Q dx dx dx A

*C

^ = ^ = Ph (T-e) , (1.2)

p dx dx

dQ/dx is the lateral heat transfer to the gas per unit length, A the gas flow

rate and P the perimeter of the conductor in contact with gas.

We now define reference values (eg. at T^) , denoted by asterisks, for

the temperature dependent quantities k, C , h. We also define dimensionless

P

k*A* distance n by dn =

^^j^ dx and.

k*A*

o where L is the length of the lead. Then the differential equations may be

written

2 2

2 L LI d ,dT, , kA . ,h , ,Ph*L > ,„ a\ , /kp . , o , _ n n a\ dTT ^dT^) - (k^) ^h^) ^k^x^) ^ ^r^) 27;:2)

^= ° (^-^^ o k* A*

de _ , Ph*L . ."^p*. .h . kA . . ,^ ^> p o where we have introduced the Lorenz constant L (=! kp/T) . Insofar as r-r-

,

^icp^i, kA

I

P

p ^nd kp are dimensionless functions

on"

of n , T and 6, the solutions

p o

4
for T and 0 must have as sole parameters the three dimensionless quantities

1/2

L ^ LI mC * pk*t2

J = ' ^' = Ph^ H = |?A^

mC *L

or, alternatively, we can choose the equivalent three parameters J, M = . Ph*L and H' = —-. ,~ . For a given conductor material, therefore, to achieve idenL ^ I o tical temperature profiles in a model and full scale lead we must preserve

the ratios IL/A*, mL/A* (or m/1) and Ph*L^/A* (or Ph*L/I)

We will consider now what the implications are for scale-up, for the

optimum operating condition and for the burn-out, or thermal runaway condition.

i) Scale-up. The first two parameters J and M give the fairly obvious

requirements that, for the same length, current scales with cross sectional

area and mass flow rate with current. Alternatively, if cross section is to

be held constant, current and mass flow must scale as the inverse of the

length. The consequences of the third parameter H may be best illustrated by

considering the scale-up of a lead without change in length. In this case

h* kg*/Djj if hydrodynamic similarity is maintained, or the the flow is

laminar, where k^^* is the thermal conductivity of the gas at reference con- ditions and D„ is the mean hydraulic diameter. Thus, one class of lead that

may always be scaled up is the matrix (eg., bundles of wires, braids, sintered spheres) simply by maintaining the same diameter filament or sphere

with the same packing density. A simple tubular conductor scales up into

many identical tubes, but not into a larger tube. A finned conductor will

scale-up if the fin dimensions and conductor radial thickness are held constant. If H is large enough, say H > 10,000, we have ideal heat transfer.

Then 9 T and a temperature profile for given material and end temperatures

is determined by J and M alone making scale-up straightforward for any such

lead.

ii) Optimum Operating Conditions. If one is not constrained by the

performance characteristics of an already existing refrigerator a lead must

be optimized by minimizing the ideal Carnot work required to both refrigerate

the cold end and supply the cooling gas. The optimum condition then corresponds to a particular temperature profile. If we neglect the electrical

energy lost in the lead through Joule heating it is a profile with zero temperature gradient at the warm end as shown by Lock [2] for self-sustained

leads with ideal heat transfer, by Pippard [3] for the more general gascooled lead with ideal heat transfer, and by Agsten [4] for a lead cooled in

a thermodynamically optimum manner. But whether we neglect the Joule heating

or not a fixed profile in dimensionless form results. However, it is not

clear that the same profile still holds for less than ideal heat transfer.

Again, if we are to run both the model and full-scale lead at optimum

conditions, this can be achieved by preserving the ratios IL/A*, m/I and

h*PL/I. The Carnot work for the cold gas is simply proportional to m which is

proportional to I, and the Carnot work for the cold end refrigeration is

dT proportional to A ^ ^_Q. The latter can be obtained from the optimum

AdT A* dT dimensionless profile: — „

-— Therefore the optimum cold end ^ dx x=0 L dn n=0

refrigeration is also proportional to the optimum I. As is well known, if

heat exchange approaches the ideal case, optimum total refrigeration per

ampere should be independent of the lead geometry. It also follows in this

5
case that the mass flow per ampere and the factor IL/A* should be the same

for all optimized leads of the same material operating between the same temperatures

.

iii) Burn-out Conditions. Burn-out of a lead corresponds to a singularity in the temperature for certain combinations of J and M. Experimentally, above a certain current for fixed mass flow rate the lead temperature

rises at ever increasing rate, or runs away, instead of approaching a limit.

The question is, for the purposes of this work, if we know burn-out conditions for a model lead can they be predicted for the full scale? Since again

burn-out is related to a property of the temperature profile we simply have

to preserve the same three scaling ratios. However, this case differs from

the consideration of optimum conditions in that burn-out can occur at some current for all values of m for a given lead, so we need to determine the

burn-out relation I vs. m. More generally, at burn-out M is a function of J

with H as parameter, but note that H may be a function of M. The general

case has not been tackled, but Bejan and Cluss[5] have recently computed the

relations for the simple model of constant thermal conductivity, and resistivity linear with temperature. In the absence of a more general solution,

burn-out must be modeled with a lead of the same material and same end temperatures with the same H. Alternatively, exact numerical calculations can

be made for a given material if the thermal conductivity and electrical

resistivity are known as functions of temperature and if the heat transfer

coefficient and its flow dependence are known.

In summary, similarity considerations lead us to the laws of scale-up

applicable to a model and full scale lead of the same material, and hence to

the prediction of the optimum operating point and the burn-out curve once

these are known for the model.

1.3. Experimental Details

A description and flow scheme for the experimental arrangement for

testing the model leads appeared in our previous annual report. [6] In view

of some minor modifications a revised flow scheme is given in figure 1.2.

The major revision is the addition of a small circulating pump to spray

liquid helium at 4 K onto the cold bus bar which connects the two current

leads under test. This was found to be necessary because of inadequate heat

exchange between the cold end of the leads and the helium supply pressure

vessel. The pump together with a heater on the pressure vessel allowed

flexibility in running the cold end of the leads from 4.5 to 15 K. The leads were instrumented with six or ten thermocouples evenly spaced

along their length and referenced at the cold end to a germanium resistance

thermometer mounted in a heavy copper fitting. This also formed the electrical end of the lead. Each lead had its own separate vacuum jacket and the

whole assembly plus the return lead assembly was soldered into a 15 liter

pressure vessel with a heavy copper bus bar connecting the two leads at the

base. Detailed drawings showing the mounting of the four model leads are given in figures 1.3, 1.4, 1.5 and 1.6 respectively. The lead no. 3 was supplied to us with two alternatives for electrical connection. In the

first, shown in figure 1.5, the electrical connection was directly to a boss

at the end of the outer copper conductor. In the second alternative, current

was supplied via an extension on which was wound a niobium-tin superconducting

wire. This was then soldered onto the helical passages for 25 cm up into the

lead as shown in figure 1.7. Both modes of connection were used in comparative tests.

Helium gas was supplied to the pressure vessel from a tube trailer via

purification and heat exchange systems. The temperature of this vessel could

be controlled by a flow splitting valve which directed up to 100 percent of

the stream through a liquid helium heat exchanger. On account of the large

6
7
Bus-bar from current supply

45.7

Thermocouple to sense | exit gas temperature

^ , ^ Terminal post, copper

Dielectric insert

(A Conductor (048 ID x 0.025 wall)

' phosphorized copper)

Thermocouple. Gold (0.07 at % iron)

vs KP (chromel)

Dewar top plate

Copper bus-bar to return current lead

Thermocouple reference junction

Radiation shield (10 cm long)

Terminal post % thermometer block, copper

Germanium thermometer

Dielectric coupling

J Helium inlet

Figure 1.3. Installation of lead no. 1

All dimensions in cm.

8
Thermocouple to $en»e

exit gai temperature

Dielectric incert

t 1

Thermocouple #10-

4.0

#9

Soft solder braid to copper

terminal over this length

—Terminal po»t. copper

0.95

.076 Annular braid space

45.7

Thermocouple #1

Reference thermocouples

#2

f 2 sJ

I

Copper bus-bar to return current lead

Stainless steel tube

0.47 O.D. X .025 wal

Phenolic rod, 0 27 D.

Copper braid) 192 Tin coated

wires of 34 gauge (0.016 D.)

1:

Copper radiation shield

(10 cm long) — Stainless steel vacuum jacket

-Germanium thermometer

Soft solder braid to copper

terminal over this length

Connection to helium inlet as for current lead #1

All dimensions in cm.

Figure 1.4. Installation of lead no. 2.

9
Thermocouple to sense

exit gas temperature

3 Separate spiral grooves

(2 mm X 2 mm cross section area) Groove lead = I.I cm

Dielectric Insert

Hollow core plugged

at both ends

Equal spacing of 10. 3 cm between Thermocouples I thru 10

Bus-bar from current supply

Rose's alloy solder

Insulating vacuum space

Three Nba-Sn Superconducting Wires

soldered inside groove of the lower Jkm? 25% of the current lead "^"^

HELIUM FLOW 100

Tapered Conductor

(Taper = 0°-9'-2l")

Thermocouple #1

1.27

Reference Thermocouples

Nbj -Sn wires

Cylindrical copper shield (24 cm long)

Stainless steel vacuum jacket

Germanium Thermometer

Copper bus-bar

return current lead

All dimensions in cm. Figure 1.5. Installation of lead no. 3. 10
Thermocouple to sense

exit gas temperatureDielectric insert

#6

6 Thermocouples equally

^

spaced at 18-5 cm

1

Bus-bar from current suppl>

1

100

Heat exchanger tubing

1.7 O.D.x 1.26 l.D.with .3 solid core

Thermocouple #1

T"

6.5

Copper shield (24 cm long)

-Germanium thermometer

---Reference thermocouples

Copper bus-bar to return current lead

All dimensions in cm.

Figure 1.5. Installation of lead no. 4.

11
Tapered current

lead conductor

4.44 O D.x .089 wall

Stainless steel tube —

Rose's alloy solder

Soft solder

Rose's alloy solder-^

154 O.D.X.05 wall

Stainless steel tube

.063 Wall phenolic

insulation

Soft solder

Rose's alloy solder-^

Stainless steel stub on

top of pressure vessel

Tapered length=IOO

Taper angle=0°-9'-2l"

0. D. @ ends=I.OI I 1.56

1. D.=.952

Germanium thermometer

Reference junctions for

current lead thermocouples

Connecting boss at base

of current lead

Current lead core with spiral grooves

Superconducting wires

Soldered to 1.27 O D.x

.089 wall Cu tube

Cu coupling, .083 wall

Hard solder

Copper bus-bar return

current lead —

Germanium thermometer

All dimensions in cm.

Figure 1.7. Alternative installation, of lead no. 3,

12
reservoir capacity and low flow rates, measurements could be made with gas

inlet to the reservoir closed. The flow through the test leads was regulated

manually by a valve in the warm gas exit piping. At this valve gas pressure

was reduced to essentially atmospheric. Finally, both ac current to 2000 A

and dc to 60 0 A could be supplied via bus bars to the warm end of the lead.

The following measurements were made during each test.

1. Conductor and helium warm, exit gas temperatures: Gold (0.07 at % iron) vs. KP (chromel) thermocouples. Estimated total error between + 0.05 K

at 4 K and +K at t 300 K.

2. Helium inlet gas temperature: Germanium resistance thermometer.

Estimated total error + 0.01 K.

3. Helium mass flow rate through lead: Thermal flow meter. Estimated

total error + 0,001 g/s.

4. Helium reservoir pressure: Magnetic reluctance transducer. Estimated total error + 0.01 MPa.

5. Pressure oscillation at lead warm end: Magnetic reluctance transducer. Frequency range as installed 0.1-200 Hz.

6. Lead current (+ 3% ac , + 1-1/2% dc).

Representative temperature profiles for lead nos. 1, 3 and 4 are given

in figures 1.8 through 1.13. In figures 1.14 and 1.15 a comparison of profiles for lead no. 3 is shown with the alternative methods of connection at

the cold end. The use of the superconducting current path does not appear to

have been of benefit; however, we feel that design of the cold end connection

m.ay not have been optimum for the heat flow in the case which utilized the

niobium-tin superconductor. Comparison of profiles for leads 3 and 4 should

take into account the much greater purity of the conductor for no. 3 as evidenced by the greater residual resistivity ratio. Further evaluation of

this lead is being conducted by the Los Alamos Scientific Laboratory.

Lead no. 2 was selected for a detailed study of thermal run-away and

temperature profiles are given and data discussed in section 1.4.

1.4 Thermal Runaway Studies

While the analysis and verification of optimum operating conditions for

helium-cooled current leads has received a great deal of attention, the

problem of thermal runaway, or the accidental burn-out of a lead, has been

neglected with the exception of the work at MIT. For current leads cooled

with supercritical helium gas at, say, 1.0-1.5 MPa, the helium flow rate may

be set independently of the lead current and so the "flow choking" mechanism

of burn-out suggested by Thullen, et al. [7] is not relevant. These authors

correctly note that pressure drop increases with reduction in mass flow rate

when currents are greater than a critical value. If the majority of the

available pressure drop is taken across the lead itself the flow is unstable

and must reduce indefinitely if the current parameter J is greater than tt. For leads designed for SPTLs , however, or any system where high pressure

helium is used, the flow will be stabilized by taking the majority of the

pressure drop across a control valve.

For the SPTL application, the analysis of Bejan and Cluss[5] is more

appropriate although this is based on the constant conductivity/linear re- sistivity model and is thus approximate for most coppers that would be con- sidered in a lead. However, one new feature of this analysis is that it is

13
Figure 1.8. Temperature profiles for lead no. 1

14
15

17
18
Figure 1.13. Temperature profiles for lead no. 4. Helium flow rate 0.070 g/s,

other conditions as for figure 1.9.

19
Figure 1.14. Comparison of temperature profiles for lead no. 3. Connection A

as for figure 1.5; connection B as for figure 1.7. Helium flow

rate 0.0168 g/s, other conditions as for figure 1.9.

20
x/L

Figure 1.15. Comparison of temperature profiles for lead no. 3. Helium flow

rate 0.0308 g/s, other conditions as for figure 1.14.

21
applied to the case of imperfect heat transfer and the burn-out predictions

for J are given as a function not only of M but H also. From these results

one can at least see how large H must be for burn-out to be independent of H;

4 practically speaking if H is greater than about 10 burn-out occurs at the

ideal heat exchange criterion

(1.6)

This relationship is only true for constant thermal conductivity, linear

resistivity, but it is helpful in interpreting experimental results. For if

in the general real property case we define the dimensionless lead length as

the criterion now becomes

,2

A*k*

o ^

(1 7) A(x)k(T) ^^''>

and if k(T) > k* for all T (RRR > = 50) then for a constant area lead Ht < 1 and burn-out occurs at higher values of J for given M. For some commercial

non-electrical coppers, k(T) < k for all T, ri-r > 1, and we may expect burnout to occur at lower J. This would be for RRR < = 9 and it leaves a grey

area between. One cannot be more quantitative without knowing the temperature profile at burn-out. This then becomes a difficult iterative process.

In order to be able to shed light on our experimental results we decided

that an accurate and general calculation should be available, which invariably means that one resorts to a numerical method. The computer program

described in the next section was, we thought, as general as one could make

it since none of the common approximations are made. Comparison with ex- perimental results, as will be seen, revealed that even this was not enough

and there may be some aspects of real leads which simply are not well enough

known until they are tested.

1.4.1. Computer Program for Lead Performance

The computer program simultaneously integrates the exact one-dimensional

momentum and energy equations given by Arp [8] for the helium gas, and the

one-dimensional heat conduction equation for the conductor. This is then

coupled to the helium energy equation by the heat transfer term. A combined

Runge-Kutta, Adams-Moulton predictor-corrector method is used with complete

property updating and error checking for both helium and conductor at every

step. Thus we have the simultaneous set of equations:

Helium pressure:

Helium enthalpy;

dP ^ p

(1-m2) - ^(-l+(j)M^) F^ - (!'M^F2j . (1.9)

dH

dx

^

2 ^{-l+<t>)M^F^ + (l-(l+(|))M^ F^. (1.10)

Conductor temperature:

(

dT

, Y . (1.11) dx

22
Conductor heat flow:

2„ ^ dk

To commence the integration at x = 0 initial values of p, H, T and Y are

given. In these equations, p is the helium pressure, H the helium enthalpy

(not to be confused with the heat transfer parameter H defined above), M

is the Mach no., C the adiabatic sound velocity, K = — | 4-^ | , and i) is the

/3t\ P ,, ^ V'^/S

Gruneisen parameter = l^r—] . F, - -ttt:— M M f With A„ = helium gas flow ^

s " ""G cross sectional area, P the wetted perimeter and f the Fanning friction

factor; = X/ihC^ with \ the heat transfer per unit length ^ . In the

usual manner the integration is iterated on dT/dx at x = 0 until at x = L the

required warm end temperature is found. It will be noted that the thermodynamic state of the helium is calculated with coordinates p and H. From

this the temperature 0 and density p„ are determined from explicit functions

8(p,H) and p^{p,E) developed especially for fast computer evaluation as is

required here. The derivatives $ , C and K are obtained from the equation of

state of McCarty. [9]

The conductor properties may be strongly temperature dependent. In view

of the large deviations from the Wiedemann-Franz Law for even moderate purity

coppers - for example 50 percent at 40 K for RRR = 100 - we decided not to

invoke the Law except for the imperfection contribution k^ , but to compute k^

and p separately. For p we used Matthiessen ' s rule, p = p^ + c.^. and calculated the residual resistivity p^ from the specified RPR and the ideal

lattice resistivity -p^ from the data given by White and Woods. [10] Our

curve, fitted to those data, is

ln(p^ X 10^) = - 9.600976 - 12.52445 InT + 8 . 309361 (InT) ^ - 1 . 583458 (InT)

+ 0.0993132 (InT)^. (1.13)

The deviation from the data is no more than two percent. This equation is

used up to 300 K, after which p^ is assumed proportional to T. For the conductor thermal conductivity k^ we used the approach followed

by Lock. [2] In this approach an analogue of Matthiessen ' s rule is used.

Thus we take — = — + — (1.14)

k k k^ • s o I This is a compromise between accuracy and computational efficiency. The

procedure is known not to be accurate above 10 K (the error can reach 10

percent) for the thermal resistivity, but for engineering purposes the possibility of computing the thermal conductivity from a single property - RRR - is very appealing. We calculate k from L T/p using an average experi-

-8 2 2 o o mental value for L of 2.1 x 10 V /K . The lattice contribution to thermal

resistivity 1/k was assumed to be independent of purity and calculated from

the data of Powell, et al. [11] for a sample of RRR = 110. Thus we compute k^

from the expression

23
5 k = T exp I C./T^

1

^ (1.15)

with C 146.587,

269.117,

16956.6,

101324. 0,

-171156.0.

This expression is used up to 150 K after which the thermal conductivity is

assumed to be constant at its 150 K value. The fit generally predicts the

thermal conductivity data of Powell to better than five percent but is low by

14 percent at 40 K. Nevertheless, considering the complex nature of the

temperature dependence of the thermal conductivity and its variation by

orders of magnitude at low temperatures, this analytical representation gives

a very realistic and economical representation. This is illustrated in

figure 1.16 where we have compared some of our computed curves with experimental data for which RRR was known and was of comparable magnitude.

In calculating the heat transfer term X we put as usual A = Ph(T-6). We

calculate h based on the mean hydraulic diameter D„ = 4A_/P and appropriate

n Kj expressions for laminar and turbulent flow in pipes as appropriate to the

calculated Reynolds number. In most cases Reynolds numbers indicated flow

would be laminar and we used the flow independent Nusselt number, Nu = hD„/k„ = 4.364. It was felt that in cases other than a simple tube the

rl yj Nusselt number would be enhanced by tortuous geometry so that this approach

should be conservative and not overestimate the heat transfer term.

As a demonstration of the computer program we first used it to calculate

the optimum operating conditions for a lead of the design of no. 2 with RRR =

50 and operating between 6 and 300 K. In figure 1.17 we plot the total

specific Carnot work vs. the specific gas flow rti/I and in figure 1.18 we show Wrp/I vs. IL/A. The procedure was to vary IL/A and find the minimum

Carnot work for fixed m. These minima are then plotted showing that the

absolute minimum W^/I for this lead is 0.270 W/A and occurs at IL/A = 2.2 5 x

5 -5

10 A/cm and rti/I = 4.62 x 10 g/As . It is interesting to compare this

result with the values given by Giisewell and Haebel [13] for a slotted finned

rod with RRR = 57, The corresponding figures are W^/I = 0.333 W/A, IL/A'

2.6 X 10^ A/cm and m/I = 4.9 x 10~^ g/As. Agsten gives W^/I = 0.340 W/A for

a Wiedemann-Franz material operating between 4.2 and 300 K as the absolute

thermodynamic optimum. The differences appear to be chiefly a result of

operating the cold end at 6 K in our case compared to 4.4 K and 4.2 K respectively. This permits a relatively greater amount of heat to be removed at

the cold end through the alleviation of the Carnot factor and thus less gas

is used.

The program was next used to investigate thermal runaway. We attempted

to duplicate the experimental data described in the next section. Here, the

procedure was to compute temperature profiles for a fixed mass flow rate,

incrementing I until a solution could no longer be obtained, or excessively

high conductor temperatures were encountered. The results will be discussed

in section 1.4.3.

1.4.2. Experimental Results

A careful study of thermal runaway was made on lead no. 2. This lead

had previously burned out during exploratory runaway tests, but was repaired

24
50 r20

10

2 -

0.5

0.2 1^ I I I

RRR=110(REF. lip

RRR= 39(REF.12)«

RRR= 4(REF.12)A

Computed —^—

J \ I I I I I I I 10 20 50 100 200

TEMPERATURE (K)

Figure 1.16. Comparison of computed thermal conductivity of copper with

experimental data.

25
26
Figure 1.18. Total Carnot work per ampere of lead current vs XL/A.
by replacing part of the burned-out stainless steel tubing and by completely

replacing the braid. We also extended the phenolic resin rod by a little

over a centimeter since there was evidence of poor thermal anchoring at the warm end

.

The lead was run at three mass flow rates: 0.0047 g/s , 0.0093 g/s and

0.0149 g/s. The middle value v/as close to what we had determined to be the mass flow rate under optimum conditions. For each flow rate the current was increased in increments of five or ten amperes. When at a particular setting

the temperature profile had stabilized it was recorded. When for a given

setting the temperature profile still continued to increase with no sign of

slowing down after 10 minutes, or an excessive temperature had been reached

(360 K) , the test was terminated and the last two currents were assumed to

bracket the critical burn-out current. In this way the burn-out current has

been established to within at most five percent. The tests were carried out

at 0.4, 1.0 and 1.5 MPa and the results are given in table 1.2. Temperature

profiles for the three flow rates at 1.0 MPa are given in figures 1.19, 1.20,

and 1.21. Pressure in fact made little difference. In all of these tests

the lead cold end temperature was held at 6 K while the gas inlet temperature

was 4.5 K.

In addition to measuring temperature profiles we also measured the lead

resistance and the values are also given in table 1.2. It is clear that this

could be a useful indicator of approach to thermal runaway, but only at a fixed helium flow rate.

Table 1.2. Thermal Runaway Tests. Lead No. 2.

Pressure

Flow

Rate M

2 Current(A)

23

M

Stable

or LeadResistance

(g/s) Unstable (y ohm)

0. 0047 3.59 115

125

1.49

1. 61

S

U

947

956

0.4 0.0093 7.18 185

190

1.16

1.20

S

U

748

750

0.0149 11.48 250

260

1.01

1.05

Su 307

588

0.0047 3. 59 110 1. 42 s 857

115 1. 49 u 1039

1.0 0. 0093 7.18 190

195

1.23

1.26

su 628

704

0.0149 11.48 265

270

1. 07

1. 09

su

374

549

0.0047 3. 59 115

120

1. 49

1.55

su 849

1010

1. 5 0.0093 • 7.18 190

195

1.23

1. 26

su

532

697

0.0149 11.48 265

270

1. 07

1. 09

su

430

574

Conditions of Tests: d.c. current , helium inlet temperature 4.5 K, conductor

cold end temperature 6.0 K.

28
x/L

Figure 1.19. Thermal runaway tests for lead no. 2. Helium flow rate 0.005 g/s,

pressure 1.0 MPa, helium inlet temperature 4.5 K.

29

x/L

Figure 1.21. Thermal runaway tests for lead no. 2. Helium flow rate 0.0149 g/s,

other conditions as for figure 1.19.

31
In the course of running a series of tests for the LASL lead, no. 3, we encountered runaway conditions and this allowed us to obtain a rather coarse

but nonetheless useful bracketing of the runaway conditions for both leads 3 and 4. The data are given in tables 1.3 and 1.4. The response time of these

two leads is very much greater than for nos . 1 and 2 making this kind of test

difficult; it was necessary to wait on the order of one hour for the establishment of equilibrium for nos. 3 and 4 compared with ten minutes for lead

no. 2. No thermal runaway tests were carried out on lead no. 1.

Table 1.3. Thermal Runaway Tests. Lead No. 3. Pressure i. (MPa)

Flow Rate

(g/s)

Current

(A)

Stable or Unstable

0. 0117 475

585

S

?

0. 0168 625

775

Su

1.0 0.0233 600

775

su

0.0308 720

930

su

0.0457 900

1160

su Conditions of Test: a.c. current, helium inlet temperature 4.5-6 K, conductor

cold end temperature 10.0 K,

Table 1.4. Thermal Runaway Tests. Lead No. 4<

Pressure

(MPa)

Flow Rate

(g/s)

M

2 Current(A)

2J

M

Stable or Unstable

0.023 1. 98 600

700

1.55

1.80

1.0 0.047 3.87 900 1.18

0.070 5. 92 1150

1378

0.99

1.186

Conditions of Test: a.c. current, helium inlet temperature 4.5-5 K, conductor

cold end temperature 10 K.

1.4.3. Discussion of Results and Comparison with Theory

The most obvious thing about the temperature profiles that we did not

anticipate is that they give evidence of incomplete thermal anchoring at the

warm end despite the modification to the phenolic resin core referred to

above. To get a really ideal experimental setup for comparison with theoretical calculations it is apparently necessary to employ water cooling or some such remedy except at the adiabatic warm-end condition. Nevertheless, from

computer simulation we do not believe in these experiments that the lack of

thermal anchoring changed the burn-out condition significantly. The second

32
point to notice is that the thermal runaway condition only involves temperature

profiles with maxima at the lower flow rates. But this, as we shall see is

not unexpected. A third point is that in all of these tests the lead was unstable at some level of current. In computer simulations we have observed

two distinct modes of burn-out. The one is a definite instability; above a certain current no solution can be found and a plot of lead temperature vs. current has infinite slope at this current. In this case the burn-out current

occurs at the adiabatic warm end condition and it is also the condition of

minimum Carnot work. In the other case, lead temperature becomes very sensitive to current and the profile has a maximum, but a solution can always be

found though the temperature may be unacceptably high. This latter case we associate with poor heat transfer, but we have not been able to explore all

parameters sufficiently to make generalizations. For the present case, com- puter simulation indicated definite instability, but we would caution that

experimentally the two would be difficult to distinguish, especially with

large capacity leads of long time constant.

The crucial point is whether these temperature profiles and thermal

runaway conditions are predictable. In figure 1.22 we have plotted the

thermal runaway condition, equation (1.6), for the simple constant conductivity, linear resistivity case with ideal heat transfer. We plot 2J/M vs. M/2 using for k*, 4.0 W/cm K which should be applicable to lead no. 2 between 150 and 300 K. The measured RRR for this lead was 49 and leads one to

expect that burn-out will occur above the line. When we plot the bracketed

burn-out points using symbols larger than the implied uncertainty we see that

this is indeed so for most of the points. We have also plotted a single

datum point for lead no. 2 for an actual accidental burn-out at high flow

rate and we see that this is quite low, but not inconsistent with the other

data. Departures from the ideal heat transfer case are accentuated at high

flow rate according to Bejan and Cluss and it is possible to draw a smooth

curve through all data for lead no. 2 which would correspond to a heat transfer factor H somewhat less than 2000. On the same graph we have plotted the

2J/M vs. M/2 relationship for adiabatic warm end according to the constant

conductivity-linear resistivity model. This shows that, indeed as observed,

temperature maxima are only expected for lower flow rates for this m.odel.

There is no unambiguous way of plotting the burn-out data for the tapered lead no. 3 on this plot, but we have plotted no. 4. As expected, for the

low resistance ratio of no. 4 (RRR = 5.4) the data fall below the computed

line

When we made computer simulations of the burn-out conditions for lead

no. 2, however, we found that the computer program overpredicted the burn-out

for given mass flow as seen in figure 1.22. Far more noticeably, for a given

current and mass flow, the temperature profiles were too low.

We apparently did not know one or more of the lead parameters well

enough for modeling purposes for this lead as tested. Realizing that others

may find themselves confronted with the same situation, we felt that it

would be worth pursuing this point to narrow down the causes. We have since

rerun computer simulations for this lead with various parameters varied in

turn to look for a plausible explanation. These included resistance ratio,

total resistance, conductor cross-sectional area, thermal impedance to the

thermal ground at the warm end, inlet temperature and, finally, the heat

transfer factor H through variation of the product Ph. Only the latter is

capable of modifying the computed results sufficiently through a reasonable

variation. We find that about a twenty-fold reduction in this factor in

needed from the flow-independent (Nu = 4.364) value for laminar flow and we illustrate this in figure 1.23 where we have shown an experimental profile

with several computed profiles for differend Ph. Taking the closest curve

it is also apparent that some further temperature dependence of the heat

33
34
transfer factor should have been allowed, probably by incorporating flow dependence of the Nusselt number; the Reynolds number decreases up the tube due

to the temperature dependence of viscosity.

In addition, we have dismantled the lead and examined it for likely

cause. The braid was in good condition. The resistance ratio was remeasured

and found to be, in fact, a little higher (RRR = 58) than our previous mea- surement on a sample from the same spool. But the total room temperature

resistance before unsoldering the braid differed from the expected braid

resistance (2.54 milli ohm cf. 2.0 milli ohm) indicating a significant contact

resistance at the soldered junctions. Certainly, heat generation at the warm

end would cause premature thermal runaway, but considering the great disproportion between the thermal conductance to the heavy copper fittings and that

to the braid it would appear that this could only be a minor effect.

One final piece of evidence we feel makes the poor heat transfer argument

conclusive. The measured helium outlet temperature for the run in figure 1.23 was 15 K below the conductor temperature compared with about 1 K for the

original calculation. For the curve with hP = 0.174 in figure 1.23 the calculated temperature difference was much more realistic at 11 K. We thus come to a not-too-surprising conclusion: a current lead is a heat exchange device and to predict its performance accurately one would be

well advised to base his estimates on measured heat transfer coefficients. In

the case studied the approach of using a mean hydraulic diameter and therefore

assuming a uniform distribution of flow channels is simply not correct.

1.5 Oscillatory Behavior

1.5.1 General Description of Observed Oscillation

Pressure oscillations were observed in all four leads tested, no. 1 being

the most prone to oscillate, the others were in the order no. 3, no. 4 and no.

2. It is interesting to note that no. 2, having suffered a burn-out failure

during testing, was repaired for further testing. The repair involved patching the outer stainless steel shedth, replacement of the braid and slight

lengthening of the phenolic resin core. After this modification, the weak

oscillations which had been observed could no longer be repeated. By these

changes we appear to have tipped the balance from marginally unstable to

stable behavior.

The dependence of pressure oscillations on the independent parameters for

a given lead may be generalized as follows:

1. Amplitude decreased with decreasing gas flow rate going to zero at some small finite value. Oscillations always ceased when the gas flow through

the lead was shut off.

2. Amplitude decreased with increasing current and could be reduced to zero above a critical current for leads 1 and 2, but not for leads 3 and 4.

3. Amplitude decreased with gas inlet temperature, again going to zero

at a critical value depending on the pressure and gas flow rate.

4. Amplitude decreased with increasing pressure for lead no. 1. The

opposite is true however for lead no. 2.

The oscillation amplitude vs. lead current is shown in figure 1.24 for

lead no. 1 operating at the given conditions, and the dependence of the

critical current (the maximum, at which oscillations occurred) on gas flow rate

with pressure as a parameter is shown in figure 1.25.

35
x/L

Figure 1.23. Comparison of experimental steady state temperature profile with

calculations for various hP.

36
OO CD ^ CN

O

O

O

O

O

37
<+-! •H o

COC

O •H

•H s

4J CO cd 14-1 tH

T-\ (U •HO H • CO CO O

)-l ^ !^

4-1 CD cfl C • U

•H CU •H0

cn

1 &6

rH • (U 4-1 O (U 4-1

4-1 CU iH

>i 4-t n CO •H

CO U

CO cu

CO &CU

(U 0 ^

o CU bO

(U 4-1 •r(

,12 u

CO 0) 5-1 c iH O

o H-l •H •H U 4-J •H s

13 (U C •H O I—

1

O QJ CU J=

01 4-1 U cfl cS o

4-1 O

CO o 4J rH rH <4-( U-( •H

1 ,C

+J d CO C c

(1) CD

(-1 t3 CU M cfl c

3 <u •H

m

CM

CU

•H

X

in

(s/B) liVUMOId i/\inn3H

38
Additional observations of importance are:

5. Pressure oscillations could not be detected in the gas supply res- ervoir . 6. Oscillations could be eliminated by connecting a warm gas reservoir

of about 1 liter capacity at the exit of the lead.

7. The oscillation frequencies were in the range 22-55 Hz. The only

dependence found for frequency was on the inlet gas density and is illustrated

in figure 1.2 6 for lead nos. 1 and 2.

8. In the case of lead no. 1 only, a low frequency modulation of the

pressure oscillation was observed with a frequency of about 0.1 to 0.5 Hz.

This was accompanied by temperature oscillations of the same frequency which

could achieve amplitudes as high as 19 K at the cold end.

9. There was no basic difference observed between oscillatory behavior

above and below the critical pressure. In the latter case it is fairly

certain that the liquid phase did not enter the lead.

For lead nos. 1 and 2, where we explored the widest range of parameters,

the existence of oscillations in the pressure-temperature plane for the helium

entering the leads, is shown in figures 1.27 and 1.28. In the case of lead

no. 2 this applied to the lead before repair as mentioned above.

1.5.2. Discussion of Oscillations

As a practical matter one might first inquire as to the amplitude of the

pressure oscillations. Maximum peak-to-peak amplitudes observed under any of

the conditions of our tests were about 0.1 MPa at 0.5 MPa for no. 1, 0.01 MPa

at 1 MPa for no. 2, 0.016 MPa at 1 MPa for no. 3 and 0.008 MPa at 1 MPa for

no. 4. We feel that although in the latter three cases these may not be

serious perturbations one should not draw comfort from the results until some understanding of the governing parameters has been obtained, especially con- sidering the limited range of the tests for nos. 3 and 4 and the possibility

of attendant heat flow which could not be measured in these experiments.

Considering the observed frequency and the configurations of the leads it

appears that the oscillations are of the thermally driven acoustic type com- monly observed in tubes inserted into helium dewars. Here, one normally has

an open cold end and a closed warm end giving rise to a standing wave with a pressure node at the open end and an antinode at the closed end - thus a 1/4 wave. In our case there are some differences, chiefly that the warm end was

not closed and there was always a flow of gas when oscillations were observed.

Nevertheless, with a considerable flow restriction in the warm gas exit piping

where most of the pressure drop to atmospheric pressure was taken, we believe

the slow flow of gas would not affect the oscillation - except insofar as the

temperature profile is determined by the flow rates as will be discussed.

The driving mechanism of such oscillations is well understood and has

been given greatest analytical clarity by Rott [14] . According to this

accepted mechanism, when an acoustic standing wave of infinitesimal magnitude

exists in the presence of a temperature gradient, an oscillating heat flux

from the tube wall to the gas accompanies it. If the temperature gradient is

such that the gas is heated in that part of the oscillation where it is moving

towards the pressure antinode and cooled when moving away, then the oscillation can be enhanced and may grow in time to become an observable sustained

oscillation. The temperature must therefore increase from pressure node to

antinode - the conditions which existed in our lead tests. When the 1 liter

reservoir was connected to the warm gas exit piping this end was apparently

converted to a pressure node so that only the 1/2 wave standing wave could

exist and the conditions for growth did not exist over half of the lead.

39
0 LEAD No. 10.2-1.2 MPa

1 LEAD No. 2,1.0 and 1.5 MPa

0 0.05 0.10 0.15

HELIUM INLET DENSITY (g/cm^)

S. Frequency of • oscillations as a function of inlet helium densi

40
2.0

1.6

1.2

1.0

0.9

o

Q.

S 0.6

U1

UJ

Q.

0.4

0.3 -

0.2-

0.1

ations

resent

Transposed

Oscillations

Absent

Liquid

J L

5 6 7 8 9

TEMPERATURE (K)

10 11 12

Figure 1.27. p,T regions in which oscillations were observed for lead no. 1.

41
2.0

1.6

1.2

1.0

0.9 —

0.6 -

0.4 -

0.3 -

0.2-

0.1

Liquid

Transposed

Critical Line^'

Oscillations

Present

Oscillations

Absent

5 6 7 8 9

TEMPERATURE (K)

10 11 12

Figure 1.28. p,T regions in which oscillations were observed for lead no. 2,

42
We would like to be able to predict for any lead design and for various

operating conditions when oscillations can be expected. The dependence of the

existence of the oscillations on current and flow rate appears to be a complex

function of resulting temperature profiles. However, the dependence on the

helium thermodynamic state - i.e., whether oscillations can exist for any

current and flow rate for given helium states - appears to be broadly predictable from the theory and stability computations worked out by Rott for a simple model. This model is of a simple tube, with an open end in a helium

dewar at T and a closed end at room temperature T„. The temperature profile

L n is assumed to be that of a discontinuous change in temperature between two

isothermal zones. Solving the one-dimensional wave equation for the pressure

under the assumptions of continuity of pressure and pressure gradient at the

temperature discontinuity, but taking into account the two-dimensional aspects

of momentum and heat transport in the vicinity of the tube wall, he was able

to plot a critical temperature ratio for neutral stability a = T„/T_, vs.

n C

the dimensionless ratio Y of the tube radius r to the ac boundary layer thick1/2

ness, (v^/oj) , where is the kinematic viscosity in the cold zone end u the angular frequency. A parameter of the plot was the ratio ^ of warm length

to cold length, A second plot gave the frequency, but von Hoffmann [15]

et al. showed that the two could be combined into a single plot of a vs. 1/2 1/2

A = Y/(cjj£^/C^) - r {C^/Z^v^} where is the acoustic velocity in the gas

at T^. Now, the frequency is no longer explicit. We have reproduced this

combined stability plot for various E, in figure 1.29 from Rott's paper making

good use of the asymptotes of the two branches of the neutral stability curve given in the paper

The physical significance of the two branches is that to the right of and

below the lower branch there is insufficient heating and cooling during a cycle to overcome the normal dissipation of the wave; this is the wide tube

limit where, in both warm and cold parts, the tube radius is many times the ac boundary layer thickness. Above the upper branch the boundary layer in the warm part fills the entire tube and the pressure wave is heavily attenuated as

in a capillary tube.

Now if we plot on the same graph the operating ranges of the ratios a and

and A at constant pressure, interpreting the tube diameter in Rott's calculations as the mean hydraulic diameter of the lead, (4 X flow cross sectional

area/wetted perimeter), we see a remarkable correspondence with our experimental observations, particularly the somewhat contrasting behavior of nos. 1 and 2 with pressure. For, considering lead no. 1, for any reasonable value of

C (we estimate E, = 0.47) at 0.25 MPa the oscillation should be excited for

cold end temperatures up to somewhere between 8 and 12 K. The experimental

value was 10 K. As the pressure is increased this temperature should reduce

to about 5 K at 1.1 MPa where 6 K was observed. Thus we find an explanation

of the behavior of figure 1.27.

Considering lead no. 2, we see that it cannot be unstable at 0.25 MPa for

any temperature and any value of C (the curve for 5 = 1 is the furthest to the

left) as observed. But at 1.0 and 1.5 MPa oscillations are possible, and for

higher cold end temperatures at 1.5 MPa than at 1.0 MPa - again as observed

and illustrated in figure 1.28. Precise stability boundaries are of course

not predictable because of the need to interpret ? for the case of helium

cooled leads, where the change in temperature between warm and cold ends is

certainly not discontinuous although it approaches it for low currents. The

behavior noted in points no. 1 and 2 above is not predicted probably because

for both high currents and low flows the temperature profiles become less

43
I I I I I—

r

OQ<

I—

CO

tn CO CO ocNj I I I I I t I I L I I I i ' t L o

ino

CMo

CSI

ID

CM

O

C5

CM

OO

O

in o

CM

in

iH

CO

4-1 c CO 4J (U 0) e C

•H •H

>-l iH

0)

di •H

CO CD 1^

p,

O c 4-J CO o C

Q>

CO

4-1 Q) 4-* 4Jo oOh (Uu CD

M-l 1> o CO < , (U Q)

to c CO •H o iH O

•H

4J TJ 4-* rt (U (J J=

CD

O CO (J TJ l_i M

CO <u a <L) >s CD

•>

(U

H CO (.1

f-l CU 4J VJ (0

•tH

B ^ 4_l o O CO

)-i

O ^ 0)

«4-l 5: •H

CO CO

O 15

(U •H CO •H O CO )^ 4-1 (-

CO (U o

T3

u •H o cu +J 3 4= CO O CO &

43 c iH o CO •H c O

+J 4-1 o CO •H CO •H O

iH > 4-J •H u •H CD ^ 01 T3 U

CO CO (U 43 O

C/3O

U

ON

CM

CUu3

00

•H

fa

44
discontinuous and the model must break down. However, the position of the

operating curves for lead nos. 3 and 4 are certainly consistent with the

observation that for all currents oscillations were observed. By comparison

with lead no. 1 one might guess that large amplitudes would have been observed

had they been tested at lower pressures.

We see from this comparison of our results with Rett's theoretical calculations that the main design feature at our disposal to avoid oscillations

for given thermodynamic states of the helium is the mean hydraulic diameter of

the lead. We have two alternatives: to make this quantity no larger than in our lead no. 2, - a proven option - or to go in the opposite direction and use

a mean hydraulic diameter at least a factor of two larger than in our lead no.

1. The former option is certainly compatible with a high heat transfer

factor; it is not certain that this could be achieved with the latter since

H « 1/(D 'D ) where D is the mean thickness of the conductor (4A /P)

The possibility of changing the boundary conditions at the lead ends

should also be considered. As we have seen, creating a pressure node at the

warm gas exit by connecting a reservoir is effective, but perhaps not con- venient in a high voltage environment. Creating a flow restriction at the

lead inlet may result in a 1/2 wave or closed-closed standing wave. In this

case heat addition and subtraction from the wall would not be effective in

exciting oscillations in the lower half of the lead and may cancel the

positive effect in the upper half. According to Rett's results some alleviation could be achieved by reducing ^ to a small value. This would imply

placing a flow restriction immediately at the warm gas exit.

On a final note some of the limitations of Rett's calculations for our purposes should be noted. Apart from the simplicity of the model and the need

to interpret his tube diameter as the mean hydraulic diameter of the complex

lead structure, his calculations were carried out assuming zero penetration of

the temperature wave in the tube wall material, or, constant wall temperature.

This can certainly not be true in the cold part of the lead. In addition,

calculations involved ideal gas equations, but the anomalous properties in the

region of the transposed critical line lead to further uncertainties in the

applicability of the results in that vicinity.

1.6. Conclusions and Recommendations

This study has been necessarily exploratory in nature and we have only

been able to make a small beginning in several areas. Nevertheless some conclusions have already emerged.

1. From the point of view of lead installation we quickly learned that

heat exchange to the entering helium gas alone was not enough to keep the

cold end of the conductor temperature down to within a fraction of a kelvin of

the entering helium. Some extraordinary cooling surface at the cold end must

be provided in addition as was accomplished in our test by cooling the bus

bar

2. To be safe, in order to avoid burn-out, one should not exceed the

adiabatic warm end condition.

3. The actual burn-out current per unit of mass flow is a function of

conductor purity and heat transfer design. Based on our experience with lead

no. 2 for RRR > 50 a safe criterion for operation seems to be 2J/M < 1.0 if

the mass flow is close to optimum or below. For flow rates above optimum this

will only be true for leads with heat transfer factor H greater than about

4 10 . Further work is needed on leads with low RRR.

45
4. The heat transfer factor for a braided lead of the design of no. 2 is much less than can be estimated from simple heat transfer concepts.

5. Accurate predictions of thermal runaway for a given lead design will

only be possible if empirical knowledge of the heat transfer factor and its

dependence on mass flow is available in addition to the RRR. The heat

transfer factor should be measurable by less expensive room temperature ex- periments .

6. Oscillatory behavior can be expected in leads where the mean hydraulic diameter is greater than about 0.018 cm unless other precautions are

taken. These precautions include the use of a warm-end reservoir or throttling

of the helium coolant at the inlet to the lead. It is recommended that actual

designs be tested for oscillations without current flow to establish the

conditions for stable operation.

7. Future work should concentrate on specific candidate designs with

demonstrated high heat transfer factors. Yet to be established is the safe

operating current at which a lead may operate with the capability of withstanding fault currents. This should be amenable to calculation.

1.7 References

[I] Trumpler, P. R. , and Dodge, B. F. , The Design of Ribbon-Packed Exchangers

for Low Temperature Air Separation Plants. Trans. Am. Inst. Chem. Eng.

43, 75 (1947).

[2] Lock, J. M. , Optimization of Current Leads into a Cryostat. Cryogenics

9, 438 (1969)

.

[3] Pippard, A. B., Continuous Refrigeration of Current Leads. Cryogenics

5, 81 (1965) .

[4] Agsten, R. , Thermodynamic Optimization of Current Leads into Low Temperature Regions. Cryogenics 13^, 141 (1973) .

[5] Bejan, A. and Cluss, E. M. , Criterion for Burn-up Conditions in Gascooled Cryogenic Current Leads. Cryogenics ]j6, 515 (1976).

[6] Arp, V. D., Daney, D. E. , Frederick, N. V., Jones, M. C. , Ludtke, P. R.

,

Parrish, W. R. , and Powell, R. L. , Helium Research in Support of Superconducting Power Transmission. NBSIR 75-823, Nat. Bur. Stand. (U.S.),

October 1975.

[7] Thullen, P., Stecher, R. W. , and Bejan, A., Flow Instabilities in Gascooled Cryogenic Current Leads. Proceedings of the 1974 Applied Superconductivity Conf., IEEE Trans, on Magnetics MAG- 11 (lEE Magnetics

Society, 1975).

[8] Arp, V. D. , Thermodynamics of Single-Phase, One-Dimensional Fluid Flow.

Cryogenics 15^, 285 (1975).

[9] McCarty, R. D. , Thermodynamic Properties of Helium 4 from 2 to 1500 K

at Pressures to 10^ Pa. J. Phys. Chem. Ref. Data 2, No. 4, 923 (1973).

[10] White, G. K., and Woods, S. B. , Electrical and Thermal Resistivity of

Transition Elements at Low Temperatures. Phil. Trans. Roy. Soc.

(London) A251 , 273 (1959).

[II] Powell, R. L. , Roder, H. M. , and Hall, W. J., Low-Temperature Transport

Properties of Copper and its Dilute Alloys: Pure Copper, Annealed and

Drawn. Phys. Rev. 115 , No. 2, 314 (1959).

46
[12] White, G. K. , and Woods, S. B. , Phil. Mag., 7, 45^, 1343 (1954).

[13] Giisewell, D. and Haebel, E. V., Current Leads for Refrigerator-Cooled

Large Superconducting Magnets. Proc. ICEC 3, Berlin, p. 187 (I.P.C.

Science and Technology Press, Guildford, UK, 1970).

[14] Rott, N. , Damped and Thermally Driven Acoustic Oscillations in Wide and

Narrow Tubes. ZAMP 20_, 230 (1969) and Thermally Drive Acoustic Oscillations, Part II, Stability Limit for Helium. ZAMP 24^, 54 (1973).

[15] von Hoffman, T. , Lienert, V. and Quack, H. , Experiments on Thermally

Driven Gas Oscillations. Cryogenics 13^, 490 (1973) .

47
2.0. HELIUM MEASUREMENTS

An important part of this report has been concerned with the measurement

of those helium properties which monitor the performance of the cooling system. The NBS work for this year is divided into several catagories. The

general problem of data transmission from the power line to ground potential

is discussed in section 2.1. Comments on thermometry are given in section

2.2, which is a follow-up on a similar discussion in last year's report.

Possible indirect methods for determining helium pressure are discussed in

section 2.3, and work on a new form of pressure transducer is dicussed in

section 2.4. Problems of impurity detection are discussed in section 2.5.

2.1 Data Transmission from the Superconducting Power Line to Ground Instrumentation

The generalized problem of transmitting the operational information (on

pressure, temperature, flow, etc.) from the power line to the final readout

condition at ground potential was discussed in some detail at the project

review meeting in Boulder on January 27, 1976. At that meeting Dr. Rabinowitz

of EPRI informed the group of an existing EPRI-sponsored contract related to

signal transmission across a high voltage gap, and as a result it was con- cluded that our superconducting power line project should not initiate any

experimental work in this direction at this time. However, the subject is

important here in relating recommended measurement systems to the generalized

constraints imposed by power line system operation.

A schematic outline of the signal path is shown in figure 2.1, assuming

that the transducer and its electrical circuit will unavoidably ride essentially at the potential of the superconductor. Since this potential may be

several hundred kilovolts above ground, special signal processing techniques

are called for. As seen in figure 2.1, the power supplies, amplifiers, etc.,

used for signal detection are enclosed in a high voltage "cage" at the power

line potential, and special equipment is used to transmit the information

outside of the cage to recording equipment at ground potential. What limitations, problems, and costs are incurred through this signal path? It is

convenient to divide the question into several categories: the power supply,

amplifier and detector specifications, transmission to ground potential, and

60 Hz noise pickup from the power cable.

2.1.1 Power Supply

The power supply should provide instrumentation power during cooldown and

electrical faults which temporarily open the line as well as during steady

state operation. This could be achieved by using a rechargeable battery

supply, and arranging a charging circuit inductively coupled to the current in

the superconductor for recharging the battery during steady state operation.

No new technology is required, but this does suggest that d.c. powered instrumentation may be most appropriate. Conversion to a.c. is possible, but it

adds to the cost of the power supply.

2.1.2 Equipment Specifications

An important question is whether conventional electronic equipment - amplifiers, a to d converters, etc. - will operate as effectively and accu- rately within the high voltage "cage" as they do at ground potential. This

point was discussed very carefully with Dr. Oskars Petersons and several of

his staff at the NBS High Voltage Section in Gaithersburg. We had been con- cerned that miscellaneous power line transients could couple into the com- ponents and perturb sensitive circuits through small stray capacitances to

ground. Dr. Petersons, a man of considerable practical experience, assures us that this will not be the case, and that ordinary components will meet their

ordinary specifications when properly "grounded" to the cage.

48
SUPERCONDUCTOR

r

SIGNAL

DETECTION

&

AMPLIFICATION

SIGNAL

MULTIPLEXER

&

ENCODER

TRANSMITTER

1

POWER

SUPPLY

|-^The "cage'

Dashed line encloses

all equipment riding at

power line potential

RECEIVER

SIGNAL

DECODER

I I I I I

(Final Readout)

Figure 2.1. Schematic outline of the transducer signal path from high

voltage to ground potential.

49
2.1.3 Transmission to Ground

The biggest question is how to transmit the transducer information out of

the "cage" to ground potential. Here there are three basic alternatives to be

considered: (1) a capacitive link, (2) a fiber-optical link, and (3) an f.m.

radio link.

2.1.3.1 Capacitive Coupling

The use of conventional overhead high-voltage power lines for transmitting communications as well as power is standard within the industry, and

techniques developed for the communication link may be applicable to our instrumentation problems. Frequencies in the range of 30 - 300 kHz have been

employed for power line carrier purposes, as this range is high enough to be

isolated from the 60 Hz power frequency and the noise it creates, yet not so high as to encounter excessive attenuation.

The accepted technique for transmitting the communications signal between

ground and the high voltage line is to use a high voltage coupling capacitor,

for which standards have been set up by the AIEE. From a discussion on this

topic with Mr. Ray Fillenberg of the U.S. Bureau of Reclamation, the cost of

such a unit is of order $8,000, though this may include certain filters and

tuning circuits. We conclude that there should be no major technological

barrier to the use of such a device for transmitting instrumentation signals

to ground potential. Several carrier frequencies could be used simultaneously.

2.1.3.2 Fiber-optics

An alternative technique is to use fiber-optical data transmission between high voltage and ground. The basic advantage of such a link stems from

its insensitivity to electromagnetic radiation from the power line, in particular, from the transients. In the early 1960 's such a fiber-optical link

was developed and marketed, but sales were disappointing (apparently) and it

has been discontinued. Since that time, the Light Emitting Diode (LED) has

been developed. In the high voltage research activities of Hydro-Quebec in

Ottawa, LEDs and fiber-optical techniques have been combined for just such

instrumentation links as are being considered here. This work was performed

with private funding and has never been described in the open literature. A

representative of a Canadian company tells us that they do not manufacture

these as standard items, but would be willing to build one on special order.

Various pulse-code modulation and/or frequency modulation techniques could be

used. Amplitude modulation is less desirable because LEDs are quite nonlinear

and also drift with age. A description of a proposed fiber-optical link for

high-voltage systems has been given by Erez [1] . Fiber-optical links have

also been used to control the firing times on the thyristors used in the d.c.

power transmission systems between Southern California and the Bonneville

Power Administration. In this case only a very small amount of information is

transmitted (one firing pulse every sixtieth of a second) , but possible

problems of surface leakage and flash-over along the fiber-optics appear to

have been solved.

2.1.3.3 Radio Transmission

A third alternative for the instrumentation link is to use radio transmission. Such a system, operating at about 100 MHz, has been described by

Morse, et al. [2]. This reference also mentions a fiber-optical link, and it

appears that the radio and fiber-optics offer comparable transmission capabilities .

50
2 . 2 Thermometry

The sensitivity of various types of thermometers to electromagnetic

fields was summarized in our last year's report [3], In discussing the subject with the Steering Comm.ittee (on May 7, 1975) prior to writing that report,

it had been suggested that it would be desirable to place them, for example,

in the high voltage insulation space, thereby exposing them to intense a.c. fields. This places stringent requirements on the thermometers, but several

types were identified which should give temperature measurements substantially

unaffected by these ambient fields.

This subject was further discussed in the Steering Committee meeting of

January 27, 1976, this tim.e in relation to the problem of induced voltages on -he electrical leads from the therm^ometer to the power supply and sensing

device. The general conclusion at this time was that thermometers would never

be placed in such high-field regions because of the great difficulty with

induced voltages on the leads and recording equipment. In other words,

thermometer selection and sensing in SPTL's would not involve electromagnetic

conditions significantly more stringent than in many other low tem.perature

systems, though small size and fast response (in milliseconds) may be very

desirable. On this basis we have not done any further study on thermometer

system.s

At the tim.e of this writing, neither our Laboratory nor the other ERDAsupported S?TL projects have had experience with the newer carbon-in-glass or capacitance thermom.eters which are now comm.ercially available. We plan to

purchase some of these for our own experimental work in the coming year, so as to become familiar with their use. The general conclusion reported here, that thermometers will probably not

be placed in high-field regions, would apply also to other types of sensors,

e.g., pressure sensors.

2.3 Pressure Determination from Density and Temperature Measurements

The pressure in a single phase fluid can be uniquely determined from

simultaneous measurements of density and temperature. This possibility, along

with that of using other pairs of thermodynamic variables (e.g., sound

velocity and temperature) was considered briefly in our previous publication

[3] but dismissed as being not as promising as direct pressure measurement as discussed in section 2.4 of this report.

Since that time. Dr. C. T. Van Degrift, who initially stimulated our microwave cavity work, has moved from the University of California to the NBS

Heat Division, and is currently pursuing this approach though inverting the

procedure to obtain a secondary temperature standard from density and pressure

r.easurements using a cavity. It is appropriate to review this question in

r.cre detail than was given in reference [3] .

2r . Van Degrift 's secondary temperature standard is (will be) obtained by

sealing a fixed quantity of heli^om. into a high pressure chamber which has one wall in comm.on with the microwave cavity, which operates in a vacuum. The

heliiim pressure is determined by the cavity frequency measurement. The helium

density is essentially constant, and is determined by calibration at a known

tem.perature . Once the density is known, other temperatures can be determined

just from cavity frequency measurement, knowing the equation of state for

helium.

In order to invert his procedure, and determine a system pressure from

density and temperature, one can irtake use of the fact that the dielectric

constant is a simple monatonic function of density and that the resonant

frequency will vary with dielectric constant if the helium permeates the

51
cavity. In this case the cavity is made with rigid walls so that the frequency does not change with pressure. To put this in quantitative terms, the

dielectric constant is related to the density of helium by

e w 1 + 4 X 10~^ p. [SI Units] t2.1)

For a cavity having rigid walls, a change in pressure will affect only the

fluid density within the cavity, and the resonant frequency, f, will be

-1/2 proportional to e , or ^ = - ^ . (2 2)

f 2 £

•

In first order, e = 1, so that

1^ « - 2 X

10""^ dp . (2.3)

To relate a small density change, dp, to small pressure and temperature

changes, dP and dT, we write

, ^, dP ^ dT /-^ A\ dp = pK p- + pa

^ , (2.4)

where .

Both K and a are equal to 1.0 for an ideal gas, but more realistic values for

helium in the compressed liquid range are

K « 0.1 and a ^ 0.5 .

2 3 Therefore, taking p w 10 Kg/m ,

• |i « -.002 ^ -0.01 ^ .

Thus, to measure pressure to an accuracy of one percent, one must determine the resonant frequency to an absolute accuracy of about one part in

50,000. This is not hard to do in the laboratory, where a good counter can be

used, but it would be difficult to do inexpensively, as, for example, with the

voltage controlled oscillator circuit described in section 2.4 of this report.

Further, the temperature would have to be measured, in this example, to an accuracy of 0.2 percent, or about 0.02 K, to obtain one percent pressure

accuracy. Because temperature effects are so large, the readings of frequency

and temperature would always have to be processed through the equation of

state, in a minicomputer, to obtain the pressure. Finally, the method is not

conveniently adaptable to differential pressure measurement, which would

always require two separate microwave cavities, perhaps beating one against

the other.

We still conclude that this would not be as practical for a field device

as are the microwave cavities discussed in section 2.4 of this report.

2.4 Microwave Cavity Pressure Transducer

In simple concept, the microwave cavity pressure transducer is a sharplytuned L-C circuit whose resonant frequency shifts with pressure. This is

accomplished by building the circuit as a reentrant microwave cavity having

one thin wall which deflects elastically when pressure is applied external to

52
the cavity. The cavity is so designed that the wall deflection causes an internal capacitance change and thus a shift in the natural resonant frequency, nominally about 1200 MHz, which can be detected by external circuitry.

Potential advantages of this design are as follows:

(1) By building the cavity of a metal which is highly conducting at helium

temperature, e.g., copper, or a superconductor, the resonant Q can be 4000 or more. This permits a very weak coupling between the resonant circuit and the

measuring circuit, with corresponding insensitivity of the resonance to ex- ternal parameters.

(2) The cavity can be made of all one metal, thus obtaining a rigid and

rugged structure and avoiding differential thermal contraction and pressure

seal problems associated with composite metal and insulating structures.

(3) Its output is essentially unaffected by external electromagnetic fields

at power-line frequencies.

(4) It can be used as a differential pressure gauge by allowing helium within

the resonant cavity itself. No loss in electrical performance results, but

the volume of contained fluid does limit the sensitivity for high frequency

pressure oscillations. Also, there is a significant zero shift as the absolute helium pressure within the cavity changes.

(5) It can be designed to be undamaged by large overpressures, though again,

this is obtained at the cost of reduced hydraulic frequency response.

2.4.1 General Circuit Concept

The use of microwave cavities of this type as transducers was initially

suggested by Dr. Craig Van Degrift [4, and private communication].

In Dr. Van Degrift 's work the active element which maintains the oscillation and provides an output signal is a tunnel diode which is physically

packaged with the cavity. The active circuit is supplied with d.c. power, and

the circuit output is monitored with a frequency counter. When operated at

low temperature, (typically below 30° K) this system can be made stable to

g one part in 10 or better, and can be applied to a variety of instrumentational

problems. The difficulty with this system, for our application, is that noise

on the d.c. power input shows up as frequency noise on the output, so that

these leads would require very careful shielding.

In our work we have separated the passive L-C circuit (the cavity) from

the active circuit which powers the oscillation and provides the analog output. We have chosen to make the active circuit a simple voltage-controlled

oscillator, which is much less expensive, but also less precise, than the

frequency counter required in Van Degrift 's version. The connection between

these 1.2 GHz components is made via a long coaxial lead. Low frequency

noise, from d.c. to a few kilohertz, induced on this lead cannot perturb the

analog output, so this system is to be preferred when the cavity and/or lead

must be placed in an electrically noisy location. The disadvantage of this

circuit is that the output signal can drift slightly in response to perturbations on the electrical characteristics of the coaxial cable if a precise

electronic counter is used as the output device. However, a simple inexpensive readout device can provide adequate sensitivity if the diaphragm is

designed for sufficient movement (w 1 micrometer) , and perturbations from the

coaxial lead should be insignificant in this case. 2.4.2 Material

In order to obtain a sharp resonance (high Q) in the electrical circuit,

and hence high resolution for small frequency changes- the cavity must be a

53
good electrical conductor. It is important also, that all of the cavity,

pressure chamber, etc., be made of the same material so that differential

thermal contractions on cooldown to 4 K will not lead to residual strains or distortions which could compromise the accuracy or calibration.

Our cavities are made of copper, chosen because it is a good electrical

conductor at room temperature, becoming much better at 4 K. Thus the transducers should give a good and usable signal at room temperature and during

cooldown, while best performance is obtained in the liquid helium temperature

range. If only cold operation is desired, even better performance could be

obtained by making them of niobium, which is superconducting below 9 K, or, as suggested by Zimmerman [5], tin-reacted niobium which is superconducting below

18 K. One problem with copper cavities must be recognized. The stress-strain

behavior of copper is strongly dependent on metallurgical condition, and

linear elastic behavior, in particular, is obtained only for very small strains.

This is discussed in more detail in section 2.4.4 of this report. Some sort

of compromise may be worth consideration for future work, by selecting a material with lower conductivity but better elastic behavior.

2.4.3 Cavity Geometry

Two different geometries have been used for the design of the cavity.

The first consisted essentially of a thin-walled pipe which was soldered into

an axial hole in a specially machined piece, as illustrated in figure 2.2

(this is the same as figure 4.3 in our report last year [3]). While this type

of cavity can be made to work, we concluded that delicate problems of machining and assembly of this type of cavity are significantly greater than that

of the second type, illustrated in figures 2.3 and 2.4. The frequency, sensitivity, nonlinearity of response, etc., of this second type of cavity, can be

calculated in a relatively straightforward manner. Regarding the cavity as a simple LC resonant circuit, neglecting secondorder contributions because of the finite cap thickness, and assuming a dielectric constant e, the capacitance (referring to figure 2.3) is

C = 15 X 10

-12 TT d X -J- — e 4 g

and the inductance is

(2.6)

L = 2 X

10~'^£ log (-) . where SI units are used throughout. Thus the resonant frequency, f^ cavity is related to the cavity dimensions by

2 1 4 q (2TTf ) LC

TT X 17.7 X lO"-"-^ e d^£ log(-)

(2.7)

of the

(2.8)

The diameter of the capacitor cap (8.38 mm in figure 2.5) is a compromise

between precise machining requirements and the goal of reducing the external

cavity dimensions while retaining high sensitivity. A large cap diameter

leads to a large capacitance between it and the closely spaced flexible wall

which it faces. Thus for a given operating frequency, limited by availability

of practical electrical components, a large cap allows a relatively small

inductance, or cavity volume, to be used. On the other hand, with ordinary

machining and assembly tolerances, the surface of this cap will not be exactly

parallel to the wall which it faces, only about 0.06 mm away. One can see

54
3.81 cm I.D

He! ium

Coupling Loop to

Electrical Circuit

1 .906 cm O.D.

Sol der

Joint 5.08 cm O.D.

Gap .0076 cm + .0004 cm ^ ^ Wall Thickness .0127 + .0008 cm

All dimensions in centimeters

Figure 2.2. 440 >fflz resonant cavity pressure sensor.

55
Figure 2.3. Geometrical parameters of the microwave cavity.

56
Flexible diaphragm

Capacitance probe

Cavity pump vent

/

high pressure

^ .06mm gap

Resonant cavity

Mechanical shield

-XL Coaxial sleeves

/ \\ provide a slip fit J K electrical connection

low pressure

Coaxial electrical lead

Figure 2.4. Assembly of the microwave cavity pressure transducer.

57
(a)

2.18 drill thru

1.5 mm drill,

1.0 mm drill

last 1.5 mm

^1111.91Q1_^\

15.62

r

Thickness determined by

desired sensitivity typically 0.4 mm

CN

CO

O

NOTE:

Because diaphragm is very thin, inside

bore is machined first, a close fitting

^brass plug is soldered in bore to prevem

flexing of diaphragm as external cuts arc made. An air vent hole must be drilled

in plug for easier removal.

11.91

(b)

All dimensions in mm

Figure 2.5. Design detail on the two pieces which are soldered together

to form the cavity.

58
that a slight misorientation between the center lines of the two pieces in

figure 2.5, will result in a closer than average spacing between some point on

the cap rim and the opposite wall. Since the total capacitance is the integrated average of the reciprocal of the gap spacing, a relatively close

spacing at the rim will have a dominant affect on the resonant frequency,

degrading the sensitivity to wall movement if the cap diameter is too large.

Our experience suggests that if the gap, g, is much less than 0.05 to

0.07 mm, one encounters annoying difficulties with machining inaccuracies,

precise control of f^ during assembly, cleaning of the cavity after assembly,

etc. These difficulties can all be overcome, but the construction costs

will rise in doing so. O'ur prototype cavities, to be installed for field

tests in our helium flow loop, have been designed with a gap of about 0.06 mm. The gap spacing is controlled during assembly by a sliding motion of part b)

inside part a) (of figure 2.5) , observing the resonance with a temporarily

installed probe connected to the electrical circuit. When the desired

resonant frequency is obtained, the probe is removed, and the pretinned parts

are heated to fix the final spacing. After this operation, the cavity is

cleaned by forced flow of filtered ethyl alcohol, inlet through the 3.81 mm

probe port.

2 The i R losses in the microwave currents broaden the resonance and reduce

the ultimate sensitivity of the device. Since the microwave currents flow in

a very thin layer over the metal surfaces, the surface finish should be kept

as fine and clean as possible. It is also possible to minimize losses by

geometric considerations, picking the cavity proportions to minimize the total

resistance through which the currents must flow. Neglecting current flow in

the facing surfaces of the capacitor plates (the current flow there is non- xmiform as a given charge density builds up or decays), the resistance will be

proportional to

Resistance 'j. — ^ ^ ~ 2^ ^ ^ . (2.9)

" a D z r a/

Num.erical experimentation with this equation (2.9) indicates that, for a given

g and f^, the losses are in fact a fairly weak function of geometry for rea- sonable variations of the parameters from the values shown in figure 2.5,

Increasing the pillar diameter, a, would reduce the losses som.ewhat, but at

the same time would result in a larger cavity, for a given frequency and

capacitor gap.

2.4.4 Diaphragm Thickness and Transducer Sensitivity

The pressure sensitivity is controlled by the diaphragm thickness, t. An

enlarged, out-of-scale sketch for calculation purposes is shown in figure 2.6.

At a certain high pressure, P^, the diaphragm contacts the opposite surface.

As the result of this additional support, subsequently higher pressures will

not linearly increase the maximum strain which the diaphragm, experiences. If

the design places this maximum strain within the elastic limit of the diaphragm

material, overpressures, unless they are extreme, should not damage the subsequent performance within its design range. For a given pressure, P, the

deflection of the center of the diaphragm is

z - 3^^-'^^ feV b ^ ^ " 256 \t ° E (2.10)

and the maximum stress in the diaphragm, which occurs at its circumference,

is

59
Figure 2.6. Schematic outline of the diaphragm distortion under pressure,

defining relevant parameters (not to scale).

60
(2.11)

In equation (2.10) v is Poisson's ratio and E is the elastic modulus.

If copper is stretched beyond its elastic limit, but not more than its

tensile stress limit, it becomes work-hardened and will subsequently exhibit

elastic behavior for stresses below the maximum value to which it has been

subjected. In practice, the stress should be cycled between zero and the peak

value several times in order to exhibit good linearity in the stress-strain

curve over a major portion of this range. Also, depending on several metallurgical parameters, it probably should not be subjected to significant

stresses in the opposite direction (leading to a phenomenon known as the

Bauschinger effect [6]). The elastic limit of OHFC copper is a very strong

function of the material condition, ranging f-rom about 32 x lo'^ Pa (46 ksi)

for material which has been cold draw down to 60%, down to 0.14 x 10"^ Pa in

the annealed condition. Inspection of data for a number of OHFC and ETP

coppers [7] shows that a reasonable tensile stress to which copper can be

stretched without exceeding the tensile strength at room temperature is in

the range 18 to 20 x lo"^ Pa. For our transducer design, it is realistic to

7 adopt a conservative limit, namely a _ = 15 x 10 Pa. Thus, after construction, the transducer should be pressurized to reach this stress and then

used only over, say, 50 percent of this range. In this way elastic, hysteresis-free performance should be obtained over its working range, provided

that the diaphragm is not pressured in the opposite direction.

If we ask that this maximum stress, 15 x lo"^ Pa, occur just when the

center deflection, Z, equals the initial gap, g, the thickness of the diaphragm must be

t = (2.12)

For our cavity dimensions, and with E = 13 x 10 Pa (copper at room temperature) , one finds

t = 0 . 15 mm

corresponding to

P = 1.3 x 10^ Pa (1.3 atmospheres) max ^

Thus, using this technique, our differential pressure transducers with a nominal working range of ^ 0.7 x 10 Pa should be undamaged by significant

overpressures. For transducers of higher working range, a would be reached

before the diaphragm touched the opposite wall, so that this overpressure

protection would not be available. This conclusion differs from that of

Neubert, in his figure 3.22 [8], where it is stated that the response of a pressure diaphragm will be linear if the displacement z is less than 0.5 of

the thickness, t: this latter can be true only if the ratio b/t exceeds a certain critical value, e.g., about 120 for copper.

The ability to sustain large overpressures without stressing the diaphragm beyond the design limit could be obtained by reversing the direction of

diaphragm motion, i.e., by letting high pressure fill the cavity, and the low.

61
or reference, pressure be external to

spaced mechanical stop could be placed

deflect beyond a preset range. The fr

adjusted independently of the gap spac

characteristics of the cavity. The di

the relatively large volume of fluid i

ical response to high frequency fluid

built any transducers on this design,

forward to do so.

it. Thus an external, very closely

to support the diaphragm if it tries to

ee travel of the diaphragm could then be

ing, g, which controls the electrical

sadvantage of this procedure comes from

n the cavity, which reduces the mechanpressure oscillations. We have not

though it would be relatively straightWhen the pressure is applied, the diaphragm distorts as sketched in

figure 2.6 (not to scale), and the capacitance change is a function of Z , g, d, and b. To a very good approximation the capacitance is

C = d/2 27Trdr ~^

(2.13)

where

6 = g - Z(l - 2p^ + p"*)

and

p = r/(b/2) ,

Combining terms.

2pdp

o 1 - w (1 - 2p^ + p^)

(2.14)

where

and

B = d/b w E Z/g

and is the capacitance when P = 0. Straightforward integration produces

the result

C = 26^/wlog r i - (1 -

Ll + (1 -

) /w 1 + /w

B ) /w 1 - /w.

(2.15)

Since w is proportional to pressure, and f is proportional to /1/C , this

expression may be used to obtain the normalized frequency shift as a function

of normalized pressure for any given diameter, 3. Writing, to second order

f- = (^f 9^;^ ^ (1 + =2") (2.16)

the calculated initial sensitivity, (1/f) (8f/9w), as a function of 3, is given

in figure 2.7. One can see that the sensitivity decreases as the diameter

ratio increases, since in this case the electrode spacing approaches a con-'

stant value, independent of pressure, near its outer rim. Also shown is the

second-order term, C2, which does not vary greatly with diameter ratio. Based

upon the yield stress limit and the cavity dimensions described above, a diaphragm designed for a maximum pressure of 10 Pa (1 atmosphere) will have a maximum normalized pressure. w s;0.2; for a maximum pressure of 20 x 10' m

^ Pa, w w 0.045.

m

From the magnitude of C2 calculated above, one can see that the

calibration will be measurably nonlinear over the allowable range of the more

62
63
sensitive gauges. In practice, the nonlinearity must be measured, since it

will vary from this theoretical value if the capacitor plates are not exactly

parallel. Also, the frequency-to-voltage circuits exhibit some nonlinearity

which adds to this mechanical nonlinearity in the final readout.

2.4.5 Electrical Circuit

Our original design utilized inductive coupling to the H field within the

cavity (as seen in figure 2.2), but we now use capacitive coupling to the E

field, a technique which appears to offer a practical advantage. With inductive coupling, it is necessary to solder the center conductor of the

coaxial pair to the inside wall of the cavity at some point so as to make a small loop perpendicular to the microwave H field. Once this operation is

done, it cannot be removed and replaced without complete disassembly of the

cavity. Further, in situ bending of the loop, to obtain a final adjustment on the coupling between the cavity and the outside circuit, will necessarily

leave small stresses in the wire. These stresses react against the insulation

holding the conductor centered within the outer conductor of the coaxial pair,

and differential contraction during subsequent cooldown may cause erratic

temperature-dependent motions of the coupling loop.

Capacitive coupling, on the other hand, involves a simple termination of

the coaxial pair in a region of the cavity where the E field is high. No

mechanical forces are applied to the center conductor other than those of the

insulation which keeps it centered. Provided that axial stresses are not

transmitted from the external lead to the short length of coax which serves as the coupling probe, there should be no tendency for relative movement between

the probe and the cavity as the temperature is cycled. (It should be noted

that the inductive probe would likewise be sensitive to axial forces along the

center conductor)

.

Our capacitive probe design is illustrated in figure 2.8, and the as- sembly of the probe in the transducer is seen in figure 2.4. The epoxy seal

at the end of the probe maintains a fairly rigid alignment between the inner

and outer conductors when cooled to helium temperature. The regular insulation is marginal for this purpose, since it contracts much more than the

outer conductor does, leaving a tiny radial gap at the inner diameter of the

outer conductor, in principle allowing slight radial movement of the inner

conductor. The conceptual design of this epoxy seal is similar to that of

standard "Housekeeper" copper to glass seals. Adjustment of the coupling to

the cavity is accomplished by axial movement of the probe assembly, which

varies the capacitance between the center conductor and the opposite wall.

This adjustment is made by monitoring the cavity resonance with an oscilloscope during assembly; we have not made any direct measurement of the spacing,

but approximate measurements suggest that it is about 0.1 mm. A schematic outline of the electrical circuit is shown in figure 2.9.

The cavity is connected to the power supply and readout instrumentation by a single coaxial lead. The power supply contains a commercially available

voltage controlled oscillator (VCO) whose output frequency as a function of

applied voltage is given in figure 2.10. A feedback circuit senses the

reflected microwave power from the cavity, and maintains the VCO input voltage

at the value corresponding to the center of the cavity resonance. The readout

circuit subtracts a controlled d.c. voltage (the zero set voltage) from the

input voltage to the VCO, multiplies the result by either 1. or 10. at the

operator's option, and transmits this to a digital voltmeter or other readout

64
a.»^

Q. o .2 E CO a

C/5 E ce E

t/i E CD E

C/3 CO CNJ o

ID

O

O

u3 o a -aeou E w E

CO o c

CD o

(=

U

•H>u

<uo

(UOU

CU

<u u

•H

(J

cdacdo

0) U

<4-l oc

bO

•Hm

<u Q

00

CNl

(U

>-l 3

W)

•H

65
p

•Hau

•Ho

iHo

•H

>-l 4JO

00

•Hao

CM

0)

t-l

66
VCO (V)

Figure 2.10. Output frequency of the voltage controlled oscillator as a function of applied voltage.

67
instrument with a range of + 10 volts. With the signal strengths available,

the circuits will lock on to the center of the resonance within 0.01 of

the linewidth. For a typical cavity Q 4000, a frequency shift of one part

in 4 X 10^ can be thus resolved. From the VCO calibration curve in figure

2.10, and operating at 1200 MHz, this corresponds to about + 0.001 volts at

the output, on the xlO range.

During the cavity assembly and setup, the lock-on circuit is disabled,

and the VCO is swept over a wider range, at the operator's selection-, so that

the resonance can be observed on an oscilloscope. When switched from set-up

to run mode, the circuit automatically hunts for the resonant signal.

The basic circuit for producing an analog voltage output, determined by

cavity frequency, has performed well in tests, but the peripheral circuits

used to isolate the cavity resonance from unwanted reflected signals have gone

through extensive revision.

Idealized curves of reflected power (measured by a diode detector) as a function of frequency are seen in figure 2.11(a). When the VCO frequency, f,

differs significantly from the cavity resonant frequency, f^, all of the

incident power is reflected. As f approaches f^, the cavity begins to absorb

an amount of microwave power which depends on the coupling strength between

the coax line and the cavity, and on the Q of the cavity. The coupling

strength is controlled by the axial position of the probe, figure 2.8, in the

cavity, figure 2.4. Assembly of the transducer involves setting the coupling

so that the cavity is almost perfectly matched at 4 K, where the Q is the

highest. Some trial and error is involved, since the adjustment requires a soldering operation on the warm transducer. Once this is set, the resonant

curve with the copper cavity at room temperature will be somewhat like the

"weak coupling" curve in figure 2.11(a).

In practice, however, the reflected power shows variations in addition to

the cavity resonance, as illustrated in figures 2.11(b) and 2.11(c). These

variations originate from the coherent addition of microwave power reflected

by non-ideal components throughout all the 1200 MHz circuit. Of these, the

most difficult for our purpose originate from signals reflected at or near the

transducer mixing with signals reflected at or near the directional coupler.

This latter mixing depends on the length of coax line, which is 0.5 meter

for figure 2.11(b) and 2 meters for 2.11(c). Experiments with 250 meters of

superconducting coaxial cable disclosed so much of this unwanted mixing that

the electrical circuit was unable to lock on to the resonance.

In the original tests a discriminator circuit was added to prevent the

lock-on circuit from operating when the magnitude of the reflected power was more than a preset value. This works well for short cable lengths, even with

the cavity at room temperature, where the discriminator circuit could be set

as the dotted line in figure 2.11(b) indicates. However, as the coax line

length was increased the increasing number and amplitude of these unwanted

reflections tended to obscure the signal from the resonant cavity at room

temperature. The further addition of a sweep limiting circuit led to a cumbersome set of controls, compromised the flexibility of the original design, and

displayed annoying transient effects. Consequently these have been recently

replaced by a simpler circuit which precisely limits the search and lock-on

circuit to operate only between two preset frequencies. Schematics for the

complete circuit are given in figures 2.12 and 2.13. This new circuit has not

been thoroughly tested as of this writing.

2.4.6. Transducer Tests

During this development, we assembled three different transducers and

subjected them to various tests, and obtained parts for two more. Basic

68
FREQUENCY

Figure 2.11. Approximate representation of the reflected electromagnetic

power sensed at the diode detector as a function of frequency.

The top curve represents the ideal case, for three different

values of the coupling strength. The middle curve represents

the curve obtained in practice with a short length of coaxial

line between the cavity and the directional coupler. The

bottom curve portrays the additional perturbations as the

coaxial line length is increased. The dashed line represents

the discriminator voltage cutoff as discussed in the text.

69
OLD

O

70
00(§)->

COCCQ <o

0) XiuC

COCo

•H

4-i O

(Uccoo

a.3

COu • > -H O 3

a. cju

4-1

4-1 60 O Cfl 3 3 O O4-1 U 3

•H

O

4-1 e o

<U rH

CJ

C

cc

cuu3

(30 ou

71
characterization of these is given in

earlier microwave cavities which were

function as pressure transducers.

table 2.1. Not included are several

only for circuit tests, and did not

Table 2,1. Transducers.

number cavity dimensions (mm) diaphragm coupling probe

length diameter thickness construction

(mm)

l.A

l.B

15.24

(cavity #1.A,

9.144

reworked)

0.25

0.13

soldered

soldered

machined and

assembled

machined and

assembled

2 10.4 0 11.91 0.14 soldered sealed coax

3 10.40 11.91 0.46 integral sealed coax

4 10.40 11.91 0.20 integral (open coax)

5 10.40 12.19 - 1.02 integral (open coax)

Transducer l.A was relatively long and thin, with a diaphragm which was

soldered in place. The solder fillet on the inside of the cavity, combined

with a relatively large ratio of d/b = 0.90 (figure 2.3), resulted in some tight radial clearances necessitating some careful polishing and fitting

during assembly. The coupling probe was assembled from several machined

washers and small pieces. The performance agreed with design predictions,

with a sensitivity of 4.3 x 10 ^ volts/Pa at 76°K, and reproducibility of

about + 0.005 volts after some initial pressure cycles. Unfortunately, it

developed a leak at a solder joint during testing and calibration in a water

bath, and it ceased resonating. Attempts to dry it out and restore resonance

did not work, though, from what we learned from subsequent tests, it is

probable that we did bake it long enough.

This transducer was then disassembled and thoroughly inspected. No

obvious problems were spotted, though discoloration on the cavity surface

suggested that improved cleaning after the final solder operation may be

helpful in maintaining a high Q.

Reassembly was made using a 0.13 mm diaphragm, about one half the

thickness of the previous one, to obtain more sensitivity for differential

pressure measurements. However, the reassembled cavity never did work well,

being quite sensitive to mechanical shock and to thermal gradients. Using

careful polishing and cleaning techniques, the Q of this cavity was about 2900

at 76°, about twice the previous value. (Screening and preliminary calibration tests were often made in a liquid nitrogen bath. The electronic

circuits used at that time often would not lock on to the cavity resonance at room temperature, where the signal was relatively low. Liquid helium was used

only after satisfactory performance in other ranges was obtained) . After

several attempts at calibration, tests of vibration, stress and thermal

sensitivity, we concluded that the diaphragm must have been distorted slightly

during soldering and that an additional support structure was necessary in

order to eliminate cavity distortion by mechanical pressure from the coax and

pressure lines.

Transducer number two was designed to be a little shorter and bigger

diameter, with a different probe construction, and improved mechanical isolation of the cavity from external forces. The different geometry should

72
theoretically reduce the losses (increase the Q) slightly, and the coaxial

probe (figure 2.8) was considerably simpler to build than was the old version.

The 0.14 mm diaphragm was soldered on, as before. Though the performance as a transducer was less than could be desired, it was subjected to a number of

tests

In spite of careful cleaning, the measured Q's were 810 at room temperature, 1400 at 76°K, and 3100 at 4°K. The measured voltage shift at zero AP

due to the dielectric constant of helium in the cavity was 4.7 x lO"'^ V/Pa at

76 °K, which is equal to the value calculated from the physical properties of

helium, within measurement errors. Due to a cold leak across the solder which

holds the diaphragm in place, the cavity was disassembled, cleaned and assem- bled again. After this operation, the measured Q rose to 2420 at 76°K, a more

acceptable value. It is evident that, within a factor of two, the Q's are influenced by rather subtle factors.

Calibration data taken at 76 °K (figure 2.14) shows good linearity over most of the range, but some hysteresis showed up as the AP approached zero.

In the first helium test with this transducer, little data were obtained due

to an external leak which slowly plugged one of the pressure lines with frozen

air; the few points which were obtained suggested a sensitivity of approxi-

-4

mately 0.53 x 10 V/Pa, though the validity of this conclusion could be ques-

-4

tioned. However, a later, more thorough test gave a sensitivity 0.787 x 10

V/Pa, with a disappointing amount of hysteresis present (figure 2.15). Two

factors from these tests suggest a fundamental problem with this transducer:

(1) the sensitivity should be slightly less at 4°K than at 76°K since the

elastic modulus of copper increases as the temperature drops, (2) more importantly, the sensitivity is an order of magnitude less than one would calculate for the ideal case. These two factors suggest that the capacitance gap

within the cavity is too large, and perhaps non-uniform, and that the thin

soldered-in-place diaphragm is afflicted with some sort of localized temperature dependent strain. Rather than attempting improvements in the technique

of soldering the diaphragm in place, it was decided to -ake a more definite

step and try machining the cavity and diaphragm out of one solid piece, as in

figure 2.5(b)

Cavity number three was constructed essentially identical to cavity

number two, except for the thicker diaphragm which was integral with the

external wall of the resonant chamber. This is machined by first turning the

inside dimensions (11.91 mm diameter by 20.32 mm deep, in figure 2.5(b)) from

bar stock at least 30 mm long. After obtaining a good 'inside surface, a close

fitting plug is soldered into this chamber to prevent flexing of the diaphragm

as the opposite end is cut off at the desired diaphragm thickness. An alternative technique, suggested by E. Ballinger, is to fill the chamber with

beeswax which provides adequate support as long as the piece is not heated by

the subsequent machine operation.

Tests on this transducer provided considerable encouragement as to its

basic design, and also allowed exploration of circuit problems which previously were smaller perturbations to the output. The circuit problems re- sulted in random background drifts in the output voltage of, roughly, .020 to

.030 volts, and random jumps of roughly the same magnitude associated with

making and breaking the coax circuit and with initiating the lock-on circuit.

These problems, which are discussed in subsequent paragraphs, made it difficult to unambiguously test for mechanical performance to an accuracy greater

than about + .02 volts. However, on several occasions the observed background

drift was opposite in direction to that expected from the effect of mechanical

hysteresis. By making the assumption that this observed drift rate was constant during a calibration run, and subtracting it from the recorded data,

we found calibrations at room temperature, 76 K, and 4 K were free from hysteresis and had standard deviations from a smooth (computer) fit of + .002 to

.003 volt, see figure 2.16. The observed sensitivities were

73
0 1 2 3 4x10

PRESSURE (Pa)

Figure 2.14. Calibration of cavity #2 at 76 K, showing some residual hysteresis.

74
5

1, 2 3 3.7x10

PRESSURE (Pa)

Figure 2.15. Calibration of cavity ^-2 at helium temperature, showing substantial hysteresis.

75
AP (Pa)

Figure 2.16. Uncorrected calibration data for transducer #3, for three different

temperatures. The apparent irreversibility is opposite in direction

to that expected from hysteresis, and is interpreted as drift in the

electronics. The discontinuity in the 76 K data was observed during

an experimental interruption.
2.085 X 10 V/Pa at room temperature (2 5 C)

1.707 X 10"^ V/Pa at 76 K

1.679 X 10~^ V/Pa at 4 K

which are consistent with the known increase in the elastic modulus of copper

with decreasing temperature [9]. These will be confirmed when better circuit

stability is obtained.

2.4.7 The Coaxial Connector

Evidence accumulated in various tests indicates that the connection of

the coaxial conductor to the transducer is a critical design element. When

the coax is cold, there is a tiny annular gap between the inside of the outer

conductor and the outside of the dielectric spacer. This gap must be sealed

to prevent residual vacuum (or pressure) leakage to the low pressure side of

the transducer. At the same time mechanical forces from bending or differential thermal contractions of the coax must not be transmitted to the center

conductor which terminates in the capacitance gap. Finally, the connector

which must be used to eliminate the above mechanical problems should not

reflect a significant amount of electrical power at 1200 MHz, i.e., it should

match the 50 ohm characteristic impedance of the coax. In one case, cooldown

stresses pushed so hard on the center conductor as to cause an irreversible

shift to a severely overcoupled condition. After some trials, the connector

in figure 2.4 was found to be satisfactory most of the time when installed in

the line external to the transducer. However, they did not always stand up

well under thermal cycling, and their construction was a slow hand operation.

More recently a commercially available hermetic coaxial seal was found which

is undamaged by fast immersion in liquid helium. It apparently uses a weakly

magnetic material as the center pin, but we do not believe this will result in

a measurable sensitivity of the transducer resonance to external magnetic

fields

Even using this commercial connector in place of the sealed coax connector, we occasionly encounter strange reflections which mix with the cavity

signal and confuse the detector circuit. Also, successive making and breaking

of the coax connections cause random variations of about + .02 V at the analog

output. These effects are still under study.

2.4.8. Summary and Conclusions

Our conclusion at this time is that the basic transducer design is working

well. We have demonstrated hysteresis-free performance, a precision in agreement with theory, and the sensitivity expected from a high-Q resonant circuit.

Some details deserve further study: performance with a simplified (less

expensive) capacitance probe assembly, reduction of random electrical reflections from coaxial line connections, and reductions of background drifts in

the electronics. Because of their insensitivity to electromagnetic noise at

power line frequencies, they should be particularly useful for SPTL studies.

Two prototype transducers are being installed in an NBS flow loop for field

evaluation, and one or two may be delivered to an active SPTL project for furthru field tests.

Potential applications go beyond the needs of the sponsor, and NBS will

be putting some funds into applying this technology for venturi flowmetering

of cryogenic fluids in the coming year.

2.5. Helium Impurity Study

2.5.1. Assitance to Brookhaven National Laboratory

BNL is evaluating * screw compressor for use with their helium refrigeration system. Since this compressor is oil lubricated, relying on the presence of oil for sealing the impeller, the oil concentration in the helium

77
discharge must afterwards be reduced to a level acceptable for the following

low temperature stages of the refrigerator. Thus in order to evaluate the

performance of the oil removal system the impurity levels must be measured.

We were asked to work with BNL in this task and have been doing so since

November 1975.

At BNL a gas chromatograph is used to analyze for all helium impurities,

including compressor oil. However, it would be extremely difficult to prepare

calibration mixtures of oil in helium with standard techniques, eg.,. preparation by weight, because the saturated concentration level is in the parts

per billion range. Therefore, we built a calibration apparatus which provides

a reproducible oil-helium mixture. The apparatus, shown in figure 2.17,

contains a windowed equilibrium cell where the recirculating helium at com- pressor discharge pressure bubbles through the oil and passes through a knockout pot; this pot serves two purposes. Firstly, the volume of the pot is

large enough for the recirculating gas to have a residence time of nearly one

hour to remove any entrained oil. Secondly, with the knockout pot and recirculation pump isolated, the large volume provides enough gas for many analyses

of a constant composition mixture. All of the apparatus excluding the equilibrium cell is at ambient temperature. The equilibrium cell is submerged in a water bath which is below room temperature to minimize possible entrainment

problems

.

The apparatus was initially tested at NHS using n-butane instead of

compressor oil. These tests showed that the apparatus provides a saturated

mixture of butane in helium if the gas is recirculated for 1-1/2 to 2 hours.

We found no evidence of entrainment as long as the equilibrium cell was kept

below room temperature.

No attempt was made to check out the apparatus with compressor oil before

shipping it to BNL since we do not yet have instrumentation for detecting ppb

level impurities. Also BNL was in the process of changing grades of com- pressor oil.

It is not possible to determine the exact concentration of oil in the

helium because the oil is such a complex mixture. However, the gas mixtures

produced by the calibration apparatus can be used as a guide to the filtration

system's effectiveness; this is because the best any filtration system can do

is to reduce the oil level to saturation conditions.

The apparatus was delivered to BNL in June, 1976. Compressor oil which

had been outgassed by heating in a vacuum for three or four days was put into

the compressor and into the calibration apparatus. (The outgassing was re- quired because any oxygen dissolved in the oil would oxidize the oil with one

of the oxidation products being hydrogen. BNL had detected a buildup of

hydrogen in a previous run when the compressor had been charged with oil

directly from the drum) . It was found that the oil level in the gas phase of

the calibration apparatus was much higher than that taken from the compressor

system which contains filters and a liquid-nitrogen-cooled adsorption bed.

Next, a sample of oil from the compressor was put into the apparatus. The

apparent vapor pressure of this oil was much lower than the original charge

since the concentration of oil in the helium dropped nearly a factor of twenty

compared with results of the first run. We suspect that the reduction would

have been greater if all of the original oil could have been removed from the

calibration apparatus. We have concluded that the oil contains volatile

materials, relative to the vapor pressure of "pure" compressor oil, which can

be removed by adsorption, but not by simply outgassing down to a pressure of

10 microns. If the vapor pressure of the oil can be reduced to the 10 mmHg

range as stated by the manufacturer the saturation level of the oil in helium

-14 would be 10 or less far below the estimated maximum allowable concen- tration level of one ppb and also below the detectable level of any analytical

equipment with the possible exception of a helium ionization detector.

78
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)-l ca

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4-1 CO

J3

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0)

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CJ3M

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79
2.5.2 Instrumentation for Air Impurities in Helium

Removal of impurities via filtration and adsorption is state-of-the-art

technology (even if it is not fully understood) , but instruments for on-line

detection are not. Therefore, we are concentrating our efforts toward the low

level detection of helium impurities with emphasis on the detection of air.

At this stage we are considering methods to detect the adsorption bed breakthrough of air while keeping in mind the need for a rugged, reliable instrument which can be used in the field. A literature search revealed little

information applicable to our particular requirements.

Both the concentration level and type of impurity dictate the kind of

detector needed. If the level of air is in the parts-per-million range a conventional thermal conductivity detector with thermistors could be used..

Also, it is possible that by replacing the thermistors with silicon diodes,

the detector could be used in situ with a substantial increase in sensitivity.

The thermal conductivity detector is a "rugged" device, insensitive to minor

pressure and temperature fluctuations but more sensitive to flow fluctuations.

Experimental work is necessary to determine how much the sensitivity would be

enhanced by operating at the lower temperatures.

If the impurity is in the parts-per-billion range there presently are only two detectors which have the necessary sensitivity: helium ionization

and electron capture. Helium ionization detectors can detect all trace

impurities in the ppb range. However, they have the disadvantage of being

very temperature and flow sensitive. Two of the three manufacturers of helium

ionization detectors have stopped selling them because the maintenance costs were so high.

A more rugged and reliable detector is the electron capture detector.

This instrument is used to detect trace quantities of oxides and halogenated

hydrocarbons in the atmosphere. It is insensitive to nitrogen but very sensitive

to oxygen. However, the detector increases sensitivity with increasing temperature and is normally operated in the 300 to 350° C range.

If air is the major contaminant in helium, the electron capture detector

could be an attractive means of detecting bed breakthrough, especially if the

packing is a molecular sieve, since oxygen breaks through the bed before

nitrogen. Some work is needed to determine the minimum detectable level (MDL)

of the electron capture detector to oxygen in a helium stream and also to

determine how sensitive the detector is to flow, temperature and pressure

fluctuations

.

Finally, the thermal conductivity detector could be used to detect ppb of

impurities if the impurities could be concentrated in some manner. One way

would be to collect the impurities on another small adsorption bed at 80 K,

warm the bed to room temperature or above and pass this gas through a detector. Also, a sample could be collected and put into a heated quartz tube

which would allow the helium to diffuse out of the container. The residual

gas could then be analyzed for impurities. Unfortunately, these concentration

techniques could be difficult to adapt to automated field operation.

Currently, the only detector which can detect hydrogen and neon in the

ppb range is the helium ionization detector. However, it might be possible to

develop a detector sensitive to hydrogen in this low concentration range.

Future work planned on this project includes the evaluation of the

electron capture detector; a manufacturer has tentatively agreed to loan us a detector for this work. Also, if it can be determined that its MDL is in the

100 to 1000 ppb range the thermal conductivity detector using silicon diodes

will be investigated. Finally, methods for detecting low levels of hydrogen

will be studied.

80
2 . 6 References

[1] Erez, A., A proposed light-coupled system for electrical measurements,

Proc. Electro-Optical System Design Conf., pp. 183-188 (Sept. 1971).

[2] Morse, R.M. , Staniforth, A., Morse, A.R., A data system for high voltage

DC test lines, IEEE Trans, on Instrumentation IM-20, 285 (1971).

[3] Arp, V.D., Daney, D.E., Frederick, N.V. , Jones, M.C., Ludtke , P.R.,

Parrish, W. R. and Powell, R.L., Helium research in support of superconducting power transmission, Nat. Bur. Stand. (U.S.) NBSIR 75-823

(October 1975)

[4] Van Degrift, C.T., A sensitive displacement transducer using an extremely

reentrant 84 MHz cavity oscillator. Rev. Sci. Instrum. 4_5, 1171 (1974).

[5] Zimmerman, J.E., private communication.

[6] Timoshenko, S., Strength of Materials, Part II, Advanced Theory and

Problems, D. Van Nostrand and Company, Princeton (1956).

[7] Reed, R.P. and Mikesell, R.P., Low temperature mechanical properties

of copper and selected copper alloys, Nat. Bur. Stand. (U.S.) NBS Monograph 101 (1967) .

[8] Neubert, H.K.P., Instrument Transducers, Oxford University Press (1963).

[9] Read, D.T. and Ledbetter, H.M. , private correspondence; also, Ledbetter,

H. M. and Weston, W.F., Low temperature elastic properties of some copper-nickel alloys. Ultrasonics Symposium Proc, IEEE Cat. No. 75 CHO

994-4SU (1975)

81
3.0 HELIUM PROPERTIES

In the early 1970 's R. D. McCarty of our Laboratory collected and cor- related all available data on the properties of helium. The output from this

work is in the form of publications [1, 2] and computer codes which have been

widely distributed throughout the world. However, some new data have become

available, and advances have been made in the equations and computer codes

which represent that data. We have maintained a low level of effort in this

field and report here some recent developments.

3.1 Improved Helium Property Computer Codes

Mccarty's basic equation of state uses density and temperature as the

independent variables, whereas for many engineering calculations, pressure,

enthalpy or entropy often occur as independent variables. New forms of helium

equations, using these latter variables, were reported by Arp [3]. However,

the residual inconsistency (typically about 0.02 K) between the McCarty and

the Arp equations sometimes leads to problems when both equation sets are used

in calculations which demand high internal consistency. In such cases it is

necessary to use slow double-iteration techniques with McCarty 's equations

rather than to use Arp's equations.

At the Cryogenics Laboratory of the Department of Engineering Sciences in

Oxford University, England, Hands revised the computer program based on Mccarty's equations to convert them entirely to SI units, and to obtain faster

computing time with very minor increase in required computer memory [4]. In

addition, both Hands [4] and Arp [unpublished] independently developed some single and double iteration subroutines to be used when faced with the internal inconsistency problems mentioned in the above paragraph.

In the meantime, several workers have uncovered an error in the equation

which McCarty used for the dielectric constant [Johnson, Acton and Kellner,

private correspondence] . Equation 25 in reference [1] should read

3M e^l L

4-n £ + 2 p The factor 3M/4tt, which is missing in reference [1], essentially causes the

published values of (e-1) to be about 4.6% lower than the true values (M = 4.0026 for helium)

.

Dr. Hands was a Guest Worker in our Laboratory during May and June of

this year. During this time various subroutines in respectively the NBS and

Oxford computer packages were compared, and the most efficient of them as- sembled into a single computer package, in SI units. The error in the

dielectric constant subroutine was corrected. At the same time, we studied

one other small problem with the basic equation of state used in [1].

The equations actually consist of a set of three mathematically separate

equations, each valid in its own region of the density-temperature plane. The

problem comes in obtaining a smooth transition from one region to another

where the equations overlap, particularly in the region from 10 to 15 K at

pressures greater than critical. Within this latter region one can find, for

example, specific heats which waver unrealistically ; more importantly, the

(^^

equations are not thermodynamically consistent in that

J dT is not exactly

10

equal to AH along an isobar. The error is negligible at 5 x 10^ Pa (5 atmos5 pheres) but about 20 per'cent at 100 x 10 Pa. Presumably one could find

related errors in other thermodynamic functions and derivatives, but we have

82
not tested this. After some study, and discussion with McCarty, we conclude

that there is no easy way to eliminate the inconsistency, but we can smooth

the functions more effectively in this range by using a quadratic interpolation scheme between 9 and 16 K rather than the previous linear interpolation

between 10 and 15 K. This has been done in the revised computer package,

mentioned above.

To obtain this revised computer code, the reader (within the U.S.) should

contact the NBS Cryogenics Division.

To put these equation of state problems in proper perspective, it is

worth noting that the equation of state for helium is probably known more

accurately and over a wider range (compared to its critical pressure and

temperature) than is the equation of state for water, the most studied of all

fluids. The next major advance in this field will probably come from the use

of non-analytical terms to more accurately describe properties near the

critical point.

3.2 Thermal Conductivity of Helium

Until very recently, no data have existed on the thermal conductivity of

helium in the range from 4 to 20 K. This interval includes the critical point

(at 5.2 K and 2.2 x 10^ Pa), in the neighborhood of which it is known that all

fluids exhibit markedly enhanced thermal conductivities. Our published tables

[1] and computer codes in this region are based upon approximate scaling laws

against C^, developed several years ago for hydrogen. Our lack of confidence

in this helium data has been stated on a number of occasions.

Acton and Kellner at Southampton University in England have measured the

thermal conductivity of helium within this range during the past year. They

do not plan to develop a data correlation, and have very kindly forwarded all

of their data to us prior to their forthcoming publication. We find that

their data is in fact generally within about 10 percent of our published

values. We have begun to develop a correlation based upon currently accepted

thermodynamic models, but have not yet completed the work. This work will be

published, and added to the computer package of helium properties, only after

the Acton and Kellner publication is completed.

3 . 3 References

[1] McCarty, R.D., Thermophysical properties of helium 4 from 2 to 1500 K

with pressures to 1000 atmospheres, Nat. Bur. Stand. (U.S.) NBS Technical

Note 631 (1972)

[2] McCarty, R.D., Thermodynamic properties of helium 4 from 2 to 1500 K at

g pressure to 10 Pa, Journal Phys. Chem. Reference Data 2, 923 (1973) .

[3] Arp, V. , New forms of state equations for helium. Cryogenics 14_, No. 11,

593-598 (Nov. 1974).

[4] Hands, B.A., HEPROP, a computer package for the thermophysical properties

of helium down to about 2 K, Cryogenics 13_, 423 (1973) .

83
4.0. FLOW FACILITY FOR MODELLING HELIUM FLOW PASSAGES OF HIGH ASPECT

RATIO (L/D)

("LONG LINE" APPARATUS)

We have now completed construction and instrumentation of this apparatus.

It is designed for investigations into steady state and time dependent fluid

mechanical and heat transfer problems related to SPTL operation. Experiments

planned include, first, measurement of friction factors during unheated steady

flow of supercritical helium at a variety of pressures. In addition to

supplying valuable data, this will serve to checkout the apparatus and prepare

for the next phase -- studies of oscillatory behavior of supercritical heliiam

flowing with uniform heating. These studies will require a systematic exploration of operating parameters and configurations. Inlet temperature, pressure, flow rate and total temperature rise of the helium will be varied, but,

in addition, it is necessary to vary the flow impedance of equipment upstream

and downstream of the heated channel. The final experiments will investigate

the response of flow parameters to point and distributed heat pulses. Total

length of channel will be variable in increments.

4.1. Apparatus

Because of the wide range of operating conditions anticipated during the course of the NBS experiments, the apparatus was designed for blow-down, or once-through operation, as illustrated in the flow schematic, figure 4.1.

High pressure (120 atm) helium gas is regulated to near the operating pressure

and then cooled to 80 K in the precooler. The helium then flows to the subcooler where it cools to near 4 K as it passes through a counter flow heat

exchanger and a coil immersed in liquid helium. The temperature of the helium

exiting the subcooler is at present controlled by a flow mixing valve. However, installation of a regulated heater is planned in order to achieve improved temperature control. After passing through the test section (in the

experimental vessel) the helium cools, in succession, radiation shields in the

experimental vessel and the subcooler transfer line. A general view of the

apparatus is shown in figure 4.2. Appearing from left to right are the subcooler, precooler and experimental vessel.

The present test section, which is located in the experimental vessel

(figure 4.3) is wound from 5 lengths of 4.008 mm I.D. x 0.41 mm wall stainless

steel tubing. The total length is 182.9 m, giving a length to diameter ratio

of 4.6 (10)^, and the diameter of the helix is 77.5 cm, giving a value of 194

for the ratio of helix diameter to tube diameter. Epoxy-glass fiber spacers

support the coil, and each coupling serves as a pressure and temperature

station. All helium-temperature components are inside a copper radiation

shield cooled by the exit gas, and this shield in turn is inside a liquidnitrogen-cooled shield.

So that the flow impedance at either end of the test section may be

varied, throttling valves and expansion pistons (7.6 cm diameter x 30 cm

stroke) are provided as shown in figure 4.1.

The nominal range of the operating parameters is 0 to 30 atm inlet pressure, 4.5 to 100 K inlet temperature and 0 to 0.7 g/s mass flow rate.

4.2. Instrumentation

The position of the pressure and temperature instrument stations is given

in figure 4.4 and the station details in figure 4.5. Although pressure taps

84

86
87
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Figure 4.5. Themometer well-pressure tap detail.

89
are provided at each station, only positions 1 and 6 are presently connected

to pressure transducers. A variable reluctance diaphragm pressure transducer

(in addition to a high precision bourdon tube gauge) is used to measure P^^. The pressure drop across the test section is measured by a high precision capacitive-diaphragm type transducer. Future plans call for the installation of microwave cavity pressure transducers.

Temperatures at each station are measured with germanium resistance

thermometers which are soldered into the wells with Wood's metal to assure

good thermal contact. An electrical shunt around each well and 5 cm of tubing

each side insures that thermal equilibrium between thermometer and helium will

not be disturbed during heating runs when the stainless steel tubing will be

directly heated by the passage of electric current.

The mass flow rate is determined by metering the exit gas at ambient

temperature. At flow rates below 0.15 g/s a thermal type meter is used, and

at higher rates a dry diaphragm type meter. Flow rate may also be measured

with a turbine flowmeter at the inlet to the test section.

Data acquisition is by means of a mini-computer programmed to give online reduction of input emf ' s to temperatures, pressure, etc. Printout intervals as frequent as 20 s are possible, the principle limitation being the

printer speed. Cathode ray tube display of the computer output is used during

the general course of experimental operation until a permanent record is

desired.

4.3. Status

Shakedown runs have been completed and have demonstrated successful

operation of the system at 5 atm pressure and 6 K.

90
NBS-114A (REV. 7-73)

U.S. OEPT. OF COMM. 1. PL' B L IC .4TI0N OR REPORT XO. 2. Gov't Accession

BIBLIOGRAPHIC DATA No. SHEET NBSIR 77-853

3. Recipient s Acces.sion No.

4. TITLE AND SUBTITLE

HELIUM RESEARCH IN SUPPORT OF SUPERCONDUCTING POWER

TDAWCMTCC TOM

5. Publication Date February 19 77

6« Performing Organization Code

275.05

7. AUTHOR(S) M. C. Jones, V. D. Arp, W. R. Parrish, D. E. Daney.

P. R. Ludtke, N. V. Frederick and B. A. Hands

i. Performing Organ. Report No. NBSIR

9. PERFOR.MLNG ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OP STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project Task Work Unit .No. 2750551

11. Contract/Grant No. ERDA E(49-l)-3800

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

U.S. Energy Research and Development Administration

Office of Conservation

Division of Electric Energy Systems

Washington, DC 20545

13. Type of Report & Period Covered

Annual Report

July 1975 - Sept. 197fi

14. Sponsoring Agency Code

IS. SUPPLEMENTARY NOTES

16. .ABSTRACT (A 200-word or less factual summary of most sigfiiticant information. It document includes a si^ificant

bibliography or literature survey, mention it here.)

This report is the second annual report of research on helium related problems in

support of superconducting power transmission development. The report is in four

sections

In the first section; results are presented from experimental and computer modeling of the performance of current leads cooled with supercritical helium gas. Performance characteristics studied are burn-out conditions and existence of oscillation in

the helium gas.

The second section, on helium measurements, reports some conclusions on the feasibility of data transmission from high-voltage regions to grounded read-out instrumentation, on thermometry and on helium impurity measurements. A major part of this section is a detailed description with test results of microwave cavity pressure transducers for use at helium temperatures.

A third section, on helium properties, reports some improvements in computer codes

for helium properties and discusses some recent data on the thermal conductivity of

hel ium.

In the final section, a description is given of a recently completed flow facility

which has been built for research on flow and heat transfer dynamics of supercritical

helium in channels of high aspect ratio modeling superconducting power transmission

line channels.

17. KEY U'ORDS fsjx to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons

)

Burn-out; helium-cooled current leads; helium flow facility; helium impurities; microwave pressure transducer; thermal -acoustic oscillations.

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