NBSIR 78-877

HELIUM HEAT TRANSFER

AND REFRIGERATION IN SUPPORT

OF MAGNETIC FUSION ENERGY SYSTEMS

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80303

Prepared for

Department of Energy

Division of Magnetic Fusion Energy

Washington, D.C. 20545

NBS!R 78-877

HELIUM HEAT TRANSFER

AND REFRIGERATION IN SUPPORT

OF MAGNETIC FUSION ENERGY SYSTEMS

V. Arp, J. A. Brennan, P. J. Giarratano, W. R. Parrssh,

W. G. Steward, T. R. Strobridge and R. 0. Voth

Cryogenics Division

Institute for Basic Standards, National Bureau of Standards

Boulder, Colorado 80303

Prepared for

Department of Energy

Division of Magnetic Fusion Energy

Washington, D.C. 20545

February 1978

*''*eAU of *

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

Sidney Harman, Under Secretary

Jordan J. Baruch, Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director
ABSTRACT

This is the first year-end report on a program of studies on cryogenic

engineering data in support of magnetic fusion energy projects. The report is

divided into four parts: (1) an assessment of the cryogenic engineering data

base used in the MFE community, and recommendations for needed work for that

base, (2) experimental progress on measurement of transient helium heat transfer;

the data are of importance for magnet stability analysis, (3) presentation of

a newly developed general technique for analyzing the efficiency of helium refrigerators of any configuration and thereby identifying sources of inefficiency,

and (4) progress towards setting up a data bank on refrigeration system reliability

The technology assessment, item (1), is a revision of a preliminary version dated

May 13, 1977, with inclusion of feedback received from both within and outside

of the MFE community.

Both the technology assessment. Part One, and the efficiency study. Part

Three have been formatted for individual publication subsequent to this report.

The technology assessment will be distributed to those who received copies of

the preliminary edition. The efficiency study will be submitted for publication in Cryogenics. Because of our fairly recent publication on transient heat

transfer (Part Two, appendix A) , it is appropriate to do further research in

this field before writing it up for a follow-up publication in an outside journal.

Key words: Efficiency; heat transfer; helium; pressure measurement; refriaeration; stability; superconducti ng magnets ; transient heat transfer.
CONTENTS

Page

Part One. MAGNETIC FUSION ENERGY CRYOGENIC ENGINEERING PROGRAM

A SURVEY 1

I- INTRODUCTION AND EXECUTIVE SUMMARY 2

II. CRYOGENIC ENGINEERING IN MFE SYSTEMS 3

A. Superconducting Magnets 5 B. Cryopumps 9

C. Refrigerator Systems 10

III. ASSESSMENT OF THE CRYOGENIC ENGINEERING DATA BASE 12

A. Helium Heat Transfer and Flow in Superconducting Magnets . 13

B. Refrigeration 17

C. Instrumentation 22

1 . Pressure measurement 25

2. Flow measurement 32

3. Helium impurity measurement 34

IV. PROGRAM PLAN 36

A. Transient Helium Heat Transfer 36

B. Helium Impurity Detectors 37

C. Refrigerator Reliability Data 37

D. Turbo Compressor Development 38

E. Cryogenic Pressure Transducer Tests 39

F. Cryogenic Flow Measurement 41

G. Other Programs 42

V. REFERENCES AND ACRONYMS 43

Appendix A. Agenda from the WORKSHOP ON STABILITY IN SUPERCONDUCTING

MAGNETS 47

Appendix B. Attendees at the Workshop on Stability in Supercondcuti ng

Magnets 53

Part Two. HELIUM TRANSIENT HEAT TRANSFER 61

1. INTRODUCTION 62

2. FORCED CONVECTION EXPERIMENTS 62

3. PLANNED CONTINUATION OF TRANSIENT HEAT TRANSFER STUDIES .... 64

4. NOMENCLATURE 66

5. REFERENCES 66

Appendix A. TRANSIENT HELIUM HEAT TRANSFER PHASE I - STATIC COOLANT

(preprint of paper to appear in Internation Journal

of Heat and Mass Transfer) 80

Part Three, MAXIMUM PRACTICAL EFFICIENCY OF HELIUM TEMPERATURE

REFRIGERATORS Ill

i. INTRODUCTION 112

II. EFFICIENCY OF DEFINED REFRIGERATOR WITH PERFECT HEAT

EXCHANGER 113

III. EFFECT OF IMPERFECT HEAT EXCHANGER ON DEFINED REFRIGERATOR

EFFICIENCY 116

IV. RESULTS 122

iii
Page

V. CONCLUSIONS . 126

VI. NOMENCLATURE . 129

VII. REFERENCES 130

Part Four. HELIUM REFRIGERATOR RELIABILITY DATA BANK 131

* 1. GIDEP 131

2. PARTICIPATION 132

i V
NBSIR 1^-^11 , Part One

MAGNETIC FUSION ENERGY

CRYOGENIC ENGINEERING PROGRAM

A SURVEY

October 1, 1977

V. Arp, J. A. BrennaHj P. J. Giarratano,

W. R. Parrish and T. R. Strobridge
I. INTRODUCTION AND EXECUTIVE SUMMARY

Cryogenic Engineering plays an important role in Magnetic Fusion Energy

because the program requires superconducting magnets and large refrigeration

systems. Further, the role of cryogenic engineering extends from conception

through execution, approximately as follows:

In this report we consider the cryogenic engineering field wh|'ch includes

helium heat transfer, flow, refrigeration, and instrumentation, as it relates

to the above requirements. It is important to note that this report complements

the recent report of Reed, Fickett, Kasen, and McHenry (March 1, 1977) of our laboratory outlining research on cryogenic materials. Both reports share in a common data base on the MFE program which has been collected within our laboratory.

The report identifies areas where the above cryogenic data or technology

will be inadequate to meet the goals of the MFE program on a timely basis and

recommends specific research programs to fill that need. Emphasis is given to

the TNS and/or EPR programs targeted for the mid-1 980' s, with some consideration

to the DEMO stage programs in, nominally, the 1990's. It is important to recognize that a significant portion of our input on this technological summary comes from other advanced technological programs as well as the MFE program, and our specific recommendations are moderated by on-going work supported by these other

programs.

Section II discusses the MFE program components whose design, fabrication,

and/or purchase heavily involves cryogenic engineering. Section III is an assessment of current cryogenic technology supporting these components. Section

IV contains our recommendations for cryogenic research in support of these

components. Section V lists references and acronyms.

basic design

operational reliability

requires

requires

materials and systems data

proven components and

reliable controls

accurate instrumentation

documentation of the

above for standard practices .

performance evaluation

industrial safety

requires

requires

2
II. CRYOGENIC ENGINEERING IN MFE SYSTEMS

The three major MFE components based in cryogenic technology are (A) superconducting magnets, (B) cryopumps, and (C) refrigeration systems. In this section we discuss these components and the input which has been received concerning them

from the MFE community and related programs.

Table II-l gives an outline of the schedule proposed by ERDA for the development of a number of the fusion devices considered here.

3
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II. A. Superconducting Magnets

Consideration of present designs for toroidal confinement fields for the

experimental power reactor (EPR) tokamak, currently thought the most promising

confinement system, suggests that approximately 20 magnets, each of about

10 meter radius and weighing about 1^000 tons, must be constructed. Superconducting tokomak toroidal field coils are being planned for peak fields in the

7 to 15 Tesla range. Mirror magnet coils include about the same range, though

perhaps extended to 20 Tesla. Ohmic heating coils will have a peak of 4 to 10

Tesla, with rise times of 1 to 20 seconds. For comparison, the largest presentday magnet, at Argonne National Laboratory, weighs an order of magnitude less,

100 tons, and reaches a peak field of less than 2 Tesla. Table II-2 gives some

general parameters of the toroidal system designs. Both design and operating

factors contribute to the complexity of furnishing experimental fusion reactors

with superconducting magnets. [Henning, 1976]

The problems of scaling up the present technology by this order of magnitude are formidable, and it is not surprising to find that a great majority of

the research effort within the scope of this study is associated with superconducting magnets.

A wide variety of design concepts have been proposed for superconducting

MFE coils, but for our purposes we divide them into two classes. The first

consists of the bath cooling designs and the second consists of the force-cooled

designs. The difference between these two types of designs can be traced, at

least in part, to the assumed size and duration of the perturbing heat inputs

which the cooling system must absorb to achieve magnet stability. A massive

perturbation requires a fully cryostable design, where the magnet can recover

without interruption of current even if the initial perturbation causes all the

current to flow in the copper over a significant portion of the winding. Pool

boiling systems can be designed to recover from substantial perturbations of this

type, the limit being set by vapor-blanketing of the conductor. Forced convection

systems, on the other hand, would react to a substantial perturbation with a con- siderable rise in local coolant pressure and temperature accompanied by strong flow

perturbations, and full recovery may not be attainable for a time at least as long

as the transit time of the helium through the cooling channel; however, if the

initial perturbation is small in total energy content, force-cooled systems can

be highly stable. Inherent in the force-cooled approach is the assumption that

research and analysis will reach the point that massive perturbations can be

prevented by good design.

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It is important to recognize, however, that the distinction between

force-cooled and bath-cooled magnets is not always clear. For example, in some designs the background heat load on a bath cooled magnet can generate

substantial convection flow which may be analyzed, from a heat-transfer

viewpoint, as forced flow.

The conservative approach to magnet design, which has served well in the

past, has been expressed by Hilal and Boom [1976]:

"Large fusion reactor magnets such as the D magnet TF (toroidal field) coils

seem to require the reliability inherently obtained with cryogenical ly stable

conductors. There is little need to employ high current density cryogenical ly

unstable conductors since the major space requirement is for massive structure.

In addition, the amount of copper used is not a major cost item; eliminating

copper to obtain a smaller unstable conductor would be a minor achievement.

Unexpected super-to-normal transitions typical of cryogenical ly unstable coils

are unacceptable in terms of potential magnet damage and lost time for a utility

company."

However, provided that the problems of reliable design and operation of

higher current density coils can be understood and controlled, there are many

potential advantages which deserve consideration. The force-cooled magnets,

e.g., under study at MIT and supported by ORNL, lie in this category. This

approach is not unique to the ORNL group, but can be found, for example in non- MFE programs as well: a forced circulation cooling system (FCCS) has also been

selected by the ISABELLE project at BNL.

In this document we must give due consideration to both bath-cooled and

forced-convection cooled systems. Both are advocated by responsible groups

within the MFE community, as well as the broader technical community, and both are reflected in our technical assessment (section III) and recommendations

(section IV).

Cryostatic stability analysis as used by Cornish, Purcell, and Boom is

based essentially upon the early work of Stekly [1965] and the modification for

end-cooling effects suggested by Maddock, James and Norris [1969]. Recently

Wipf and Sikora at LASL have been improving and generalizing this analytical

approach, but a publication is not yet available.

Stability considerations for smaller perturbations, such as flux jumps or

small frictional conductor movements, have been summarized by Wilson [1976].

In particular, he develops the concept of transient stability, or recovery on the

millisecond time scale. .

An important trend which was mentioned both by Purcell (GA) and Boom (U.

Wise.) is towards operation at less than 4 Kelvin. Purcell believes that re- quired fields in the 10 Tesla range can be obtained more assuredly by NbTi coils

at (nominally) 3 Kelvin than by the more difficult engineering of NbjSn coils at

4 K. Boom feels that as magnets become very tall, pool boiling may have to be

7

\
abandoned because of excessive vapor collection at the top of the dewar during

periods of high heat loads. In this case, superfluid cooling may look attrac1 tive. His comments apply more to huge Energy Storage coils than to MFE systems

of the 1980's, but they may be applicable to DEMO's in the 1990's.

The problems, then, which the magnet designers face may be listed as follows

1. What is the maximum magnitude and duration of the initial perturbation

which the helium cooling system must be designed to handle?

2. What are the correct numerical values of the heat transfer parameters

which appear in the Stekly, Maddock-James-Norris, and Wilson types of

stability theories? When is it necessary to use transient rather than

steady state heat transfer correlations in these theories? How much difference is involved?

3. For pool -boiling systems, will natural convection provide enough fluid

motion to avoid vapor-blanketing during the course of an instability? Can

the geometry and orientation effects be predicted? If operation at less

than 4 K is desired, how does the attainable heat transfer depend upon

whether pool boiling at reduced pressure, or pressurized subcooled liquid

is used?

4. For forced-convection cooling systems, how large are the pressure,

temperature, and flow perturbations following an initial instability of

a given size? Will these require that the coil be de-energized in order

, to avoid burnout?

The first question (above) requires performance analysis on real magnet

systems, which is beyond the scope of this report, though it does influence the

answer to the succeeding questions. In section III we assess the current technology related to these questions, leading to specific recommendations in section

IV.
II.B. Cryopumps

Multiampere neutral beam injectors will require vacuum systems capable of

pumping large quantities of hydrogen and deuterium (greater than 10 Torr-liters

per second) while maintaining a vacuum in the 10"'* Torr range during neutral

beam pulses. In addition, large quantities of these gases have to be pumped out

without interrupting the operation of the system. Cryopumps offer some outstanding advantages for this application [Halama and Bamberger, 1975].

This subject was discussed by telephone both with H. Halama at BNL and

T. Batzer at LLL. The conclusion is that design data for such pumps is largely

available, the exceptions being:

1. If one chooses to use a gas which melts at a relatively high temperature (e.g., CO2) to smother and trap a low temperature gas at higher-thanusual cryopump temperature, performance data is lacking.

2. If the trapped gases contain tritium, its radioactive decay will re- lease heat, which may degrade the trap performance.

Though these were viewed as relevant and timely problems, they were under study,

and it was felt that additional work under this program was not necessary for

the MFE program. Accordingly, we have neglected further consideration of this

topic, and will not discuss it further in this report. However, we will be glad

to receive further input from interested reviewers, especially if other actions

are recommended.

9.
II. C. Refrigerator Systems

Essentially all of the existing work on refrigeration for MFE is connected

with the Large Coil Project Test Facility (LCPTF) at ORNL. Basic to this facility

will be a large refrigeration system with a complex of outputs to up to six

different test coils which can be mixed pool -boiling and/or force cooled. Considerable planning is going into the installation, and it appears that the

system will utilize the best of present cryogenic technology in an effective way.

However, it is important to realize that, at least at this time, it is not being

designed with the redundancy of critical components, advanced computer control,

etc., which will be demanded of an EPR for which unscheduled refrigeration

down-time may be disastrous.

The refrigeration problems listed by the LCPTF project review (April 14-15,

1977) include the following:

1) Impurities from the magnets and support structure are a problem.

Adequate impurity analysis and purification procedures need to be found.

2) Liquid helium pumps may be needed for the forced circulation systems.

However, long and expensive development programs may be required to obtain

pumps of documented reliability. ;

3) All magnets and leads will require accurate measurement of helium

coolant flow. Flow measurement systems must be calibrated since there is no experimental data which will allow accurate prediction of flowmeter performance

at the anticipated helium states and flow rates.

4) Thermal oscillations can cause excessive heat leaks into the cryogenic

system. This phenomenon is not well understood, requiring careful design and

construction of liquid helium transfer systems to try to avoid this type of

difficulty during operation.

Important input has been also obtained from the wel 1 -researched reports of

Kadi and Longsworth, and Manatt, et al . , prepared for ERDA-Power Transmission.

Much of the information they contain is applicable as well to MFE refrigerator

plans. The interested reader is encouraged to review these two reports very

carefully, and we shall not attempt to discuss them here. Suffice it just to

list the conclusions and recommendations which are applicable to MFE as well

as SPTL systems

.

1. Both reports recognize the need for on-line instrumentation to monitor

impurities in the gas stream, and recommend that research be done in this field.

See section III.C for further discussion of the topic, and IV. A. for our recommended work.

2. Both reports recommend that a Refrigerator Reliability data bank be es- tablished and operated. This is further discussed later in this section, and

recommended work is outlined in IV.C.

3. Both reports give considerable discussion to the room temperature com- pressors which are used for helium refrigerators. The screw compressor is the first

choice at this time because it is readily available and probably very reliable.

10
However, its efficiency is low (38 to 42%) and it presents severe oil removal

problems for which long time answers are unknown. Manatt, et al . go on to present

a strong case for the development of centrifugal compressors. This is discussed

in more detail in III ,C.

It is important to note that the helium refrigerator capacities required for

the EPR installations are estimated to be 10^ to 10^ larger than the largest ex- isting helium refrigerator (about 1500 watts at 4.2 K), or roughly the size of

large industrial air separation plants.

11
III. ASSESSMENT OF THE CRYOGENIC ENGINEERING DATA BASE

In this section we assess various cryogenic engineering technologies which

are available to the MFE community. The material has been collected from a number of sources and personal contacts. The assessment leads to the identification

of several areas where we recommend that DMFE support research towards a clearly

identified goal of the MFE program. We encourage direct discussion with any

readers who wish to raise questions concerning the completeness or accuracy of

the material presented.

12
III.A. Helium Heat Transfer and Flow in Superconducting Magnets

During the week of July 25-29, 1977 a most productive and useful workshop

on stability in superconducting magnets was held at Los Alamos (not funded by

this project). The agenda for the technical program is listed in Appendix A,

and the attendees are listed in Appendix B. During this workshop the problems

of heat transfer and flow were thoroughly discussed, both in a separate subgroup,

and then before the whole workshop. Typed notes from this workshop are expected

in the near future, but only handwritten notes of the participants are available

now. This section of our technology assessment strongly reflects the conclusions

of that workshop.

Heat Transfer

Accurate heat transfer data are necessary in order to analyze, understand,

and predict the stability of superconducting coils. For conservatively designed,

low current density coils, e.g., the ANL bubble chamber magnet and the Stanford

LASS magnet, existing steady state heat transfer data suffice, though tests on representative sections may always be necessary when special geometries are used.

But as current densities, operating fields, stresses, etc., are pushed upwards,

evidence from the work of Dressmer [1976] and Wilson [1976, and recent unpublished work] indicates that one must invoke some estimate of time-dependent heat

transfer in order to understand the recovery process and the border between stable

and unstable magnet performance. Toward this end, the conclusions and recommendations of the workshop may be summarized as follows:

(1) There is experimental evidence (Dressner, 1977] that steady state heat

transfer data are inapplicable during the typical 10"^ to 10"^ second recovery

time of a stable, or quasi-stable, magnet. To understand the recovery process,

transient heat transfer data, such as being developed by Dr. Steward at NBS and

Dr. Iwasa at MIT, are urgently needed. The experimental measurements should

provide accurate heat transfer data on a millisecond time scale; some special

applications will require 1 to 10 microsecond response. The NBS measurements

are the most comprehensive, with response time of 20 microseconds, but the mea- surements to date have utilized a carbon film surface which was found to inadequately represent the heat transfer of practical superconductors at short

times (a Kapitza conductance effect). The MIT measurements complement the NBS

work by utilizating practical superconductors, but with some slower response

times ( 1 millisecond), and more limited fluid conditions. Proposed work for

continuing the NBS studies with representative surfaces is outlined in IV. A.

(2) The "cold end recovery" process, first discussed by Maddock, James and

Norris [1969], requires knowledge of the complete q (heat flux) vs AT curve,

including any regions where 8q/8(AT) is negative . This "negative resistance"

region is often undocumented -- i.e., it has never been documented for poolboiling helium -- because q rather than AT is usually the controlled parameter

in heat transfer experiments. In a q-controlled experiment, this negative re- sistance region is masked by hysteretic effects. This documentation is needed,

both for steady state and, more generally, for controlled values of dT/d(time),

both positive and negative. This is a most important area of transient heat

transfer research.

13
(3) The effects of confining geometries in bath cooling are not easily I predicted, but some correlations developed for other fluids should be investigated for applicability to helium systems. As part of any large magnet

design, it is probable that specific model tests will be required. However,

the question of generalized analysis and guidelines was seriously discussed.

For encapsulated force-cooled conductors, e.g., those under study by Hoenig and

Montgomery at MIT, such analysis rightfully accompanies the experimental

work and will be a valuable guide to future design and development. For shortchannel flow perpendicular to a conductor, e.g., parallel to the plane of a pancake winding, less systematic work is in progress. First, such flow can be

induced by background heat leaks in bath-cooled systems and the possible enhancement of such flows by "chimneys" or other thermosyphon effects have apparently

not been considered for helium-cooled coils as they have for other fluids and • systems. Secondly, heat transfer in low L/D channels (e.g., pancake coil flow

passages) have perhaps not been adequately explored. For supercritical helium i this is essentially a transient heat transfer problem, related to boundary layer

formation. For bath-cooled magnets, this is related to item (4) below. Finally,

limitations, if any, to such analyses in very tall systems where substantial

gravitational pressure differences exist have not been considered.

(4) Fluid flow has relatively little effect on nucleate boiling, but there

is some evidence that it has relatively substantial effect on the transition to

film boiling. This, in turn, would have a substantial effect on the col d-end-recoverj

stability analysis, and in fact may make forced two-phase flow an especially

attractive cooling mode from a stability viewpoint. Study and documentation is | needed.

j (5) There is considerable interest in using varnished or coated conductors | in order to secure favorable manipulation of the nucleate boiling characteristic

and the transition to film boiling. This approach, first reported by Butler et al

.

[1970], has been recently given more study by Iwasa at MIT. It was felt that the

perturbation in heat transfer due to the insulating film probably could be estimated knowing the conductivity and thickness of the film, at least for steady

state. This needs to be tested for both steady state and transient heat transfer

conditions

.

! (6) Interest is reviving in the possible use of superfluid helium for rapid

removal of large heat pulses; maxima in the range 4 to 20 W/cm^, as limited by

helium properties, motion, and past history, were mentioned. Very little data

on this high-flux regime exist in the literature.

It is our considered opinion that, of all the problems outlined, the one

which requires the most lead time (because it is a difficult experiment) and which

is most useful for stability studies without having to^ specify a conductor geometry,

is the measurement of transient heat transfer. Our recommendations for this work

are in section IV. A.

14
System Behavior

System behavior in bath cooling would be related to single phase natural con- vection or thermosyphon effects at low heat flux, and displacement of liquid by

vapor formation at high heat flux. Natural convection effects can be treated

analytically in simple geometries, but the complex of passage sizes, shapes, and

orientations which one can expect in a real magnet vitiates the usefulness of

generalized analysis not tied to that particular magnet geometry. The same is

true, approximately, of vapor blanketing or two-phase fluid motion in a coil under

high heat flux. Thus we do not recommend any separately funded, generalized research

in this area for pool boiling systems.

In forced flow systems, on the other hand, time-dependent coolant flows,

pressures, and temperatures from a general analytical method can be related

quite unambiguously to the predicted performance of a specified coolant system.

A number of published stability or recovery calculations belong in this category: Green and Saibel, 1968; Bald, 1970; Mori and Inai, 1974; Todoriki and

Agatsuma, 1974; Tsukamoto, 1974; Hoenig, et al . , 1975. However, all of these

have been performed in the approximation that the coolant flow and pressure

remain unchanged as its temperature does change in response to a heat input.

In this sense they are applicable to small, fast transients, e.g., Wilson's

transient stability analysis, but they cannot begin to explore the degradation

in performance which would accompany larger or longer-lasting perturbations.

The reason that coolant flow and pressure perturbations have previously not

been included is probably because of the mathematical complexity of the problem.

At the present time it appears that the MIT group, supported by ERDA-DMFE and

OCR (Office of Coal Research)-MHD, is making good progress on the problem, and we do not see the need for additional funding through our program. It is strongly

recommended that their existing funding for this work be continued. This point

is outlined in IV.G.3.

The Possibility of System Oscillations

The tendency of liquid helium cooling systems to develop flow, pressure

and temperature oscillations has been referred to in the LCPTF study, and it

has become of sufficient concern to superconducting power transmission line

programs that our laboratory has been funded for review and study of the phenomenon.

Three distinct modes of oscillation in forced flow systems can be identified from a large body of literature on other cryogenic and high temperature

fluids. These are:

(a) density wave oscillations, which have a period approximately equal to

the residence time of the fluid in the channel (Friedly, et al . , 1967; Jones

and Peterson, 1975).

(b) acoustic or "organ pipe" oscillations with a standing wave set up

between flow discontinuities (e.g., a valve) (Thurston, Rogers and Skoglund

1967; Krishnan and Friedly, 1974).

15
(c) pressure drop or Helmholz oscillations, analogous to a mass on a spring (Maulbetsch and Griffith, 1969).

Provided the time constant for radial thermal diffusion is small compared to

the period of oscillation, these can all be studied by the application of

classical linear stability theory to the one dimensional conservation equations of mass, energy and momentum describing flow in the heated channel. But

of equal importance to the dynamics of the channel itself are the external

flow impedances or acoustic impedances at inlet and outlet of the channel

since they determine the manner in which propagated disturbances are reflected

or transformed.

We have made calculations for type (a) for supercritical helium systems.

We find that under certain circumstances (e.g., low pressures and proximity

to the transposed critical temperature) these oscillations are a real possibility in supercritical helium (Jones and Peterson, 1974). Stability computations for acoustic oscillations, type (b), are considerably more difficult

and those that have been reported (Krishnan and Friedly, 1974) have not permitted generalization of the results. We, therefore, still lack criteria

for this type of oscillation that can be applied to helium systems.

We have studied the possibility of negative differential flow resistance

in helium cooled channels and find that this is only a possibility at inlet

conditions of below about 3.5 K and below about 3 atm for short channels with

high heat flux. Thus, oscillations of type (c) may be of very limited possibility in applications of superconductivity, but other possibilities for sus- taining this mode of oscillation need investigation.

We recommend that the possibility of system oscillations should be investigated as part of the development of any specific system design, but we do

not recommend a generalized study at this time.

16
III.B. Refrigeration

Over the past fifteen years, there have been many improvements and advances

in low temperature refrigeration technology. First of all, the analytical

helium equation of state [Mann, 1962; McCarty, 1972,1973] has permitted fast,

convenient thermodynamic cycle studies yielding accurate results. Material

and fabrication developments have improved the quality of the machinery while

better instruments and controls have made refrigerator operation less labor

intensive. All of these events have been the result of the growth in the

use of superconductors for one application or another. The early superconducting

devices must have seemed large to the scientists at that time but the MFE superconducting magnets, power transmission cables and particle accelerators under

development now will cost millions of dollars and are part of progress important

to the nation. Therefore, in our opinion, it is imperative that the supporting

refrigeration equipment be the best that modern technology can produce.

Two important problems exist today; first, we must strive to make the

refrigerators as reliable as any other part of a system and second, we must

continually emphasize that for maximum efficiency the refrigerator and the

device to be cooled must, insofar as possible, be designed to complement one another. We divide this technology element into several subcategories.

Refrigeration Cycles .

The optimum design of a cryogenic refrigerator or liquefier is a complex

thermodynamic problem, primarily due to the non-ideal properties of the working

fluid. The variable specific heats of the real gas can yield designs unknowingly involving negative heat exchanger temperature differences which are

not uncovered by a simple first-law analysis. Variable properties also require

optimization of pressure ratios and inlet temperatures for the various components

used in the system. The usual approach to cycle studies is to make a parametric

study of the cycle using a computer and then choosing one of the calculated

cycles that is compatible with available components and the operational requirements of the particular installation. Although this approach gives exact results,

one is never quite sure that a different (and unconsidered) cycle might not yield

a superior efficiency.

The NBS Cryogenics Division has developed a method for calculating a maximum

possible efficiency for liquefiers as a function of component efficiency. Reductions in efficiency due to system pressure drop and heat exchanger temperature

differences were also included in the analysis. Recently, under this contract

we have expanded this work to include refrigerator as well as liquefier efficiencies, and to include temperatures down to 4 K. This work is presented in

Part Three of this report, and will be submitted to Cryogenics for outside

publication. This analysis can serve as a guide to the real gas analysis by

providing an ultimate refrigerator or liquefier efficiency based on the available component efficiency. It also shows how the efficiency of each component

in the cycle effects the overall efficiency so the economics of additional re- search to improve the efficiency of a particular component can be evaluated.

17
We do not believe that further work needs to be done in this field for MFE

systems at this time.

Refrigeration Reliability .

In 1973, NBS was commissioned by the Fermi Laboratory to visit the five

large industrial helium plants in Kansas and Oklahoma to gather data and assess

the operating history of helium expansion machines. Two of the facilities

separate helium from natural gas only, while there are four helium liquefiers at

the other three locations. These, the world's largest helium liquefiers, are roughly equivalent to a refrigerator of up to 3 kW capacity at 4.2 K. The un- published result was one of the first known surveys of operating and failure

history for helium liquefier expansion engines, heat exchangers, compressors,

purification equipment and auxiliary equipment of this size.

Recently, ERDA-Power Transmission let two identical contracts to private

industry to assess and study existing concepts and methods of cryogenic re- frigeration for superconducting transmission cables. As part of the work, many

of the major laboratories in this country and Europe were visited or contacted

to gather operating and failure histories of helium refrigerators and liquefiers

[Kadi and Longsworth], [Manatt, et al.]. Thus, the failure rate data has been

extended and much other useful information was presented. In addition, both

reports concluded that a refrigerator reliability data bank should be established to collect as much information as possible so that the maximum benefit

may be realized from the reliability engineering sciences. NBS has also informally suggested on past occasions that this data bank should be established.

NBS began work on this data bank during the present fiscal year. The

project is specified in more detail in Section IV. C. of this survey, and a year

end report is given in Part Four of this report.

Compressors

Most helium compressors are converted from air or refrigeration service.

Presently those in use are oil lubricated reciprocating or screw and non-lubricated reciprocating compressors. The lubricated machines always present the

potential for contaminating the refrigerator cold box internals even though oil

removal systems are installed. The non-lubricated reciprocating compressors

typically have piston ring life of 90 days.

The screw compressor is a rotating, positive displacement machine capable

of respectable pressure ratios in a single stage. However, for such machines

there is no long term history in helium refrigeration service and, at present, screw compressors appear to be less efficient than reciprocating compressors.

The high pressure ratio per stage without excessive heating is accomplished by

pumping relatively large amounts of cooled oil through the machine along with

the gas being compressed. Therein lies a potential problem. For continuous

operation without fouling heat exchangers, piping, turbine blading and valves,

the helium working fluid must be very pure. Various estimates of the allowable

oil impurity fall in 10 to 100 ppb range. While tests show that current oil

18
removal apparatus can achieve those low levels, oil carryover into the cold box • would have the following effects: The oil will foul or poison the adsorbers

intended to remove water and CO2 so those constituents may pass on and poison

the lower temperature adsorbers intended to remove N2 and O2 and so on until the

entire purification stream is ineffective. The potential for oil contamination

along with all the steps taken to eliminate that possibility simply must be

weighed against the advantages of smaller size, lower cost, turndown capability

and the probable long life that has already been demonstrated by screw com- pressors in fluorinated hydrocarbon applications. Screw compressors have only

recently been used for helium compression and should be attractive for refrigerators requiring a few hundred kilowatt input power (the largest screw air

compressor that we are aware of is about 800 kW). Screw compressor and the

required oil removal development for helium service should be carefully watched.

However, the much larger capacities which will be demanded of EPR refrigerators

forces consideration of much larger compressors, i.e., turbine compressors, as used in industrial plants.

At room temperature and up to the moderate pressures used for helium re- frigerators, helium behaves very nearly as an ideal, monatomic gas. The ratio

of specific heat at constant pressure to that at constant volume is 1.67 for

helium a compared to 1.40 for the diatomic gases such as oxygen, nitrogen and

hydrogen. The high specific heat ratio and low molecular weight of helium

present two compression problems.

1) The specific heat ratio determines the temperature rise of the gas

upon adiabatic compression. The larger the specific heat ratio, the

higher the temperature for a given compression ratio. In all com- pressors, the temperatures must be kept below the oil cracking temperature if lubricated or below material limits in any case. Therefore, the permissible compression ratio per stage for intercooled

helium compressors is limited relative to other gases and more stages

of compression are required for the same final pressure.

2) Rotating centrifugal compressors for high flow rates are attractive

because of demonstrated long service life. To the first order, the

pressure ratio obtainable per centrifugal stage is dependent on the

molecular weight of the gas. For helium, with molecular weight 4,

from 8 to 14 intercooled stages would be required for centrifugal

compression from one to 15 atmospheres. , ,

At this time, both centrifugal and regenerative compressors have been proposed

for helium service in the larger capacities. A regenerative compressor is a rotating machine that is capable of higher compression ratio per stage than a centrifugal machine and like the centrifugal can be oil free in the helium

spaces. Improvements in the aerodynamics and a preliminary design for 1270 kw

helium compressor modules has recently been reported by Sixsmi th. [1 976]

In many ways the helium centrifugal or regenerative compressors will be

very similar to existing machinery. The bearings, seals, lubrication and control systems should be state-of-the-art; however, the aerodynamic design and

performance with helium have not been proven for machines of the size that will

be required for MFE systems.

Our recommendations for compressor development are given in IV. C.

19
Liquid Helium Pumps

t The use of low temperature helium pumps in force-cooled systems is planned i in the LCPTF project and has been carefully analyzed by Bejan [1976]. The f general conclusion from his work is that economic cooling systems must employ

efficient cold pump recirculators in which the large flow demanded by the forcecooled superconducting magnet is confined to the cold end of the refrigerating j column; further, if the liquid helium pump efficiency is less than 40%, systems i employing elevated temperatures are more economic. Centrifugal helium pump | ( efficiencies in the 40% range have been reported by Ludtke [1975]. i

Several reports of low temperature helium pumps are available [Sixsmith , and Giarratano, 1970; McConnell, 1973; Ludtke. 1975; Morpurgo, 1977]. In addition,

Fermi Lab has obtained both centrifugal and reciprocating helium pumps on a commercial basis, and ORNL will be using a commercial reciprocating pump in the

LCPTF. Aside from the NBS work (McConnell and Ludtke), very little documentation

of pump performance exists, and cumulative operating experience is limited. The

NBS work did show that centrifugal pumps behave quite predictably in liquid

helium, and can hardly be made to cavitate. They should be quite reliable,

though several stages may be required to obtain a useable pressure ratio for MFE

systems.

Complete documentation and records of pump performance, reliability, etc. will

be invaluable to the MFE goals. We do not recommend any separate pump testing

program under this contract, but we do recommend that, when a particular project

requires a pump, sufficient funding and effort be planned to faithfully document

both its initial performance and long-term service. This should go into the

refrigerator reliability data bank, and would be an important input to subsequent program planning.

Ejectors

Most helium refrigerators operate at 4.3 K or above. However, if the

trend indicated by Purcell toward lower temperature operation is followed,

future MFE systems may be required to operate at 3 K (nominally) in order to

achieve higher field. The approach taken in the few refrigerators operating

below 4.2 K has been to install several 'vacuum' pumps in series upstream of the

normal compressor suction. Since the pressure is low, these 'vacuum' pumps must

have high volumetric capacity -- typically Roots blowers are used. Because of

this added equipment, both capital and operating costs are greater than for a conventional system. More importantly, the potential for contaminating the

helium stream is much greater since air will enter the reduced pressure system

at any leak in the equipment (e.g., at rotating seals). Freezing of contaminant

species is a leading cause of cryogenic refrigerator failure. Reliability is

also compromised with the use of mechanical blowers.

An expansion ejector is a small device which functions as a low temperature

pump driven by the potential energy of the high pressure stream, which is

normally wasted in a Joule-Thomson valve. The ejector, which has no moving

parts, is in the vacuum space of the cold box as is all the subatmospheric

20
piping. Therefore the risk of contamination is reduced to the ordinary problems

associated with low temperature refrigerators operating at pressures greater

than one atmosphere. Use of an ejector would also eliminate the need for the

large room-temperature vacuum pumps mentioned above and therefore improve system

reliabil i ty.

The ejector appears to offer substantial practical advantages for refrigeration systems operating below about 4.5 K. However, very little has been published on the performance of helium ejectors -- certainly not enough to reliably

calculate their expected performance on large refrigeration systems. The most

extensive low temperature performance data published reports work done at NBS

with nitrogen as the working fluid [Daney, McConnell and Strobridge, 1973]. Recently Quack [1977] has suggested that ejectors may advantageously replace pumps

in many applications.

We have recently designed and installed two small ejectors on one of the

refrigerators for testing ISABELLE magnets at Brookhaven National Laboratory.

With it, this particular refrigeration unit reached 3.6 K (c.f. a predicted

value of 3.5 K) whereas before the ejector was installed, the minimum attainable

temperature in the same test had been 6 K [Strobridge, 1977].

The important milestone in this work is to determine whether the straightforward ejector design now being pursued produces satisfactory performance and

design data for refrigeration below 4.2 K. Several more tests will be required

to reach the milestone, a convenient stopping point if all goes well. It is

important that this work be completed, but we do not recommend additional work

over that already planned for ISABELLE. This project is listed in section

IV. G. 5.

Control

s

The refrigeration control system for a complex, multi-magnet MFE system

will have to be more reliable than for, e.g., a bulk liquefaction plant, and it seems quite clear that advanced process-control technology will be required.

There are two parts to this problem. The first relates to the availability of

accurate pressure, temperature, liquid level, flow, etc. sensors, as is discussed

in section III.C. The second part involves the rapidly developing technology of

minicomputer or microprocessor controls. This latter topic has been discussed

in the refrigeration reports of Kadi and Longsworth, and Manatt, et al . The

conclusion from these reports is that the problem must be recognized and appropriately funded when refrigeration systems are designed, but that, in the absence of a specific application, a separate study of control systems would not

be appropriate at this time.

21
III.C. Instrumentation

The development of MFE systems will require extensive stage by stage testing in order to evaluate the design, performance and safety of superconducting

magnet systems, and it is quite apparent that accurate and reliable instrumentation will be required to monitor and evaluate these tests.

The instrumentation considered in this survey is essentially that associated with the refrigeration. Thus it does not include the important topic of

quench detection (local normal region detection) in the superconducting magnets.

This latter topic deserves separate consideration, as it was given, for example,

pit the Los Alamos workshop on magnet stability.

The environmental conditions to which the measurement devices will be subject are high magnetic fields (8-10 T), if located on or near the magnet, and

low temperatures 2.2-20 K) as a result of the liquid or supercritical helium

coolant. Walstrom [1976] has discussed typical variables to be measured and

has classified them as: (1) controlled or input variables which may be set by

the test engineer for normal operating conditions, (2) diagnostic variables

which indicate the response of the system to the input variables and (3) alarm

variables which are a "subset of the input and diagnostic variables and which

provide input for initiating a protective shutdown of the coil system in the

event of a quench or failure of a critical facility or test coil component."

Table III.l is a summary of the above variables to be measured with the

appropriate instrumentation. In Table III. 2 we present the anticipated ranges

of operation and accuracies desired for all the instrumentation which has been

considered in this MFE study.

The assessment of instrumentation for pressure, flow, liquid level, and

density measurements at helium temperatures has been based on our experience

with these types of measurements here in the Cryogenics Division of NBS-Boulder

plus extensive evaluations and surveys of instrumentation already done for other

cryogenic fluids. The publications which contain these previous evaluations

are:

(1) Arvidson, J. M. and Brennan, J. A., ASRDI Oxygen Technology Survey,

Vol. Ill: Pressure Measurement, NASA SP-3092, 1975.

(2) Sparks, L. L., ASRDI Oxygen Technology Survey, Vol. IV: Low Temperature Measurement, NASA SP-3073, 1974.

(3) Roder, H. M., ASRDI Oxygen Technology Survey, Vol. V: Density and

Liquid Level Measurement Instrumentation for the Cryogenic Fluids

Oxygen, Hydrogen, and Nitrogen, NASA SP-3083, 1974.

(4) Mann, D. B., ASRDI Oxygen Technology Survey Volume VI: Flow Measurement Instrumentation, NASA SP-3084, 1974.

22
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24
(5) Brennan, J. A., Stokes, R. W., Kneebone, C. H. and Mann, D. B., An

Evaluation of Selected Angular Momentum, Vortex Shedding and Orifice

Cryogenic Flowmeters, NBS Tech. Note 650, March 1974.

The above instrumentation assessments are comprehensive in scope and contain

e.g., descriptions and operating principles of all the traditional and more

recent devices known to provide the previously mentioned measurements in cryogenic applications. Performance capabilities, encompassing probable precisions

to be expected, ranges of operation, potential problem areas in application of

the devices and other information of interest to potential users are also included.

Careful consideration of these documents and our experience with measurement systems operating at cryogenic temperatures lead to the following:

(1) The state of the art in thermometry and liquid level or phase detection

is sufficiently advanced for applications in MFE systems. Arp, et al

[1975] have investigated the sensitivity of thermometers to electromagnetic fields and Sample and Rubin [1976] have provided a later

summary of errors in thermometry induced by magnetic effects. These

papers, along with the evaluations of Sparks [1974] and Roder [1974]

provide an adequate basis for selection of all but the newest sensors,

e.g., carbon-in-glass or capacitance types. It is recommended that

some of these new types be tried by ongoing projects, but a separately

funded study of them is unnecessary.

(2) The state of the art for density determination is somewhat limited

both in documented cryogenic (particularly liquid or supercritical

helium) experience with densimeters and in commercial availability

of these instruments. However, simultaneous temperature and pres- sure measurements along with an adequate equation of state, e.g.

McCarty [1972], provide the density. Therefore, the lack of significant development for densimeters poses no particular problem for

successful MFE system instrumentation (provided, of course, adequate

thermometry and pressure transducers are available).

(3) Pressure and flow are the measurements that are potentially the most

difficult to achieve accurately and reliably in MFE systems. Although

the difficulties, primarily in calibration and shifts in calibration

due to temperature effects, are not unique to helium temperatures,

they nevertheless are unresolved.

In view of the above we have concentrated our attention in this report on pressure and flow measurements and have provided recommendations to alleviate

the problems areas.

III.C.l. Pressure measurement . Pressure measurements in helium cooling systems

may be accomplished in either of two ways, (1) by bringing pressure taps out

from the low temperature region to room temperature, where a wide variety of

conventional instrumentation may be used, or (2) by placing a transducer at low

temperature, in the helium system, and bringing an electrical signal out to room

25
temperature. It is common to experience difficulties with method (1), e.g., the

temperature gradients in the pressure taps promote thermoacoustic oscillations,

variable thermal gradients show up as a drift in the output, and lengthy pres- sure lines may easily reduce the frequency response to below acceptable limits.

Coolant supply line pressure, vapor bath pressure, magnet system coolant inlet

and downstream pressures may also be monitored. In addition, an important and

practical method for measuring helium flow is to use a venturi or orifice plate

in conjunction with an appropriate pressure transducer.

Temperature effects .

Changing temperature conditions are the origin of many of the difficulties

associated with cryogenic pressure measurement. Dean and Flynn [1966] have

analyzed the temperature dependence of the performance of non-self generating

pressure transducers (any that require external power to carry the information,

e.g., strain gauge and capacitance bridges, linear variable differential transformers, and potentiometers). The essence of the analysis is that:

(1) The temperature dependence of the sensitivity shift (change of slope

of output vs pressure curve) is due primarily to the change in Young's

modulus upon which the spring constant of the force summing element

of the transducer depends. Secondarily, the expansivity of materials

contributes to the sensitivity shift.

(2) The temperature dependence of the zero shift (bias of the output vs pressure cruve) is a function of the size and expansivity of the

transducer components.

(3) Both sensitivity and zero shifts are worse under the condition of a temperature gradient across the transducer body.

Our experience with cryogenic pressure measurement and some commercially

available transducers lead to the following suggestions for avoiding temperature

effects on pressure transducer performance.

(1) First, we note that little can be done about Young's modulus temperature dependence for a given material but --

(2) Proper selection of materials for transducer components can minimize

this effect as well as expansivity effects.

(3) Proper design can minimize the expansivity effects even though ex- pansivity changes are unavoidable (e.g., in a potentiometric device

described by Dean and Flynn the wiper arm acts as a cantilever spring

to absorb the differential contraction (or expansion) instead of being

moved along the variable resistor which would thus cause the sensitivity and or zero shift.

(4) Avoid extreme environmental conditions (but with a corresponding loss

in response) by remote location of the sensor. For single phase fluid

in the sensing lines the frequency response may be estimated from the

following considerations:

26
a) tubing and tranducer may act as:

i) organ pipe (volume of pipe large compared to the transducer cavity) with the resonant frequency given by

f - ^

where v is the velocity of sound in the measurand and

I is the tube length.

ii) Helmholz resonator (volume of transducer cavity large

compared to the pipe, e.g. close-coupled installation)

with the resonant frequency given by

2

^

27T (il + 1.7 r)V

where

r = pipe radius

V = volume of transducer cavity.

(5) Locate one flush mounted transducer in the extreme environment and

another transducer at the end of a tube in a more temperate environment. The flush mounted transducer will provide the high frequency

response (dynamic levels, e.g., with small piezoelectric transducer)

and remote mounted transducer will provide static levels. The anticipation is that the temperature shock contribution to dynamic response

will be small or of low enough frequency to separate easily.

(6) Employ a close coupled temperature regulated pressure transducer installation. A heater and thermostat would be required to regulate

the transducer temperature, and pressure tap lines must be optimized

so that sufficient response is obtained but with heat leak to cryogen

minimized. Further, the transducer must be oriented such that the

cryogen (if in liquid state) does not flow into the transducer cavity

and insulation around the transducer should be employed. It is possible

to accomplish the same result (i.e., minimize temperature gradients)

without the heater and regulator but using the insulation. The effect

is to increase the time constant of the transducer (it will eventually

reach the temperature of the cryogen) and sensitivity and zero shifts

will still be experienced but thermal gradient effects will be reduced.

Survey of commercial availability of pressure transducers for low temperature

application .

In order to assess the instrumentation industry capability to provide pressure transducers operable at the temperatures of interest (< 20 K) we have sent

requests for technical literature on cryogenic pressure transducers to more than

27
70 different suppliers. The source of potential suppliers was: The Thomas Register

"Cryogenics and Industrial Gases Buyers Guide," and the listing of exhibitors at

the Instrument Society of America Exhibit and Conference, Houston, Texas, 1976.

Response has been sporadic Cfar from complete) and not always relevant to the

inquiry. However, a report completed in 1972 assessing pressure measurement

system for LOX and LH2 systems [Hayakawa, 1972] has provided an indication of

commercial availability of these devices.

Our impressions, as determined by our own experience with commercial pres- sure transducers and as determined by the responses to our request for technical

literature, are that since 1972 (time of Hayakawa report) industry has not | increased their capability in the area of cryogenic pressure measurement. Even j more important, the test data and conditions of test for substantiating advertised performance is not readily available for assessment.

As mentioned previously, temperature effects on transducer performance are difficult to predict but are one of the largest sources of pressure measurement

errors in cryogenic systems. There are several reports of tests in independent laboratories of commercially available pressure transducers (e.g.:

Smelser, 1962; Kinzie and Murphy, 1967; and Horn, 1965). All these reports point

out and substantiate the susceptibility of the devices to temperature effects.

Furthermore, Kinzie and Murphy noted a substantial difference in performance of

transducers tested when sensing the pressure in a flowing cryogen compared to

pressure in the same transducers sensing the pressure of a static cryogen. ;

Typical thermal sensitivity shifts (TSS) and thermal zero shifts (TZS) for

commercial pressure transducers range from 0.009% F.S./K (C.005% F.S./°F) to

0.05% F.S./K (0.03% F.S./°F) [Hayakawa, 1972]. Under the assumption of a temerature change of 296 K (i.e., from 300 K to 4 K) then combined zero shift and

sensitivity shift errors could range from aout 5% to 31% if no effort was made

to recalibrate the transducer at the expected temperature of operation. Transient

and thermal grandient effects would most likely increase the error.

Resolution capability and other performance parameters of commercial pres- sure transducers are generally adequate for purposes of MFE systems but space

limitations may restrict the use of some of the instruments.

Survey of literature for pressure measuring devices designed for/or operated at

liquid helium temperatures .

We have reviewed the technical literature for "laboratory" pressure mea- suring devices which have been designed for or operated at liquid helium temperatures but which are not commercially available. A summary of the review , is given in tables III-3 and III-4. It was not possible to determine from the

| papers all the parameters which generally represent performance capability of |- pressure transducers; thus, the tables are not complete for all parameters.

All of these devices appear to fall into two general groupings. The first

| group (table III-3) are instruments employing pressure ports conducive to engineering application. This group also all use capacitance as the general operating principle (either in direct measurement of capacitance change with pressure, or frequency change of LC circuit as capacitance changes with pressure).

28
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The second group (table III-4) are "devices" whose physical properties

change with strain (or pressure), e.g., carbon resistors and tunnel junction

resistance, or devices which rely on pressure dependence of phenomena such as superconducting transition temperature or Faraday magneto optical effect.

Generally, these phenomena are weakly dependent on pressure and/or require

sophisticated equipment for their operation (e.g., the change in superconducting

transition temperature of tin with pressure, dTc/dP = -4.5 x 10"^ K/atm and

the need for optical system for Faraday magneto-optic device). For these

reasons Group 2 pressure measuring devices are not considered pratical for the

application at hand and only the operating principles are given in table III-3.

Returning to a discussion of table III-3, we shall note first that they all

have been designed and built specifically for liquid helium temperature operation.

A careful selection of materials and carefully controlled fabrication techniques

have been employed in their construction. The matter of fabrication is not

trivial since leaks that could develop if care is not exercised in construction

would render the transducer useless for pressure measurement.

This group is also characterized by low power dissipation (attractive for use in helium systems) and good resolution for high accuracy measuremenis

While "all these instruments have been designed for and have been demonstrated

to be successful in operation at helium temperatures, we note that temperature

effects on sensitivity and zero bias are still a potential source of error in

pressure measurement if room temperature or liquid nitrogen calibrations are

used for the helium temperature operation. (For changes of temperature between

300 K and 4 K, the sensitivity variation ranges between 2.5°^ to "quite large"

for these instruments.)

Finally, the small size of instruments (except for #3) is compatible with

the anticipated space restrictions imposed when used in MFE systems.

Magnetic field effects .

Although magnetic effects were not to be considered in great detail in

this preliminary instrumentation assessment it is apparent that since some of

the pressure transducers may be required to operate in a high magnetic field

region (8-10 T) this effect will have to be considered in the selection process.

Briefly, the magnetic effects may be considered as intrinsic or extrinsic.

There is an intrinsic effect if the operating principle of the sensing device

is inherently dependent on magnetic field (e.g., superconducting transition

temperature, magnetic reluctance, and resistance). There may be an extrinsic

effect if proper material selection is not exercised in the construction of

the device. For example, even though the operating principle of the sensing

device is insensitive to magnetic field (such as capacitance devices) the

improper use of magnetic materials in the construction may distort the field

as well as introducing errors due to unanticipated body forces (and strains)

induced by the high magnetic field on the transducer components. Thus possible

magnetic effects are:

31
Intrinsic effects of magnetic field on:

Superconducting transition temperature

Magnetic reluctance

Resistance

Extrinsic effects of magnetic field:

May induce unanticipated body forces if improper materials used

Eddy current heating may introduce error if device is temperature

sensitive

(A.C.) Electromagnetic noise may introduce signal processing

difficulties (e.g., trying to record low-level output signal

in el ectromagnetical ly noisy environment.

Either of the above effects will introduce error in pressure measurement

and therefore instruments susceptible to them must be avoided.

Summary of Pressure Measurement Problems

The primary obstacles to selection and/or application of pressure transducers for operation in MFE systems are:

(1) Important detailed calibration data are not available on the few

instruments which may be suitable, e.g.,

:-; a) Temperature change effects on sensitivity and zero,

b) Temperature gradient effects.

(2) Insufficient information on optimization of installation configurations to reduce temperature effects and maintain required time

response.

(3) Lack of standardization of cryogenic test procedures for uniform

evaluation of existing or proposed cryogenic pressure transducers.

(4) No readily available source of potential commercial suppliers of

cryogenic pressure transducers with accompanying test data to substantiate advertised performance; no readily identifiable commercial

pressure transducers with proven performance.

(5) Lack of information on extent of magnetic field effects on transducer

operation. (For applications where the transducer would experience

high magnetic fields.)

Our recommendations for work in this area are given in IV. D.

1 1 I.e. 2. Flow measurement . Helium flow through MFE systems may often be split

into parallel channels so as to keep the overall hydraulic resistance or temperature rise within acceptable bounds. In such cases, it is necessary to monitor

the individual flows since the development of a plugged flow channel may not

otherwise be evident until a portion of the magnet unexpectedly goes normal.

32
Measurement of helium flow in the laboratory is not new, but the various

measurement instruments and techniques have never been systematically documented and evaluated as they have, for example, for oxygen or LNG [Mann, 1974;

Brennan, et al . , 1 974]

Very little data are available which are directly applicable to flow mea- surement of liquid or low temperature gaseous helium. Almond, et al . , (1972)

and Niinikoski, (1972) have reported flow measurement of helium gas (less than

1.5 mol/s in He^-He"* dilution refrigerator systems) using thermal flowmeter

techniques, but the information presented is generally too brief to allow an overall evaluation of the flowmeter and the accuracy claims. For applications

where long time constants and low accuracy requirements are sufficient, these

flowmeters are probably quite adequate but at the present time there is no acceptable flow facility available for providing base line information on helium

flow measurements.

For liquid helium flow mec^surements , there is a considerable amount of

indirect data available from the space program to indicate some flowmeters

that might be appropriate for liquid helium service. Predominate among the

flowmeter types are the turbine and the head type. Turbine flowmeters have been

used for nearly all the cryogenic fluids and enjoy wide acceptability in the

space program for measurements of liquids hydrogen, oxygen and nitrogen. They

are delicate instruments, however, and can be severely damaged during cooldown

if allowed to spin at high speeds in vaporized gas. Turbine flowmeters nominally have a flow range of 10 to 1 and are available in a wide variety of sizes.

To prevent large errors, the calibration of a turbine flowmeter must be

done in the fluid in which it will be used. Little success has been reported in

efforts to calculate or otherwise predict the calibration of a turbine flowmeter

using surrogate fluid calibrations. The same is true to some degree with all

flowmeter types but more effort has been placed on this capability for turbine

flowmeters than on most other types and with little success.

Vortex shedding and ultrasonic flowmeters might also be appropriate for

helium service but no data is presently available for these flowmeters at

temperatures lower than that of liquid nitrogen.

Head type flowmeters (Venturis and orifices) are the most promising candidates for use in large refrigeration systems. However, for all but possibly the

largest liquid helium flows, the meter diameter will probably be less than the

7.5 cm which is the smallest size for which ASME codes exist. ASME experience

indicates that the calibration of smaller sizes cannot be reliably calculated

from their physical dimensions, presumably because of normal machining inaccuracies.

Thus, absolute measurement errors of approximately 5% would not be un- expected in the use of this type of helium flowmeter, though measurement precision would be much better, nominally ~ 2% over its design range. When desired, the accuracy could be brought within usual limits by calibration, which

normally for this type of meter is expected to be independent of fluid properties. However, helium has a high compressibility and extremely low viscosity

33
compared to water, which is used for most industrial calibrations, and no ex- perimental work has shown that an industrial meter calibration will apply to a head meter used in liquid helium. A small bit of theoretical work is being done

under this contract on the expected sensitivity of venturi flow measurement to

helium properties, especially near the transposed critical line, where neither

incompressible fluid nor ideal gas properties are even remotely valid.

Head type flowmeters would benefit greatly if the required pressure measurements could be made at the flowmeter instead of at the room temperature end

of the impulse lines. Pressure oscillations in the lines are one of the most

serious problems with this type of flowmeter when used in cryogenic systems.

Normally this type flowmeter is limited to about a 3 to 1 flow range by the

range of the pressure measurement instrumentation.

The lack of a helium flow facility seriously impairs the state-of-the-art

determination of helium flow measurements. As requirements for these measurements become more critical, the more serious the need for a facility will be- come. It would be well advised to develop a suitable facility sufficiently far

in advance of the need to allow flowmeter manufacturers the necessary lead time

to make any required modifications to their flowmeters. However, the expected

cost of such a facility would be 0.5 to 1.0 M$, as judged by the cost of the

NBS liquid nitrogen flow loop, and even with cost-sharing this may be difficult

to justify from the DMFE budget.

Furthermore, as with pressure instrumentation, a readily available source

of potential commercial suppliers of helium flow measurement devices does not

exist.

Our recommended program to determine the accuracy of helium head-type flow

calibrations using industrial water calibration facilities is outlined in IV. E.

III.C.3. Helium impurity measurement . A major cause of unscheduled downtime in

helium refrigerators is the plugging by impurities. The LCPTF report and both

refrigeration studies (Kadi and Longsworth, Manatt, et al.) point out the need

for reliable impurity monitors. Accurate impurity level measurement requires a system which collects a representative sample and a detector which measures the

impurity level. The sampling system must be designed to eliminate any change in

the sample composition by addition of impurities or by precipitation of impurities. Ideally, the monitoring device should be in situ to avoid sampling

problems. This device should be a continuous, on-line detector which can be set

to sound an alarm if the impurity level exceeds a certain preset value; simple

calculations show that this preset value must be in the range of 1 ppm or less

if impurity plugging is to be eliminated.

Selection of the optimum type of impurity detector depends upon the impurity and whether it is in the gas phase (i.e., homogeneous mixture) or in a liquid or solid phase (i.e., heterogeneous mixture). Jensen and Harrison at BNL are attacking the problem of detecting oil droplets in the helium discharge from

a screw compressor used for SPTL studies. They have a private company developing a light scattering device to detect oil droplets (or any other liquid or

1

34
solid impurity) in the 0.1 to 10 micron size range. The system is promising and

should be an important input to the MFE community; however, the project is

moving along and requires no additional funding through this contract at this

time.

To our knowledge, no one is actively trying to develop an on-line monitor

for detecting impurities present as a gas. Duval at LLL uses a gas chromatograph containing a helium ionization detector to detect impurities in the 10 to

20 ppb range. However, the instrument requires a trained technician and frequent maintenance to keep it operating properly. Similar instruments are in use

at BNL (SPTL project) and at SLAC.

There is an obvious need to evaluate the existing techniques for detecting

trace gas impurities in helium to see if they can be adapted for continuous on- line service. Based on the need for this type of work we recommend initiating

the research program described in section IV. A.

35
IV. PROGRAM PLAN

The following program plan is based on our assessment of the cryogenic

engineering data base for MFE systems, as discussed in section III. It assumes

a substantial DMFE committment totaling about $300,000 per year.

IV. A. Transient Helium Heat Transfer

The need for transient heat transfer with temperature controlled (rather than

heat flux controlled) surfaces in order to understand and predict superconducting

magnet stability was a dominant theme at the LASL workshop, July 25-29, 1977.

When heat is suddenly applied to a surface cooled by helium, a finite time is

required to build up the fluid convection or vapor content to its steady state

value. This time interval ranges from 0.01 to several seconds during which the

temperature difference and heat transfer rate will differ significantly from the

values finally achieved in steady state. The problem is that flux jumps, training

steps, etc. typically occur on a much shorter time scale, so that regime is somewhat independent of fluid motion, i.e., pool boiling vs forced convection, so that

this type of data are readily applied to both cryostabil i ty analysis [Dressner,

1976] and transient stability analysis [Wilson, 1976].

Relevant heat transfer measurements for these analyses have been reported by

Iwasa and Apgar [1976] at MIT, and Steward [1977] at NBS. The MIT measurement

used both bare and insulated copper for their heat transfer surface, pool boiling

helium at 1 atmosphere, and relatively slow response, about 1 millisecond. Recently, at the LASL conference, they indicated that some of their measurements

were of uncertain accuracy because of a thin film of helium leaking into the

insulation between their heater and their heat transfer surface. The reported

NBS measurements covered a much wider range of helium conditions, geometry effects,

and times (down to 20 microseconds), but used a carbon film heat transfer surface.

The current NBS activity is devoted to producing a temperature-controlled copper

surface in contact with the helium.

The temperature-controlled heat- transfer surface, as recommended by the LASL

workshop, is required so that the negative slope portions of the q (heat flux) vs. AT (temperature rise) curves can be traced out as a function of time. This "negative resistance" portion of the heat transfer curve is of particular importance

in understanding the recovery process of superconductivity following a localized

energy release. It can show up as hysteresis if q rather than AT is controlled

in the experiments, and there is some question as to its magnitude during transient

conditions. Its magnitude during steady state conditions possible can be inferred

from existing measurements, but it has never been well documented for liquid helium

The AFOSR has indicated that they will support this project with $40K in

FY 78, which is one-half of the full time cost for the project. Their interests

and objectives are coincident with those of the MFE stability analysis, with one exception: they are interested in faster response times (20 microseconds or less,

simulating fast magnet pulses) than are necessary for MFE work (approximately 1 millisecond). We have already achieved this fast response.

36
We propose that ERDA-DMFE support this work with an additional $40K, so that

full progress can be made in this important and timely project.

IV. B. Helium Impurity Detectors

The objective of this work is to develop an on-line system which continuously

monitors trace gaseous impurities in helium refrigeration systems.

As a guide to selecting the proper detector, those developed for gas chromatography are the prime candidates, but other detectors also will be evaluated.

Each detector will be judged on its sensitivity to composition, temperature,

flow and pressure fluctuations. Based on the results of this phase, the most

promising detectors will be evaluated experimentally.

Methods for obtaining a representative sample from a flowing helium stream

and for controlled concentration of impurities will be considered. If concentration

techniques are feasible, it will offer the possibility of using more reliable and

rugged detectors. These sampling problems will be cost-shared where feasible with

an on-going project supported by a consortium of LNG (liquefied natural gas) manu- facturers, who face the same type of problems in custody-transfer of LNG.

Cost (3 year project)

Preliminary screening and evaluation

Laboratory evaluation of most promising

detectors, development as needed

Study of performance on a real system:

the problem of obtaining a representative sample

$170K

IV. C. Refrigerator Reliability Data

This project is intended to gather information on the reliability of helium

refrigeration systems as they relate to needs and problems of the MFE program

plan. Its output will be used to further define and delineate the critical elements

in the operation of such systems, and thus to recommend procedures and/or necessary

research.

The program is proceeding as follows:

1) Determine the extent to which NBS can legally be involved in such a program.

(We are precluded, for example, from recommending a specific product.) Presuming that our proposed actions are judged appropriate, we would:

2) Determine exactly the data needed by reliability experts. At this time we

intend that the reliability analysis be done by established experts in the

field;

3) Determine how much interaction there can or should be with the GovernmentIndustry Data Exchange Program (GIDEP) or other similar programs. We will

take as much guidance as possible from the existing failure rate program.

$ 5K

80K

95K

37
These three steps have been completed in FY 77 (see Part Four of this report for

a detailed discussion). The next steps are as follows:

4) Solicit the cooperation of managers of all possible helium refrigeration/

liquefaction facilities;

5) Design the data collection system, taking into account suggestions from

both GIDEP and facility managers;

6) Start the program with heavy NBS involvement which will be phased out if It

is found that one of the other established failure rate progrms can suc- cessfully assimilate this project.

Progress in this program will have to be carefully assessed at intermediate

stages in its development. We will also try to obtain the financial involvement

of other agencies which would use the data when the system is operational.

Cost: $10K/yr

IV. D. Turbo Compressor Development

Based upon the potential for significantly increased efficiency and reliability of refrigeration systems, as compared with those using screw compressors,

we recommend that a substantial program be initiated for room temperature helium

turbo compressor development. This could take two directions:

1) Develop machines specifically for helium, a low density, high gamma

gas. These could either be centrifugal, or of the regenerative type (which produce a higher compression ratio and might be easier to fabricate, though they are not so well known). The developed unit should be at least two stages with

a throughput equivalent to an 800 kW compressor operating between 1 and 15 atm.

The units should be fully instrumented to obtain all pertinent data and should

be supplied with a test stand suitable for continuous operation.

Output from this program would be the design data necessary for private

industry to build a practical, full size helium turbo compressor when needed for

refrigeration service.

2) The Union Carbide-Linde Division presently holds a patent for a com- bined helium-freon refrigeration cycle which uses conventional centrifugal com- pressors. This has never been used to obtain 4 K refrigeration, to our knowledge

and one naturally wonders whether the helium-freon separation (which must occur

at some intermediate temperature in the cycle) will be complete enough to avoid

plugging the 4 K section with solid freon during extended operation. It is con- ceivable that phase equilibria data may be required to answer these questions.

Our inclination is to opt for the compressor development work, based upon

the seemingly higher probability of reliable long-term operation. However,

38
before making such a decision, the freon-helium approach needs much more careful

consideration than we have given it to date.

We propose, therefore, first to enter into discussions with Linde concerning

their degree of assurance that the freon-helium compressors will remain impurityfree for extended 4 K refrigeration service. Then, because we in the Cryogenics

Division lack special expertise in turbo compressors, we will convene a panel of

appropriate specialists from government laboratories and industrial experts to assess the complete picture and recommend a course of action which will best fit

with the program goals. This panel will provide steering and guidance for the

program throughout its life.

Cost: ,

Preliminary discussions, panel recommendations $ lOK

Subsequent development:

Quite uncertain, perhaps $200K/yr for 4 years.

This project would be by far the largest single item in the proposed program

plan. Because it would be of considerable value to other large superconductive

programs as well, every effort should be made to obtain cost sharing with them.

Further, it is reasonably possible that industry, sensing the potential value,

may be willing to engage in cost-sharing with this program.

IV. E Cryogenic Pressure Transducer Tests

This project has two parts (1) development of a calibration facility, and

(2) screening and evaluation of transducers for liquid helium service.

The development of the calibration facility would be done at NBS, where

design has already begun using NASA funding. This ERDA project would be re- sponsible only for the incremental cost of necessary modifications to make the

facility usable down to 4 K, rather than just down to 76 K, and to modify the

external control system to achieve calibrations at up to 20 Hz at low base

pressure appropriate to helium cooling systems.

In the screening and evaluation work, candidate commercial transducers would

be tested. If it appears that small production changes, e.g., use of an alternate

material for a given component, would result in superior cryogenic performance,

these changes sould be tried and evaluated. Because of the relatively small

market for cryogenic transducers, this stage of work probably would involve small

"development" contracts as well as outright purchases.

The output from this project would be a quantitative assessment on cryogenic

pressure transducers for use in MFE systems.

39
Cost:

(Prior NASA support $ 55K)

Assembly, system tests, procedure development ^^j,

(ERDA share) '^'^

Transducer purchases; small contracts; experimental evaluations; accuracy 2%

(Additional work to prove the ultimate system

accuracy -- not requested from DMFE

50K

60K)

40
IV. F. Cryogenic Flow Measurement

The absence of any facility within the U.S. (and probably the world) for

evaluating liquid helium flowmeters is due, at least in part, to the probable

expense in setting up such a facility. The throughput of an 800 kW (1000 H.P.)

helium refrigerator is about 2 liters/s at liquid density, or 300 g/s. A

facility for calibrating liquid flowmeters at this flow rate would cost 0.5 to

1.0 M$, as judged by the cost of the NBS liquid nitrogen flow loop, and even with cost-sharing could hardly be justified out of the DMFE budget.

We propose a much more modest approach based upon head-type flowmeters with

in-situ pressure measurement. The objective will be to show that an industrial

water calibration can be used to predict a liquid helium calibration (hopefully)

to within the nominal precision of about 2%. We do not anticipate that, for this

accuracy appropriate to the MFE program, it will be necessary to test at more

than one helium state, since the relevant properties (viscosity and conductivity)

do not vary greatly except near the critical point and transposed critical line;

in practice, the meters generally would not be used within this region. For con- venience, the helium state can be chosen as boiling (or near boiling) liquid.

The flow rate can be smaller than quoted above, provided that the meter size does

not get too small for standard fabrication techniques to be used in building it;

a design flow rate of 30 g/s represents a reasonable compromise.

A non-trivial problem with head-type flowmeters for liquid helium is that

of obtaining an accurate AP reading. Long pressure taps running to an external

(warm) transducer tend to oscillate (as discussed in III.C). The cure for this

often involves very high fluid impedances with consequent long run times and

large helium use. Greater testing efficiency would be obtained if these tests were made subsequent to those in IV. E., so that a calibrated cold pressure transducer could be used. Likewise, this combination would be very practical for a working installation.

The procedure will be to build a modest helium flow loop capable of accepting the meter under test. The helium calibration can be obtained by liquid level

measurement, motion of a piston or bellows, or possibly by warming to room temperature. A nominal accuracy of 2.1 would be a design goal. The installation

would not be intended as a permanent standard.

Only a venturi meter should be tested, since these are to be recommended

over orifice meters for their generally higher accuracy. These would be patterned

after ASME standards for larger sizes. The meters would be calibrated in both

an industrial water calibration facility and the helium flow loop.

The estimated cost of this project is as follows:

Design, build, and check out liquid helium flow loop

*gQI^

for flowmeter tests for 2-3% accuracy.

Build one venturi with cold pressure transducers,

evaluate performance in helium flow and water $40K

calibration loops.

41
IV. G. Other Programs *

It is important to explicitly recognize and encourage certain on-going programs

which are presently funded by other agencies, or already supported by DMFE. Their

output and goals are of considerable importance to the MFE program, within the

framework of this cryogenic engineering assessment. The previous recommendations

IV. A to IV. F are made assuming these projects will continue. These include:

1) The Low Temperature Materials Program, also funded through our laboratory. Our program does not overlap with it, but shares many common interests

and resources.

2) The helium impurity monitoring instrumentation being evaluated at BNL

for the Superconducting Power Line project. This instrument is designed to detect

entrained liquids or solids, rather than impurities in solution as recommended

for study in IV. A. Obviously, these are complementary projects and close com- munication between the projects must be maintained.

3) The MIT studies on forced-cooling system response to a pulse heat input.

These studies are most valuable in completing the more general analysis of cooling

system requirements for magnet stability. Complementary activities within this

area are funded by DMFE (through ORNL) and by OCR-MHD.

4) Low temperature pump usage and studies at ORNL for MFE, and at Fermi

Lab for HEP. Both centrifugal and reciprocating pumps are in use at Fermi Lab.

Pumps may be an important element in the LCPTF and other cooling systems in the

future. Input to the GIDEP program is recommended.

5) The ejector studies now underway within the BNL ISABELLE program.

These simple devices have the potential of obtaining increased system performance

by lowered operating temperature, without the complications and expenses of subatmospheric refrigerator suction pressure.

The regular project review which will be held to assess the performance

and direction of this cryogenic engineering program should explicitly include

consideration of the above items.

42
V. REFERENCES AND ACRONYMS

Almond, D. P., Lea, M. J., Pickett, G. R., A simple He^/He^ gas flowmeter for use with a dilution refrigerator. Cryogenics 12^, No. 6, 469 (Dec. 1 972).

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 (Oct. 1975).

Arp, V. and Frederick, N. V., Helium Research in Support of Superconducting

Power Transmission, Annual Report, NBSIR 77-853, CONS-3800-1 (Feb. 1977).

Also, private communication.

Arvidson, J. M. and Brennan, J. A., ASRDI Oxygen Technology Survey, Vol. Ill:

Pressure Measurement, NASA SP-3092, 1975.

Bailey, R. L., Fifth International Conference on Magnet Technology, Frascati

(1975).

Bald, W. B., BNL Report 15805 (1970).

Bejan, A., presented at the 1976 Applied Superconductivity Conference.

Bejan, A., Cryogenics 97 (1 977).

Brennan, J. A., Stokes, R. W., Kneebone, C. H. and Mann, D. B., An Evaluation

of Selected Angular Momentum, Vortex Shedding and Orifice Cryogenic Flowmeters, NBS Tech. Note 650, March 1974.

Butler, A. P., James, G. B. Maddock, B. J., and Norris W. T., Int. J. Heat Mass

Transfer 13, 105 (1970).

Chester, M., Choudhery, S. C, Jones, B. K., and Williams, D., Simple liquid

helium pressure transducer, Rev. Sci . Instrum. 39, No. 5, 719-722 (May 1968).

Daney, D. E., McConnell, P. M. and Strobridge, T. R., Advances in Cryogenic

Engineering 18, 476 (1973)

Dean, J. W. and Flynn, T. M., Temperature effects on pressure transducers, ISA

Trans. 5, 225-252 (1966).

DeSorbo, W., Magneto-optic pressure measuring device, U.S. Patent 3,564,924

(Feb. 23, 1971).

Friedley, J. C, Mangan, J. L., and Kroeger, P. G., Final Report NAS8-21014,

NASA Marshall Space Flight Center, Huntsville, AL (1967).

Fujii, Genshiro and Nagano, Hiroshi, Superconducting pressure gage at high

pressure and low temperature. Progress in Refrigeration Science and Technology, 1, Proc. XIII International Congress of Refrigeration, Washington,

DC, 1971. AVI Publishing Co., Inc., Westport, Conn., pp. 499-502.

Gonano, Roland, In situ vapor pressure measurement for low temperature thermometry. Rev. Sci. Instrum., £!_, No. 5, 716-719 (May 1 970).

43
Green, W. J. and Saibel, E., Advances in Cryogenic Engineering U, 138 (1968).

Halama, H. J. and Bamberger, J. A., paper presented at the Sixth Symposium on Engineering Problems of Fusion Research, San Diego, CA (1975).

Hayakawa, K. K., Cryogenic Pressure Measurement Technology, High Pressure Flange

Seals, Hydrogen Embrittlement of Pressure Transducer Material, Close Coupled

vs Remote Transducer Installation and Temperature Compensation of Pressure

Transducers, Vol. I. Space Div. North American Rockwell, SD 72-SA-0156-1

(Oct. 1972).

Henning, C. D. , presented at the 1976 Applied Superconductivity Conference.

Hilal, M. A. and Boom, R. W., presented at the 1976 Applied Superconductivity

Conference.

Hoenig, M, 0., et al., presented at the Fusion Power Conference, San Diego, CA

(Nov. 19, 1975).

Hoenig, M. 0. and Montgomery, D. B., IEEE Trans. Magnetics MAG-11 , 569 (1975).

Horn, Leon., ISA Preprint No. 13.3-4-65, Oct. 1965.

Itskevich, E. S., Voronovskii, A. M., Gavrilov, A. F., and Sukhoparov, V. A.,

High pressure chamber for pressures up to 19 kbar intended for working at

helium temperatures. Cryogenics 7^, No. 6, 359-61 (Dec. 1 967).

Iwasa and Apgar, presented at the 1976 Applied Superconductivity Conference.

Jackson, J., Cryogenics^, 103 (1969).

Jones, M. C, and Petersen, R. G., ASME Publication 74-WA/HT-24 (1975).

Kadi, F. J. and Longsworth, R. C, Final Report to ERDA under contract

E(ll-l)-2552 (1976).

Kinzie, P. A. and Murphy, J. A., Final Report, June 1967. Prepared for Georce C.

Marshall Space Flight Center, Huntsville, Alabama, by Rocketdyne Division of

North American Rockwell Corp. under contract No. NAS8-20394.

Krishman, V. S. and Friedley, J. C, Proceedings of the 5th International Heat

Transfer Conference, Tokyo (1974).

Lubell, M., private Communication (1977).

Ludtke, P. R., NBSIR 75-816 (1975).

Maddock, B. J., James, G. B. and Norris, W. T., Cryogenics 9, 261 (1969).

Manatt, S. A., Wapato, P. G., Stanko, J. and Baumgartner, J. P., Report prepared for ERDA under contract E904-3)-1061 (1976).

44
Mann, D. B., NBS Tech. Note 154 (1962).

Mann, D. B., ASRDI Oxygen Technology Survey Volume VI: Flow Measurement Instrumentation, NASA SP-3084, 1974.

Maulbetsch, J. S. and Griffith, P., Proceedings of the Symposium on Two-Phase

Flow Dynamics, Vol, 1_, 799 Euratom Report 4288e, Brussels (1 969).

McCarty, R. D., NBS Tech. Note 631 (1972); J. Phys. Chem. Ref. Data 2, 923 (1973).

McConnell, P. M., NBSIR 73-316 (1973).

Morpurgo, M., Cryogenics 1_7, 91 (1977).

Niinikoski, T. 0., Rev. Sci . Instr. 43, 430 (1972).

Quack, H., Paper CD-6 presented at the Cryogenic Engineering Conference, Boulder,

Colo. (1977).

Roder, H. M., ASRDI Oxygen Technology Survey, Vol. V: Density and Liquid Level

Measurement Instrumentation for the Cryogenic Fluids Oxygen, Hydrogen, and

Nitrogen, NASA SP-3082, 1974.

Sample, H. H., and Rubin, L. G., Instrumentation and Methods for Low Temperature

Measurements in High Magnetic Fields, Proc. of the First Biennial Symposium

on Cryogenic Instrumentation, Houston, Texas (Oct. 1976).

Sixsmith, H., and Giarratano, P. J., Rev. Sci. Instr. 41_, 1 570 (1970).

Sixsmith, H., private discussions at BNL (1976).

Smelser, P., Adv. in Cry. Engr., Vol. 8, 378-386 (1963).

Snyder, N. S., Cryogenics 10, 89 (1970).

Sparks, L. L., ASRDI Oxygen Technology Survey, Vol. IV: Low Temperature Measurement, NASA SP-3073, 1974.

Stankiewicz, J. and White, R. L., Carbon resistors as pressure gauges. Rev. Sci.

Instrum., 42, No. 7, 1067-1068 (July 1971).

Stekly, Z. J. J. and Zar, J. L., IEEE Transactions NS-12 , 367 (1965).

Steward, W. G., Submitted to International J. of Heat and Mass Transfer (1977).

Thurston, R. S., Rogers, J. D. and Skoglund, V. J., Advances in Cryogenic

Engineering, Vol. 12, 438, Plenum Press, NY and London (1967).

Todoriki, I. and Agatsuma, K., Fifth International Cryogenic Engineering

Conference, Kyoto (1974).

Tsukamoto, 0. and Kobayashi, S., J. Appl . Phys. 46 (1975).

Walstrom, P. A., Instrumentation in the Cryogenic Industry, Vol. 1, Houston, TX,

Oct. 11-14, 1976.

45
Wilson, M., presented at 1976 Applied Superconductivity Conference.

AFOSR Air Force Office of Scientific Research

ANL Argonne National Laboratory

BNL Brookhaven National Laboratory

DEMO . a successor to EPR, nominally planned for 1995

DMFE Division of Magnetic Fustion Energy

EPR Experimental Power Reactor, nominally planned for 1985

FCCS Forced circulation cooling system

GA General Atomics Company

HEP High Energy Physics (program of ERDA)

ISABELLE Intersecting Storage Accelerator at BNL

ITR Ignition Test Reactor

LASL

' Los Alamos Scientific Laboratory

LCP Large Coil Program at ORNL

LCPTE Large Coil Program Test Facility

LLL Lawrence Livermore Labotatory

LNG :; Liquefied Natural Gas

MFE Magnetic Fusion Energy

MHD Magnetic hydrodynamic (program or power generation)

NBS National Bureau of Standards

OCR

' Office of Coal Research

ORNL Oak Ridge National Laboratory

PB Pool boiling

PPPL Princton Plasma Physics Laboratory

SLAC - - Stanford Linear Accelerator Center

SPTL . Superconducting Power Transmission Line

TFTR Tokomak Fusion Test Reactor

TNS The Next Step, a reactor prior to EPR

TSS thermal sensitivity shift (of transducers)

TZS thermal zero shift (of transducers)

U. Wise. University of Wisconsin

UWMAK University of Wisconsin Tokomak

46
Part One, Appendix A

Agenda from the

WORKSHOP ON STABILITY IN SUPERCONDUCTING MAGNETS

JULY 25-29, 1977

NATIONAL SECURITY AND RESOURCES STUDY CENTER

LOS ALAMOS SCIENTIFIC LABORATORY

TECHNICAL PROGRAM

MONDAY

Chairman: H. Laquer

8:30 Welcome to Los Alamos H. Agnew

Introductory Remarks W. V. Hassenzahl

9:00 Review of Cryostabi 1 ity Z. J. J. Stekly

9:45 Stability in Terms of Minimum Propagating

Zones and Disturbance Spectrum S. L. Wipf

10:40 Coffee

11:00 Magnetic Instabilities and Quenching

Currents in Superconducting Composites B. Turck

11:30 Theoretical Research on Stability

at ORNL L. Dresner

12:30 Lunch

Chairman: S. L. Wipf

1:30 Modus Operandi S. L. Wipf

2:00 What More is There to Know? M. N. Wilson

2:40 Tea

3:00 Working Session (Contributed Papers)

- The Disturbance Spectrum

Role of Acoustic Emission in Monitoring Stability

in Superconducting Magnets 0. K. Mawardi

Mechanical Perturbation Studies

at ANL .S. T. Wang

Stability and Mechanisms Which Give

Rise to Instabilities in Superconducting Magnets J. H. Murphy

Measurement of Frictional Heating Y. Iwasa

4:00 Working Session (Seminar)

- Can We Predict a Quench?

5:15 Bus Leaves (Bus returns in time for Cocktail Party)

6:00 - 7:30 Cocktail Party and Buffet,

National Security and Resources Study Center

47
TUESDAY

Chai rman 8:30

9:15

10:00

10:20

12:00

12:30

Chairman:

1:30

3:30

3:50

5:30

John D. Rogers

Facing the Stability Problem in

Magnet Construction J. Purcell

Development and Operating Experience

on NbTi and Multi filamentary Nb^Sn

Magnets D. N. Cornish

Coffee

Working Session (Panel Discussion)

- Stability in Large Coils:

Engineering Implications

Panel

R. D. Bradshaw J. H. Murphy

D. N. Cornish J. Purcell

H. Desportes Z. J. J. Stekly

S. T. Wang

Stability Considerations and Safety Factors in

Large Systems with Thermal Runaway Problems

(Example: Fission Reactor) To Be Announced

Lunch

R. I. Schermer

Working Session (Contributed Papers)

- Structure Versus Stability -^^ Pulsed Dipole Performance and Coil

Structural Constraints W. S. Gilbert

The Fermi lab Tevatron Magnets ..M. Kuchnir

The Isabelle Coils A. D. Maclnturff

Design Tradeoffs between Stability and

Structural Requirements C. L. Linkinhoker

Stability of the 300-kJ METS Magnets J. J. Wollan

Tea

Working Groups (Parallel):

Coil Monitoring (Wilson)

Heat Transfer (^^)

Other groups to be established based on the interests of the attendees.

Bus Leaves

48
WEDNESDAY

Chairman: J. J. Wollan

8:30 Review of Heat Transfer Including

Transient Effects V. Arp

9:30 Working Session (Contributed Papers)

- Response of Coil to Disturbances:

Heat Transfer

Transient Heat Transfer Measuren^ents Y. Iwasa

Superconductor Stability in Narrow

Cooling Channels J. W. Lue

10:30 Coffee

10:50 Working Session (continued)

Stability as a Problem of Heat Transfer. . .G. Claudet

Recovery of a Cryostable Magnet Following

a Mechanical Perturbation L. R. Turner,

Boiling Heat Transfer Induced Instabilities in

Large Superconducting Magnets J. W. H. Chi

Cryogenic Stability Using Supercritical

Helium as a Coolant../. M. A. Hilal

12:30 Lunch

Chairman: M. N. Wilson

1:30 Working Session (Contributed Paoers)

- Hollow Conducior Systems

Analysis on the Stability of Hollow ForceCooled Superconductors 0. Tsukamoto

A Computer Simulation of fJormal

Zone Propagation J. K. Hoffer

Forced Flow Conductor Development

at ORNL J. R. Miller

Cryostabi 1 i ty in Forced Flow Cooled

Cabled Superconductors M. 0. Hoenig

Performance of SIN Muon Channel G. Vecsey

3:20 Tea

3:40 Working Session (Seminar)

- Can We Improve Stability?

Higher Temperatures for Better

Stability D. Atherton

5:15 Bus Leaves

49
THURSDAY

Chairman:

8:30

10:00

10:20

12:30

Chai rman

1:30

!I 3:10

:| 3:30

i1 4t30

j

' '

'

' Chairman

;i ;vv\ 8:30

;/ 10:10

10:30

i 12:30

:! . 1:30

3:10

3:30

4:45

K. D. Williamson

Working Session (Seminar)

- Can We Learn From Coil Failures? L Operational Experience of Existing Super- } conducting Magnet Systems S. Y. (David) Hsieh

Coffee \ Working Groups (Parallel)

Lunch

S. L. Wipf i! Working Session on Terminology and Definitions i Introduction A. F. Clark

. Small Discussion Groups

(Group Leaders: Clark, Dresner, Laquer, Oberly,

Strauss, Wilson, Wipf)

Tea

Synthesis of Definitions (All Discussion Groups)

Buses leave for ai fresco supper at

Bandelier National Monument

FRIDAY

W. V. Hassenzahl

Review of Working Group Results ! Coffee

Consensus (?) on Terminology and Definitions

Lunch

Sunmaries of Conference • Monday: S. L. Wipf

j Tuesday: J. D. Rogers i Wednesday: H. Laquer

Tea

General Conclusions: Wilson '

Bus Leaves

50
NOTES CONCERNING THE PROGRAM

The working sessions provide a forum to present the most up to date reports

of current work relating to stability. They range from sessions for prepared

contributed papers to the less formal seminars for plenary discussions, with

panel sessions somewhere in between. All participants are encouraged to

contribute in the working sessions.

Suggested topics to be dealt with in seminars include, but are not limited

to» the following.

V. Can we predict a quench:

Operational prediction versus design prediction.

How can coils be monitored?

Can something be done after a woul d-be-quench has already started?

What is the use of distinguishing between quench current below and above

short sample critical current?

What information can be obtained from test coils?

Can we use the propagation velocity?

2. Can we improve stability?

High or low operating current versus stability. The role of electrical

insulation.

Advantages and disadvantages of hollow conductors and forced flow.

Structural material versus copper.

Fiddling around with parameters (spec, heat; heat transfer; thermal

conduction; mechanical parameters; superconducting parameters.)

Is training useful? Can it be achieved without quenching?

3. Can we learn from coil failures?

Failure (quenches) of large coils: are they related to stability?

Can we gain any criteria on disturbances related to coil size?

Can we define a safety factor? If so, how does it depend on size?

Small working groups will be formed on Tuesday afternoon to discuss special

topics under the leadership of a group chairman. These groups will work informally during the week and on Friday morning the group chairmen will sum- marize the results. Suggested topics for the working groups include the

following:

Coil monitoring

Heat transfer

1.8 K operation of coil systems

Low field-high frequency or pulsed operation

Propagation velocity

Sources of disturbances in coils
The work on Terminology and Definitions will begin by dividing existing

and new terms into three classes:

. Class 1. Terms which are useful and well defined.

Class 2. Terms which are useful but not well defined.

Class 3. Terms which are either not useful, confusing, or

to be avoided.

We want to establish a consensus on the definitions of the terms in

Class 1

Please bring relevant material.

52
Part One, Appendix B

Attendees at the

rtorkshop on stability in Superconducting Magnets

Joiy 25-29 , 1977

Lui5 klh>:QS :iCIENTIFIC LABCRRATORY

Los Alamos, NM 875^5

[A] VISITUnS

Erik Acara

Airco, Inc.

IQO riountain Ave.

Murray Hill, NJ 07974

(20 1)46^-2400 Ext. 35 1 FTS fa-2 12-344-8580

John Alcorn

General Atomic

P.O.Box. ttlbOd

San DiefeO, CA 92138

Vincent Arp

Cryogenics Division, NES

Bouluer, CO 80302

(303)4i>y-1000 Ext. 3422

FTS 8-323-3422

Bob Braasiiaw

Convair Division

General Dynamics

San Diego, Ck

8^1-89uu Ext. 2779

/

Wilkie Y. K. Chen

General Atoiaic

P. 0. Box 8150(5

San Diego, CA 92138

53
John W.H. Chi

Fusion Power Systems Dept.

Westinghouse Electric Corp.

Large Site, P.O. Box 10864,

Pittsburgh, PA l'523b

(412)892-5600 Ext. 5486

Al Clark

Cryogenics Division

National bureau of Standards

Boulder, CO 80302

FTS 8-323-3253

G . Claudet

Centre d'Etudes Nucl^aires de Grenoble

Service des Basses Temperatures

85X

F38041 GhhNOBLE-CEDEX (FRANCE)

D.N. Cornish

Lawrence Liverniore Laboratory

P.U. Box 506

Livermore, CA 94550

(415)447-1100 Ext. 4663

FTS 457-4663

Henri Desportes

CEA/SACLAY - DPh/PE-STIPE

B.P. n'^ 2 - GIF/sur/YVETTE (FRANCE)

941 .80.00

Lawrence Dresner

Oak Riage National Laboratory

P. 0. Box Y

Oak Ridge, TN 37830

(615)483-8611 Ext. 3-5476

54
'Ailliam S. GilDert

Lawrence Berkeley Laboratory - Bldg.

University of California

berKeiey, CA ^4720

(415)0^3-2740 Ext. b37^

FI^ 0-451-6372

Joe Heim

F e r in i 1 a D P. 0. box 500

Batavia, IL o5010

Honamea A. Hilal

541 LKt5, 15UU Jonnson Drive

Maaison, UI 53705

(bOo)236-b Ibo

(Office) 2b3-23b8

hit Che 11 0. Hoenig

Francis Bitter Magnet Laboratory MIT

170 Hloany St., Carr.oridge, l-iA 0213^

(617)253-5503

5.Y. (Da via) hsieh

Blag. 12':^ brookhaven National Laboratory

Upton, NY 11973

^ . Iwasa

hoom Nrtl4-3212

MIT

CaniDridge, hA 02 139

(617) 253-5540

Douglas Koop

Alcoa Laboratories

Alcoa Center, PA 15006

(41^) 33b'-6D:3l

55
Moyses Kuchnir

Fermilab

P.O.Box 50

Batavia, IL 50510

FTS 8-370-3368

Henry L. Laquer

Consultant

Ht.l Box 445

Espanola, NM 87532

(505)753-3788

J. W. Lue

Oak Riage National Laboratory

P. 0. Box Y

Oak Ridge, TN 37830

(515)483-8511 Ext. 3-5474

FTS 8-850-5476

0. K. Mawardi

Case Western Reserve University

University Circle

Cleveland, OH 44106

(216)368-4571

FTS 8-292-457 1

A. D. hclnturff

Brijokhaven National Laboratory

Upton, HY 11973

FTS 8-654-4845

John h. Miller

Oak Ridge National Laboratory, Box Y Blag. 9204-1

Oak Ridge, TN 37830

(615)483-8611 Ext. 3-7733

John H. Murphy

Westinghouse Reserch Laboratories,

Beulah Roaa

Pittsburgh, PA 15235

(412)256-5163

Ctiarles Oberly

V»right-Patterson AFb, OH 45433

FIS 5-7 7 2-

56
J.5hn Puree 11

ueneral Atomic

P . U . box 0 1 oOc

ban L 1 e g o , C A 9 2 1 3 o

Guncer Hies

InstiTLut fUr lecr.niscne PnysiK

Posufach 3640

D75 Karlsrune

•rtest Germany

P. Seyi'erl

Centre a'Etuaes Nucliaires ce Grenocle

Service aes Basses Te:np^ratjres

o5a

F3oG41 G.iLriOELE-CELLX (Fr.nNC£)

(76)97 .41 . 11

Z. J. J. Stekly

nag-netic Corporation oi A.Tierica

179 bear hill R o a a

rt alt nam , N;A 02 154

( 0 17 ) 090-4242

Bruce Strauss

FNAL

FTS 0-37^-3671

Michael J. Superczynsxi

Coae 2722

U . 5 . M a V a 1 Snip n ct i: Center

Annapol i s , l-iL 2 14 0 1 (301)267-214-

Usami Tsukamoto

Faculty of Engineer iiig

Yokonama National University

Tokiwacai 15c, hOGOgay£-:<u

Y OK Oil aina, Japan

D err. arc Turck

Clm/;:)«CLAY - LPn/Pc-STIPc:

D.P. no. 2 GIF-s jr- 1 V ETTc. (FnANCi:)

9^ 1 .dO .00

Larry n . i u r n e r A r £, 0 n n e a t i o n a x LcL, oratory, Blag. 3o2

57
Argonne, IL 60439

(312)739-7711 Ext. 3^37

FTS - 388-3537

George Vecsey

SIN

Switzerland

Wayne Vogen

4 12 Hhb

University of California

Berkeley, CA 94720

(41b) 642-3804

M. S. Walker

Interniagne t i cs General

Charles Industrial Park

P.O. Box 566

Guilderland, NY 12085

(518) 456-5456

S. T. Wang

Argonne National Laboratory, Bldg. 362

Argonne, IL 60439

(312)739-77 11 Ext. 4156

FTS 8-388-4156

fiei'nha-r-dt L. Willig

Magnetic Engineering Associates (MEA)

247 Third St.

Cambridge, MA 02142

(617)868-2550

Martin N. Wilson

Francis Bitter National Magnet Laboratory MIT

170 Albany St.

Cambridge, MA 02139

(617)253-5547

FTS - 8-835-5547

58
LB] LASL PAHTICIPANTS

Los Alamos Scientific Laboratory

P. 0. Box 1663 Itn

Los Alamos, NM 875M5

D. B. Colyer

Q-10

(505) 667-7672

FTS 8-843-7872

E. F. Hammel

ADE

(505) 667-7300

FTS 8-843-7300

W.V. Hassenzahl

g-10

(505) 667-5026

FTS 8-843-5626

J.K. Hoffer

Q- 10

(505) 667-4U49

FTS 8-843-4049

W.E. Keller

Q-10

(505) 667-4838

FTS 8-843-4838

J.D.G. Lindsay

CTR-9

(505) 667-4404

FTS 8-843-44U4

J.D. Rogers

CTH-9

(505) 667-5427

FTS 8-843-5427

H.I. Schermer

g- 1 U

(505) 667-6093

FTS 8-843-6093

R.D. Turner

g-10

(505) 667-5458

FTS 8-843-5458

R.W. Warren

CTR-9, 'westinghouse

(505) 667-65bl

FTS 8-843-656
K . D . Williamson

CTri-9

(505) 567-4096

FTS b-bM3-4U96

5 . L . W i p f

g- 10

(305) 667-7960

FTS 8-843-7960

J.J. U t3 1 1 a n CTh-9

(505) 667-6636

FTS 8-843-6686
NBSIR 78-877, Part Two

HELIUM TRANSIENT HEAT TRANSFER

October 1 , 1977

W. Gene Steward

61
1. INTRODUCTION

Transient helium heat transfer experiments were carried out at the NBS Cryogenics Division for two years under sponsorship of the Air Force Office of

Scientific Research. In these tests thin carbon films deposited on flat quartz

substrates served both as fast time response heaters and thermometers. The first

phase of these tests, transient heat transfer to a static helium bath, is covered

in detail in Appendix A which is a preprint of a paper to be published in the

International Journal of Heat and Mass Transfer. A summary of these tests is as fol lows

SUMMARY OF TESTS

Temperature Rise and Heat Transfer Coefficient as a Function of Time

Static Helium Forced Flow

Heat flux range .05 to 20 W/cm

Ambient pressure

Elevated pressure to 0.3 MPa

Variable orientation

Repeated pulses

Transient decay

to 1 MPa

XX (in progress)

X

X

X

Forced convection transient heat transfer tests, which began under the AF

funding, were continued under ERDA funding, as was the development of computer

plotting of the data. Currently, work has started on a new type of heaterthermometer which employs heat transfer surfaces other than carbon. Looking

toward a method of obtaining transient data in the form desired for stability

analyses we have worked out a preliminary design for an electronic system to

vary heater power with time in such a way as to produce scheduled rate of sur- face temperature change (either rising or falling).

The final section of this report discusses plans for future work.

2. FORCED CONVECTION EXPERIMENTS

The forced convection experimental apparatus is shown schematically in

figure 1. The objectives of the apparatus design are to provide transient heat

transfer data under conditions of flow and heating similar to those of forced

flow cooled superconductors. Flat plate heaters of 0.5 cm wide x 1.0 cm long

(in the flow direction) are inserted into a rectangular flow channel 0.15 cm deep. A length of uniform channel approximately 18 times the channel depth

precedes the test heater to ensure a fully developed momentum boundary layer.

The pre-heater shown is intended to establish the thermal boundary layer and

thus avoid entrance effects. Power is supplied to the pre-heater at approximately

the same power density as in the test heater. Further specifications are:

1) Liquid helium maximum flow rate with test loop installed -- 18 g/s.

(maximum Reynolds number, Np . in the test section = 1.7(10)^)

62
2) Heater area - 0.1637 cm , pre-heater area = 0.322 cm .

3) Heater surface material -- vapor deposited carbon.

All of the instrumentation and data acquisition equipment are the same as

in the static helium tests with the addition of the pump, variable frequency

power supply, and flowmeter frequency counter. Data were recorded on magnetic

tape and plotted by computer.

Test results shown in figures 2 through 9 are for heat fluxes of 0.1, 0.5,

2 5, and 10 W/cm , usually with nominal mass velocities of 0, 40, 120, and 240

2 5 5 g/s*cm which correspond to liquid Reynolds numbers of 0, 2.8(10) , 8.3(10) , and

1.7(10) . Reynolds number is shown as the parameter here because some form

of Reynolds number is the principle flow related correlating factor for forced

convection heat transfer in steady state.

At the lowest heat flux of 0.1 W/cm~ the positive slope portion of the temperature rise curve in figure 2 is where fluid conduction and Kapitza resistance

are the most important factors in heat transfer (see Appendix A). It is apparent,

however, that convection also plays an important role in the conduction and

nucleate boiling region, since there is a significant temperature reduction (and

heat transfer coefficient increase in figure 3) as flow increases. It is also

apparent that the increasing turbulence due to increasing flow leads to nucleate

boiling at lower surface temperature (superheat) and correspondingly higher heat

transfer coefficients. Figures 4 and 5 show the temperature rise and heat

2 transfer coefficients at 0.5 W/cm (peak nucleate boiling heat flux in steady

2 state occurs between 0.5 and 1 W/cm ). The steady state nucleate boiling tem-

-4

perature and heat transfer coefficient are reached by 10 s, the conduction period

_ c (Appendix A) being nearly finished at the earliest measurement time of 2(10) s. Flow has little effect at any time. Apparently the bubble nucleation effect is

strong enough at this heat flux to overshadow the convection effect on heat

transfer. The ineffectiveness of forced convection at the high nucleate boiling

heat flux was noted previously by Giarratano [1] in steady state helium heat

transfer experiments.

In figures 6 through 9 the heat flux is above the steady state nucleate

boiling limit, but the final steady state film boiling temperatures and heat transfer

coefficients are reached only after one second or more of heating. The level

- 5 -4

region extending from 2(10) to 2(10) s in figures 6 and 7 is the "transient

nucleation" period in which nucleate boiling exists briefly. As in nucleate boiling

of figures 4 and 5 forced flow has little effect in the transient nucleation

period; however, the curves of constant flow rates begin to diverge and the

_ 3 difference becomes significant by 10 s. Finally, in steady state, the heat

2 transfer coefficients at mass velocities of 235 to 238 g/S'cm are factors of

four to five times that of a static fluid. The steady state variation of heat

2 transfer coefficient at 5 W/cm with Reynolds number is shown in figure 10 along

with a correlation of Hendricks, et al . [2] developed from liquid hydrogen data.

63
Heat transfer coefficients predicted by the Hendricks correlation have been

reduced by 20 percent since that reduction has been observed previously for

rectangular channels heated on one side only [3]. This correlation is not

entirely satisfactory since it predicts h = 0 at H^^ ^

= 0, so that its applicability at best is limited to high flow rates. Other steady state correlations

are being investigated. Further flow tests will be carried out at pressures up

to 1 MPa.

3. PLANNED CONTINUATION OF TRANSIENT HEAT TRANSFER STUDIES

Thin carbon films vapor deposited on quartz substrates act as both heater

and thermometer for all the experiments to date. These films have the advantage

of extremely small mass relative to the surface area, hence, low heat capacity

- 5 and fast time response, a thin cross section (5(10) cm) which makes thermal

gradients through the film very small, and absence of any possible temperature

difference between the heater and thermometer since one film performs both

functions. The films are sensitive as thermometers and have proven to be stable

in spite of severe thermal shocks.

While the carbon film heaters have many advantages, the results obtained

in the first phases of this study indicate that some materials on the surfaces

of superconductors may have heat transfer characteristics substantially different

from the carbon film. Figure 11 shows that the kind of material in contact with

the helium is particularly important in the early stages of transients when

kapitza resistance is the predominant thermal resistance (see Section 3 in

Appendix A). This figure compares measured heat transfer coefficients with those

estimated for a hypothetical material with negligible Kapitza resistance. Since

this resistance is believed to vary inversely as the third power of temperature,

_3

the two curves of figure 11 essentially merge after approximately 10 s when the

temperature begins to rise rapidly. It should be emphasized that other surface

material and surface condition effects exist even in steady state boiling and

convection heat transfer. However, these other surface effects which have been

observed in steady state heat transfer, are so unpredictable that no attempt has

been made to illustrate them in figure 11. Data of reference 7 in Appendix A

would place heat transfer coefficient curves for metals somewhere between the

two curves of figure 11; however, the Kapitza resistance is not well enough established to allow a good approximation for a particular metal.

To establish helium transient heat transfer coefficients for metals and other

surface materials of interest, experiments would be conducted in a manner similar

to the present experiments, but with metallic and other surfaces used instead of

carbon. The considerations for the design of test heaters are as follows:

1) Extremely thin vapor or vacuum deposited films do not necessarily have

the same heat transfer characteristics as the bulk material even though the

chemical composition may be the same. Therefore, the heat transfer surface

should be of significant thickness capable of being prepared in the same way

as bul k material s

64
2) Metals or other materials of interest generally are not suitable as resistance thermometers, particularly near liquid helium temperature. For

this reason it will be necessary to separate the thermometer from the heater.

Carbon films seem to be the best choice as thermometers because of their

fast response, and because their small cross sections relative to their

surface area minimizes error due to heat conducted from the electrical

leads, and maximizes the thermal coupling to the heater surface.

3) With the thermometer electrically insulated from the heat transfer

surface it is not suitable to use the same film as thermometer and heater

because a large heat flux through the electrical insulation would cause

an unacceptable temperature difference across the insulation.

4) Because of their low electrical resistance, materials such as copper would not function well for resistance heating. This means that

the heater must also be separate and electrically insulated from the heat

transfer surface. Other materials with higher resistivity, such as stainless steel, could double as heater and heat transfer surface; however, the

thickness would have to be very small to provide sufficient electrical re- sistance and prevent excessive temperature differences across the thickness.

The configuration of the proposed test heaters which takes these factors

into consideration is shown in figure 12. The configuration shown is for a material

of low electrical resistance and high thermal conductivity such as copper or aluminum. The carbon film below the surface would be divided into zones of

which the larger area would be devoted to resistance heating and the smaller

area would act as a thermometer. Lateral direct conduction of heat between the

surface and the thermometer is small, but thermal contact between the surface and

the thermometer is large. Since the heat flux to the thermometer is small, the

measurement error should be small. The size of the space between heater zones which is devoted to thermometry should be made as small as possible to reduce the

temperature non-uniformity at the surface at high heat flux. The surface material

must be thick enough to distribute the heat uniformly, but not so thick as to

produce large temperature differences across the thickness. For example, a copper

surface material 0.05 cm thick has an approximate thermal time constant of

T = 5.2(10"'^ s at 4 K, and

T = 2.4(10)"^ s at 10 K.

For a 0.1 cm thermometer gap in the heater, and a surface-to-helium heat transfer coefficient of 3 W/cm K an estimate of the maximum deviation in the surface

temperature at the gap is.

= 0.96.

65
The temperature difference due to conduction of heat through a 0.05 cm thickness of copper may be as high as 10 percent of the surface-to-helium temperature difference being measured. Though this is a significant difference it

is possible to calculate the correction with reasonable accuracy.

Surface materials suggested for testing are copper which may be cleaned,

oxidized, tinned, or coated with a typical insulating material such as enamel.

Other materials may be tested as consistent with time and funding.

Future tests planned include the following:

a. Forced flow near critical and supercritical pressures with the present

carbon film heaters.

b. Use of metal and other heat transfer surfaces.

c. Transient tests in which the heater power is controlled in such a way

as to produce a uniform rate of temperature rise or fall (rather than

constant heat flux)

d. Further tests in which temperature recovery following a pulse is measured.

e. Shorter power supply rise time so that steady power is reached and con- stant heat flux measurements may be started at earlier times Ips).

;, This is of particular interest in the design of fast response superconducting switches, and for better understanding of the conduction heat

transfer period and Kapitza resistance.

4. NOMENCLATURE

specific heat at constant pressure, j7(g-K) mass flow rate per unit cross sectional area (or mass velocity),

g/(s'cm^)

channel height, cm, heat transfer coefficient W/(cm *K) "

thermal conductivity, W/(cm*K)

Reynolds number based on liquid properties - Gh/y

surface temperature if heating were uniform, K bulk liquid temperature, K temperature at greatest deviation from T, K dynamic viscosity, g/cm's)

density, g/cm"^

thermal time constant, s

5. REFERENCES

1. Giarratano, P. J., Hess, R. C. and Jones, M. C, Forced convection heat

transfer to subcritical helium I, Advances in Cryogenic Engineering, Vol. 19,

Plenum Press, New York, (1974).

Re,L

min

66
2. Hendricks, R. C, Graham, R. W., Hsu, Y. Y., and Friedman, Experimental

heat transfer and pressure drop of liquid hydrogen flowing through a heated

tube, NASA TND-765 (1961).

3. Rohsenow, W. M., and Hartnett, J. P. (ed.) Handbook of Heat Transfer , McGraw

Hill, Inc. (1973), Ch. 7.

Fifjiire Captions

Figure 1. Schematic of the forced convection transient helium heat transfer

apparatus.

Figure 2. Effect of helium flow on heater temperature transient rise for a heat flux of 0.1 W/cm^.

Figure 3. Effect of helium flow on transient heat transfer coefficient for a heat flux of 0.1 W/cm^.

Figure 4. Effect of helium flow on heater transient temperature rise for a heat

flux of 0.5 W/cm2.

Figure 5. Effect of helium flow on transient heat transfer coefficient for a heat flux of 0.5 W/cm^.

Figure 6. Effect of helium flow on heater transient temperature rise for a heat

flux of 5 W/cm2.

Figure 7. Effect of helium flow on transient heat transfer coefficient for a heat

flux of 5 W/cm^.

Figure 8. Effect of helium flow on heater transient temperature rise for a heat

flux of 10 W/cm2.

Figure 9. Effect of helium flow on transient heat transfer coefficient for a heat

flux of 10 W/cm2.

Figure 10. Effect of Reynolds number on steady state heat transfer to liquid

helium at 4 K with a heat flux of 5 W/cm"^.

Figure 11. The effect of heat transfer surface Kapitza resistance on transient

helium heat transfer at a heat flux of 2 W/cm^.

Figure 12. Configuration for transient heat transfer tests with surfaces other

than carbon films.

67
r

\.V-)\ \\.) 1 i KI.U: M ^ K •! U.T 1 .\ i (l.tULU; a 1 KOoi. N otilt: 1,1) OJ .->riC)U .\)

1 l:u f.ixe

Ol) 1 KK l.IQbllJ

HKLIUM

i;}-;si:k \ on;

'fArMOSl'Hi-;!UC

JXXKK CELL

(['1U:SSI_-R F".S O t MPa)

KST SLCTIOX

'i;mi' l\lk

KOLXD-TO-R E;CT ANGULAR

r R A xs I f lo X s r:c t io x

figure 1. Scliematic of the forced convection trruisient liclium lieat transfer

appiiratus.
o

Q£ CM E B CO

01 a.

3 CT> Ol O r»» —« O — CM

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69
70
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71
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72
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01 o

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c~

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r~

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-(->

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(O

c:o3oE

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.£= 3

4- ir> O

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zS'.'a 3c:*^?a3c

73
74
a

0 3 0

U • 3 > > >

-3 S 3 2 2 5

S: + o a

\

•m

<5» I"

75
QD tiO ^ o -r

U.

—

—

^ (J V lU % >>> J ,n M « = "Si d TM TM

S, 2 2 2 QQQQQaQQaa

CDO * I

CD

CO

_ o » O

+-> fd

&-oc

<u

•r-U

»r

»4- O)oo

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"rTM

t/) C

fO

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76
77
78
HI': AT 'J RANSFl'lU SUJM--ACI':

l''igur(^ IZ. l'"x])crimfntaT ar r anf^emont for t.ransic:nt hctit transftir tests on

sin-faces c-lIkm; UiZin c arboji film.

79
Part Two, Appendix A

(preprint of paper to appear in the International Journal

of Heat and Mass Transfer)

TRANSIENT HELIUM HEAT TRANSFER

PHASE I - STATIC COOLANT''^ .

W. G. Steward

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

I Transient heat transfer data have been obtained for flat heating surfaces in static liquid and supercritical helium. MeasureI -5

1 ments start 2(10) s after step power inputs, and cover a heat

' 2 flux range of 0.05 to 20 W/cm , pressures from 0.09 to 0.3 MPa, and

f four different heater orientations. Initial heat transfer coefficients, being limited primarily by the Kapltza resistance, are 10

i to 100 times greater than steady state, and the time to reach steady

i state varies from 10 ^ s to 1 s. For heat flux below the steady

state peak nucleate boiling limit the temperature follows calculations based on pure conduction to the steady state nucleate

boiling level. Above that limit the transient conduction period

leads to an apparent metastable nucleation period followed by a

transition to film boiling.

I

. 1. INTRODUCTION

I Many superconducting devices employ large power pulses or are subject to

power transients due to overloads or other causes. It is usually desirable,

and often imperative, that the superconductor not quench during the pulse.

In other cases one is concerned about the temperature recovery following heat

release from flux jumps or micromechanical movements. Fortunately, heat transfer from the superconductor surface to the helium coolant can be many times

This work was carried out at the NBS Cryogenics Division, Boulder, CO under

tlie sponsorstii p of the Air Force Office of Scientific Research, .Wash i n;.', t on

,

DC. Continuation of the work is sponsored by ERDA-DMFE.

80
higher during a short pulse than in steady state for a given surface temperature rise. Thus, even though rapid pulses or transients cause relatively high

losses, the ability of the helium coolant to dissipate pulsed heat rapidly is

a compensating factor.

At the beginning of a heating transient, as the power rises rapidly, the

temperature gradient in the fluid adjacent to the solid surface may be extremely steep since a finite time is required for an equilibrium temperature

field with less steep gradients to be established. During this period the

thermal resistance is so low that Kapitza solid-liquid interface resistance,

which is usually ignored at temperatures above the lambda point, may be the

dominant resistance. Since nucleation, growth of bubbles, and formation of

vapor films also require finite time, liquid exists at least briefly in nonequilibrium states above the saturation temperatures and boiling can be

delayed significantly. As the thermal boundary layer thickens through prolonged application of power or repetition of pulses, convection effects become

important and eventually steady state convection or boiling heat transfer are

established. In a chain of pulses the surface temperature rises while the

power is being applied and decays asymptotically toward the starting value

while the power is off; therefore, the entire level of temperature rises with

each pulse. The temperature ultimately reached after a large number of pulses

is related to the average heat flux for the cycle and the steady state heat

transfer coefficient.

It is apparent that the use of steady state heat transfer rates when

designing for heating transients in superconducting devices is overly conservative. Yet, for lack of more appropriate data, steady state boiling helium

81
liedL transfer data [lave been used in designs involving transient processes.

An exception to the use of steady state data was the pulsed high energy inductive storage system designed by AERl'^"''^ for the Air Force Aero Propulsion

Laboratory. Heat flux for this design was estimated from the transient ex-

[2] perimental data of Jackson. This data is not sufficiently detailed; howeve

for general application. More recently, experimental results have been report

[3] by two Japanese authors at Yokohama National University, for the pulse

heating of a thin lead film in a liquid helium bath. This study, however,

gives only one point at the superconducting transition temperature of the lead

film at a given power level, whereas, the entire temperature development is

desired. The temperature rise after the onset of nucleate boiling until the

establishment of steady temperatures has been experimentally determined by

[41

R. L. Bailey at the Rutherford High Energy Laboratory, but the time scale

was on the order of seconds and tenths of seconds and power densities were

2 less than 25 mW/cm . In the wide range of existing and contemplated superconducting applications, pulses as brief as 0.05 ms and repetition rates as high

as 200 per second are of concern. '^'^ In most applications, for economy of

weight, material and liquid helium, the power levels will be pushed to the

maximum safe limits as dictated by the ability of the coolant to dissipate

2 heat due to transients, and these maximum heat fluxes may exceed 10 W/cm

for short pulses.

; Therefore, it is clear that designers of superconducting equipment need

a more complete picture of transient helium heat transfer phenomena than was

available before the commencement of this study. This total program will

provide experimental data and verified predictive methods covering heat fluxes

2 up to 25 W/cm , measurement times starting at less than 0.05 ms extending to

82
steady state, varying orientation of the heated surfaces, varying helium

pressures and temperatures, and both static and forced flow of coolant. This

paper covers the first phase of the program, namely the experimental results

for a static coolant.

1.1. Acknowledgement

The author wishes to acknowledge the contributions of Dr. M. C. Jones

who recognized the need for helium transient heat transfer data and supplied

many of the ideas for measurement methods. Thanks are due also to Dr. Vincent

Arp for many helpful suggestions and preliminary studies.

2. IDEALIZED TRANSIENT CONDUCTION HEAT TRANSFER

At the beginning of a heating transient, before convection currents or

boiling have been established, conduction is the principle mechanism of heat

transfer. The following set of assumptions and boundary conditions for

solution of the heat conduction equation apply to the configuration of the

test apparatus.

1) The heat source is an infinite plate in the X-Y plane at Z = 0.

(This is equivalent to assuming that the thermal boundary layer is very thin

relative to the width of the plate.)

2) Heating occurs only on the surface at the liquid solid interface.

3) Stagnant liquid helium is present at Z > 0.

4) A solid substrate exists at Z < 0.

5) Time during which heat is applied is short enough that convection

currents are not established. This means there is not time for bubble Corm.u Ion and temperatures may exceed the equilibrium saturation temperaturi>

6) Radiation heat transfer is negligible.

83
7) Fluid and solid properties are constant.

8) Temperature difference due to Kapitza resistance may be considered

separately.

The Fourier heat conduction equation in one dimension of the solid substrate

[6] IS

9^AT , 8 AT

s _ 1 s ^% ^9

Z < 0; (1)

and for the fluid

8^AT^

^ 8AT^

Z > 0. (2)

Boundary conditions for the substrate for a square wave heat pulse are.

AT^ (Z,0) =0 (3)

Llm AT^ (Z,0) = 0 (4)

Z -> -oo

^s = \(^L. ' O<0<ep (5)

is the part of the heater power being dissipated in the substrate. Boundary

c-onditloiiK fur the fluid are.

84
AT^ (Z,0) = 0

Lim AT^ (Z,e) = 0

Z 00

-.(Si)., • « < e <

(6)

(7)

(8)

= 0 , e > ep .

is the part of the heater power being dissipated in the fluid. Also, the

total heater power is

P = P^ + P

f s where

P = const 0 < 6 <

(9)

P = 0 ^ Sp, and

AT^ (0,e) = AT^ (0,0) . (10)

The solution for the fluid is. [6]

-z

AT

f K

4a^0

-Z erfc for 0 < 6 < Bp (11)

and

AT

-Z

f K,

-Z erfc

-Z

V

af(O-Op) e 4a^(0-0p)

-Z erfc

\2/^0-0p)/

for 0 > 0 . (12)

85
The equations for the substrate temperature field are exactly analogous to

(11) and (12) except that a subscript "s" appears instead of "f".

At the surface (Z = 0) we have

and

(13)

(14)

Equation (J3) gives the surface temperature rise during the heat pulse and

equation (14) gives the temperature decay following the pulse.

Obviously the assumptions drastically limit the time during which

equations (13) and (14) are valid. Comparison with experimental data help

determine the conduction period duration.

3. KAPITZA RESISTANCE

Whereas fluid thermal resistance builds up as the boundary layer develops,

Kapitza resistance apparently occurs the instant heat flow begins. Consequently

Kapitza resistance initially dominates the heat flow process at liquid helium

Lemperal ures , Lliougli only briefly. How long Kapitza resistance dominates

depends upon its magnitude in relation to the increasing conduction resistance,

and this in turn depends on the kind of material and physical condition of

the surface as well as the temperature.

Since tlie exact nature of a surface is difficult to characterize^''^ and

may change with time, the influence of material and surface condition is

86
dLlficiUL Co predict. Flowever, for a given heater tlie Kapitza conductance li

3 (reciprocal of resistance) has been observed to obey approximately a T

relationship

Vref id'^1 •

where h^^ is a constant Kapitza conductance at reference temperature T^^^

is the power being dissipated into the helium bath.

The heater surface temperature is \

bath f K

Therefore,

P

= Vref

ref

3 ( bath + AT^ + AT^ I AT^^ . (16)

For low power levels AT can be estimated from the experimental data as K

the asymptote of the temperature curve as 6 0. In this way it was determined that

2 h = 0.7 W/cm K at approximately 4 K.

Also,

= 0.011 (17)

ref cm K

If ~ 1.9 K, = 0.075 which is well within the possible range

of values of h^^ at 1.9 K shown by Snyder

87
4. EXPERIMENTAL PROGRAM

The apparatus used is shown schematically in figure 1. A carbon thin

[81 film resistor whose resistance is a known function of temperature serves

both as an extremely fast response heater and thermometer. Its dimensions

are approximately 5 mm parallel to current by 10 mm wide. The carbon film

is deposited on a quartz substrate which has an extremely low heat capacity

at liquid helium temperature so that most of the heat flows into the fluid

and very little flows into the substrate. Voltage across the film and

voltage across a standard resistor are recorded at intervals as close as

0.5 microsecond by a digital recording oscilloscope; 2048 points are recorded

on each sweep of the oscilloscope on each of two channels. Power may be

supplied as a step function or as pulses of any desired width or frequency.

A minicomputer data acquisition system has been added to facilitate data

processing and to permanently record the data. The heater surface temperature, power level and heat transfer coefficients axe printed as a function

of time.

' For these tests the heater was submerged in a static helium bath. Insofar as possible power was applied as a step input or square wave pulse and

ramp times of less than 1.5 (10) ^ s were achieved. However, at high power

levels resistance change due to large temperature rise causes power variation

of up to + 10 percent from the average. Tests covered the following parameter variations:

2

1. Power densities ranging from 0.05 to 20 W/cm repeated for each

variation of the other parameters.

2. Four orientations of the heater — vertical, horizontal facing up,

horizontal facing down, and 45° down.

88
Position

Controlle

DIGITAL

STORAGE

OSCILLOSCOPE

He Fill

AMPLIFIER

Bock

Pressure

Regulator

FUNCTION

GENERATOR

N2 Fill

& Vent

Tempering

Ring - Variable

Position

Ht Trans.

Surface

Germanium

Thermometers

Figure 1. Pool pulsed heat transfer experimental apparatus.

89
3. Variations in bath pressure and temperature.

4. Repetition of square wave pulses with variations of pulse width

and time between pulses. Finally, temperature decays following single pulses

were recorded.

4.1. Preliminary Liquid Nitrogen Transient Heat Transfer Experiments

We performed tests at four power levels with liquid nitrogen for a

check on measurement techniques and predictions for a fluid which is markedly

different from helium in heat transfer related properties, and for which

Kapitza resistance is negligible. Figure 2 shows the temperature rise of a

vertical heater surface in liquid nitrogen as a function of time from the

start of a step power input. The heat flux per unit area was nearly constant

2 during the measurements and the four tests were at 5, 10, 15 and 18 W/cm . -2

The temperature rise in the first 3(10) s closely follows the non-boiling

transient heat transfer calculations.

2 For 5, 10 and 15 W/cm the heater temperature continued to rise in proportion to (time)^''^ and for 9 to 15 K above the steady state nucleate boiling

level, then dropped back down to that level. The time required to finally

reach steady state nucleate boiling was approximately 0.2 s, A power density

2 of 18 W/cm exceeded the maximum nucleate boiling heat flux and the heater

began a slow transition to a film boiling temperature which might have burned

out the heater if the heating had been allowed to continue.

The heater temperature rise during the non-boiling transient heat transfer period was calculted from equation (13) in which ~ + = 0.0906 for LN^

/li^ /{Xf cm K

90
91
and quartz substrate. At liquid nitrogen temperature Kapltza resistance is

negligible; therefore AT = AT^.

4.2. Liquid Helium Transient Heat Transfer Experiments with a Vertical

" :i Heater

) A. 2.1. Accuracy

'

: The carbon thermometer resistance at helium temperature increased

by 1.3 percent over the duration of the test program; however, this slow

drift did not affect the test results since the thermometer was calibrated

at room temperature and by nitrogen and helium vapor pressure measurements

each time the apparatus was cooled from room temperature and the helium

temperature calibration was re-established before each test. Systematic

error in calibration is no more than + 0.01 K, and experience with pulses

repeated seconds apart and months apart indicate that heater-to-bath temperature difference measurements were repeatable within + 5 percent. While

instantaneous heater power is accurate to + 0.1 percent, the variation during

a pulse from the stated average power due to heater resistance change ranges

2 2 from + 1 percent at 0.1 W/cm to a maximum of + 8 percent at 20 W/cm .

Time measurements are initiated by the triggering of the digital oscilloscope at the beginning of a pulse; however, the time to achieve essentially

steady power is approximately five microseconds later than the triggering

time. No corrections wore made in the plotted data for the finite power

rise time

The liquid helium results are complicated by a significant Kapitza

resistance which is added in series with the thermal boundary layer resistance,

Because of the Kapitza resistance the heater temperature does not rise simply

92
in proportion to during the non-boiling transient hfat transfer period

For helium " " * AT = AT^ + AT^,

where AT, is given by eq. (13). For helium the quantity

K K 1/2

-L. + -I- = 0.0118 ^^^-^y- . . (18) a /a^ • cm K

s f

The surface temperature step AT due to Kapitza resistance is given by

equation (15).

The non-boiling trnasient heat transfer calculations using eqs. (13),

(•2), (3) and the constant (4) are shown in figure 3. The earliest time at

which temperature can be measured is about 20(10) ^ s. It can be seen that

except at low power levels the non-boiling transient heat transfer period

is over by 20(10) ^ s; however, the calculated temperatures at 6 < 20(10) ^ s

are consistent with the measured temperatures. The non-boiling transient

period leads either to steady state nucleate boiling as in the 0.1 to 0.6

2 W/cm curves or to a metastable nucleation period at higher power levels.

The length of time the metastable nucleation period lasts decreases as the

93
power level increases. (The steady state nucleate boiling at power levels

2 below 0.6 W/cm might be thought of as a "metastable" nucleation period which

2 lasts indefinitely). At power levels of 1 W/cm and above, the metastable

nucleation period leads to a period of relatively slow transition to stable

2 film boiling. At the highest power levels of 9, 15 and 25 W/cm the transition to film boiling has already begun at the earliest possible measurement

times. The transition times become longer at higher power so that steady

state film boiling is reached at about 0.1 s for all power levels.

The steady state nucleate and film boiling power level versus tempera-

[9] ture rise relationships are shown along with data from previous research

in figure 4. The new data, including the value of the peak steady state

nucleate boiling heat flux, compares well with the previous data and Kutateladze correlation but with nucleate boiling AT slightly on the high side. It

is interesting to note that the metastable nucleation data falls on an

extension of the nucleate boiling curve above the steady state peak heat

flux. The curve of Kapitza AT versus P is also shown on this curve. Obis. viously the nucleate boiling AT must be greater than AT , so the very large

K. Kapitza resistance of the carbon film may account for the slightly higher

than average nucleate boiling AT of the new data.

Heat transfer coefficients derived from these data are discussed in

Section 4.6.

4.3. The Effect of Heater Orientation on Transient Helium Heat Transfer.

The orientation of the heating surface is known to be an important factor

in steady state natural convection and boiling heat transfer because of the

effect of buoyancy and orientation in inducing currents and in expelling

94
o

+->

(T3

jQ3

-o

(U

+-> ra

i- -M

n3

I/)

<aO-•->

s-O)e

(O

s- -M . J4-> OJ <0 -M

•IO

(T3 S- S- 0) I— >

CO

(U3

01

95
vapor bubbles. Particularly In boiling, bubbles formed on a vertical or

upward facing surface tend to be expelled and swept away; whereas, bubbles

forming on a downward facing surface tend to be trapped and form a thick

insulating film which results in poorer cooling. Data such as shown in

figure 3 for the vertical heater have been obtained for the horizontal facing

up, facing down, and 45 degree facing down positions. Comparisons are shown

2 in figure 5 for power levels of 0.1 W/cm in the steady state nucleate boiling

2 regime, and 1.0 and 5 W/cm in the film boiling regime. Little effect of

orientation is seen in the transient regime; however, the downward facing

heater finally rose markedly above the other orientations particularly in steady

state film boiling. The effect is less pronounced in the nucleate boiling

range but still exists there. In general, vertical and horizontal facing up

positions are nearly the same; 45 degree tilted downward gives the medium

temperature, and a heater facing straight down produces the highest temperatures.

4.4. Pressure Effects

In all tests of figure 6 the bath bulk temperature was maintained at

approximately 4.02 K; the pressure is raised to 0.151 MPa (A symbol) and

0.310 MPa ( symbol) as opposed to atmospheric pressure, 0.086 MPa (o symbol).

2 For a heat flux of 0.1 W/cm the temperatures for the three pressures follow

nearly the same transient conduction curve; however, the higher pressures

lead to decidedly higher steady state temperatures as would be expected; the

increase in surface temperature in going from 0.086 MPa to 0.151 MPa is approximately the same as the increase in saturat ion temperature between those pressures

It should be noted that the highest pressure, 0.310 MPa is above the critical

pressure, yet the steady state temperature is consistent with the boiling curves

96
11 I l U J J.^J J

^uj3/M <i V3av/a3ivdissia >i3M0d

C3>

o

<o +^

m

+J

</)

•rfO +J

fTS u

T3

OJ

S-0) 4JC <D

0) cc: _E c X •p— a;X QJ

OJ (/) LO -M 13 C u

Oi 1/1 00

Oi -o Q. to ^ +-> o

J- c <o

(U

OJ to

d) +J s_

<D jD (/) c o

o

(/I >

•r— OJ

s-

(13 Q.

Q.

E -O O c o (T3

J97
o« •a

o

u

Q

U

•

• a

Q) u

> ••

• o

• ° a

o

o

•

• o

°

» •Q • o• 8°

'

l«3

C_J C_J ^ u- u- H

8

B

if o^

^oVo

" K>

B

QQ

to

ft

toc

o

(t3 ->C

<u

*r—o

o

Mo

to

0)

i-

:3

+-> 03

OJCLE

o

co

(/I

Q.

Oo

O)

:3

CDO

M iV3Sm 3BniVH3dlAI3i d3iV3H

98
<D +J

•O

1—o

•r—

-»-> S-(U>

S-o

14->,c

0+J

(/) •r- x:

O)

j_

ra E

S- 13

a> r—

Q.

E O)

a> XI

-M

OJ x: 0

t-> •1

4-> c0 1-U

<u

i_rj Q.

to

(/) to

OiQ. x:

+J -0

fO cr

<+-0

-MU

O) H-

>+- LU

OJ3

>l iV'aSm 3HniV«3dlM3i «3iV3H

9Q
X

CNI

CNl ZD CO o • <3

CO

CNl O

II o ^ c

Q_ CD

o

CO

oo

II

CD

— CD

— CO

Lxi OO

<

CO o

<

o <] •

O •

O <•

O ^»

ooooc

_ ro

O 4

CD tr5 CO oCO CM oCNl in

>l iV '3Sia 3bniVy3cJIAI3i a3iV3H

CD

100
lower pressures. The same phenomena can be seen in the metastable nucleation

2 region to the left of the transition to film boiling at 1 and 5 V.'/ cm . However, at 0.310 MPa the transient conduction curve leads directly to the transi2 tion to film boiling with no period of apparent nucleation. At 5 W/cm only

-4

a small flat region can be observed at times less than 10 s for 0.086 MPa,

and the two higher pressures are already in transition to film boiling at the

earliest measured time.

It is interesting that the trend of steady state film boiling temperatures

at 5 W/cm^ is reversed from that of nucleate boiling. This is perhaps due to

improved heat transfer through the vapor or supercritical fluid film at higher

pressure and that the film would be thinner at higher pressure. This does

2 not explain the behavior at 1 W/cm in which the highest pressure produces the

highest temperature as in nucleate boiling. However, the fact that the steady

2 state heat transfer at 1 UV cm exhibits some characteristics of nucleate boiling

is not particularly disturbing since then heat flux is only slightly above the

nucleate boiling peak and some regions of the surface are probably undergoing

nucleate boiling.

4.5. Repetition of Pulses and Temperature Decay

v-Then heating pulses are repeated the temperature rises while the power is

being applied and decays as>Tnptotically toward the bath temperature while the

power is off. If succeeding pulses begin before the surface has fully regained

equilibrium with the bulk fluid the entire level of temperature rises with each

pulse until a stable condition is reached in which the total energy of each

pulse is dissipated during each cycle. Figures 7 is an example of one of these

9 temperature histories during 1 ms heating pulses of 5.2 W/cm and an off-time

-3

of 8.16(10) s. These tests show, for this example, that sustained pulses would

produce a peak temperature of approximately 30 K above the bath temperature.

10]
TIME, e.s

Figure 8. Temperature decay following 10 W/cm pulses of 10" , 10" , and 10" s duration.

102
Figure 8 shows a typical temperature decay following square wave heat

pulses. The nature of these decay curves depends on not only the heat flux

during the pulse but on the temperature reached at the end of the pulse and

the duration of pulse. All of these factors affect the boundary layer temperature profile as the decay begins. These figures confirm the previous

interpretation of the initial temperature step as being due to Kapitza resistance. Following the pulse, when the power is essentially removed (except for

a small measurement current) and Kapitza resistance should disappear, the

temperature does in fact return to the bath temperature.

4.6. Heat Transfer Coefficients

Heat transfer coefficients were calculated for all of the experimental

data. Some of these are shown in figure 9 for a vertical heater and atmospheric pressure, along with curves calculated from,

c K

where AT and AT,, are from equations (13) and (17).

C K

2 At the lowest power levels of .051 and 0.1 W/cm the experimental data is

seen to follow the transient calculations until steady state nucleate boiling

is reached. The steady state condition is preceeded in both those cases by

a dip in heat transfer coefficient corresponding to the slight temperature

overshoot seen in figure 3. It is seen that the time to reach steady state

varies all the way from 10 ^ s to 1 s

2 At 0.6 W/cm a vestige of the transient curve is seen just before nucleate

boiling begins at about 3(10) ^ s; however, at higher power levels transient

conduction curves apparently intersect the boiling or metastable boiling curves

103
before the earliest measurement time. The assumed metastable nucleation curve

2 2 2 for 24 W/cm , and extrapolation of those curves for 2 W/cm and 8 W/cm to the

transient conduction Intersections are shovm in dashed lines in figure 9. It

should be emphasized that the extrapolated curves are rough estimates..

The Kapitza resistance of the carbon heat transfer surface may be representative of materials used for insulation but is considerably higher than

typical base metals. The heat transfer coefficients for heaters with negligible

Kapitza resistance has been estimated by subtracting AT from equation (17) K

from the observed AT. The transient conduction heat transfer coefficient then

becomes independent of P. These curves in figure 10 indicate that heat transfer

coefficients at very early times may be three times as large for materials with

negligible Kapitza resistance as for the carbon film.

Negligible Kapitza-resistance heat transfer coefficients such as shown

[2] in figure 10 are compared with the experimental data of Jackson in figure 11.

The 0.031 cm manganin wire used as heater and thermometer by Jackson may

account for the rather large differences between the two sets of data since

surface curvature has a pronounced effect on bubble nucleation when the radius

of curvature is on the order of the bubble radii. Even though there are points

2 of agreement on the curves of 1.5 to 2.2 W/cm the slope dh/d6 is considerably

steeper in the present data and Jackson's data indicates larger dependence of

h on heat flux.

-

, 5. CONCLUSIONS

Transient heat transfer experiments using fast response carbon thin film

heater thermometers in a static helium bath show the effects of heat flux,

heater orientation and pressure on surface temperature and heat transfer

104

rT—

1

1 I I I I I—

J /l I 1 1 I I 1 1 I I L 1 1 I I I I

I I I

o

fO

(

SCO

sO) Murc

<a

i~ -M

-l-> »a

s-o

M-

• to cu

+-> o

c

<u fO

•r— 4-> u </)

• 1

•r—

«4- 00 4- CU

O) s- O

O <T3 N

S- (-) cu •r- ^- Ql

1/1 03

03

s- <U

4-> JD

+-> •f— "3 cn

O) •r— x: , cn

+j <U c c

O)

•1

cn

I/) c

>

i- 1—or—

O)

i.cn

•r- U.

oin o cp

)\ • lua/w q'iNHIOIdJBOO a3dSNVUi iV3H

106
T

o00< o

+-> to -aM

c

(U

•r—U

CO

<+-Q)o

o

I-(U

«4-inc

fO

S-

+-> -M

CU fO

•r- 4-> c -o

-t->

-t-> c

CD

fO CU

O) s- ^ Q.

<u a»

c -o

OJ c

-M I 1 O) oo

JD I 1 c c

o o

(/) CO

•1— ^

S- o

n3 fO CL'-D EO 4- O O

(U

>!• iu3/'v\'q'll\l3l3ldd303 a3dSNVyi 1V3H

107
coefficients. For heat flux below the steady state peak nucleate boiling

limit the surface temperatures follow calculations based on pure conduction

(allowing for Kapitza resistance) until, after a small overshoot, the temperature reaches the steady state nucleate boiling level. Above the steady state

nucleate boiling peak the transient conduction period leads to an apparent

metastable nucleate boiling period followed by a transition to film boiling.

All of these events vary systematically with heat flux. Steady state temperatures

fall within the range of previous experimental data which, granted, allows considerable latitude.

A downward facing orientation of the heater increases the steady state

temperature rise 20 to 60% over the other orientations, but has little effect

during transients. Likewise, raising the coolant helium pressure has little

effect during the transient conduction period but raises the steady state and

metastable nucleate boiling temperature levels. The effect of pressure is

reversed in steady state film boiling, apparently because measured pressure

improves heat transfer through the vapor (or low density supercritical fluid)

film.

Kapitza resistance for materials such as the carbon heat transfer surface

used in these experiments may override much of the benefit which might have

been gained from the initial high helium conduction heat transfer. For that

reason the type of electrical insulating material or metal covering the superconductor and contacting the helium coolant wil] have a signif Leant effect on

transient cooling. Investigation of surface material effects as well as forced

convection effects are planned for a continuation of this program.

-5 -2 Steady state nucleate boiling is reached in 10 to 10 s, but steady state

film boiling conditions are not reached until 10 to 1 s from tfie beginning of

108
a heat pulse. Since the initial heat flux is at least an order of magnitude

higher than steady state, the time variations in heat transfer coefficients

should be considered in the design and analysis of superconducting equipment

if these transition times comprise a significant part of anticipated heating

pulses.

6. NOMENCLATURE

c heat capacity, j/g*K

2 h heat transfer coefficient, W/cm 'K

2 h Kapitza conductance, W/cm 'K

K

h,, - Kapitza conductance at T _ K,ref ref

thermal conductivity for the fluid, W/cm*K

K thermal conductivity for the substrate, W/cm'K

2 P power going into the fluid, W/cm

2

power going into the subst;rate, W/cm

T temperature, K

T, , bath temperature, K

bath

Z coordinate normal to the heating surface, cm

2 thermal diffusivity of the fluid, cm /s

2 a thermal diffusivity of the substrate, cm Is

s

y >

AT^ temperature of the fluid above the bulk temperature, K

AT^ temperature of the substrate above the bulk fluid temperature, K

6 time from the start of a heating pulse, s

6p duration of a pulse, s

109
7. REFERENCES

1. J. Teno, et al. , Development of a pulsed high-energy inductive energy

storage system. Final Report, AVCO Everett Research Laboratory, Inc.

prepared for Air Force Aero Propulsion Laboratory, Wright-Patterson

Air Force Base (1973).

2. J. Jackson, Transient heat transfer and stability of superconducting

composites, Cryogenics 9^, 103 (1969).

3. 0. Tsukamoto and S. Kobayashi , Transient heat transfer characteristics

• of liquid helium, J. Appl. Phys. 46, No. 3 (1975).

4. R. L. Bailey, Heat transfer to liquid helium in pulsed heated channels.

Experiment Report RL-73-089, Department of Engineering Science, Rutherfor High Energy Laboratory, Chilton Didcot Berkshire, England (1973).

5. V. Arp, P. J. Giarratano, R. C. Hess, and M. C. Jones, Heat transfer

in pulsed superconducting magnets, Nat. Bur. Stand. (U.S.) NBSIR 74-363

(1974).

6. H. S. Carslaw and J. C. Jaeger, Conduction of heat in solids, Oxford

University Press (1950).

7. N. S. Snyder. Thermal condu _Laace at the mtertace of a '.'.t.d and helium

II (Kapitza conductanci ) , Nat. -^^ur. Stand. (U.S .) led. Net' 385 (1969).

8. R. S. Ccillitjr, L. L. Sparks, and I. t<. 6trob^id^e , Larbcn thij- film

thermometry Nat. Bur. Stand. (U.S.' N3t.lR /4- 5.5 (1V7 ).

9. E. G. Brentari, P. J. Ciairatano, an.l R. V. Sii.itli, I'/oi.ini', he/it Is ia'tfer

for oxygen, nitrogen, hydrogen and l)2lium, M t. lur. S.'atid. (II.:-;.) ['(M-ii.

Note 317 (1965).

110
NBSIR 78-877, Part Three

MAXIMUM PRACTICAL EFFICIENCY OF HELIUM TEMPERATURE REFRIGERATORS

R. 0. Voth

Cryogenics Division

National Bureau of Standards, Institute for Basic Standards

Boulder, Colorado 80302

ABSTRACT

An ideal refrigerator using a perfect gas v/orking fluid is defined which gives the efficiency of a refrigerator as a function of

compressor and expander efficiency, heat exchanger temperature difference, and heat exchanger pressure drop. Although not suited to

detailed hardware design, this approach clearly relates the overall

cycle efficiency to component efficiencies. In contrast, computer

studies of specific cycles using real fluid properties are usually

such that the details tend to overshadow major trends. The results

of the study show that in an efficient cycle the major losses are

in the compressor and the cold end expansion device. For current

compressor and expander efficiencies the maximum practical helium

temperature refrigerator efficiency is about 37 percent of Carnot.

Key words: Component efficiency; cryogenics; efficiency; helium; maximum

refrigerator efficiency; perfect gas analysis; refrigerator.

Ill
MAXIMUM PRACTICAL EFFICIENCY OF HELIUM TEMPERATURE REFRIGERATORS

R. 0. Voth

Cryogenics Division

National Bureau of Standards, Institute for Basic Standards

Boulder, Colorado 80302

I . Introduction

We know the efficiency of large helium temperature refrigerators is limited

by the efficiency of the available compressors, expanders, and heat exchangers.

But which component arrangement or cycle will yield maximum refrigerator efficiency

and which component in the refrigerator contributes most to the refrigerator

inefficiency? Answers to these questions can be found by using a computer and

real gas properties to investigate various refrigeration cycles that use components with variable efficiencies. However, the results of such a computer

study are usually clouded by the fact that a) One is never quite sure that a

totally different and unconsidered cycle may not yield higher efficiencies and

b) The computer studies are usually such that the details tend to overshadow

the more obvious major trends. We describe here a simple defined refrigeration cycle using an ideal gas working fluid which determines both the major

sources of inefficiency in a refrigerator and the maximum refrigerator efficiency

based on the efficiencies of available components. Although this approach

overlooks many of the details required to design a real refrigerator such as

heat exchanger temperature pinches, non-ideal fluid refrigeration effects and

economic and reliability considerations, it offers the advantage of clarity and

simplicity. The overall refrigerator efficiency is directly related to the

individual component efficiencies and the maximum efficiency obtainable for a

given set of component efficiencies is given.

Contribution of the National Bureau of Standards, not subject to copyright.

112
II. lfficiemcy of defiwed refrigerator with perfect meat exchanger

A defined refrigerator cycle -- the schematic is shown on figure 1 -- has

an isothermal compressor, a theoretical isothermal expander as the final expansion

device, heat exchangers, and a variable number of expanders located at intervals

in the heat exchanger to absorb heat exchanger losses. The isothermal expander

is a theoretical device to obtain cooling from a perfect gas and it is analogous

to the final expansion device of a refrigerator where the constant temperature

cooling is a result of evaporating liquid at constant pressure. The expanders

used to intercept heat exchanger losses are situated so the inlet temperature of

each expander is higher than the discharge temperature of the preceeding expander

by the temperature difference of the heat exchanger. The heat exchanger temperature difference is proportional to the absolute temperature as defined by a

constant, C = AT/T. This arrangement minimizes heat exchanger losses [1].

The isothermal expander is supplied by a separate compressor so its operating

pressure ratio can be maintained constant while the number of loss absorbing

expanders in the heat exchanger can be varied along with their operating pressure ratio.

Refrigerator efficiency is defined as the ideal compressor power divided

by the actual compressor power minus the recoverable power from the expanders,

or

n =

p . p X 100, (1)

a e where

n

= refrigerator efficiency,

Pj = ideal compressor power required to produce a set amount of

refrigeration,

P = actual compressor power required to produce the same amount of

refrigeration,

and Pg = recoverable power from the expanders.

113
Figure 1. Defined refrigerator cycle.

114
The ideal power required by the compressor in an ideal refrigerator is.

Pi = Q(T„-T^)/T,, (2)

where

Q = refrigeration load,

= rejection temperature (300 K),

and = refrigeration temperature (4.2 K).

Actual compressor power required to supply the isothermal expander is

rfi,RT^. In Pr^

P = -i-iL ^ (3)

^ ^c . where

rtij = isothermal expander mass flow rate,

R = gas constant,

Pr^ = compressor pressure ratio (Pf^/P|_)>

and = compressor isothermal efficiency.

When the heat exchanger is perfect, no loss absorbing expanders are required so

the total mass flow rate (rfi) equals the mass flow through the isothermal expander

or

•^I " mRT, ?n Pr/

I c I

where

rij = efficiency of the isothermal expander,

and Prj = pressure ratio of the isothermal expander.

Recoverable power from the isothermal expander is equal to the refrigeration

load or Pj = Q. Since no expanders are required to absorb the heat exchanger

losses, the cycle efficiency can be found by combining equations (1), (2), (3)

and (4). This combination results in the equation

115
I

- 1

X 100. (5)

This equation reduces to 100 percent when Pr^ = Pr^ (resulting from no pressure

drop in the system), and when the compressor and isothermal expander are 100 percent efficient, or n^, = nj = 1- The effect of pressure drop on refrigerator

efficiency is found for this simple system by increasing the compressor

pressure ratio above the isothermal expander pressure ratio or Pr^ = APrp

where A = (1.0 + AP/P)/(1.0 -AP/P), and AP/P is the ratio of the pressure difference between the warm and the cold end to the absolute pressure in that stream.

The efficiency of a simple refrigerator including some pressure drop effects but

perfect heat exchange is shown in figure 2. From the general trend of the

efficiency versus pressure ratio, the maximum refrigerator efficiency occurs at

an infinite pressure ratio. Because an infinite pressure ratio is unrealistic

and because a pressure ratio of 50:1 is well above the knee in the curve, we

chose to maintain the isothermal expander pressure ratio constant at 50 for the

remainder of the investigation. From equation (5), it is apparent that the compressor and isothermal expander efficiency have an equal effect on refrigerator

efficiency.

III. EFFECT OF IMPERFECT HEAT EXCHANGER ON DEFINED REFRIGERATOR EFFICIENCY

The effect of heat exchanger temperature difference on refrigerator efficiency is found in a similar fashion. The additional compressor power required

when expanders are needed to absorb the heat exchanger losses (loss absorbing

expanders) can be found by using equation (3) but the compressor mass flow

rate is now a function of the number of expanders and their pressure ratio.

Pressure ratio is determined by the number of loss absorbing expanders because

they are placed end to end.

116
eo

CO

CO

CO

LU

CM o

+^

fO

QJ f—

CD 1 ^

ex. f—

OJ

QJ (J c:

(-^ o Oii o

o.

CJ

(j

fD

QJ

00 -o

ra>>u s- c (O

CDu +-> ^- QJc

Ol o

Q.

S- E o O -(-> <_) 031. S-

(1) o

•r— ro Q.

s- SO

M- QJ S- <D cn "O

Ss_ QJ -a SQJ zs

S- 00

00

QJ

i.

Q 5

CM

<U

117
The temperature drop across an expander operating with a perfect gas is

Tin -Tout = Vln(l " V'-

where

T^.^ = expander inlet temperature,

^out

~ expander outlet temperature,

rip = isentropic efficiency of the loss absorbing expander,

Pr^ = pressure ratio across the expander (P^/'^l^'

a = (k - l)/k (0.4 for helium),

and k = ratio of specific heats (1.667 for helium at ambient temperature and zero pressure).

From figure 1, expression for T^, T^,

^n+1 ' written based on

equation (6) as,

T2 = T^[(l - rip + npPrL"°')/(l - C)], (7)

T3 = T^Ld - rip + ripPrL"'')/(l - C)], (8)

and T^^^ = T^[(l - rip + ^''l''^^^^

'

where C = constant used to define heat exchanger temperature difference,

C = AT/T,

Assuming the same efficiency for all the precooling expanders and substituting to

maintain in the equation yields,

Vi = h

- ThCO - + npPO/d - ClA'^P (10)

where N^^p = number of loss absorbing expanders.

118

I
Solving for Pr-j yields,

^ lAx

Pr = I —

(1 - C)(VT,) f np - 1 'P

Note that the value of the fraction within the brackets must be positive so that

as the efficiencies of the heat exchanger and expander are decreased more loss

absorbing expanders are required.

The mass flow through the compressor supplying the loss absorbing expanders

is the sum of the mass flows through the individual expanders. The total expander

flow rate is determined by an energy balance around each heat exchanger expander

combination as follows:

% ='^i r^' ^^^^ exp

exp exp

'^N ^% =

"'I'^n 1 C)^ ^^^^ exp exp-1 ^

'

and ifij,. + rfi|^ + tfij^ = m^N ^ \

- 1]. (14)

exp exp-1 exp-2 ^

'

Thus, the total compressor flow for a variable number of loss absorbing expanders

located end to end becomes,

rfi = m, + rfip rfi + tfi^ + rfi,. = rfi, ^ 1 (15)

^ 2 ^xp-2 ^exp-1 ^xp L Ml - C)^exp J

where

ifi = total compressor mass flow rate

ifij = isothermal expander mass flow rate from equation (4),

and subscripts

1 = first or warmest loss absorbing expander ,

'^exp

~ ^^'^^^ coldest loss absorbing expander.

119
The most efficient refrigerator will use the expander output power to

reduce the power required by the compressors. The maximum recoverable power

from the isothermal expander is equal to the refrigerator capacity, or

Pi = Q. - (16)

The maximum recoverable power from the loss absorbing expanders involves

a summation. The power from a single expander operating with a perfect gas is,

P = rtiCpAT, (17)

or for expander number 1, figure 1, the recoverable power is,

where

P-j = power output of the warmest loss absorbing expander,

rfi-| = mass flow rate through the warmest expander,

and Cp = specific heat at constant pressure.

By substituting the temperatures from equations (7,8,9) and the expander mass flow

rates from differences between equations (12, 13, 14, 15) into equation (18) while

substituting to maintain T^, the expander output power becomes:

(1 -C) exp

C

(1 -O^exp

C

(1 -C)'^exp

C

(1 -O^xp

3

- ,N l^pThfnp-npPOd -np-npPO^- . (21)

where

P = expander power output,

rip = expander isentropic efficiency,

Prj^ = pressure ratio across the expanders.

120
^exp

~ total number of loss absorbing expanders,

and subscripts

1 = warmest loss absorbing expander —

,

^exp

^ coldest loss absorbing exapnder.

Equations (19, 20, 21, and 22) are summed to establish the total recoverable

power from the loss absorbing expanders or

'^'^ ^d^^/' '"*^p

" 'p"" tr *^ ^p'

''p""

Total compressor power required by the refrigerator is the sum of the compressor power required for the isothermal expander loop (Pr^ = 50) and the compressor power required by the loss absorbing expanders or

RTu In Pr,

a I

~ M 1 .Ml - Cfexp

RTu In Pr, —2 ^ (24)

The total recoverable expander power is the sum of the recoverable power from

the loss absorbing expanders and the isothermal expander, or

(1-C) exp n=l ^

*^

The overall refrigerator efficiency can now be calculated by

c ^ P^(equation 24) - P^(equation 25) (26)

This equation includes the effect of component inefficiency and heat exchanger

temperature difference. The effect of pressure drop can also be included by

changing the compressor pressure ratios as they were changed previously. Increasing the pressure ratio across the compressor that supplies the isothermal

expander by the factor A = (1.0 + AP/P)/(1.0 - AP/P) above the pressure ratio of

the isothermal expander and increasing the pressure ratio of the compressor that

121
supplies the work absorbing expanders above that calculated by equation (11)

results in the maximum effect of pressure drop on cycle efficiency. This maximum effect was used in the data presented herein. A minimum effect of pressure

drop on refrigerator efficiency would be to reduce only the pressure ratio of

the isothermal expander below that of the compressor supplying it. The minimum

and maximum effect were very nearly the same differing in the third significant

digit.

IV. RESULTS

The ideal refrigerator efficiency with a perfect heat exchanger -- no

pressure drop and no temperature difference is shown in figure 3. The

pressure ratio of the isothermal loop was maintained at 50 during the calculation and no loss absorbing expanders were required. The maximum refrigerator

efficiency was very nearly equal to the product of the compressor and isothermal

expander efficiency.

The effect of pressure drop and heat exchanger temperature difference on

refrigerator efficiency is shown on figures 4 and 5. Again the refrigerator

efficiency is plotted versus expander efficiency but in this instance the refrigerator has loss absorbing expanders and their adiabatic efficiency is

numerically equal to the isothermal expander efficiency. The cycle efficiency

varied slightly with the number of loss absorbing expanders and if the maximum

efficiency occurred with six or less expanders that point was plotted, otherwise the cycle efficiency for the minimum number of expanders required to produce ^a positive fraction in equation (11) was plotted.

The refrigerator efficiency depends primarily on the compressor and cold

end expander efficiency in contrast to a liquefier where the efficiency of

the precooling expanders has a significant effect [2]. The dependence of the

122
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125
refrigerator efficiency on the efficiency of the loss absorbing expanders is shown

on figure 6. In this figure, the compressor efficiency and the isothermal expander efficiency are constant while the efficiency of the loss absorbing

expanders varies. Comparing figure 6 with figures 4 and 5 shows that the

efficiency of the isothermal expander affects the refrigerator efficiency more

than the efficiency of the loss absorbing expanders.

V. CONCLUSIONS

The efficiencies derived using the relatively simple defined cycle can

serve as a guide for the efficiency study of cycles using real components with

real gases. If the efficiency of the real gas cycle is near the maximum found

by the defined cycle, the real cycle is using the available components in their

optimum operating locations and pressure ratios. On the other hand a significantly lower real gas cycle efficiency indicates that a new cycle may achieve

higher efficiencies. Of course, the design of any practical unit is a compromise, so the real refrigerator efficiency may be lower than the defined cycle

efficiency due to reliability or capital cost considerations.

The defined cycle also shows that the overall refrigerator efficiency

is most directly effected by the efficiency of the compressor and the isothermal

final expander. Thus, it is apparent that the efficiency of the cold end of a

real refrigerator has a significant effect on overall refrigerator efficiency.

Replacing the final Joule Thomson expansion valve with a more thermodynamical ly

efficient expansion ejector or a wet expander (liquid in its exhaust), should

significantly increase the efficiency of a refrigerator. If it is preferrable

to keep the Joule Thomson expansion valve (JT valve), its inlet temperature

should be reduced as closely as possible to the refrigeration temperature by

the use of a cold expander. Using a cold expander-JT valve combination will

also reduce the optimum pressure ratio of the JT valve.

126
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127
The compressor efficiency has a major influence on the refrigerator

efficiency, but because the compressor has a long development history,

increasing its efficiency above the current 60 percent of isothermal is probably

difficult and expensive.

The results from the defined refrigeration cycle indicate that with a

60 percent isothermal compressor, 80 percent adiabatic precooling expanders,

and an 80 percent isothermal expander the maximum refrigerator efficiency

attainable is 48 percent (figure 2). Losses due to system pressure drops and

heat exchanger temperature differences reduces the expected maximum efficiencies

to 42 percent (figure 3) and 37 percent (figure 4). A reasonable maximum

efficiency for a practical refrigerator may lie between 37 to 42 percent

without considering other losses such as heat leak to the cold components or

possible irreversibilities resulting from the properties of a real gas. Even

with the irreversibilities of a real gas a refrigerator efficiency of 37 percent

should be possible if the refrigerator is designed to achieve the maximum

efficiency without compromises to reduce the capital cost of the refrigerator

or to increase the reliability of the refrigerator.

128
VI. NOMEilCLATURE

A proportionality constant used to define pressure drop

C proportionality constant used to define heat exchanger temperature

difference, C = AT/T

Cp constant pressure specific heat

k ratio of specific heats, Cp/Cv

rfi mass flow rate

fflj isothermal expander mass flow rate

N number of expanders

exp

P recoverable expander power

P actual compressor power

a Pg recoverable power from the expanders

Pgj recoverable power from the isothermal expander

P^ high pressure

P|^ low pressure

P|^^ total recoverable power from the loss absorbing expanders

Pr^ compressor pressure ratio

Prj pressure ratio of the isothermal expander

Pr|^ pressure ratio of the loss absorbing expanders

Q refrigeration load

R gas constant

refrigeration temperature (4.2 K)

rejection temperature (300 K)

T^^ inlet temperature

"'"out "temperature

a (k - 1 )/k (0.4 for hel ium)

n refrigerator efficiency

129
compressor isothermal efficiency

efficiency of the isothermal expander

adiabatic efficiency of the loss absorbing expanders

VII. REFERENCES

Von P. Grassmann and J. Kopp, "On the Choice of Temperature Difference

and Heat Transfer Coefficient in Heat Exchangers, Kaletechnik 9^, No. 10

(1957), pp. 306-308

Voth, R. 0. and D. E. Daney, "H2 Liquefaction: Effects of Component

Efficiencies," 10th Intersociety Energy Conversion Engineering Conf.,

IEEE, Inc., New York, Catalog No. 75CH0 983-7 TAB (1975).

130
NBSIR 78-877, Part Four

Helium Refrigerator Reliability Data Bank

T. R. Strobridge

In 1973, NBS was commissioned by the Fermi Laboratory to visit the five large

industrial helium plants in Kansas and Oklahoma to gather data and assess the

operating history of helium expansion machines. Two of the facilities separate

helium from natural gas only, while there are four helium liquefiers at the other

three locations. These, the world's largest helium liquefiers, are roughly equivalent to a refrigerator of up to 3 kW capacity at 4.2 K. The unpublished result

was one of the first surveys of operating and failure history for helium liquefier

expansion engines, heat exchangers, compressors, purification equipment and auxiliary equipment of this size; however, the data was only used as a basis for recom- mendations, and was never published.

Recently, ERDA-Power Transmission let two identical contracts to private

industry to assess and study existing concepts and methods of cryogenic refrigeration for superconducting transmission cables. As part of the work, many of the

major laboratories in this country and Europe were visited or contacted to gather

operating and failure histories of helium refrigerators and liquefiers [Kadi and

Longsworth], [Manatt, et al.]. Thus, the failure rate data has been extended and

much other useful information was presented. In addition, both reports concluded

that a refirgerator reliability data bank should be established to collect as much information as possible so that the maximum benefit may be realized from the

reliability engineering sciences. NBS has also informally suggested on past

occasions that this data bank should be established.

At the start of this program, our technical staff was uncertain about possible legal problems in the collection and public dissemination of this type of data.

In discussing this broad legal question, we found that such a program already

exists: The Government-Industry Data Exchange Program (GIDEP).

1 . GIDEP

The GIDEP (Government-Industry Data Exchange Program) is a cooperative activity

between Government and Industry participants seeking to reduce or eliminate expenditures of time and money by making maximum use of existing knowledge. The program provides a means to exchange certain types of technical data essential in the

research, design, development, production and operational phases of the life cycle

of systems and equipment.

In 1959, the Commanders of Army, Navy and Air Force ballistic missle programs

began the Interservice Data Exchange Program (IDEP) which also included the ballistic

missle contractors. The National Aeronautics and Space Administration (NASA) and

the Canadian Department of Defence noted the value of the program and joined in 1965.

In 1970, by agreement of the Joint Logistics Commanders of the Military Services and NASA, the Navy assumed the responsibility for consolidated program

management of the renamed Government-Industry Data Exchange Program (GIDEP).

Pursuant to that same agreement, the program has been greatly expanded both in

scope and quantity of data handled. Today, GIDEP is an efficient, working system

for communications and the exchange of specialized information among government

and industry activities.

131
The program is centrally managed and funded by the Government. Its participating organizations are: The United States Army, Navy, Air Force, Marine

Corps, Defense Supply Agency, General Services Administrations, National Aeronautics and Space Administration, Federal Aviation Administration, Energy Research and Development Administration, and the National Security Agency as well

as the Canadian Department of Defence and hundreds of industrial organizations.

Previously GIDEP was restricted to government activities and government con- tractors. Now, because of government emphasis on commercial off-the-shelf items,

any activity which uses and/or generates the types of data GIDEP exchanges may

be considered for membership. The program specifically excludes classified and

proprietary information.

Participants in GIDEP are presently provided access to the four major data

banks listed below. The proper utilization of these data banks can assist in

the improvement of quality and reliability and reduce costs in the development

and manufacture of complex systems and equipment.

Engineering Data Bank Metrology Data Bank

Failure Rate Data Bank Failure Experience Data Bank

2. Participation

It was suggested to us that the best way to get our Refrigeration Reliability operation started is to attend the scheduled GIDEP meeting and workshop during the first week of October, 1977. This was attended not only by the

GIDEP staff, but also by industrial and government users of their services. We

arranged to meet with the GIDEP Officer-in-Charge during this meeting.

In the meantime we have visited the GIDEP representative of a local aero- space company who has been very cooperative in discussing details of the program

operation and in opening his files for our study. We further have discussed the

operation with the NBS legal advisor, who tells us that collecting the data for

input to the reliability sciences and even disseminating the data with generic

names is permissible within the context of the GIDEP program.

After attending the workshop during the first week of October, the next

step of our work will be to design the data collection system for this particular

project within the GIDEP framework and format.

132
NBS-114A (REV. 7-73)

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET

1. IMUU.K A 1 ION OK Ki:i'OK 1 NO. NBSIR 78-877

2. Gov 'l Ac i ( s s 1 on No.

3. Kt i ipu iu's Ai c < s s lull Ni).

4. rn i.i; and sum i n.i HELIUM HEAT TRANSFER AND REFRIGERATION IN SUPPORT OF

MAGNETIC FUSION ENERGY SYSTEMS

5. Pub! i< ai ion Dau- February 1978

6. Performing Organization C,ode

275

7. AUTHOR(S) V. Arp, J. A. Brennan, P. J. Giarratano,

W. R. Parrish, W. G. Steward, T. R. Strobridge, and R. 0. Voth

8. Performing Organ. Report No.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No. 2750555

11. Contract/Grant No.

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Department of Energy

Division of Magnetic Fusion EnergyWashington, D.C. 20545

13. Type of Report & Period

Covered

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

16. ABSTRACT (A 200-word or less factual summary of most si^ificant information,

bibliography or literature survey, mention it here.)

This is the first year-end report on a program of

ing data in support of magnetic fusion energy projects

four parts : (1) an assessment of the cryogenic engine

community, and recommendations for needed work for tha

gress on measurement of transient helium heat transfer

magnet stability analysis, (3) presentation of a nev/ly

for analyzing the efficiency of helium refrigerators o identifying sources of Inefficiency, and (4) progress

on refrigeration system reliability. The technology a vision of a preliminary version dated May 13, 1977, wi

ceived from both within and outside the MFE community.

If document includes a significant

Studies on cryogenic engineerThe report is divided into

ering data base used in the MFE

t base, (2) experimental pro-

; the data are of importance for

developed general technique

f any configuration and thereby

towards setting up a data bank

ssessment, item (1), is a reth inclusion of feedback reBoth the technology assessment. Part One, and the efficiency study. Part Three,

have been formatted for individual publication subsequent to this report. The

technology assessment will be distributed to those who received copies of the preliminary edition. The efficiency study will be submitted for publication in

Cryogenics. Because of our fairly recent publication on transient heat transfer

(Part Two, appendix A), it is appropriate to do further research in this field before

writing it up for a follow-up publication in an outside journal.

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the {[rst hotter of the first key word unless a proper name; separated by semicolons

)

Efficiency; heat transfer; helium; pressure measurement; refrigeration; stability;

superconducting magnets; transient heat transfer.

18. AVAILABILITY rXXUnlimited

[23! Official Distribution. Do Not Release to NTIS

Order From Sup. of Doc, U.S. Government Printing Office Washington, D.C. 20402, SD Cat. No. C13

[Xl Order From National Technical Information Service (NTIS)

Springfield, Virginia 22151

19. SECURITY CLASS

(THIS REPORT)

UNCLASSIFIED

20. SECURITY CLASS

(THIS PAGE)

UNCLASSIFIED

21. NO. OF P.AGES

138

22. Price

$7.25

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