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UNITED STATES DEPARTMENT OF COMMERCE

C. R. Smith, Secretary NATIONAL BUREAU OF STANDARDS • A. V. Astin, Director

R. H. Kropschot, B. W. Birmingham, and

D. B. Mann, Editors

Institute for Basic Standards

National Bureau of Standards Boulder Laboratories Boulder, Colorado 80302

^ , National Bureau of Standards, Monograph 111 , Issued October 1968

For sale by the Superintendent of Documents, U.S. Government Printing Office Washington, D.C. 20402 - Price $2.00
rSlonal Bursa-j cf !'::.nd3rds NOV 2 9 1968

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Abstract

The discovery of the element helium was made just 100 years ago in 1868. Today helium is produced in large quantities and

used in many technological applications. This treatise is a source document containing information on helium resources, production,

conservation, thermodynamic properties, liquefaction and re- frigeration techniques, transportation and storage of liquid and

safety requirements. It also contains a discussion of uses for liquid and cold gas in cryoelectronics, superconductivity, bubble chambers,

cryopumping and missile and space systems. The book brings

together articles by noted authorities in cryogenic technology in which they discuss their specialized field in great depth.

Key Words: Conservation, cryoelectronics, cryogenics,

cryopumping, helium, liquefaction, magnets,

purification, refrigeration, safety, storage, thermodynamic properties, transportation.

Library of Congress Catalog Card No. 68-61553

I

II
Foreword

One hundred years ago, in 1868, J. N. Lockyer discovered a mysterious element in the chromosphere of the sun. Since Lockyer and other

well-known scientists of the time thought that it existed nowhere else,

the element was given the name "helium" from the Greek word "helios"

or "sun." Less than 25 years later, however, W. F. Hillebrand of the U.S. Department of the Interior obtained an unidentified gas from the

mineral uraninite. Sir William Ramsay, learning of Hillebrand's work,

obtained a similar gas from the mineral cleveite. Ramsay removed

the oxygen and nitrogen he found present in this gas and examined the

residue. This residue, helium, proved, after 27 years of uncertainty,

that the sun element did indeed exist on earth. But the work of these three pioneers of science remained relatively

unheralded until a method was developed to obtain and use this sub- stance in volume. The key to this development came in 1905, when a

natural gas discovered in Kansas was found to contain nearly two percent

helium. In 1918, exactly 50 years after its discovery, helium was first extracted in volume from natural gas. From then until today, the use and

importance of hehum has expanded at an ever-increasing rate. Helium has many uses; for example, as a gas in lighter-than-air

ships, as a pressurant and control gas in aerospace vehicles, as a shield- ing gas in welding, and as a coolant gas in reactors. One of the most

interesting uses is in the field of cryogenics.

Because of its very low boiling point, 4.2 K, helium is needed in

liquid form to provide the refrigeration for all types of cryoelectronic

and other devices, especially those involving the phenomenon of superconductivity. Helium itself is scientifically interesting because of its transformation to hehum ii at 2.18 K and the absence of a triple point.

The unusual behavior of hehum ii, such as superfluid creep, abnormal

III
viscosity, and sound propagation, has been studied extensively by low

temperature physicists.

This book, however, is aimed at articulating the developments

in the last decade in the technology of liquid helium, for much has been

written previously on the physics of liquid helium. In fact it is now

becoming apparent that the 1960's will be remembered as the era of a developing liquid heHum technology.

I am pleased, therefore, to have this book available at this time

not only to share in the Helium Centennial Celebration but hopefully

to act as a stimulus to an increasingly rapid development of the

technology.

Washington, D.C., W. R. HiBBARD, Director,

March 1968. U.S. Bureau of Mines.

IV
Preface

When one thinks of liquid helium and its associated technology,

the names of two pioneers come immediately to mind. The first, of course, is the Dutch scientist KamerUngh Onnes, since he first suc- ceeded in hquefying hehum at the University of Leiden in The Netherlands in 1908. Several attempts were made prior to 1908 to liquefy helium, notably

by Olszewski and Dewar, but these were unsuccessful. In fact, Kamerlingh Onnes was to make several abortive attempts to hquefy heUum by

using a technique of cooUng compressed helium gas with liquid hydrogen

and then making a rapid expansion. He finally used a circulation tech- nique where compressed hehum gas was cooled to near the triple point

temperature of hydrogen and then piped through a regenerative heat

exchanger prior to expansion. By this "boot-strap" cooling in the regenerative heat exchanger he eventually obtained helium cold enough to form liquid upon expansion.

It was not until the early and middle 1920's that several other

laboratories reported success in liquefying helium. Simon and Kapitza

made contributions in the early 1930's, but their developments did not seem to have the impact on liquid heHum technology as did our second

pioneer — Samuel C. ColUns of the Massachusetts Institute of Technology. Collins developed a highly successful helium liquefier after World War II, and over 350 have now been sold commercially, putting

liquid helium within reach of virtually every low-temperature physicist.

ColUns was successful, with what we now know as the "Collins

Cryostat," because he perceived that there were two major problems

associated with earlier attempts to produce liquefiers, and he provided

solutions to these problems.

One problem pertained to imperfect vacuum technology, which

would probably be considered trivial today but was most troublesome

two decades ago. Collins circumvented the need for a high vacuum,

and thus leak-tight joints, by suspending the regenerative heat exchangers of the liquefier in an atmosphere of gaseous helium with the cold

end in the downward position. In this manner he was able to depend on

convective equilibrium for insulation, and small leaks to the helium

atmosphere made little difference.

Another problem pertained to the hfe and effectiveness of the

expansion engine. Earlier developers had used material combinations

such as a bronze piston in a stainless steel cylinder, and we now realize

that these combinations were relatively poor from the standpoint of V
both wear and clearance. Collins chose nitrided steel for both the piston

and cylinder, with the result that not only was the metal better from an

operational standpoint but clearances could be reduced by an order of magnitude, yielding an overall improvement in performance.

I would hke now to identify another individual who is not nearly

as well known as Kamerlingh Onnes or Collins but who I feel had a profound influence on large-scale Hquid hehum technology as we know

it today. This individual was the late Ray E. Brown, who spent his entire

professional career in the United States Department of the Navy. Since

his contributions may be less well known than those of Kamerhngh Onnes

or Collins, I would hke to highhght just a few of them in this Introduction.

In 1927 Ray Brown became associated with the lighter-than-air

ship program of the Navy Bureau of Aeronautics, and the rest of his working life was synonymous with the growth of helium as a useful

material — first, as a lifting gas for airships, and finally, as an important

commodity for many military, medical, industrial, and scientific applications.

Ray Brown played a prominent role in the design of railway tank

cars, automotive trailers, and other methods of transportation for compressed helium. He participated in the development of storage facilities for helium at the Navy's lighter-than-air facilities. During World War II, he managed almost singlehandedly the procurement and distribution of

helium for the Navy's dirigibles and for other military operations.

By virtue of his association with the gaseous helium field, he main- ' tained a continuing awareness of all aspects of helium behavior including

that of liquid helium, and after World War II he conceived the idea of

using helium in the liquid phase for economy in transportation. He

pointed out that when helium is shipped in the compressed form in steel cylinders the mass of helium is only one to two percent of the total mass

of the container; therefore, most of the transportation charges are for transporting steel, and only a small portion is for transporting hehum.

Thus, he suggested that hehum either in the liquid form or as a cold,

dense gas could be used to increase this percentage so that a much

greater portion of the transportation cost would be for the hehum.

Ray Brown came to the National Bureau of Standards in the mid1940's to propose the development of a suitable hquid hehum technology

to support his suggestion. However, it was not until about 1954 that the

National. Bureau of Standards was in a position to assist him in this venture at its then relatively young Cryogenics Laboratory in Boulder,

Colo. During the next several years the National Bureau of Standards

with Navy support worked on many aspects of the hquid hehum problem

and laid the groundwork for the large-scale hquid hehum ventures which

are operating today and explained in more detail in the main text of

this book.

VI
Kamerlingh Onnes, Collins, and Brown are notable for their in- dividual efforts. In addition, the team effort of the National Bureau of Standards, and certainly the U.S. Bureau of Mines, whose many years

of work resulted in the development and perfection of a cryogenic

process for separating helium gas, have provided the basis for the

industrial interest in a large-scale Hquid heUum technology. Another

significant stimulus to this technology is the Bureau of Mines' conservation program for helium, which makes it economically attractive for industrial firms to consider large-scale liquefaction. Accordingly, the

editors of this Monograph feel that its publication during this centennial

year celebration of the discovery of the element helium is most timely.

I will not attempt to summarize either the entire book or portions

of it in this Introduction; however, I would like to point out that the

editors deliberately omitted discussion of superfluid helium since this topic has been amply covered in other books and articles authored by

many notable scientists working in the field of low-temperature physics.

We have also omitted a discussion of hehum heat transfer, both superfluid and supercritical, since the data are so scanty at this time. Research

efforts are currently underway to provide such data, but it will probably

be at least one to two years before comprehensive articles can be

written. The appHcations described in chapter 8 were selected on the

basis of significant uses and are certainly not intended to be all-inclusive; however, it is hoped that they do include the applications which involve

large expenditures.

The use of proprietary names, products, and processes in this Monograph, though contrary to the usual NBS practice, is essential to the proper understanding of the work presented. These proprietary

references are made with no intent to endorse or condemn specific

products or processes, but are made solely with the intent to be representative and comprehensive in the field of liquid helium technology.

The National Bureau of Standards has officially adopted the Inter- national System of Units, abbreviated SI for Systeme International.

It has been the Bureau's policy to employ the SI in all its publications

except where the use of these units would obviously impair communication or would reduce the usefulness of a publication to its primary re- cipients. This Monograph does not consistently use SI units because

much of the work in the field of liquid helium technology has been based

upon other systems of units. Appropriate conversion factors are included

where necessary. We urge that speciaHsts and other users of data in this field ac- custom themselves to SI units as rapidly as possible. B. W. Birmingham, Deputy Director,

Institute for Basic Standards/Boulder, Colo.

VII

Contents

Page Foreword Ill Walter R. Hibbard, Department of the Interior, Bureau of Mines,

Washington, D.C. 20240

Preface v Bascom W. Birmingham. NBS-Institute for Basic Standards,

Boulder. Colorado 80302 Chapter 1. Helium Resources, Production, and Conservation 1 William M. Deaton and Paul V. Mullins, HeHum Activity, Bureau

of Mines, Amarillo, Texas 79106 Helium resources 3 Helium production history 6 Helium uses and markets 8 Helium conservation 10 Helium production process 12 Helium production by the Bureau of Mines 16 Chapter 2. Properties of Helium 33 Douglas B. Mann, Cryogenics Division, NBS-Institute for Basic Standards, Boulder, Colorado 80302

Pressure-volume-temperature surface 35 Thermodynamic functions 40

Transport properties 51 Discussion 52 Chapter 3. Liquefaction Techniques 55 Samuel C. Collins, 28 Village Hill Road, Belmont, Massachusetts

02178

Principles of refrigeration 56 Common helium hquefiers 62 Hehum liquefaction facilities 66 Chapter 4. Refrigeration 77 Thomas R. Strobridge, Cryogenics Division, NBS-Institute for Basic Standards, Boulder, Colorado 80302

Introduction 78 Thermodynamic relationships 79

Refrigeration cycles and methods 83 Components and present-day refrigerators 122

Efficiencv, weight, and volume of 4 K refrigerators 142 Chapter 5. Storage, Distribution, and Handling 153

L. E. Scott, R. C. Van Meerbeke, B. S. Kirk, and G. C. Nubel, Air Reduction Company, Murray Hill, New Jersey 07971

Introduction 155

Storage systems 156 Vented containers 158 Nonvented containers 166

Insulation 170

Sloshing and thermal oscillation heating effects 192 IX
Contents — Continued

Page Chapter 5. Storage, Distribution, and Handling — Continued

Container design 196

Transfer of liquid helium 213

Large scale liquid and gaseous helium distribution 239 Chapter 6. Safe and Efficient Use of Liquid Helium 255 Hugh M. Long and Paul E. Loveday, Linde Division, Union Carbide

Corporation, Tonawanda, New York 14150

Introduction 256 Handling liquid heUum 257 Helium liquefaction 258

Transport and storage 261 Transfer 274

Safety in cryostat design and use 279 Summary 286 Appendix A. Sizing of relief devices 288 Chapter 7. Cryoelectronics 299 James W. Meyer, MIT Lincoln Laboratory, Lexington, Massachusetts 02173

Robert A. Kamper, Cryogenics Division, NBS-Institute for Basic

Standards, Boulder, Colorado 80302

Introduction 300 Quantum devices 301 Semiconductor devices 306 Superconductor devices 310 Working devices 322 Chapter 8. Applications 327 Robert A. Kamper, Richard H. Kropschot, and Alan F. Schmidt,

Cryogenics Division, NBS-Institute for Basic Standards, Boulder, Colorado 80302 Magnets 329

Linear accelerators 337 Bubble chambers.... 339 Cryopumping 343 Non-condensable gas injection 351 Other applications in missile and space systems 357 Author Index 363

Subject Index 369

X
CHAPTER 1

Helium Resources, Production, and

Conservation

William M. Deaton ' and Paul V. Mullins 2

Page 1.1. Helium Resources 3

1.1.1. Helium occurrence 3

1.1.2. Bureau of Mines helium survey 3

1.1.3. Estimated helium resources 4

1.1.4. Geographical location of resources 5

1.2. Helium Production History 6

1.2.1. Production by the Bureau of Mines 6

1.2.2. Production by private industry 7

1.2.3. Foreign production 7

1.3. Helium Uses and Markets 8

1.3.1. Principal uses 8

1.3.2. Principal consumers 8

1.3.3. Geographical distribution of market 10

1.4. HeUum Conservation 10

1.4.1. Helium conservation program 10

1.4.2. Volumes to be conserved 11

1.4.3. Geographical location of plants, pipelines, and storage

field 11

1.4.4. Future helium supply 11

' Formerly Chief Technologist (now retired). Helium Activity, Bureau of Mines, U.S. Department of the Interior, Amarillo, Tex. 79106.

* General Manager, Helium Operations, Bureau of Mines, U.S. Department of the Interior, Amarillo, Tex. 79106.

1
2 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

Page 1.5. Helium Production Process 12

1.5.1. Bureau of Mines helium production plants 12

1.5.2. Private industry helium production plants 13

1.5.3. Private industry helium conservation plants 14

1.5.4. Other helium separation methods 15

1.6. Helium Production by the Bureau of Mines 16

1.6.1. Description of Bureau of Mines helium plant process 16

Gas pre-treating 16

Crude helium separation 18

Helium purification 19

Nitrogen refrigeration cycle 21

1.6.2. Helium quality and quality control 22

Composition of produced helium 22

Production monitoring methods 22

Methods used to prevent contamination after purifica- tion and to assure purity of helium shipped 23

Helium trace-impurity analysis method 24

Quality control at point of use 25

1.6.3. Helium handling and storage 25

Helium piping, valves, and compressors 25

Fixed storage 26

1.6.4. Hehum transportation 26

Railway tank cars 28

Highway tube trailers 28

Small compressed gas cyhnders 28

Liquid helium transportation 28

Hehum transportation by pipehne 29

1.6.5. Cost of hehum shipment 29

Cost of gaseous hehum shipment 29

Cost of liquid helium shipment 30

1.6.6. Helium prices and regulations governing sale 31
HELIUM RESOURCES 3

1.1. HELIUM RESOURCES

1.1.1. Helium Occurrence

Helium is present in the earth's atmosphere to the extent of about

1 part in 200,000 or 0.0005 mole percent. It is abundant in the sun and

stars, thus is not a rare element so far as the whole universe is con- cerned. It is a rare element on the earth only in the sense that it is generally so widely dispersed that it would be difficult to recover it from the

atmosphere as a concentrated gas. Helium is found in minor amounts in most natural gases produced in the United States. In many cases, the

helium content is only a few hundredths or a few tenths of 1 percent;

but natural gas produced in large volumes in some areas contains 0.4

to 0.6 percent, and some natural gas produced in smaller but significant

volumes contains up to 2 percent. In a very few instances, the helium content exceeds 2 percent. The highest content observed is about 9 percent.

Helium-bearing natural gases usually contain appreciable amounts of

nitrogen and thus are relatively low in heating value. Helium-bearing

gas from most sources, however, is marketable as fuel gas. Helium is also found in natural gases produced outside the United

States. No thorough survey of foreign natural gases has been made.

Natural gases produced at some locations in Canada contain helium in significant amounts and some helium is produced there. Helium has

been reported in natural gases in the U.S.S.R., in Japan and may be

expected to be found, at least in small amounts, in many natural gases

throughout the world.

1.1.2. Bureau of Mines Helium Survey

For more than 40 years, the Bureau of Mines has conducted a con- tinuous "helium survey" to determine and evaluate occurrences of helium in natural gas in the United States. The survey is conducted by

obtaining and analyzing samples of natural gas from new wells drilled

in the exploration for oil and gas. Several hundred samples are obtained

and analyzed each year, and complete analyses with well data are published periodically [1-7].^

' Figures in brackets indicate the literature references at the end of this chapter.
4 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

1.1.3. Estimated Helium Resources

Helium contained in the helium resources of the United States was

estimated in 1966 to be about 200 billion cubic feet. (See table 1.1.)* The resources considered were limited to those where the helium content

of the gas is 0.3 percent or more, and the estimated amount of helium at a particular location is 100 million cubic feet or more. The total recoverable

helium in these resources, however, is less than 200 billion cubic feet.

Table 1.1. Estimated volumes of helium in the principal helium resources in the United

States {as of January 1966)

Geographical location Billion cubic feet of helium

(at 14.65 psia and 60 F)

Rocky Mountain area: Arizona, Colorado, Montana, New Mexico, Utah,

Wyoming 25.5 Midcontinent area: 169.2 Total 194.7

9.0 Total 203.7

Many factors affect the recovery of hehum. Hehum is an inherent

part of the natural gas and, in most instances, becomes available for recovery only under the conditions controlling the production of natural gas for fuel purposes. Thus, availabihty of helium for recovery is generally dependent on the rates of production and other conditions established in marketing natural gas. Also, it is important that the source

gas be available for a long enough time to justify building a plant.

To accommodate practical and economical recovery, it is desirable

and, in most cases, necessary to select a focal point where large volumes

of hehum-bearing natural gas have been brought together, after gathering

from the source wells, and thus are available at a single point for hehum

extraction. These desired conditions are provided at some points on gas transmission lines that transport helium-bearing gas to market. Also,

*Editors' NOTE: The system of units was chosen by the authors and represents current industiial convention. Conversion tables to the mks system are given at the end of the chapter.
GEOGRAPHICAL LOCATION OF RESOURCES 5

the potential net recovery at any point depends on optimum plant size and efficiency, on uniformity and continuity of gas availability, and on

other factors. Taking into account these various factors, it is estimated

that in 1966 the amount of recoverable helium in the United States was

about 180 billion cubic feet.

1.1.4. Geographical Location of Resources

The geographical location of the principal sources of helium in the

United States is shown in figure 1.1. Table 1.1 indicates that about 87

percent of the estimated hehum resources is in the Texas Panhandle,

Oklahoma Panhandle, and Southwestern Kansas. Natural gas from parts

of the fields in this area has been transported to market for various

periods, ranging from 15 to 35 years, and withdrawal is continuing at a

rate that is expected to cause major depletion of the gas in these fields in 20-25 years. Hehum extraction plants existing prior to 1962 had

capacity to extract only about 10 percent of the more than 4 billion cubic

feet a year of helium that was contained in the gas being withdrawn from

this important area. As a result of the Government's helium conservation program (see section 1.4), much of this helium is now being extracted

and saved as a conservation measure.

Figure LL Geographical location of principal helium resources in the United States.
6 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

1.2. HELIUM PRODUCTION HISTORY

Helium has been produced in the United States in appreciable quantities since 1918. The quantities produced were relatively small except

during World War 11. In 1948, a sharp upward trend began as new uses were made of helium in some rapidly expanding Government and civilian

activities. This trend has continued. The annual production of helium

from 1920 to 1966 is shown in figure 1.2. 1.2.1. Production by the Bureau of Mines

The Federal Government pioneered helium production in 1918.

Virtually all helium production in the period 1918-61 was by the U.S. Department of the Interior's Bureau of Mines and was to meet the needs

of Federal agencies. Bureau plants are situated near Amarillo, Tex.;

Exell, Tex. (30 miles north of Amarillo); Keyes, Okla.; Shiprock, N. Mex.;

and Otis, Kans. All helium production has been accomphshed by use

of cryogenic processes, and the helium has been extracted from hehumbearing natural gas. The enabling legislation for early helium operations of the Bureau

of Mines was the Hehum Act of March 3, 1925 (43 Stat. 1110). This Act was amended September 1, 1937 (50 Stat. 885), to authorize sales of

1940 I94G 1950 TIME, colendor yeors Figure 1.2. Helium produced and sold by the Bureau of Mines, 1920-66.
PRODUCTION BY PRIVATE INDUSTRY 7

helium for medical, scientific, and industrial uses. A further amendment of September 13, 1960 (74 Stat. 918) provided for long-term helium

purchase contracts to accommodate an extensive helium conservation

program, which is described later. 1.2.2. Production by Private Industry

Private industry helium production plants were few in number and

contributed Mttle to the total volume of helium produced in the period

1918-61. Only one private company operated plants in this period. This

company had two small plants — one in Kansas and one in Colorado — for short periods during 1928-37. But in the early 1960's, private industry

again entered the helium production field. A private plant with a potential

output of 70 million cubic feet annually was constructed in eastern Arizona and started operating in 1961. Another plant was completed

near Otis, Kans., in 1966, with a planned annual output of 150 million

cubic feet. A third plant near Elkhart, Kans., which has a production

capacity of 140 million cubic feet, was completed and placed in operation

in 1966. In 1962-63, five large helium extraction plants were constructed in the Texas Panhandle and western Kansas by private industry under

contracts with the Federal Government to extract helium from natural

gas being transported to fuel markets. The product from these plants is a crude (unrefined) helium-nitrogen mixture containing 50-85 percent

helium, and is purchased by the Federal Government under long-term

contracts. Most of it is stored, for future use, in a natural gas underground

storage reservoir near Amarillo, Tex. (see sec. 1.4). 1.2.3. Foreign Production

One helium plant with a potential output of 12 miUion cubic feet per

year is in operation in Saskatchewan, Canada. This plant was built in 1963. Except for the production from this plant, there is no significant

helium production outside the United States. Small quantities have been

produced in Europe as a by-product of air separation. Some reports

indicate hehum has been produced from natural gas in the U.S.S.R., but

few details are available. Announcement has recently been made of expected production of hehum in Europe [8]. This helium would be obtained from natural gas produced from wells in the North Sea off the coasts of England and The

Netherlands. Also hehum may be available in the uncondensed gas from

a natural gas hquefaction plant supplying liquefied natural gas for shipment from Algeria to France and England.

306-922 0-68—

2
8 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

1.3. HELIUM USES AND MARKETS

1.3.1. Principal Uses

With minor exceptions, all helium produced during the first 20 years

of its production was used as buoyant gas for inflation of blimps, dirigibles, and balloons — principally for military operations. Small quantities were used in synthetic breathing mixtures in deep sea diving operations.

Following authorization of non-Federal sales in 1937, small quantities were used in synthetic breathing mixtures for temporary relief of respiratory disorders and for reduction of explosive hazards in anesthesia.

Shortly thereafter, the use of helium in shielded-arc welding began and

then helium found an important use in the early nuclear energy investi- gations and developments. Nevertheless, about 99 percent of the annual

hehum production before 1946 was used in bhmps, dirigibles, and

balloons. Non-Federal sales were small. In the late 1940's, heUum sales increased sharply and a considerably different pattern of helium use began to develop. A major factor was

the Bureau of Mines production of helium of better than 99 percent

purity, rather than the former 98.3 percent purity. The higher purity was of appreciable benefit in some uses. Helium users demanded still higher purity. The purity of produced helium was raised, first to about

99.5 percent and, shortly thereafter, to nearly 100 percent. This was a strong stimulus to helium usage. Larger quantities were used for shielded- arc welding and nuclear energy development, and new uses developed

and grew rapidly. In the 1950's, large volumes of helium were used in the

development and operation of ballistic missiles and in the initial program

for space exploration. At the same time, helium was finding a variety of uses in rapidly expanding research programs. The usefulness of helium

in these appUcations is due to its unique physical properties, especially

its inertness, its low density, and its very low liquefaction temperature.

A survey in 1962 indicated that the approximate division among various uses was as shown in figure 1.3. 1.3.2. Principal Consumers

Federal agencies have always been the principal consumers of hehum. Prior to and during World War II, the Federal agencies — and

primarily the Navy Department — used virtually all hehum produced.

This pattern began to change shortly after World War II as the Atomic

Energy Commission and commercial consumers began to use appreciable volumes of hehum. After 1950, the development programs for

ballistic missiles and space exploration resulted in the Air Force and
PRINCIPAL CONSUMERS 9

Figure 1.3. Percentage breakdown of

principal helium uses {1962).

LEAK

DETECTION 2%

Other 0.1%

-Weather Bureau

2.0 7o Figure 1.4. Percentage breakdown of principal users of helium

{1966).

National Aeronautics and Space Administration, as well as the Atomic

Energy Commission, becoming consumers of large quantities of helium.

During the period 1950-66, a fairly consistent pattern prevailed of about

two-thirds going to Federal consumers and one-third to non-Federal consumers. Surveys indicated that 50 percent or more of the helium

sold to non-Federal distributors and consumers was used on Federal
10 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

contracts, and this made Federal consumption approach 90 percent of the total market. The 1966 survey indicated the division among Federal

agencies and commercial consumers was as shown in figure 1.4. 1.3.3. Geographical Distribution of Market

From the beginning of production, the geographical location of major

markets has been on the east and west coasts of the United States. In

1966, about 25 percent went to points on the east coast and 45 percent to points on the west coast. The remaining 30 percent was used at locations

in the Gulf Coast and interior States.

1.4. HELIUM CONSERVATION

1.4.1. Helium Conservation Program

Rapidly growing importance and use of helium prompted a critical study in the late 1950's of its future supply and demand. This led to a conclusion by the Federal Government that it should engage in an extensive helium conservation program. In September 1960, the program was

approved and legislation was passed by the Congress to accommodate it. Plans for the conservation program were based primarily on the concept

that helium extraction plants could be constructed at strategic points on

the principal pipelines transporting helium-bearing gas to market, and

that except for that portion needed to meet current demand the helium

would be stored underground by the Government for future use when

needed. Private industry was to be encouraged to finance, construct,

and operate the plants and would sell its output to the Federal Government under long-term contracts. The product produced would be a "crude" or unpurified helium, with a composition of 50-85 percent

helium and the remainder chiefly nitrogen.

The conservation program got underway in 1961 with the execution

of four contracts with private industry to build and operate five plants

situated near the towns of Bushton, Ulysses, and Liberal in Kansas; and

in Sherman County, and near Dumas in the Texas Panhandle. All of

these plants were built and in production by mid-1963. The Bureau of

Mines built a 425-mile pipeline gathering system with connections to the

above plants to take their full output and to deliver it to one or more of

the Bureau's plants for purification and marketing, or to an underground

storage reservoir in the Clififside Gas Field near Amarillo. The storage

reservoir is a partly depleted natural gas reservoir which, for over 30

years, has been the source of helium-bearing natural gas processed at the Amarillo Hehum Plant.
VOLUMES TO BE CONSERVED 11

1.4.2. Volumes to be Conserved

By the conservation program, it is estimated that about 78 bilhon

cubic feet of hehum will be extracted and saved over a 22-year period

from helium-bearing gas en route to fuel markets. It is estimated that about 36 billion cubic feet of this helium will need to be purified to meet

market demand during this period; and the remainder, or about 42 billion cubic feet, will be placed in the storage reservoir for use beyond 1985.

HELIUM PIPELINE

O-US BUREAU OF MINES HELIUM PLANTS

I - AMARILLO

2- EXELL

3- KEYES

4 - OTIS

5 - NAVAJO

-CONSERVATION HELIUM PLANTS

A -PHILLIPS PETROLEUM CO, DUMAS

B - PHILLIPS PETROLEUM CO, SHERMAN

C- NATIONAL HELIUM CORP \

D- CITIES SERVICE HELEX, INC

E - NORTHERN HELEX CO C'-AREAL BOUNDARIES OF PRINCIPAL HELIUM-BEARING NATURAL GAS FIELDS SERVING PLANTS

Figure 1.5. Geographical location of Bureau of Mines helium plants, conservation helium

plants, pipelines, and helium storage reservoir. 1.4.3. Geographical Location of Plants, Pipelines, and Storage

Field

The geographical location of helium plants, pipelines, gathering

system, and the storage field for conserved helium is shown in figure 1.5. 1.4.4. Future Helium Supply

Future helium supply for important uses would seem to be assured

by the helium conservation program for 30 to 50 years, and perhaps

longer. The volumes extracted and saved might be increased appreciably

if conditions justify and action is taken before about 1970, for there are

additional quantities available which could be recovered economically.
12 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

1.5. HELIUM PRODUCTION PROCESS

1.5.1. Bureau of Mines Helium Production Plants

All of the Bureau of Mines helium production plants extract helium

from natural gas and use a cryogenic process. A block diagram of the

process is shown in figure 1.6. The process is described in detail later

Helium - bearing natural Processed natural gas 450psi

Condensate Removal

coz

Removal

Water

Removal

Tube Trailer

Treated Gas 435psi Processed Gas 190 psi —t \—: —

L

Crude Helium

Separation

''Cold Box"

Crude Helium Compre ssion 2700 psi

Helium

Purification

Grade -A Helium Compression 4000 psi

Railroad Tank Car

450psi

Gas

Compression 450 psi

Nitrogen

Refrigeration 600 psi

Standard Cylinders Figure 1.6. Block diagram of Bureau of Mines helium plant process.

but briefly it consists of the following. The natural gas is supplied from

the source pipeline at a pressure of about 435 psi (pounds-force per

square inch gauge or Ibf/in^) or, if not available at this pressure, is compressed to 435 psi upon entering the plant. The gas is pretreated to remove hydrocarbon condensate, carbon dioxide, and water before it enters the low-temperature equipment. In this equipment, called "cold

boxes," the entire gas stream is cooled to about —250 F and liquefied

except for the helium and a small percentage of the nitrogen; and after separating the uncondensed helium and nitrogen, the liquefied natural gas is returned through heat exchangers in counterflow to the incoming

stream to the cold box outlet. Leaving at near ambient temperature, the

processed gas is returned to the gas pipeline from which it was taken

and goes on to fuel markets.
PRIVATE INDUSTRY HELIUM PRODUCTION PLANTS 13

The helium-nitrogen product, called "crude heHum," generally has

a composition of about 70 percent helium and 30 percent nitrogen. It returns through heat exchangers against the incoming natural gas stream

and leaves the cold box at ambient temperature and usually at 150-200

psi. This stream is further processed, as later described, to purify it to Grade-A hehum quahty (99.995 mole percent purity) by a high-pressure

(2700 psi) and low-temperature (—320 F) cryogenic process. The Bureau of Mines helium plants differ widely in natural gas processing capacities, which range from 6 to 130 million cubic feet of gas per day. The helium contents of the entering natural gas streams

also differ greatly, ranging from 0.9 to 5.8 percent. The hehum production

capacities range from about 35 milHon to over 350 million cubic feet of hehum per year.

1.5.2. Private Industry Helium Production Plants

Processes used in private industry helium plants are low-temperature processes using the same general principle as that used in Bureau

of Mines plants. Just prior to 1966, the only private industry hehum plant producing

Grade-A hehum was the plant of Kerr-McGee Oil Industries, Inc., at Navajo, Ariz. [9]. The gas processed at this plant comes from a small

gas field nearby that produces an inert gas containing about 91 percent

nitrogen and 9 percent hehum.

In this plant, separation of hehum from the natural gas is first done

at relatively low pressure (about 50 psi). After cooUng to near —320 F

with liquid nitrogen, the uncondensed crude hehum product is warmed,

compressed to 2900 psi, and then purified in the same manner as later described for Bureau of Mines plants. The designed natural gas processing rate is 2.5 million cubic feet per day and the designed purified

hehum production rate is 70 million cubic feet per year. Equipment is installed at this plant to liquefy helium for shipment.

A new hehum plant near Otis, Kans., owned and operated by the

Kansas Refined Hehum Company, that processes hehum-bearing natural gas produced in nearby gas fields, was placed in operation early in 1966. The designed processing capacity of this plant is 25 million cubic

feet of natural gas per day. Its helium production output is estimated

at about 430 thousand cubic feet per day (150 milhon cubic feet per

year), and it contains a hehum liquefier of sufficient capacity to hquefy

the entire output of the plant. Shipment of the hehum is scheduled to
14 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

be made in liquid helium semitrailers of about 10,000-gallon capacities

(see chapter 5). The Alamo Chemical Company, a subsidiary of Phillips Petroleum

Company, completed and placed in operation in 1966 a helium plant near Elkhart, Kans. [10]. It processes natural gas produced from the

Greenwood Field. The designed processing capacity of this plant is 76 million cubic feet of gas per day and its heHum production capacity

will be about 140 miUion cubic feet per year. Most of the output of this plant can be hquefied and shipped in this form.

1.5.3. Private Industry Helium Conservation Plants

All the helium extraction plants producing helium for delivery to the Bureau of Mines under the Government's helium conservation pro- gram produce only a crude helium mixture of helium-nitrogen-methane

having the following range of analyses: helium 50-85 percent, nitrogen

15-50 percent, methane 0-5 percent. The plants also use a low-temperature process. Their refrigeration requirements are supplied principally by Joule-Thomson cooling obtained in throttling the main natural gas stream from a pressure of 300 or 450 psi to a return-gas pressure

of 150 to 200 psi, although some of the plants use auxiliary refrigeration

to some extent. UnUke the Bureau of Mines plants, the private industry

helium recovery plants are designed to extract liquefied petroleum gas (LPG) products from the gas being processed. These products then are

Figure 1.7. Flow diagram of National Helium Corporation s helium plant.
OTHER HELIUM SEPARATION METHODS 15

sold in the LPG market. This adds somewhat to the equipment needed

and the refrigeration requirements. A description of the National Helium

Corporation's plant, one of the conservation helium plants, showing the

processing cycle has been pubUshed [11]. Figure 1.7 is reproduced from

this article. The helium content of the gases processed by these plants ranges

from 0.40 percent to 0.71 percent, and the processing capacities of the

plants range from 200 million to 870 million standard cubic feet of natural gas per day. The amount of helium extracted by the plants

ranges from 450 million to 1,050 million standard cubic feet per year.

1.5.4. Other Helium Separation Methods

Although no plants in operation in 1966 use other than low-temperature processing to produce helium, two other methods have been suggested. One is a diffusion method involving differential diffusion of helium, in the gas to be processed, through permeable membranes or porous diffusion barriers [12-15], and the other is selective diffusion of

the helium through silica glass tubing [16]. The Bureau of Mines examined both of these methods many years ago (1921-22), but at that time

the problems associated with diffusion processes and equipment for them

made this approach less attractive than low-temperature processing. An

examination of the diffusion process, utilizing high-silica or quartz

tubing, showed that this method would be workable; but to be practical

would require a large area of high-silica or quartz media, which at that time was not available. With the progress made in diffusion processes and equipment — for example, the separation of uranium isotopes in the diffusion plant at Oak Ridge, Tenn. — and the development of synthetic diffusion media and

the large-scale production of fine capillary high-silica and pure quartz

tubing, which is selectively permeable to helium, renewed consideration

of these processes is being made.

An experimental unit using high-silica tubing has been tested by

the Research Council of Alberta, Edmonton, Alberta, Canada. Its find- ings have been reported in the literature [17]. The Standard Oil Company of Ohio has patented a process for separating helium from natural gas that uses micro size (40-50 micron)

silica-glass hollow spheres [18]. No tests using this process have thus

far been reported.
16 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

1.6. HELIUM PRODUCTION BY THE

BUREAU OF MINES

1.6.1. Description of Bureau of Mines Helium Plant Process

Bureau of Mines helium plants have various gas processing capacities ranging from 6 million to 130 million cubic feet of natural gas per day.

The capacities are such that the plants can in most cases process the

total amount of gas available to the plants. There are only a few days each

year during the winter when the natural gas pipelines are carrying more gas than the plants can process. Details of the process used follow. Gas Pretreating: Natural gas entering the plant at 435 psi passes

through a separator to remove any hydrocarbon condensate or liquid

water carried by the gas. At most plants, two gas treating units (fig. 1.8) are used to remove carbon dioxide and water vapor from the inlet gas. Two units are provided in order to give flexibihty of operation, to accom- modate seasonal fluctuations in gas supply, and to permit the Bureau to meet contractual commitments to process a fixed minimum quantity

of gas. The gas processed at the various plants of the Bureau contains

carbon dioxide ranging from 0.4 to about 3.0 percent. A solution of about

15 percent monoethanolamine, 76 percent diethylene glycol, and 9 perFlGURE 1.8. Carbon dioxide and water removal equipment at Keyes Helium Plant.
BUREAU OF MINES HEUIUM PLANT PROCESS 17

Figure 1.9. ''Cold boxes' at Keyes Helium Plant. cent water is used to remove carbon dioxide. The carbon dioxide content

of the natural gas is reduced to 0.005 mole percent or less. Hydrogen

sulfide, if present, is also removed.

Although the primary purpose of the amine-glycol solution is to remove carbon dioxide, the gas is also partially dehydrated because of the high glycol and low water content of the solution. The gas is dried

further by passing it through soUd desiccant beds where the water dewpoint of the gas is reduced to about — 100 F, which at 435 psi is equivalent to about 0.0025 pound per million cubic feet. Various desiccants

and combinations of desiccants have been used, including activated

bauxite (a naturally occurring aluminum oxide ore), activated alumina

(a synthetic product), silica gel beads, and molecular sieves. The latter

is used only as a trimmer bed to assist in lowering the water content of the gas to the above value. All of these desiccants have performed satis- factorily. The choice depends on their cost, performance, and length of service before having to be replaced. Inadvertent fouhng of the desic-
18 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

70 % Crude helium 200 psig

Processed

natural gas 190 psig Crude helium 55-175 psig

Treated natural gos_ 435 psig Figure 1.10. Typical crude helium extraction cycle of Bureau of Mines helium plant. cant beds by carryover of amine and glycol (particularly the gas entrant

portion) sometimes occurs and may justify the use of a lower capacity,

but less expensive, type of adsorbent, such as bauxite.

Crude Helium Separation: After pretreating the gas to remove

carbon dioxide and water, it goes to low-temperature equipment housed

in cold boxes (fig. 1.9) where separation of helium from the natural gas

is made. The gas stream separated is a crude helium product. Steps

of the process are illustrated schematically in figure 1.10. Treated

natural gas enters the main gas heat exchangers, 1, at about 435 psi

and is cooled first to about —60 F. It then enters a separator, 2, where

the condensed heavy hydrocarbons are separated and returned through

heat exchangers to the processed gas stream. Removal of heavy hydrocarbons from the inlet gas prevents the formation of plugs later in the

process, due to possible solidification in the colder parts of the process

equipment. The combination of adequate gas pretreating and heavyhydrocarbon separation assures virtually trouble-free operation from

freezeups. Units have been on stream without need for shutdown be- cause of freeze-plugs for as long as one year's continuous operation.

After separation of heavy-hydrocarbons in separator, 2, the vapor

phase is directed to additional heat exchangers and cooled further to about —250 F. This gas stream is then throttled into a separator, 3, maintained at a pressure of 100 to 225 psi, depending on optimum proc- essing conditions at individual plants. The gas phase from this separator

is cooled and rectified with hquid nitrogen refrigeration in the rectifier,
BUREAU OF MINES HELIUM PLANT PROCESS 19

Figure 1.11. Helium purification units at Keyes Helium Plant. 4, before being removed as a crude helium product consisting of about

60-70 percent helium, 30-40 percent nitrogen, small amounts of methane

and hydrogen, and a trace of neon. The separator pressure is maintained at the lowest pressure consistent with good helium separation

efficiency and adequate to accommodate return of the processed gas through heat exchangers and redelivery to a pipeline, or other disposition.

The natural gas liquids from the gas separator, 3, and rectifier, 4, are

throttled to the return path of the main gas heat exchangers for counter- current heat exchange with inlet gas. The crude helium product is also returned, at about 100-200 psi, through the main gas heat exchangers in counter-current heat exchange with inlet gas. Helium Purification: Purification of crude helium is accomplished

by additional low-temperature processing steps. Prior to purification,

the hydrogen content of the crude helium stream is reduced to about 0.001

percent by catalytic oxidation. This step is accomplished at 50-175 psi
20 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

and ambient temperature by the addition of a small amount of air to the

crude helium stream just ahead of the hydrogen removal equipment.

Addition of air is controlled automatically to provide consumption of all oxygen and to leave a residual hydrogen content in the crude helium of about 0.001 percent. The provision for all oxygen to be consumed is to avoid the possibility of its accumulation in the low-temperature system

where it might inadvertently become mixed with Jiydrocarbons to create

a potentially hazardous mixture. After the hydrogen content is reduced,

the crude helium is dried with activated alumina and compressed to about 2700 psi. Lubricating oil, unavoidably introduced by oil-lubricated

compressors, is removed in a charcoal filter at ambient temperature and

the stream is passed again through alumina-filled driers. Clean, dry,

high-pressure crude helium is then fed to the purification units. These

are shown in figure 1.11.

Recovered helium

to compressor Crude helium

2,500 psig

Nitrogen makeup

Nitrogen 3 psig

Heat

Exchanger

Figure 1.12. Typical cycle for crude helium purification to Grade-A helium in Bureau

of Mines Helium Plant. Figure 1.12 shows the helium purification cycle schematically.

Crude helium at 2700 psi and ambient temperature enters main heat ex- changer, 1, and is cooled by counter-current heat exchange with three

effluent streams. Sources of these streams are described later. The

inlet crude hehum stream is cooled to about —320 F, and its temperature is stabihzed at this temperature by a Hquid nitrogen bath in heat
BUREAU OF MINES HELIUM PLANT PROCESS 21

exchanger, 2. The Hquids formed are separated in separator, 3. The

vapor phase has a composition of 98.2 percent heHum, 1.8 percent

nitrogen, about 0.004 percent hydrogen, and 5-22 ppm of neon. This

stream is cooled to about —340 F by a hquid nitrogen bath, 4, maintained at 3 psia (pounds-force per square inch absolute). Another phase

separation is made, and the vapor phase has a composition of 99.5 per- cent heUum, 0.5 percent nitrogen, 0.001 to 0.004 percent hydrogen, and a

trace of neon. Condensed liquids in separator, 3, and separator, 4, are

principally nitrogen, containing about 3.7 percent of dissolved hehum.

The hquid streams from these two separators are throttled to about

50 psi in separator, 5, where virtually all dissolved helium is released.

The vapor phase from this separator is composed of about 37 percent

helium and 63 percent nitrogen and is returned to main heat exchanger,

1, where it is warmed by counter-current heat exchange with inlet highpressure crude helium. This helium-rich stream is then directed to the

suction of the crude helium compressors for recycling and recovery. Liquid nitrogen from separator, 5, is also returned to the main heat

exchanger where it is evaporated and warmed to ambient temperature

by heat exchange with the inlet high-pressure crude helium. This stream

is about 99.9 percent nitrogen and 0.1 percent hehum and is the source

of makeup nitrogen for the nitrogen refrigeration system.

Final purification of the 99.5 percent helium outlet stream from

separator, 4, is accomplished by adsorption of the remaining impurities

in one of two sets of charcoal-filled adsorbers, 6. These adsorbers are maintained at —320 F by a bath of hquid nitrogen, and the hehum passed

through them reaches a purity of at least 99.995 mole percent. Hydrogen

remaining after the catalytic hydrogen-removal step described previously

serves the purpose of signaling saturation of the charcoal used in final purification. Hydrogen is the first detectable impurity to appear in the

effluent stream from the adsorbers. Detection of the first trace of hydrogen in the parts-per-million range by sensitive monitoring instruments

signals the need to switch to regenerated charcoal in other adsorbers.

After passage through the charcoal, the pure helium stream is directed

to the main heat exchanger, warmed to ambient temperature, and sent

directly to shipping containers or to diaphragm-type compressors when

higher shipping container pressures are required.

Nitrogen Refrigeration Cycle: Nitrogen refrigeration required for both the crude helium separation units and the helium purification units

is supphed by a Claude-type refrigeration cycle. Separate nitrogen

refrigeration equipment is provided for the crude helium extraction and
22 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

the crude helium purification cycles. These are shown in figures 1.10 and 1.12. The nitrogen compressors are common to both cycles, and

the return low-pressure nitrogen streams from both cycles go to a com- mon suction to the compressors.

In the nitrogen refrigeration cycle, nitrogen is compressed to about

600 psi, passed through a soHd desiccant drier, and then the stream is divided, with part of it going to the crude helium separation cycle and

part of it going to the helium purification cycle. Each part enters a heat

exchanger, 7, where it is cooled to about —150 F by nitrogen streams

leaving the cycles. Each stream is divided again with about 80 percent

diverted to expansion engines or turbo expanders, 8. The remaining 20

percent goes to heat exchanger 9, where it is cooled to about — 300 F. Liquid nitrogen is produced when this portion of the nitrogen stream is throttled from 600 psi to the low pressure of about 15 psi in the nitrogen

accumulator, 10, (fig. 1.10) and the nitrogen baths around vessels, 2, 4, and 6 (fig. 1.12). Liquid nitrogen from accumulator, 10, flows by gravity

head into the condensing section of crude helium rectifier, 4 (fig. 1.10). Flash vapors combined with nitrogen vapors from rectifier, 4, in the crude

helium separation cycle and those produced from the various hquid

nitrogen baths in the purification cycle join the respective expansion

engine exhausts, return through nitrogen heat exchangers, 9 and 7, and

are warmed to ambient temperature by counter-current heat exchange

with the high-pressure inlet nitrogen.

1.6.2. Helium Quality and Quality Control

Many uses of helium require that it be of very high purity. Careful attention is given the production of helium and to the controls used to assure its quality at all of the Bureau of Mines helium plants.

Composition of Produced Helium: All heHum produced at Bureau

of Mines plants is 99.995 or higher mole percent hehum, and is designated Grade-A hehum. The total impurity is held to less than 50 ppm

and is usually in the range of 20-25 ppm. The greater portion of this

is neon, which is not completely removed in the purification process.

It is present to the extent of about 15 ppm or less and is not considered

objectionable in most helium uses. Production Monitoring Methods: To assure that only Grade-A

quahty hehum is produced, continuously recording instruments for checking on the presence of impurities are used. Specifically, these are supersensitive calorimeter-type instruments for detecting hydrogen in
HELIUM QUALITY AND QUALITY CONTROL 23

the range of 0-10 ppm and supersensitive thermal conductivity recorders

capable of detecting 5 ppm nitrogen and measuring nitrogen in the range

of 0-500 ppm. Chromatographic analyses of the pure helium stream

with automatically programed chromatographic analyzers are also used

and this use is increasing. These are capable of detecting and measuring

neon, argon-oxygen, nitrogen, and methane impurities in helium when

present to the extent of 5 ppm or less. The usual total impurity content

of hehum from the plant purification process is 20-25 ppm.

A check for water is made with a continuously recording electrolytic moisture analyzer [19]. Methods available for measuring very low

water contents are not too exact, particularly near the zero ppm level; but tests made on hehum from the helium purification process with the

electrolytic moisture analyzer have consistently shown values of less than 1 ppm. This is to be expected, as 1 ppm is equivalent to a water

dewpoint (ice point) of —105 F in a gas at atmospheric pressure, or about —55 F at 2700 psi. In the purification process, the helium emerges

from charcoal adsorbers operated at a temperature of —320 F or lower

while it is at a pressure of 2700 psi. Water vapor content at these conditions is less than any presently known method of testing can detect. Methods Used to Prevent Contamination After Purification and to Assure Purity of Helium Shipped: Although hehum from the purification

process has a very low impurity content, careful handhng is necessary

to avoid introducing impurities during shipment, storage, and use. To assist in controlling the quality of helium shipped, all shipping

containers are required to be returned with a residual charge of Grade-A

helium of not less than 15 psi pressure as a precaution against their

becoming contaminated with air, with other gases, or with moisture.

Railway tank cars and highway tube trailers returned for filling are checked before filling by analyzing the contained residual hehum.

Supersensitive thermal conductivity instruments are used to determine

the total impurities in the residual helium. Supersensitive gas analyzers

check for traces of hydrogen and water. In addition, a sensitive chromatographic gas analyzer is used to doublecheck the presence or absence of gases other than hehum.

If the results of the above checks are satisfactory, the container

is filled with pure helium. If tests of residual helium in a returned tank

car or trailer container show contamination, or if it is returned with

valves open, the container is evacuated and purged with Grade-A helium

until tests on the purging helium show that it will fill with Grade-A quality

helium.

306-922 0-68—

3
24 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

In the case of small cylinders returned for filling, it is not practical

to check each cyUnder before it is filled because of the large number of cylinders filled each day. In filling cylinders, they are placed on a fiUing

rack, the residual helium contents are blown down — and recovered if tests show that it is 90 percent hehum or better — and the cyHnders are then evacuated to an absolute pressure of 1.25 inches of mercury. Without removing them from the rack, they are purged three times by fiUing

them to 50 psi and then blown down to atmospheric pressure. They are then filled to about 300 psi pressure and the hehum in the rack of cylinders is tested. To assure obtaining a representative sample, about

25 pounds pressure is discharged from the rack and the discharging

hehum is tested for impurities. If the tests show Grade-A quality hehum,

the cyHnders are then filled to their rated shipping pressure of 1800-

2600 psi. All purge hehum is recovered.

To check further on the purity of hehum shipped, a random sampling of about 10 percent of the cylinders filled is made, and their contents are analyzed on the aforementioned sensitive instruments. Statistical analysis applied to the results in 1966 showed that more than 99.7

percent of 208,485 cylinders shipped averaged less than 50 ppm total impurities and 91.1 percent averaged less than 25 ppm.

Helium Trace-Impurity Analysis Method: A further routine check

of helium shipped in small cylinders is made by analyzing each month

about 40 of these cylinders by a more precise and specific method

developed by the Bureau of Mines [20] . In this method, all of the impurities contained in a sample of 1 or more liters of the helium to be

tested are solidified and recovered in a trap maintained at liquid helium

temperature. The impurities thus isolated and concentrated are ana- lyzed with a mass spectrometer. From the analysis and a knowledge

of the degree of concentration of the impurities, it is possible to compute

impurities in the original sample in the parts-per-million range. The

data shown in table 1.2 were obtained by this method. They show the

average impurities found in the 402 cyhnders analyzed under this program in 1966.

Table 1.2. Average impurities in Grade-A helium, ppm

Number

cylinders

analyzed

H2O H2 Ne N2 O2 A CO2 Total ppm

402 2.4 0.1 14.2 2.1 0.4 0.05 0.05 19.3
HELIUM HANDLING AND STORAGE 25

The moisture content of the shipped helium depends on the internal

condition of the shipping container with respect to moisture. As the

helium put into the container is completely dry, any moisture in the container, even though it may be only adsorbed moisture on the container's

internal surface, will be picked up by the helium until the partial pressure

of water vapor in the helium comes into equilibrium with the vapor pres- sure of the adsorbed water on the container wall surfaces. Analysis for moisture in the helium is made with an electrolytic moisture analyzer

[19] mentioned earlier. Quality Control at Point of Use: In transferring helium from shipping

containers or from storage, adequate precautions against its contamination should be taken. All connecting pipelines should be purged with

hehum. This should be done at low pressure to minimize wastage of

helium and to achieve a high lineal velocity of gas flow to aid in blowing

dust or scale from the Une. Lines should not be left open to the at- mosphere after purging. Atmospheric air always contains water vapor, and the dry internal surface of pipe that has been in Grade-A helium

service will adsorb water vapor if exposed to air and will release it to the dry helium when the line is returned to helium service.

Care should be taken to avoid backflow of gas from systems to which

helium containers are connected because it could cause contamination of

the helium and containers. Obviously, where the standards of helium

purity are high, it takes only a very small amount of another gas to exceed the allowable impurity level.

1.6.3. Helium Handling and Storage

It is not difficult to handle Grade-A helium without contamination

or loss if careful attention is given to procedures and equipment.

Helium Piping, Valves, and Compressors: Despite helium's leakage

rate, which is higher than most other gases, through minute cracks or openings, it may be shipped, stored, and handled by distribution lines with negligible loss, provided good equipment and materials are used

and reasonable care is taken. Maximum use of welded, brazed, or sol- dered connections, choice of good quality valves, and systematic and

frequent tests for leakage are required. In the past, packless valves

have been used on many shipping containers; but with the development

of more suitable valve stem packing materials and valve designs, packed

stem valves are coming into greater use because they offer advantages

of easier operation and lower maintenance costs. Valves having a soft seat serve best where tight close-off" is needed. Generally, they are not
26 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

as serviceable, however, as valves having metal-to-metal plug and seat where throttling from a higher pressure to a lower pressure is involved.

Conventional compressor equipment can be used if choice is made

of first class equipment and proper consideration is given to packings

and other areas of potential leaks. Because of the necessity of avoiding

contamination of high-purity helium, special attention to this matter is required. Oil-lubricated compressors may be used providing adequate

cleanup equipment is used after compression. Some users of helium have

experienced high-temperature breakdown of compressor oil, which

results in hydrocarbon gas contamination of the helium. Such contamination may not be easily and effectively removed by the cleanup equipment; therefore, careful attention should be given this matter. Compressors are available that use Teflon or Kel-F piston rings which need no

lubrication. Compressors of the diaphragm-type are used by the Bureau

of Mines to compress pure helium. This method of compressing Grade-A

hehum has been completely satisfactory. Careful maintenance attention

is given and effective warning and safety devices are provided to pre- clude carryover of oil, should there be any rupture of the diaphragms

used in the compressors.

Fixed Storage: Storage vessels of seamless pipe, with welded heads

and welded connections, and pressure vessels of conventionally welded

construction are in common use for heHum storage. Most helium storage

pressures are in the range of 2000 to 6000 psi. Higher pressures have

been used or contemplated at some points of helium distribution or use.

1.6.4. Helium Transportation

Large investment costs for shipping containers and high transportation costs have prompted strong consideration of other methods or improved equipment for hehum shipments as possible means of reducing

costs. The three most promising methods are (1) shipment in compressed

gas containers of greater capacity, (2) shipment of heHum as a Hquid,

and (3) transportation of heHum by pipeHne. Some significant developments have been made with regard to ah three methods and more may

be expected in the future. Three types of containers — railway tank cars, highway tube trailers,

and smafl cylinders — are used for shipping compressed gaseous heHum.

These are shown in figure 1.13. In 1966, about 80 percent of the heHum

shipped from Bureau of Mines heHum production plants was in railway

tank cars. In some instances, the cars are shipped to points where the

heHum is distributed to users. At these points, heHum is transferred
HELIUM TRANSPORTATION 27

Figure 1.13. Equipment used in transporting gaseous helium — railway tank car, highway

tube trailer, and standard compressed gas cylinders.

from the railway cars to tube trailers or cylinders for delivery to indi- vidual customers at points where it is used, or is placed in storage

vessels for later use as needed. Shipment of helium from production

plants in tube trailers has some economic advantages when the shipping

points are within about 700 miles of the helium production plants, and

trailers are commonly used for that service. General use of trailers is growing because they offer advantages of mobility and other convenience at point of use. Small cylinders are used to make deliveries of helium where the quantities are small or where the use of small cyl-
28 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

inders is of advantage to the customer. Most small cylinder shipments

are made by motor freight.

Railway Tank Cars: Railway tank cars used for shipping compressed

gaseous helium range in capacity from 240,000 to 345,000 cubic feet and

have service pressures ranging from 2200 to 4000 psi. In the late 1950's

and early 1960's, the capacity of railway tank cars for gaseous helium was

increased about 20 to 30 percent by improved design and higher allowable

filling pressures. A further increase in capacity of up to 30 percent has

been predicted by use of a single large cylinder of welded multilayer

construction. In 1966, the Bureau had 233 railway tank cars in service. Their cost

at that time was about $100,000 each. A few are leased to large commercial purchasers of Bureau-produced helium. They use them to ship

helium from Bureau of Mines plants to distribution points where the

helium is unloaded into highway tube trailers or small compressed gas

cylinders for delivery to their customers, or is put into fixed storage

for later sale. Highway Tube Trailers: About 12 percent of the helium sold by the

Bureau of Mines in 1966 was delivered in highway tube trailers. The

Bureau owned seven trailers, ranging in capacity from 98,000 to 137,000

cubic feet each. The service pressure rating of these trailers is 2640 to 2810 psi. Many helium users also own and use helium tube trailers. The

higher capacity trailers became available in the early 1960's as a result of improved design and use of higher strength materials, which allowed

higher filling pressures, and of higher permissible weight loads on

highways.

Small Compressed Gas Cylinders: In 1966, the Bureau of Mines had

approximately 103,000 small standard compressed gas cylinders in Grade-A helium service. A large number of additional cylinders, owned

by various helium users and gas distributing companies, were in use for transporting hehum. The service pressure rating of the small cylinders

ranges from 1800 to 2640 psi. Somewhat more than 8 percent of the he- hum sold by the Bureau in 1965 was shipped in small cylinders.

Liquid Helium Transportation: The Bureau of Mines has not shipped

any hehum as a hquid, but some commercial companies hquefy hehum

and ship it in that form. The new hehum plant of the Kansas Refined

Helium Company at Otis, Kans., is equipped with a helium liquefier

large enough to liquefy the entire product of the plant, about 450,000

cubic feet per day. Most of the helium produced at this plant is shipped

as a hquid. The Alamo Chemical Company plant near Elkhart, Kans.,
COST OF HELIUM SHIPMENT 29

also has a liquefier large enough to liquefy all the helium it produces,

and most of the hehum will be shipped as a Hquid. Further information on liquid helium shipment is given in chapter 5. Helium Transportation by Pipeline: Relatively small and relatively short pipelines of 2-inch size and up to 90 miles in length have

been used by the Bureau of Mines to transport Grade-A helium between some of its helium production faciUties. These pipelines are operated

at 2500 psi pressure. Economic use of pipelines for transporting large

quantities of hehum from points of production to markets would be

greatly dependent on the stability and continuity of such markets served

by the fixed pipeline transportation system, the load factor, and length

of amortization period for the investment. The principal artery of such

pipehne systems would need to serve strategically located terminals

where the helium would be used, or from which further distribution

could be made. The cost of such further distribution would need to be

added to obtain the overall cost of this form of transporting and delivering

helium. Where the aforementioned conditions are favorable, a pipeline

system could offer economic benefits in the transportation and marketing of hehum.

1.6.5. Cost of Helium Shipment

Helium production plants must be situated where a suitable helium

supply is available and cannot, hke plants producing oxygen or nitrogen,

be arbitrarily located close to the point of use. All of the present helium

plants in the United States are situated in the central part of the country,

while the largest hehum users are on the east and west coasts. The

relatively long distances between points of production and markets,

together with some other conditions, make shipping time and costs major considerations in helium distribution. Cost of Gaseous Helium Shipment: Shipment of helium as a compressed gas requires the movement of heavy containers. Railway tank

cars weigh about 240,000 pounds each. The shipping charges are based

on a lading weight of 40,000 pounds. The cars contain an average of 275,000 standard cubic feet of helium. Highway tube semitrailers are transported on a mileage basis without regard to the weight of the trailer. Their helium-carrying capacity ranges from 98,000 to 137,000 cubic feet. Standard compressed gas cylinders weigh about 130 pounds each and

contain 213-242 cubic feet.
30 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

Table 1.3. Typical transportation costs in 1966 for transportation of gaseous helium

{costs per Mcf of helium shipped in tank cars, tube trailers, and small cylinders to se- lected destinations from Bureau of Mines plants)

Destination Tank cars Tube trailers Small

cylinders

Approximate

round-trip miles

#I.O I ijpi.UO tl '^^ z. lO 9 i^f^ i'*. / o oUU

9 "^9 9 f\'\ 30.32

o.^ 7^I o o.xjo OO.O 1 1 7^0

o.oU 9 nnn

Huntsville, Ala 4.27 6.28 21.63 2,200 Los Angeles, Calif 5.94 6.58 25.01 2,250

Knoxville, Tenn 4.74 6.83 48.19 2,500 Cleveland, Ohio 4.77 7.62 35.44 2,650 Sacramento, Calif 6.17 8.39 25.74 3,000 Cape Kennedy, Fla 5.57 8.54 54.90 3,200

Portland, Oreg 7.24 9.95 43.29 3,375 Newark, N.J 6.15 9.90 19.21 3,400 Boston, Mass 6.54 11.34 64.66 4,100

Note: Capacities of shipping containers assumed to be 275 Mcf for tank cars which was average capacity of cars in use; 125 Mcf for trailers which was representative of late-model trailers in use; and 213 cf for small cylinders. Cylinder

quantities are less than truckload. Transportation costs include "use charge" of 11 cents per mile for tank cars and

$15 per day for trailers. (M = Roman numeral 1000.)

The shipping costs differ greatly, depending on the type of shipping

container used, so the shipping costs per unit volume of hehum shipped

are greatly different. Table 1.3 gives typical transportation costs in 1966

for transportation of gaseous hehum in tank cars, tube trailers, and

small cyhnders to selected points from Bureau of Mines hehum pro- duction plants.

Cost of Liquid Helium Shipment: Chapter 5 covers shipment of helium as a hquid. When large volumes of hehum are required at a single

destination or where the hehum can be offloaded at convenient points en route on a particular trip, shipping costs by this method can be lower

than by ^ any other means. Where hehum is shipped as a hquid only to save on transportation costs, the hquefaction cost has to be taken into consideration. Sometimes additional benefits accrue to help justify this type of transportation.

A study made by the Bureau of Mines on the economics of shipping

hehum as a hquid showed that where reasonably large quantities of
HELIUM PRICES, REGULATIONS GOVERNING SALE 31

helium were needed and the use rate at a given location was high,

savings in shipping costs by this method would more than offset the cost

of liquefaction. Confirmation of this is indicated in chapter 5.

1.6.6. Helium Prices and Regulations Governing Sale

Prices paid for helium by the ultimate consumer vary with the

users' geographical locations, quantities purchased, and other conditions. In 1966 the delivered price was in the range of $40 to $90 per

thousand cubic feet. The prices charged by the Bureau of Mines f.o.b.

its production plants from 1955 to 1960, which were based on total cost

of production (including amortization and interest), were $15.50 and

$19.00 per thousand cubic feet to Federal and non-Federal purchasers,

respectively. In 1961, the price to both Federal and non-Federal con- sumers was increased to $35 per thousand cubic feet in order to provide

additional revenue required for long-range financing of the comprehensive hehum conservation program that was initiated at that time.

This price was still in effect in 1966. Regulations governing sale of helium by the Bureau of Mines are issued periodically by the Secretary of the Interior. These regulations cover conditions of sale and establish prices for helium, rental rates for shipping containers, and various service charges.

Conversion Table

1 psi = 6894.8 newton/meter^

1 ft3 = 0.02832 meter3

1 gal = 0.003785 metei^ Tk= (n' + 459.67)/1.8
32 HELIUM RESOURCES, PRODUCTION, AND CONSERVATION

REFERENCES

[1] C. C. Anderson and H. H. Hinson, Helium-bearing natural gases of the United States —

Analyses and analytical methods, Bureau of Mines BuU. 486 (1951), 141 pp.

[2] W. J. Boone, Helium-bearing natural gases of the United States — Analyses and analytical methods. Supplement to Bulletin 486, Bureau of Mines Bull. 576 (1958), 119 pp.

[3] R. D. Munnerlyn and R. D. MiUer, Hehum-bearing natural gases of the United States —

Analyses, second Suppl. to Bull. 486, Bureau of Mines BuU. 617 (1963), 93 pp.

[4] R. D. Miller and G. P. Norrell, Analyses of natural gases of the United States, 1961, Bureau of Mines Information Circ. 8221 (1964), 148 pp.

[5] R. D. Miller and G. P. Norrell, Analyses of natural gases of the United States, 1962, Bureau of Mines Information Circ. 8239 (1964), 121 pp.

[6] R. D. Miller and G. P. Norrell, Analyses of natural gases of the United States, 1963, Bureau of Mines Information Circ. 8241 (1965), 102 pp.

[71 B. J. Moore, R. D. Miller, and R. D. Shrewsbury, Analyses of natural gases of the United States, 1964, Bureau of Mines Information Circ. 8302 (1966), 144 pp.

[81 Canadian helium to lose present market in Europe, Oil Week, Nov. 1, 1965, p. 14.

[9] P. K. Smith and H. S. Pylant, HeUum plant in commercial operation. Oil and Gas J., Oct. 29, 1962, pp. 136-139.

[101 High on HeHum, Chemical Week, May 8, 1965, p. 26.

[11] G. D. Kinney, World's biggest helium plant opens, Oil and Gas J., Sept. 30, 1963,

pp. 54-61.

[12] S. Weller and W. A. Steiner, Engineering aspects of separation of gases — Fractional Permeation Through Membranes, Chem. Eng. Prog. 46, 585 (Nov. 1950).

[13] S. A. Stern, T. F. Sinclair, P. J. Caries, N. P. Vahldieck, and P. H. Mohr, HeUum re- covery by permeation, Ind. Eng. Chem. 57, 49 (Feb. 1965).

[14] D. R. Huffman and R. J. Robinson, Method of separating hehum from other constit- uents of natural gas, U.S. Patent 3,239,996 (Mar. 15, 1966).

[15] E. I. duPont de Nemours & Co., Inc., Private communication (Oct. 31, 1966).

[16] K. B. McAfee, Jr., Bell Labs. Record 38, 354 (1960).

[17] N. Melnyk and H. W. Habgood, Extraction of heUum from natural gas. The diffusion- through-glass process. Can. Mining and Metallurgical Bull. (Oct. 1961), pp. 768-774.

[18] David Frazier, Hehum separation, U.S. Patent 3,184,899 (May 25, 1965).

[19] F. A. Keidel, Determination of water by direct amperometric measurement. Anal. Chem. 31, No. 12, pp. 2043-2048 (Dec. 1959).

[20] C. G. Kirkland, L. W. Brandt, and W. M. Deaton, Determining trace impurities in grade-A hehum. Bureau of Mines Report of Investigations 5644 (1960), 12 pp.
CHAPTER 2

Properties of Helium

D. B. Mann '

Page 2.1. Pressure-volume-temperature surface 35

2.1.1. Origin of data and correlation methods 35

2.1.2. Pressure-density diagram 37

2.1.3. Pressure-temperature diagram 40

2.2. Thermodynamic functions 40

2.2.1. Derived functions 40

2.2.2. Temperature-entropy diagram 43

2.2.3. Enthalpy-entropy diagram 43

2.2.4. Pressure-enthalpy diagram 50

2.2.5. Pressure-internal energy diagram 50

2.2.6. Specific heat-temperature diagram 50

2.3. Transport properties 51

2.3.1. Viscosity 51

2.3.2. Thermal conductivity 51

2.4. Discussion 52

' Cryogenics Division, NBS Institute for Basic Standards, Boulder, Colo. 80302.

33
34 PROPERTIES OF HELIUM

Critical analysis of helium refrigeration and liquefaction processes

requires an extensive knowledge of the variation of the thermophysical

properties. The ideal gas relation, PV= nRT, gives a good approximation

of physical properties at low pressures and ambient temperatures, but

when dealing with processes which extend over broad ranges — as in the case of a cryogenic process, from ambient to near absolute zero — it is necessary to know property values to a high degree of accuracy. Internal

consistency of the data is of equal importance to the engineer charged

with equipment design and evaluation.

It is the purpose of this chapter to supply the necessary thermophysical properties over an entire range of temperatures (3 to 300 K)

and pressures (0.5 to 100 atmospheres*) useful to a critical description

of processes involving liquid hehum and liquid helium temperatures. It should be specifically mentioned that this chapter does not deal with

superfluid properties. Additional engineering properties such as vis- cosity and thermal conductivity are included where broad range correlations exist. These data are most useful in support of design calculations

of chapters 3 and 4.

* Atmosphere (normal = 760 torr)= 1.01325 X 10^ N/m^.
PRESSURE VOLUME TEMPERATURE SURFACE 35

2.1. PRESSURE-VOLUME-TEMPERATURE

SURFACE

2.1.1. Origin of Data and Correlation Methods

The work of Keesom [1]- assembled and in most cases correlated

the experimental property data through 1941. Akin [2], using the

Beattie-Bridgeman equation of state and data derived from Keesom,

presented thermodynamic properties of helium for temperatures ex- tending from 590 K down to about 12 K and to quite high gas densities.

Lounasmaa [3] provided experimental data which indicated that the virial form of the equation of state used by Keesom was somewhat

inadequate in representing the properties for densities greater than

the critical density at temperatures below 20 K. Mann and Stewart [4]

correlated these data of Lounasmaa with those of Keesom [1], Berman

and Mate [5] , and Domb and Dugdale [6] . The results of this correlation were presented in graphical form for the temperature region 3

to 20 K. The work of Edeskuty and Sherman [7] provided experimental

data at 4.2 K and below and up to densities in excess of three times the

critical density.

Correlation of the applicable data of these more recent investi- gators with those of Keesom will give a continuous representation of properties extending from 3 to 300 K and pressures to 100 atm. The following is a list of symbols and definitions used in this chapter:

molar specific heat of helium gas at very low pressures. Assumed

to be numerically equal to 5/2 R J/g-mol-K.

Csat specific heat of hquid helium at saturation conditions = J/g.

mol-K.

p mass concentration (density) = g-mol/liter.

H enthalpy = J/g-mol.

Ho enthalpy value of liquid helium at the conditions T=0 K,

P = zero. Hl enthalpy value at the conditions r= 4.2144 K (normal boiling

point, T^t>e) P= \ atm, saturated liquid.

P absolute pressure = atmospheres.

5 entropy = J/ g-mol-K.

So entropy of liquid helium at the conditions T= 0 K, and P — zero, assumed to be numerically equal to zero.

^ Figures in bracicets indicate the literature references at the end of this chapter.
36 PROPERTIES OF HELIUM

Sl entropy value at the conditions r= 4.2144 K (normal boiling

point, T'^se) , P=l atm, saturated liquid.

T temperature = degrees Kelvin.

U internal energy = J/g-mol.

V molar volume = liters/g-mol.

P

Z compressibility factor —

pKl

The temperature scale is that designated by the symbol Tsse [Clement, Logan, and Gaffney [8]]. Molar volumes are used to correlate

data and then are converted to units of density. Constants and conversion factors are as follows: One atmosphere = 1.013250 X 10^ N/m^

R (gas constant) =8.20575 X 10-'^ liter-atm/g-mol-K

(chemical scale) = 8.31424 J/g-mol-K

One calorie (thermochemical)~4.1840 J One hter (old) = 1.000028 X lO-^m^

The equation of state developed to describe the pressure-densitytemperature surface is based on the equation of Benedict-Webb-Rubbin

[9] (BWR) as modified for nitrogen by Strobridge [10] and further modified here for hehum. With pressure as the dependent variable, the

equation is

+ + ^+ ^) p5e-«..P^/- + n,,p^ + nup'. (2.1)

Values for the constants are R = 0.820575 (10-^

ni= 0.1627693557(10-1)

/i2 = - 0.3700903492 (10-1)

n3= 0.7567259965(10-2)

follows:

^4 = -0.1887566673 (10-1)

^5= 0.4792344640(10-1)

/Z6= 0.1560344984(10-3)

117= 0.9274298542(10-3)
PRESSURE-VOLUME-TEMPERATURE SURFACE 37

n, =-0.2833997045(10-4)

=-0.5002871627(10-2)

^,0- 0.3451283837(10-1)

nu= 0.1008924438(10 «) nv2= 0.65 (10-2)

Aii3 = - 0.1851562535 (10-4)

ni4= 0.4204351888(10-4)

;ii5 = -0.1712445150(10-3)

nie= 0.7739390202(10-6)

rii7 = - 0.2870693948(10-8)

A least squares method developed by Jones and Gallet [11] and

programmed for use on a digital computer was used to determine the

values of these constants. Because the constant rivz appears in the

exponential term, it was necessary to select several values for this constant and solve for the set of constants which gave the minimum

in the sum of squares of deviations. The resulting best fit for the data was only shghtly affected by changes in the value of the constant rivz, but it was found necessary to carry the maximum number of places in the derived constants to assure internal consistency and reproducibility.

Data used for the least squares treatment consisted of 695 pressure,

density, and temperature points. These points were taken from the

following published works.

Berman and Mate [5] : Thirty-eight data points were used to describe the saturated vapor and saturated liquid conditions. Temperatures range from 2.2 to 5.18 K. Edeskuty and Sherman [7] : Eighty-eight data points were used to describe the compressed

Hquid region from 2.2 to 4.2 K. Pressures range from 1 to 140 atm and the highest density

values are in excess of three times the critical density.

Hill and Lounasmaa (HL) [12] : A total of 258 experimental data points was used to describe

the rather broad temperature and pressure region from about 2.5 to 20 K. Pressures range from 1 to 102 atm with density values extending to three times the critical density. Keesom [1]: The virial equation of state and the values for the coefficients adopted by Keesom were used to obtain 311 pressure-density-temperature points along 29 isotherms

extending from 3.1 to 574 K. Pressure-density-temperature points having densities greater than the critical density were excluded because of the divergence of the equation of state values for pressure from the experimental values of (HL) [12] in the region of the 20 K

isotherm. Except for the two-phase region described by Berman and Mate [5] , overlapping of the data regions was extensive.

2.1.2. Pressure-Density Diagram

Figure 2.1 is a plot of pressure versus density for isotherms from 3

to 30 K. The density changes most rapidly in this temperature range. The graph was plotted using eq (2.1).
Figure 2.1. Pressure-density diagram.
PRESSURE-TEMPERATURE DIAGRAM 39

TEMPERATURE, "K

Figure 2.2. Pressure-temperature diagram.

306-922 0-68—

4
40 PROPERTIES OF HELIUM

2.1.3. Pressure-Temperature Diagram

Figure 2.2 is a plot of the data of figure 2.1 on the coordinate sys- tem pressure versus temperature. The pressure and temperature data cover the same ranges as figure 2.1 and with specific volume varying

from 5 cm^/g to 20 cm^/g. The two-phase liquid-vapor boundary is also indicated.

2.2.1. Derived Functions

The property functions of enthalpy, entropy, and internal energy for the single-phase fluid region were calculated using only the equation

for the pressure-density-temperature surface, eq (2.1). This procedure was necessary in the absence of specific heat measurements over the

entire range of interest. Equations used to calculate these properties

are shown below. (For a complete derivation of the equations, the

reader is referred to Beattie [13].)

The values of //° (36.5582 J/g-mol) and S° (9.2609 J/g-mol-K) are the

reference enthalpy and entropy of the ideal gas evaluated at the normal

boihng point temperature (4.2144 K) and 1 atm. The values of enthalpy, entropy, and internal energy of the satu- rated Hquid given in table 2.1 were taken from Mann and Stewart [4].

These values were found by graphically integrating the expressions

2.2. THERMODYNAMIC FUNCTIONS

(2.3)

(2.2)

(2.4)

(2.5)

(2.6)
DERIVED THERMODYNAMIC FUNCTIONS 41

Table 2.1. Properties of helium, two-phase liquid-vapor region

Tem- Saturated vapor Saturated liquid pera- rProc res- ture sure

V A u 5 V h u s K Atm cm ^Ig Jig Jig Jlgm-K cm^lg Jig Jig JIg-K

3.00 0.241 224.11 28.96 23.49 10.247 7.0852 5.30 5.13 2.356

3.20 0.320 174.83 29.48 23.81 9.911 7.1849 5.85 5.62 2.527

3.40 0.416 138.70 29.93 24.08 9.606 7.3025 6.47 6.16 2.690

3.60 0.529 111.48 30.34 24.36 9.322 7.4388 7.18 6.78 2.883

3.80 0.661 90.50 30.59 24.53 9.021 7.5965 7.93 7.40 3.059

4.00 0.814 73.86 30.72 24.63 8.736 7.7785 8.81 8.17 3.255

4.20 0.990 60.61 30.79 24.71 8.443 7.9885 9.85 9.05 3.460

4.40 1.190 49.80 30.65

.

24.65 8.150 8.2305 10.96 9.97 3.674

4.60 1.417 40.72 30.32 24.47 7.845 8.5470 12.33 11.10 3.920

4.80 1.672 32.87 29.49 23.92 7.473 9.0334 13.88 12.35 4.188

5.00 1.959 25.67 28.02 22.92 7.000 9.9404 16.00 14.03 4.598

The specific heat of saturated liquid (Cgat) was taken from Hill and

Lounasmaa (HL) [14]. Equation (2.1) was developed starting, with the work of Lounasmaa

[3]. This pubHcation presented the derived constants for the BWR equation of state using the data obtained experimentally by HL [12]. As the

data range was extended, the BWR equation was also modified as necessary. The temperature-pressure region 4.8 to 5.8 K and 1.8 to 2.8 atm

was less accurately reproduced as this modification proceeded to the

form of (2.1). The reasons for this decrease in accuracy was the inability

of the developed equation to represent exactly the form of the critical isotherm at or near the critical pressure. The effect of this problem was twofold. Properties defined by (2.1),

(2.2), (2.3), and (2.4) in this temperature and pressure region were found

to have substantial errors. In addition, the integration, indicated by (2.2) and (2.3), for the values of derived properties at densities greater than

the critical density and at pressures greater than 2.8 atm resulted in substantial errors at temperatures near the critical point.

To ehminate these errors, the properties were derived using (2.1),

(2.2), (2.3), and (2.4) for temperatures of 6 K and above for densities

greater than critical density and for temperatures approaching the satu- rated vapor conditions at densities less than the critical density.
42 PROPERTIES OF HELIUM

Equations (2.2) and (2.3) were then modified so as to substitute for the lower hmit of integration the values for the properties along the satu- rated Hquid region. Equations (2.7) and (2.8) indicate this modification

and, with (2.4), are the equations used to derive properties at pressures

greater than the critical pressure and at temperatures less than 6 K:

P T (dP\ 1 , , ,

Isobars were then plotted graphically on temperature-entropy and

temperature-enthalpy coordinates. The entropy change along the

isochore 30.93 g-mole/liter was computed from 4.75 to 10 K using the

specific heat data of HL [12] and the expression

r^2 C

AS = y^r, (2.9)

where Cv is the specific heat at constant volume. The intersections of the

isobars with this isochore given by HL [12] were noted as the temperature was increased.

A correction for the properties along the 5 and 6 K isotherms was de- termined and the tabulated data were adjusted to be consistent with the

above-described specific heat data of HL [12]. In addition, the 2-atm

isobar was plotted from HL [12] in the region of the critical values and

the tabulated properties obtained directly from this plot.

Values taken from the pressure-density-temperature surface

described by (2.1) and the derived properties found from (2.2) through

(2.9) are tabulated by Mann [15] as functions of pressure and temperature. Forty isobars are presented covering the range 0.5 to 100 atm. One

degree increments of temperature are given except at the two-phase

boundaries, and, in general, extend from 3 to 300 K. Values fisted in the

tables afe specific, not molar, quantities. The molecular weight of helium was taken as 4.0028.

It was found that a better representation of the pressure-densitytemperature surface could be achieved by considering a number of restricted ranges of temperature and pressure when defining the constants of (2.1), but this procedure would then require the fitting and

dp (2.7)
TEMPERATURE-ENTROPY DIAGRAM 43

smoothing of several property surfaces. It was felt that the surface

matching would not increase the accuracy of the tabulated values over

the entire range of interest by a factor that could justify the extensive

data manipulation required.

The pressure-density-temperature surface described by (2.1) fit the

original data to varying degrees of accuracy depending on the temperature and pressure region considered. In general, the ability of (2.1) to reproduce the original data is less accurate in the regions of phase

change, and in particular, in the region of the critical point. As mentioned

above, the liquid-vapor region was handled graphically in an attempt to reduce the error in this region. In the temperature region 10 to 300 K the accuracy of the values of pressure, density, enthalpy, and entropy is estimated to be within 3 per- cent while internal energy values may vary somewhat more. Below 10 K and at low pressures the properties of enthalpy, entropy,

internal energy, and pressure are accurate to an estimated error of

3 percent, while density may have an uncertainty of as much as 5 percent. At higher pressures the density is defined to within the estimated

3 percent, but the pressures may be in error as much as 5 percent.

Values for the specific heats of helium derived from the graphs are estimated to be within 5 percent in the temperature range 10 to 300 K.

It is not recommended that specific heats be extracted from the graphs

at temperature levels below 10 K.

2.2.2. Temperature-Entropy Diagram

Figures 2.3 and 2.4 are plots of the physical and thermodynamic

properties of hehum on temperature-entropy coordinates. Two charts are used to give clarity in presentation of the data below 25 degrees. The

charts are overlapping in temperature between 25 and 15 K. It will be

noted that the area in the region of the critical has not been defined. This

is caused by the lack of data in this region and the inability of (2.1) to predict properties in this pressure temperature region.

2.2.3. Enthalpy-Entropy Diagram

The enthalpy-entropy diagram of figure 2.5 will aid in the design of equipment operating below 25 K. Where the temperature-entropy dia- gram of figures 2.3 and 2.4 aid in describing thermodynamic cycles, the

design parameters for the critical low-temperature heat exchangers can

best be found using the enthalpy-entropy plot.
44 PROPERTIES OF HELIUM

Figure 2.3. Temperature-entropy diagram, 15-300 K.
TEMPERATURE-ENTROPY DIAGRAM 45

Figure 2.4. Temperature-entropy diagram, 3-25 K.
46 PROPERTIES OF HELIUM
Figure 2.6. Pressure-enthalpy diagram.
Figure 2.7. Pressure-internal energy diagram.
SPECIFIC HEAT-TEMPERATURE DIAGRAM 49

2 4 6 8 10 12 14 16 18 20 TEMPERATURE,»K

Figure 2.8. Specific heat-temperature diagram.
50 PROPERTIES OF HELIUM

2.2.4. Pressure-Enthalpy Diagram

A third cross plot of the physical and thermodynamic properties of hehum is shown in figure 2.6. These data are also derived from the

original equation and are of aid to the engineer and designer of heat

exchanger equipment for hefium service.

2.2.5. Pressure-Internal Energy Diagram

The properties charted in figure 2.7 are especially useful in describing the process of constant volume heat addition. Such processes occur

in storage and transport of liquid hehum where the storage vessel is sealed and a pressure rise caused by heat leak into the vessel is desired

as a function of time. The internal energy function is defined by eq (2.4).

2.2.6. Specific Heat-Temperature Diagram

This diagram (fig. 2.8) describes the specific heat function as it relates to pressure and temperature changes near the critical point.

These data are taken directly from HL [12], as the state equation is not capable of showing specific heat variations in the region near the critical point.

Table 2.2. Dynamic viscosity of gaseous helium [1] T fjL (micropoises) T fx (micropoises)

300 201.2 100 98.8 280 192.4 80 85.6 260 183.4 70 78.5 240 174.2 60 71.0 220 164.6 50 63.1

200 154.8 40 54.6 180 144.6 30 45.4 160 134.0 20 34.9 140 122.9 10 22.3 120 111.2 5

14.2
TRANSPORT PROPERTIES 51

2.3. TRANSPORT PROPERTIES

2.3.1. Viscosity

It can be shown [1] that viscosity is a very weak function of pressure

at temperatures near room temperature. As the temperature is lowered

to near the two-phase hquid-vapor region, viscosity becomes a strong

function of pressure. At the present time, no correlation of a continuous

nature has been performed on viscosity of gaseous or Hquid hehum. The

engineer and designer are therefore forced to use the functional rela- tionship derived by Keesom [1] for viscosities of heHum gas at low

pressure. This relationship is

/x = 5.023P-647. (2.10)

Viscosity (ix) is given in micropoises. Selected values of viscosity are

given in table 2.2.

2.3.2. Thermal Conductivity

The situation in regard to data on thermal conductivity of helium gas

is similar to that of viscosity. Thermal Conductivity is also a weak function of pressure near ambient temperature conditions. As the temperature is lowered to the region near the two-phase liquid-vapor boundary,

thermal conductivity becomes a strong function of pressure. Again cor- relations of a continuous nature are not available. The engineer is again

forced to rely on the data of Keesom [1] for thermal conductivity. These

data are presented in table 2.3.

Table 2.3. Thermal conductivity of gaseous helium [1]

T(K) k (WaUs/cm-K) X 10^

273 1.44 260 1.39 220 1.24 180 1.08 140 0.90 100 0.70 60 0.48 20 0.21
52 PROPERTIES OF HELIUM

2.4. DISCUSSION

The compilation and correlation of the thermophysical properties of

helium presented in this chapter are only one in a series of such presentations begun by Keesom, Bijl, and Monte [16]. As investigative and

measurement techniques improve, so do the data necessary to revise

the correlations. Zelmanov [17] published data w^hich affected properties

correlations below 20 K. This information was the basis for a revision to the original Keesom, Bijl, and Monte [16] data and was published by

Scott [18]. New experimental measurements by Lounasmaa [3] for the

temperature region 3 to 20 K and the correlations of this information by

Mann and Stewart [4] pointed up the need for a revision of the properties

charts over the entire temperature range of 300 K. The widespread use of the digital computer has aided significantly

in the task of correlating experimental data. Numerical methods, pre- viously found impractical, could be applied to the development of a correlating function or equation of state. McCarty and Stewart [19] applied

these techniques to the temperature region 20 to 300 K. Mann [15] using

the work of Strobridge [10] developed the modified BWR equation (with

18 constants) that was used in this chapter. As indicated previously, this state correlating function met the requirements of the process engineer

interested in equipment and process design over the temperature range

3 to 300 degrees. This equation does have a number of disadvantages.

To name two, it is impractical to use without the aid of the digital computer and it becomes quite inaccurate for helium near its critical point.

Additional experimental data and correlations will certainly follow to improve the precision and accuracy of helium property data. Glassford

and Smith [20] have already extended the experimental data to 1300 atm

below 20 deg. It is the opinion of this author that the experimentalist

should not neglect the region near the critical point of helium. Lounasmaa

[3] did provide some experimental data in this extremely difficult region,

but more data are necessary. It is also suggested that a new approach to a

correlating function should be made; one that takes into account quantum

effects of helium at these low temperatures.

With additional development of high-precision measurement tech- niques as well as the progress in computer technology, additional ad- vances in correlating methods will lead to even more accurate and

precise property representations.
REFERENCES 53 REFERENCES

[1] W. H. Keesom. Helium (Elsevier Publ. Co., Amsterdam. The Netherlands, 1942).

[2] S. W. Akin. The thermodynamic properties of helium. ASME Trans. 72, 751 (1950).

[3] 0. V. Lounasmaa. Specific heats at low temperatures. Thesis submitted for degree of Doctor of Philosophy. University of Oxford, Great Britain (1958).

[4] D. B. Mann and R. B. Stewart, Thermodynamic properties of helium at low temperatures and high pressures, NBS Tech. Note 8 (1959).

[5] R. Berman and C. F. Mate. Some thermal properties of helium and their relation to the temperature scale. Phil. Mag. 3A, 461 (1958).

[6] C. Domb and J. S. Dugdale. Solid helium. Progress in Low Temperature Physics

(ed. C. J. Gorter. North Holland Publ. Co.. Amsterdam. The Netherlands, 1957), Vol. II, p. 338.

[7] F. J. Edeskuty and R. H. Sherman. PVT relations of hquid He^ and He^, J '-w Temperature Physics and Chemistry (ed. J. R. DiUinger, Univ. of Wisconsin Press, Madison. Wise, 1958). p. 102.

[8] J. R. Clement, J. K. Logan, and J. Gaffney. Liquid hehum vapor and pressure-temperature scale, Phys. Rev. 100, 743 (1955).

[9] M. Benedict. G. B. \^ ebb. and L. C. Rubin. An empirical equation for the thermodynamic properties of hght hydrocarbons and their mixtures. J. Chem. Phys. 8, 334

(1940).

[10] T. R. Strobridge. The thermodynamic properties of nitrogen from 64 to 300 K between

0.1 and 200 atmospheres. NBS Tech. Note 129 (1962).

[11] W. B. Jones and R. M. Gallet. The representation of diurnal and geographic variations

of ionospheric data by numerical methods, Telecomm. J. No. 5, 29 (1962).

[12] R. W. Hill and O. V. Lounasmaa. The thermodynamic properties of fluid hehum. Phil. Trans. Roy. Soc. (London) 252A, 357 (1960).

[13] J. A. Beattie. Thermodynamic properties of real gases and mixtures of real gases. Thermodynamics and Physics of Matter (ed. F. D. Rossini. Princeton Univ. Press,

Princeton. N.J.. 1955), p. 240.

[14] R. W. Hill and O. V. Lounasmaa. The specific heat of liquid helium, Phil. Mag. 2A,

143 (1957).

[15] D. B. Mann. The thermodynamic properties of hehum from 3 to 300 K and from 0.5 atmospheres to 100 atmospheres. NBS Tech. Note 154 (1962).

[16] W. H. Keesom. A. Bijl. and L. A. J. Monte. Le diagramme log T, S de Thehum, xA.ppl. Sci. Res. A4, 25 (1953).

[17] J. Zelmanov. The entropy diagram for hehum at low temperatures, J. Phys. (USSR)

8, (1944).

[18] Russell B. Scott. Cryogenic Engineering (D. Van Nostrand Company, Inc., Princeton,

N.J., 1959).

[19] R. D. McCarty and R. B. Stewart, An equation of state for calculating the thermody- namic properties of hehum at low temperatures. Progress in International Research

on Thermodynamic and Transport Properties (ASME, New York, N.Y., 1962), p. 107.

[20] A. P. M. Glassford and J. L. Smith. Jr.. Pressure-volume-temperature and internal energy data for helium from 4.2 to 20 K between 100 and 1300 atmospheres, Proc.

International Institute of Refrigeration, Commission 1, Boulder, Colo. (1966), p. 343.
i!

i
CHAPTER 3

Liquefaction Techniques

Samuel C. Collins ^

Page 3.1. Principles of refrigeration 56

3.1.1. Joule-Thomson effect 56

3.1.2. Liquid hydrogen cooling 58

3.1.3. External work machines 59

3.1.4. The Stirhng cycle 59

3.1.5. Simon expansion liquefier 61

3.2. Common helium Uquefiers 62

3.2.1. Helium hquefaction cycle 62

3.2.2. Management of the heat exchanger 63

3.2.3. The Joule-Thomson heat exchanger 64

3.2.4. Expansion machines 65

3.3. HeHum liquefaction facihties 66

3.3.1. Otis, Kans., helium Uquefier (Ch. Trepp, Sulzer Brothers

Ltd., Winterthur, Switzerland) 66

3.3.2. 500 Incorporated (S. C. Collins and Arthur Post) 69

3.3.3. Union Carbide, Linde Division (A. J. Westbrock) 70

3.3.4. Air Products and Chemicals, Inc. (W. A. Snow, A. R. Winters, R. B. Currie) 74

3.3.5. CVI Corporation, Helium add-on liquefier (C. B. Hood)... 76

' Village Hill Road. Belmont, Mass. 02178.

306-922 0-68—

5

55
3.1. PRINCIPLES OF REFRIGERATION

3.1.1. Joule-Thomson Effect

The specific heat of real gases is a function of the pressure. Over a wide range of temperature and pressure, the specific heat increases with

the pressure. The Joule-Thomson coefficient, (~) , will, therefore,

increase in magnitude as the temperature falls, passing from negative

values at higher temperatures to zero at the inversion temperature and,

finally, to positive values at lower temperatures. This property is illus- trated in the enthalpy-temperature diagram of figure 3.1. Isothermal compression, state 5 to state 1, is attended by a substantial decrease in specific enthalpy. If the compressed gas undergoes

isenthalpic expansion, 1 to 1', there is a significant drop in temperature,

a positive Joule-Thomson effect. If the compressed gas is led through one

channel of a counterflow heat exchanger before expansion and the colder

expanded gas is returned through a second channel of the heat ex- changer, the colder gas will absorb heat from the incoming high-pressure

stream, with the result that the latter will arrive at the expansion valve

Editors' Note: The contributions to helium hquefaction and refrigeration made by

Samuel C. Collins have resuhed in widespread impact on technology. In 1958, on the

fiftieth anniversary of the first hquefaction of hehum, he was awarded the Kamerhngh- Onnes Gold Medal for his development of the Colhns Hehum Cryostat. In 1951 he was

awarded the John Price Wetherill Medal of The Frankhn Institute and in 1965 was the I first recipient of the highest award given by the Cryogenic Engineering Conference,

The Samuel C. CoUins Award.

This chapter was prepared by Prof. Colhns, at the request of the editors, to present

his views on Hquefaction techniques for hquefaction of hehum. Our decision to separate

liquefiers and refrigerators into two chapters is not unique, since all hquefiers are refrig- erators but all refrigerators are not necessarily hquefiers. Recent industrial practice,

however, has prompted dividing the subject into two parts. 56
JOULE-THOMSON EFFECT 57

at a progressively lower temperature until the liquid phase appears. Referring to the diagram of figure 3.1, the state of the high-pressure gas changes from 1 to 2 in the heat exchanger. After expansion at 2, there

are two phases: liquid at 3, and saturated vapor at 4. The low-pressure

stream is warmed from 4 to 5 by heat received from the high-pressure

stream during its passage through the heat exchanger. It has been

assumed that the heat exchanger is well insulated and highly efficient. Otherwise, the temperature at 5 would be appreciably lower than that

at 1.

TEMPERATURE —

Figure 3.1. Joule-Thomson refrigeration process.

Assuming no leakage of heat from the surroundings and no work

done by the gas, the fraction, e, of the stream which remains as hquid

is given by the relation:

^5 ~ hi € = T 7- (3.1)

fi^ — 1I3

If there is no accumulation of liquid, the maximum refrigerative effect,
58 LIQUEFACTION TECHNIQUES

Q, is expressed by the equation:

Q = h-h,. (3.2) Note that the amount of refrigeration available from this simple system

is independent of the temperature level at which it is applied so long as that level is chosen to lie between the boiling point of the liquid phase

at the lower pressure and the temperature of the warm end of the heat

exchanger. The amount of refrigeration is greatly affected by the temperature level at the warm end of the heat exchanger or, more specifically, the temperature of the compressed gas entering the warm end

of the heat exchanger. The lower this temperature, the greater will be

the magnitude of ^ in eq (3.2) and of e in eq (3.1). The usefulness of the Joule-Thomson effect as a means of liquefying gases depends upon the possibility of starting with compressed gas

at a temperature well below the inversion point. In the case of helium,

this implies a temperature well below 30 K. Actually, for useful yields

of liquid, temperatures in the range of 8 to 15 K are necessary. There

are two accepted methods of precooling hehum to the desired level; one by the use of liquid hydrogen which goes down to 14 K under reduced

pressure, and the other by expanding a part of the helium stream in an external-work machine. More elaborate counterflow heat exchangers

are required for the appHcation of these coolants.

3.1.2. Liquid Hydrogen Cooling

A hquid hydrogen bath is placed in the high-pressure flow stream

just prior to the inlet of the Joule-Thomson heat exchanger. The temperature of this bath is lowered by pumping to a pressure near the triple

point of the Hquid hydrogen (14 K). Provision must be made for recovering the refrigerative value of the effluent hehum, which leaves the warm

end of the Joule-Thomson heat exchanger at about 15 K on its way to the conipressor operating at room temperature. This can be done easily

in a second heat exchanger for coohng the high-pressure stream by

transferring its heat to the outgoing gas. Partial use of the hydrogen

vapor can be made.

Because of the fire and explosion hazard introduced by the pres- ence of hquid hydrogen, its use in the Hquefaction of hehum is lagging.
EXTERNAL WORK MACHINES 59

3.1.3. External Work Machines

When a compressed gas expands adiabatically in an engine, the

temperature decreases chiefly because of the decrease of internal energy. The work is expressed by the relation: J^=RToj^ [1-/"^^], (3.3)

where To is the temperature of the gas entering the engine, r is the

ratio of final to initial pressure, and A- is the ratio of specific heats, —. The refrigeration generated by the expansion resides in the cold

gas discharged by the engine. It is usefully applied by allowing the cold

gas to absorb heat from the material which is to be cooled; in this case,

a stream of helium. Whereas the work done by the engine equals the

decrease in the enthalpy of the gas, the available refrigeration is the

product — CpAr where Cp is the specific heat at constant pressure and

AT is the change in temperature of the gas as a result of the expansion.

The work is exactly equivalent to the available refrigeration only at the

inversion temperature for the pressure range involved. Note that the

work done is directly proportional to the temperature of the gas entering

the engine. It is desirable, therefore, to perform the expansion at the

highest temperature at which the resultant refrigeration can be used.

A cold stream of expanded hehum is preferred to an evaporating

liquid for coofing a stream of compressed hehum. Assuming the transfer of heat to take place in a counterflow heat exchanger, the loss of entropy by the stream being cooled is more nearly equal to the gain of entropy by the fluid being heated and the process is, therefore, more

nearly reversible.

3.1.4. The StirHng Cycle

A special variety of the external work machine which is sometimes used to cool helium to the 15 K level or lower for partial liquefaction in a Joule-Thomson process is the reversed heat engine of

Stirling. In this machine the refrigerative effect is produced by the adiabatic expansion of compressed helium just as in the ordinary expansion
60 LIQUEFACTION TECHNIQUES

engine. Instead of the counterflow heat exchanger with its pair of channels for high- and low-pressure streams respectively, however, the

Stirling system employs a regenerator, a cylindrical vessel filled with

porous material of substantial heat capacity and extensive surface. Compressed helium flows momentarily through the regenerator into the

expansion chamber. Expansion occurs and, thereafter, the cold ex- panded hehum returns through the regenerator. A thermal gradient is

(a)

da

. Compression Chamber

Expansion KChamber

(b) (c) (d) Figure 3.2. The Stirling cycle.

soon established in the regenerator with the result that each new charge

of compressed helium arrives at the expansion chamber with a rela- tively low temperature. After expansion, the temperature is lower still. The principle of operation is illustrated by the four sketches of figure

3.2. At a, the charge of hehum has been compressed but it still resides

in the compression chamber — the displacer D being in its lowest position. At 6, the displacer has been moved upward, thereby forcing the

hehum to flow through the regenerator R into the expansion chamber.

At c, both piston P and displacer have moved upward to create a larger

volume in the expansion chamber. The pressure and temperature of the

helium fall to the minimum level. At d, the expanded gas has been

forced to return to the compression chamber by way of the regenerator.

Provision is made for the absorption of heat by the cold helium along the

path between the expansion chamber and the regenerator.

The advantages of the Stirling cycle over the conventional expansion engine-counterflow heat exchanger assembly are the relative simphcity of both the regenerator and displacement apparatus, the absence

of valves, and a greater tolerance for impurities. The disadvantages are
SIMON EXPANSION LIQUEFIER 61

the large clearance volume presented by the voids in the packing of the

regenerator and the difficulty of providing adequate heat capacity,

since the specific heat of all soUd substances approaches zero at very

low temperatures.

3.1.5. Simon Expansion Liquefier

No summary of means for Hquefying helium would be complete

without mention of Simon's expansion Hquefier. In this liquefier, gaseous

hehum is pumped into a metal container which can be cooled to the

temperature of hquid or solid hydrogen. When the container, together

with its charge of helium, now at a pressure of 100 to 150 atm, has at- tained the desired temperature, it is thermally isolated from the sur- roundings. The pressure is then reduced slowly by valving the helium

to an external gas holder. At any stage of the expansion, the hehum left in the vessel will have expanded reversibly. It does work on the gas being pushed out. The overall expansion is rather efficient and the

amount of hquid left in the container when the pressure has returned to

1 atm is approximately 60 percent of the quantity possible with a completely reversible expansion. If the pressure and temperature are 150

atm and 11 K respectively at the beginning of the expansion, the container wiU be about four-fifths full of liquid when the expansion is complete.

An aU-important factor contributing to the success of the Simon

expansion hquefier is the vanishingly small heat capacity of the vessel

into which the helium is compressed. Substantially none of the liquid

formed is eyaporated in coohng the container from the starting temperature to 4.2 K. The charging half of the cycle of the Simon liquefier is relatively

wasteful of Hquid hydrogen. The heat absorbed by the hydrogen bath is the sum of the decrease of the internal energy of the hehum after it enters the container and the flow work of compression. The latter term

is ^mpv where m is the mass of a small quantity of entering gas, p the

pressure, and v its specific volume. If all of the gas enters at the same

temperature, Tu the total flow work is NRT, where A'^ is the number of moles of helium entering the vessel and R is the gas constant. The

amount of hquid hydrogen required to absorb the heat produced by the

flow work may actually exceed that needed to bring about the reduction

of the internal energy of the hehum. It is a case of comparing the magnitudes of two products; namely, RTi and CiTi — To), where To is the

final temperature before expansion.
62 LIQUEFACTION TECHNIQUES

3.2. COMMON HELIUM LIQUEFIERS

3.2.1. Helium Liquefaction Cycle

Nearly all of the liquid helium produced at the present time is made

in liquefiers which utilize liquid nitrogen for all of the help it can give

and, in addition, expansion engines at one or more temperature levels. In figure 3.3, a liquefaction cycle with nitrogen cooling and adiabatic

expansion of a portion of the helium at each of two temperatures with

isenthalpic expansion of the remainder is shown. The counterflow heat

exchanger is made up of six sections, although fewer sections are often

used. It is possible to construct a helium liquefaction cycle with adiabatic

expansion at only one temperature level when liquid nitrogen is used,

but the yield of liquid per mole of helium circulated is greater with

multiple level expansion. Ideally, the number of thermal stages of ex- pansion should be such that the entire temperature interval between

300 K and 4.2 K is spanned by the engines, just as sections 3 and 5 of the heat exchanger of figure 3.3 are bracketed by engines Ei and Ez,

respectively. The fraction of the total stream of helium which is to be

Li quid

Nitrogen Compression

Figure 3.3. Helium liquefaction cycle.
MANAGEMENT OF THE HEAT EXCHANGER 63

liquefied could then be cooled all the way from 300 K with the highest

thermodynamic efficiency.

Referring again to the cycle of figure 3.3, liquid nitrogen cooling

may be supplanted by a third expansion cylinder. Sometimes liquid

nitrogen is not readily available, or is expensive, or requires too much

attention. In such circumstances, moderate efficiency can be attained

with adiabatic expansion at only two levels and high efficiency with

expansion at three temperature levels. 3.2.2. Management of the Heat Exchanger

Broadly the function of the heat exchanger is to shuttle the re- frigerant to and fro between the compressor — which does work at room temperature — and the engine and throttle valve, which operate

at very low temperatures. This process should be performed with

minimum loss of refrigeration. The heat exchanger tends to be large and

costly. It is sometimes desirable to accept certain deficiencies on the

part of the heat exchangers in order to conserve space, shorten cool- down time, or save capital outlay. It is often desirable to squander liquid

nitrogen or expanded hehum at higher temperature levels in order to conserve refrigeration generated at lower temperatures.

Consider, for instance, section 1 of the heat exchanger of figure 3.3.

It carries the major part of the heat load of the entire heat exchanger

system. Note the effect of liquid nitrogen, generously applied, upon

the performance of this section. Without liquid nitrogen, the incoming

stream of compressed hehum must be cooled by the outgoing stream,

whose mass rate of flow is probably no more than 92 percent of that of

the incoming stream. Even if the heat exchanger is highly efficient, the

temperature difference between streams at the cold end will be at least 20 K. Although little or no refrigeration is being discarded to the sur- roundings, the 20 K deficit at the cold end of section I must now be made

good by refrigeration generated at a lower temperature. By the use of

either liquid nitrogen or expanded helium from an engine operating

at an elevated temperature, the mass rate of flow of outgoing fluid can

be made to exceed that of the incoming gas, and the temperature

"pinch" (minimum temperature difference between high and low pres- sure stream) can thereby be shifted from the warm end to the cold end

of section 1 so as to trade inexpensive high-level refrigeration for the more expensive low-level variety. In section 2, the temperature difference must grow larger toward

the cold end because of the unbalanced flow, but rtothing can be done
64 LIQUEFACTION TECHNIQUES

about it. In section 3, however, the opportunity recurs to change the

balance of flow by adjusting the portion of the high-pressure stream

which passes through engine Ei. The temperature of the compressed

hehum entering section 4 can, therefore, be brought very close to that of the expanded helium leaving the warm end of that section. In section 5, the provision for optimum flow through section 6, the Joule-Thomson

heat exchanger, automatically shifts the balance of flow heavily in favor

of the outgoing low-pressure stream so that the AT at the end of section 5

is very small. When refrigeration rather than liquid helium is the desired service,

the mass rate of flow of the outgoing stream of helium is greater, and

the requirement for Uquid nitrogen and/or expanded hehum is greatly

reduced.

3.2.3. The Joule-Thomson Heat Exchanger

For any chosen temperature at the warm end of the Joule-Thomson

exchanger, there is an optimum pressure. Reference is made to the temperature-entropy diagrams, figures 2.3 and 2.4 of Chapter 2. Note that the lines of constant enthalpy are curves having maxima at pressures

ranging from 35 atm at 20 K to 3 atm at 4.5 K. Isothermal compression

at lower pressures is attended by a decrease of enthalpy and at higher

pressures by an increase of enthalpy. At any given temperature there is a pressure at which the enthalpy is a minimum. At 8 K, for instance, the

minimum comes at 15 atm. If helium in this state is processed in a heat

exchanger and throttle valve, the maximum fraction liquefiable, as computed by eq (3.1) with data taken from figure 2.3, is 0.5, and the maximum refrigeration by eq (3.2) is 21.97 J/g. Had the pressure been either

higher or lower, the results would have been less favorable. It should be

observed at this point that in order to secure the performance indicated

above the pressure of the incoming stream cannot remain at 15 atm. It must be reduced to a lower value (4 to 6 atm) partway along the Joule- Thomson heat exchanger. Otherwise the temperature difference be- tween high- and low-pressure streams becomes zero, and heat transfer

stops before the enthalpy of the high-pressure stream reaches the re- quired value. The enthalpy of 15 atm hehum at 4.5 K is approximately

17 J/g, whereas at 4 atm and 4.5 K the enthalpy is only 11 J/g. In general,

it can be said that continuous adjustment of the pressure of the stream

of compressed hehum, so as to maintain the condition of minimum en- thalpy at every temperature level, will provide a maximum AT throughout the length of the heat exchanger for driving heat from one stream to
EXPANSION MACHINES 65

the other. This is especially important when refrigeration at 4.2 K is wanted. In practice, continuous adjustment of the pressure is not feasible, but two or three expansion valves in series will accomplish essentially the same result. The lower the temperature at the warm end of the Joule-Thomson

heat exchanger, the greater is the fraction of the J-T stream hquefied,

I or the greater is the refrigerative effect at 4.2 K if there is no accumu-

' lation of liquid. This does not mean, however, that the lower engine,

E2, should be operated at the lowest possible temperature for the sake

I of more intense precooling of the J-T stream. It must be remembered

j that the refrigerative power of the gas expanded in the engine decreases

I linearly with the temperature at which the expansion begins. Optimum

conditions are reached by due consideration of the two effects discussed

above, as well as the interaction with the remainder of the heat exchange

system. Precooling of the Joule-Thomson stream is rarely carried below

8 K and may stop at a level several degrees higher.

3.2.4. Expansion Machines

There are many varieties of reciprocating expansion engines in use. The most significant differences relate to the piston, its construction,

and the material of which it is made. The piston may be equipped with

conventional nonmetalhc piston rings which seal against leakage and

prevent contact between the metallic part of the piston and the cylinder

wall. The piston may be encased in a plastic sheath which is so closely

fitted to the cylinder that leakage is acceptably low. The piston and cylinder may be made of alloy steel of great hardness. The piston rod is very slender and is made to operate in tension so as to avoid sidewise

forces on the piston.

Another type of piston which shows considerable promise is an

elongated plastic plunger of uniform diameter, one end cold, the other

at room temperature. The cylinder is longer than the piston and of uni- form bore. Sealing is accomplished by a single O-ring at the warm end.

There is no leakage and little danger of seizure. There is, however,

appreciable clearance volume.

Turbines for expanding helium have been developed, and they are now being used in a few installations. It is expected that they will find extensive use in large liquefiers and refrigerators. The following sections

describe large-scale helium liquefiers. Not all manufacturers are repre- sented here, but the installations are typical of the state of the art in

this rapidly expanding field.
66 LIQUEFACTION TECHNIQUES

3.3. HELIUM LIQUEFACTION FACILITIES

3.3.1. Otis, Kans. Helium Liquefier (Ch. Trepp, Sulzer Brothers, Ltd., Winterthur, Switzerland)

It may be shown that it is more economical to transport hehum in hquid form than as compressed gas in high-pressure cyhnders. For large

delivered quantities (e.g., in trailers of up to 30 m^ capacity) the ratio of packing weight to contents is more favorable with liquid than with gas transports. For this reason, a liquefier processing helium was added

to a new helium extraction plant at Otis, Kans. The nominal liquefaction

capacity of this unit is 800 1/hr.

107.4 g/s

16.2 otm

81 K

32 K

15.12 K

10.2 K(

16 atm

6.1 K

228 q/s 8.6 atm

81 K

q- Qim 32 K

4 atm

.5.9 K

4 atm

5.65 K

283. 15 q/s

307.1 g/s

I.I otm. 79.6 K

I 8.5 atm : IT

52,1 g/

24.1 g/s

1.25 otm 445 K

1.25 aim

4,45 K

4 atm

14.47 K 12

28 g/s (800 Itr/h ) Figure 3.4. Flow diagram for helium liquefaction.
OTIS. KANS.. HELIUM LIQUEFIER 67

The flow diagram for the heHum liquefaction process is shown

schematically in figure 3.4. The Hquefaction process can be split into three temperature ranges corresponding to the different methods employed for cooling.

First range: 300 — 80 K: precooling with liquid nitrogen.

Second range: 80 — 10 K: cooling with work performing expansion.

A medium-pressure stream flows through two series-connected cold

helium gas turbines which operate at different temperatures.

Third range: 10 — 4.45 K: Joule-Thomson stage, double throttling

of hquefaction stream. The advantages of double compared with single

throttling result in smaller, more efficient low-temperature heat ex- changes.

In the design of a liquefaction cycle incorporating turbines, a dif- ference must be made between losses incurred through dissipation of mechanical work performed and those due to non-isentropic expansion.

The latter assume increasing importance as the temperature becomes

lower. In the case of heat exchanger design, an attempt was made to differentiate between the exchanger losses and the transfer losses. An

exact separation is not possible. The former are caused by the finite size of the heat exchanger surfaces, the latter by the difference in the

products of the mass rate and specific heat of the heat-exchange streams. The sum of all individual losses is equal to the difference between

the effective and the theoretical performance. If effective performance

is expressed as 100, we obtained for this hquefaction cycle losses equal

to 74 percent and a thermodynamic efficiency of 26 percent.

Since in this plant the feed gas is admitted at 80 K and 16 kg/cm-,

the above efficiency cannot be equated with that of a plant taking in helium gas under ambient conditions. A comparison has been made by

including the losses and thermodynamic efficiency of a process of iso- thermal compression at 310 K and isobaric coofing from 310 to 80 K.

It may be shown that plant efficiencies of 35 percent may be obtained

assuming isothermal compression.

The plant may be regarded as a closed cycle since it is not con- nected with a gas holder on the low-pressure side. In normal operation

the large hquid helium vessel does, in fact, have a similar influence on

the cycle as a gas holder. The low pressure (Hquefaction pressure) is maintained constant

with suitable control arrangements. Reduction of the feed rate would

result in a decrease in the low pressure, and for this reason the discharge

rate of the medium-pressure compressor is regulated with bypasses
68 LIQUEFACTION TECHNIQUES

Figure 3.6. Turbines for the Otis plant.
500 INCORPORATED 69

according to the liquefaction pressure. The regulating valve in the feed

line remains fully open.

At nominal load the compressor bypass valve is fully closed. Any

increase in the feed rate beyond nominal load capacity would result in the liquefaction pressure rising. In this case, the feed rate is reduced

and the excess gas let off to storage vessels. Constant pressure is maintained in the high-pressure stream by the bypass of the fourth compressor stage.

A photograph of the plant structure is shown in figure 3.5. All components below 80 K are enclosed in one evacuated cold box. Heat

exchangers are of the plate fin type.

A photograph of the cold gas turbines is shown in figure 3.6. Oil- lubricated high-speed centripetal turbines are employed. In air tests, efficiencies of 85 percent have been attained. Wheel diameters are 50

and 45 mm. Maximum speed is 100,000 rpm.

3.3.2. 500 Incorporated (S. C. Collins and Arthur Post)

The first helium hquefier to be manufactured in quantity was

produced by Arthur D. Little, Inc. The original model of this machine,

as developed by Collins in 1946, had a capacity of about 2 1/hr. Numerous

modifications have led to much higher rates of liquefaction. A later model, illustrated in figure 3.7, liquefies at the rate of 9 1/hr with vir- tually no operator attention. A large external dewar is employed, and the

Figure 3.7. Collins helium Hquefier.
70 LIQUEFACTION TECHNIQUES

helium is liquefied within the dewar by means of a remote expansion

valve. This system is made by 500 Incorporated, a subsidiary of ADL.

Hehum refrigerators based on a Stirling-type cycle developed by

McMahon and Gifford [1] - are also manufactured by 500 Incorporated.

These machines are designed for long periods of unattended operation and are used for cooling masers, superconducting magnets, etc., as well as for liquefying helium.

3.3.3. Union Carbide, Linde Division (A. J. Westbrock)

In 1962-63 the Linde Division of Union Carbide Corporation in- stalled what were then the largest helium liquefiers in the world [2].

These units were of the same basic design. Two installations were made

under contract with NASA at the Lewis Research Center, Cleveland,

Ohio, while the third was placed in commercial operation by Linde at Amarillo, Tex. Capacity of these machines exceeded 100 1/hr measured

in 7000-1 storage dewars which were installed with each unit. Liquid production is accomplished by means of a Claude process

as illustrated in figure 3.8. The Grade-A helium feed gas is compressed

from slightly above atmospheric pressure to 270 psig in three stages of compression.

The gas then flows through a water-cooled aftercooler into the

liquefier cold box, where the high-pressure gas is cooled to approximately 80 K against recycle helium gas and nitrogen vapor from the

nitrogen forecooler. The gas then flows through the nitrogen forecooler,

after which it passes through a silica gel adsorption trap which removes

trace impurities, principally nitrogen, from the helium feed gas. The

gas is then further cooled to a temperature of approximately 24 K against

low-pressure helium return gas. Next the gas is divided into two streams; one fraction is expanded in the expansion engine to 13 K while the other,

still at high pressure, is cooled further against expanded gas and then

to approximately 6.0 K by heat exchange against the low-pressure re- cycle stream. This very-low-temperature, high-pressure gas is then ex- panded through a Joule-Thomson throttle valve to about 4.5 psig with

resultant liquefaction of approximately 15 percent of the expanded

stream. The Hquid yield and non-fiquefied gas are separated in a phase

separator. The hquid made is transferred to the storage dewar, while

the low-pressure gas is passed countercurrent to the high-pressure

2 Figures in brackets indicate the literature references at the end of this chapter.
UNION CARBIDE, LINDE DIVISION 71

HELIUM

MAKEUP

COMPRESSOR

GAS

HOLDERy

P=270

T=300

SURGE

SNUBBER

T=279

GEL " TRAP

EXPANDER

T=24.5 P=266

T=297 HXI

LIQUID

NITROGEN

NITROGEN

FORECOOLER

SEPARA TOR

P = 4 NOTE:

TEMPERATURE, K

PRESSURE. PSIG

HXn

Hxm

T'12.9 Hxm

T=I3.4 C:=<h- LIOUID HELIUM

Figure 3.8. Schematic of Linde helium liquefier. helium gas progressively through each of the heat exchangers from the

cold end to the warm end, with the exception of the forecooler. This

recycle stream is joined by makeup helium and flows to the inlet of the

compressor first stage. A gas holder floats on the suction side. Both oil-lubricated and dry-lubricated three-stage reciprocating

compressors have been used in these installations. With the oil-lubri- cated machine, the system also incorporates an effective oil-removal

trap at the warm end. All of the compressors have special oil-pressurized packing gland seals to reduce helium loss to about 0.1 percent of

helium feed flow. The cold boxes are of double wall dewar type construction using

evacuated perlite insulation between walls. Heat leak is absorbed by

the use of liquid nitrogen coils on the inner vessel wall, on the gel trap,

and in the forecooler. Heat leak by conduction was minimized by the use of stainless steel in the inner vessel, piping, and supports.

306-922 0-68—

6
72 LIQUEFACTION TECHNIQUES

Heat exchangers in the cold box are of the brazed aluminum plate

and fin type. These provide maximum surface with minimum space

requirements and desirable physical properties at the extremely low

temperatures encountered.

The low-temperature expanders are reciprocating units of high

efficiency. Thermal efficiencies in excess of 70 percent were obtained.

The expanders are insulated by double-walled permanent vacuum perhte

insulated bell jar type vessels that can be easily removed for routine

maintenance of the expander cylinder and valves. Connection between

the liquefier cold box and the expanders is by means of super-insulated

piping.

Figure 3.9. Linde helium liquefier installation at NASA-Lewis Research Center.

Liquid hehum produced in the separators within the cold boxes

is transferred automatically by liquid level control through superinsulated piping into the 7000 liter super-insulated liquid helium dewars.

The normal evaporation rate is approximately 0.3 percent/day. To reduce

conductive heat leak to these dewars, their design incorporates a number

of design features, including re-entrant load rods, re-entrant piping,

helium-vapor-cooled radiation shields, in addition to the super insulation.
UNION CARBIDE. LINDE DIVISION 73

The Linde liquefier at Amarillo, Tex., is used for the commercial

production of liquid helium which is shipped throughout the country.

Demand for liquid helium increased to the extent by 1965 that the

AmariUo unit was modified in that year to utilize liquid hydrogen fore- cooling in order to increase production. This change has upgraded

production from the unit to more than 250 1/hr. The Hquefiers installed at the NASA Lewis Research Center are used in conjunction with a nominal 6-ft diam by 12-ft high space simulator also built by Linde. The installation is shown in figure 3.9. Space

simulation is accomplished within the chamber, which is maintained at a vacuum of 10~^^ torr through diffusion pumping and cryopumping.

The test object radiates to high emissivity walls maintained at liquid

helium temperature.

HeHum vaporized during testing is warmed to ambient temperature

and flows to the 92-ft-diam helium recovery gas bag shown to the left in the picture. Recovered helium is reprocessed through the liquefier

and stored in the dewars which are a part of the system.

An additional use of these helium liquefiers by NASA has been the

production of liquid helium for distribution to other NASA installations throughout the country.
74 LIQUEFACTION TECHNIQUES

3.3.4. Air Products and Chemicals, Incorporated (W. A. Snow,

A. R. Winters, R. B. Currie)

In 1958 Air Products and Chemicals, Incorporated, under contract

to the U.S. Navy, produced a portable helium liquefier for installation

at the Naval Air Station at Lakehurst, N.J. This liquefier was part of

a program supported by the U.S. Navy to reduce costs in the servicing

of the blimp fleet stationed at Lakehurst. The liquefier was capable of delivering to storage an equivalent of 35,000 scf per day (59 1/hr). The

liquefier is shown in figure 3.10 together with a Navy blimp it was to service.

A multipurpose helium refrigerator-liquefier was developed and

built by Air Products under contract with the RCA Service Company

for NASA Langley Research Center. The system was designed to meet

the following requirements:

Figure 3.11. Multipurpose liquefier-refrigerator.
AIR PRODUCTS AND CHEMICALS, INC. 75

1. Provide 1.32 kW of refrigeration below 5 K to a helium cryopanel

located inside a high-vacuum space chamber.

2. Maintain the cryopanel below 88 K in the presence of a 3 kW

heat load inside the chamber.

3. Produce 140 net 1/hr of helium when operating as a Hquefier.

A photograph of the apparatus is shown in figure 3.11. The process cycle

is quite complex and reflects the dual requirements of the refrigerator

liquefier system. A complete description is given by Snow, Winters,

and Currie [3].

Figure 3.12. CVI, Incorporated, add-on liquefier.
76 LIQUEFACTION TECHNIQUES

3.3.5. CVI Corporation, Helium Add-on Liquefier (C. B. Hood)

The CVI Add-On Liquefier was developed to provide an in-house

liquefaction capability to laboratories already operating refrigeration

equipment at the 20 K level. The design capacity is 34 1/hr when operating with a refrigerator providing 1.2 kW of refrigeration at 20 K. Helium

at 20 atm, precooled to 80 K, is cooled in three successive steps by the

20 K load stream and low-pressure flash gas before it is expanded into a collection tank. The warmed flash gas is recompressed in an auxiliary

compressor and returned to the refrigerator low side at 2.6 atm. The heat

exchanger and collection tank assembly is suspended in an evacuated

vertical cylinder. The exchanger assembly is shown in figure 3.12 with

the control panel/skid assembly shown beside it.

REFERENCES

[1] H. O. McMahon and W. E. Gifford, A new low-temperature gas-expansion cycle, Parts

I and II, Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, N.Y., 1960), Vol. 5, p. 354.

[2] J. L. Aberle and A. J. Westbrock, Liquid helium and nitrogen supply systems for space

simulators. Advances in Cryogenic Engineering (ed, K. D. Timmerhaus, Plenum

Press, New York, N.Y., 1963), Vol. 8, p. 190.

[3] W. A. Snow, A. R. Winters, and R. B. Currie, A multi-purpose high capacity hehum

refrigerator-liquefier, Proc. Institute of Environmental Sciences, April 21-23, 196.5; A. R. Winters and W. A. Snow, Capacity and economic performance of a large 5 K

hehum refrigerator. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus,

Plenum Press, New York, N.Y., 1966), Vol. 11, p. 116.
CHAPTER 4

Refrigeration

T. R. Strobridge ^

Page 4.1. Introduction 78

4.2. Thermodynamic relationships 79

4.2.1. First law of thermodynamics 79

4.2.2. Second law of thermodynamics 81

4.3. Refrigeration cycles and methods 83

4.3.1. The relationship of refrigerators and hquefiers 83

4.3.2. Liquid helium baths 84

4.3.3. The Simon process — adiabatic expansion 86

4.3.4. The Joule-Thomson process 88

4.3.5. The Brayton cycle 99

4.3.6. The Claude cycle 106

4.3.7. The Stirling cycle 116

4.3.8. The Gififord-McMahon cycle 119

4.4. Components and present-day refrigerators 122

4.4.1. Compressors 122

4.4.2. Heat exchangers 130

4.4.3. Expanders 132

4.4.4. Modern refrigerators 138

4.5. Efficiency, weight, and volume of 4 K refrigerators 142

' Cryogenics Division, NBS Institute for Basic Standards, Boulder, Colo. 80302. 77
4.1. INTRODUCTION

Much early low-temperature research was devoted to condensing the more stubborn of the normally gaseous elements. Liquefaction of the

lowest temperature species, helium, was finally accomplished by

H. Kamerlingh Onnes on July 10, 1908 [1].- For this contribution to the

technology, Onnes was awarded the Nobel prize. Many of the succeeding

experiments concerned the behavior of helium itself and led to the

discovery of the lambda transition and the unique, interesting properties

of liquid helium II. Investigations rapidly branched out to other materials

and superconductivity was observed in mercury by Onnes in April of 1911 [2]. It is interesting to note that in the short communication reporting the evident disappearance of resistance, Onnes remarked on the

advantages of having superconducting "films of molecular thickness,"

although he had in mind applications different than those encountered

today. Many of the unique properties of materials at helium temperatures

are currently being utiHzed for technological applications. Thus pulsed

or steady-state refrigeration is required simply to maintain the lowtemperature environment. Refrigerators are differentiated from liquefiers

because in the latter, the desired effect is the removal of energy from the

helium itself so that it wiU condense, while the refrigerator absorbs

energy at low temperatures from heat sources external to the working

fluid.This chapter is devoted to refrigeration methods encountered

today — evaporating liquid hehum baths, the Simon process, the JouleThomson, Brayton, Claude, Stirling, and Gifford-McMahon cycles. The

very special He^-He^ dilution refrigerators are excluded. Examples

of analysis techniques using the appropriate thermodynamics are given

for some of the cycles along with descriptions of typical modern refriger- ators and their components. Finally, data are presented on existing

closed-cycle refrigerators which allow prediction of the weight, volume, ' and power requirements of 4 K refrigerators of various capacities.

- Figures in brackets indicate the literature references at the end of this chapter.
THERMODYNAMIC RELATIONSHIPS 79

4.2. THERMODYNAMIC RELATIONSHIPS

4.2.1. First Law of Thermodynamics

The first law of thermodynamics relates the mass and energy fluxes which occur in natural events. Thereby it is used to predict the outcome

of processes, providing a tool for the analysis of the performance of

refrigerators and their components. Derivation of the first law has been

made in detail many times over and once more would serve no useful purpose. The nomenclature for the thermodynamic relations in this chapter has been adapted from Van Wylen and Sonntag [3], and the

reader is referred to their treatise for a thorough modern treatment of

classical thermodynamics. A control volume is defined as a volume en- closed by a control surface which surrounds the material to be studied.

Mass, work, heat, and momentum can flow across the control surface

into and out of the control volume. If attention is to be focused on a fixed

quantity of mass, the imaginary boundaries drawn about that mass define

a thermodynamic system. Work and heat may cross the boundary of a system, but mass may not. Referring to figure 4.1, the first law for uni- form state, uniform flow processes is

e + 7712 "2 H H Z9 — 77^1 \U\-\ h Zi

L \ 2gc gcJ \ 2gc gel

(4.1)

where ^c.v. and JFc.v. are the heat and work crossing the control surface,

771 is mass, h is specific enthalpy, u is specific internal energy, V is velocity, and Z is vertical position. The subscripts i and e refer to the

incoming and exiting flows and 1 and 2 refer to states 1 and 2 in the con-

"c.v- ^ c. J 1

2 2c

Figure 4.1. Pictorial description offirst law ofthermodynamics for uniform state, uniform

flow process.
80 REFRIGERATION

trol volume indicated by the subscript c.v. The proportionaUty constant

that relates force, mass, time, and length in Newton's second law is gc and the local acceleration of gravity is g. (In the SI system,

gc=lkg-mlN-s\)

The convention is adopted that heat transferred to the system is con- sidered positive, as is work done by the system. The terms uniform state, uniform flow mean that:

"1. The control volume remains constant relative to the coordinate

system.

2. The state of the mass crossing the control surface is constant with time and uniform over the various areas of the control

surface where the flow occurs.

3. The state of the mass within the control volume may change with

time but at any instant of time the state is uniform over the

entire control volume (or over several identifiable regions that

make up the entire control volume)." [3]

Equation (4.1) is an especially useful form of the first law. For example,

if there is no mass flow across the control surface, then the control volume

becomes the system and eq (4.1) reduces to

iQ2 = ilV2-\-m {u,-u,)V-\-^+{Z2-Z,) ^1- (4.2) 2gc gel

For a system undergoing a process from state 1 to state 2, eq (4.2) simply

shows that in the absence of mass transfer, the difference between the

heat transfer and the work done on or by the system is equal to the

change in internal, potential, and kinetic energies associated with the mass within the system. When both the amount and the state of the mass

within the control volume remain constant with time, eq (4.1) becomes

3)

Such a process is commonly referred to as a steady flow process where

^mi = ^me. If we imagine now that the control volume surrounds a stationary vessel that is to be filled by a pipe that passes through the
SECOND LAW OF THERMODYNAMICS 81

c.v. (4.4)

where the heat transfer, velocity, and work terms may or may not be

negligible. In many instances the velocity terms in the relations given

above will indeed be negligible. When those terms are included, all velocities must be relative to the same coordinate system. Care must be

taken to make the necessary incremental solutions if any of the terms

vary with time. An example of this situation is the cooldown of a lowtemperature refrigerator wherein temperatures throughout the entire

refrigerator do indeed vary with time until steady-state operation is reached. During cooldown, primary contributions to the internal energy

terms, contained within the brackets describing the contents of the control volume, come from the material in the piping, heat exchangers, and

other low-temperature apparatus in the refrigerator, and not from the

working fluid itself. Applications of the first law of thermodynamics to various refrigerator components will be given in later sections of this chapter.

4.2.2. The Second Law of Thermodynamics

The second law of thermodynamics leads to the definition of entropy

and is the basis for the concepts of the thermodynamic temperature

scale, the Carnot cycle, reversibihty, and irreversibility. As stated, the

first law predicts the outcome of an assumed process. However, the first law holds equally well for the process proceeding in the reverse direction.

The permissible directions for processes are given by the second law.

A corollary of the second law is that during a process, the total entropy

change of a system or control volume and the surroundings (the universe)

must always be greater than or equal to zero. A reversible process will cause zero entropy change of the universe while there will be a net entropy increase as a result of an irreversible process. An assumed

process can occur only if the net entropy change is zero or positive.

According to Van Wylen [3], "A reversible process for a system

is defined as a process which once having taken place can be reversed

and in so doing leaves no change in either the system or the surroundings." Reversible processes and cycles are ideals to which actual proc- esses and cycles can be compared to measure their performance. The
82 REFRIGERATION

thermal efficiency of an actual refrigerator is compared to that of a Carnot

refrigerator operating between the same temperature levels. The ratio of the amount of power required by a Carnot machine to the refrigeration

produced is given by

n ~ T (4-5) VCarnot ^ where Wn is the net power, Q is the refrigeration produced, T is the

temperature, on the thermodynamic temperature scale, at which the

cycle is producing refrigeration, while Tq is the temperature at which

heat is being rejected by the machine — usually about 300 K, the nominal

temperature of the earth's atmosphere or cooling water supply. Thus,

for a helium temperature refrigerator operating between 4.2 and 300 K

the best performance that could be achieved is 70.4 units of power re- quired per unit of refrigeration. As will be shown in section 4.5, it is not unusual for an actual machine to consume 100 times the Carnot power

per unit of usable refrigeration. The efficiencies of various processes

are also calculated by comparing them with appropriate, assumed,

reversible processes.

Wc Qc W, Q, Gas

at T Liquid at T

REFRIGERATOR LIQUEFIER Figure 4.2. Comparison ofenergy and mass transfers across control volume for a

refrigerator and a liquefier.
RELATIONSHIP OF REFRIGERATORS, LIQUEFIERS 83

4.3. REFRIGERATION CYCLES AND METHODS

4.3.1. The Relationship of Refrigerators and Liquefiers

In 1967 during the Eleventh International Congress of Refrigeration

in Madrid, Spain, Professor Kurti of the Clarendon Laboratories at Oxford University made some very interesting remarks about the subject

at hand. He noted that although there are many natural spontaneously

occurring events which result in temperatures above that of the sur- roundings, the creation of temperatures below that of the surroundings

requires the intervention of man. Very low temperatures are produced

both in helium refrigerators and helium liquefiers. Although the terms

refrigerator and liquefier bring to mind devices that are dissimilar in nature, a sHghtly closer examination reveals that in many ways they are much alike. Figure 4.2 illustrates the energy and mass transfers across

the control surfaces around a refrigerator and a liquefier operating at steady state. The effect desired from a refrigerator is the ability to absorb

energy at low temperatures. This refrigeration effect is shown as Qr in figure 4.2. The other energy transfers across the control surface are the

work, Wc^ that is necessary to produce the cooling; the work, We^ that is produced by expansion engines if any, heat leak through the insulation

to the low-temperature portions of the refrigerator, Qhi\ and the heat

rejected to the atmosphere, Qc, which is the algebraic sum of all of the

other energy transfers. Here it is indicated that the refrigeration effect,

Qr, is produced at a single temperature, T. A liquefier can be thought of

as a refrigerator and an independent side stream of gas which is to be

cooled from ambient temperature to its liquefaction temperature. One

essential difference between the two cooling devices is that in the liquefier, cooling for the side stream must be produced at all intermediate

temperatures from ambient to the liquid temperature. In practice liquid

is indeed produced internally in many helium refrigerators. The material

to be cooled is then attached to a vessel within the refrigerator which

contains the liquid helium. Any energy evolved by the refrigerated material itself or transferred to it from the ambient surroundings tends to raise its temperature. The energy is transferred to the then colder liquid

helium, a portion of which evaporates, and the cold gas is returned to the

refrigerator for recycling. If the object to be cooled cannot be placed

within the refrigerator cold box itself, the liquid can be piped to a remote

location through insulated channels and the cold' gas returned to the
84 REFRIGERATION

refrigerator. In either case, there is no mass transfer into or out of the

control surface. As the control surface is drawn around the liquefier in figure 4.2, there is mass transfer across that surface. A different situation

could be imagined wherein the liquid is drawn out of a reservoir in the

liquefier, used to cool some object, and warm, room-temperature gas returned to the liquefier. The control surface could be drawn around the

entire system and there would be no mass transfer across the control

surface. The essential difference is, however, that warm gas is returned

to be cooled and reliquefied.

There is one more point to be made about the situation described

by figure 4.2. There are no intermediate temperature thermal shields

indicated between ambient temperature and the lowest temperature

parts of the apparatus. In that situation, any heat leak into the cold box

reduces the useful refrigeration at the refrigeration temperature by an

equivalent amount or, in the case of the liquefier, reduces the hquefaction

rate. 4.3.2. Liquid Helium Baths

By far the most common method of providing cooling at 4.2 K is by using liquid hehum, supphed from a remote hquefier, as the heat sink. The material which requires this low-temperature environment is placed

in thermal contact with the liquid helium bath either by immersing

it in the liquid or attaching it to the container. The heat load imposed

upon the low-temperature sink consists of the energy evolved, if any, by

the cold object and heat leaks to the low-temperature apparatus via support structures, through insulations, and by radiation. As soon as a temperature difference providing the driving potential is established,

energy will be transferred to the colder liquid. If the vessel containing

the liquid is maintained at a constant pressure, i.e., allowed to vent, the

liquid will evaporate at a rate sufficient to maintain equilibrium temperature at that pressure. The cold saturated helium vapor leaving the vessel f has a refrigeration potential and in some instances the vent pipe is attached to intermediate temperature shields which reduce the amount

of heat transfer from the surroundings to the low-temperature space, thus decreasing the liquid evaporation rate. The heat of vaporization of liquid helium (see chapter 2) is 0.725

W-hr/1 at 1 atm. The rate of liquid evaporation is found by dividing the

total heat load (supphed to the Hquid) by the heat of vaporization of the

hquid. Thus a 1-W heat load will evaporate 1.38 1 of hquid per hour and
LIQUID HELIUM BATHS 85

large heat loads require substantial amounts of the relatively expensive

hquid. Each new cooling requirement forces a choice between refrigerating with bulk liquid helium or with a closed cycle refrigerator. Winters

and Snow [4] describe an economic analysis of providing refrigeration

at liquid heUum temperatures for large heat loads. The analysis takes into account the capital and operating costs, both direct and indirect, for var- ious duty cycles. Their results show that for a 100-W heat load, unless the

system is to be operated in excess of 95 hr per year over a 10-year period,

it is economically favorable to provide the refrigeration with bulk liquid

hehum. At higher heat loads the economic break-even number of operating hours per year becomes smaller.

V| =4 Liters

C.v. (I)

t

V =4 Liters

V, = I Liter

(2)

Figure 4.3. Liquid helium bath used for example calculation of the heat transferred

to the liquid in an insulated enclosure.

Since the heat leak to insulated low-temperature liquid containers

is often estimated by measuring the amount of gas that is evolved, let us examine the 5-1 hquid hehum vessel in figure 4.3. Initially the vessel

is shown to contain 4 1 of saturated hquid helium at 1 atm pressure with

a 1-1 ullage volume filled with saturated vapor. In addition to heat transfer across the control surface there is mass transfer at the vent. After a period of time, 3 1 of liquid have evaporated because of the heat influx,

leaving 1 1 of saturated liquid and 4 1 of saturated vapor. The remainder

of the helium has passed out through the vent. Data from chapter 2 show
86 REFRIGERATION

that 7827.7 J are required to effect a phase change for 3 1 of liquid helium

and that 374.5 g of hquid will evaporate. Misinterpretation of the situa- tion seems to occur often enough that it is worthwhile to point out the

errors which can be made. The proper form of eq (4.1) for the container is

^c.v. = ruehe H- mi^ui^ -h niv^Uv^ — ruf^ui^ — niv.Uv^. (4.6)

In order to evaluate eq (4.6), two types of dynamic measurements,

the mass flow out of the vessel arid the liquid level, must be made and

the container geometry must be well known. Obviously, eq (4.6) is not

the most convenient approach for determining the heat transfer, since

if the change in liquid level is known, the mass of liquid evaporated can

be determined from the vessel geometry and the product of the heat

of vaporization, and the mass of liquid which has changed phase gives

the heat transfer. Often it is more convenient to measure the mass flow,

TUe, than it is to monitor the liquid level. All too often it is assumed that the internal energy-mass products in eq (4.6) are negligible and ^c.v.

is equated to niehe. If this is done for the example in figure 4.3, the cal- culated heat leak is 28 percent too high. The product nieLv, where

is the heat of vaporization, which is sometimes used for an approximate

calculation, leads to a computed heat input that is 13.3 percent too low

for the example, because only 86.7 percent of the mass that changes

phase passes through the vent. Equation (4.7), „ nieLi' Qc..=- pr, (4.7)

where vi and Vv are the specific volumes of the saturated liquid and

vapor, takes into account the amount of mass that changes from liquid

to vapor but is retained in the vessel and is the most convenient path to correct results providing the assumption that the liquid and vapor are saturated is valid for the situation in question. If there is temperature

stratification in the liquid or vapor, the measurement of rue will not be

sufficient to determine the heat load to the liquid.

r 4.3.3. The Simon Process — Adiabatic Expansion

Before the development of the Collins liquefier, the Simon process was often used to batch-produce small amounts of liquid helium. When

a high-pressure gas is allowed to vent, the fluid in the thick-walled con-
THE SIMON PROCESS -ADIABATIC EXPANSION 87

tainer expands isentropically and its temperature decreases as the

pressure is lowered. Although the expansion process is isentropic, as the

fluid cools there will be heat transferred to it from the warmer walls of the vessel, resulting in an entropy increase of the fluid. Fortunately at liquid hehum temperatures the specific heat of most container materials

is very low and such entropy contributions may be negligible. In practice, Uquid helium is produced by the Simon process in the

following manner. The apparatus is designed so that the high-pressure

vessel can be either thermally isolated from its surroundings or brought

into contact with one of two heat sinks, usually liquid nitrogen and

liquid hydrogen. The vessel is filled with helium, usually at a pressure in excess of 100 atm, and that pressure is maintained by introducing more

helium as the vessel and the hehum are cooled to liquid nitrogen temperature. Then hquid hydrogen is used to cool the apparatus to about 20 K. The Hquid helium yield is increased substantially if the pressure over the

hydrogen is reduced so that the temperature of the vessel approaches

10 K. As soon as the lowest temperature has been attained, the highpressure vessel is isolated from the surroundings and the expansion

process is carried out by reheving the pressure in the cold vessel. The amount of liquid helium produced depends upon the state of

the fluid prior to the expansion process. The technique is described by

Simon [5, 6], and Pickard and Simon [7]. Scott [8] develops the correct

relationship for predicting the liquid yield when using this technique.

Primarily because of economic advantages, the Simon process has

been used recently by Taylor [9] to intermittently refrigerate sodium

magnet coils at approximately 7 K. The system produces 15 kW of

refrigeration for a period of 1 min and is capable of repeating that duty

every 6 hr. The high-pressure helium and the vessel are first cooled to 12 K by a closed-cycle helium gas refrigerator. The flow path conducts

the fluid from the vessel to the sodium magnet coils and then through

thermal regenerators into ambient temperature low pressure helium

storage. Just prior to refrigerating the magnet coils, the helium in the

vessel is at 12 K and about 160 atm. Since the coils must be maintained

at about 7 K during operation, the vessel must be blown down to about

70 atm in order to reduce the temperature prior to the experiment and

that helium is used to precool the magnet coils and the regenerators.

The thermal regenerators are used to cool the helium gas as the vessel is being recharged in preparation for another refrigeration cycle, thus

minimizing the loss of coohng capacity.

306-922 0-68—

7
88 REFRIGERATION

4.3.4. The Joule-Thomson Process

The lowest temperature stage in virtually all liquid-helium-temperature refrigerators consists of a counterflow heat exchanger, an

expansion valve, and an evaporator. The expansion process through the

valve is irreversible and the fluid performs no external work, although

prior to passage through the restriction it has the potential for doing so by virtue of its elevated pressure. In section 4.3.1, the energy balance

around an entire refrigerator showed that the refrigeration capacity

would increase if the work produced by expansion machines were ex- tracted from the control volume. Therefore, if the pressure reduction

of the fluid occurred in a work-producing expander, rather than irre- versibly through a valve, the cycle efficiency would increase. In most

instances, the valve has been chosen over a mechanical expander be- cause the increased cost and complexity of the expander were judged

to overshadow the benefits to cycle performance.

The Joule-Thomson process is defined as the adiabatic isenthalpic

reduction in pressure of a fluid flowing through a restriction in a passage.

Although this process takes place only in the expansion valve, the additional energy transferred in the heat exchanger and evaporator will

60

0 10 20 30 40 PRESSURE , atmospheres

Figure 4.4. Helium inversion curve.
THE JOULE-THOMSON PROCESS 89

also be considered. The Joule-Thomson process has been extensively

analyzed by Dean and Mann [10] and much of the material in this section

is from that work. Because the Joule-Thomson process is used so ex- tensively in hquid-helium-temperature refrigerators, the differences

between the refrigerators can be thought of as primarily the various

methods used for precooHng the fluid before entering the Joule-Thomson

exchanger.

The temperature change as a result of an adiabatic isenthalpic

process depends upon the properties of the fluid, the inlet pressure

and temperature, and the final pressure. The Joule-Thomson coefficient is

This coefficient may be positive, negative, or zero and the locus of points

where the Joule-Thomson coefficient is zero is called the inversion curve and is shown on temperature-pressure coordinates in figure 4.4. The

vapor pressure curve for helium appears in the lower lefthand corner of

that figure.

Figure 4.5. Joule-Thomson expansion process.

Regions of heating and cooHng upon expansion are shown in figure

4.5. A Joule-Thomson expansion process which has initial and end states within the inversion curve envelope where fx is positive will lower the

temperature of the fluid, while such a process occurring outside the

envelope will raise the temperature of the fluid. Expansion from a tem-
90 REFRIGERATION

perature above the maximum inversion temperature will always result

in heating. If the fluid is expanded from state A to state B, the temperature of the fluid will rise. Further expansion to state C results in cooling

to the original temperature and an end state of D will lower the temperature below the initial temperature. Since the same net temperature

drop is obtained by expanding the gas from either states C or ^4 to state D, there is no reason for expending the energy required to compress the

fluid to any pressure higher than the pressure at C. Lowering the initial temperature results in operation in the portion

of the temperature-pressure plane where the slopes of the isenthalps are greater and eventually where the isenthalps intersect the vapor pressure

curve. This intersection is the extension of the isenthalps into the two- phase liquid-vapor region and means that a fraction of the gas expanded

into this region is liquefied. State D' is the intersection of an isenthalp with the two-phase

liquid-vapor region and E' is the intersection of the isenthalp with the

isobar corresponding to the exit pressure of the expansion valve. The

choice of the final pressure corresponding to E' establishes the temperature of the Hquid following the expansion process. The pressures

at D' and E' should be as low as possible if maximum temperature drop

is desired.

The Joule-Thomson refrigeration system is a counterflow regenerative heat exchanger connected in series with an expansion valve and

an evaporator as shown schematically in figure 4.6. The compressor for the system operates at ambient temperature while the heat exchanger

in figure 4.6 operates generally in the cryogenic temperature range. Intermediate heat exchangers and expansion engines are not considered

as part of the thermodynamic system.

In order to cool the system down to the refrigeration temperature, an outside source of refrigeration must be used to precool the helium at station 1 to about 55 K and preferably much lower. With the system at the

uniform precooling temperature, the refrigeration process is started by

supplying a quantity of compressed cold gas to station 1. The refrigera- tion made available by the expansion process is used to cool the mass of metal of the evaporator and the return stream side of the heat exchanger.

The external heat load is withheld in order to expedite cooling. Cooling

the heat exchanger return stream side, between stations 4 and 5, results

in cooling the inlet gas stream between stations 1 and 2. For the ideal heat

exchanger considered here, stations 1 and 5 remain at the initial temperature while the temperature of stations 2, 3, and 4 is depressed. As the
THE JOULE-THOMSON PROCESS 91

T

S Figure 4.6. Schematic of a Joule-Thomson refrigeration system. temperature of station 2 is depressed, the temperature of stations 3 and 4

also becomes colder due to the Joule-Thomson effect. Thus the refrigeration process is self-perpetuating until an equilibrium condition is reached.

The stable equihbrium condition at which a Joule-Thomson refrig- erator is normally operated is shown by stations 1, 2', 3', 4', and 5 on

the temperature-entropy surface of figure 4.6. Temperature depression

at the expansion valve continues until liquid is condensed into the

evaporator. Refrigeration is obtained isothermally by the evaporation of

the hquid. A prerequisite for this type of operation is that the return stream pressure be less than critical pressure, since the temperature

at which the evaporator will operate is fixed by the vapor pressure of the

refrigerant.

This type of operation tends to be stable in temperature due to the

presence of Hquid and the magnitude of the heat of vaporization. Mass

flow through the evaporator must be matched to the heat load for continuous stable operation, as an excessive heat load will result in reducing

the quantity of liquid rather than in changing the system temperature. If corrective action is not taken, such as increasing the compressor flow

rate, the evaporator will eventually go dry, forcing the evaporator

temperature to increase into the non-isothermal operating region. In

contrast, insufficient heat load will result in flooding of any type of evaporator. Thus, the refrigeration process must be controlled in either

mode, but the case of isothermal operation is easier to control.
92 REFRIGERATION

Dean and Mann [10] made extensive calculations to obtain the

characteristics of the Joule-Thomson system. If the coefficient of per- formance method of rating refrigerators is applied to cryogenic systems,

the results are performance figures less than unity. Here the performance

has been rated using the reciprocal of the coefficient of performance and

is the compression power requirement divided by the refrigeration

absorbed. The power required for the compressor was assumed to be

equal to the power required to compress an ideal gas isothermally be- tween the same states: —m^^ln (P1/P5), (4.9)

where Wc is the compressor power requirement, m is the mass ffow rate of the fluid, R is the universal gas constant, Tamb is the compression

temperature, M is the molecular weight, and Pi and P5 are the compressor discharge and suction pressure. To obtain realistic refrigeration

power requirements, the power term as calculated by eq (4.9) must be

divided by the isothermal compression efficiency — a figure of 50 to 60

percent is common for reciprocating compressors using piston rings.

The useful refrigeration Q, absorbed at 4.2 K, was calculated from a first law energy balance. Performance is rated by the ratio WdQ, the dimensionless figure of merit. In order to put the calculations on a common

basis, division by the isothermal efficiency has not been performed and

is left to the reader who must consider the merits of his compressor.

When a refrigeration system is considered that utilizes other energy

sources, such as an auxiliary refrigerator, that power term must be

added to the numerator to obtain an overall performance figure.

Application of the first law of thermodynamics to the refrigeration

system of figure 4.6 where frictional losses due to velocity effects are neghgible yields the following energy balance:

Q=m{h^-hi) (4.10)

where Q is the refrigeration power. The externally useful refrigeration

is equa[l to the diff"erence in the total enthalpy of the refrigerant entering

and leaving the system at the top of the heat exchanger stations 1 and 5. The figure of merit, WdQ, is found by forming the ratio of eqs (4.9) and

(4.10).Calculations were made for a wide range of parameters consisting

of the heat exchanger operating pressures, precooling temperature, and
THE JOULE-THOMSON PROCESS 93

temperature difference at the top of the heat exchanger. The majority

of the calculations were made with a return stream pressure of 1 atm,

allowing isothermal refrigeration at the normal boiling point temperature.

Inlet stream pressure and precooling temperature were varied from

near the critical to near the maximum inversion pressure and temperature. The temperature difference at the top of the heat exchanger was

varied by reducing the temperature at station 5. The figure of merit

and the flow rate in gram moles per second required to produce 1 W

of useful refrigeration were calculated for each set of parameters.

Temperature differences between heat exchanger refrigerant streams introduce irreversibilities, but they are necessary to effect heat

transfer. Consider the equations which describe the heat transfer

between the two gas streams of the heat exchanger.

heat exchanger load = m j Cp^_^dT=m I Cp^^dT (4.11)

where a and b are intermediate temperatures of the high and low pres- sure gas streams (see fig. 4.7), and Cp is specific heat at constant pressure. The temperature and pressure dependency of the specific heat allows

HEAT EXCHANGER LOAD, Q

Figure 4.7. Heat exchanger cooling curve —general case.
94 REFRIGERATION

the possibihty that b> a for small values of Ti — T5. Since b represents

the temperature of the available refrigeration of the low-pressure

stream, the condition 6 > a is not consistent with the second law concept

of heat energy flow and is therefore indicative of an invalid choice of operating parameters. Either Ti— T5 must be increased or other operating pressures and precooling temperatures must be chosen. Increasing

T1 — T5 without changing the precoolant temperature will always in- crease WclQ because of the temperature dependence of the term (h^—hi)

in eq (4.10). The designer should investigate the temperature distribution

within the heat exchanger for both gas streams using eq (4.11). Even

though a minimal temperature difference occurs within or at the cold end

{T2 — T4) of the exchanger, the temperature difference T1 — T5 may be

used as a parameter to indicate heat exchanger energy losses. Once

this temperature difference is fixed, then the internal temperature dif- ferences become a function of the properties of the gas.

It is desirable to determine the combination of parameters that gives both the maximum obtainable refrigeration for a given compressor

size and the most economical operation. However, no practical single

set of parameters satisfies both requirements.

An examination of figure 4.5 shows that the maximum refrigeration

is obtained by an expansion process when the initial state of the gas

before expansion is on the inversion curve. The effect of inserting an

evaporator and a heat exchanger in the system is not so easily seen. Consider eq (4.10). The pressure, temperature, and mass flow rate at station 5 and the temperature and flow rate at station 1 may be con- sidered constant for an equilibrium condition. Thus, the only variable

is hi=f{P), and the condition for maximum refrigeration becomes ^-^^ = 0. This may be restated as

(-

, . , dTip \dPl„

or

Cf,'fJi=0.

The Joule-Thomson coefficient must be zero in order to satisfy

eq (4.12), as Cp is never zero through the operating range of the Joule- Thomson process. Thus, to obtain maximum refrigeration for a fixed mass

^) f^l =0 (4.12)
THE JOULE-THOMSON PROCESS 95

INLET PRESSURE (P|),atm. —

Figure 4.8. Generalized performance surface for a Joule-Thomson refrigeration system.

flow rate and precooling temperature, the inlet stream pressure must lie on the inversion curve. Figure 4.8 is an unsealed representation of the performance surface

for a Joule-Thomson refrigeration system giving the figure of merit and

flow rate as a function of inlet pressure for values of Ti — with F5 = 1 atm. Continuous curves of WdQ are shown for each precooling temperature (Ti) from near the normal boiling temperature to near the maximum

inversion temperature. The locus of the minimum WdQ is marked

"optimum," indicating the best operating pressures from the standpoint

of power requirement. The inversion curve has been superimposed on the

performance curve to indicate the states where maximum refrigeration

is achieved as a function of mass flow. Flow rate information, ~qj^^ where M is molecular weight, included

with the performance curves by dashed lines, gives the flow in gram
96 REFRIGERATION

moles per second required to produce 1 W of refrigeration. Since this is a flow rate divided by an energy rate, the time units cancel causing the flow

information to have units of gm-mole/J. Multiplying these values by

47.5 converts the flow information to cubic feet per minute (standard of

1 atm and 0 C) per watt of refrigeration — convenient units when dealing

with American compressor manufacturers.

Results of eq (4.11) are given by a line marked ''second law viola- tion," indicating a limit to the range of possible operating parameters.

Calculations of heat exchanger operation on or below this line result in T4 > T2 for isothermal refrigeration. Attempts to operate in this region

result in Ti — T^ becoming larger than the designer anticipates with a

INLET PRESSURE {F;),atm.

Figure 4.9. Joule-Thomson refrigerator performance - (Ti - T.-,) = 0.25 K.
THE JOULE-THOMSON PROCESS 97

corresponding reduction of performance — possibly to a point where no

useful refrigeration is obtainable. One exception to this statement may

be made if the refrigeration system is modified by inserting another heat

exchanger and expansion valve in series as discussed later in connection

with the Claude cycle.

Figures 4.9 and 4.10 are representative of the results of the calculations for two different heat exchanger temperature differences and

show the effect of the irreversibilities in the exchanger. For example,

at a precool temperature of 26 K, the optimum power required per unit of refrigeration increases by 32 percent as the heat exchanger temperature difference {T\ — T^) spreads from 0.25 to 0.50 K.

INLET PRESSURE { F> ) ,atm. Figure 4.10. Joule-Thomson refrigerator performance — {Ti — T^) = 0.50 K.
98 REFRIGERATION

Figure 4.12. Section offinned-tube heat exchanger.

(Courtesy of Air Products and Chemicals, Inc.)
THE BRAYTON CYCLE 99

Dean and Mann [10] present the results of similar calculations for parahydrogen and for nitrogen. Thus, information is available which

permits the study of a cascaded Joule-Thomson helium refrigerator

where the helium stream is precooled by a parahydrogen Joule-Thomson

system which is in turn precooled by a nitrogen Joule-Thomson circuit. Just such a refrigerator is shown in figure 4.11. This small unit provides

0.5 W of refrigeration at 4.4 K and can be supplied with vacuum shrouds

which are suitable for experiments in spectroscopy, x-ray diffraction Mossbauer effects and other current research. The compressed gases

may be supplied to the unit either from high-pressure storage cylinders,

in which case the low-pressure gases are vented to the atmosphere in an open cycle mode of operation, or the gases may be recirculated by a compressor in a closed cycle mode. One of the miniature heat exchangers

from this refrigerator is shown in figure 4.12.

4.3.5. The Brayton Cycle

The hydrogen and nitrogen precooling circuits in the cascaded

Joule-Thomson hehum refrigerator remove energy from the helium as

it is progressively cooled from room temperature to the inlet of the final heat exchanger. The same energy removal is often accomplished by

utilizing the refrigeration produced by one or more mechanical ex- panders of either the reciprocating or turbo-machinery varieties de- scribed later. In 1873 a Boston engineer, George Brayton, conceived an

ideal heat engine combining constant pressure and constant entropy

processes. During the past several years the Brayton cycle, operating as an open cycle refrigerator, has been used extensively in aircraft cooling,

and the cycle is being considered for space power systems. The lowtemperature refrigeration performance of the cycle has been studied by

Muhlenhaupt and Strobridge [11] with nitrogen, parahydrogen, and

helium as the working fluids. A schematic flow diagram of the modified Brayton refrigeration

cycle is shown in figure 4.13, and the process paths are illustrated on

temperature-entropy coordinates in figure 4.14. The processes are

isobaric cooling in the high-pressure side of the heat exchanger between

stations 1 and 2, adiabatic expansion between 2 and 3, isobaric heat

absorption of the refrigeration load between 3 and 4, isobaric heating

between 4 and 5, and isothermal compression at Ti between the high and

low pressures. The process path shown in figure 4.14 between stations

5 and 1 is not isothermal because a finite temperature difference wiU
ENTROPY

Figure 4.14. Process paths of modified Brayton cycle refrigerator.
THE BRAYTON CYCLE 101

exist in the heat exchanger of any refrigerator. It is assumed that the

temperature of the fluid at the inlet to the compressor will be Ti due

to heat transfer with the surroundings through the interconnecting piping.

The modification to Brayton's original cycle is the assumed isothermal

rather than the isentropic compression. The compression power re- quired will be calculated from the same relationship used for the Joule- Thomson cycle discussed in the previous section. The figure of merit,

WclQ, is given by

RT, In

£_5 M{h4-hs) (4.13)

The expansion device does produce a certain amount of power which

presumably could be routed back to the compressor drive, reducing

the amount of power required from outside sources. In actual practice,

the amount of power produced by the expander is very small compared

to that required by the compressor, and is not recovered. Therefore,

the expander power is not reflected in eq (4.13). The Brayton cycle does

not produce refrigeration at a single temperature as does the JouleFlGURE 4.15. General performance of Brayton cycle refrigerator as a function of compressor exhaust pressure for a constant temperature differential and expander efficiency.
102 REFRIGERATION

Thomson process; rather, the heat sink temperatures vary from up

to w,hich will be referred to as the refrigeration temperature for this cycle since it is the highest at which heat is absorbed from the load. In order to see how the heat exchanger and expander efficiencies affect the figure of merit, assume that the compression temperature, the pressures, and the refrigeration temperatures are fixed. Neglecting for the

moment any heat leak to the heat exchanger, eq (4.1) yields the following
THE BRAYTON CYCLE 103

I 2 3 4 6 8 10 20 30 40 HIGH PRESSURE , atm. Figure 4.17. Brayton refrigerator performance — 50 percent expander efficiency.

enthalpy balance for the heat exchanger:

hi-\- h4 = h:^-\- h2. (4.14)

The assumptions have fixed the enthalpies at stations 1 and 4. Any losses

in the exchanger will depress the temperature and hence the enthalpy

at 5, forcing an enthalpy and temperature rise at 2. Even though the

compressor power requirement stays constant, any inefficiencies in the heat exchanger raise the operating temperature of the expander

and therefore decrease the enthalpy change across the evaporator as

/i.3 becomes larger. The isentropic efficiency of the expander is defined

306-922 0-68—

8
104 REFRIGERATION

20 40 60 80 100 120 T, , °K

Figure 4.18. Expander performance— 70 percent efficiency.

as

h'z —

hi — h!y (4.15)

where the prime indicates the enthalpy following an isentropic expansion

from the inlet to the expander. Losses in the expander have the same

effect as those in the heat exchanger, since the temperature at station 3

is increased and the amount of useful refrigeration depends on the en- thalpy at the outlet of the expander.
THE BRAYTON CYCLE 105

Figure 4.19. Brayton cycle helium flow requirement per unit of refrigeration as a function of the evaporator inlet and outlet temperatures.

If the refrigeration temperature and the expander and heat ex- changer efficiencies are selected, assuming ambient temperature compression and 1-atm low pressure, the power required to produce a unit

of refrigeration is a function of the high pressure. Figure 4.15 shows

a typical plot of eq (4.13) evaluated at various high pressures for a given

refrigeration temperature, heat exchanger temperature difference, and

expander efficiency. This curve shows that under the assumed conditions,

there is in fact an optimum compressor discharge pressure at which the

cycle should be operated if maximum thermal efficiency is desired.

Fixing the expander and heat exchanger efficiencies and repeating the

calculations for other refrigeration temperatures will produce a family

of curves similar to that in figure 4.15. The locus of the minima in those curves then gives the optimum operating pressures for that particular

expander efficiency and heat exchanger temperature difference over a range of refrigeration temperatures. Going one step further; and finding

the locus of the optimum pressures at other heat exchanger temperature

differences, then produces the charts shown in figures 4.16 and 4.17,

which are for 70 percent and 50 percent expander efficiencies, respectively. Performance charts such as these quickly show the effect of

component efficiencies upon the power required to produce the refrigeration. For example, at a refrigeration temperature of 20 K and a 70
106 REFRIGERATION

percent expander efficiency, increasing the heat exchanger temperature

difference from 2 to 4 K would require a 27 percent increase in compressor power to produce the same amount of refrigeration. At a 4 K

heat exchanger temperature difference, a change in the expander effi- ciency from 50 to 70 percent would divide the power requirement in

half. Other information needed for preliminary design work, such as the

expander power output, the expander inlet temperature, and the flow

rate, is given in figures 4.18 and 4.19. It should be noted that the minima

in the curves illustrated in figure 4.15 are quite wide, so that within

reasonable limits, the Brayton cycle refrigerator can be operated at non- optimum pressures without suffering a severe loss of efficiency. The

further effects of precooling, pressure drop, and reheat are discussed

in the reference given.

4.3.6. The Claude Cycle

The Claude cycle, as it has become known after the Frenchman

who constructed one of the first really successful low-temperature

expansion engines, has been used for many years in air liquefaction and

separation plants. This thermodynamic cycle has also found use in other applications at lower temperatures and has been incorporated in several different liquid helium temperature refrigerators. The cycle can be thought of as a Joule-Thomson system precooled by a Brayton

cycle refrigerator. In practice, however, the cycles are superimposed

and share the same warm heat exchangers and compressor, as shown

in figure 4.20 and companion figure 4.21. The compressor circulates the helium at mass flow rate rhi. As the

high-pressure stream passes through the warmest of the counterflow

heat exchangers HX I, it is cooled by the low-pressure stream to temperature T'z. At that point, some of the fluid is diverted through the expansion

engine at rate m2. The remainder of a flow, rhi — m2, is further cooled in HX II to below the inversion temperature by both the low-pressure

stream returning from the load and by the cold exhaust of the expansion

engine, and then passes through the final heat exchanger before entering | the Joule-Thomson valve. A portion of the stream is liquefied during the i expansion process, and this liquid is evaporated as the stream passes

through the heat absorber. The vapor then returns through the low

pressure side of the heat exchanger.
THE CLAUDE CYCLE 107

Figure 4.20 shows two expansion valves, one of which spHts the

Joule-Thomson heat exchanger into two sections. At times, intermediate

valves are necessary to produce the proper temperature profiles in the Joule-Thomson heat exchanger and, in fact, three valves are used in a 300-W, 1.85 K refrigerator which will cool a superconducting linear

accelerator at Stanford University [12].

Figure 4.20. Schematic of single-engine Claude refrigerator. The presence of an expansion valve between 4 and 5 will change cycle to the curve 5-6'

.

Steady-State conditions are assumed throughout for analysis. Potential and kinetic energy contributions to the fluid energy are negligible.

With these considerations, application of the first law of thermodynamics

to the various components of the system yields the following equations:
108 REFRIGERATION

For the upper heat exchanger HX I,

hi — h2 = hv2~hxx. (4.16)

For heat exchanger HX II,

{rhi — m-z) {hz — h:i) = Thiihu — hio) . (4. 17)

For heat exchanger HX III,

h3 — h4 = hio — hi). (4.18)

For the lower heat exchangers HX III and HX IV,

h:i-h = hio-h^. (4.19)

For the expansion valve,

/i7 = /i6. (4.20)

For the heat absorber evaporator,

Q= {mi-m2){hs-h-j). (4.21)

The figure of merit is

In (P.//>,.)

.

Rearranging eq (4.17) and substituting equahties gives the expression for percent of total mass flow through the expander

^=(^'~\"+;'°~^'^)xiOO. . (4.23)

rhi \ ko — /

Using the above equations and the expression for isothermal

compression work, Muhlenhaupt and Strobridge [13] have made calculations which assist in obtaining the most efficient performance from the
THE CLAUDE CYCLE 109

Figure 4.22. Cooling curve for heat exchanger with secondary valve as shown in figure 4.21.
110 REFRIGERATION

Claude cycle refrigerator. For a given set of heat exchanger and expander

efficiencies and a high pressure, there is, in terms of the amount of power

required per unit of refrigeration, an optimum engine inlet temperature.

Equation (4.23) shows that the flow distribution is a function of the engine

inlet temperature. The Joule-Thomson portion of the cycle becomes

more efficient as the inlet temperature of the exchanger (governed by

the inlet temperature of the expansion machine and its efficiency) is lowered. However, at lower temperatures the precooling refrigeration

from the expander decreases so the expander flow rate must increase

and more compressor power for precooling is required. The optimum

engine inlet temperature occurs at a balance between these two opposing

effects. For each set of conditions, the best value of T2 was found to within 0.05 K and the curves in figures 4.23 through 4.26 are all for optimum T2.

It is possible to select numerous sets of independent parameters

which are invalid for a single expansion-valve refrigerator due to implied

violations of the second law of thermodynamics in the Joule-Thomson

heat exchanger as discussed in section 4.3.4. The fluid entering the lowpressure side of that heat exchanger is assumed to be saturated vapor

at 1 atm, station 8, figure 4.20 and state 8, figure 4.21. Optimization of T2 determines state 10, selection of gives state 3, and solution of

eq (4.19) determines state 6. All these manipulations are straightforward

applications of the first law of thermodynamics. However, because the

specific heats of the high- and low-pressure streams are different, it may

be that under the assumed conditions, the calculated temperature at state 6 is lower than at state 8. This impossible situation is illustrated schematically in figure 4.22 where the temperatures of the high- and lowpressure streams are plotted as a function of the heat transferred be- tween them. This cooling curve shows, with the dashed high-pressure

line, the invalid temperature profile in the heat exchanger calculated

from the first law of thermodynamics for a single-valve arrangement.

The effect of a secondary valve between 4 and 5 is to lower the pressure

to some intermediate level where the specific heat of the fluid will give a positive temperature difference throughout the exchanger. The heat

balance ' below the expansion machine is not affected by the second

valve. During the computations, if a negative temperature difference was found between 6 and 8, a second valve was included in an attempt

to correct the situation. The region where such valves are required is indicated on the performance charts.
THE CLAUDE CYCLE 111

Typical results of the performance characteristic calculations of

the single-engine Claude-cycle refrigerator at 4.2 K are shown in figures 4.23 through 4.26. Figure 4.23 shows the compression power per unit

of refrigeration WdQ as a function of high-pressure Pi for various values

of expander inlet temperature T? and upper heat exchanger inlet temperature (low-pressure side) Tu. The values on the charts represent

the conditions for which the expander inlet temperature Tz is at an optimum level and the cycle inlet temperature Ti is at a temperature

of 300 K. The solid line marked "two valves" indicates the limit to the

range of possible operating pressures for the single-valve refrigerator.

Hence, an additional valve must be included in the system before higher

pressures are valid. The power requirement per unit of refrigeration

reaches a minimum value; and the cycle can be optimized for high

pressure as well as for expander inlet temperature. Figure 4.24 gives

expander power per unit of refrigeration at 4.2 K, WelQ, as a function

of high-pressure Pi for various values of expander flow m-z. Notice the

relatively small power output of the expander compared to the power

input to the compressor.

Figure 4.25 shows the expansion engine outlet temperature Tout as

a function of the expander inlet temperature T^^ for several values of helium inlet pressure Pjn, while figure 4.26 shows total volumetric flow

Ft as a straight-line function of the compression work per unit of refriger- ation WclQ for several values of high-pressure Pi.

The shapes of the curves in figures 4.23 and 4.24 reflect the characteristics of the Brayton cycle and the Joule-Thomson process, which in turn depend upon the thermodynamic properties of the working fluid. For illustration, at a t^Tz of 1.0 K, the minimum WdQ occurs at about

25 atm high pressure with an engine inlet temperature of about 27 K. The outlet temperature of a 70 percent efficient expander with an inlet

at 27 K will be ~ 12 K (see fig. 4.25). Since ^T2 is 1.0 K, the inlet to the

Joule-Thomson heat exchanger will be at 13 K. From figure 12 of Dean

and Mann [10], the optimum high pressure for the Joule-Thomson

process, with the inlet at 13 K and a 1 K temperature difference, is ~ 23 atm with the performance decreasing as the high pressure is decreased. The optimum high pressure for a Brayton cycle with the

discharge of the expander at 12 K is greater than 23 atm for an ambient

temperature difference in the heat exchanger of 10 K. Therefore, the

best high pressure for the combination of the cycles falls somewhere

in between — in this case ~ 25 atm.
112 REFRIGERATION

200' I I I I I I I I I I I I I I I I I I I I I I I I I 0 10 20 30 40 50 HIGH PRESSURE ,atm Figure 4.23. Claude refrigerator. performance — compressor power.

Consider now the effect of changing the value of while holding

the high pressure constant at ~ 25 atm. The performance of the cycle

increases as ^Ti becomes smaller because the Joule-Thomson process

efficiency increases. The; optimum engine inlet temperature rises shghtly, since the most favorable inlet temperature for the Joule- Thomson process varies in the same way. Helium behaves almost as an ideal gas at moderate pressures and down to fairly low temperatures.

This means that the temperature difference at the cold end of HX I will be about the same as at the warm end, and the curves for constant values

of T\\ will roughly parallel, or at least have the same general characteristics as, those for constant T2. At lower pressures, the performance of both cycles becomes

progressively unfavorable and the power requirement per unit of refrigeration increases rapidly. There is still a general trend of increasing
THE CLAUDE CYCLE 113

4 I I I I I I I I I I I I I I I I I I I I I I I I I I 0 10 20 30 40 50 HIGH PRESSURE , atm Figure 4.24. Claude refrigerator performance — expander power.

I optimum engine inlet temperature with better Joule-Thomson heat

j exchanger performance at constant pressure (see the curve for T2 = 22 K)

I in figure 4.23. In figure 4.24 the expansion engine power output, i.e., the amount of

|fj precooling performed by the expander, is given as a function of pressure

' for several values of Ar2. At low pressures, two effects are evident.

The relative amount of flow through the expander increases as does the

amount of energy removed from the cycle per unit of refrigeration. The

j total flow rate is proportional to WdQ, so the amount of precooling

' required is also proportional to WdQ as reflected by the shape of the

j curves in figure 4.24. The relative flow through the expander decreases

at higher pressures where a greater amount of work is produced per unit

' flow.
114 REFRIGERATION

For discussion purposes, consider a set of realistic component

efficiencies (i.e., realistic in the sense that no advance in technology

would be required to design and fabricate such components today).

Depending on the mass flow rate, expander efficiencies between 50 and

80 percent can be achieved. The temperature difference at the warm

end of HX I could be between 5 and 10 K, while ^Tz across HX III could

be less than 1 K, with 0.5 K considered to be a reasonable value. Choosing

as 5 K, t^Ti as 0.5 K and r)e as 70 percent, figure 4.23 shows a power

requirement of about 345 W/W for a 100 percent efficient compressor

and no pressure drop. Typical isothermal compressor efficiencies are about 60 percent. Estimates can be made of the effect of fluid pressure

losses. For example, if the pressure drop were 10 percent of total pres- sure in both the high- and low-pressure streams, the intake and dis- charge at the compressor would be 0.9 and 27.5 atm, respectively, for optimum conditions. Ideally, this would raise WdQ to about 365. If the

compressor efficiency is included, WdQ would be about 610. In addition.
THE CLAUDE CYCLE 115

Figure 4.26. Claude refrigerator helium flow rate requirement.

there are other losses, such as heat leak to the low-temperature space, which become critical in smaller capacity refrigerators and which

should be considered when translating the information given here to real

situations. Inspection of the periormance charts shows the possible improvement in thermal efficiency if the components could be slightly upgraded.

Consider making the following changes in efficiencies given in the

example.

1. Decrease AT-z from 0.5 K to 0.25 K,

2. Increase r]e from 70 to 80 percent, and

3. Increase rjc from 60 to 70 percent while holding ATi constant

at 5 K.
116 REFRIGERATION

From Muhlenhaupt and Strobridge [13] , WdQ is about 245. The assumed

pressure drop increases that value to about 260, and the compressor

efficiency increases it still further to a final value of 371, which is a considerable improvement.

About the same improvement in performance as in the previous

example is available by precooHng the cycle with liquid nitrogen. Precooling involves depressing Ti to the temperature of an evaporating liquidnitrogen bath and providing an additional heat exchanger to communicate

with the ambient temperature compressor. The liquid nitrogen con- sumption is determined by the efficiency of the warmest heat exchanger.

The cooling provided by the liquid can be converted into an equivalent

power requirement to be added to the compressor power. In most

instances it is estimated that the increase in WdQ would not exceed

10 percent and is considered to be negUgible. Inclusion of the equivalent power required for precooling would involve additional parametric

studies beyond the scope of this work. Typical conditions could be:

Ti = 77.36 K,

Ari = 2.0K,

Ar2=1.0K,

7)e = 70 percent,

r}c— 100 percent, and

WclQ = 24S at a high pressure of 19 atm with no pressure drop.

Many existing Claude cycle refrigerators are liquid nitrogen precooled,

but only the largest of these approach the power requirement of 468

W/W of refrigeration (see section 4.5) which comes from adjusting

the above conditions for 10 percent pressure drop and 60 percent compressor efficiency.

4.3.7. The Stirling Cycle

The operating principles of the Stirling cycle have previously been

discussed in Chapter 3. This periodic flow cycle, incorporating one or more regenerators, lends itself to compact construction with one piston

serving for both compression and expansion in a single-stage unit. A two-stage Stirling cycle refrigerator is shown in figure 4.27. The

top flange is the coldest of the heat sinks. The slightly larger flange is the intermediate temperature station and the bottom flange is the at- tachment point for the insulating enclosure. This particular refrigerator

will absorb about 80 W at 20 K.
THE STIRLING CYCLE 117

Figure 4.27. Two-stage Stirling cycle refrigerator which can simultaneously provide about 80 W of refrigeration at 20 K and 250 W at 75 K.

The power consumption at 12 K, the lowest temperature which can be attained, is 11.5 kW or 15.4 hp. (Photo courtesy of \orth American PhiUips Co.)
118 REFRIGERATION

Stirling cycle refrigerators are not in themselves suitable for re- frigeration at liquid helium temperatures because of the rapidly dis- appearing specific heat of the regenerator materials at low temperatures.

The lowest temperatures reported for pure Stirling cycle two-stage

refrigerators are about 12 K. However, helium Joule-Thomson circuits are effectively precooled by this type of refrigerator. The precooling

stations at two separate temperature levels are shown in figure 4.28,

which is a schematic of such a refrigerator suggested by Rietdijk [14].

Compr essor

Precooling

Stations

Figure 4.28. Schematic of Stirling precooled refrigerator with an expansion ejector.

(Courtesy of Philips Research Laboratories.)

Energy is transferred from the helium to the Stirling cycle refrigerator

through heat exchangers attached to the cold heads of the cooler. The

configuration shown is analogous to a two-expander Claude cycle. After

leaving' the lowest temperature precooling station at about 15 K, the he- lium enters the high-pressure side of the Joule-Thomson heat exchanger.

An expansion valve and evaporator at the low-temperature end of that exchanger would complete the cycle and provide refrigeration at about

4 K if the compressor suction pressure were at 1 atm. A rather more

elegant arrangement is shown in figure 4.28 which allows the compressor
THE GIFFORD-McMAHON CYCLE 119

suction pressure to be above the vapor pressure of the heHum at the

refrigeration temperature. The refrigerant is held at three different

pressure levels by an expansion ejector. At the highest pressure, full compressor flow enters the nozzle of the ejector where it is accelerated

to high velocities with a corresponding decrease in pressure. This region

of low pressure induces flow, m-zy from the low-temperature evaporator

through the low-pressure side of the final heat exchanger. The two flow streams mix and, as the velocity is decreased, experience a rise in pres- sure to P2. A fraction of the fluid emerging from the ejector may be

liquefied. The stream is divided with rhi returning to the compressor

through the low-pressure side of the heat exchangers and m-z is diverted

through the high-pressure side of the lowest temperature heat exchanger

to an expansion valve and the evaporator. As mentioned in section 4.3.4,

the cooling capacity of a Joule-Thomson system depends only upon the mass flow rate and enthalpy difference at the warm end of the Joule- Thomson heat exchanger and is independent of the refrigeration temperature. Therefore, the expansion ejector can offer advantages in two

ways. First, lower refrigeration temperatures can be achieved without

any decrease in compressor suction pressure; second, cooling may be

produced at 4 K with a compressor suction pressure in excess of 2 atm. Savings in compressor and heat exchanger size are made in either case. Haisma and Roozendal [15] report refrigeration temperatures as low

as 1.75 K with Pi at 31 atm and Pz at 1.15 atm. A production model of

this type of refrigerator is not ready as yet, but a 10-1/hr helium liquefier

featuring a noncontaminating roUing-diaphragm sealed-compressor,

heat exchangers laminated from metal gauze, paper, and resin, and an

expansion ejector which permits a 2.5 atm compressor suction pressure

has been marketed in Europe [16]. It should be noted that several

thousand Stirling cycle refrigerators are in the field in the United States today for cooling electronic equipment at temperatures higher than

liquid helium.

4.3.8. The Gifford-McMahon Cycle

By itself the Gifford-McMahon refrigeration cycle has about the same low-temperature limit as the Stirling cycle, since there is periodic

reversing flow through one or more regenerators. Since this cycle lends

itself very readily to staging, refrigerators precooled at three different

306-922 0-68—

9
120 REFRIGERATION

I (g) 4 COMPRESSED GAS SUPPLY

(g) EXPANDED GAS RETURN

CYLINDER

DISPLACER

HEAT ABSORPTION STATION Figure 4.29. Schematic ofGifford-McMahon cycle.

temperature levels have been built for liquid helium temperature service. Figure 4.29 schematically shows a three-stage unit. The three cylinders are made of low-conductivity material and are fitted with dis- placers, sealed against the cylinders at the top, whose movements are controlled from outside the cryostat. The three thermal regenerators

are indicated as are the three heat-absorption stations at 14, 35, and

80 K. The regenerators are packed with a material such as metal screens

to produce the desirable characteristics of high heat capacity per unit volume, low pressure drop, small void volume, and low axial thermal

conductivity. Imagine now that all three displacers are in their lowest

position and that steady-state operating conditions have been reached

so that the proper thermal gradient exists in the regenerator material
THE GIFFORD-McMAHON CYCLE 121

from room temperature to 80 K in the first regenerator. The compressed

gas supply valve is opened, and helium enters the system, raising the

pressure in the volume at the top of the cylinders above the displacers.

When the pressure in the regenerators and the volume above the dis- placers reach the maximum pressure, the displacers are all moved to the

upper position, forcing the gas above them through the regenerators and

into the volumes which now appear below the displacers. The gas is cooled as it gives up heat to the cold regenerators and the decrease in specific volume allows more fluid to enter the system from the compressed

gas supply. The supply valve is closed and the exhaust valve is slowly

opened. Refrigeration is now produced as each of the elements of fluid do

work on preceding elements as they pass out of the system. The fluid is warmed as it passes through the regenerators and through the heat

absorption stations. When the pressure in the system reaches its lowest

level, the displacers are moved to the lower position, forcing the re- maining cold gas out through the regenerators and to the compressor

suction. At this time the temperature distribution in the regenerators

has returned to its original profile. This refrigerator inherently has several

mechanical advantages. Most of the low-temperature apparatus is passive

except the displacers, which are shown as sealed at room temperature.

Later models have displacer seals at low temperatures, but the problem

is not serious since the pressure difference across the seals is only that

Figure 4.30. Internal components of Gifford-McMahon refrigerator.

(Courtesy of 500 Inc.)
122 REFRIGERATION

which is required to induce flow through the regenerators. The displacers are operated at low speed, typically less than 100 rpm, and the light

duty drive mechanism is used to operate the room-temperature inlet and exhaust valves. Difficulties encountered with contaminants in the

refrigerant are alleviated somewhat, since the exhaust phase of the

cycle tends to flush impurities back out of the system.

The nonmetallic displacer piston and small piston rod are in the

upper left-hand corner of figure 4.30, which shows the internal components of a prototype for a Gififord-McMahon helium refrigerator. In

order, to the right of the displacer, are the cylinder assembly, small

regenerator matrixes and spacers, the internal shell of the Joule-Thomson

heat exchanger, other regenerator matrixes, the external shell of the

Joule-Thomson heat exchanger, a radiation shield, and the mounting

flange. Refinements of this design are incorporated in a 2 W liquid

helium temperature refrigerator described by Stuart et al. [17]. In this unit the regenerators are inside the nonmetallic displacers, and the

internally finned tubing which constitutes the high-pressure side of the

Joule-Thomson heat exchanger is wrapped on a larger diameter tube

which is filled with perforated disks and spacers. Very close temperature regulation, ± 0.01 K at about 4 K, is reported.

4.4. COMPONENTS AND PRESENT-DAY

REFRIGERATORS

4.4.1. Compressors

Gas compression equipment is required for aU helium refrigerators.

The compressors directly affect the performance of the refrigerator in several different ways. In many applications, reliability and continuous

operation of the refrigerator are of paramount importance. There are other components in refrigerators whose partial failure or degradation

of performance may be overcome. There is no way, however, that the

operation can continue in the event of a compressor failure unless

standby compressors are available or perhaps high-pressure gas storage

has been provided. The compression equipment is usually the most

expensive single component of a helium refrigerator, at times accounting for about one-third of the capital cost. In addition to being expensive,
COMPRESSORS 123

the compression equipment is often responsible for the major portion

of the inefficiencies of low-temperature refrigerators.

Conventional oil-lubricated reciprocating compressors are often

used for helium refrigerators. Although the design of oil-lubricated

reciprocating compressors has been established through many years of experience and a wide variety of capacities is available, such compressors do present the problem of oil contamination of the refrigerant.

It is entirely possible to remove the oil contaminants from the refrig- erant stream, but this involves additional capital and operating expense,

and can be bothersome, to say the least. The oil can be removed from

the helium as shown in figure 4.31. First the high-pressure stream from

: --TOR FILTER ADSORBENT

BED

Figure 4.31. Removal of oil contaminants from refrigerant stream.

the compressor is passed through a separator which removes large

droplets entrained in the stream. Next, a filter is used to remove smallersized droplets before the stream enters the absorbent bed. The filter might consist of the container packed with glass wool. The absorbent

bed, if properly sized and designed, will remove essentially all of the

oil vapor from the stream. Care must be taken that microscopic droplets in the range from 0.01 to 3 /xm in diameter do not enter the absorbent

bed. Particles in that size range may pass directly through the absorbent bed because they do not possess enough Brownian motion to insure contact with the absorbent. Usual absorbents are siHca gel or activated

charcoal. There can be other associated problems. For example, when

one hydrogen temperature refrigerator was put on stream some years

ago, the low-temperature heat exchangers periodically became plugged

by frozen contaminants. It was found that while all of the oil was being

removed from the refrigerant, the contaminant was air, which had been

dissolved in the oil and was carried into the system.

There are several types of compressors which are used to avoid oil contamination. Figure 4.32 shows a diaphragm compressor, and figure 4.33 is an internal view. Here the refrigerant is completely separated

from any possible oil contamination except in the event of diaphragm

failure. The diaphragm is moved back and forth by, a column of oil which is driven by a reciprocating piston. Although long life has been
124 REFRIGERATION

Figure 4.32. Two-stage diaphragm compressor rated for 15 scfm at 245-atm discharge

pressure with 1-atm suction pressure and a 11.2-kW drive motor.

(Photo courtesy of American Instrument Co., Inc.) reported for the diaphragms, it is generally recommended by the manufacturers that routine replacement of the diaphragm be scheduled at 1500 to 2000 hr of operation. A diaphragm failure is commonly detected

in two ways; the first of these is a photoelectric device sensitive to naphthalene quinoline, which is present in most lubricating oils or can be

added if necessary. The detector is located in the high-pressure discharge

of the compressor and is sensitive to two parts per million of oil con- taminants. If oil is detected, the device puts out an electrical signal

which can be used to activate the soleniod valves, switches, or alarms

so that appropriate action can be taken. The other method of detecting

diaphragm failure is to sense the pressure in between the triple dia- phragms which are used in these compressors. Pressure-activated

switches then produce the needed electrical signals.

Other reciprocating compressors do not provide any lubrication

in the cylinder and separate the cylinder from the crankcase oils by
COMPRESSORS 125

Figure 4.33. Section view of diaphragm compressor.

(Courtesy of American Instrument Co., Inc.)

specially designed distance pieces. The sealing of the piston in the

cylinder is accomplished with either Teflon-filled rings or by using a

special labyrinth seal in which there is no contact between the piston

and the cylinder (see fig. 4.34). These labyrinth-sealed pistons are

usually run at high rpm. Since there is not a positive seal between the

two members, there is some leakage past the piston, decreasing slightly the efficiency of the machine. Tests have shown (Ziircher and Meier

[18]) that little is gained by putting plastic piston rings in a labyrinthsealed compressor. While there are leakage losses incurred in the labyrinth-sealed compressor, the additional friction losses caused by the

plastic rings rubbing on the cylinder walls result in about the same compressor efficiency, especially at higher pressure ratings. The higher

cylinder wall temperatures encountered when using the plastic piston

rings also tend to decrease the efficiency. Notice in figure 4.34 the oil wiper rings situated above the cross head and the nonlubricated piston
126 REFRIGERATION

Figure 4.34. Section view of labyrinth piston compressor.

(Courtesy of Sulzer Brothers, Ltd.)

rod packing at the lower end of the cyhnder. Oil can be prevented from

creeping up the connecting rod into the packing in several ways. First,

the distance between the oil wiper rings and the packing can be made

greater than the piston stroke so that the wetted area never enters the

packing. Second, a baffle plate installed on the piston rod in the space

between the wiper mechanism and the packing helps to prevent oil from moving up the rod. Some plastic piston ring designs use a broad

rider ring which supports the piston in the cylinder, accompanied by one or more compression rings which are spring-loaded against the

cylinder wall. An interesting combination of two types of nonlubricated

reciprocating compression devices is shown in figure 4.35. The volume

displacement of diaphragm compressors is limited to about 110 cfm

because of the large diaphragm diameters needed at low pressures. In
COMPRESSORS 127

Figure 4.35. Combination of plastic piston ring and diaphragm compression stages. A 94-kW drive motor is required for this 140 scfm helium compressor with 245-atm discharge pressure and 1-atm suction

pressure. (Photo courtesy of American Instrument Co., Inc.)

figure 4.35 the first three stages of the compression are accomplished

by plastic-ringed pistons and the final compression is via a diaphragm

stage.In comparison to reciprocating devices, turbomachinery has the

reputation for high reliability. This factor, in addition to the possibihty

of alleviating the oil contamination problem through the use of gas- lubricated bearings, has prompted the development of turbomachinery

for miniature hehum refrigerators. In these refrigerators the small
128 REFRIGERATION

turbocompressor rotors (see figs. 4.36 and 4.37) are supported on bearings lubricated by the refrigerant gas itself. Theoretically, these very

high-speed compressors should have indefinite life since there is no

metal-to-metal contact and hence no rubbing wear. The development

of these compressors presented a formidable task, since turbomachinery

is best suited to high flow rates. At the low flow rates required by the i refrigerators for which these compressors were developed, very high

rotational speeds are necessary to achieve reasonable efficiencies. Since

speeds approaching 120,000 rpm are common, gas-lubricated bearings \ become mandatory because conventional types of bearings are not suit- !

Figure 4.36. Miniature centrifugal helium compressor.

This two-stage compressor is rated at 35 scfm (3.02 g/s) with 0.82-atm inlet and 1.29-atm discharge. The power con- sumption of the 16-kg compressor is 1.43 kW at 85,000 rpm. (Photo courtesy of Airesearch Manufacturing Co.)
COMPRESSORS 129

able. Further complication arises because, in general, more stages of compression are required to obtain the same pressure ratio with turbocompressors as compared to positive-displacement compressors. The

development of suitable powder supplies to drive the compressor motor,

which is integral with the rotor shaft; the design of the motor itself with the associated heat-dissipation problems; the establishment of

bearing designs which allow startup, the acceleration through critical shaft speeds, and the shutdown — all add to the difficulties in these

development programs.

Separate approaches have been taken in the design of the compressors shown in figures 4.36 and 4.37. The centrifugal compressor

shown in figure 4.36 is supported on foil-type gas-lubricated bearings.

Both the thrust and journal bearings consist of alternating foils of Teflon and spring steel. Further details about the compressor and

bearings may be found in Wapato [19]

.

Tilting-pad journal bearings support the shaft of the four-stage

regenerative compressor shown in figure 4.37. These journal bearings

and the stepped-thrust bearings are self-acting, as are the foil-type

bearings. This particular 4-hp compressor has a pressure ratio of 1.75

BEARING

Figure 4.37. Regenerative helium compressor. A 3-kW induction motor drives shaft at 47.500 rpm. The compressor pressure ratio is 1.75 at a flow of 22.9 cfm. (Photo courtesy of General Electric Co.)
130 REFRIGERATION

I

and operates at 47,500 rpm (Colyer and Gessner [20]). The two compressors both suffer from the losses characteristic of turbomachinery.

These are electrical inefficiencies in the drive motor, windage losses at the wheels, aerodynamic losses in the flow passages, pressure drop in the intercoolers, and friction losses at the bearings. All these ineffi- ciencies require additional input power to accomplish the required

purpose.

The power required for compression of the refrigerant can be con- veniently estimated as follows. First assume that the refrigerant is an ideal gas and that it will be compressed isothermally. Given the flow

rate and the inlet and outlet pressures, calculate the ideal power rejj quired from eq (4.8). The next step is to apply a correction factor which I relates the performance of real compressors to that predicted by eq

(4,8). Experience has shown that JFc as calculated from eq (4.8) should

be divided by a factor somewhere between 0.5 and 0.6 for conventional

reciprocating compressors. That correction factor is somewhat lower |( for the miniature turbomachinery just described.

4.4.2. Heat Exchangers

The importance of highly efficient counterflow heat exchangers was

seen in earlier sections, where the effect of the heat exchanger perform- ance upon cycle efficiency was demonstrated. The first law of thermo- ' dynamics, apphed to counterflow heat exchangers, neglecting heat ' transfer from the surroundings, states simply that the total enthalpy I change of the warm streams must equal the total enthalpy change of the cold streams. The second law tells us that in order to minimize the

entropy increase associated with the operation of the heat exchanger,

the temperature difference between streams must be as small as possible. In principle, at least, temperature differences can be made smaller

and smaller by increasing the length of the fluid flow paths. However, ' this means that the pressure drop in the fluid streams will increase and

hence more compressor power will be required to move the fluid through

^ the system. The choice of an optimum heat exchanger for a particular I refrigerator depends upon the opposing effects of the heat transfer and

pressure drop characteristics of the exchanger in addition to the cost and the volume of the exchanger. The simplest type of counterflow heat

exchanger consists of two concentric tubes. The high-pressure stream

flows in one direction in the inner tube and the low-pressure fluid moves

in the opposite direction in the annulus between the inner and outer
HEAT EXCHANGERS 131

tubes. Modern heat exchangers are much more complicated and fall into two broad groups — the shell-tube, and the plate fin. In general, the

shell-tube type exchanger can be used at higher pressures while the plate

fin exchanger has the characteristic of a large amount of heat transfer

surface per unit volume.

Equation (4.24) gives the heat transferred from the warm fluid to the cold fluid across a differential area dQ = UATdA (4.24)

where U is the overall heat transfer coefficient and AT" is the temperature

difference between the fluids. It can be seen that an increase in either

the heat transfer coefficient or the area will result in the same heat

transfer at a smaller AT", which is desirable from the viewpoint of overall cycle efficiency. Efforts to improve heat exchanger performance have

led to a proliferation of different designs. Tubes finned both inside and

out are extensively used to increase the effective heat transfer surface,

as are the extended surfaces employed in plate fin exchangers. Accurate

Figure 4.38. Hampson type heat exchanger being wound.

(Courtesy of Air Products and Chemicals, Inc.)
132 REFRIGERATION

prediction of the characteristics of a new heat exchanger design is very

difficuk and usually heat-transfer coefficients and pressure-drop friction factors are determined from test models. Experimental correlations for 88 different surface configurations are given by Kays and London [21].

Treatments of the basic principles involved are contained in McAdams

[22] and Jakob [23]

.

The Hampson type of shell-tube exchanger consists of a number of helices of small tubing through which the high-pressure stream flows. Cylindrical members both inside and outside of the helices form an

annular space for the low-pressure stream. One of the difficulties that may be encountered with this highly efficient exchanger is the fact that several parallel paths (see fig. 4.38) are used for the high-pressure stream

and even flow distribution is difficult to achieve. On the low-pressure

side improper spacing of the high-pressure tubes will cause flow channeling, and again, uneven distribution. For maximum efficiency each of the high-pressure streams should transfer the same amount of heat over

the same temperature difference to the low-pressure stream. Nonuniform flow distribution and its effect on heat exchanger performance

are treated by Fleming [24]

.

The arrangement of the miniature counterflow coaxial tube heat

exchanger in figure 4.39 is one way to reduce the flow channeling problems. Each of the high-pressure tubes is contained in its own lowpressure channel, assuring a constant hydraulic diameter for the fluid flowing in the annuli. Another highly effective, Hghtweight miniature

heat exchanger of the plate fin variety is shown in figure 4.40. The fins in the core of this exchanger, which is brazed in a hydrogen atmosphere,

are 0.051 cm high and 0.00254 cm thick and are spaced 0.063 cm apart.

The design flow rate is 2.16 g/s of helium with a 3 percent pressure drop

in both the high- and low-pressure sides. At the inlet of the exchanger,

about 63.8 K, the design temperature difference between the high-pres- sure and low-pressure streams is about 1.10 K and the effectiveness of

the 5-kg exchanger is 0.978, where the effectiveness is defined as the

ratio of the heat actually transferred to the cold stream in the exchanger

to the maximum amount that could be transferred consistent with the

second law of thermodynamics.

4.4.3. Expanders

Even though counterflow heat exchangers are used to conserve the

cooling capacity of the low-pressure stream, additional cooling is re-
Figure 4.40. Miniature plate fin heat exchanger.

Flow rate is 2.16 g/s and a temperature difference of 1.1 K at the warm end of the heat exchanger leads to an effective- ness of 0.978. The exchanger mass is about 5 kg. (Courtesy of Airesearch Manufacturing Co.)
134 REFRIGERATION

Figure 4.41. Two-cylinder reciprocating expander which operates at about 20 K with ^ 72 percent efficiency. Speed is between 150 and 450 rpm and the unit produces 1.7 kW of refrigeration. (Courtesy of CVI, Inc.)

quired in refrigerators using the Joule-Thomson process. This additional refrigeration, called precooling, is often supplied by work-producing expansion machines. A portion, or all, of the high-pressure refrigerant

is passed through a mechanical expander where its potential for doing

work is exploited. The temperature of the fluid decreases as it gives up

energy in the form of mechanical power. The low-temperature, lowpressure exhaust stream of the expander is then used to cool other

refrigerant streams as they pass to lower temperature portions of the

refrigerator. As seen in figure 4.2, the cooling capacity of a refrigerator

depends directly upon the amount of work produced by the expansion

device. . If the amount of energy crossing the control surface can be

made greater because of an increase in the output of the expansion

machine, the refrigeration capacity is enlarged by an equivalent amount.

An expansion machine operating between two given pressure levels

will produce the greatest amount of work and the lowest exhaust temperature if the expansion process is isentropic, that is, adiabatic and

reversible. Any losses incurred in the expansion machine will decrease
EXPANDERS 135

Figure 4.42. Sectional view of reciprocating expander.

(Courtesy of Air Products and Chemicals. Inc.)

its refrigeration capacity and increase the exhaust temperature. The

efficiency of expansion machines is defined as the ratio of the actual

enthalpy change of the fluid as it passes through the machine to the

enthalpy change which would occur if the gas were expanded in an

isentropic process starting at the same initial conditions and ending

at the same final pressure. Typical expander efficiencies lie between

50 and 80 percent.

Since these mechanical expanders operate at depressed temperatures, unique problems are encountered. Ordinary lubricants are prohibited, and a means must be provided for transmitting the power

produced by the expander to the room temperature surroundings without

306-922 0-68— 10
136 REFRIGERATION

I

producing excessive heat leaks. In the reciprocating expansion machine

described in chapter 3, heat leak to the low-temperature expansion

cylinder is minimized by operating the piston and valves with long,

slender rods in tension. The self-centering piston is lubricated by a thin

film of gas leaking past it. Distance pieces of up to 6 ft can be used to isolate the low-temperature cylinders from the crankcase in the ex- pander shown in figure 4.41. This two-cylinder unit will produce about

1.7 kW of refrigeration at 20 K, and operates between 150 and 450 rpm.

The measured isentropic expansion » efficiency at about 20 K is 72

percent (Hood [25]). A larger reciprocating expander described by Snow

et al. [26], and shown in figure 4.42, uses wax-impregnated leather for piston rings. In this expander the room-temperature crosshead assembly

and the low-temperature cylinders are separated by 8-ft distances pieces.

The exhaust of the warmer expansion cylinder is normally about 12 K,

while the cold cylinder is about 6 K.

In recent years much attention has been devoted to the development

of turboexpanders, and a 100,000-rpm oil-lubricated centrifugal turbine

Figure 4.43. Miniature gas-lubricated turboexpander.

About 225 W of refrigeration were produced by this turbine operating between 20 and 1 atm. An efficiency of about 40 percent was measured with the inlet temperature at 30 K and a flow of 60 scfm. Improvements at lower pressure

ratios are mentioned in the text.
EXPANDERS 137

was shown in figure 3.6 of chapter 3. The theory of the externally pres- surized gas-lubricated bearings which were used to support the turbine

shown in figure 4.43 is presented by Sixsmith and Wilson [27]. This

miniature turboexpander revolving in excess of 500,000 rpm dissipates

the energy extracted from the refrigerant in a helium brake circuit via the impeller located at the opposite end of the 0.794-cm-diam titan- ium shaft. Titanium was chosen for the shaft material because of its high strength to weight ratio, and low thermal conductivity. When as- sembled, the 30 K turbine wheel is only about 1.27 cm from the nearest room-temperature gas-lubricated bearing, so a low conductivity shaft

is essential. At a pressure ratio of about 4.9 to 1, the turbine efficiency was 65 percent, and improvements to later versions yielded efficiencies approaching 80 percent. Self-acting tilting-pad journal and gimbaled

spiral-grooved thrust bearings aline the shaft of the 77 K turboexpander

shown in figure 4.44. As can be seen, the nozzle is bladed for partial

admission to the 1.59-cm-diam impulse turbine wheel. A three-phase

Figure 4.44. Disassembled miniature turboexpander which produces about 80 W ofpower

at 200,000 rpm operating at 77 K.

(Courtesy of General Electric Co.)
138 REFRIGERATION

permanent magnet alternator converts the mechanical energy to electrical energy for dissipation outside the low-temperature environment.

For further design and performance details see Colyer [20].

A third helium turboexpander is shown disassembled in figure

4.45. The nominal inlet temperature to this expander is reported by

Wapato [19] as 22 K. A full-admission nozzle directs the fluid to the

shrouded impulse turbine driving the alternator. The rotating assembly

is supported by self-acting conical foil bearings which absorb both

radial and thrust loads at the operating speed of 196,000 rpm.

Figure 4.45. Helium turbo-alternator.

Electrical power output— 13.1 W. Overall efficiency — 32.4 percent. (Courtesy of Airesearch Manufacturing Co.)

4.4.4. Modern Refrigerators

A number of liquid heUum temperature refrigerators are in operation

today. Refrigerators with capacities as small as about 1 W and very

large units (capacity in excess of a kilowatt) are on stream. A photograph

of a Gifford-McMahon cycle refrigerator during assembly is shown in figure 4.46. The unit is inverted in this view with the low-temperature

section extending upward. The displacer drive mechanism is seen at the bottom. Several refrigerators of this type are currently being used to cool low-noise maser amplifiers in satellite communication ground sta-
MODERN REFRIGERATORS 139

Figure 4.46. Gifford-McMahon cycle refrigerator during assembly.

(Courtesy of 500 Inc.)

tions. This particular unit produces about 1 W of refrigeration at 4.2 K. The refrigeration temperatures spanned by Claude cycle refrigerators

range from 1.85 K up to near the critical temperature of helium. Components of two lightweight Claude cycle refrigerators being developed

for the U.S. Air Force were shown in previous sections. The space

application for these small refrigerators demands lightweight, reliability,

and low power input. Gas-lubricated turbomachinery was selected to enhance the reliability even though it is difficult to obtain the highest

efficiency at low flow rates with this class of machinery. In contrast

to alternators which generate electrical power that is dissipated in a

load resistor outside the cryostat, the energy abstracted from the re-
140 REFRIGERATION

Figure 4.47. Cutaway view of Claude cycle refrigerator.

Forty-watt capacity will be used to cool a superconducting quadripole magnet. (Courtesy of AIRCO Co.)

frigerant as it passes through the turboexpander shown in figure 4.47

is dissipated in a separate hehum brake circuit. The brake circuit piping and the hehum-water heat exchanger are seen at the very top of fig- ure 4.47, which shows a mockup of the 40-W refrigerator which will be

delivered to Brookhaven National Laboratories. The turboexpander is located directly beneath the brake assembly in the cryostat. The canister

just below the expander is a filter, which points out the necessity of insuring that particulate matter does not enter the small passages in the nozzle
MODERN REFRIGERATORS 141

and turbine blading. The cylindrical heat exchangers and low-temperature apparatus are protected from the surroundings by laminar insulation and a gas-cooled thermal shield at an intermediate temperature.

A 2-W, two expander Claude cycle refrigerator has been designed

for cooling electronic equipment [28] . One of the features of this small- capacity unit is that, although the valves in the expansion engine are operated by bellows-sealed tension rods, the piston is sealed by carbonfilled PFE rings and operates the piston rod, which is made of stainless

steel tubing, in compression. The crosshead and motor-generator as- sembly are hermetically sealed. The heat exchangers are constructed

by laminating a series of metal disks with epoxy. The disks he in a plane

perpendicular to the direction of flow, and are separated by a thin film of epoxy. Heat is conducted from the warm to the cold stream via the

metalHc members, yet the axial conduction is reduced because of the

relatively low thermal conductivity of the epoxy. Multilayer insulation

occupies the vacuum space of the double-walled cold box and the

refrigerator can operate with helium at atmospheric pressure in the

inner vessel. All of the equipment, including the compressors, gas holder, cold

box with control console and expander assembly, gas supply manifolding

and liquid helium collection dewar, comprising a liquefier-refrigerator, are seen in figure 3.7 of chapter 3. This 10-1/hr liquefier, using the two

15-hp 35-scfm compressors, can deliver 35 W of refrigeration at 4.5 K.

In addition to the two expansion engines exhausting at approximately

60 K and 15 K, liquid nitrogen precooHng can be used to produce maximum cycle efficiency. Larger capacity units are available, producing

up to 400 W of refrigeration at 4.5 K precooled by liquid nitrogen and

slowly operating expansion engines with pistons made of a plastic

material sealed at room temperature.

The 800-1/hr helium liquefier at Otis, Kans.. discussed by Trepp in the chapter on hquefaction, is the largest capacity hquid hehum temperature facility at this time. Since that facihty is not designed to function as a refrigerator, the 1.4-kW helium refrigerator (see Snow et al. [26]) designed for the NASA Langley Research Center, Materials Research Space Chamber, is the largest capacity liquid helium temperature

refrigerator in service. Both plate fin and shell-tube heat exchangers are used in this two expansion engine Claude cycle refrigerator. Figure 3.11

of chapter 3 is a photograph of the cold box containing the heat ex- changers and expanders. Depending upon the operating mode selected,

this refrigerator will provide cooling at 4.5, 20, or 80 K, and has a liquefaction capacity in excess of 150 1/hr with 600 installed horsepower.
142 REFRIGERATION

Two Claude cycle refrigerators are in operation at temperatures

below the normal boiling point of helium. The 80-cm hehum bubble

chamber at the Rutherford High Energy Laboratory, England, is cooled

by a refrigerator which will absorb 73 W at 3 K in addition to a 500-W

80 K load arising from the bubble chamber and cryostat heat shield. The

plant features dry-lubricated compressors with carbon PTFE piston

rings, aluminum plate fin heat exchangers, and externally gas lubricated

high-speed turboexpanders. The design and fabrication of this refrig- erator are fully described by Firth et al. [29]. As mentioned, a 300-W

1.85 K refrigerator to cool a superconducting electron linear accelerator

has been delivered to Stanford University. Here, precooling is accomplished by the use of liquid nitrogen and two reciprocating expansion

engines. The Joule-Thomson heat exchanger in this unit has three ex- pansion valves in order to maintain the proper temperature profiles,

but the situation is further compHcated by the very low pressure in the return stream and by the problems of film creep and superleaks pre- sented by the superfluid helium.

4.5. EFFICIENCY, WEIGHT, AND VOLUME OF

4.2 K REFRIGERATORS

The size and power requirements of hquid helium temperature

refrigerators were determined in a survey by Strobridge and Chelton

[30]. These characteristics are important because in practice they may

limit the applicability and operating time of the equipment. Although

the data acquired through the survey show considerable scatter, the

expected trends can be developed and attempts were made to establish

guides by which the weight, volume, and power requirements may be

predicted. The data on liquid helium temperature refrigerators available

at that time are given in table 1. The first 13 of the 26 units existed

and, presumably, the performance data given were measured. The

remaining refrigerators were under development or were proposed and

the values shown may be subject to change. Capacities range from less than 1 W up to several kilowatts. Liquid nitrogen requirements are shown for the precooled refrigerators. The additional power necessary

to provide the precooling in lieu of liquid nitrogen was not added to the

values of WdQ for lack of a common basis for comparison. However, it is estimated that the average increase in total input power would be less than 10 percent.

Study of figures 4.48, 4.49, 4.51, and 4.52 shows that the functions

that have been defined are subject to arbitration because of data varia-
4.2 K REFRIGERATORS 143

OS CO CSI lO

1—I CO lO ON

l-H Tj<

S3 =t: 33^33^33 -T3 -O T3 rt"-0 -O -O "O 3 = 333>33"33 W\J --W ^

~a -a 3 3

OS d d

00 CO eg 00 o Ost^CO-^COCNlOrOfOLCO^t^fOi—ICO

lC ^ O ON O O O O eg O lC rf; CNJ O

On O fC OCM Tf" (N r-TtrOOCOOO^LOLOC^OOr-ONON ONvO-^r-HTf-CSrOTtTfTfONCOCMTfcOi—

I

^ CO CO O l-H CO

CS O-'^'OnCOOOi-hCMi—ih-fOOOOO

LOr-ONONOOfOCOO^t^cOLC CO-^tJ- vOCOCS—(OO-^OONt^O-^r-OvOOuCOOO CMt—iCOO cooN-^vcocO'-^ONCOONOr^fOOr-HLOt^oofOOOOvOLOTf

fo Tj-' csi d r-n' d d d

^ d

\o eg lc

ON lO CO

ON Li:;c-^dddddddd-«^<cooNdod " CM CO 1—I CM CS Tf OON CM OuOO-^T^-oO ONTf-CMOCMTfCVOTf-CM-^vO^ONTfO lOCMOCM-^ cm COOnOn\D

CM 00 CM lO CO CO CO ON

1^ r^CMOcOCOCOCOC--.—lOi-ONDLOt^LOONCOTtLOCMCOr^i—ilC-^O O rfCMr- lOOCOCM

vO CO I—I lO OCM o o CM o o o ^

lO O CM

_ _ _ CM lO_ d d CO irj d d d CM r-^ lO CO O CO CM o o o o o o lo

r-H O O

1^

CMCMLr^LOCMu-DLOTfLOrtcOCOcOiOC^jOO

Tf"" Tt <^" Tj^ Tj^ CM Tj<' -"^ CM CO CO ^
144 REFRIGERATION

100

7o CARNOT

0.1

x22

19

•^I7(3.6»K) #1

•^A|8 ^16(3.6

a23

-aI3-

A10

25

0.1 10 100 1000 CAPACITY (WATTS ) 10,000

Figure 4.48. Efficiency of liquid helium temperature refrigerators as a function of re- frigeration capacity.

tion. The dashed lines shown in the figures were established through

the authors' best judgment and were substantiated by simple least squares treatment. Clearly, extrapolations of the dashed lines should

be avoided or made with extreme care, especially for lower refrigeration

capacities. The general shape of the curves can be predicted from a knowledge of the performance of low-temperature machinery. However,

since the data are minimal, it is difficult to define the curves quantitatively.

The power required per unit of refrigeration for a Carnot refrig- erator was given by eq (4.5), and it is informative to form the ratio (-)

Percent Carnot = ^^-(^^ X 100. (4.25)

( ) V ^ / Actual
4.2 K REFRIGERATORS 145

100

WC/Q

i

2.5 )

AiJ22

N,a9(2.5''K) -X N - x8 - \

- 18 ' ii6(3 .6' <) WATT) •

?I7(3.6"K)

1

1

\

-

•A9

r -A' — A 0

2

- A 24 a25 - -a3

L

10 100 CAPACITY ( WATTS

)

1000 10,000

Figure 4.49. Specific power requirements of liquid helium temperature refrigerators as a

function of refrigeration capacity.

This ratio indicates the extent to which the real refrigerator deviates

from ideal performance and is given in table 1 and plotted in figure 4.48

as a function of refrigeration capacity. Since the coordinates of the plot are logarithmic, the scatter in the points shown amounts to considerable

deviation in the values. The units under development have been differ- entiated from those which have been constructed and tested by underlining the point numbers which refer to the entries in table 1. The dashed line in figure 4.48 was established using the data for existing refrigerators only for capacities of 10 W or less, and all of the

data for the larger units. This line was then used as the basis for all other curves relating to efficiency.

There does not appear to be any significant difference, within the

scatter of the data, in the performance of the refrigerators using the

different cycles. The efficiencies of the Claude and the Gifford-McMahon

refrigerators seem to be comparable although the majority of the experimental units are of the Claude type. The performance of both cascade
146 REFRIGERATION

4.2° K REFRIGERATORS

0.5

0.2

— TURBOMACHINERY

AS REPORTED AND CH ELTON

L

15 20 30 CAPACITY , WATTS

Figure 4.50. Comparison ofturbomachinery power requirements to figure 4.51

.

(Courtesy of General Electric Co.)

Joule-Thomson refrigerators is lower than the general trend, as is expected.

The adverse effects of miniaturization are apparent from figure 4.48. The heat leak to the low-temperature regions of the smaller units will be proportionally higher than in the larger ones and the net refrigeration

potential is decreased. In addition, it is more difficult to construct highly

efficient components in the small sizes, and certain other losses do not decrease in direct proportion to refrigeration capacity.

The input power required per unit of refrigeration is shown as a function of refrigerating capacity in figure 4.49. This graph shows the same information as figure 4.48, but the required specific power can be

read directly. It should be noted that all of the input energy and the
4.2 K REFRIGERATORS 147

energy- from the refrigerated load must be dissipated at ambient temperature. In both figures 4.48 and 4.49 it is apparent that the majority of the

small-capacity proposed refrigerators are expected to have poorer

thermal performance than that of the existing ones. An examination of the weight and volume data shows that, in general, the low-capacity

proposed refrigerators are smaller and lighter than the existing ones. Some sacrifice in efficiency has been necessary to eff'ect these gains.

Indeed, Colyer [31] subsequently presented the data for low-capacity

lightweight turbomachine refrigerators in figure 4.50, which gives specific power requirements 1.5 to 2 times higher than figure 4.49. Figure 4.51 shows that weight reductions of nearly a factor of 10

are expected for some of the low-capacity refrigerators. Two possible reasons for this significant potential gain are: First, it is probable that minimum weight was not a design criterion, since low mass is not of primary importance in a number of applications. Secondly, optimistic

estimates may have been made for the weights of the proposed units. However, it is reported that the weights and volumes of two of the experimental units wiU not vary significantly from the original estimates. The

dashed line was fitted only to the data for existing refrigerators and is a good representation even for the points at higher capacities (23 through

26). The low-capacity proposed Claude cycle units are generally fighter

than those of the other types. Since the majority of the data for the pro- posed refrigerators are for a capacity of about 1 W, it would not be

expedient to make general conclusions about the weight advantage of the Claude cycle for aU refrigeration capacities.

The dashed fine shown in figure 4.52 was established in a similar

fashion to those for refrigerator performance and weight. This line is a reasonable representation of the volume characteristics which may

be expected from present-day designs. The large-capacity refrigerators are intended to operate at a permanent installation and no effort was made to reduce significantly the weight and volume. The conclusions

which may be drawn from an inspection of figure 4.52 are that the potential for significant volume reduction is great and that the proposed

lower capacity Claude refrigerators again appear to be attractive. The data for the efficiencies of the various refrigerators are more

consistent than for their weights and volumes. This indicates that efforts have been made to utilize highly efficient components and it is not

likely that major improvements in the performance of the cycles will be

forthcoming. The efficiency of low-temperature refrigerators wiU slowly

improve as the components are refined. It is not unreasonable to expect
14« REFRIGERATION

lO"

10^

-• GIFFORD -MCMAHON ! SOLVAY

-A CLAUDE

-X CASCADE JOULE -THOMSON

. Underlined numbers indicate machines under development or proposed.

-

LL-

(\_ 1 —L- " T" /'

/

al3

/ '1

23 —i1.

-L- 1.

K 22 —A

/ 12 w

1

—A 1- /

T"

-L9(2 \A

4( 2,5 "Klfl 1^

19

5°/'1

- >l --tt{

'il8 .16 36 "K

1 T~

aI7(36 "K

10"'

I 10 10^ lO' 10* CAPACITY, wotts Figure 4.51. Weight of liquid helium temperature refrigerators as a function of capacity.

that significant improvements can be made over the weights of existing

units through careful design and, indeed, the majority of the proposed

refrigerators are expected to be considerably hghter. Substantial re- ductions in volume are predicted for the machines under development.

Based upon the trends established for the existing refrigerators,

it appears that a 4.2 K unit with a capacity of 10 to 50 W would require

700 to 1000 W of input power per watt of refrigeration. Such a unit would
4.2 K REFRIGERATORS 149

lo'

• GIFFORD - MC MAHON

- SOLVAY

-A CLAUDE

X CASCADE JOULE -THOMSON

• Underlined numbers indicate machine

under development or proposed.

I 1 I" AV /'

/

-/- / t~

6 25//

,/ —±.f

/ T

T

T

/

iA 12 TfT

X 22 /

T / T /

til

/ II 5 y

'

-•1 T"T ^ T

T

! /

\- c J. 5 2 9 us "K)

A

)

j

10"' I 10 10^ lo' 10* CAPACITY, watts Figure 4.52. Volume of liquid helium temperature refrigerators as a function of capacity.

weigh about 87 to 130 kg/W of refrigeration and occupy a volume of about 0.28 m^/W of refrigeration. If these figures are biased by the values

given for the proposed refrigerators in the 1 to 5-W range, the power

requirement would be about 1000 to 1500 W/W of refrigeration. Assuming that the weights and volumes can be cut to one-half of that given

by the present estimates through further development, then the specific

weights and volumes become 43 to 65 kg/W and 0.14 m^, respectively.
150 REFRIGERATION

REFERENCES

[1] H. Kamerlingh Onnes, The liquefaction of helium, Commun. Phys. Lab., Univ. Leiden, No. 108 (1908).

[2] H. Kamerlingh Onnes, V. The disappearance of the resistance of mercury, Commun.

Phys. Lab., Univ. Leiden, No. 122b (1911).

[3] G. J. Van Wylen and R. E. Sonntag, Fundamentals of Classical Thermodynamics

(John Wiley & Sons, Inc., New York, N.Y., 1965).

[4] A. R. Winters and W. A. Snow, Capacity and economic performance of a large 5 °K

helium refrigerator. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus,

Plenum Press, New York, N.Y., 1966), Vol. 11, p. 116.

[5] F. Simon, Helium liquefaction with performance of work, Z. Ges. Kalte-Ind. 39, 89

(1932).

[6] F. Simon, On the expansion method for liquefaction of helium, Proc. 7th Intern. Congress of Refrigeration 1, 367 (1936).

[7] G. L. Pickard and F. Simon, A quantitative study of the expansion method for lique- fying helium, Proc. Phys. Soc. (London) 60, 405 (1948).

[8] L. E. Scott, Ideal yield of a Simon hquefier. Cryogenics 3, 112 (1963).

[9] C. E. Taylor, Sodium cryogenic coils, Proc. 1965 Symp. on Engineering Problems of Controlled Thermonuclear Research, CONE 650512 (1965), p. 152.

[10] J. W. Dean and D. B. Mann, The Joule-Thomson process in cryogenic refrigeration systems, NBS Tech. Note. 227 (1965).

[11] R. C. Muhlenhaupt and T. R. Strobridge, An analysis of the Brayton cycle as a cry- ogenic refrigerator, NBS Tech. Note. 366 (1968).

[12] S. C. CoUins, R. W. Stuart, and M. H. Streeter, Closed cycle refrigeration at 1.85 K,

Rev. Sci. Instr. 38, 1654 (1967).

[13] R. C. Muhlenhaupt and T. R. Strobridge, The single-engine Claude cycle as a 4.2 K

refrigerator, NBS Tech. Note. 354 (1967).

[14] J. A. Rietdijk, The expansion-ejector, a new device for liquefaction and refrigeration

at 4 K and lower, Intern. Inst, of Refrigeration, Commission 1, Boulder, Colo. (Annexe 1966-5), p. 241.

[15] J. Haisma and K. Roozendal, Investigation of the behavior of an ejector in the low temperature region beyond the X-transition of helium. Intern. Congress of Refrigeration, 12th meeting, Commission 1, Madrid, Spain, Paper No. 1.37 (1967). To be published.

[16] G. J. Haaruis, A new type of helium liquefier. Intern. Congress of Refrigeration, 12th meeting. Commission 1, Madrid, Spain, Paper No. 1.36 (1967). To be published.

[17] R. W. Stuart, W. H. Hogan, and A. D. Rogers, Performance of a 4 K refrigerator. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, N.Y., 1967), Vol. 12, p. 564.

[18] A. Ziircher and H. Meier, Reciprocating compressors with non-lubricated cylinders

and labyrinth pistons or plastic piston rings, Sulzer Tech. Rev. 49, No. 1 (1967).

[19] P. C. Wapato, Development of a 4.2 K cryogenic refrigeration system, Garrett Corp., Los Angeles, Calif., Airesearch Div., Final Rept. No. 66-0347, Contr. No. DA-36-

039-AMC-03725(S), DDC AD486509 (1966).

[20] D. B. Colyer and R. L. Gessner, Miniature cryogenic refrigerator turbomachinery. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, N.Y., 1968), Vol. 13, in press.
REFERENCES 151

[21] W. M. Kays and A. L. London, Compact Heat Exchangers (McGraw-Hill Book Co., New York, N.Y., 1964), 2d ed.

[22] W. H. McAdams, Heat Transmission (McGraw-Hill Book Co., New York, N.Y., 1954), 3ded.

[23] M. Jakob, Heat Transfer (John Wiley & Sons, Inc., New York, N.Y., 1949), Vols.

I and II. [24] R. B. Fleming, The effect of flow distribution in parallel channels of counterflow heat exchangers. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum

Press, New York, N.Y., 1967), Vol. 12, p. 352.

[25] C. B. Hood, CVI Inc., Columbus, Ohio. Private communication (1968).

[26] W. A. Snow, A. R. Winters, and R. B. Currie, Multipurpose high capacity hehum

refrigerator-liquefier, Proc. Inst. Environmental Sciences Annual Technical Meeting (1965), p. 305.

[27] H. Sixsmith and W. A. Wilson, The theory of a stable high speed externally pres- surized gas-lubricated bearing, J. Res. NBS 68C, 101 (1964).

[28] C. E. Witter, Design of a closed-cycle helium temperature refrigerator. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, N.Y.,

1966), Vol. 11, p. 107.

[29] M. Firth, D. E. Ward, and J. B. Gardner, A 73 watt, 3 K helium refrigerator using gas bearing turbines. Intern. Inst, of Refrigeration, Commission 1, Boulder, Colo. (Annexe 1966-5), p. 227.

[30] T. R. Strobridge and D. B. Chelton, Size and power requirements of 4.2 K refrigerators. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, N.Y., 1967), Vol. 12, p. 576.

[31] D. B. Colyer, General Electric Co., Private communication (1968).

Trade names and units: The appearance of trade names in this chapter was required to give adequate credit to those who contributed

pictorial material used to illustrate typical devices. Such appearance in no way implies approval, endorsement, or recommendation of a specific

product by the National Bureau of Standards. Deviations from SI units occur where needed for clarity because of prevailing practice in certain

segments of the technology.

306-922 0-68— 11

CHAPTER 5

Storage, Distribution, and Handling

L. E. Scott,' R. C. YanMeerbeke,^ B. S. Kirk,' G. C. Nubel '

Page 5.1. Introduction 155

5.2. Storage systems 156

5.3. Vented containers 158

5.3.1. Liquid nitrogen-shielded containers 158

5.3.2. Vapor-shielded containers 161

5.3.3. The vaporization process , 163

5.3.4. Atmospheric pressure variations 164

5.4. Nonvented containers 166

5.4.1. Pressure rise 167

5.4.2. Supercritical venting 169

5.5. Insulations.. 170

5.5.1. Types of insulation 170

5.5.2. Simple high-vacuum insulation 173

5.5.3. Evacuated porous insulations 183

5.5.4. Evacuated multilayer insulations 186

5.6. Sloshing and thermal oscillation heating effects 192

5.6.1. Sloshing 192

5.6.2. Thermal oscillations 193

5.7. Container design 196

5.7.1. Materials of construction 196

' Air Reduction Company, Murray Hill, N.J. 07971. Present address: Columbia Gas Systems, New York, N.Y. 153
154 STORAGE, DISTRIBUTION, AND HANDLING

5.7. Container design — Continued

5.7.2. Vacuum retention 198

5.7.3. Internal supports 199

5.7.4. Neck plug hazards 202

5.7.5. Air transport and absolute pressure relief valves 207

5.7.6. Catastrophic loss of vacuum 211

5.8. Transfer of liquid hehum 213

5.8.1. Transfer line design 213

5.8.2. Cooldown losses 220

5.8.3. Pressure transfer 223

5.8.4. Gravity and pump transfer 231

5.8.5. Depressurization 233

5.8.6. Flash gas recovery 237

5.9. Large-scale liquid and gaseous helium distribution 239

5.9.1. The Airco distribution system 240
INTRODUCTION 155

5.1. INTRODUCTION

Many of the physical principles used today in liquefying and storing

helium were known to Kamerlingh Onnes [1] ^ and his colleagues when

they first Hquefied hehum in 1908. The technology and commercial

impetus to make liquid helium an everyday commodity developed in only the last 15 years. In 1908, as today, workers in low-temperature physics have devel- oped liquid hehum technology primarily as a means to obtain an en- vironment suitable for their experiments. Until recent years, it was

necessary for a low-temperature experimenter to be proficient in cryo- genic techniques, as well as in his own field of interest, in order to successfully conduct experiments in the helium temperature range. The

introduction of the Collins cryostat in 1949 greatly reduced the experimental difficulties. This machine provided a working environment

refrigerated to helium temperatures; a concept which was not utilized extensively. Modifications were made to allow liquid helium to be trans- ferred from the machine, stored, and used at will, thus eliminating the

machine as part of the experimental apparatus.

Since the development of the Collins machine, the demand for liquid hehum has grown rapidly. By 1959, several industrial concerns were marketing liquid helium to scientific communities. The development of large liquefiers to support these markets has made liquid helium

a readily avahable commodity. In 1958, 700,000 hters (700 m^) of hquid

helium [2] were produced in the United States; essentially 100 percent was produced by Collins machines. In 1964, an estimated 1,700,000 1 (1700 m^) were produced [3], of which only 50 percent was made in Collins machines. By 1966, the total hquefaction capacity exceeded

12,000,000 1 (12,000 m3) per year. The largest single boost to liquid helium production has been the

realization that the production and shipment of liquid helium are the

least expensive way of distributing gaseous helium [4]. Lower capital

investments and lower distribution costs are possible with liquid dis- tribution as compared to gaseous distribution, in much the same way

that these benefits have been achieved for many years in distributing

oxygen and nitrogen as liquids. The typical liquid helium consumer can now be supplied from large central liquefiers. Distribution is now made

routinely by air and truck.

Figures in brackets indicate the literature references at tiie end of this chapter.
156 STORAGE, DISTRIBUTION, AND HANDLING

This chapter describes the equipment and techniques used in the

handhng of hquid helium. Since many of the techniques are commonly

used in handling other cryogenic fluids, emphasis is placed on methods

which are unique with liquid helium. Many excellent publications contain information on the properties of insulation and construction materials

suitable for low temperatures. This information will be presented only

briefly, for the convenience of the reader. The last portion of the chapter

describes a liquefaction and distribution system for handling liquid

helium on a large scale.

5.2. STORAGE SYSTEMS

The low boiling point of heUum (table 5.1) would appear to be the

most difficult design criterion for liquid helium storage containers. It is only when we also consider the very small heat of vaporization of helium

that the true problem of liquid helium storage becomes apparent: namely,

limiting the heat transfer from the surroundings to the helium contents. The real significance of this is appreciated upon considering that the

heat of vaporization of water is some 450 times that of helium on a vol- ume basis. Without elaborate thermal protection, helium can exist as

liquid for only the most brief periods.

Table 5.1. Comparison of the physical properties of helium with other cryogenic fluids and lA^ater

Property He H2 Nz H2O

Density (Ib/ft^) 7.798 4.418 50.46 62.43

0.12 0.07 0.81 1.00 Normal boiling point (°F) -452.10 -423.2 -320.36 212

9.0 192.7 85.67 970.3

(Btu/gal) 9.4 113 579 8090

Most of the problems have been successfully overcome in present

containers, but not without high cost. Most small containers require
STORAGE SYSTEMS 157

an investment of about $25 per liter of capacity, making liquid helium

storage very expensive compared with other products that are commercially shipped. When compared to equivalent gas storage, the in- vestment for small liquid containers is six times greater than for standard gas cylinders. Very large (10,000-gal) Hquid storage systems may cost only half as much as equivalent gas storage. Figure 5.1 illustrates the

approximate cost of hquid containers now in existence.

Figure 5.1. Approximate cost, per unit storage capacity, for various size liquid helium

containers. Two different design criteria are commonly applied to liquid helium

containers depending upon the desired final product state, liquid or

gas. If the product is to be used as liquid, the container is designed for constant venting of the vapor evolved from liquid evaporation because

of heat leak. If a gas or vapor requirement exists, and the gas can be

recovered at the destination or use point, then the container is designed

for nonvented storage and transport, allowing the heat leak to cause a pressure increase in the container. Significant differences exist in the

design and operation of the two systems.
158 STORAGE, DISTRIBUTION, AND HANDLING

5.3. VENTED CONTAINERS

5.3.1. Liquid Nitrogen-Shielded Containers

The most common container in use today is the double dewar.

This design is best summarized by the description of Wexler [5]. Containers of this type are made by numerous manufacturers and have

been produced in sizes ranging from 5 to 1200-1 capacity, the more common sizes being 25 to 50-1 capacity (see figs. 5.2 and 5.3). The insulation system shown is high-vacuum insulation. Liquid

nitrogen surrounds the liquid helium insulation space to reduce the

heat transfer by reducing the temperature difference across the helium

insulation system. The liquid nitrogen "shield," in turn, is insulated

from ambient conditions by a similar high-vacuum insulation. Typical

performance and specifications of these containers are shown in table 5.2. An important contribution to the excellent performance of these

containers is the refrigeration effect of the cold helium vapor vented

through the neck tube. Helium is unique in that the heat of vaporization is relatively small compared to the specific heat of the vapor. As

an example, a 4 K temperature rise of the cold vapor requires as much

heat addition as the heat required to release the vapor from the liquid

phase originally.

Figure 5.2. Typical construction of small liquid nitrogen- DOLLY shielded containers.
LIQUID NITROGEN-SHIELDED CONTAINERS 159

Figure 5.3. Typical 25-liter, nitrogen-shielded, liquid helium container ready for shipment.

Note the foam rubber protective rings around circumference and metal cylinder to protect the relief valve assembly. Table 5.2. Representative capacities, evaporation rates and weights of nitrogen-shielded

liquid helium containers

CAPACITY EVAPORATION RATE WEIGHT -FULL

Helium Nitrogen Helium Nitrogen (lbs)

(liters) (liters) {liters/24 hr) (liters/24 hr)

10 12-15 0.20-0.4 P/4-3 100-125

25 33-35 0.30-0.45 2V4-4 175-190

50 42-60 0.45-0.5 23/4-4 290-360

100 48-65 0.5 -0.65 4V2-5 460-500
160 STORAGE, DISTRIBUTION, AND HANDLING

While venting, the cold vapor intercepts the heat being conducted

down the neck tube, reducing the heat reaching the liquid vessel to a

practically negligible amount. A mathematical analysis of this process can be found in Scott [6 (p. 239)] . The significance of the effect is shown

in Wexler [5] and L. Scott [7], where a tenfold increase in container

heat leak was found in nonvented containers versus vented containers.

Seemingly minor heat sources can have a measurable effect on the

loss by evaporation. Radiation down the neck tube can be significant.

Wexler [5] found that blackening the neck tube reduced evaporation

losses in experimental dewars. Increase in evaporation rates has been

observed by slightly tilting the container. This results from the bumper

supports located in the vacuum annulus touching the warmer wall in the tilted position, which does not occur when the container is exactly

vertical.

VACUUM SEAL-OFF AND

POSITIVE SAFETY VALVE

LIQUID HELIUM RESERVOIR VACUUM SPACE UMINATED INSULATION

NORMAL VENT LINE AHACHED TO COPPER

SHELL (Gas Shield) COPPER SHELL

PLASTIC BUMPER

VACUUM SPACE

Figure 5.4. Construction of a vapor-shielded container, using only one

refrigerated shield.
VAPOR-SHIELDED CONTAINERS 161

5.3.2. Vapor-Shielded Containers

The concept of using the refrigerating capacity of the vapor has

been used to ehminate the need for Uquid nitrogen shielding. Two

methods have been used. The first of these (fig. 5.4) provides an inter- mediate heat shield between the ambient wall and the hehum vessel. This shield is refrigerated by directing the escaping cold vapor through

coils attached to the shield. The neck tube heat conduction is also inter- cepted at two or more points: at the bottom, the vapor flows up the neck

for a short distance before entering the shield; and again near the top,

the neck is refrigerated by continuing the tube coil from the shield to a physical attachment on the neck to provide heat exchange between the

vapor and the neck. The insulation systems commonly used are: high vacuum alone between the hehum vessel and the shield and evacuated

multilayer insulation between the shield and the ambient shell. In figure 5.5, an alternate method [8] is shown for utilizing the

refrigeration of the vapor. The vapor does not pass through coils attached

Figure 5.5. Multiple vapor-cooled shields as described by Paivanas [8].
162 STORAGE, DISTRIBUTION, AND HANDLING

Figure 5.6. A 100-liter container using

the multiple vapor-cooled shield concept

{see fig. 5.5). The specifications of this container are: Normal evaporation rate= 1.25%/day

Weight, full =250 lbs. Inner and outer shell construction — stainless steel; Dimensions, Diameter= 20 inches; Height = 63 inches; Shock provisions = shock mount base. (Courtesy Linde Division, Union Carbide Corporation).

Figure 5.7. A 100-liter container using

the single vapor-cooled shield concept

{see fig. 5.4). The performance and dimensions are: Normal evaporation rate=< 1.5%/day

Weight, full =258 lbs. Inner and outer shell construction — stainless steel; Dimensions, Diameter= 22 inches; Height = 66 inches; Shock provisions= shock mount base. (Courtesy Cryogenic Engineering Company).
THE VAPORIZATION PROCESS 163

to a shield, but passes directly up the neck. Numerous shields are placed

in thermal contact with the neck. Multilayer insulation is used between

each shield. While each shield is cooled at only one point, the area of neck attachment, temperature gradients along the shield are reduced to acceptable levels by using high thermal conductivity materials (such as copper) in the fabrication of the shield. In this system, heat passing

radially through the insulation is intercepted and conducted along each

shield to the neck tube, where it is absorbed by the escaping cold vapor. Vapor-shielded containers presently have sHghtly higher rates of evaporation than nitrogen-shielded containers, but they have the distinct advantage of eliminating the requirement for liquid nitrogen. Containers

of this general type have been made in sizes from 50-1 to 2000 gal capacity. Some representative illustrations are shown in figures 5.6 and 5.7. 5.3.3. The Vaporization Process

In a constant pressure vaporization process, such as a vented

hehum dewar, the relationship between the mass vaporized and the

heat input is normally given by: mt = j- (5.1)

riv where mt = mass vaporized

^ = heat input

hv = heat of vaporization.

This relationship is commonly used to solve for the heat imput q by

measuring the mass expelled from^ the container — either by gas meter

or container weight loss — and equating this value to /7^^ Several unique

properties of helium may lead to the incorrect application of this simple

relationship. Since the heat of vaporization is relatively small compared

to the specific heat of the vapor, a small degree of superheat in the vapor

leaving the container wiU mask the true heat leak of the container. The

degree of superheat may vary with liquid level. Hence, the rate of gas

evolution from the container will also vary with the content level. Also,

the ratio of the saturated liquid density to the vapor density at 4.2 K

is only 8:1 at 1 atm. Therefore, the vapor which fills the void resulting

from liquid vaporization is not a negligible mass and must be added to an external measurement of vaporization losses,

A correction for vapor superheat may be made by measuring the

temperature of the vapor at the point of exit from the liquid container

and the true enthalpy change substituted for hi^ in eq (5.1). A much

simpler method is to confine loss rate measurements to full containers.
164 STORAGE, DISTRIBUTION, AND HANDLING

The vapor density error can easily be corrected by:

(5.2)

where mt — total mass vaporized

m = mass vented from container

Fjr;= average specific volume of vapor in the container

Vl = specific volume of the saturated liquid.

At 1 atm, assuming zero superheat of the gas, this reduces to: mt= 1.15 m. 5.3.4. Atmospheric Pressure Variations

The apparent heat input calculated from loss rate measurements

is greatly influenced by pressure changes on the evaporating fluid. Be- cause of this, it is sometimes difficult to measure dewar loss characteristics when the container is vented and subject to atmospheric

pressure variations. A correction for this process, for a full container,

may be approximated [9] by equating the heat capacity of the fluid and

the cooling effect of the vaporization loss due to flashing:

VCs /dT\ dP_ dm

where V = container volume

Cs = liquid specific heat at saturation

P = container pressure

t = time

(^^pj ~ slope of the temperature-vapor pressure curve.

This equation may be solved for the rate of vaporization resulting from

small pressure changes:

drtit dp

where

Vi.h, \dPI
THE VAPORIZATION PROCESS 165

If the values for the fluid properties of hehum are used to solve for K,

with convenient units,

c = 7.9 X 10-3 (5.5) where c = vaporization rate due to change in barometric pressure,

scfh

F= total container volume, liters \^) ~ ^^^^^^ barometric pressure, mm/hr

For example, if the flash loss rate from eq (5.5) is compared with the

normal evaporation rate of a full 100-1 vapor-shielded container (about

1.5 percent per day), a change in atmospheric pressure at the rate of

only 1 mm per hour will result in a 50 percent change in the measured

liquid evaporation rate. For this reason, meaningful loss-rate measure- ments can be obtained only if the test is conducted over a long period

of time, to average out the effects of pressure variations. The use of a

precision regulator, controlling absolute pressure on the container, is an alternative.

life

Figure 5.8. 7800-gal liquid helium semi-trailer constructed of 5083-0 aluminum by- Cryogenic Engineering Company.

This trailer, built in 1%2 for nonvented transport, is designed for 100 psig operating pressure. Heat leak is approxi- mately 95 Btu/hr. Insulation is of the multilayer type. A tubular heat exchanger is attached to the inner vessel, for use

during inital cool-down. (Courtesy: Cryogenics Division, NBS Institute for Basic Standards.)
166 STORAGE, DISTRIBUTION, AND HANDLING

5.4. NONVENTED CONTAINERS

Low-pressure, vented containers are the preferred method of stor- ing and shipping hquid hehum when the ukimate use requires a Uquidphase product. Low container pressures result in lower hquid transfer

losses at the destination (see sec. 5.8.6) and the refrigeration available

in the vented vapor may be used effectively to reduce vaporization

Figure 5.9. This 5300-gal semi-trailer, built in 1959, was the first large-scale unit ever constructed for nonvented liquid helium transport.

The inner stainless steel tank is designed for a maximum operating pressure of 180 psig. An additional inner copper

shell reduces thermal stratification of the contents and also is used to direct the vapor generated during filling into close contact with the heavy pressure shell. This latter feature reduces liquid consumed during cooldown by making maximum

use of the gaseous reuseable heat. (Courtesy: Linde Division, Union Carbide Corporation.)

losses during transport. When the ultimate use of the product requires

gaseous helium or both gas and hquid, it is more economical to conserve

the vapor by not allowing the container to vent during the period of

transport, and by allowing the temperature and pressure of the contents

to increase. Containers designed for this service are less expensive than

vented containers. This cost reduction is usually accomplished by

omitting refrigerated shields at the sacrifice of greater heat input. Some

liquid can usually be recovered at the destination (see sec. 5.8) by

venting the container into a gaseous recovery system. The insulation

systems used are the multilayer type in a high vacuum. Figures 5.8, 5.9, and 5.10 show several containers now in use.
PRESSURE RISE 167

40-0

NS. 8 GA. (.164") SA-212 CARBON STEEL

EMPTY WT. 43,350 LBS. LOADED WT. 53.700 LBS.

Figure 5.10. Construction details of 10,000-gal liquid helium semi-trailer. The trailer 'utilizes multilayer insulation to provide a lOOBtu/hr heat leak rate. Copper baffles are used to reduce thermal

stratification and sloshing.

5.4.1. Pressure Rise

The heat added in a nonvented or constant-volume process may be

equated to the change in internal energy of the contents [10]:

Q=^u=^H-v^p (5.6)

where Q = total heat added to the system

f/ = internal energy function

// = enthalpy function = specific volume = container volume divided by total mass of hquid and vapor

AP = the pressure rise at constant volume.

The pressure-internal energy relationship may be conveniently plotted

for various specific volumes as shown in figure 2.7 and described in chapter 2, if the contents are in thermal equihbrium (without temperature gradients). The process can be illustrated by following a line of

306 922 0-68— 12
168 STORAGE, DISTRIBUTION, AND HANDLING

constant volume (isochor) for the horizontal distance corresponding to Af/ or Q. If the vessel is initially partially full, a constant heat input will result in a hnear pressure increase with time until the liquid-vapor dome

is intersected. After that point, the pressure again increases at a nearly constant but much faster rate. It is interesting to note that at specific

volumes less than that corresponding to the critical point, the liquid

miniscus steadily rises until a single phase exists. At specific volumes

greater than the critical, the liquid miniscus lowers until only a single

phase exists.

20

18

16

14

12

10

CD ao

.J

CO

8.6 psig

(PHASE ENVELOPE

INTERSECTION) -

12 16 20 TIME (HOURS)

24 28

Figure 5.11. Pressure rise in a nonvented, 25-liter, nitrogen- shielded helium dewar.

The container initially contained 22V4 liquid liters. Note the sharp inflection in the curve, as the contents became single phase [10].

In the absence of temperature gradients (thermal stratification),

the pressure rise can be accurately predicted and the transition from two

phases to a single phase is sharply defined as shown in figure 5.11.
SUPERCRITICAL VENTING 169

Thermal stratification usually will result in faster rates of pressure rise than a system in thermal equilibrium [10]; also, the phase transition

point will be less clearly defined. These effects can usually be avoided

by the use of construction materials of high thermal conductivity to distribute the energy input uniformly throughout the helium contents. A nonvented transport container would be designed for a specified, no-loss transit time corresponding to an allowable pressure rise and

initial specific volume or filling condition. The container would arrive

at the destination with its contents at an elevated temperature and

pressure, and possibly in the form of a single-phase, cold, dense, gas. Several discharge techniques for liquid gas recovery will be discussed

later.

5.4.2. Supercritical Venting

If a nonvented vessel is delayed in transit, excessive gas losses can

occur when the container pressure reaches the relief valve setting. An

approximate expression [11] for the loss experienced in venting a single

phase container is shown below:

where mf = final mass in container

//ii = initial mass in container

Q = total heat input to expell mass (mj — mf)

P = venting pressure

V = volume of container

In = natural logarithm.

Considering a large vessel such as the semi-trailer shown in figure

5.10, a 30-W energy input would result in a 2 percent/day loss of gaseous

product if the container was vented at 1 atm (see sec. 5.4.1). The same

energy input would result in a 4.6 percent/day loss if the container were

vented at 100 psig. However, if the container was initially shipped 100

percent full of liquid at 1 atm, a pressure of 100 psig would not be

reached for a period of 7 days. During the 7-day period of pressure rise, no product would be lost.

In {mflmi)={l-K)QIKPV (5.7)

K
170 STORAGE, DISTRIBUTION, AND HANDLING

5.5. INSULATIONS

As has been pointed out earlier in this chapter, the very small heat

of vaporization and the low boiling temperature of liquid helium place severe requirements on insulation systems for this fluid. As an example, the heat leak into a high-quality liquid helium container is such that it boils away only about one percent of the total capacity per day. The same heat leak into an equivalent volume of water

would require more than five months to raise the water temperature one

degree Celsius. In addition, liquid helium is a comparatively expensive cryogenic

fluid. Thus, the use of relatively complex and expensive insulation sys- tems is economically justified. In fact, a common and effective method

of insulating liquid helium storage containers involves the sacrifice of a

less expensive cryogenic fluid, usually liquid nitrogen, to absorb part of the heat leaking into the container (sec. 5.3.1). Part of the heat leak into any liquid helium container takes place

through the necessary mechanical support members and plumbing.

These aspects of liquid helium containers are treated elsewhere in this chapter. This section is concerned only with insulation materials and

systems.

5.5.1. Types of Insulation

There are two broad categories of insulations: the evacuated and

non-evacuated. The evacuated insulations can conveniently be subdivided into three major types: simple high-vacuum insulation, evacuated porous insulations, and evacuated multilayer insulations.

The non-evacuated insulations are porous materials such as ex- panded silicate powders or ceramic fibers in which the interstitial spaces

are filled with atmospheric pressure gas. Because of their low installed

cost, they are widely used in less demanding cryogenic insulation systems

such as air liquefaction plants. However, their relatively low thermal

effectiveness, which is due mostly to the high heat leak through the

atmospheric pressure gas phase, generally precludes their use with

liquid helium. Only evacuated insulations are normally used with

liquid helium.

The simple high-vacuum insulation system is composed of a well

evacuated space bounded by highly reflective walls. It is exemplified

in the classical Dewar flask and the Thermos bottle and is illustrated
TYPES OF INSULATION 171

— HIGH VACUUM

COLD

REFLECTIVE SURFACES

(a)

WARM

MODERATE VACUUM

COLD

POROUS SOLIDS (POWDER, FIBER, OR FOAM)

(b)

ffftf

(c)

REFLECTORS

—HIGH VACUUM

FIBEROUS -SPACERS

VACUUM

LAYERS OF CRINKLED OR EMBOSSED,

J METAL COATED, -PLASTIC FILM

Figure 5.12. Three types of evacuated insulation: (a) simple high vacuum insulation; (6) evacuated porous insulation; (c) and (d) two general types of evacuated multilayer insulation.

schematically in figure 5.12a. It is, structurally and conceptually, the

simplest of the evacuated insulation systems. It has the least weight and

least total heat capacity of any of the evacuated insulations, a feature

which minimizes cooldown losses in equipment which undergoes fre- quent thermal cycling (transfer lines, for example). This insulation has

the unique property that the heat leak is independent of the insulation

thickness (at least for nonmicroscopic thicknesses), a characteristic

particularly useful in insulating small equipment. When used between

liquid nitrogen temperature and liquid helium temperature, simple

high-vacuum insulation is normally the most effective insulation avail- able. It is, however, not nearly so effective when used between room

temperature and Hquid helium temperature.

The evacuated multilayer insulations consist of a well evacuated,

laminar assembly of many thin, highly reflective layers which are spaced

apart in such a manner that poor thermal conduction takes place be- tween adjacent layers. Two commonly used evacuated multilayer insulations are shown schematically in figures 5.12c and 5.12d. The insulation
172 STORAGE, DISTRIBUTION, AND HANDLING

shown in figure 5.12c consists of multiple layers of thin metal reflectors separated each from the others by a spacer material. In figure 5.12d is shown an insulation composed of multiple layers which are crinkled or embossed in such manner that adjacent layers touch one another only

at small, discrete, fairly widely spaced points. Between room temperature and liquid helium temperature, evacuated multilayer insulations are the most effective insulations yet developed. They are also the most

expensive and difficult to install properly. In addition, they have rela- tively large total heat capacities which can introduce cooldown problems.

The heat leak through an evacuated multilayer insulation is inversely

proportional to its thickness. The total heat leak can, in principle, be

reduced to any desired extent by using a sufficient thickness of the in- sulation. Thus these insulations can be characterized by an apparent

thermal conductivity (kn) , which is, however, highly anisotropic, being

very small in the direction perpendicular to the layers but being several

orders of magnitude higher in directions parallel to the layers. This

anisotropy causes increased heat leak at any discontinuities in the

insulation, such as penetrations by supports and plumbing.

In spite of their several drawbacks, the unexcelled thermal effec- tiveness of the evacuated multilayer insulations makes them extremely

useful for applications between room and liquid helium temperatures,

especially for storage containers in which auxiliary cooling is not employed.

The evacuated porous insulations, shown schematically in figure

5.12b, consist of finely dispersed solids — powders, fibers, or even foams — similar to those used in nonevacuated insulations, but which are enclosed within impermeable walls so that the interstitial gas can be

evacuated to a moderately low pressure, 0.01 to 0.001 torr. This degree

of vacuum is considerably poorer than that required for the other types

of evacuated insulation, a factor in the comparatively low total cost of

these insulations. In addition, they are simple to install, usually being

poured directly into the evacuatable insulation space. However, they

have roughly the same, relatively large, total heat capacity as do the

evacuated multilayer materials.

The evacuated porous insulations can be characterized by an

apparent thermal conductivity which is isotropic. Their apparent thermal

conductivities are, however, roughly ten times greater than those of the

better multilayer insulations. For this reason, these porous insulations are seldom used as the only insulation in liquid helium vessels. They
SIMPLE HIGH-VACUUM INSULATION 173

are, however, sometimes used to insulate the Hquid nitrogen vessels

incorporated in shielded helium containers. In the following sections, experimental heat transfer data, the

mechanisms of heat transfer, and other properties of the evacuated

insulations will be discussed, the greater emphasis being given to the

simple high-vacuum and the evacuated multilayer insulations. It will become evident that there is a notable absence of published experimental data for these insulations at liquid helium temperature. Almost

all of the data presented, therefore, will be for liquid hydrogen temperatures or higher.

5.5.2. Simple High-Vacuum Insulation

Heat transfer through a simple high-vacuum insulation system takes

place by radiation transfer between the bounding walls and by free molecular conduction through the residual gas in the vacuum space. The radiation and free molecular conduction processes do not interact

with one another. They can, therefore, be treated separately, and the

total heat transfer rate taken as the sum of those evaluated for each

process. Radiation and free molecular gas conduction heat transfer are important in all evacuated insulations, but because they occur in their

least complicated form in the simple high-vacuum insulations, they

will be discussed in some detail here. Radiation Heat Transfer. Any two surfaces separated by a trans- parent medium will interchange thermal energy by a process of emitting

and absorbing electromagnetic radiation. Basic to any discussion of this heat transfer process is the concept of the perfectly black surface, or blackbody, which is one that completely absorbs all of the electro- magnetic radiation that impinges upon it. Such a blackbody must also emit radiation according to the Stephen-Boltzmann relation.

qlA = (TT\ (5.8)

In eq (5.8), qjA is the rate of thermal energy radiated per unit normal

area, T is the absolute temperature of the blackbody, and cr is a fundamental physical constant having the value 5.67 X 10"'- W/cm--K^ (1.712

X 10--' Btu/hr-ft^-R4).
174 STORAGE, DISTRIBUTION, AND HANDLING

10'

10'

10-5

1-6 <

5000

5 10 50 100 WAVE LENGTH, MICRON500 1000

1^'IGURE 5.13. Blackbody spectra for several temperatures.

The area under each curve is a-T^".

Figure 5.14. Band of wavelength of sig- nificance in radiation heat transfer at low temperatures.

The shaded band includes the range of wavelength

that accounts for 90 percent of the radiant energy

emitted by a black or grey body at the indicated temperature. The dashed curve describes the wavelength

of maximum radiation intensity as a function of temperature.

10 50 100 TEMPERATURE, «K 300
SIMPLE HIGH-VACUUM INSULATION 175

From eq (5.8) it follows that the rate of heat transfer between two

extended perfectly black surfaces at temperatures Ti and T-z is

qlA = (T(Ti^Ti). (5.9)

The radiation emitted from a blackbody has a well-defined distribution of wavelengths, termed the blackbody spectrum for a particular

temperature. Such spectra for several temperatures are shown in figure

5.13. For each temperature, these spectra extend over a wide range of wavelengths. As an example, the significant band of wavelengths can

be selected as those which account for 90 percent of the total radiation

intensity. The dashed curve in figure 5.14 shows the wavelength of maximum radiation intensity as a function of temperature, and the two solid curves define the band of wavelengths which accounts for 90 percent

of the radiation intensity. It is important to recognize that these wavelengths are very long. They occupy the intermediate and far infrared

regions of the electromagnetic spectrum and are far longer than the

wavelengths of visible light (0.3 to 0.7 fim). Wavelengths as long as 40

/xm are significant at room temperature, and wavelengths up to 150 fxm

(0.006 in) are significant at hquid nitrogen temperatures. These are dimensions easily measurable with a machinist's micrometer.

Practical surfaces do not emit or absorb radiation as effectively

as would a blackbody; some part of the radiation incident upon practical

surfaces is reflected. The radiation properties of real surfaces are characterized by their emittance, absorbance, and reflectance. The reflec- tance is the fraction of the incident radiation reflected by the surface;

the absorbance is the fraction of the incident radiation absorbed by the

surface; and for opaque materials, the sum of the absorbance and re- flectance must be 1. The emittance is the ratio of the intensity of radiation emitted by a

real, isothermal surface to that which would be emitted by a blackbody

at the same temperature as computed from eq (5.8). Kirchhoff s law re- quires that, for any given wavelength of radiation, the absorbance be

numerically equal to the emittance. Most of the surfaces encountered

in cryogenic insulations are "grey": their absorbance and emittance

do not change much over the range of significant wavelengths. KirchhofTs law can, therefore, be applied with good accuracy to values of absorbance and emittance averaged over the entire range of significant

wavelengths.

In order to compute the rate of radiation heat transfer between two

extended surfaces, eq (5.9) must be modified as follows to take into
176 STORAGE, DISTRIBUTION, AND HANDLING

account the emittances of the two surfaces, ei and €2:

qlA = a

€162

€1 + 62 — €16-^(Ti-n). (5.10)

This equation is exactly correct for semi-infinite, parallel plates. It is also correct for semi-infinite, coaxial cyhnders and concentric spheres

when specular (rather than diffuse) reflection from the surfaces is as- sumed; in these cases qJA is referred to the inner surface area.

If diffuse reflection is assumed, the correct form for spheres and

cylinders is eq (5.11), in which the subscripts 1 and 2 refer to the inner

and outer surfaces, respectively.

€2+ (^1/^2) (ei- €162)

.

(Ti-Ti). (5.11)

Generally, reflections in cryogenic insulations have both diffuse and

specular components. The assumption of diffuse reflection yields the

higher qIA value and is, therefore, the conservative assumption.

In most cases encountered in cryogenic insulations, the factor

A1IA2 is nearly 1, the emittances Ci and €2 are approximately equal, and

both emittances are much less than 1. Under these conditions, both

eqs (5.10) and (5.11) can be approximately simplified to the following

form in which Cavg is the arithmetic average of Ci and €2. ^M=(l/2)o-eavg (Tt-TI). (5.12)

This equation clearly shows that there are two ways to reduce the

rate of radiation heat transfer: by reducing either the differences in the

fourth powers of the surface temperatures or the average emittance of

the surfaces. In most high-vacuum insulation systems, the temperature of the

colder surface is unimportant, for if the temperature of the colder sur- faces is less than 31 percent of the warmer surface temperature, the

fourth power of these temperatures will differ by a factor of more than

100. Radiation heat transfer in cryogenic insulations is, therefore,

approximately proportional to the fourth power of the warm boundary

temperature, a fact which explains the greatly enhanced effectiveness

of simple high-vacuum insulations at low temperatures. In liquid helium

storage vessels, for example, the reduction of the warm surface tempera-
SIMPLE HIGH-VACUUM INSULATION 177

ture from 300 to 77 K, by using either a liquid nitrogen or helium vapor

cooled shield, will reduce the radiation heat transfer by 99.58 percent.

Once the boundary temperatures of the insulation are fixed by the

overall design of the equipment, the lowest practical surface emittances

must be obtained to minimize radiation heat transfer. The lowest emit- tance surfaces are clean, undistorted surfaces of pure metals, especially

the metals having low electrical resistivities. Such surfaces can be

routinely fabricated which have emittances of 0.02 or less, and eq (5.12)

shows that the radiation heat transfer between two such low emittance

surfaces will be only 1 percent of that which would take place between

black surfaces.

There are a number of published compilations of experimentally

measured surface emittance data. Among these are the compilations by

Dickson and Jones [12]; Purdue University [13]; Reynolds, Corruccini,

Fulk, and Burley [14]; Goldsmith, Waterman, and Hirschhorn [15];

and Gubareff, Janssen, and Torborg [16].

Some illustrative values of the emittances of various metal surfaces

have been selected from these published compilations and are shown

in table 5.3. Both monochromatic and total emittances are given, and

no distinction is made between normal and total hemispheric emittances.* In addition to showing values for practical surfaces which can

be produced by standard manufacturing methods, values are also shown

for laboratory-quality surfaces in order to illustrate the very low emit- tance values that can be obtained by the most careful fabricating tech- niques. In fact, the room-temperature emittance shown for ultra-high vacuum deposited silver is the lowest experimental value yet reported

for any room-temperature surface.

The values shown in table 5.3 make one point abundantly clear: emittance is not a true property of the bulk material; rather, it is a property of the surface itself. The emittance values are strongly dependent on the exact manner in which a surface is prepared. Because surface

emittances are dependent on so many parameters, some of which are very difficult to measure, the literature data for emittances are valuable

mostly as qualitative guides for the selection of materials and fabricating

techniques which will yield low emittance surfaces. These qualitative

guides, based both on experimental data and theoretical considerations,

have been formulated by Scott [6] and Dickson and Jones [12].

*Normal emittances are those measured perpendicular to the surface, and total hemispheric emittances are averaged

over the entire hemisphere from which the surface can be viewed. For low emittance surfaces, these two kinds of emit- tances differ by only about 20 percent.
178 STORAGE, DISTRIBUTION, AND HANDLING

Table 5.3. Emittances of selected metal surfaces

A/To f^»-i ol iviaienai Surface temperature K

Radiation

Aluminum

Vacuum deposited at 10"*" torr ~ 300 10/Ltm 0.0124 Vacuum deposited at 10"^ torr freshly deposited ~ 300 10/Ltm 0.0155 20pLm 0.0122 SOfxm 0.0104 aged in air several weeks ~ 300 10/xm 0.0188

20)Lim 0.0148 0.0130 Clean, annealed bulk metal 300 300 K 0.038 76 300 K 0.022 4 300 K 0.014 Household foil 80 300 K 0.043-0.065 H-19 tempered foil 76 300 K 0.056

dull side 76 300 K 0.063 Commercially vapor-deposited on polyester 76 300 K 0.04-0.05 Copper

Vacuum deposited at 10"^ torr 295 10/Am 0.0099 82 lO/xm 0.0076

4.2 > l.S/xm 0.005

4.2 300 K 0.0042 Mechanically polished 4.2 300 K 0.0147 Commercial copper sphere 76 300 K 0.03

GoldVacuum deposited at 10"** torr 300 10/Ltm 0.0061 20/Am 0.0060

30/>tm 0.0058 295 10/xm 0.0101 82 lO/xm 0.0076 0.0015 inch' foil 76 300 K 0.01 Electroplated on stainless steel (1% silver in gold) 0.00005-in thick 76 300 K 0.028

0.0001-in thick 76 300 K 0.027

0.0002-in thick 76 300 K 0.025 Vapor deposited on polyester in plastic film 76 300 K 0.02
SIMPLE HIGH-VACUUM INSULATION

Table 5.3. Emittances of selected metal surfaces — ConXmned

179

Material Surface temperature K

Radiation"

Silver Vacuum deposited at 10"*^ torr

Vacuum deposited at 10~^ torr

Electropolished

Electroplated

Chemically deposited

Alloys 316 Stainless Steel

finished to 2 fx'm rms

finished to 15 fx'm rms

Yellow brass, shim stock Yellow brass, hand polished Yellow brass, electropolished Yellow brass, mechanically polished

Soft solder

300

295

82

4.2 7690

83

83

76

77

4.2 90

76

10/U,m 20/>im

30/>tm lO^tm

10/xm

1.5^tm 300 K

293 K

300 K

273 K

300 K

293 K

300 K

0.0047 0.0044

0.0042

0.0113

0.0093 0.0044 0.007-0.012

0.023

0.027

0.045 0.029

0.10 0.018

0.046

0.03- 0.047

*Entries in "Radiation" column given in micrometers (ju.m) refer to monochromatic

radiation of that wavelength; entries given in K refer to blackbody radiation corresponding

to that temperature.

Most importantly, the visual appearance of a surface is not a reliable indication of its emittance for the intermediate and far infrared radiation involved in cryogenic insulation. For example, organic lacquers

and paints, even though they may be transparent to visible light, are very black to infrared radiation. And a highly polished stainless steel surface looks more reflective than the rather dull, matte surface of chemically deposited silver, but the silver has the lower emittance for thermal radiation.

Emittances generally decrease somewhat with decreasing surface

temperatures. This is evident from some of the data in table 5.3 and is further illustrated by the data in figure 5.15 which show the measured

emittances [17] of several surfaces as a function of temperature.
180 STORAGE, DISTRIBUTION, AND HANDLING

100 200 300 TEMPERATURE, ''K Figure 5.15. Experimentally measured total emittances of several practical surfaces as a function of surface temperature illustrating the significant reduction of emittances at lower temperatures [76]. A is for chemically cleaned copper; B is for aluminum foil; C is for chemically deposited gold; D is for chemically deposited silver. Surface contaminants, especially organic materials such as lubri- cants or even fingerprints, increase the emittance. Rough surfaces

have higher emittances than do smooth surfaces; the characteristic

dimensions of roughness are, however, related to the wavelength of the

radiation. Irregularities small in comparison to the radiation wavelength

cannot be "seen" by the radiation.

Any mechanical deformation (cold working) of the surface, such as occurs in mechanical polishing or burnishing, will increase the emittance.

Dickson and Jones [12] have generalized that emittances decrease in the following order of surface preparation techniques: mechanical

polishing, high-vacuum vapor deposition, electropolishing, and ultra- high-vacuum deposition.

Gas Conduction. Heat can be conducted from one surface to another

by gas molecules. The abihty of the gas to conduct heat is proportional

to the heat capacity of the gas, to its density, to the mean-free-path of the gas molecules (the average distance a molecule travels between

collisions either with other gas molecules or with the surfaces), and to the velocity of the molecules.

In roughly atmospheric pressure gases, the mean-free-path is entirely determined by collisions between gas molecules. As the pressure
SIMPLE HIGH-VACUUM INSULATION 181

is reduced (but not too far), collisions between molecules become less frequent and the mean-free-path increases. This increase is, however,

almost exactly offset by the decrease in the gas density, and the ability of the gas to conduct heat is unchanged. This is the region of viscous

gas conduction.

As the gas pressure is further reduced, the collisions between gas molecules and the surfaces become more frequent than collisions between molecules, and the mean-free-path becomes limited by the

distance between the surfaces. Still further reductions in pressure

cannot increase the mean-free-path, but it does decrease the gas density;

thus, the ability of the gas to conduct heat becomes proportional to the gas density or to its pressure. This is the region of free molecular gas conduction.

Evacuated insulations are always maintained at a pressure suf- ficiently low that free molecular gas conduction prevails. In insulation

systems in which one wall is at liquid helium temperature, only helium

and a small amount of hydrogen can exist as gases; all other materials

are, for practical purposes, completely frozen out. The vapor pressure

of hydrogen at 4.2 K is 7 X 10"'^ torr [18]. The vapor pressure of nitrogen

at 10 K is calculated to be about 1 X IQ-^o torr [19]. Corruccini [20] has presented an illuminating review of free molecular conduction in cryogenic insulations. He shows that the rate of heat

transfer by this mechanism is given by the following relation:

In this equation, qIA is the rate of heat transfer per unit normal area

(watts/cm-), p is the gas pressure in the vacuum space measured by a gage at 300 K (torr), Ti and Tz are the temperatures of the two surfaces

in degrees Kelvin, a is the average accommodation coefficient which is discussed below, and C is a numerical constant determined by the kind

of gas involved. Corruccini gives the. value for C as 0.0280 for helium,

and 0.039 to 0.053 for hydrogen, depending on the temperatures involved.

The average accommodation coefficient, a in eq (5.13), is defined

as follows:

glA = Cap{T,-T,). (5.13)

«1«2 a = (5.14) a2-i-ai{l-a2){AilA2)

Here, ai and a> are the accommodation coefficients for the inner and

outer surfaces, respectively, and AJA-y is the ratio of the surface areas.
182 STORAGE, DISTRIBUTION, AND HANDLING

The accommodation coefficient is a measure of the efficiency of thermal energy interchange that occurs when a gas molecule collides

with the surface. It is less than or equal to 1, and its exact value depends

upon the kind of gas, the temperature of the surface, and very strongly

upon the exact conditions of the surface. Based on the very limited

experimental data for the low-temperature accommodation coefficients on the sort of surfaces commonly used in simple high-vacuum insulations, Corruccini [20] has given suggested values for several gases. Those for helium and hydrogen are listed in table 5.4. He also points out that increased surface roughness and the presence of adsorbed gas

films tend to increase the accommodation coefficient. Table 5.4. Suggested accommodation coefficients

Temperature K

Accommodation coefficient

Helium Hydrogen

300 0.3 0.3 80 0.4 0.5 20 0.6 1 4 1 1

It is important to note that the distance between the two surfaces

does not appear in eq (5.13). In free molecular gas conduction, the

rate of heat transfer is independent of the thickness of the gas layer.

The spacing between the surfaces is important, though, in defining the

pressure at which the transition from viscous to free molecular gas conduction takes place; the greater the spacing, the lower is the transi- tion pressure. For spacing on the order of an inch, the transition pressure

is 10~^ to lO""* torr. For the microscopic spacing of surfaces found in some evacuated porous insulations, however, this transition pressure

is near 1 atm. Equation (5.13) shows also that the heat transfer by free molecular

gas conduction is directly proportional to the gas pressure. In principle,

then, this mode of heat transfer can be reduced to any desired extent by a sufficient lowering of the gas pressure. In practice, however, it is only necessary to reduce the gas pressure to such a level that gas conduction is negligibly small compared to the heat transferred by

other modes, particularly by radiation.
EVACUATED POROUS INSULATIONS 183

For example, consider a simple high-vacuum insulation, the walls

of which have an average emissivity of 0.02. Between 300 and 77 K,

conservatively taking a=l and assuming helium gas, eqs (5.10) and

(5.13) show that the free molecular gas conduction is reduced to 10

percent of the radiation heat transfer at a pressure of about 7 X 10~^

torr. Considering wall temperatures of 77 and 4 K, as would be found

in a liquid nitrogen-shielded helium container, the gas pressure must

be reduced to about 2 X 10~^ torr before gas conduction is reduced to 10 percent of the radiation heat transfer. The techniques for producing, maintaining, and measuring the vacuums required in evacuated insulations are far too broad an area

to be summarized here. Dushman [21] and many other texts offer reviews

of the entire area of vacuum technology, and Scott [6] gives a review

aimed particularly at evacuated insulation technology.

5.5.3. Evacuated Porous Insulations

Radiation heat transfer is a major component of the total heat leak

through evacuated insulations, particularly at temperatures not far below room temperature. We have seen in the preceding section that one effective way of lessening radiation heat transfer is to reduce the

emittances of the two boundary surfaces in simple high-vacuum

insulation.

Another effective method of reducing radiation heat transfer is that of filling the evacuated insulation space with a finely divided solid material such as powders or fibers. These porous materials reduce

radiation heat transfer either by scattering or by absorbing and re- emitting radiation. At the same time, however, some additional heat

transfer by conduction from particle to particle through the insulation

is introduced. (This will hereafter be called solid-phase conduction.)

For not-too-thin layers of evacuated porous insulations, the resist- ance to radiation heat transfer is directly proportional to the total number

of scattering and absorbing particles between the warm and cold boundaries; this is, of course, directly proportional to the thickness of the

insulation layer. Similarly, the resistance to sohd-phase conduction

is proportional to the insulation thickness. Consequently, these insulations, in contrast to the simple high-vacuum insulations, can be characterized by an apparent thermal conductivity, A:«, which for one-dimensional, rectihnear heat flow is defined by the following integral form of the Fourier heat transfer law:

306-922 0-68— 13

qlA = ka{n-Tc)IL. (5.15)
184 STORAGE, DISTRIBUTION, AND HANDLING

The value of ka is not a constant; it depends strongly on the values

of Tw and Tc, the warm and cold boundary temperatures. But, for insulation thicknesses (L) of greater than V2 in or so, ka is practically independent of thickness. Therefore, at least in principle, the total rate of heat transfer through evacuated porous insulations can be reduced to any desired extent by sufficiently increasing the thickness of the insulation layer.

The ka of evacuated porous insulations is, again excepting very thin

layers, independent of the boundary surface emittances. In addition,

it is isotropic; that is, it has the same value in any direction. Thus, the

value of ka determined for particular boundary temperatures can be

properly appUed to other geometric cases (cylinders and spheres, for example) involving similar boundary temperatures by using standard,

integral form, conduction heat transfer equations which are available

in any standard heat transfer text. Some representative values for the conductivities of evacuated

porous insulations are listed in table 5.5 together with the boundary

temperatures associated with each value. There is a notable lack of data for the cold boundary at liquid helium temperature; the value shown

for perhte between 20 and 4 K appears to be the only one yet reported.

From comparing the tabulated ka values of several insulations for 77

and 20 K cold boundary temperatures (and the same warm boundary

temperatures), however, one would expect that ka values for 4 K cold

boundary temperatures should be somewhat smaller than those shown

for the 20 K cold boundary temperature.

The relative merits of evacuated porous insulations can be illus- trated by comparison with a simple high-vacuum insulation having

surface emittances of 0.02. Considering as the first case boundary

temperatures of 300 and 77 K, the simple high-vacuum insulation would

exhibit a heat leak of 4.6 X 10~'* W/cm^. A good evacuated powder having

a ka of 4 /xW/cm-K would require the very reasonable thickness of only

1.94 cm (0.76 in) to attain an equally low heat leak. Upon comparing a liquid-nitrogen-shielded helium container with an unshielded helium container insulated only with an evacuated porous

insulation, however, the picture is completely changed. The heat leak

through the high-vacuum insulation from the nitrogen shield to the

helium vessel is only 2X 10~^ W/cm-. Taking a fairly low ka value of

2 /xW/cm-K for an evacuated powder between 300 and 4 K, eq (5.15)

shows that a 296-cm (116-in or 9.4-ft) thickness of insulation on the

powder-insulated container would be required to equal the performance
EVACUATED POROUS INSULATIONS 185

Table 5.5. Apparent thermal conductivity values for some evacuated porous insulations

Boundary

temperature

Material Density Reference Warm Cold

ka

p.W IxBtu A A. cm-K hr-ft-R

Iblff'

Perlite, —30 mesh 304 76 10 578 8.7 24 304 OA20 6.5 376 8.7 24

to on 2 115 8.7 24

76 4 0.8 46 8.7 24 Expanded oU4 lb 6 350 12.5 23 carbon powder "inn on 3.5 200 12.5 -

Silica aerogel + 300 76 4.75 274 _ 22 50%, 5/Lt dia alu- minum 300 20 3.75 217 — 22

Silica aerogel + 305 90 3.7 214 27 50% fine copper

flake 294 20 3.1 180 28 "AA" Fiberglas 303 77 5.7 330 8 29

isocyanate 300 77 100 5800 30 foam 300 20 69.6 4020 30

of the liquid nitrogen-shielded vessel. Such a thickness is, of course,

impractically large.

Thus, this illustration shows that evacuated porous insulations

are not useful as the sole insulation in liquid helium vessels. Their

use is normally restricted to the less stringent task of insulating the

liquid nitrogen or vapor-cooled shields from room temperature in shielded helium containers.

Because the uses of evacuated porous insulations in liquid helium

handling equipment are so restricted, they will not be discussed in any

further detail here. Instead, the reader is referred to the reviews of Scott [6]; Beck, Kreith, and Kropschot [22]; Kropschot [23]; Fulk [24];
186 STORAGE, DISTRIBUTION, AND HANDLING

Glaser [25]; and to the more theoretical treatment by Larkin and

Churchill [26]. 5.5.4. Evacuated Multilayer Insulations

A very effective method of inhibiting radiation heat transfer between two extended, parallel surfaces has been known for many years:

the use of floating radiation shields. Radiation shields are thin sheets

of reflective material, metal foil for example, which are placed between

and parallel to the two surfaces. Floating radiation shields are those

which are supported in such a tenuous manner in a high vacuum that each can exchange energy with its surroundings only by radiation heat

transfer. The term "floating" refers to the temperatures of the shields,

which are determined entirely by the condition of radiation equilibrium.

In order to illustrate the effectiveness of floating radiation shields,

let us consider a high-vacuum insulation system into which has been

placed a number, n, floating shields and assume that all surfaces have

the same low emittance, Cavg- With the same assumptions involved in eq (5.12), the radiation heat transfer through this system is as follows: gM = l/2eavgO-(rf-r|)/(M+l). (5.16)

A comparison of this equation with the analogous equation for a simple vacuum insulation, eq (5.12), shows that the presence of n floating

radiation shields reduces the radiation heat transfer by the factor

1/(^H- 1). Thus, the introduction of one floating shield reduces radiation

transfer by 50 percent; three, by 75 percent; and nine, by 90 percent.

Needless to say, it is enormously difficult to fabricate any large

number of truly floating radiation shields within a reasonable insulation

thickness. Some systems incorporating only a few shields have been

built, but these have generally been used only for very specialized

apphcations.

The evacuated multilayer insulations, however, offer a practical

method of approximating a large number of floating radiation shields

within a reasonable insulation thickness. This is accomplished by

supporting each thin radiation shield (reflector) more or less uniformly over its surface, each shield being supported by its neighbors, and the

whole array being supported by one or both walls of the insulation space. Obviously, this mechanical support is accompanied by some thermal

contact, and evacuated multilayer insulations are effective only because

such thermal contact can be kept small.
EVACUATED MULTILAYER INSULATIONS 187

Broadly, there are two methods by which thermal contact between

adjacent reflectors is kept small in evacuated multilayer insulations.

One method is to place a spacer sheet made of a low-density, usually

fibrous, low thermal conductivity material between each pair of reflectors. Such has been illustrated in figure 5.12c. The other method

requires that the reflectors be made of a low thermal conductivity

material: polyester plastic film coated with a very thin (few hundred

Angstroms) layer of metal such as aluminum. These reflectors are either

crinkled or embossed so that when they are placed against one another,

they touch only at small, discrete, comparatively widely spaced points.

Such an insulation is shown schematically in figure 5.12d. There are, naturally, many variants of the multilayer insulation

theme, many of them protected by patents, and many different materials

for reflectors and spacers have been tried. Most multilayer insulations use either aluminum foil or aluminum-coated polyester film for reflectors. The most common spacer material is fine glass fiber in the form

of thin mats, cloth, or paper. In addition, synthetic fiber netting, thin

sheets of plastic foam, powders, and embossed plastic sheets are sometimes used as spacers [31]. In experimental insulations, even a multitude

of tiny springs of plastic monofilament has been attached to the reflectors

to act as spacers [32]. Multilayer insulation normally is applied to the inner vessel of a cryogenic container, and subsequently, the outer vacuum shell is assembled around it. Most commonly, the insulation is applied by

spirally wrapping a continuous strip of reflector or reflector-spacer

pair onto the object being insulated. Less common and generally less

effective techniques include the prefabrication of multilayer assemblies

quilted together with needle and thread [33] and the application of

reflective plastic film reflectors in a shingle fashion [34]. Another tech- nique involves supporting a fairly rigid, prefabricated, multilayer

assembly between widely spaced, stand-off supports [35]. Spiral wrapping is most easily accomplished on cylindrical sur- faces. The application to compound-curved surfaces, such as eliptical

tank heads and spheres, usually requires that the layers be cut in a gore pattern and carefully interleaved to prevent thermal shorts between different layers of the insulation. In this respect, the insulations

using only crinkled or embossed plastic film reflectors have the advantage of being more conformable to compound-curved surfaces than are

the aluminum foil-glass fiber insulations. Furthermore, when multilayer

insulations, particularly those using metal foil reflectors, are spirally
188 STORAGE, DISTRIBUTION, AND HANDLING

wrapped on small-diameter objects, there can be significant heat con- duction along the spiral path of aluminum. This can be prevented by

slitting small gaps in the foil at intervals. FinaUy, for reasons we will get to shortly, the tension under which multilayer insulations are wrapped

is critically important to obtaining optimum performance.

From this discussion, it should be clear that there is a considerable

amount of art involved in the proper application of multilayer insulations. Much of this art is proprietary. In addition, it can be seen that the

proper installation of multilayer insulation can be complicated and

time-consuming, an important factor in the comparatively high installed cost of this type of insulation. In a well-evacuated multilayer insulation (vacuums of 10~^ torr or better are required), heat is transferred from shield to shield through

the insulation by two processes, radiation and solid-phase conduction.

These two processes interact strongly with each other. Equation (5.16) offers only a crude approximation for the radiation heat transfer: first, the shield temperatures are not the same as they would be if they were

truly floating radiation shields, and second, the spacer materials in many

insulations further inhibit radiation transfer by the same scattering and

absorption-reemission processes that occur in evacuated porous insulations. There is even less quantitative understanding of solid-phase

conduction in these insulations. It exists and it is significant in the

overall heat transfer through the insulation. In truth, the relative contributions of radiation and solid-phase conduction to the total heat

transfer in this type of insulation have yet to be measured in an exact and unambiguous manner. Although the quantitative understanding of the heat transfer mechanisms in evacuated multilayer insulations is quite limited, they are at least qualitatively understood.

A uniform multilayer insulation has a fixed number of reflectors per unit thickness. The resistance to radiation transfer is proportional

to the total number of reflectors and, hence, to the total insulation thick- ness. The resistance to solid-phase conduction is, by the nature of the process, also directly proportional to the insulation thickness. Therefore, as is the case with evacuated porous insulations, the multilayer

insulations can be properly characterized by an apparent thermal

conductivity, ka, which is independent of the insulation thickness, but is definitely dependent on the warm and cold boundary temperatures.

Furthermore, if the sample is thick enough to include a large number of

reflectors (50 or more, say) the ka value is practically independent of the

emittance of the boundaries. In contrast to the evacuated porous insula-
EVACUATED MULTILAYER INSULATIONS 189

tions, though, the ka values for evacuated multilayer insulations are extremely anisotropic; for some insulations using metal foil reflectors,

the value oikn in the direction parallel to the reflectors has been reported

to be W times greater than the kn in the direction perpendicular to the

reflectors [36]. The total heat transfer rate through a multilayer insulation is markedly increased by compressive forces resulting either from compression between the bounding surfaces of the insulation space or from

the tension applied during the wrapping of the insulation. Most of the

increase presumably takes place in sohd-phase conduction. According

to the best available theories, the increase in heat leak should be pro- portional to between the 1/2 and 2/3 power of the applied compressive

load. This has been strikingly confirmed experimentally for many

insulations by Black and Glaser [31] for compressive loads ranging

from 15 psi down to 0.01 psi, which is little more load than that produced

by the weight of the insulation itself. Although compressive loading of any given evacuated multilayer

insulation sample increases the total heat leak, this same compressive

loading also decreases the thickness of the sample. For small compressive loads, the sample thickness decreases more rapidly than the total heat leak increases. This results in a net reduction in ka. Beyond some

optimum compressive load, the situation reverses, and ka begins increasing with increased compression. Curves illustrating this effect for several

multilayer insulations are shown by Glaser, Black, and Doherty [36]. Rather than describing an optimum compressive load for a particular

multilayer insulation, however, this optimum is usually specified by an optimum insulation density expressed in number of reflectors per unit

insulation thickness (reflectors per inch). For maximum insulation

effectiveness, then, the insulation density must be very carefully controlled during installation. The published ka values for several representative evacuated multilayer insulations are given in table 5.6. As was found for evacuated

porous insulations, there is a notable lack of published data for helium

temperature cold boundaries [37-39]. Noting the general decrease in ka as the cold boundary temperature is reduced from 77 to 20 K, it may

be surmized that the ka for boundary temperatures of 300 and 4 K

should be somewhat less than those shown for 300 to 20 K.

In discussing the evacuated porous insulation, it was calculated

that an impractically large thickness of insulation would be required

to equal the insulating performance of a nitrogen-shielded liquid helium
190 STORAGE, DISTRIBUTION, AND HANDLING

Table 5.6. Apparent thermal conductivities of some evacuated multilayer insulations

Material Layersper

inch

Boundary

temperature

k„ Density Reference Warm Cold

fxBtu A A

cm-K hr-ft-R

Ib.jft^

l-mil Al foil —0 9 mm nnnHf^H 66 iO U. / 4U 7.5 37

^ICloo crlpcc rini=*r mpt IIICIL* • 66 300 20 0.5 28 7.5 37

l-mil Al foil -8- m1 1 0"lii 50 300 76 0.5 28 8.7 37 300 20 0.4 23 8 7

Al foil — glass

fiber 50-100 300 76 0.38 22 5.5 38 paper 50-100 300 20 0.36 21 5.5 39

(proprietary) 75-150 300 20 0.17 10 7.5 28

V4-mil, crinkled aluminum coated

polyester film 106 289 Ti 0.42 24 32 no spacer

(proprietary)

container. If one makes the same calculations for a container insulated

only with a very good evacuated multilayer insulation having a ka of

0.2 /xW/cm-K, one finds that only a 29.5-cm (11.6-in) thickness of insulation between 300 and 4 K is required to equal the performance of the

liquid-nitrogen-shielded container.

As was stated earlier, the residual gas pressure within an evacuated

multilayer insulation must be reduced to lO""^ to 10~^ torr, a somewhat

lower pressure than is required by evacuated porous insulations. The

proper evacuation of multilayer insulations can be quite time-consuming.

Some materials, particularly plastic films and organic binders in spacer

materials, often contain sizable quantities of absorbed gases which are only slowly liberated. And since either metal foil or plastic film reflectors
EVACUATED MULTILAYER INSULATIONS 191

are impervious to gas flow, the gases being pumped away must flow along

the narrow, obstructed channels between the reflectors until they can

escape at an edge or other discontinuity in the insulation. In cases where a high evacuation rate is required in a multilayer

insulation, the impediment to gas flow can be significantly reduced by

perforating the reflectors so that gas can flow outward through the thick- ness of the insulation [36, 40]. Naturally, the perforations decrease

the eff"ectiveness of the reflectors; so the size, number, and spacing

of the perforations must be judiciously selected so as to attain adequate pumping speed without too seriously reducing the insulating

effectiveness.

One final problem in properly applying evacuated multilayer insulations will be mentioned: the effects of mechanical penetrations through

the insulation layer. Penetrations by mechanical supports and fluid conduits can not be avoided, but their presence can cause a deterioration in the performance of these highly anisotropic insulations. This

results because the comparatively large amount of heat being con- ducted along the penetrating member can be transferred to the radiation

shields, and this heat can then be conducted along each shield for a considerable distance around the penetration. The result is the increase

of the shield temperatures above their equilibrium value and a net increase in heat transfer. This effect can be lessened by thermally

decoupling the insulation from the penetration. One method of accomplishing this is to leave a gap between the insulation and the penetration which is then filled with an isotropic insulation such as glass fiber. Of course, if the penetration is a member which can absorb heat,

a cold-vapor vent line for example, the penetration can actually improve

the performance of the insulation by reducing the shield temperatures

below their equilibrium values. In these cases, good thermal coupling

rather than decoupHng is desirable. A multilayer insulation system in which this situation is exploited is described in section 5.3.2. Here much emphasis, perhaps too much, has been given to the

problems that may be encountered in properly applying evacuated

multilayer insulations. These problems are, however, quite real, and

they should be anticipated in designing and selecting insulations for liquid helium containers. In spite of the difficulties and expense of using

evacuated multilayer insulations, the fact remains that they are the only

insulations effective enough to make practical the insulation of liquid

helium storage vessels without recourse to auxiliary cooling in the form

of liquid nitrogen or helium vapor-cooled shields.
192 STORAGE. DISTRIBUTION, AND HANDLING

5.6. SLOSHING AND THERMAL OSCILLATION

HEATING EFFECTS

The commercial distribution of liquid helium in large capacity

semitrailers, described in section 5.9, demands a relatively sophisticated

cryogenic design. While multilayer insulation techniques have played

a very important role in reducing heat leak, other effects must be

considered. Thermal stratification, for example, produces excessive

pressure rise for a given heat leak. The energy of sloshing in the vessel

and thermal oscillations in the interconnecting piping also contribute to the overall heating effects.

5.6.1. Sloshing

Accelerations of the trailer, associated with speed or direction

changes and road disturbances, will produce sloshing of the liquid

within the vessel. Energy thus gained by the liquid will eventually be

converted to heat through friction. This heat will cause product vaporization or pressure rise additive to that caused by heat leak. The energy

of the sloshing liquid, due to given input accelerations, is therefore a

significant design consideration of a liquid-helium vessel. The literature on energy of sloshing is very scant, particularly that applicable to a horizontal cylindrical vessel. Arthur D. Little, Inc. [41] found that basic relationships for surface waves developed by Lamb [42] could be used if the cylindrical vessel was approximated by an equivalent

rectangular tank. From this basis, it can be shown [43] that vertical accelerations (road disturbances) will produce no sloshing. Rough

estimations also show that sloshing energy due to rotational accelerations

is small compared to that of longitudinal or latitudinal accelerations.

Equation (5.17), developed by Arthur D. Little, Inc., gives the

specific sloshing energy input to the liquid for sloshing in the direction

of a given acceleration. Note that a deceleration has the same effect as an acceleration and that starting and stopping of a trailer have additive effects:
THERMAL OSCILLATIONS 193

where dq = energy input in Btu/lb a = acceleration in g's

L = total length of tank in feet A = height of the Uquid in feet n = number of full transverse baffles J = work-energy equivalent, 778 ft-lb/Btu.

An order of magnitude approximation of the energy input for a helium trailer can be obtained as follows [43]. A typical cryogenic trailer on a 30-hr trip receives approximately 1600 accelerations and decelerations, averaging 0.04 from starting, stopping, and routine speed

changes. For the heHum trailer described in section 5.9 with eight full transverse baffles, the total energy gained for such a trip is about 15 Btu

for a 90 percent full load. While this heat gain is significant, it is roughly

equivalent to V200 of the normal heat leak while stationary. This estimate

is only for longitudinal sloshing, the type with the greatest effect. The

sloshing energy for the same trailer without baffles would increase by a

factor of 80, an indication that baffles are necessary. Sloshing may be

eliminated by transporting single-phase fluid, which eliminates the

Hquid-vapor interface. In practice this is normally the case, since

nonvented transport containers are shipped liquid full. The singlephase state will occur during transit in partially full vessels as the

rising pressure approaches the critical point of 2.26 atm. Sloshing is observed with other cryogenic fluids in trailer transport,

but a quantitative correlation of the effects has not been made. The mixing effect of sloshing is apparent in transport vessels subject to temperature stratification (see sec. 5.4.1); the mixing serves as a temperature

equahzing mechanism which causes the pressure to drop. The handling

of trailers, as reported by various drivers, is adversely affected by the

sloshing motion of the liquid.

5.6.2. Thermal Oscillations

Thermal oscillations may occur spontaneously in tubes connecting a

liquid hehum reservoir to ambient temperatures. The existence of this phenomenon is quite well known (see. for example Clement [44], Scott [6], and Wexler [5]) and is commonly used to detect liquid levels in helium

dewars. Thermal oscillations, or pressure oscillations, as they are often

called, in a storage or transport container cause a heat pumping effect which may increase the heat leak into a heHum reservoir up to 1,000

times that of normal conduction in a tube.
194 STORAGE, DISTRIBUTION, AND HANDLING

A recent paper by Bannister [45] discusses pressure oscillations

in some detail, presenting quantitative as well as qualitative results. Physically, the oscillation is a traveling wave phenomenon which trans- ports energy from the warm end of the tube to the cold end and which

ejects discrete pulses of warmed vapor from the cold end. Several results are mentioned which may be helpful in liquid helium equipment design.

The frequency of oscillations has been computed by Norton [46] for certain tube configurations. Bannister correlated the pressure amplitude

as a function of the slenderness ratio (length to diameter) shown in figure 5.16. The significant result is the heat pumping rate shown

in figure 5.17 as a function of oscillation intensity (amplitude times

frequency).

The oscillation amplitude is maximum when the cold end of the

tube is 5 to 15 cm above the liquid surface. Below the liquid surface,

both the amplitude and the frequency are reduced. Bannister noticed

that the oscillation intensity could be increased by heating the tube

along a portion of its length and, thus, could be reduced by cooling a portion of its length. This is currently done in dewar design by thermally

shorting the vent lines to the liquid nitrogen or helium vapor refrigeration

0 100 200 300 400 500 SLENDERNESS RATIO - L/0

Figure 5.16. Thermal oscillation— pressure amplitude vs. slenderness ratio for tube exposed to a liquid helium reservoir [45].
THERMAL OSCILLATIONS 195

HEAT PUMPING RATE - WATTS

Figure 5.17. Thermal oscillation — heat pumping rate vs. oscillation intensity (amplitude

times frequency) for tube exposed to a liquid helium reservoir [45].

shield. It was also noted that the increased oscillation intensity served to depress the temperature profile of a tube at the cold end of the tube.

Finally, the heat pumping is significant only when the cold end is in the

vapor phase. This seems reasonable, since more energy is required to pump hquid up and down a tube than to move cold vapor. Experience with various types of liquid helium dewars indicates

that oscillations are not self-sustaining in a freely venting tube. For this reason, oscillations are not usually a problem in a properly venting

dewar. It is apparent that restricting the flow, particularly in the cold

end, with foam, sintered metal, fibers, beads, etc., will reduce the effects of thermal oscillations, but this also restricts the flow. Another technique

is to place expanded sections of tubing or accumulators in the connecting

piping to dampen the oscillation energy in the cold regions.
196 STORAGE, DISTRIBUTION, AND HANDLING

5.7. CONTAINER DESIGN

5.7.1. Materials of Construction

There is little need to provide details on materials and construction

for cryogenic equipment. This is covered quite well in the literature,

e.g., Bell [47] and Kropschot [48]. Usually, materials suitable for cryo- genic service with other low-temperature fluids are also appropriate for liquid helium storage and transport vessels. Some special distinctions

exist, however, between 4.2 K service and 77-90 K service, and rather

than attempting to discuss the entire area of cryogenic materials selection, the special requirements for 4.2 K service are emphasized.

Generally, copper, aluminum, nickel, their respective alloys, and

austenitic stainless steels may be used for vessel construction. Support

members may be metallic (stainless and high-strength alloys, such as Hastalloy B or A-286 stainless) or nonmetallic (fluorocarbon polymers,

polyester fiber, etc.). Data compilations for these and other materials

are given by McChntock and Gibbons [49] and by Schwartzberg, Good,

Keys, and Kieter [50] among other sources. Actually, there has been

very little mechanical testing done at helium temperatures, and most

published curves are extrapolated from 77 K or 20 K down to 4.2 K. A

comparison of common cryogenic materials suitable for helium service

is given in table 5.7 as the ratio of yield stress to thermal conductivity.

Copper has special application for small helium dewars because

of its ductility for spinning and its low surface emissivity in the polished

condition. Electrolytic tough pitch copper is used for anti-slosh (sec.

5.6.1) and de-stratification (sec. 5.4.1) baffles because of its high thermal

conductivity. Aluminum has not been very widely used for helium equipment. Since stainless steel is usually used for interconnecting piping

(low conductivity), aluminum to stainless transition joints are required.

These have traditionally been avoided because of their susceptibility

to vacuum leaks; however, more reliable joints are becoming available.

Aluminum does have an economic and weight advantage and there

exists a successful 7,000-gal helium trailer designed for air transport,

manufactured by the Cryogenic Engineering Company (fig. 5.8) using a 5083-O aluminum alloy.

Type 304 and 304L austenitic stainless is commonly used. Type 304

has better properties in general, but 304L has better impact properties

in the weld heat affected zone. This is attributed to its lower carbon

content. The common 9 percent nickel steel, while much used for liquid
MATERIALS OF CONSTRUCTION 197

oxygen and nitrogen service, has very poor impact properties below 77 K,

especially in the heat affected zone. Carbon steel, such as SA-212, is often used for outer shells which are not subjected to low temperatures.

Cryogenic transport vessels, regardless of operating pressure, are considered as pressure vessels, and therefore must satisfy pertinent state and local codes where they apply. Most of these codes incorporate the

ASME Pressure Vessel Code. Vessels subject to interstate usage, as is most transport equipment, come under the jurisdiction of the Interstate Commerce Commission. There are presently no general regulations for cryogenic transport, and, although ICC specification MC331 generally

applies, special permits must be obtained for hazardous fluids and for operation above 25 psig. Liquid hydrogen, for example, fits into the first category and the liquid helium trailers described in section 5.9.1 are in the second. Apart from general safety and transport restrictions, the

Table 5.7. Comparison of materials used for support members*

Material ay k a-ylk

{1,000 psi) (BtulhrlfilF) Aluminum 2024 55 47 1.17 Aluminum 7075 70 50 1.4 Copper, ann 12 274 0.044

Hastelloy=^ "B" 65 5.4 12 HastelW "C" 48 5.9 8.1 "K" Monel" 100 9.9 10.1 Stainless Steel 304, ann 35 5.9 5.9 Stainless Steel (drawn 210,000 psi) 150 5.2 29

85 21 4.0 Titanium alloy (4Al-4Mn) 145 3.5 41.4 Dacron ' 20 t0.088 227 Mylar ^ 10 t0.088 113 Nylon ^ 20 0.18 111 Teflon^ 2 0.14 14.3

*cry is the yield stress and k is the average thermal conductivity betvi^een 20 and 300 K. tRoom-temperature value.

^ Haynes Stellite Company.

^ International Nickel Company.

' E.I. du Pont.
198 STORAGE, DISTRIBUTION, AND HANDLING

ICC requires vessel design according to Section VIII of the ASME

Boiler and Pressure Vessel Code. To satisfy the Code for vessels operating below —325 F, notch impact tests of vs^eld samples must be performed

at the operating temperature, which is —450 F for liquid helium. The

value of impact tests at this temperature is questionable. With current techniques, the sample must either be cooled on the anvil or transported

from a helium bath to the anvil during the 5-s period allowed by the Code.

Since the specific heat is extremely low in the range of 4 to 20 K, and the

temperature difference so great (300 K), it is probable that the sample

would warm considerably, possibly to 30 or 40 K. The Compressed Gas Association represents another endeavor to regulate uniformly the design of cryogenic transport equipment. This is an effort of the industrial gas industry to provide for the safe transport

of compressed and liquefied gases. Two pertinent documents published

by the CGA are Pamphlet S-1.2, Safety Relief Device Standards for Cargo and Portable Tanks for Compressed Gases, 1966 edition [51]; and

the "Proposed Insulated Cargo Tank Specification MC—341 for Low

Temperature Liquefied Gases" [52]. As of this writing, the latter document is not final, but wiU be proposed to ICC for inclusion in the appropriate regulations.

5.7.2. Vacuum Retention

Vacuum insulation techniques and their application to helium dewar

and equipment design are discussed generally in sections 5.3 and 5.5. Because of the cryopumping in the insulation space, about 11,000

1/s/ft- for air, the initial or warm level is not as important as the vacuum

integrity, and since helium is the only gas that will not cryopump,

the leak rate between the helium vessel and the vacuum space has the

greatest effect. Most liquid helium equipment is fabricated under the

scrutiny of mass spectrometer leak testing of both the individual components and the final piece of equipment. A purchase specification of a

1,000-1 helium dewar, for example, might require mass spectrometer

leak testing with no detectable leak at a sensitivity of 10~^ atm-cm^/s.

Such a design specification would vary with the size of the vessel and the

amount of re-evacuation that could be tolerated. An activated charcoal

or molecular sieve adsorbent attached to the helium vessel will adsorb

helium and hydrogen outgassing from the metal walls and the insulation

material.
INTERNAL SUPPORTS 199

There are several types of vacuum sealoff devices commonly used

for liquid helium equipment. The leaded pinchoff tube is most used for small dewars where the need for re-evacuation is infrequent. Larger

vessels and transfer lines usually include a high-vacuum valve, and

many of these have locking devices to prevent accidental opening. The

Richards [53] sealoff valve, often used, is equipped with a removable

bonnet. Burst disks or vacuum-rehef devices are included on larger transfer

Unes and hehum vessels to prevent over-pressurization of the vacuum

space. The rehef capacity of the device should handle the maximum

boiloff from a ruptured inner vessel. A combination of a Richards-type

sealoff and vacuum-rehef valve is available from several manufacturers.

Current vendor literature on these and similar devices is often very

helpful in making final component selections.

Liquid hehum transfer hues may require re-evacuation several

times a year. A poor vacuum will show up during use as a uniformly cold,

wet, or frosted transfer hne. Individual cold spots, particularly at elbows

and fittings, indicate poor construction or failure of internal spacers. A well designed and constructed transfer line will always feel warm to the touch and any signs of moisture or frost will indicate excessive heat

leak. A helium transfer hne containing air may be re-evacuated to below

1 X 10~^ torr in the warm condition. It wiU then cryopump to below

1 X 10~^ torr when used in helium service.

5.7.3. Internal Supports

Low heat leak is a prime design criterion for liquid helium vessels. Generally, the more rugged the support, the greater the heat leak. Therefore, the design of the support system requires a careful balance

between low heat leak and adequate strength. Consequently, low-loss

helium containers cannot be made invulnerable to damage resulting

from rough, improper handhng. Repair charges are high, up to 30

percent of initial cost for nitrogen-shielded dewars and up to 50 percent

of cost for the vapor-shielded type.

The first liquid helium dewars of nitrogen-shielded design were

intended for use within the laboratory and were handled rather carefully.

Centralized production facilities have since created the need for longdistance trucking and air shipment of helium. Common carrier shippers

and commercial truckers, however, are unpredictable and often unreliable in their treatment of fragile equipment.

306-922 0-68— 14
200 STORAGE, DISTRIBUTION, AND HANDLING

Figure 5.18. Impact-O-Graph recording g-meter, three-directional, 0 to 15 g range.

(Courtesy Impact-O-Graph Corp.) While firm g-load design limits are still a gray area, one guideline

has been established by the Air Force. Military Specification MIL-A8421B (USAF) 5 May 1960 provides the following design criteria for air transportable equipment. For flight and taxing loads, a structural design

suitable for 3 gs in any direction is required; and for emergency landing

conditions, equipment must withstand 4V2 gs in the vertical direction

and 8 gs longitudinally for 0.1 s without external breakage. It is the

authors' opinion that this design limit should be met for vacuum retention to prevent release of cold vapor in a crash condition.

A preliminary field study was performed by Air Reduction Company [54] to determine the actual ^-loads encountered during various

phases of transportation of liquid helium dewars. A recording ^-meter

(fig. 5.18) was used to record the number and magnitude of shocks in three directions. Results show that a dewar, secured in place during

trucking, receives very small shocks — less than 2 ^s in the vertical direction and less than 1 g in either the lateral or longitudinal direction.

Miscellaneous handling, however, which includes loading and unloading
INTERNAL SUPPORTS 201

operations, contributes many shocks ranging from 1 to 3 ^s in all directions, with higher single shock up to 8 gs in any direction. As a result of this study, the following g-\oad specification was established by Airco

for all Hquid heHum dewars up to 500-1 in size:

g Load Design Limit

Static loads (sustained) Short duration loads (about 0.1 s)

Vertical Lateral

5 6 25 15

Figure 5.19. 500-liter vapor-shielded liquid helium dewar {Airco Cryogenics Division).

Note the rubber shock mount system and the vacuum jacketed liquid withdrawal valve.
202 STORAGE, DISTRIBUTION, AND HANDLING

There are some techniques used by various manufacturers to reduce

the impacts of handhng. Rubber ring cushions in the base will absorb a

portion of the shock before it reaches the vessel. Sponge rubber rings

around the circumference of a vessel protect it from lateral bouncing.

The rubber shock mount system shown on the dewar in figure 5.19

reduces the impact of short duration shocks by a factor of 2V2 to 5. One

manufacturer (Airco Cryogenics Div.) uses the Impact-O-Graph meter

to monitor the handling of dewars as part of the warranty conditions.

The appearance of a shipping package is another facet. The initial trend of transporting helium dewars in a rugged crate seems to be

diminishing. Many suppliers have the opinion that the crates create an impression of ruggedness, which is not warranted. "Careful handling"

and "fragile" warning signs are also useful in reducing structural failures. Trailers, on the other hand, are not subjected to the unpredictable

handling received by small dewars. This is because they are in constant

contact with the road and attached to a limited acceleration vehicle.

Sufficient experience has been gained in trailer transport of other

cryogenic fluids to provide adequate design data. The Compressed Gas

Association (1966) makes the following recommendation on ^-load

design as an addition to existing ICC Regulations.

5.7.4 Neck Plug Hazards

Helium, hydrogen, and neon are the three air-solidifying cryogenic

fluids. The boiling points of these fluids — 4.2, 20.4, and 27.2 K, respec- tively—are all below the solid point of air, 62 K, and the nitrogen-oxygen

system eutectic, 50 K. Air that is drawn into a dewar vent hne or neck

tube wiU condense and freeze, forming a solid plug. This will prevent

further venting and cause pressure to build to the burst point of the container. Helium vessels, being the coldest, are most susceptible to this hazard, and because of the increased heat leak in the nonvented condition, the pressure will rise more rapidly. The potential hazard of neck

Minimum* Static g Loadings

Vertical downward.

Vertical upward....

Longitudinal

Lateral

2.0 LO

LO

0.5 *Based on V4 ultimate strength
NECK PLUG HAZARDS 203

plugging is increased with normal barometric pressure fluctuations,

air transport (see sec. 5.7.5), and by opening the dewar to the atmosphere

for product transfer. During 1961 and 1962, several malfunctions of liquid helium containers caused a temporary embargo to be placed by the airlines on the

air shipment of this cargo. In 1963, dual-neck, nitrogen-shielded helium

dewars were introduced and readily accepted by the industry. Work

done by L. Scott [7] on pressure rise rates and burst pressures strongly

supported the use of dual-neck dewars. As of this writing, all major

producers of liquid helium distribute it in vessels equipped with some

type of dual rehef passage. Experience has shown that double-neck or dual relief passage design

has effectively reduced the hazards of neck plugging. If a plug forms

in the normal vent line or fill connection, a second relief passage with a higher relief valve setting wiU prevent excessive pressure buildup. This

is achieved in most dewars with a double concentric neck design (figs. 5.20 and 5.21); the inner tube serves as the access port and normal vent

Figure 5.20. Dual neck and relief valve arrangement for a typical nitrogen- shielded helium dewar.
204 STORAGE, DISTRIBUTION. AND HANDLING

Figure 5.21. Dual vent and relief valve arrangement for a typical vapor-shielded dewar. and the outer annulus acts as an emergency relief passage. Loss rate measurements for nitrogen-shielded dewars indicate a 5 to 15 percent

increase in heat leak for a dual-neck dewar, compared to a single-neck

dewar, with slightly lower loss rates when venting from the inner tube

(L. Scott, [7]). The cold vapor venting up the neck tube keeps the heat

leak increase to a minimum (see sec. 5.3.1). Vapor-shielded dewars generally have an independent secondary

vent, rather than a double-neck design, to achieve the dual relief capacity.

There are actually three relief passages for a typical vapor-shielded

helium dewar (figs. 5.21 and 5.22). The shield vent, with a relief valve

set for the desired operating pressure (0.5 to 3.5 psig), is dependent on

the neck vent or access tube, as they are interconnected a short distance

up the neck from the inner vessel. The secondary relief passage runs

directly from the inner vessel to an external relief valve set to relieve

below the design pressure of the vessel. Other relief valves would have

intermediate pressure settings.

A technique for reducing the possibility of a neck plug occurring is described by Neary [55]. A closed-system operation of the dewar elimi-
NECK PLUG HAZARDS 205

Figure 5.23. Ball valve transfer tube access port for helium deivar.

(Linde Divisinn. Union Carbide Corporation.)
206 STORAGE, DISTRIBUTION, AND HANDLING

nates any portion being opened to the atmosphere. The neck tube access, terminating in a ball valve and O-ring adapter, as shown in figures 5.23

and 5.24, will receive a transfer tube while maintaining a positive pres- sure on the vessel. In this case, since the vessel need not be blown down

Figure 5.24. Top view of a 100-liter multi-shielded helium dewar showing ball valve access port, relief valves, and accessories.

(Linde Division, Union Carbide Corporation.)

to atmospheric pressure prior to a transfer, the operating pressure should

be maintained at the desired transfer pressure, 2 to 3 psig. The open type

dewar is maintained at low pressures, 0.5 to 1.0 psig, to reduce flash losses of depressurization, which are treated in more detail in section

5.8.
ABSOLUTE PRESSURE RELIEF VALVES 207 A permanently installed liquid line on a dewar will eliminate the

need for access to the inner vessel. The line must terminate with a vacu- um-insulated valve and appropriate bayonet connection and will serve

for both filling and withdrawal. A vessel using this construction is shown

in figure 5.19. Neck plugs do occur in spite of careful handhng during normal

dewar usage, and they must be removed. A neck plug may be noticed

by lack of venting when the pressure gage shows a higher reading than

the relief valve setting, or by finding an obstruction on insertion of a

transfer tube. The standard practice for removing neck plugs is to insert a copper or brass rod or tube down the neck. This source of heat will easily melt the ice plug. Care must be taken in this technique to avoid

burns from the escaping cold helium vapor. Gloves and safety glasses

are appropriate. A plug saw is available which is inserted into the neck

and rotated. The saw drills a hole in the plug to relieve the pressure and

then removes the plug in a manner similar to a typical hardware hole saw. One other technique is to blow warm helium gas down the neck

or vent line which may melt the plug. This was tried in one case by the

author on a small curved vent line (one of several vent lines on the dewar)

without success. In this case, the entire dewar was warmed to ambient

temperature to remove the plug.

5.7.5. Air Transport and Absolute Pressure Relief Valves

The variations in ambient pressure during air transport have particular importance as they affect shipments of liquid helium. During

flight, the aircraft cabin pressure decreases to values ranging from 12.2

to 8.3 psia, which corresponds to equivalent altitudes of 5,000 to 15,000

ft, respectively (fig. 5.25). This reduction in external pressure imposed

on a helium dewar causes depressurization of its contents and resulting

very high flash rates. A typical air shipment of liquid helium would lose 15 to 20 percent of its contents (fig. 5.26) if allowed to vent to the reduced

pressure. A subsequent increase in cabin pressure occurs as the aircraft descends, creating an external pressure which is higher than the dewar

pressure. This back pressure causes an influx of air into the neck tube

or vent lines through valves unintentionally left open or through relief valves which may not reseat to reverse flow. A neck plug may occur or

the dewar may arrive at its destination under partial vacuum. The
208 STORAGE, DISTRIBUTION, AND HANDLING

economic loss of high flash rates is coupled with the hazard of neck

plugging and venting in a closed space. The potential of a neck plug occurring can be reduced by venting

through either an absolute pressure relief valve or a differential relief valve preceded by a heat exchanger. An absolute pressure relief valve

0 I J 1 1 \ \ 1 0 10,000 20,000 30,000 40.000 50.000 ALTITUDE iN FEET

Figure 5.25. Ambient pressure vs. altitude, showing range of cabin pressures for commercial aircraft in cargo service.

(fig. 5.27) controls the absolute pressure within the dewar by sensing the

pressure with respect to an evacuated bellows within the relief valve. The net effect of the absolute pressure relief valve is to maintain a constant venting rate, proportional to the heat leak. This prevents a decrease

of pressure in the vessel below that of the absolute pressure setting,

which, for this service, ranges from 15.2 to 18.5 psia.
ABSOLUTE PRESSURE RELIEF VALVES 209

24

6 8 10 12 14 16 Po-AMBIENT PRESSURE IN PSIA Figure 5.26. Helium dewar depressurization losses, loss in percent liquid contents vs. po when vessel is depressurized from 15.2 psia to reduced ambient

pressure po.

INLET portFigure .5.27. Absolute pressure relief valve maintains constant absolute pressure ofventing

vessel.

(Courtesy of Tavco, Inc.)
210 STORAGE, DISTRIBUTION, AND HANDLING

The external heat exchanger does not control the vessel pressure, but only serves to preheat the venting vapor prior to the relief valve. The heat exchanger has sufficient surface to prevent freezing of the valve

for large evaporation rates encountered in air shipments. It is assumed

that a differential pressure relief valve maintained at ambient temperature will always reseat and seal tight in the reverse direction; even if this is not true, the heat exchanger is of sufficient length and volume

so that small quantities of air that may leak through the valve will not reach the cold portions of the neck or vent lines. The airlines, through their regulatory organizations, are in accord

with either of these solutions for preventing general malfunction of a dewar in air transport. The International Air Transport Association and

the Air Traffic Conference of America make the following statement in the latest I ATA Regulations [56]: "Changes in pressure and temperature, resulting from altitude variations, can

cause malfunction of safety relief devices and plugging of vent lines unless

the relief valves are specifically designed for these conditions. Failure of a liquid compartment to vent excess pressure can be prevented by placing safety relief devices to the vapor space on lines connected separately to the hquid compart- ment, by use of an absolute pressure safety relief device, or by use of a differential pressure relief device preceded by a heat exchanger."

Large quantities of helium escaping into the aircraft cabin reduce

the oxygen concentration, and because of the high velocity of sound in helium, it also changes the pitch of the human voice. This problem is more serious in larger helium shipments. For example: A 500-1 dewar,

not equipped with a functioning absolute pressure relief valve, will vent more than 2500 scf of gaseous product into the aircraft cabin in a typical

air shipment. A pitch of voice increase is easily noticed with helium concentrations as low as 5 percent [57], but the change in voice pitch does

not cause severe disruption of communication until helium concentrations above 50 to 60 percent are reached [58]. While communication intelligibility should not be a problem, the

increase in voice pitch does serve as a warning of decreased oxygen concentrations'. Figure 5.28 illustrates that a 5 percent helium + 95 percent

air mixture at a pressure of 10 psia causes a sufficient reduction of oxy- gen to impair human performance [59]. Helium reduces the partial

pressure of oxygen, already reduced because of the altitude, leading to symptoms of hypoxia. The tissues most sensitive to oxygen deficiency,

such as the central nervous system (brain and eyes), are the first to be
CATASTROPHIC LOSS OF VACUUM 211

affected. The above comments apply equally well to a venting dewar

in any enclosed space. See chapter 6 for more detailed information on

this safety aspect.

A detailed analysis of a specific situation of helium venting is dif- ficult because of the volume and air circulation rate for a particular

aircraft, but the above discussion does point to the need for absolute

pressure control of a venting helium vessel. It is important to note that the heat exchanger-relief valve combination does not prevent the depressurization flash losses.

I 40 30 20 10 0 HELIUM CONCENTRATION % IN Hg-AIR MIXTURE

0 5 10 15 20 25 30 35 40 45 VOLUME 7o OXYGEN IN Hg- AIR MIXTURE

Figure 5.28. Helium concentrations in helium-air mixtures at subatmospheric pressures, showing effect of reduced oxygen concentrations on human performance [59].

5.7.6. Catastrophic Loss of Vacuum

The insulation systems in liquid helium storage vessels must be

maintained at a high vacuum in order to be effective. In well constructed

vessels, the insulation vacuum undergoes very little deterioration with

time since all residual gases except helium are effectively cryopumped

by the 4.2 K surface in the insulation space.
212 STORAGE, DISTRIBUTION, AND HANDLING

Of greater importance is the effect of instantaneous or sudden loss of vacuum caused by accidental structure failure. Most failures in helium

vessels are attributable to mishandling of the equipment. The failure often occurs as a rupture of the neck tube and inner vessel support

members, but it may sometimes occur as a piercing of the outer shell. The effect, in either case, is a drastic increase in heat leak and evaporation rate with a subsequent rapid pressure rise — possibly to the point

of rupturing the vessel. Adequate venting capacity must be provided to prevent excessive

pressure buildup in the event of accidental loss of vacuum. High capacity

relief valves on the helium vessel and burst disks or vacuum-relief

devices on the insulation space are important design considerations for liquid helium equipment. Emphasis must be placed on selection of the

reliable relief valves on the inner vessel because of the high rate of pres- sure rise resulting from increased heat leak. One solution is to calculate the lost-vacuum heat leak from non- evacuated insulation characteristics and to provide a venting capacity

which will prevent excessive pressurization of the vessel. Experience

has shown, however, that air condensing on the inner vessel will give a much higher heat input than predicted. A theoretical solution becomes

rather involved and difficult because of the lack of data on air flow through

various insulations.

Experimental tests offer a practical solution to the problem. Such

a test was performed by Airco and by Superior Air Products on a 25-1, nitrogen-shielded, helium dewar [7] and by Cryogenic Engineering Company [60] on a lOO-l, vapor-shielded dewar. In both of these instances,

the tests resulted in the accepting of a particular size relief valve, rather

than the actual determination of a venting capacity.

These results were later extrapolated for larger size vessels, up to 1,000-1 capacity, based on the assumption that a similar design would

have a lost-vacuum heat leak directly proportional to the surface area

of the inner vessel. The subsequent occurrence of structural failures and

loss of vacuum in commercial liquid helium container transport has

confirmed the experimental and extrapolated value selections.

A method of relief valve sizing, advocated by the Compressed Gas

Association, is applicable and commonly used for larger cryogenic vessels (liquid semi-trailers). A formula based on predicted insulation

properties from Pamphlet S-1.2 is generally accepted as adequate by

the member CGA companies. The reader is referred to that publication

for full details [51].
TRANSFER LINE DESIGN 213

5.8. TRANSFER OF LIQUID HELIUM

The transfer of liquid helium has always required a high degree

of sophistication in equipment and procedures. The properties of helium

which affect hquid transfer are the low temperature and low latent heat. The small difference in the saturated vapor and liquid densities is also important. Since some vaporization may occur during transfer, and the

two phases, liquid and vapor, differ in density by a factor of only 8, they

do not always separate well. The problems, techniques, and equipment

vary greatly with the particular instance of transfer. The end use may

be liquid, cold vapor, or compressed gas; the flashed vapor may be

recovered or vented to the atmosphere; and product purity may or may

not be important.

The losses which occur during transfer may be divided into equipment related losses and procedure i elated losses. The equipment most

commonly used is described in the following section. The losses which

are inherent with various transfer procedures are described starting in section 5.8.3.

5.8.1. Transfer Line Design

Liquid helium transfer lines may be made from rigid or flexible (convoluted) metal tubing; copper, bronze, and stainless steel are the

most common materials. A typical example is illustrated in figure 5.29. Small transfer lines are usually made as a single unit, but larger ones may be fabricated as individual segments with special couplings.

Simple high-vacuum insulation is commonly used to minimize

cooldown losses even at some sacrifice in heat leak, especially in small

lines. Evacuated multilayer insulation is used in some larger lines. In any vacuum-insulated transfer line, the outer jacket must be

sealed to the inner helium line at both ends in order to form the vacuum

space. When the inner line is cooled, it contracts while the warm outer

jacket does not. This differential thermal contraction is easily taken up

in flexible transfer lines, but in rigid lines, some means of accommodating

this contraction must be incorporated. Some rigid transfer lines use

Invar or similar alloys for the inner tube to minimize the thermal contract. Most commonly, though, bellows-type expansion joints are placed

in either the inner line or outer jacket.

Transfer lines of any appreciable length require additional mechanical supports (spacers) to prevent the inner Hne from directly
214 STORAGE, DISTRIBUTION, AND HANDLING

lOL

lATiON

LIQUID DISCHARGE

LEG

LIQUID WITHDRAWAL

LEG

Figure 5.29, Composite 3116-in-I.D. flexible liquid helium transfer line. Note 1/4 to 1/2-in helium-hydrogen bayonet and 1/4-in vacuum jacketed control valve. (Linde Division, Union Carbide

Corp.)

touching the outer jacket. These supports are comparatively large sources of heat leak. They are, therefore, made of low thermal conductivity plastic or thin metal alloy sheet, and they are designed to have

small, near point contacts with the inner Hne and the outer jacket.

Flexible lines require more supports for a given length than do rigid

lines, but even flexible lines are made with heat leaks as low as 1 Btu/

hr-ft ( ~ O.OlW/cm) for sizes up to V2-in inside diameter.

Heat Leak. Liquid helium is always transferred at, or very near, saturated conditions. Sub-cooling can be achieved for short periods by

pressurization, but for reasons which will become apparent in section

5.8.3, excessive pressurization is undesirable. The transfer line heat

leak will therefore vaporize a portion of the liquid transferred. The vapor

thus formed creates two-phase fluid flow with resulting increased pres- sure drop. This is not as serious with helium, however, as it is with most

other fluids. The relative similarity of the liquid and vapor densities and

viscosities allows mixed flow to occur with little interference of the two

phases. Pressure drop for liquid helium, for example, with 10 percent
TRANSFER LINE DESIGN 215

vapor by weight, may only be two or three times that for 100 percent

hquid flow. Valves. Vacuum jacketed, extended stem valves are generally

suitable for liquid helium. Thin-wall stainless steel tubing is commonly

used for the stem material, and the hollow stem may be evacuated to eliminate gaseous conduction along its length. A double-stem design

similar to that shown in figure 5.30 allows all welded construction in the vacuum space and thus avoids O-ring or mechanical seals. Several

valves used in helium service are shown in figure 5.29 (small size) and

figures 5.30 and 5.31 (larger sizes). These have relatively low heat leaks,

aU welded construction in the vacuum space, and easily replaceable

seats. Either O-rings or packed stems at the warm end are used on the

helium space. The valve in figure 5.30 includes a connection to the helium

space for a purge valve or safety rehef device.

Other types of specialty liquid valves are described in the literature

[6, 61]. These two references also provide excellent background for cryogenic transfer and for design and selection of related equipment.

Couplings. Vacuum jacketed, bayonet couplings are available in a wide variety of sizes and configurations. The ever-present problem of the

industrial gas industry, that of mating connections, is thus very serious.

The use of adapters, in addition to being costly, is often prohibitive

because of increased heat leak. Several commonly used bayonet couplings will be described.

Two bayonet couplings used in early helium service are shown in figures 5.32 and 5.33. The first is widely known as the Air Force hydrogen

bayonet and is used extensively for that fluid. It has a short length, about

7 in, compared to its 2-in diameter and has a relatively high heat leak

for liquid helium. The NASA bayonet (fig. 5.33) has a length of about 14

in for its 1-in-diam inner fine, which results in very low heat leak. Check

valves at the liquid end and a bleed valve on the vapor space are advantageous operating features. This particular device, however, is more

elaborate than is necessary for most applications.

The current trend is toward simplified construction and longer heat

leak paths, up to 15 in. Two examples are shown in figures 5.34 and 5.35; both are supplied by the manufacturer with added length for helium

service. An O-ring at the warm end provides an adequate seal on the

annular gas space to prevent leakage. The cold end may or may not have an additional sealing surface. If a liquid seal is not used, the mating

annular passage is very narrow to minimize convection currents in the

306-922 0-68— 15
216 STORAGE, DISTRIBUTION, AND HANDLING

JACKET

Figure 5.30. Typical vacuum jacketed liquid helium valve, showing thin- wall evacuated stem design with warm end connection on the helium gas

space.

(Courtesy Consolidated Precision Company)

gas space. It is desirable to orient the bayonet in a horizontal or cold

end down position. If this is not done, convection currents and possibly hquid seal leakage will greatly increase the heat leak and cause

the O-ring seal to freeze and possibly leak. Experience with the Cryenco

bayonet (fig. 5.34) shows that it performs well with or without a Teflon

liquid seal when tilted at least 25° from the horizontal position.

One more bayonet is shown in figure 5.36, the Linde helium-hydrogen bayonet. This has a relatively short heat path (about 4 in), but has a

sufficiently low heat leak for use with small helium transfer fines such
TRANSFER LINE DESIGN 217

Figure 5.31. Typical vacuum-jacketed liquid helium valve.

(Courtesy CVI Corporation)

TEFLON

-INNER LINE ^S. S. TUBING LIQUID SEAL OUTER LINEOUTER LINE S. S. TUBING 0 RING^ INNEf GAS SEAL LINE

FEMALE MALE

Figure 5.32. Air Force type hydrogen bayonet used for helium service, 2-in diam and 7-in length.

Note Teflon liquid seal and O-ring gas seal.
218 STORAGE, DISTRIBUTION, AND HANDLING

as figure 5.29. It has an advantage of being interchangeable in V4-in and

V2-in hne sizes, and its use is well established in both helium and hydrogen service.

A slight, intentional gas leak in a bayonet actually reduces the total liquid evaporation resulting from bayonet heat leak [62]. The escaping

helium vapor intercepts the normal heat conducted in the coupling,

much the same as does the vapor vented through the neck of small

dewars.

CHECK VALVES - OPEN Figure 5.33. NASA helium bayonet, 1-in diam and 14-in length.

Note liquid and gas product seals, spring operated check valves on the liquid line, and bleed valve on the gas space.

Figure 5.34. Liquid helium bayonet, male half, l-^A-in diam, 15-in length.

Note Teflon seal on liquid end and O-ring seal at the warm end. (Courtesy Cryogenic Engineering Company)
TRANSFER LINE DESIGN 219

Figure 5.35. Liquid helium bayonet assembly, a warm end O-ring seal is shown, but a

narrow mating annulus eliminates the need for a liquid end seal.

(Courtesy CVI Corporation)

GAS SEAL FEMALE MALE

Figure 5.36. Liquid helium-hydrogen bayonet V4-in or V2-m diam 4-in length, metal-to- metal liquid seal, and O-ring gas seal. Liquid Nitrogen Shielding. For more stringent heat leak requirements, the transfer Hne may be shielded with liquid nitrogen. It is possible to construct a three- or four-wall transfer tube, consisting of a vacuum jacketed helium line completely enclosed within a nitrogen line. The nitrogen portion may or may not be vacuum insulated, depending

on economic considerations. Nitrogen-shielding of valves and couphngs

is also possible and perhaps more desirable since these are sources of

high heat leak. A simpler method of shielding is by nitrogen tracing.

Small diameter copper tubing is thermally bonded to the transfer line
220 STORAGE, DISTRIBUTION, AND HANDLING

jacket, valve, or bayonet connection. Liquid nitrogen is then circulated

through the tubing to reduce the outside jacket temperature.

5.8.2. Cooldown Losses

Cooldown of equipment with a cryogenic fluid consumes a quantity

of product porportional to the mass and specific heat of the vessel and

inversely proportional to the heat capacity, latent plus sensible, of the

cryogen. Helium has a low latent heat and large cooldown losses are common. On the other hand, if the high sensible heat of helium vapor is utilized, the losses may be reduced by a factor of 15 or 20. For example,

if only the latent heat of hehum (9.02 Btu/lb) is used, 15.3 1 of Hquid will vaporize in cooling 1 lb of stainless steel from 300 K to 4.2 K, whereas,

if it were possible to use all of the available refrigeration capacity of the vapor, only 0.36 1 would be required. In examining the heat capacity of structural materials versus temperature, it is clear that most of the refrigeration is required at the

higher temperatures. Table 5.8 gives the specific heat and enthalpy

versus temperature for several metals [6, 63]. By precooling a piece of

equipment to intermediate temperatures with less expensive cryogenic

fluids, in particular nitrogen, considerable helium can be saved. While

the latent heat of 15.3 1 of liquid helium is required to cool 1 lb of stain- less steel from 300 K to 4.2 K, that of only 0.70 fiters is required to cool

the same mass from 77 K (nitrogen boiling point) to 4.2 K. A similar

Table 5.8. Specific heat and enthalpy vs. temperature for several common metals

Specific heat — Btu/lb -F Enthalpy -Btu/lb

Temperature K Cu [63] Al [63] 18-8 S.S. [6] Cu [63] Al [63] 18-8 S.S.*

4.2 0.000024 0.000068 0.00031 0.0000646 0.000231 0.000085 20 0.00184 0.00213 0.0011 0.0146 0.0206 0.012 77 0.0466 0.0803 0.038 2.35 3.61 1.74 100 0.0607 0.115 0.057 4.57 7.63 3.76 300 0.0923 0.216 0.114 34.3 73.3 38.1

*Determined by graphical integration — j Cpdt. Jo
COOLDOWN LOSSES 221

10

0 80 120 180 240 300 INITIAL TEMPERATURE , To (^K)

Figure 5.37. Specific liquid helium requirement for cooldown vs. initial equipment temperature. Maximum liquid requirement, tTmax from eq (5.18) and minimum liquid requirement

o-^in from eq (5.19) [64].

savings is possible if advantage is taken of the high sensible heat of

helium, and for cooling other structural materials. In any event, it is clear that some degree of precooling is desirable.

Complete cooldown of a piece of equipment using only the latent heat is an extreme which may be considered as the maximum liquid

requirement. The cooldown consumption in this case may be found by

dividing the enthalpy change of the equipment from initial to final temperature and dividing by the latent heat of heHum (eq 5.18). Results

are shown as maximum specific liquid requirement versus initial temperature in figure 5.37 for cooling to saturated liquid temperature.
222 STORAGE, DISTRIBUTION, AND HANDLING

^max (5.18)

where CTmax— maximum specific liquid requirement in pounds of cryogen

A minimum liquid requirement may be defined as the liquid re- quired to cool a system to its operating condition if all of the refrigeration available in the liquid is utilized. A unique treatment of this problem by Jacobs [64] gives a reliable estimate of the consumption by

considering the ideal case. The technique simply assumes that the temperature of the leaving fluid is equal to the temperature of the warmest

part of the equipment. This is the limiting case of effective heat exchange.

The minimum specific liquid requirement is given by eq (5.19) in terms

of initial and final equipment temperature:

where (Tmin = minimum specific liquid requirement,

Ta= initial equipment temperature,

Tc = final equipment temperature,

Ts = saturation temperature of fluid, Te = equipment temperature,

Ce = equipment specific heat,

\ = latent heat of fluid,

Cpf= isobaric specific heat of fluid vapor.

In addition to the double integration above, we require information

regarding the specific heats of the equipment and vapor. Since the

specific heats are not constant and are not available as analytic functions

of temperature, numerical or graphical integration is necessary. This was done using tabulated specific heat data for helium, specific heat and

enthalpy data for copper, aluminum, and 18-8 stainless steel, and subject

to the condition that the final equipment temperature is equal to the

saturated liquid temperature [64]. Figure 5.37 gives the minimum specific liquid requirement, CTmin, as a function of initial equipment temperature using helium as the cooling fluid. Curves are shown for aluminum,

copper, and stainless steel materials and comparing the maximum conper pound of metal.

hi = initial specific enthalpy,

h2 = final specific enthalpy.

k = latent heat of cryogen.

CpfdTf \ die (5.19)
PRESSURE TRANSFER 223

sumption from eq (5.18) to the minimum consumption from eq (5.19). While the minimum Uquid requirement represents an ideal situation,

this may be approached by careful cooldown techniques.

Several techniques for minimizing cooldown losses follow directly

from the above. They are: use of thin-wall construction and materials

having low specific heat for parts in contact with helium; precooling with

liquid nitrogen; and utilization of the sensible heat of helium during cool- down. There are other considerations also deducible from the above,

but more directly resulting from experience. A slow transfer rate during

cooldown, especially for transfer lines, will take advantage of longer

fluid equipment contact and higher overall heat transfer rates. Observ- ance of the vapor flow leaving the equipment and estimation of its temperature (condensation, frost, liquid air, etc.) will aid in judging the rate of liquid flow for cooldown.

The liquid discharge should be located at or near the bottom of the

receiver. The liquid thus vaporized will displace the warm vapor up- wards and contact the vessel wall in an approximately counterflow

arrangement. If liquid is discharged at the top of a warm vessel, the

large quantity of helium vapor can entrain liquid as droplets and carry them out through the vent. Also note that inherent in the cooldown with

liquid helium, the vessel must be filled with saturated vapor at 4.2 K. This requires 0.13 ft^ of liquid for every cubic foot of volume, which is later displaced as a flash loss when the vessel is fiUed with liquid.

Finally, heat exchange between the incoming liquid and the exist- ing warm vapor should be prevented, as this will vaporize product. The

transfer tube should be completely vacuum jacketed, including the

portion that extends into the inner vessel. The effect is also evident in liquid discharge, where the warm pressurization gas can vaporize the

liquid being withdrawn. In one case, a 20-in length of uninsulated line,

exposed to the warmer gas above the liquid, completely vaporized 80

Ib/hr of helium being withdrawn.

5.8.3. Pressure Transfer

We shall for the purpose of discussion, consider transfer of liquid

helium in several separate categories; there is considerable overlap of these categories. Several methods of transfer are illustrated in figure

5.38. The conventional pressurized transfer will be considered first; gravity, liquid pumping, and cold vapor compression methods are dis- cussed in the following section. Liquid and vapor may be transferred
224 STORAGE, DISTRIBUTION, AND HANDLING

WARM GAS VAPOR

LHe TRAILER

LIQUID

RECEIVER

VAPOR COMPRESSOR

VAPOR RETURN

IQUID -=r

LHe TRAILER RECEIVER CONVENTIONAL PRESSURE

TRANSFER COLD VAPOR COMPRESSION

VAPOR

LHe TRAILER

V

AH

i_

LIQUID

RECEIVER GRAVITY TRANSFER

VAPOR. ^ VAPOR

r RETURN

LHe TRAILER J3

RECEIVER

LIQUID PUMP TRANSFER

Figure 5.38. Comparative liquid transfer techniques, illustrated for transfer from a

liquid semi-trailer to storage vessel.

without external w^ork or additional pressurization if the storage vessel

pressure is sufficiently high. Several techniques are illustrated in figure

5.39. While the actual liquid transfer is straightforward (see sec. 5.8.1), the behavior of the product remaining in the storage vessel during withdrawal is given special consideration as depressurization (see sec. 5.8.5). An energy addition is required for pressurization; this may be vessel

heat leak, warm hehum gas addition, or provision for a heat source

within the container. Normally, heat leak provides a portion of the ini- tial pressurization, as well as a minor contribution during the transfer. The relationship of heat leak to pressure rise is treated in section 5.4.1. The heat added by insertion of warm transfer tubes is of some importance iri product transfer from small dewars. The warm transfer tube

vaporizes liquid as it is placed in the dewar, and if a gas seal is provided

at the outlet of the neck tube, the vapor thus generated provides a source

of pressurization. The pressurization achieved in this manner may be

sufficient to transfer from 20 to 30 1 of helium without the use of additional

pressurizing gas.
PRESSURE TRANSFER 225

VAPOR

LIQUID YIELD LHe TRAILER SIMPLE BLOW DOWN

INTERNAL HEAT EXCHANGER

VAPOR

LIQUID YIELD LHe TRAILER JOULE THOMPSON

HEAT EXCHANGER

LIQUID YIELD LHe TRAILER

VAPOR

LHe TRAILER

LIQUID YIELD RECEIVER

LIQUID DRAW DE PRESSURIZ ATION EXTERNAL J-T EXCHANGER

Figure 5.39. Comparative liquid extraction and trailer depressurization techniques.

Use of an internal or external J-T exchanger increases the liquid yield and reduces flash losses. For substantial liquid transfers, helium gas pressurization is usually required, if not to achieve initial pressure buildup, then to maintain

the necessary pressure during liquid withdrawal. The source of warm

gas is usually from high-pressure gas cylinders with the helium gas at essentially ambient temperature. A mass and energy balance yields

the following relations for such an addition of mass to a constant-volume

system:

miV\ = {mi-\- mg)v2

hgTUg -\- miUi = m2U2

where hg = specific enthalpy of gas added,

mi = initial system mass, mg = mass of gas added,

Vi and i;2 = initial and final specific volumes,

Ui and U2 = ^^^^^3l and final specific internal energies,

^= volume of system.

(5.20)

(5.21)
226 STORAGE, DISTRIBUTION, AND HANDLING

Generally, rug and Wz are unknown. A further simplification may be

made by assuming that the mass of warm gas added is small compared

to the system mass, an assumption which is valid for most purposes.

where enthalpy of saturated vapor at final conditions and U2 is the

internal energy at the final pressure and initial specific volume of the

system.

In either case such a pressurization often results in an end state which Hes in the subcooled hquid region, i.e., to the left of the saturated

liquid line on a temperature-entropy chart (see fig. 2.4). Once the desired transferring pressure is obtained, a continuing

supply of heat or warm gas is necessary throughout the transfer period.

As liquid is withdrawn, cold vapor must be generated to fill the void space. An optimum or ideal transfer efficiency may be calculated under

the assumption that the equilibrium exists at saturated conditions. For

every unit volume of liquid withdrawn, a quantity must remain behind

to fill the void as a unit volume of vapor. Transfer efficiency = 100

{pL — p\')lpL, where pL and p\ are the liquid and vapor densities at the

transfer pressure P2 [5]. At 1 atm pressure, a transfer efficiency of 86.8 percent is calculated from the above. This also assumes that the mass of the pressurization gas is neghgible. Figure 5.40 shows the heat

addition in Btu per pound of liquid transferred as a function of the

transfer pressure, and the equivalent heat addition in terms of warm

pressurization gas [65].

Lowering the temperature of the pressurization gas reduces the

amount of vaporization, although larger quantities of gas are required.

Two techniques to lower the temperature of helium used for pressurization are: use of a liquid nitrogen pre-cooler; and returning the flash gas with a cold vapor compressor (see fig. 5.38). A more complete mass, volume, and energy balance then becomes:

mg{hg — hv) = mi{u2 — Ui) (5.22)

m\V\ = m2f2, (5.23)

and rrighg -\- miUi = m2U2-\- niehe. (5.24)

where subscripts g, 1, 2, and e represent gas added, initial conditions,

final conditions, and exit conditions, respectively.
PRESSURE TRANSFER 227

10 20 SATURATION PRE SSUR E - P S I

A

Figure 5.40. Liquid helium transfer, required heat addition, q, for constant pressure liquid withdrawal and equivalent warm

press urization gas consumption vs. transfer pressure [65].

The greatest source of error in transfer efficiency calculations is thermal stratification. If no heat transfer takes place between the gas and liquid phases, a 100 percent efficient transfer is possible. This situa- tion may be approached by quick pressurization with rapid liquid withdrawal. Many theoretical treatments of this topic can be found in the

Advances in Cryogenic Engineering. A recent paper by Epstein [66]

gives a generalized technique for calculating the pressurization gas requirement for cryogenic fluid transfer with confirming experimental

results. As stated earlier, liquid may be transferred under pressure, without

external pressurization if the initial container pressure is sufficiently

high. This technique is used in the system described in section 5.9. As liquid is withdrawn, the vessel pressure decreases to a point where

either no liquid remains or atmospheric pressure is reached. An analysis

of this system was made by Canty [65] assuming no thermal stratifica-
228 STORAGE, DISTRIBUTION, AND HANDLING

tion. His results are shown in figure 5.41, which shows the pressure

changes as a function of the percent hquid withdrawn for various initial pressures. The limit of withdrawal or "liquid empty" curve is also shown.

3 5

I

1 1 1 1 1 Pcr3 0

5

0 20 40 60 80 100 PERCENT WITHDRAWN

Figure 5.41. Pressure history vs. percent withdrawn for a helium

vessel initially liquid full without external pressurization; the dashed line indicates the pressure at the liquid empty point.

Note that the quantity of liquid that may be withdrawn down to 14.7 psia is a function

of the initial pressure [65].

Throttling Losses. The transfer of a liquid at or near saturated conditions involves what is commonly called a throttling flash loss. This

flash is attributed to the isenthalpic pressure drop. The properties of helium are such that this flash or loss rate is unusually high. Because of helium's low critical temperature, the degree of subcooHng that can be

achieved by quick pressurization of the liquid is only 1 or 2 K. The

throttling loss is predicted by an isenthalpic expansion from storage

pressure Pi to receiver pressure Pz (usually 1 atm). We may properly

consider the subcooling by taking the initial enthalpy as that of the
PRESSURE TRANSFER 229

saturated liquid at the storage temperature. Equation (5.25) gives the

percent flash when transferring liquid from Pi to Pzwt % flash = 100 hnull^

(5 25)

Ilv2 — hl2

where hn = enthalpy of liquid at Pi.

/i/2 = enthalpy of saturated Uquid at P2.

hi-2 = enthalpy of saturated vapor at P2.

Figure 5.42 presents the flash loss versus storage pressure for saturated liquid conditions. This is not the total transfer loss, but only

that due to throttHng the flow. Even so, the advantage of low transfer

pressures is obvious. There are instances when fluid throttling is essential for collecting liquid in the receiver — for example, if the storage

Figure 5.42. Helium throttling flash losses.

Loss in percent liquid transfer vs. storage pressure p, for an isenthalpic expansion from p, to 14.7 psia.
230 STORAGE, DISTRIBUTION, AND HANDLING

vessel contains supercritical fluid. This expansion is the same phenomenon as that across a Joule-Thomson valve in a liquefier. Pressure Drop. In view of the potential for high flash losses, it is indeed fortunate that pressure drop with liquid helium is generally

lower than for other cryogenic fluids. The flow in transfer lines is almost always turbulent. The Fanning

equation, in the form of eq (5.26), is thus appropriate with friction factors

taken from a Moody diagram.

Ap = 2/^^ (5.26)

For liquid helium in long smooth tubes, the following correlation by Koo

for high values of the Reynolds number, Nr = DupIix^ is applicable.

0.00140+ 0.125/A^^o-32) (5.27)

The viscosity of helium is quite small. The table below compares

values for liquid and gaseous helium to those for nitrogen.

Viscosity* Ib/ft-hr

Gas Liquid

Helium 0.0030 0.0046

Nitrogen 013 .39

*These values for saturated conditions at 1 atm. The resulting high Reynolds numbers for a particular flow rate give low

friction factors for helium. For a given volume flow rate, the pressure

drop of liquid helium is an order of magnitude less than for nitrogen.

This helps in keeping flash losses to a minimum. Experience has shown,

for example, that adequate transfer rates for small experimental apparatus may be achieved with dewar pressures of only 1 or 2 psig when

transferring through a typical Vs-in transfer line. The Fanning equation is also applicable for flow through corrugated

metal hose, but it is difficult to determine the correct friction factor. An equivalent roughness (e/Z) on a Moody diagram), taken as the ratio of depth of convolution to maximum diameter, gives poor results.
GRAVITY AND PUMP TRANSFER 231 A more practical method of predicting pressure drop for flow through

corrugated hose was developed by Hawthorne and von Helms [67]. Their correlation assumes fluid flow through uniformly spaced orifices,

where every convolution represents an orifice. The results of this work

are presented in eq (5.28).

where Ap= pressure drop in psi

A^= number of corrugations

Z)= inside diameter, inches

S = corrugation pitch, inches

F= fluid velocity — ft/s and p= fluid density — Ib/ft^.

Experimental results with water, and qualitative results with liquid

nitrogen give good agreement with this expression. While there is no

experimental evidence to verify its use for liquid or gaseous hehum,

reasonable agreement is expected. Additional details are available in the

original paper [67].

5.8.4. Gravity and Pump Transfer

Any comparison of gravity with pumping transfer methods is, for the present, academic. A review of the literature on cryogenic pumping

reveals no commercial applications of liquid hehum pumping. However,

pump transfer techniques may well be developed in the near future. A fairly large difference in the elevation of the source and receiving

vessel is needed to create the pressure differential necessary for the

gravity transfer of hquid hehum. This is caused by the small hquid

density of 0.78 Ib/ft^. At the normal boiUng point, a 1-ft Hquid head

exerts a pressure of only 0.054 psi. When flashed or displaced cold vapor

(4.2 K) is returned to the storage vessel, the driving pressure for liquid transfer is reduced to only 0.046 psi/ft of head. This occurs when liquid

is transferred in a closed cycle from a storage vessel to a receiver (see

fig. 5.38). In both cases, the driving pressure differential is reduced with

increased saturation pressure as illustrated in figure 5.43. One instance where gravity transfer may play an important role is in a large liquid distribution system. If the storage vessel is placed above

(5.28)

306-922 O-bb— 16
232 STORAGE, DISTRIBUTION, AND HANDLING

0.06 Q

<

UJI

0.03 ' 1 1 1 1 1 1 \ 1 1 1 0 2 4 6 8 10 LIQUID SATURATION PRESSURE - PSIG Figure 5.43. Available head pressure vs. saturated transfer pres- sure for gravity liquid helium transfer. Warm gas return implies external pressurization and saturated vapor return implies

a closed cycle transfer.

the receiving dewar, the transfer is essentially one of displacement;

the hquid displaces cold vapor from the trailer, which in turn fills the

void space of the liquid withdrawn from storage. Additional refrigeration

is needed only to compensate for heat leak and flash losses which are normally small. The present state of cryogenic pumping precludes the use of centrifugal pumps for liquid helium, but we should not ignore the potential

for such use. A centrifugal pump coupled with an active inducer section,

such as that designed by Pesco [68] for hydrogen service, seems to be

a reasonable way to overcome the low NPSH requirement of Uquid

helium. Also, one should note that the pump discharge would generally

be in the supercritical region and possibly the suction as well. If such a pump is operated in the range of subcooled liquid to supercritical vapor,

the fluid process would not likely enter the two phase region at all but

pass to the left of the saturated liquid line on a temperature-entropy

diagram. The operation is then similar to a cold vapor compressor

because of the high compressibihty of helium. A cold vapor compression
DEPRESSURIZATION 233

liquid transfer scheme is shown in figure 5.38. The vapor compressor

provides the driving force for the transfer — essentially producing a pressurized transfer. This technique is especially appropriate for large

(1000 to 10,000 gal) Hquid hehum transfers. One further liquid pump will be mentioned. A positive displacement

pump described by Barrel and Schock [69] is driven by a superconducting solenoid completely immersed in the helium reservoir. The oscillat- ing solenoid pump has a capacity of less than 0.5 1/min, but presumably

this capacity could be increased. An advantage of this pump is the relative ease of automatic flow control. There has been some work on high pressure hehum pumps intended

for cylinder filhng. A pump designed by Linde for heUum-hydrogen

service [70] was intended for use in conjunction with a trailer mounted

vaporizer and liquid trailer; it is capable of producing 6000 psig gas for high-pressure storage [71]. The design was proved in hydrogen service,

but never tested with liquid helium.

Insofar as positive displacement pumps for liquid hehum are con- cerned, several factors must be considered. The isentropic compressibility of liquid helium is about twice that of nitrogen, requiring very

small clearance volumes to achieve good volumetric efficiency. Frictional heating within the pump is best removed on the discharge side,

to avoid interference with the suction flow. The discharge stream should

contact the cold end of the pump barrel for efficient heat interception.

Finally, because of the low NPSH requirement, the use of a fore-pump

is desirable.

5.8.5. Depressurization

We are considering vessel depressurization separately to avoid

the necessary complexity of an overall transfer analysis and to more

closely observe the behavior of the contents of a vessel during the

process. A depressurization may be attributed to liquid or vapor withdrawl, or simply a blowdown to gain access. The two aspects receiving

special attention are the liquid yield and product flash losses. The depressurization process is shown on a temperature-enthalpy

chart for helium in figure 5.44. This adiabatic process is essentially

reversible and, therefore, isentropic. We may assume that the vessel was initially filled at 1 atm to a given percent fuU (1) and allowed to pressurize by heat leak along a constant volume line to (2) and to (3). The depressurization then begins at (3), at a given specific volume and
234 STORAGE, DISTRIBUTION, AND HANDLING

Figure 5.44. T-S diagram illustrating the condition of a liquid helium vessel: filled to 10% ullage at 1.0 atm.

(1), pressure rise during transit to liquid fi ll point (2), and to 6 atmospheres (3) isentropic product withdrawal to liquid full point (4), and liquid (5B) or vapor (5A) withdrawal to 1 atmosphere. No gas addition during cycle.

pressure. A constant entropy line (average entropy of the fluid remaining)

determines the point at which the vessel is completely. full of saturated

liquid. Below the saturation curve, a phase separation occurs. If vapor

is withdrawn (from the top), the product leaving has a higher energy

than that within the vessel and the process ends at (5A) with decreased

entropy and increased hquid. When Hquid is withdrawn, the product

leaves at a lower energy level than that within the vessel and the process

ends at (5B) with less liquid remaining. More product may be withdrawn from the vessel in process (3)-(5B), however, without addition of heat or warm gas to the vessel. An entropy balance is convenient because of the isentropic process, but the variation in leaving conditions requires a stepwise calculation.

The process, however, may be simpHfied by taking the expansion in two

steps [72]. The first step from (3) to (4) is strictly isentropic, where 53 and 54, as specific properties, are equal. From the saturated liquid Une

(4) to either (5A) or (5B), average leaving conditions are used, rather than

successive steps. The result for vapor withdrawal is given by eq (5.29)

which gives the quantity of liquid remaining after depressurization [72].

Vb L Vg{Se — Si ) -hVi{Sy — Se) . (5.29)
DEPRESSURIZATION 235

where 5^= (5d + 5g)/2

Y= liquid yield as percent of initial volume,

1^?= specific volume of saturated liquid at final pressure,

i;g= specific volume of saturated vapor at final pressure,

specific volume of saturated liquid at initial pressure, s/= entropy of saturated liquicj at final pressure,

Sg = entropy of saturated vapor at final pressure,

5ft= entropy of saturated liquid at initial pressure,

Sd = entropy of saturated vapor at initial pressure.

Results of the above are presented in figure 5.45 for depressurization of

a vessel full of hquid at pressure Pi down to 1 atm, giving the percentage liquid loss as a function of Fi.

4 6 8 10 12

p. - INITIAL PRESSURE PSIG

Figure 5.45. Helium depressurization losses.

Isentropic depressurization (expansion) of dewar contents from storage pressure p,

to 1 atm.

If the depressurization takes such a long time that heat leak be- comes a factor, the heat addition may be treated separately, as shown in section 5.4.1, again, as a repetitive step calculation. Should a mass addi-
236 STORAGE, DISTRIBUTION, AND HANDLING

tion also be involved, a First Law analysis, as shown in section 5.8.3, is more effective. It may be possible to assume average leaving conditions

in some cases, rather than resorting to successive step calculations.

Two techniques are useful for increasing the liquid yield: (1) where

the liquid remaining in the trailer is maximized, and (2) where more

liquid is recovered from the product leaving. Their use, of course, de- pends on the ultimate aim of the helium transfer. A heat exchanger and expansion valve located within the storage

vessel is shown in figure 5.39. The flow is throttled to 1 atm and 4.2 K

by the valve preceeding the heat exchanger. The existing fluid, now

colder than that within the vessel, absorbs heat energy with the net

effect of increasing the quality of the liquid remaining after the vessel

is depressurized to some pressure below Pc = 33.2 psia. A comparison

of this depressurization with an ordinary vapor blowdown is shown in figure 5.46 for a trailer which was initially filled with fiquid at 4.2 K and

later depressurized from a pressure Pi to 1 atm. The yield is shown as percentage recovery of liquid in the trailer.

100

tr 90

80

60

FOR TRAILER INITIALLY LIQUID FULL AT I ATM.

3 4 5 6 7 INITIAL PRESSURE — ATMOSPHERES

Figure 5.46. Percent liquid recovery in trailer vs. initial pressure for depressurization

to 1 atm.

Ordinary blowdown vs. internal J-T heat exchanger (increased liquid yield).
FLASH GAS RECOVERY 237

Greater efficiency of liquid transfer can be achieved with the ex- ternal Joule-Thomson exchanger [73] , as shown in figure 5.39. Liquid

is withdrawn from the trailer and, after heat exchange with helium vapor

at 4.2 K. is throttled to 1 atm. The flashed vapor is used to precool the

incoming Hquid. Ideally, the vapor will leave at the same temperature as the trailer; but, because liquid is being withdrawn, rather than vapor, more product will be transferred. In addition to being more efficient, the

external Joule-Thomson exchanger may be used as part of the transfer

system. This may be done by using the annular space of a three wall

transfer tube as the heat exchanger. The performance of this type of exchanger may be compared to the Joule-Thomson exchanger of a helium

Hquefier. Again, stepwise calculations are necessary because of the

variation of inlet conditions to the exchanger as trailer pressure drops.

5.8.6. Flash Gas Recovery

The cost of helium, whether liquid or gas, and the indication of the

order of magnitude of flash loss in the previous sections clearly illustrate the importance of giving serious economic consideration to a flash gas recovery system. The greatest volume of liquid hehum handhng is that of small dewar filling, and in all commercial installations for this purpose, some type of gas recovery system is available. For a hquid-gas helium distribution network described in the fol- lowing section, flash gas recovery is an integral part of the gaseous

product handling. Flash gases from dewar filling at hquid-gas conversion

stations are heated further in the normal vaporizer and then compressed

for gas cyhnder fiUing. Where liquid transfer is a part of a liquefier or

cryostat production, any flash gas is recovered and returned as makeup

to the liquefier.

The two significant problems for hehum gas recovery are capacity

of the gas system and contamination of product with air and moisture.

Liquid helium transfers generate large volumes of flash gas from pre- coohng, heat leak, and liquid displacement. A relatively high-volume

compressor is required to maintain adequate fill rates unless lowpressure gas storage is used to take up slack between gas generation

and compressor capacity. Low-pressure gas storage equipment is generally of the dry type with flexible rubber or plastic composite fining, and

compressors are preferably nonlubricated. This equipment is chosen to eliminate or minimize helium contamination. A gas recovery system
238 STORAGE, DISTRIBUTION, AND HANDLING

may include a charcoal bed adsorber at liquid nitrogen temperature to remove impurities, as in many helium liquefiers. In general, and specifically for systems without adsorbers, considerable attention is given to purging all lines and receiving vessels

with gaseous helium before introducing liquid. There are two techniques

which have proven to be satisfactory in purging contaminants. A pres- sure purge is a periodic buildup of pressure with warm gas and blow

down to atmosphere. This must be done several times to achieve any

degree of success. A vacuum purge is more certain, provided that the gaseous helium back fill is sufficient to vaporize any condensed

or frozen impurities. This should also be repeated two or three times for best results. Finally, the liquid nitrogen precool described in section 5.8.2 may

cause serious difficulties in purging. As much liquid nitrogen must be

removed as possible and the remainder vaporized with warm helium

gas before purging can be successfully achieved. A pressure transfer

is used to remove liquid nitrogen from a dewar, similar to a helium

transfer. As an added note of caution, flexible transfer lines should not be precooled with nitrogen, as liquid will collect in the bellows sections,

which are difficult to purge. Cleanliness is also important, and all pieces

should be cleaned, dried, and capped — especially bayonets and lines that might be exposed to the weather and condensed moisture.

HELIUM EXTRACTION AND

LIQUEFACTION PLANT

GASEOUS

HELIUM CYLINDERS

GASEOUS

HELIUM TRAILER

LIQUID HELIUM DEWARS

1X2

SCALE ^

A. DISTRIBUTION METHOD AT EXTRACTION AND LIQUEFACTION PLANT.

GASEOUS

HELIUM TRAILER oo 1

COMPRESSOR

HEAT

EXCHANGER

\ LIQUID .

] HELIUM ( r—i TRAILERS ^ SCALE

B. TYPICAL LOCAL DISTRIBUTION CENTER.

Figure 5.47. Schematic presentation of bulk liquid helium distribution systems.
LARGE-SCALE HELIUM DISTRIBUTION 239

5.9. LARGE-SCALE LIQUID AND GASEOUS

HELIUM DISTRIBUTION

Traditionally, helium has been shipped as a compressed gas in high-pressure tubes mounted on truck trailers or rail cars, as described

in chapter 1. The most common tube trailers typically haul up to 40,000

scf* of helium compressed to approximately 2,400 psig. More recently,

designs using longer trailers and tubes made of higher strength steels have increased the carrying capacity to nearly 130,000 ft^. The U.S. Bureau of Mines has operated gas rail cars for many years. During this time, many design improvements have been made; at present, the maxinum capacity of newer units is about 345,000 ft^ at pressures to 4,000

psig.

Table 5.9. Comparison of helium shipping containers

Typical units Heliumcapacity

Helium

weight

{Wh)

Container's weight

{Wc)

Ratio

scf lbs lbs Standard cylinder 285 3 135 45 Standard trailer 40,000 415 26,800 65 Jumbo trailer 130,000 1,345 56,600 42

Rail car 300,000 3,100 234,000 75 Dewar (500-1) 13,350 138 1,100 8 Liquid trailer (10,000-gal) 1,000,000 10,350 43,350 4

Commercial shipment of liquid helium was started in the late 1950's

with shipments to various consumers of Hquid helium for research

purposes [74]. Recent technological advances in cryogenic insulation

and heUum hquefaction coupled with rapidly growing markets for helium

and the liberalized Helium Conservation Act of 1961 has made practical

the transport of liquid helium even when the end use is gaseous helium.

The distribution of helium as a Hquid rather than a gas offers potential

economic advantages comparable to those now achieved in distributing

and handling oxygen, hydrogen, argon, and nitrogen in liquid form.

The weight advantage offered by shipping liquid is shown in table

5.9 and results in much lower transportation charges. The capital cost

*Industry standard of measurement is cubic feet at 1 atm and 70 F (scf).
240 STORAGE, DISTRIBUTION, AND HANDLING

of a liquid helium trailer is approximately one and one-half times that for a gas rail car; however, compared to a rail car, a liquid trailer can complete twice the deliveries in a given time and has three times the capacity. Therefore, liquid trailers require only one-fourth of the capital investment of gas rail cars to distribute the same amounts of helium. This

savings in transportation cost and capital investment is only partially

offset by the investment and operating cost of a helium liquefaction

system. The ability to supply either liquid or gas from distribution centers

is a further advantage of liquid transport systems.

5.9.1. The Airco Distribution System

In May 1966, the Air Reduction Co., Inc., started transporting liquid

helium in semi-trailers from a new helium separation and liquefaction

facility to four distribution centers within the United States [75] . Figure

5.47 is a block diagram showing the major components of and the products handled in this system.

Production Facility. Helium for commercial bulk distribution is obtained from natural gas in an integrated purification and liquefaction

plant. The Kansas Refined Helium Company plant in Otis, Kans. [75-

FlGURE 5.48. Flow diagram illustrating, in simplified form, a typical helium extraction and liquefaction process.
THE AIRCO DISTRIBUTION SYSTEM 241

77], (sec. 3.3.1) produces more than 800 1/hr of 99.995 percent helium.

This plant currently has the largest single storage tank for liquid helium —

a 32,000-gal, liquid nitrogen-shielded unit with a 1 percent per day

evaporation rate. Basically, the plant process involves the following

steps: (1) preliminary gas cleanup, (2) heavy hydrocarbon removal,

(3) methane and nitrogen stripping, (4) final helium purification, and

(5) helium liquefaction. Figure 5.48 shows a simplified flow schematic;

a more detailed process discussion can be found in chapter 1. System Operation. The distribution sequence starts with the liquid

helium trailer being completely filled with liquid (at 16 psia) at the

liquefaction plant. FiUing and billing are done by weight (fig. 5.49).

Extreme accuracy in weighing is important. Liquid helium is valued as high as $15 per pound. Thus a full trailer contains over $150,000 of product. Weighing errors, due to scale inaccuracies, snow and ice buildup on the trailer, etc., can be costly.

The liquid trailers are normally refilled in about 4 hours. Upon

completion of the fill, the trailer vent is closed and the trailer is transFlGURE 5.49. The 10,000-gal liquid helium semi-trailer is weighed on a truck scale when

making custody transfers of its 11,000-lb lading.
242 STORAGE, DISTRIBUTION, AND HANDLING

0 2 4 6 8 10 12 14 TIME (DAYS)

Figure 5.50. Pressure rise in 10,000-gal liquid helium trailer as a function of time, showing percent full dependency relation.

ported, within 48 hours, to any of several transfiUing locations. During

transport, heat leak into the trailer will cause a pressure rise rate of

about 0.5 psi per hour (see fig. 5.50).

With the nonvented mode of operation for the liquid helium trailer,

its contents will be a single-phase, cold, dense, fluid when it arrives

at the transfilling location. Normal procedure is to depressurize the

vessel to reestablish two-phase product. This is accomplished by withdrawing the dense fluid from the trailer through a manifold, heat ex- changer, and compressor into gas trailers and cylinders. During this
THE AIRCO DISTRIBUTION SYSTEM 243

Figure 5.51. At a transfilling facility, liquid helium is pressure

transferred from the liquid semi-trailer into various size liquid dewars, as required.

All transfilling hoses are vacuum insulated. Gas generated in the operation is recovered and compressed into cylinders or trailers. The dewars are filled on a scale and contents checked by level probe. Normal trailer pressure is 5 psig. (Courtesy Air Reduction Co.,

Inc.) depressurization period, liquid containers are not normally filled. Optimum trailer discharge pressure has been found to be 4 to 5 psig.

This permits Hquid withdrawal rates of up to 6 gal/min and minimizes

flash generation while filling liquid containers.

It is possible to obtain liquid helium from the trailer immediately

upon its arrival at a local distribution site, even if the product is in the

supercritical condition. This can be accomplished by transferring the

product directly into dewars, utilizing a transfer tube incorporating a

throttling valve. The efficiency of this method can be increased by

including a heat exchanger [73] in the transfer tube as discussed in
244 STORAGE, DISTRIBUTION, AND HANDLING

section 5.8.5. This method is not normally used because varying trailer head pressure makes automatic flow control more difficult, and because

all vapor generated must flow out through the limited vent passage

of the dewar neck tube; this limits the transfer rate. When liquid dewars are being filled, as in figure 5.51, all gas gen- erated in the operation is collected, warmed, and compressed. Since

the compressor is a constant capacity machine (7000 scfh in the system

being described), it limits the rate at which liquid transfers can be

accomplished while still maintaining a no-loss system. When dispensing

500,000 ft^ per week of product from a 10,000-gal liquid trailer, as much

as 65 percent of the product can be recovered as liquid.

If approximately 95,000 scf of product is withdrawn daily, the

trailer pressure will remain constant. Should product demands exceed

this figure, the 100 Btu/hr trailer heat leak wiU not sustain the trailer pressure and external pressurization would be required. Since the

trailers are not equipped with pressure building coils, gas from the

fourth stage of the compressor is used to maintain a minimum pressure

of 5 psig.

Liquid Trailers. The liquid helium trailer is the key component

in a bulk liquid helium distribution system. It must be designed to transport and dispense the product without loss. Therefore sloshing

effects during transportation (sec. 5.6.1) and thermal oscillations (sec.

5.6.2) must be minimized, requiring costly design features. One of a fleet of tractor-trailer units is illustrated in figure 5.10. The tanks of these trailers have a nominal capacity of 10,000 gal. Heat

leak is limited to 100 Btu/hr by the use of evacuated multilayer insulation

and a special support system. Eight baffles in the tank restrict the motion

of the liquid, which makes the trailers more manageable on the road

and reduces the mechanical damage that would result from unrestrained

surges during abrupt changes of speed or direction. The baffles reduce

sloshing of the liquid when the trailers are in motion and temperature

stratification when they are stationary.

The trailer design also incorporates a pneumatic type suspension on the rear tandem for smooth-riding characteristics to minimize

mechanical energy input into the liquid during transport.

The support system for the inner vessel is designed in accordance

with proposed ICC specification MC 341 given in section 5.7.3. The trailers are normally operated in the 100 percent fuU condition

(0 percent ullage) and sealed for shipment. As is shown in figure 5.50,
THE AIRCO DISTRIBUTION SYSTEM 245

the rate of pressure rise is nearly constant and more rapid than that experienced with trailers less than 100 percent Hquid full. Since the

trailers have a 125 psig maximum operating pressure, the nonvented

hold time, when 100 percent filled, can be as much as 9 days. Trailers

filled to 90 percent of capacity (or approximately 9,600 gal of liquid)

have a slower rate of pressure rise until the density change creates a

hquid-full condition, at which point the contents become single-phase

and the rate of pressure rise approximates that of a trailer initially fiUed

to 100 percent of capacity. The nonvented hold time, when initially filled to 90 percent is 14 days. Actual operating data proved the validity of the

internal energy method for calculating heat leak rate or rate of pressure

rise, as discussed in section 5.4.1 using the data of figure 2.7. Saturated hehum vapor is one-eighth the density of Hquid (at 1 atm);

therefore, a 10,000-gal trailer may contain 1500 lb or more of cold helium

vapor when emptied of hquid. The remaining dense vapor can be ex- tracted by warming, either from trailer heat leak or external pressurization, to raise the residual gas temperature. Trailers are normally returned

to the liquefaction plant with a 250 to 400-lb residual. This requires

that about 11,000 Btu be added after the trailer is empty of liquid.

Ambient temperature helium pressurizing gas will provide approximately

TIME - DAYS

Figure 5.52. Graphical presentation of the variation of residual lading pressure and temperature with time, for various residuals,

in 10,000-gal liquid helium semi-trailer with a 100 Btulhr total heat rate.
246 STORAGE, DISTRIBUTION, AND HANDLING

Figure 5.53. Compression equipment in a liquid helium transfiUing

facility. The four-stage, nonlubricated compressor delivers 7200 scfh of 3000 psig helium gas with a 1 psig inlet. The trace oxygen, dewpoint and hydrocarbon analyzers (shown left center) continuously monitor the product and automatically shut down the compressor when contaminants in the 4 ppm range are detected.

650 Btu/lb per pound; therefore, about 17 lb of pressurizing gas is required for this operation.

The unloaded trailers are returned to the liquefaction plant without

further venting. The heat leak into the trailer warms the residual helium

gas and thereby increases its pressure. Figure 5.52 shows the rate of increase of both temperature and pressure for several values of the

amount of cold vapor left in the unloaded trailer. It is noteworthy that the rate of pressure rise is substantially independent of the amount of

cold gas left in the trailer. Upon the trailer's return to the liquefaction plant, it is cooled to approximately 5 K with liquid helium and filled for another shipment.

Refrigeration used to cool a trailer directly reduces the quantity of
THE AIRCO DISTRIBUTION SYSTEM 247

helium liquefied by the plant during the cooldown period. A compromise

is made on the residual product returned in the trailer to maximize

both trailer payload and plant production.

Gas Handling and Compression. A second major component of the

bulk Uquid distribution system is the gas recovery compressor. The

compressor (see fig. 5.53) is a four-stage machine which boosts 15.5

psia, 70 F, helium to 3000 psig for gas trailer and cylinder filling. To

prevent product contamination, the first three stages of the compressor are of the nonlubricated piston type with Teflon seals and rings.

The fourth stage is a hydrauhcally actuated diaphragm. The fourth stage

could not be of the Teflon ring type because current Teflon ring tech- nology is limited to about 1500 psig. Temperature instrumentation with

TO GAS TRAILERS — 1

JO GAS CYLINDERS

Figure 5.54. Flow schematic for liquid helium transfilling station. The facility operates automatically once started. All gas generated in the transfilling operations is recovered and

compressed into suitable gas distribution containers.

a high-temperature alarm and shutdown is installed on each of the first three stages to protect the Teflon rings from temperatures above 300 F. Ring and diaphragm life should be in excess of a thousand hours. The

machine has a capacity of approximately 7000 scfh when operating within

design conditions.

Figure 5.54 shows the piping for the major components at the local

transfilling stations. The operator has the option of:

306-922 0-68— 17
248 STORAGE, DISTRIBUTION, AND HANDLING

1. Removing either liquid or dense vapor from the trailer for compression into gas cylinders and trailers (after warmup to ambient conditions in a heat exchanger).

2. Fining of liquid dewars from the trailer and gas recovery from the

dewars.

3. Combination of 1 and 2. After the operator has selected the desired operational pattern,

pressure actuated control valves automatically control flows from the

trailer and to the compressor. Once the controls are set, liquid dewars

can be filled as required without upsetting the remainder of the system.

Gas Purity. Local hquid helium transfilling stations have no facilities for repurification should the helium become contaminated during shipping or handling. Therefore, as the helium is compressed into cylinders or gas trailers, it is constantly monitored for moisture and oxygen,

which would indicate product contamination from the atmosphere, and

for total hydrocarbons, which indicate a malfunction in the compressor

diaphragm or seals. Moisture is monitored on an electrolytic type analyzer with a 0 to 10 ppm range; oxygen is monitored electrochemically also over a 0 to 10 ppm range; and hydrocarbons are detected by ultraviolet means using

a fluorescent additive in the compressor hydraulic fluid. The system is shut down autom.atically should either the oxygen or hydrocarbon

contaminant level increase to 4 ppm or the moisture to 2 ppm.

In the electrolytic moisture analysis [78], a sample gas stream is passed through a capillary tube containing phosphoric anhydride (P2O5)

which absorbs all the moisture. Rhodium electrodes, embedded in the

P2O5, are electrically energized, electrolyzing all the absorbed moisture.

According to Faraday's law, the amount of current needed to effect the

electrolysis is directly proportional to the absorbed water. The electrochemical oxygen analyzer [79] is based upon the principle

that the flow of electrical energy to an anode (cadmium or lead) from a cathode (silver) is proportional to the quantity of oxygen consumed in a galvanic oxygen reaction in alkaline solution.
REFERENCES 249

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I
250 STORAGE, DISTRIBUTION, AND HANDLING

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[44] J. R. Clement and John Gaffney, Thermal oscillations in low temperature apparatus. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1960), Vol. 1, p. 302.

[45] J. D. Bannister, Spontaneous pressure oscillations in tubes connecting helium reservoirs to 300° K environments. Liquid Helium Technology (International Institute of Refrigeration, Commission I, Boulder, Colo., 1966), p. 127.

[46] M. T. Norton, Cryogenics Division, NBS, Institute for Basic Standards, Boulder,

Colo., Private communication (1965).

[47] J. H. Bell, Jr. Cryogenic Engineering (Prentice-Hall, Englewood Cliffs, N.J., 1963).

[48] R. H. Kropschot, Mechanical properties of materials. Applied Cryogenic Engineering

(ed. R. W. Vance and W. M. Duke, John Wiley & Sons, New York, 1962b), Chapter

3, p. 152.

[49] R, M. McClintock and H. P. Gibbons, Mechanical properties of structural materials

at low temperatures, NBS Monograph 13, (Supt. of Documents, U.S. Govt. Printing

Office, Washington, D.C., 1960).

[50] F. R. Schwartzberg, S. H. Good, R. D. Keys, and T. F. Kiefer, Cryogenic Materials Data Handbook (Martin Company, Denver, Colo., 1964). (Contract AF33 657)-9161; OTS, PB171809.

[51] Compressed Gas Association (500 Fifth Avenue, New York, 1966), Safety rehef device standards for cargo and portable tanks. Pamphlet S-1.2. See also: Compressed

Gas Association, Handbook of Compressed Gases (Reinhold, New York, 1966),

p. 210.

[52] Compressed Gas Association (500 Fifth Avenue, New York, 1964 revision). Proposed

insulated cargo tank specification MC-341 for low temperature hquefied gases. Docket 1204.

[53] R. J. Richards, A high-vacuum seal-off valve. Rev. Sci. Instr. 25, 520 (1954).

[54] R. C. Van Meerbeke, Commercial handhng of hquid helium dewars. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1967), Vol. 12, p. 646.
252 STORAGE, DISTRIBUTION, AND HANDLING

[55] R. M. Neary, Air solidifying cryogenic fluids, Paper 64-WA.SAF-l, ASME Safety

Conference, New York (1964).

[56] International Air Transport Association Regulations (1965), Sec. IV, p. 35, 10th ed.. Item No. 865, Montreal.

[57] R. G. Biel, Frequency analysis of vowels produced in a helium-rich atmosphere,

J. Acoust. Soc. Am. 34, 347 (1962).

[58] J. P. Cooke, Verbal communication intelligibility in oxygen-helium and other breathing mixtures at low atmospheric pressures. Aerospace Med. 36, 1167 (1965).

[59] National Aeronautics and Space Administration, Bioastronautics Data Book, Washington (1964).

[60] Cryogenic Engineering Company, Private communication (1963).

[61] P. D. Fuller and J. N. McLagan, Storage and Transfer of Cryogenic Fluids (ed. R. W.

Vance and W. M. Duke, John Wiley & Sons, New York, 1962), Chapter 9. [62] D. B. Chelton, U.S. Patent 3,034,319 (1962).

[63] A compendium of the properties of materials at low temperatures (ed. V. J. Johnson), WADD Technical Report 60-56, Part II. [64] R. B. Jacobs, Liquid requirements for the cooldown of cryogenic equipment. Advances

in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1963), Vol. 8, p. 529.

[65] J. M. Canty, Pressure phenomena during transfer of saturated cryogenic fluids. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1961), Vol. 6, p. 272.

[66] M. Epstein, Prediction of liquid hydrogen and oxygen pressurant requirements.

Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1965), Vol. 10, p. 303.

[67] R. C. Hawthorne and H. C. Von Helms, Flow in corrugated hose. Product Engineering

34, 98 (June 10, 1963).

[68] C. A. Schalla, Liquid hydrogen booster pump for aerospace ground-support systems. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1964), Vol. 9, p. 521.

[69] B. Darrel and K. Schock, An electrically pumped liquid helium transfer system. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1966), Vol. 11, p. 607.

[70] C. F. Gottzmann, High-pressure liquid-hydrogen and helium pumps. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1960), Vol. 5, p. 289.

[71] J. W. Marshall, Logistic supply and handhng of hquid hehum. Paper 61-AV-lO,

ASME Aviation Conference, Los Angeles, Calif. (1961).

[72] L. E. Scott, Ideal yield of a Simon liquefier. Cryogenics 3, 111 (1963).

[73] L. E. Scott and J. Kleinbaut, U.S. Patent 3,306,061 (1967).

[74] H. M. Long, The commercial production, storage and distribution of liquid helium. Liquid Helium Technology (International Institute of Refrigeration, Commission I, Boulder, Colo., 1966), p. 187.

[75] G. C. Nubel and H. Hover, Description of a bulk liquid helium distribution system. Liquid Helium Technology (International Institute of Refrigeration, Commission I, Boulder, Colo., 1966), p. 329.
REFERENCES 253

[76] S. Ergenc and Ch. Trepp, A large-scale helium liquefier, Sultzer Tech. Rev. 481 1966,

No. 4, 11215.

[77] Ch. Trepp, A large-scale helium liquefier, Liquid Helium Technology (International

Institute of Refrigeration, Commission I, Boulder, Colo., 1966), p. 215.

[78] F. A. Keidel, A novel, inexpensive instrument for accurate analysis for traces of water. The Pittsburgh Conference on Analytical Chemistry and Applied Spectros- copy (1956).

[79] P. A. Hersch, Galvanic determination of traces of oxygen in gases, Nature 169,

792 (1952); see also British Patent 707,323.

CHAPTER 6

Safe and Efficient Use of Liquid

Helium

Hugh M. Long ^ and Paul E. Loveday ^

Page 6.1. Introduction 256

6.2. Handling liquid helium 257

6.2.1. General precautions 257

6.3. Helium liquefaction 258

6.3.1. Safety in Hquefier design 258

6.3.2. Safety in Hquefier operation 260

6.4. Transport and storage 261

6.4.1. Codes and regulations 261

6.4.2. Materials and methods of fabrication 262

6.4.3. Insulation 263

6.4.4. Conduits and piping components 265

6.4.5. Operating environment 271

6.5. Transfer 274

6.6. Safety in cryostat design and use 279

6.6.1. General design considerations 279

6.6.2. Mechanical considerations 281

6.6.3. Glass apparatus 282

6.6.4. Fill and vent procedures 284

6.7. Summary 286

Appendix A. Sizing of relief devices 288

' Linde Division, Union Carbide Corporation, Tonawanda, N.Y. 14150. 255
6.1. INTRODUCTION

Liquid helium is used in a wide variety of research and engineering

projects, and personnel without previous experience with the unique

characteristics of this fluid niust frequently plan the operation. Chapter

2 emphasized the very low latent heat of vaporization, the low critical pressure, the low density ratio of saturated hquid to saturated vapor, as well as the high volume ratio of 530 R (294 K) vapor to saturated liquid.

These properties of helium require careful planning, since one small

oversight can cause complete failure to provide the desired liquid helium

environment, in addition to exposing personnel to unwarranted hazards.

This chapter outlines some of the considerations necessary to achieve safe and efficient use of liquid helium. It presents general safety

precautions and contains specific sections on liquefaction, transport,

storage and transfer, and cryostat apparatus. A summary and conclusions section is also included. The reader is cautioned to study his specific operation, because the material presented herein can serve only as reminders rather than firm recommendations. Liquid helium

technology is so young and applications so varied that existing rules and

regulations cannot provide complete answers, though the referenced

design guides should be helpful.

Designers of liquefaction and transport equipment use new tech- nology not yet available in established design guides and regulations.

For research applications, only the prudence of the designer can insure

safe performance. It is hoped that the material in this chapter will pro- vide adequate stimuli toward the development of safe design procedures.

Zabetakis [1] ^ has summarized the physiological, physical, and

chemical hazards, as well as additional commentary on laboratory,

plant, and test site safety. He also includes a Safety Data Sheet for each

cryogenic fluid, plus many pertinent references.

Industrial Safety Data Sheets are available from the National

Safety Council, 425 North Michigan Avenue, Chicago, lU. 60611, and

Chemical Safety Data Sheets can be obtained from the Manufacturing

Chemists Association, Inc., 1825 Connecticut Avenue, N.W., Washington, D.C. 20009. Such sources of information can be very helpful in planning operations using other materials associated with helium

liquefaction and its applications.

Figures in brackets indicate the literature references at the end of this chapter. 256
GENERAL PRECAUTIONS 257

6.2. HANDLING LIQUID HELIUM

The potential hazards in handling liquid helium stem mainly from

four important properties:

a. The hquid is extremely cold (helium is the coldest of all cryogenic

liquids).

b. The ultralow temperature of the hquid hehum will condense

and solidify air.

c. Very small amofints of liquid are converted into large volumes

of gas.

d. Hehum is non-life supporting.

In addition, there are potential hazards due to the mandatory use of vacuum-insulated vessels and transfer hnes, and to changes in physical

properties of materials between ambient and low temperatures.

6.2.1. General Precautions

Prior to using hquid hehum, the specific operation should be

reviewed with respect to the following list of precautions.

Cover eyes and exposed skin. Accidental contact of hquid helium,

or cold issuing gas, with the skin or eyes may cause a freezing injury

similar to a burn. Protect your eyes and cover the skin where the possibility of contact with cold fluid exists. Keep air and other gases away from liquid helium. The low temperature of liquid helium or cold gaseous heltum can solidify any other

gas. Sohdified gases and liquid allowed to form and collect can plug

pressure-relief passages and foul relief valves. Plugged passages are hazardous because of the continual need to relieve excess pressure

produced when heat leaks into the cold hehum. Therefore, always

store and handle liquid helium under positive pressure and in closed

systems to prevent the infiltration and solidification of air or other gases. Keep exterior surfaces clean to prevent combustion. Atmospheric

air will condense on exposed helium-cooled piping. Nitrogen, having

a lower boiling point than oxygen, will evaporate first from condensed

air, leaving an oxygen-enriched liquid. This liquid may drip or flow to nearby surfaces. To prevent the possible ignition of grease, oil, or other

combustible materials which could come into contact with the air- condensing surfaces, such areas must be cleaned to "oxygen-clean"

standards. Any combustible foam-type organic polymer insulations

should be carefully apphed to reduce the possibility of this air condensation-oxygen enrichment process — which if followed by an impact could

set off an explosive burning of the foam.
258 SAFE AND EFFICIENT USE OF LIQUID HELIUM

Pressure-relief devices must he adequately sized. Most cryogenic

liquids require considerable heat for evaporation. Liquid helium,

however, has a very low latent heat of vaporization. Consequently, it evaporates rapidly when heat is introduced, or when liquid helium is first transferred into warm or partially cooled equipment. Failure of the container vacuum can occur, adding appreciable heat. Pressurerelief devices for liquid helium equipment must, therefore, be of adequate capacity to release helium vapor resulting from such heat inputs,

and thus prevent excessive pressure hazards. A quantitative method

for determining the required relief device capacity for liquid helium

containers is discussed in section 6.4.4. Keep equipment area well ventilated. Although helium is non-toxic,

it can cause asphyxiation in a confined area without adequate ventilation. Any atmosphere which does not contain enough oxygen for breathing can cause dizziness, unconsciousness, or even death. Helium, being

colorless, odorless, and tasteless, cannot be detected by the human

senses* and will be inhaled normally as if it were air. Without adequate

ventilation, the helium will displace the normal air and give no warning

that a non-life-supporting atmosphere is present. Store liquid containers

in well-ventilated areas.

6.3. HELIUM LIQUEFACTION

6.3.1. Safety in Liquefier Design

Liquefier and refrigerator design has been covered in chapters 3

and 4. Safety features to be considered and incorporated in the design

and operation of liquefaction equipment (because of the low temperature)

have already been enumerated. In addition to these requirements, the

design requirement for pressure vessels as outlined by ASME Codes [2] should be met. The usual safety considerations associated with compressed gases and machinery vibrations must also be incorporated in the design [3, 4]. The ever-present possibility of plugging due to impurities and the

limitation on the length of operating periods imposed by removal of these impurities must be kept in mind during the design of liquefiers.

Proper venting and incorporation of relief devices for traps are mandatory. Adsorbent traps must be designed to prevent adsorbent dust from

entering the system and to contain the adsorbents even under the

*The pitch of the human voice is raised by high helium content, but this should not be used to gauge safe environ- ments.
SAFETY IN LIQUEFIER DESIGN 259

sudden depressurization due to emergency pressure relief. An adsorbent

that generates minimum dust should be chosen.

Pressurized gas when cooled to low temperatures must be treated

as liquid, and pressure relief devices (or expansion volumes) must

be provided to cope with the pressure buildup caused by warming. For

the same reason, all piping that passes into low-temperature regions

and that may be isolated by valving must have a safety relief device

or expansion volume.

Vacuum-insulated enclosures should be provided with safety-relief

devices on the vacuum space, even if they are continuously pumped.

Some details of relief design will be given in the following section on

transport and storage equipment.

Liquefaction systems incorporating hehum recovery from cryostats

require carefully designed purification systems. In addition, if oil-sealed gas holders are employed, care must be taken to prevent discharge of

oil into the system.

The following philosophy of Dr. A. J. Croft [5] was employed in the

design of a Uquid hydrogen facility and might well be adopted in design

and operation of other cryogenic installations.

Any plant which embodies compressed gases or low-temperature liquids may explode from various physical causes, including faulty design and faulty operation. Where hydrogen

is involved, there is the added possibility of chemical explosion. These dangers have to be thoroughly guarded against and a balanced approach to them is one of the hardest

tasks for a designer of a hquefaction system. It is very easy to obscure real dangers by a mass of impressive precautions against improbable ones. An order of priority for looking into the safety of a system may be made out as follows:

(a) In principle, one could eliminate danger altogether by paying enough attention

to the design, materials, workmanship, and testing which go into a liquefaction plant. This is the first and most important line of defense.

(b) Next, still assuming a perfect system, one must look for every possible error on

the part of the operator and allow for it if an accident might result.

(c) Then one must assume that either or both of these hues of defense have failed and that an accident occurs. The plant must be designed and precautions taken so as to reduce as far as possible any hkelihood of injury to the operator or other people or serious damage to the building.

In particular, the last injunction — which can be paraphrased as "fail-safe design" — should be strictly followed in helium facilities even though

the inert nature of helium renders it far safer than the other common

cryogens. Indeed, a real concern with the design of helium apparatus

is the inherent safety of its use, which can easily lead to complacency

and possibly inadequate attention to safety.
260 SAFE AND EFFICIENT USE OF LIQUID HELIUM

6.3.2. Safety in Liquefier Operation

The general safety procedures for a liquefier parallel those of a chemical plant [6]. The moderate pressures in helium liquefiers and the

chemical inactivity of helium can lead to complacency. However, the

possibility of frozen impurities causing plugs in low-temperature piping

and heat exchangers is reason enough to guard against this complacency.

Unattended operation of liquefiers should be avoided because of the

great hazard due to unobserved contamination during periods when

pressures may inadvertently fall below atmospheric pressure. The

principal source of danger is the inherent time-limitation on operation

due to the impurities and the need to remove the impurities in limitedcapacity traps for proper operation. Care must be exercised in venting

traps which may contain hydrogen, oxygen, and hydrocarbons in mixture

ratios which are dangerous. Although traps are sometimes provided at different temperature levels in an attempt to selectively remove oxygen

and hydrogen, this selectivity is difficult between oxygen and the hydrocarbons, which liquefy at nearly the same temperature. Vent lines from

various traps should not be manifolded, and each trap should be provided

with a safety device, preferably a blowout disk. Traps should be purged

separately unless a very high dilution is used to ensure low contaminant

levels in the purge gas. Helium gas must be used for low-temperature

purges. It is desirable that helium be employed for high-temperature

purges as well, since this will eliminate the possibility of freezeup during

cooldown. If nitrogen is used for high-temperature purge and thaw, then

a further helium purge is required before cooldown.

The safety procedures for transferring liquid helium will be covered

in a subsequent section of this chapter. One condition however, which

is usually confined to the liquefier, is the cooling of warm containers

prior to fiUing. In order to conserve liquid helium, the containers are usually precooled with liquid nitrogen and then purged with helium gas. There is danger that a residue of nitrogen will be frozen in the bottom

of the container and subsequently plug the transfer fines or possibly

the apparatus into which the helium is transferred. A safer procedure

is to precool with gaseous helium which has been cooled with liquid

nitrogen, thus eliminating free nitrogen from the helium compartment

altogether. The next best procedure is to employ gaseous nitrogen only,

although this is difficult in practice.
CODES AND REGULATIONS 261

6.4. TRAxNSPORT AND STORAGE

Usucilly, the "use point" for helium is physically separated from the

Hquefier, sometimes only by a hallway but at other times by a continent

or an ocean. This factor, coupled with the characteristics of helium which

result in a sizable loss of the available liquid when removed from a

storage vessel, has led to combining transport and storage into a single

piece of equipment. Such equipment ranges in capacity from a tiny

flask of a liter or two to a large semitrailer containing 40,000 1. Some of the small equipment (up to about 25-1 capacity) is fabricated

from copper using techniques described by Wexler [7]. Larger equipment

is usually fabricated using applicable provisions from the ASME Code

or Department of Transportation Regulations and Specifications. To

insure safe and economical storage and transport of liquid helium, close

attention must be paid to equipment design.

6.4.1. Codes and Regulations

Until recently most domestic liquid helium shipments were made

with a very slight positive pressure in the container, up to about 3 psig.

Recent trends to bulk liquid helium distribution from large commercial

liquefaction sources, as described by Nubel and Hover [8], have led to shipment of pressurized liquid helium. In such shipments the initial pressure is quite low (about 1 psig), and the heat leak to the helium

during transportation is permitted to raise the internal energy rather

than evaporate liquid. Thus the pressure increases during shipment,

sometimes well in excess of the 25 psig (40 psia) value which places

the shipment under regulatory control of the Department of Transportation [9] (DOT). Such regulatory control is independent of the quantity

shipped. Ultimately, equipment used in such shipments must be de- signed, fabricated, and tested according to approved DOT Specifications; there are additional DOT Regulations which govern the conditions

under which such DOT Specification Containers may be shipped. Since

commercial pressurized liquid helium distribution is a new operation,

the Compressed Gas Association is working with the DOT to obtain

formal approval of the Specifications and Regulations [10]. For both cargo and portable tanks, the liquid helium container

must meet the requirements of the ASME Code for design, materials,

fabrication procedure, and test. Until such time as specific Regulations

and Specifications are published by DOT, pressurized liquid helium
262 SAFE AND EFFICIENT USE OF LIQUID HELIUM

shipments above 25 psig will be handled by Special Permit from DOT.

Each shipper must obtain such a permit.

While most of the transport and storage containers having less than

1000-lb water capacity are operated at pressures so low that existing

DOT and ASME rules do not apply, DOT Specification 4L, Paragraph

78.57 [9] can be used as a guide to the types of material specifications,

joining procedure requirements, tests, and product identification which

are desirable. (This Specification includes a formula for calculating

vessel thickness and service pressure, hence does not reference the

ASME Code.)

Air shipment of liquid helium is permitted under conditions defined

by the International Air Transport Association (lATA) [11]. Liquid

helium is not accepted for shipment on passenger aircraft, but when

properly packaged and labeled is shipped on cargo aircraft. Nonpressurized liquid helium (i.e., in containers open to the atmosphere) is never accepted because of the solid air plugging hazard discussed in chapter 5. Present regulations permit shipment of up to 500 kg (4000 1) of low-pressure liquid helium ( < 25 psig), and up to 140 kg (1120 1) of pressurized liquid helium ( > 25 psig). Section 5.7.5. discusses the

characteristics of relief devices and absolute pressure relief valves

in the air transport environment, and summarizes current lATA requirements for venting of helium containers. These requirements are of a

qualitative nature. Recommendations for quantitative relief-device

sizing procedures are summarized later in this chapter. lATA does not require use of containers fabricated to specification, but it is strongly

recommended that the specification guides described above be used in designing equipment for air transportation service.

6.4.2. Materials and Methods of Fabrication

Hansen [12] and Zenner [13] have emphasized the need to avoid

low-temperature embrittlement in selecting suitable materials for fabrication of equipment for the low-temperature fluids. The ASME Code at present requires impact tests for SA-240 stainless steel used for liquid

helium service; such tests are not required for some wrought aluminum

alloys. Many other nonferrous materials are no doubt suitable for use

in liquid hehum equipment, but the ASME Code requires the designers

to conduct suitable tests to satisfy the ductility requirements. Data on

other materials are becoming available through extensive work done

by various Government agencies [2, 14, 15].
INSULATION 263

While many materials will eventually prove suitable for use in liquid helium equipment, to date most existing equipment employs

SA-240 Type 304 or 304L stainless steel, or SB-209 Alloy 5083 aluminum

for material exposed to these extremely low temperatures. In addition

to the base metal considerations, close attention must be paid to weld

procedures, requirements for which are outlined in Section IX of the

ASME Code, in order to achieve safe and reasonable equipment. In- adequate welds are not only hazardous from a strength viewpoint, but

may also make it impossible to achieve the high vacuum required for

efficient insulation of the container.

6.4.3. Insulation

Thermal insulations have been described in detail in chapter 5.

It is obvious that careful design and application must be achieved to attain reasonable performance, whether the insulation system is based

on liquid nitrogen shielding, cold gaseous helium shielding (single or

multishield), or multilayer insulation. From a safety standpoint, the

characteristics of the insulation system are tremendously important.

Vacuum failure may occur in either the outer jacket or inner container.

If a flaw should develop in the vacuum jacket, then air would enter

the vacuum space. Any exposed metal surface which is cooled by liquid

hehum will serve as a condensation point for liquid or sohd air, adding

to the vessel heat flux. Similarly, a helium leak will flood the insulation

space with helium, also leading to much greater heat flux across the

insulation system. If helium gas generated by this increased heat influx

is not removed from the liquid container at a sufficient rate, the pressure

will increase to the burst point. Both types of heat flux behavior in addition to fire exposure should be checked for the specific insulation

system, and the capacity of the conduit and relief devices of the inner

container sized to assure adequate flow. Since gases leaking into the vacuum space can be cryopumped or cryosorbed and thus retained within

the space, it is extremely important to protect the insulation system

with an adequate pressure-relief device. Otherwise the equipment

could be destroyed either by implosion of the inner vessel or by rupture

of the outer vessel when the container is warmed.

Typical heat leak performance for various currently available

designs is summarized in table 6.1. One point needs to be emphasized,

and that is the large difference in heat leak to the inner vessel for venting and nonventing operation, which may vary markedly, depending

306-922 0-68— 18
264 SAFE AND EFFICIENT USE OF LIQUID HELIUM

!
CONDUITS AND PIPING COMPONENTS 265

on the specific type of insulation system. For the small liquid nitrogenshielded dewars, the nonvented heat leak is approximately ten times

the vented heat leak; for multishielded dewars the ratio is over 15 to 1; while for the 40,000-1 semitrailer the increase is only a few percent.

Thus, the method of operation of the equipment should be based on

the overall system requirements. If the end use is primarily in the

Hquid state, a continuously venting system may be more efficient.

6.4.4. Conduits and Piping Components

Liquid helium transport and storage containers are used in a wide

variety of operations; hence many factors can influence their safe and

efficient handling. Small containers of 500 1 or less are designed to provide the user with a great deal of flexibility in removing the product

from the container. A major feature is the neck access tube, which

permits a wide variety of transfer lines to be inserted. This versatility can lead to unsafe conditions if proper operating procedures are not

followed.

Neck plug hazard. A neck plug occurs when air condenses and

freezes in a neck tube or vent line when either is open to the atmosphere.

The pressure in the dewar is now unable to vent and will rise to the point

of rupturing the vessel. The potential hazard is increased with fluctua- tions in barometric or ambient pressure and is also increased while

making product transfers. The dual relief passage is the most effective method of eliminating

neck plug hazards. If a plug occurs in the normal vent line or fill connection, the second relief passage with a higher relief valve setting will prevent excessive pressure rise. This is achieved in most dewars with

a double concentric neck design, as shown in figure 5.20 of chapter 5. From a safety viewpoint, great care must be taken to keep air from

entering the container when inserting and removing transfer lines, to prevent formation of a solid air plug. It is mandatory to close the aperture

to the container immediately after removing the transfer hardware; also,

it is mandatory to minimize the time period when opening for insertion. While raising the tube, air will condense on the cold outer tube, and

unless restricted, will run down the tube into the conduit. A commonly

used technique to prevent this is to provide a suitable wiping ring

between the tube and the conduit opening, such as an absorbent cloth. The cloth will then collect the major portion of condensed air, particularly

if the removal of the tube is done quickly. A more positive method of

preventing air from entering the container is the use of the ball valve
266 SAFE AND EFFICIENT USE OF LIQUID HELIUM

ACCESS I

I PORT^ I

! VACUUM-INSULATED

HOSE OPEN BALL VALVE WITH

DIP TUBE HERE

COOLDOWN POSITION OF DIP TUBE

— LIQUID CONTAINER

LIQUID WITHDRAWAL

POSITION (TIP ABOUT

'/Z'FROM BOTTOM)

Figure 6.L Insertion of transfer hose dip tube.

[16] arrangement shown schematically in figure 6.1. This technique has

the advantage of maintaining a slight positive pressure in the container

at all times, since the O-ring seal to the tube is effected before the baU

valve is opened for insertion of the tube, or vice versa on removal.

Most of the early liquid helium dewars had only a single aperture

to the container, and some near disasters occurred due to plugging of

this path by solid air. For this reason, dual relief paths have come into common use, and most manufactureres now supply containers with

such an arrangement. An adapter is frequently used on single neck tube

dewars to provide two relief paths, as discussed by Scott [17] and

shown in figure 6.2. For larger equipment, it is also recommended that dual relief paths be used. Often other design requirements wiU favor the use of

two separate lines. One of these lines should be sized for the catastrophic

requirement (either fire exposure or air condensation), while the second

line should handle the relief requirement for loss of vacuum with

helium, as described below.
CONDUITS AND PIPING COMPONENTS 267

By combining the operating techniques described above with the use of dual paths to suitable relief devices, the incidence of plugged

containers has dropped to virtually zero. However, careless handling

such as leaving one or both of these paths open to the atmosphere

invites disaster. This is true for containers holding only cold gas as well

as those holding liquid, because air can reach the low-temperature

regions by diffusion, or by a reverse pressure differential caused by

barometric or altitude variation. It is easy to overlook the fact that saturated helium gas at 4.2 K and 1 atm, if isolated by a solid air plug, can build to approximately 300 psi at 60 K, the approximate temperature

at which a solid air plug might break. If a small amount of liquid is present in the container when the air plug is completed, the pressure can build to a much higher level, and thus rupture the container with

extreme hazard to personnel. If a plugged container is suspected, move

it to an isolated area, evacuate the area, erect warning signs, and notify

the suppher immediately.
268 SAFE AND EFFICIENT USE OF LIQUID HELIUM

For those containers having an access path for insertion of transfer

hardware, test for a plug can be made by inserting a V4-in-diameter by

2 ft-long rod into the opening. If the path is plugged an obstruction will be felt. Means should be provided to prevent any inserted object from

falling into the container. Also, should a plug be broken the outrush

of vapor might expel the tool in a hazardous fashion if a method of

limiting its travel is not applied. One method of relieving such a plug

is to place a common crankcase heater into the neck and let it rest on

the plug. Connect the heater to a 110-V circuit and from a distance turn on the heater. The heater will melt the plug and relieve the pressure. The remainder of the plug should then be reamed out, and the contents promptly emptied so the vessel can be returned to the supplier for checking.

Relief devices. To assure adequate sizing of relief devices and conduit paths to the inner container, it is necessary to predict the amount

of heat which might be added to the helium. This becomes a matter of

judgment as to the type of environment to which the insulation system

will be exposed. The loss of vacuum insulation due to helium or air leakage has already been mentioned. For liquid nitrogen-shielded containers, a mechanical failure could admit liquid nitrogen to the insulation space, where it would come in intimate contact with the helium

container. The container could be exposed to a fire, while simultaneously

losing vacuum.

Scott [17] discusses the effect of failure on relief capacity requirements for 25-1 liquid nitrogen-shielded dewars. Unpublished data by

Schiffhauer [18] also indicates that a liquid nitrogen failure mode

required greater relief capacity than vacuum deterioration with helium.

The CGA has a recommended guide to sizing relief devices for cryogenic containers [19]. The CGA equation estimates the total heat

input to a vessel under fire conditions (including the effect of some

insulation deterioration), considers the latent heat of vaporization of

the lading, and converts the required relief rate to Qa in ft^/min of air at 60 F and 1 atm. The relief rate for air is one of the standard methods

for rating relief devices, hence selection of a device with the required

capacity from manufacturers' listings is straightforward.

A relief device should be sized to provide a flow of
CONDUITS AND PIPING COMPONENTS 269

where

:

Q(i = minimum rated flow capacity of the rehef device in ft^/min of air at 60 F and 1 atm,

G,= factor based on characteristics of the gas _ 73.4 X (1200-0 iZT\

LC \M] ' U= total thermal conductance of the container insulating material at 100 F (Btu/hr-ft^-F) when saturated with gaseous nitrogen at atmospheric pressure. (Since the thermal conductivity of heUum is much

greater than that of nitrogen, the value of U should be based on hehum for containers in helium service.)

A = total outside surface area of the container in square feet,

L = latent heat at flow conditions in Btu/lb,

C= constant for gas or vapor related to ratio of specific heats (See ref. [19], fig. 2), Z= compressibility factor at flow conditions,

r= temperature of gas at flow conditions in R,

M= molecular weight of the gas, t= temperature of gas at flow conditions, F. The CGA provides G, values for some cryogenic fluids. The conventional

CGA formula cannot usually be used for calculating G, for helium, since

the design pressure of the transport and storage container is usually

well above helium's critical pressure. However, a d value of 52.5 has

been adopted by the CGA for helium vessels with design pressures

between 35 and 200 psig. This Gj value of 52.5 for helium is based on

calculations and discussion given in Appendix A. The size of a spring-loaded relief valve for loss of vacuum (i.e., when helium gas is in the insulation space) can be determined using

the formula

(?„ = 4.5 iUA),

where Qa and A are in the units above, and U = thermal conductance in —;

^ , ^ of the insulation at 1 atm of helium at 290 R. Where the insulahr-ft--F tion is such that appreciable convection does not exist, U can be es- timated by taking the ratio of the thermal conductivity to the insulation

thickness. The factor 4.5 varies somewhat with pressure, so for safety a conservative value is used. Additional detail for this equation is given

in Appendix A.
270 SAFE AND EFFICIENT USE OF LIQUID HELIUM

Figure 6.3. Estimated total heat flux versus area for air condensation and fire conditions in multilayer (SI) insulated liquid helium containers.

It is quite possible that a flaw could develop in the outer jacket,

admitting air to the insulation space. Depending on the specific insulation

system, it is conceivable that the condensation of this air could add heat

to the helium at a greater rate than that occurring with fire exposure.

To illustrate, for one type of insulation system only, Frainier [20]

showed that air condensation on a liquid hydrogen container insulated

with 1 in of multilayer insulation has a heat flux of 400 Btu/hr-ft^. Figure

6.3 summarizes the total heat flux as a function of liquid container area

for both air condensation and fire exposure. Heat flux due to fire exposure

is plotted for 1, 3, 6, and 12 in of multilayer (SI) insulation, and the

resulting comparison with heat flux due to air condensation for 1 in of

multilayer insulation shows that the fire condition represents the greater

relief device sizing requirement for all cases except large containers

having 6 to 12 in of insulation. For the latter case, it is also likely that the 400 Btu/hr.-ft- would decrease appreciably, though this should be

determined experimentally.

The two dotted lines of figure 6.3 indicate the potential for very

large heat flux due to air condensation on a bare container. It can be
OPERATING ENVIRONMENT 271

concluded that both air condensation and fire exposure should be

carefully evaluated for the specific insulation system when sizing

relief devices.

6.4.5. Operating Environment

In the last few years, the nature of liquid helium transportation has

changed from an experienced scientist gently cradling a small flask while crossing the hallway between the liquefier and his research equipment, to the shipment by air, land, and sea of a wide variety of containers

Figure 6.4. 500-1 liquid helium container.

(Courtesy of Air Reduction Co.).
272 SAFE AND EFFICIENT USE OF LIQUID HELIUM

handled by workers who have never seen liquid helium before and who

have not the slightest notion of its physical characteristics.

This vast change in handling procedure placed a severe challenge on designers of helium transport and storage equipment. At the start of commercial supply of liquid helium in 1959, the only available containers were the liquid nitrogen-shielded copper dewar containers typically

used for laboratory operations. Fortunately, the 25-1 size could be adapted

for transport use by placing the container in a suitable shipping carton. Fiberboard shipping drums with shredded foam packing between the

container and drum, and wooden boxes have been used, along with

various types of shipping frames having rubber shock mounts. Experi- ence in common carrier shipment of these packaged containers has

been satisfactory.

New designs of containers have become available in recent years, and these generally are based on design and fabrication procedures

consistent with DOT Specifications. The smaller containers usually use some form of packaging to reduce the loads imposed in common

carrier handling. Figure 6.4 shows the shock mount arrangement used on a Hofman 500-1 container.

Larger containers such as the Linde 7000-1 Transport are mounted

on semitrailers which provide appreciable cushioning, particularly

when the new air-ride type of shock absorbing suspension is used.

Whenever a container of helium is shipped, make certain that (a) the liquid and gas phase valves are fully closed and capped; (b) all piping connections are tight; (c) the relief devices are properly in place

and in satisfactory operating condition; (d) there are no frost buildups

[which indicates container vacuum or insulation deterioration]; (e) container pressure is within an acceptable range; (f) container decals are properly in place and legible. Label the container with the following

handling precautions and identification notices:

LIQUEFIED GAS CONTAINER

LIQUID HELIUM -LOW PRESSURE (up to 40 psia)

. [or LIQUID HELIUM -PRESSURIZED (greater than 40 psia)] KEEP CONTAINER UPRIGHT

DO NOT DROP

DO NOT BUMP

DO NOT LAY ON ITS SIDE;
OPERATING ENVIRONMENT 273

(g) shipping papers are in order, including: necessary bill of lading,

receiving instructions, delay enroute instructions, and the name and

phone number of person(s) to be contacted in the event of a need of instructions or information; and (h) packaging of the container is pro- vided, according to the needs of the specific container design.

Whenever a container of helium is received, inspect the container

immediately to determine whether there are any breaks or leaks in the

piping, or unusual frost formation. Check the pressure in the container

to be certain that it is in the proper range for the specific container

design, and that the relief device is operating properly. Test the smaller

containers for neck plugs. Higher than normal loss of helium from the

container can occur due to loss of container vacuum, low liquid nitrogen

level (if liquid nitrogen shielded), or thermal oscillations (see chapter 5). Altitude variation. Air shipment of liquid helium is a common occurrence, due to the large distances encountered between liquefier and

customer. Such shipments can cause an appreciable variation in the

atmospheric pressure immediately surrounding the helium container

with typical variations from 8.2 to 12.2 psia at altitude, as pointed out by Van Meerbeke [21] and discussed in chapter 5. Care must be taken

during descent of the aircraft that air does not flow into the container

piping with consequent formation of a solid ice plug in the cold portion

of the conduit.

There are several approaches to solving this problem, the most

common being the use of an absolute pressure-relief device. Another

approach consists of a heat exchanger coil which warms the gaseous

helium prior to reaching the relief valve, thus greatly increasing the

ability of the valve to resist backflow of air when the pressure difference

is reversed. StiU another common approach is to use a conventional

spring-loaded reHef valve with a setting sufficiently high that the container pressure is always greater than atmospheric. A 25-psia setting

or greater is required, since the altitude variation can typically cause

a 6 to 7 psi change in the ambient pressure (i.e., from 14.7 to 8.2 psia).

This procedure has the potential of not releasing any helium from the

container, so no-loss shipments are possible with suitably designed

containers.

While the alternate approaches described above have been used

successfully, the preferred arrangement for air shipment of small containers is the absolute pressure-relief valve. It works equally well with

liquid nitrogen-shielded dewars, gaseous helium-shielded dewars, and

gaseous helium-cooled multishielded dewars. The continuous evapora-
274 SAFE AND EFFICIENT USE OF LIQUID HELIUM

tion permitted by this device also tends to reduce losses, and the pressure

is controlled at the desired set point.

Weather. All transport containers are exposed to the elements,

which can cause problems due to rain, snow, and freezing rain, etc.

It is appropriate to protect the relief devices from these weather effects. This is sometimes accomplished by using weather shields, or protective

housings. By specifying suitable positions of the relief device (i.e., outlet pointed down), adequate protection against moisture penetration and

later freezeup can also be obtained. Frangible disks are usually fitted with a plastic blowofif cap on the outlet side to prevent undue corrosion

with consequent premature failure. The plastic cap must be applied in such manner that it does not impede this relief path.

Transfer operations provide a much sterner challenge to safe and

efficient use of liquid heHum than storage, because of the heat input

experienced at the interfaces of the connecting conduit between the

storage and receiving apparatus, and in the conduit and valves. Cooldown

of the transfer elements and the receiving apparatus also contributes

significant losses, even when precooled with liquid nitrogen. Evolution

of a large amount of helium vapor when making such transfers requires

adequately sized flow passages and relief devices to prevent over- pressurizing the apparatus. Figure 6.5 shows a typical transfer arrange- ment with a 100-1 container supplying liquid helium to a cryosorption

array located within a vacuum chamber. Note the vacuum-insulated

transfer line used to transport the liquid helium from the supply container on the right to the array on the left. (The smaller container in front supplies liquid nitrogen to provide auxiliary refrigeration to minimize liquid helium consumption.)

Proper cooldown and purge of liquid helium equipment are essential

to efficient and safe operations. Potential loss of liquid helium is best

emphasized by examining the loss of liquid helium in cooling 1 lb of

stainless steel, assuming no sensible heat recovery:

6.5. TRANSFER

From 298K to 78 K-14 liters From 78K to 4.2K- 0.67 liters
TRANSFER 275

Figure 6.5. Liquid helium transfer from a Linde vessel (LS He-102) to cryosorption array

in vacuum chamber.

(Courtesy of Linde Division, Union Carbide Corporation). Therefore, the weight of material to be cooled should be kept to a reasonable minimum, and precooling with liquid nitrogen should be

used when feasible. Some principles and cautions to keep in mind in the precool operation are:

(a) Use of liquid nitrogen will greatly reduce the liquid helium cool- down loss, but the liquid nitrogen must be completely removed, fol- lowed by a helium gas purge. If not removed, the nitrogen will freeze

at helium temperature with consequent operational problems and safety

hazard due to plugged lines.

(b) Saturated nitrogen vapor at 139 R (77 K) in some cases may

prove to be a good alternate cooldown medium since it avoids the need

for liquid nitrogen removal.

(c) In some cases, particularly where a gaseous helium recirculation

system is available, it may be advantageous to precool using helium

gas cooled to 78 K by liquid nitrogen in a heat exchanger. This limits the purge operation to one initial step, and avoids any warmup which

might occur when a gaseous helium purge follows a nitrogen cooldown

operation.

(d) A trace system employing liquid nitrogen or cold gaseous

nitrogen might be used to advantage, especially where intermittent
276 SAFE AND EFFICIENT USE OF LIQUID HELIUM

100

UPPER LIMIT

4 8 12 TRANSFER TIME, HOURS

Figure 6.6. Helium loss when transferring from a lOO-l supply container into a receiver

vessel; transfer components have an assumed 15 Btu/hr heat flux, equivalent to a loss of

6.2 Uhr.

Dotted line — indicates maximum recovery at transfer rates from 6.2 1/hr and up, with no cooldown of transfer line or

receiver. X - Experimental net recovery of 70 1 = 100 — (.4 + — Experimental net recovery of 43 1 = 100 — (A-\-C). A — Residual vapor in container. B — Accumulated losses due to single cooldown of transfer line, and three 25-1 containers from 78 K, plus steady

heat leak to containers and transfer line. C — Accumulated losses due to 14 repetitive transfers to a cold 3-1 receiver with transfer line cooldown and

steady heat leak.

transfers are required. The nitrogen should be removed and gaseous

hehum purged prior to Hquid hehuni addition.

Chapter 5 discussed in considerable detail the many factors in- volved in the transfer of liquid helium including their effect on the loss of liquid helium during the process. Figure 6.6 illustrates some of the

problems involved in the efficient use of liquid helium in typical opera-
TRANSFER 277

dons. In this highly simpHfied transfer, conventional vacuum-insulated

components, having a steady-state heat leak of 15 Btu/hr are used to provide a flow path from a 100-1 supply container to the receiving vessel. For the purposes of this illustration, it is assumed that over the limited

time period for the operation there is zero-heat leak to the supply and

receiving vessels, no pressurizing or depressurizing effects (pressure

remaining at 14.7 psia), and no cooldown loss in either the transfer

components or the receiving vessel. The dotted line shows the upper

limit of liquid helium captured in the receiving vessel for various rates

of transfer (based on the amount of hquid hehum leaving the supply

container).

Also shown on figure 6.6 are the initial contents of the 100-1 supply

container, the 13.4 1 which remain in the container as saturated vapor

at 14.7 psia, and the loss due to the 15 Btu/hr steady-state heat leak to the transfer components for the total time of transfer. The two crosses represent results obtained in several typical operations. The upper

cross shows that 70 1 are recovered in the sequential fiU of three 25-1 nitrogen-shielded containers. In this operation, the transfer line is cooled only once, requiring a little less than 3 1, and three copper compartments are cooled from 78 K to 4.2 K, requiring about 8 1. For these

low-pressure transfers (1 to 2 psig), depressurization and Hne losses

described in chapter 5 are small compared to the precoohng losses,

and therefore are not indicated separately in these typical results. The lower cross represents the loss incurred when a series of 14 refills are made to a relatively small receiving vessel in which the

transfer components must be recooled for each transfer. This 3-1 cool- down loss per transfer adds up to a sizable overall loss for such a recurring operation.

Generally speaking, hquid helium transfers should be made in a reasonably short time period, measured in minutes. Continuous transfers over a few hours should be avoided. The conventional transfer hoses,

valves, and bayonets fit in well with short-term transfers, and have been

designed for ruggedness and good adaptability at reasonable cost. When liquid helium must be transferred over a long time period or repeatedly transferred wherein the components warm up between

transfers, design of special components should be considered.

The addition of necessary safety devices to containers and transfer

lines may sometimes be detrimental to the thermal performance since

they may cause thermal oscillations (chapter 5) in the piping which can
278 SAFE AND EFFICIENT USE OF LIQUID HELIUM

WITHDRAWAL HOSE

Figure 6.7. Experimental arrangement

used to avoid thermal oscillations. contribute excessive amounts of heat. Short [22] describes a thermal

oscillation damper for cryogenic containers. This device has been very

useful in eliminating thermal oscillations in containers having concentric

dual rehef passages. Shal'nikov [23] describes the use of a short thread

at the base of a withdrawal tube to remove thermal oscillations. A technique for removing thermal oscillations, which sometimes

occur in the insulated withdrawal tube path during periods of nonwithdrawal when retained in the container, is to provide a short, small- diameter jumper line between the liquid withdrawal tube outlet and the gas phase of the container as shown in figure 6.7. Both the liquid phase

valve and the gas phase valve are then cracked or opened as required

to remove the oscillations. The normal evaporation of helium from

the container passes through the relief device provided for that purpose.

Using this technique, overall loss from a container with the line inserted

is only slightly greater than the loss from a container without the line inserted. This method of operation reduces the loss incurred in recooling

the parts inserted into the container. The most efficient operation is therefore dependent on the frequency of withdrawal. Similar techniques can be used on transfer hardware that does not contain the liquid-line valve.
GENERAL DESIGN CONSIDERATIONS 279

6.6. SAFETY IN CRYOSTAT DESIGN AND USE

The principles already elucidated for liquefaction, storage, and

transport equipment should be adhered to in cryostat design. Safe

operation is usually dependent upon safe design. The small-scale re- search apparatus holding only a few milliliters of hquid helium may

be a bomb in disguise if the liquid is trapped. Therefore, this apparatus

needs a safety review during the design as surely as does the largevolume apparatus holding hundreds of Uters. Once again, however, firm

recommendations are difficult, but some pointed reminders can be

enumerated.

Standard procedures for laboratory safety are a required step

in the safe operation [24, 25]. In addition to the usual considerations of

safe laboratory practice with experimental apparatus, those intended

for cryogenic service should have critical reviews in two areas: first, the materials of construction should be reviewed to ensure that materials

are employed properly and that allowance has been made for the effect of temperature gradients on dissimilar materials, during cooldown,

steady-state, and warmup; and second, the venting of the vapors evolved

during all phases of operation should be determined to ensure adequate

size of vent apertures and safety-rehef valves and proper installation of vent lines. Obviously, if the apparatus employs Uquid nitrogen as well

as helium, the review should encompass this cryogen as well.

^ 6.6.1. General Design Considerations

I

I As a first step toward the design of safe apparatus, the ASME

I Boiler and Pressure Vessel Code [2] should be consulted. The various

j sections of this code provide information on material selection, joints,

I and formulae with safety factors for the design of a limited number of

equipment configurations. The research equipment designer may find

that his experimental requirements are in conflict with the codes and

that his design must be non-code. These codes then give him a reference

point from which he can gage the extent of his departure from sound

engineering design practice. More importantly, they allow him to employ

j proven engineering practice for that portion of his apparatus which

! can be designed to code specifications.

I The design of cryostats for helium service is covered in some

detail by White [26] and by Rose-Innes [27] and in somewhat less detail

in some parts of Hoare, Jackson, and Kurti [28] and by MacKinnon [29], Scott [30], and Din and Cockett [31]. Numerous designs of specific

I 306-922 0-68—19

II
280 SAFE AND EFFICIENT USE OF LIQUID HELIUM

apparatus and design tips are reported in Review of Scientific Instru- ments, Journal of Scientific Instruments, Pribory i Tekhinka Eksperimenta, and Cryogenics, as well as other journals. Few of the authors

give specific reference to safety features, although safety is implicit

in many of the instructions.

The apparatus designer is frequently left to his own counsel or to locally prevailing procedures, either of which may or may not be adequate. In particular, as apparatus evolves and becomes more sophisticated, some simple, very effective safety aspects are sometimes

designed out of the apparatus. When any experimental apparatus

undergoes evolutionary development, a special check of safety features

is in order, since safety is likely to be overlooked in the search for functional improvement.

The general desire in cryostat design is for an isothermal system

at some specified temperature level having a minimum heat leak. These

considerations usually lead to deep, narrow vessels which are of either

glass or metal construction with high-vacuum thermal insulation. There

are usually closed experimental spaces within these vessels which are suspended from a top cover by long, small cross-section support members, either tubular or solid. Frequently, these members, if tubular,

form the principal opening to a boiling cyrogen and may represent a reasonable obstruction to the discharge of vapors from the system. In

addition, sometimes there are radiation shields within or around the

supports, which further obstruct the discharge path, particularly in the case of emergency venting. If the experimental requirements are such that venting may not be adequate, then the apparatus should be

placed behind a shield to protect personnel in case of misadventure.

The designer should consult some of the standard works on equipment design [2, 32] and on methods of joining metals [33, 34, 35] in order

to assess the strength and suitability of the particular joints to be used.

When the designer employs a new technique of his own it must be

thoroughly tested to insure that it will not contribute to or create an

unsafe situation. An admonition by Dr. Croft [5] relating to metal apparatus, that "nothing is safe for use until it has been hydraulically tested with water,

oil, carbon tetrachloride, or some other suitable liquid," should be

adhered to, although carbon tetrachloride is not a recommended testing

liquid.

A final reminder is that the properties of materials found in many

data books are measured values of a particular specimen or small
MECHANICAL CONSIDERATIONS 281

numbers of specimens and do not, therefore, represent the variabihty

in properties which may be encountered In practice. A safety factor

must be used, and then the apparatus should be tested at above the

working stress.

6.6.2. Mechanical Considerations

The configuration of components which are subjected to low

temperatures and those which traverse large temperature gradients

must be reviewed carefully. The forces resulting from thermal contraction must be considered, since the thermal contraction of most metals

when cooled to or below liquid nitrogen temperature is sufficient to produce forces greater than the yield strength of the metal, if the member is restrained to its room temperature length. For example, consider

a 304 stainless steel rod or tube restrained to its room temperature

length and cooled to liquid nitrogen temperature [14, 36, 37]. The thermal

contraction is L(139R)-L(528R) , ^ . , ^ ' -285 X 10-- in/in. \ 1(528 R)

The modulus of elasticity of 304 stainless at 139 R is F=29X 106 psi^

so that the resulting stress is

/ L(139 R)-L(528 R) \ contraction- Y

L(528 R) / = - (29 X 10^) (- 285 X 10-5) 82,500 psi

while the yield strength of annealed 304 stainless ranges from 35,000

to 68.000 psi.

Other metals, such as 410 and 17-4 PH stainless, have sufficiently

high yield strength that they can be restrained without yielding, but

unfortunately they become brittle and should not be used in tension.

There are, however, some metals, such as A-286 stainless and Invar,
282 SAFE AND EFFICIENT USE OF LIQUID HELIUM

whose properties are such that they also can be restrained without

exceeding their yield strengths, and they are not brittle. If the member is in a thermal gradient, then the contraction will be less than for full cooling, and the stress level will be correspondingly

lower, in many cases low enough to be safe. However, the designer

must check to be sure. This determination requires knowledge of the

thermal gradient in the member in order to arrive at an estimate of the

length. The elasticity modulus is nearly constant with temperature, so

that a single value can be used to determine the stress level once the

contraction is known; this stress must then be compared to the lowest

yield strength in the temperature range in question, in order to assess

the design.

Dissimilar metals in a thermal gradient may also give rise to forces

which will exceed the yield strength of the metals or which will distort the apparatus, thus increasing the heat leak or breaking other supports

or instrumentation connections. Even when dissimilar metals are avoided or designed for, problems may arise during cooldown or emer- gency venting due to the asymmetrical thermal gradients which can

exist if only one of several parallel tubes is used as a vent. Gaskets

held by bolted flanges may loosen due to differential contraction of

the components, i.e., bolts, flanges, and gaskets. Safe practice dictates

that some form of elastic member be employed to ensure sealing.

Support members which are hollow and sealed must be used with

caution, since small leaks into the sealed volume may result in a con- siderable volume of gas being condensed at the low temperature region

with subsequent high pressure on warmup. Where slow warmup is expected because of massive components, small vent holes may suffice

to release pressure, but for rapid warmup restricted volumes should

be avoided.

6.6.3. Glass Apparatus

For many years the most common equipment for handling the

lower bx)ihng cryogens was glass apparatus, and even today many lab- oratories find it convenient and economical to employ glass in the construction of cryostats. Rose-Innes [27] discusses a typical glass cryostat

and its use in some detail. Safe practice dictates that large pieces of

glass be coated with a plastic film or taped with a plastic, cloth, or paper

tape to restrain the fragments in event of breakage. While glass dewars

usually implode upon breaking, as a result of the insulating vacuum
GLASS APPARATUS 283

in the jacket, they should nevertheless be protected, either by the above means or by placing them in a metal or plastic enclosure. When clear

strips are left in glass dewars for observation of the liquid level, some

protective cover between the observer and the strip should be provided.

Large glass apparatus should be annealed in a furnace before

installation. Glass tubing that must be installed at the apparatus should

be carefully annealed in situ to remove internal strain which may cause unexpected failure during use. Inspection with polarized light is recom- mended. If annealing is not possible, then connections should be made

by flexible plastic tubing, rubber, or by ground glass joints.

Dewars which are moved frequently should be handled with extreme

care as they may get scratched and break unexpectedly, since they are always under stress due to the vacuum. The sealoff tips should have a protective cover to guard against accidental bumping.

Most hard glasses such as Pyrex are quite permeable to gaseous

helium at room temperature and above. The permeation rate is suffi- ciently high that, if warm dewars are left with hehum in them, the vacuum deteriorates in a very short time and permanently sealed

dewars are rendered useless. The common practice is to provide a pumpout connection at the top of the dewar and to evacuate them

periodically. White [26] reports that dewars may be used some 100 hr

or so between evacuations. The vacuum achieved need not be very high,

since the heHum-cooled inner sleeve of the dewar will cryopump the

residual gases other than hehum. In fact, many experimenters intentionally introduce a small quantity of air into the vacuum in order to enhance the higher temperature heat transfer and speed the cooldown.

As indicated by both White and Rose-Innes, the connection between

the dewar and the pumpout tap is best made with a short piece of plastic

or rubber tubing. This serves as a safety relief as well as a flexible joint.

If the dewar vacuum is permanently fused to a glass system, some form

of blowout rehef device should be provided between the dewar and the

evacuation tap.

The common practice of attaching glass dewars to the dewar cap

by a rubber sleeve provides very effective overpressure relief, since the

thin rubber can sustain only sHght internal pressures without ballooning.

If, however, the glass dewars are sealed to the cap with O-rings or by

the use of glass pipe flanges, then a safety rehef device should be

provided in the dewar cap, since these seals can sustain considerable

internal pressure.
284 SAFE AND EFFICIENT USE OF LIQUID HELIUM

Glass manometers are frequently used in cryogenic apparatus for pressure measurement and control. They should always be constructed

with catch basins at the ends to prevent fluid from being discharged

into the apparatus or the room. Care should be exercised to avoid rapid

pressure changes in systems with mercury manometers because of the

considerable momentum of a rapidly moving column of mercury. As

with all glass apparatus, safety shields should be employed.

6.6.4. Fill and Vent Procedures

Economical and safe transfer of Hquid helium is in part a function

of the design of the cryostat. In most transfer operations the total losses can be reduced by carrying out the transfer at low pressures and at moderate flow rates. There is a clear distinction, however, between

filling an open volume such as a transport container and filling an

experimental cryostat where the free volume for liquid is small and the

surface of the included apparatus is large. There is a further distinction

between the initial fill of any apparatus and the refill operation.

For initial fill, the cryostat should be designed to allow the vapors

from the transferred liquid to pass over the apparatus to be cooled, in order to utilize the sensible refrigeration in the vapors and to warm the

vapors up before discharge from the cryostat. The sensible heat to warm

the helium vapor to room temperature is over 75 times the latent heat

of vaporization. The recovery of this refrigeration is accompanied by

an increase in the vapor volume, and the principal hazard in transfer

is the large expansion of the cold vapor or liquid when warmed by the

apparatus in the cryostat. The principal losses in transfer result from

the entrainment of liquid in this large volume of vapor and the failure | to recover the refrigeration in the effluent vapors. Thus, for the initial

fill into a cryostat, the liquid should be introduced at the bottom of the

liquid container and in a relatively open volume which will allow low \ vapor velocities to aid disengagement of the vapor and liquid phases | during the early stages of transfer. ! During refill of the liquid helium, the apparatus cooling problems

do not exist, but there is an entrainment problem resulting from vapors I which accompany the transferred hehum. If the transfer line on refill ! is allowed to go to the bottom of the cryostat, this entrainment can be { very large, mainly resulting from first rush of vapor that is produced

in a warm transfer line. The entrainment can be greatly reduced by

precooling the transfer fine prior to insertion into the cryostat and then
FILL AND VENT PROCEDURES 285

i further reduced by inserting the transfer Hne only as far as the antici- pated filled helium level, so that none of the two-phase liquid is dis- charged below the liquid level. Sizing the vent opening in a cryostat is subject to some variability.

For example, when large masses of metal must be cooled, the largest

I vapor evolution may occur during cooldown, and the amount of vapor will

I not necessarily be related to the liquid volume contained when the appaI ratus is full. Conversely, in some situations (such as cryostats for

I superconducting magnets), there may be considerable evolution during

j cooldown and, in addition, periods of high evolution during magnetic

\ field discharge — both accidental and intentional. In other cases, the vent

I need handle only the small volume resulting from the usual low heat

il leak to well-designed cryostats.

One aspect of venting which should not be overlooked is warmup

after an experiment is completed, since the 4.2 K saturated vapor in an "empty" cryostat will expand to about 100 times its volume at room

temperature and thus could result in large pressure buildup, if confined.

In laboratories where gaseous helium is recovered, there is no problem,

since the cryostats are connected to a gas holder, but where helium is

j vented to the atmosphere a problem can develop. While it is fairly com- mon practice to employ a narrow vent tube with a small wad of cotton

or glass wool to prevent back diffusion of air, this is not a safe procedure,

since barometric pressure changes can cause air transport into the

j cryostat. A positive relief valve is preferred, and it should be so ar- ranged that the vapor passing through it is sufficiently warm that no

' atmospheric moisture can freeze on the valve and impair its opening

or closing. The use of a slit rubber tube or "Bunsen" valve is not recomI mended, since the slits are quite prone to freezing open due to the

( rubber becoming stiff when only slightly cold gas is passing through the

valve. They are totally unacceptable as emergency relief valves. The rate of vapor evolution is dependent upon the geometric

I arrangement and construction of the apparatus and the properties of

i the fluid. The heat transfer rate to boiling cryogens has been determined

for simple geometries by a number of investigators, and a recent review

1 by Brentari and Smith [39] provides data for use in equipment design,

j The data indicate that the peak nucleate pool boiling heat transfer in

j helium is about 1 W/cm^ of surface area, which corresponds to an

I evaporation rate of 1.41 liquid liters per hour per square centimeter

1 or about 10 1/hr of vapor at helium temperature, which could expand to 1000 1 at room temperature, if an additional 80 W are available to warm
286 SAFE AND EFFICIENT USE OF LIQUID HELIUM

the vapors to room temperature. An object 10 cm in diameter by 10 cm

long could thus produce an evaporation rate of 440 liquid liters per hour.

A liquid container of these dimensions would contain 785 cm^ and would

be emptied in about 6 s if boiled at the peak nucleate flux rate producing

about 5.5 1 of vapor at helium temperature and about 550 1 if warmed

to room temperature. The warming of the vapor would require about

160 X 103 J (25kW for 6 s) compared to about 2 X 10^ J (315 W for 6 s) needed for vaporization. Since in most apparatus this amount of heat

is not available, emergency venting invariably produces cold vapor, and

the design of vents must be suitable for cold vapors.

6.7. SUMMARY

The uses of liquid helium are extremely diverse, making it difficult to provide specific recommendations, but the following general principles will be helpful in planning safe and efficient operations:

a. Select an adequate thermal isolation system for the supply and

receiving vessel, and transfer components.

b. Make certain that the degree of vacuum in the thermal isolation

system is such that adequate performance wiU be achieved.

c. Consider precooling and purge requirements in the design of equipment, and use materials compatible with the operating temperature.

d. Where feasible, use adequate helium-content-monitoring devices.

The low density of liquid helium, coupled with a varying liquid

nitrogen weight, can completely mask the true helium content. Point sensors, or diff'erential pressure gages, have been used

successfully. If liquid helium equipment is overfilled, very high

losses result.

e. Slowly insert transfer hardware, sensor probes, etc., to permit

coohng with hehum vapor rather than liquid helium. The energy

input with the transfer hardware may be used to build the

pressure required for transfer. Fast insertion will result in high

vapor discharge accompaned by Undesirable hquid entrainment.

f. When reducing the pressure of the liquid helium container, bleed

off slowly from the gas phase to avoid liquid entrainment.

g. Use sufficient pressure difference to effect the transfer in a reasonably

short time. However, use the lowest pressure consistent with

the time requirement to minimize added losses due to pres-
SUMMARY 287

surization (through addition of gaseous hehum to maintain re- quired pressure), transfer loss, pressure drop and depressurization, as discussed by Van Meerbeke [21]. For many operations,

0.5 to 1.0 psig is sufficient, though other operations require

3 psig or more.

h. Design equipment with adequate vapor vent paths and suitable

safety devices. In sizing relief devices, include the additional

heat input which can result from loss of vacuum, exposure to

fire, air condensation, and when applicable, the energy input

potential from other system characteristics such as that due to a cryogenic magnet going normal, etc.

i. Avoid transfer arrangements which permit gaseous helium to bubble

through a pool of Uquid.

j. Where periodic refills are required, some means of preventing

entry of warm gaseous helium into the Uquid hehum container

is desirable. This can be accomplished in a variety of ways,

including use of liquid-nitrogen shielding and venting of gaseous

helium from a location near the receiving vessel until the transfer

Une is cold.
288 SAFE AND EFFICIENT USE OF LIQUID HELIUM

APPENDIX A. SIZING OF RELIEF DEVICES

The Compressed Gas Association [19] provides information for sizing of relief devices for compressed gases, including pressurized

liquid oxygen, nitrogen, and argon and liquefied hydrogen, in cargo,

portable, and storage tanks. A value of d (the parameter used to size

relief devices) is listed for each commodity and is dependent on the

latent heat of vaporization and other physical properties of the specific

fluid.Sizing of relief devices for liquid helium containers is somewhat

more complex, since the relieving pressure is usually above the critical pressure where no latent heat value exists. A method for handling

helium is presented herein, a value being computed in typical fashion,

but based on L', an equivalent value used in place of the latent heat of

vaporization. The units of L' are Btu/lb, where Btu's represent the

total heat added to the vapor between temperatures Ti and Tr per

pound of efflux gas leaving the container at constant pressure, P. The L' varies significantly with Tx, and has its lowest value shortly after start of relief.

Sizing for Fire Condition

The basic equation for sizing relief devices is

where Qa = scfm of air at 520 R (289 K) and 14.7 psia, required capacity

of the relief device

W= Ib/hr, required flow of the gas Ca= Gas constant = 356 for air C= Gas constant of the gas Za = Compressibility factor = 1.0 for air at STP

Z= Compressibility factor for the gas at the flow condition

Ma = Molecular weight of air = 28.97

M= Molecular weight of the gas Ta = Temperature of air at STP= 520 R

T= Temperature of the gas at the flow condition.
APPENDIX A. SIZING OF RELIEF DEVICES 289

Also,

(6A-2)

where H= Heat input in Btu/hr/ft^

A = Total outside area of the liquid container in ft^ L = Latent heat of gas in Btu/lb.

Source information for this earlier work [19] has shown that as the

size of the container increases the fraction of the surface exposed to the open fire decreases, hence ^0.82 jg ugeJ rather than A, and HA^^'^

then represents the total heat passing to the fluid in the container under

fire conditions. For insulated containers.

where U= Total thermal conductance of the container insulation system,

Btu/hr/ff^ F. Substituting (6A-2) and (6A-3) for W in (6A-1), along with the other

values defined.

HA^-^'' = kA^-^Ht2-t)/l, in Btu/hr,

where /c= Thermal conductivity of the insulation, Btu/hr ft F

r2= 1200F

^= Temperature of the gas at flow condition, F

/= thickness of insulation, feet.

This can be expressed as

HA^-^'' = UA^-^Ht2-t), in Btu/hr, (6A-3)

13.1 UA'^-^HUOO-t) 356 / ZT 28.97 \ 60 L C W(1.0)520;

18.3 UA''-^Hl200-t)

LC

{ZTIMyi\ (6A-4)

A factor G, is defined as

G,=

73.4 (1200-0

LC

{ZTIM)"\ (6A-5)
290 SAFE AND EFFICIENT USE OF LIQUID HELIUM

Table 6A-1. Gj values for helium, for insulated containers

(Calculated from: Gj = 0.097 ( 1200 -tx)VZ]rlyi')

Relief pressure psia mi Vi A, Vx tx

Btu

lb lb 'R

lb

lb

lb °R °F lOOX 1.0 0.1259 10 7.66 0.793 0.1589 12 — 448

.1589 12 10.92 .67 .2368 14 — 446

.2368 14 15.71 .742 .3200 16 — 444

.3200 16 19.91 .81 .3948 18 -442

.3948 18 23.47 .85 .4642 20 -440

.4652 20 26.72 .78 .5940 24 — 436

.5940 24 32.70 .765 .1165 30 — 430

.7765 on30 A 1 AO 41.Oo TOO l.UOOl /I A40 — 420

1.0661 /I A40 C A OO 54, 3o TAO 1 O /I TO 1.34/3 C A50 — 410

1 OA no dU AT O 1 0/.31 Tl 1 QOO 7n- OAA

TH z. / Id 1 nn — OOK)

£. i ID inn in^ .OOo 't.UOD i Ol A

4.0657 150 192.77 .60 Tc r

o. 755 250 — 210

6.755 250 317.09 .627 10.78 400 -60

10.78 400 503.29 .742 14.54 540 + 80 50O 1.0 0.1208 8 5.04 0.76 0.159 10 -450

.159 10 8.19 .372 .4284 12 -448

.4284 12 16.08 .724 .5913 14 -446

.5913 14 19.79 .808 .7321 16 -444

.7321 16 23.0 .848 .8628 18 -442

.8628 18 25.97 .874 .9878 20 -440

160A 1.0 0.1148 10. 8.21 0.892 0.129 12 -448

.129 12 10.47 .84 .1537 14 -446

.1537 14 13.49 .802 .1918 16 -444

.1918 16 17.30 .812 .2361 18 -442

.2361 18 21.12 .844 .2803 20 -440

.2803 20 24.67 .87 .323 22 -438

.323 22 27.98 .887 .3645 24 -436
APPENDIX A. SIZING OF RELIEF DEVICES 291

Table 6A-1. G, values for helium, for insulated containers — Continued

(Calculated from: C, = 0.097(1200-(x) V^ZxT^j.//.')

Relief pressure psia Ax

\ni\ ^ TTlx

)

2 (mi — nix)

Btu

lb lOOX 10.92 0.896 0.207 3.26 14.1 0.52 28.3

15.71 .835 .33 4.79 12.1 .6 38.2

19.91 .871 .258 4.20 14.2 .73 38.4 23.47 .905 .19 3.56 16.9 .8 35.8 26.72 .925 .15 3.25 20.0 85 32.8 32.70 .89 .22 5.98 24.2 .92 30.8 41.08 .883 .235 8.38 31.4 .95 26.9 54.38 .814 .272 13.30 39.8 1.0 25.0

67.31 .896 .208 12.93 55.7 19.8 92.72 .855 .29 25.41 75.0 17.2 130.39 .85 .30 37.67 107 14.1 192.77 .834 .332 62.38 157 11.5 317.09 .80 .40 124.32 249 8.68 503.29 .813 .373 186.20 406 6.02 676.99 .871 .258 173.70 585 4.30 50O 8.19 0.88 0.24 3.15 11.6 0.33 25.2 16.08 .686 .628 7.89 8.6 .63 51 ;1 19.79 .862 .276 3.71 11.6 .69 42.7

23.0 .904 .192 3.21 15.1 .83 38.4 25.97 .924 .152 2.97 18.0 .86 34.8 28.82 .937 .126 2.85 21.2 .92 32.2 160A 10.47 0.946 0.108 2.26 19.8 0.65 20.6 13.49 .92 .16 3.02 17.4 .67 26.0

17.30 .901 .198 3.81 17.3 .7 28.9 21.12 .906 .188 3.82 18.4 .77 30.4 24.67 .922 .156 3.55 21.0 .81 28.8 27.98 .935 .13 3.31 23.8 .84 27.4 31.12 .943 .113 3.14 26.2

L- 88 26.6

2 (mi —mx)
292 SAFE AND EFFICIENT USE OF LIQUID HELIUM

and also.

(6A-6)

Note that the value of G, in (6A-5) is four times that indicated by comparison of (6A-4) and (6A-6). This factor of 4 has been introduced to permit use of a t/ value determined at 100 F instead of at the higher

average temperature incurred under fire conditions, and to allow for some insulation performance deterioration under fire conditions.

For the specific case of sizing relief devices for insulated helium

containers, it is proposed, where necessary, to substitute U (as pre- viously defined) for L, so that

Gi (Helium)

73.4 (1200-0 (ZTVl'^

L' X 378 U.OO/

(6A-7)

where C=378 for a monatomic gas at . k= CplCi = 1.67 at standard

conditions, and

Gi (Helium)

0.097 (1200-0 (ZD 1/2 (6A-8)

where t, Z, T, and L' are determined at the flowing conditions. In the rehef of helium vapor above the critical pressure, the following

system can be visualized:

mi = lb helium

vapor, at: Pu psia

Ai,Btu/lb

rrix lb hehum vapor

remaining in container, at: P^ = Pi,psia

ha:. Btu/lb

i^x, ft^/lb

(wi — rrix) lb helium vapor

relieved through

safety device, at: Px = Pu psia

Ai-^ /i^, Btu/lb

The heat absorbed by the helium vapor for any such incremental step can be approximated by

q=mx(K hi) -\- {mi — rrix) r,^^\ ^t^-
APPENDIX A. SIZING OF RELIEF DEVICES 293

The heat absorbed per pound of heHum vapor leaving the container is then:

TTixihx — h\)-\-{m\ — rux) L' = ql (mx — mjc) =

(mi—mx)

{h,-h,)

{hx — h\) {mi + TUx) Btu

2 (mi — TRx) lb (6A-9)

Table 6A1 summarizes, for several pressures, L' and d values as determined by (6A-8) and (6A-9) for a number of incremental steps.

The Zx. Tx. and tx values have been used in the calculation of d.

Figure 6A1 shows a plot of those Gi values as a function of Tx- Note that the d value peaks below 20 R for each of the several pressures

investigated. Figure 6A2 shows a plot of L' versus Ti for each of several

pressures.

20 30 50 100 200 300 400 500 600

X., deg. Rankine

Figure 6A-1. Values of 0\for sizing helium relief devices.

Relief pressures of 50. 100 and 160 psia are indicated.
294 SAFE AND EFFICIENT USE OF LIQUID HELIUM

A Gi value of 52.5 for pressurized liquid helium is being added to table 1 of the CGA Pamphlets S-1.2 and S-1.3 for a set pressure of

200 psig. Per the footnote to this table, the same value could be used

for lower settings of the relief device.

T|, deg. Rankine

Figure 6A-2. Heat absorbed per pound efflux for a helium container relieving above the

critical pressure.

Relief pressures of 40, 50, 88, 100, and 160 psia are indicated.
APPENDIX A. SIZING OF RELIEF DEVICES 295

The values of U and A used in the Qa — GiUA^-^^ formula would

depend on the characteristics of the specific design with U being deter- mined at 100 F with helium in the insulation space, and A the area in square feet of the Hquid container.

Since most safety devices are rated in terms of free air flow (at 60 F

and 14.7 psia), selection of the proper size of rehef device is straightforward. Size of the piping between the container and the relief device

should be no less than that of the device. Since no provision has been

made in the calculation procedure for warmup of the cold vapor, the

piping should be kept to minimum length.

A frangible disk would usually be used to meet this relief device

requirement, and the pressure for rating the flow of the device should

not exceed its nominal setting, or 1.2 times design pressure (plus 15

psi if vacuum insulation is used), whichever is less.

Loss of Vacuum (No Fire)

CGA Pamphlets S-1.2 and S-1.3 also specify that a safety relief valve be used on cryogenic containers, with a flow capacity sufficient

to handle loss of vacuum in the insulation space with air or lading, whichever is greater. The following example for loss of vacuum with hehum

shows the recommended method for determining this requirement.

From (6A-1) and substituting Ca = 356, Ma = 28.97, Za= 1.0, ra = 520 R, we get

Qa =

is.sr

c

{ZTIMyi\

But for this condition,

W=-Y-=UAaOO-t)IL,

and for helium

= 0.0244^/^(100- ?) [Zni^lL.

At a relief pressure of 10 psig= 25 psia

306-922 0-68—20
296 SAFE AND EFFICIENT USE OF LIQUID HELIUM

L = 6.72 Btu/lb of helium vaporized, or = 6.72/

(^'^ 5 95

'^^

) ^-^^ ^^"/^^ \^e\\um efflux

(where 0.725 = at 25 psia) and T=8.7R, t = -45l F,Z= 1.0, P/ (?,-0.0244(f/^)(551)[1.0(8.7)]»/V9.28

At a rehef pressure of 70 psig=85 psia (this is above critical point),

from figure 6A2, let U = 12 Btu/lb efflux, and r=15 R,^=-445 F,Z=1.0;

thus

Qa = 0.0244(f/^) (545)[1.0(15)]i/2/12 = ^.2UA.

At a relief pressure of 145 psig= 160 psia (above critical point), from

figure 6A2 letL' = 17 Btu/lb efflux and T=16 R, i=-444 F,Z=0.93; thus

^a = 0.0244(t/.4)(544)[0.93(16)]i/Vl7 = 3.10f/^.

The Qn requirement will vary with pressure. However, to streamline

calculation procedures, assume

where A is area of the liquid container.

Since the capacity of most relief valves is given as ^a=cfm of air

at 520 R and 14.7 psia, selection of the required device from manufacturers' listing is straightforward. The calculation procedure does not consider warmup of the cold vapor, so piping length to the relief valve

should be kept to a minimum.
REFERENCES 297 REFERENCES

[1] M. G. Zabetakis, Safety with Cryogenic Fluids (Plenum Press, New York, N.Y., 1967).

[2] ASME Boiler and Pressure Vessel Code: Section II — Materials Specifications; Section VIII — Rules for Construction of Unfired Pressure Vessels; Section IX — Welding Qualifications (American Society of Mechanical Engineers, 345 East 47th Street, New York. N.Y. 10017).

[3] J. L. Cost. Design and construction for safety, A Summary Report of Cryogenic Safety

Conference, Air Products, Inc., Allentown, Pa. (July 1959), pp. 31-40.

[4] G. M. Kintz and Frances C. Hill, Safety at gas processing plants, Bureau of Mines

Bulletin 588, USGPO, Washington, D.C. (1960).

[5] A. J. Croft, Ancillary equipment for the production of liquid hydrogen and liquid helium. Experimental Cryophysics (ed. F. E. Hoare, L. C. Jackson, and N. Kurti, Butterworths, London, 1961), p. 115.

[6] Howard H. Fawcett and William S. Wood, Safety and Accident Prevention in Chemical Operations (John Wiley & Sons, New York, N.Y., 1965).

[7] A. Wexler, Storage and transfer of liquefied gases, Experimental Cryophysics (ed. F. E. Hoare. L. C. Jackson, and N. Kurti, Butterworths, London, 1961), p. 138.

[8] G. C. Nubel and H. K. Hover, A bulk liquid helium distribution system. Liauid Helium

Technology (Intern. Inst, of Refrigeration, Commission 1, Bulletin 19oo-5), p. 329.

[9] Interstate Commerce Commission Regulations, Tariff No. 19.

[10] CGA Docket 1204: Insulated Cargo Tank Specification for Low-Temperature Liquefied Gas, Requirements for semitrailers and truck-mounted tanks which remain

on the chassis; CGA Docket 1127A: Insulated Portable Tank Specification for LowTemperature Liquefied Gases, Requirements for tanks having gross volume in excess of 1000-pounds water.

[11] Restricted Articles Regulations, International Air Transport Association, Section

IV, articles 865, 866 and 867; Section VI, articles 252 and 808.

[12] O. A. Hansen, Low temperature characteristics of metals. Refining Engineer (Dec.

1959), p. C-21.

[13] G. H. Zenner, Low temperature metals: A roundup, Chemical Eng. (Oct. 31, 1960).

[14] Cryogenic Materials Data Handbook, Technical Documentary Report No. ML-TDR64—280, Air Force Materials Lab., Research and Technology Division, Air Force Systems Command, Wright-Patterson Air Force Base, Ohio.

[15] Aerospace Research Apphcation Centers (ARAC), Indiana University, Bloomington,

Indiana.

[16] J. A. Paivanas and P. E. Loveday, U.S. Patent No. 3,272,374.

[17] L. E. Scott, Neck plug hazards in liquid helium shipment, Advances in Cryogenic

Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1964), Vol. 9, p. 379.

[18] J. H. Schiffhauer, Relief device requirements for nitrogen-shielded liquid helium

dewars, unpublished.

[19] Safety Relief Device Standards, Pamphlet S-1, Parts 2 and 3, Compressed Gas Association, 500 Fifth Avenue, New York, N.Y. (also see chapter 5). [20] R. J. Frainier, Liquid hydrogen containers: loss of vacuum with air, unpublished.

[21] R. C. Van Meerbeke, Commercial handling of liquid hehum dewars. Liquid Helium

Technology (Intern. Inst, of Refrigeration, Commission 1, Bulletin 1966-5), p. 303.
298 SAFE AND EFFICIENT USE OF LIQUID HELIUM

[22] A. J. Short, U.S. Patent No. 3,298,187.

[23] A. I. Shal'nikov, The prevention of sonic vibrations in hehum siphons. Cryogenics 6, No. 5, 299 (October 1966).

[24] H. A. J. Pieters and J. W. Creyghton, Safety in the Chemical Laboratory (Academic

Press, New York, 1957).

[25] The General Safety Committee of the Manufacturing Chemists Association, Inc., Guide for Safety in the Chemical Laboratory (D. Van Nostrand Co., Inc., New York,

1954).

[26] G. K. White, Experimental Techniques in Low Temperature Physics (Oxford Univ.

Press, 1959).

[27] A. C. Rose-Innes, Low Temperature Techniques (D. Van Nostrand Co., Inc., London,

1964).

[28] F. E. Hoare, L. C. Jackson and N. Kurti, Experimental Cryophysics (Butterworths, London, 1961).

[29] Lachlan MacKinnon, Experimental Physics at Low Temperatures (Wayne State Univ. Press, Detroit, Mich., 1966).

[30] R. B. Scott, Cryogenic Engineering (D. Van Nostrand Co., Inc., Princeton, N.J., 1959).

[31] F. Din and A. H. Cockett, Low Temperature Techniques (Geo. Newnes Ltd., London,

1960).

[32] J. Strong, Procedures in Experimental Physics (Prentice-Hall, New York, 1938).

[33] American Welding Society Inc., Brazing Manual (Reinhold Publ. Corp., New York,

1963).

[34] H. H. Manko, Solders and Soldering (McGraw-Hill Book Co., New York, 1964).

[35] J. F. Lancaster, The Metallurgy of Welding, Brazing and Soldering (Elsevier Publ.

Corp., New York, 1965).

[36] Victor J. Johnson (ed.), A Compendium of the Properties of Materials at Low Temperatures, Part II — Properties of Solids (Wright Patterson Air Force Base, Ohio,

WADD Tech. Note 60-56, 1960).

[37] R. M. McClintock and H. P. Gibbons, Mechanical Properties of Structural Materials

at Low Temperatures — A Compilation from the Literature, NBS Mono. 13 (1960).

[38] G. K. White, Experimental Techniques in Low Temperature Physics (Oxford Univ.

Press, 1959).

[39] E. G. Brentari and R. V. Smith, Nucleate and film pool boiling design correlations for O2, N2, H2, and He, Advances in Cryogenic Engineering (ed. K. D. Timmerhaus,

Plenum Press, New York, 1965), Vol. lOB, p. 325.
CHAPTER 7

Cryoelectronics

James W. Meyer ^ and Robert A. Kamper -

Page 7.1. Introduction 300

7.2. Quantum devices 301

7.2.1. The solid-state maser 301

7.2.2. The photon counter 304

7.3. Semiconductor devices 306

7.3.1. The parametric ampHfier 306

7.3.2. Radiation detectors 308

7.3.3. The cryosar 309

7.4. Superconductor devices 310

7.4.1. Some useful properties of superconductors 310

7.4.2. Superconducting switches 312

7.4.3. Superconducting amplifiers 315

7.4.4. Superconducting DC instruments 316

7.4.5. Superconducting radiation detectors 317

7.4.6. Passive superconducting circuit elements 318

7.4.7. Superconducting magnetometers and gravimeters 319

7.5. Working devices 322

7.5.1. Working masers and paramps 322

7.5.2. Working radiation detectors 323

7.5.3. Conclusion 323

' MIT Lincoln Laboratory, Lexington, Mass. 02173 (operated with support from the U.S. Air Force).

^ Cryogenics Division, NBS Institute for Basic Standards, Boulder, Colo. 80302. 299
300 CRYOELECTRONICS

7.1. INTRODUCTION

At present, cryoelectronics show great potential for the future and some solid achievement in the past. There is a promising future because

of many cryolectronic devices of comparatively recent invention, and

recently refrigeration has become sufficiently convenient to be practicable in many situations quite outside an elaborately equipped lowtemperature laboratory. Convenient as it has become, however, refrigeration still has a nontrivial cost in money, maintenance, and power, so there has to be good reason to choose a device working at a low temperature rather than some other device performing the same function at ambient temperature. Such reasons do exist: the maser amplifier has won acceptance because of its exceptionally low noise level; the super- conducting linear accelerator is being built in the hope that its electrical

loss will be so small that it can run at a power level unattainable at ambient temperature; and superconducting computer memories are being developed to an advanced state in the hope that their sheer simplicity will make them so much cheaper to fabricate than their present

day rivals that it will become worthwhile to use them when a sufficiently

large memory is required. Some recent developments in this field promise

an advantage in speed also. In this chapter we will describe briefly some of the physical phe- nomena which seem to be particularly significant for electronics, and

outline the principles of operation of some of the presently more important devices. We will also give some examples of successful applications.

There are some devices, such as the tunnel diode and the field effect transistor, which are normally used at ambient temperature but still work perfectly well when cold. We have not discussed these unless,

like the paramp, they have had extensive use at low temperatures or involve principles which are important for other purely cryoelectronic

devices. We have to admit that our choice has been a little bit arbitrary.

We have grouped the cryoelectronic devices under three headings:

Quantum, Devices, which we take to mean devices involving transitions

between the energy levels of paramagnetic impurities in crystals;

Semiconductor Devices; and Superconductor Devices. This is not a profound way to categorize them (for example, many of the superconductor devices are distinctly quantum mechanical) but it serves to intro- duce some order into the chapter. In a final section we describe a few

cryoelectronic devices which are presently in use in the field.
QUANTUM DEVICES 301

7.2. QUANTUM DEVICES

7.2.1. The Solid-State Maser

This is a low noise microwave amplifier which can significantly

improve the sensitivity of a microwave receiver when it is used as a pre- amphfier. It has been used in radio and radar astronomy systems, in space communication ground terminals, in deep space instrumentation,

in missile range tracking equipment, in ballistic missile defense radars,

and in reentry physics research radars. In each case the improved

system performance justified the inconvenience of low-temperature

operation.

The physical phenomenon upon which the maser depends [1-5] ^ is the paramagnetism of unpaired electrons in the metalHc ions of some

salts and oxides. This is associated with the presence of closely spaced

energy levels which can be split apart by the action of an appHed magnetic field. A flux density of the order of 0.1 to 1 tesla (T) causes a splitting

corresponding to a quantum of microwave energy in most paramagnetic

materials. Such a field is conveniently available from either an iron-core

electromagnet or a superconducting solenoid. The occupation of these

energy levels is governed by Boltzmann statistics. At high temperatures,

energy levels spaced as closely as these are very nearly equally occupied.

It is only at temperatures in the liquid helium range that there are substantial differences in energy level occupation. Figure 7.1 depicts the

relative population of three energy levels in thermal equilibrium at a temperature T. The lower the temperature, the greater is the preference

for E\ rather than Ez and E^. If we apply microwave power at a frequency

/ which satisfies the resonance condition hf=E3 — Ei, we can induce

transitions between these energy levels. As we increase the applied

power we disturb the occupation of £"1 and Es until we reach a point where

both levels are equally occupied and no further absorption of power can

take place. We say that we have "saturated" the transition. Depending

upon the details of the way the system relaxes toward equilibrium, it is possible in this way to achieve a situation in which the population of

£"3 exceeds that of E-z. Now microwave energy at a frequency resonant

with £"3 — £"2 will stimulate radiation of energy stored in the excess popu-

' Figures in brackets indicate the literature references at the end of this chapter.
302 CRYOELECTRONICS

(b)

0 0.5 1.0 RELATIVE POPULATION

q:

ijj z

0 0.5 1.0 RELATIVE POPULATION

Figure 7.1. Distribution of spins in a

system of three energy levels.

(a) In thermal equilibrium at a temperature T=(E2—Ei)/k.

(b) With the transition E3—E1 saturated, assuming

no relaxation into E^.

lation of £3. The available power clearly depends upon the difference in population of E3 and E2, since stimulated emission and absorption

through the same transition are equally probable. If this available power

exceeds the loss in the microwave hardware, we get coherent amplification of the stimulating wave. Spontaneous emission, which is incoherent and generates noise in the system, depends upon the occupation

of E3 and the frequency of the radiation.
SOLID-STATE MASER 303

From these elementary considerations we can draw several important conclusions. Clearly we improve the operation if we increase the

equilibrium occupation of the lowest level £"1. We do this by going to lower

temperatures. Increasing the separation of the energy levels would help

too, but this requires that we raise the frequency of the saturating power

into a range where high power sources become very expensive. We also observe that at a lower temperature the population of £"3 required to maintain a given excess over that of Ez is smaller, so noise from spontaneous

emission can be reduced. The power required to saturate a resonance

is proportional to the relaxation rate, which also falls when the temperature is reduced.

The reduction of these simple principles to practice requires ingenuity on the part of the microwave engineer. We shall not go into the

details of his problems here, except for some specific examples of hard- ware which has been used in practical applications. We point out, however, that refrigeration requirements for a maser are quite small,

typically of the order of one watt at the low temperature including all heat leaks and the power dissipated by the maser itself. The maser is one of the most sensitive amplifiers we can construct. We talk about the noise power developed by an amplifying system in terms of an effective noise temperature r„. This means that the noise

power from an amplifier with bandwidth B is equivalent to that from a

noise source at the input terminals of power Pn = kTnB. Contributions

to this noise temperature come from the black and grey body absorbers

and emitters in the system as well as the active device itself. Microwave

plumbing, for example, which absorbs 10 percent of the microwave

energy and is in equihbrium at a temperature of 300 K can contribute

30 K to the system noise temperature. This example shows the importance of reducing circuit losses or the temperature or both. A carefully constructed waveguide termination which absorbs

essentially all incident power and is maintained at a known low temperature can be used for an absolute standard noise source [6, 7] for measuring the performance of very low noise devices such as the maser. Some ingenuity is required to construct a radiometer to make full use of the available accuracy of this standard of very small power levels. In the microwave region, Planck's radiation formula can be simplified, and the minimum achievable noise temperature as limited by spon- taneous emission noise is [8] Tn = hflk. Figure 7.2 shows the variation

of this minimum noise temperature with frequency. Throughout the
304 CRYOELECTRONICS

100

FREQUENCY , MHz

Figure 7.2. Minimum theoretical noise temperature

plotted against frequency.

radio and microwave region of the spectrum it does not exceed a few

degrees, but in the visible region it rises to some thousands of degrees.

A number of masers have been operated below the lambda point of

liquid helium. Not only does this very low temperature help to improve

the overall performance in terms of gain, bandwidth, and noise, but also the superfluid maintains a very stable temperature and eliminates

potential problems resulting from bubbling or rapid boiling of liquid

helium.

7.2.2. The Photon Counter

Bloembergen [4] and Weber [3] independently suggested a device

which in principle can detect radiation with arbitrarily small noise.

Although not yet realized in practice, there is no reason to suppose that

it cannot be developed. In principle it is similar to the maser. Its action

is to replace each incoming photon with another of higher energy which

is therefore easier to detect. Consider the three-level system shown in
PHOTON COUNTER 305

Pump Output

i Signal

Figure 7.3. Energy level system

illustrating the action of the quantum counter.

figure 7.3 in which the spacing between E-z and the ground state Ei is resonant with the incoming signal. The temperature must be low enough

so that in equilibrium only Ei is significantly occupied. A powerful micro- wave oscillator supplies "pump" power at a frequency resonant with

E3 — E2, but since E-z and E3 are normally almost unoccupied it induces

transitions at a negligible rate. When an incoming photon is absorbed by

the system it induces a transition from Ei to Ez. The pump is then able

to induce a further transition up to E3. Provided that a system has been

chosen with favorable transition probabilities, most of these events will result in the system relaxing by spontaneous emission of a photon whose

energy is equal to E3—E1. Thus the net result of the whole process is the

absorption of photons of energy Ez—Ei followed by the emission of t^A'^ photons of energy E3—E1 , where r) is the quantum efficiency of the device.

One might therefore regard this device as an amplifier with power

gain r){E3 — E\) I (Ez — El) . However, it has the property that the phases

of the output and input signals are independent. This raises an inter- esting question which has been discussed by Heffner [8], who pointed

out that according to the quantum mechanical uncertainty principle

the phase of a wave and the number of quanta associated with it can

never be determined simultaneously with arbitrary precision. Thus the

noise in an amphfier, which represents fluctuations in the number of quanta, is necessary in order to preserve information about the phase.

This sets a lower limit to the attainable noise temperature of a phase

sensitive amplifier. The photon counter we have been discussing avoids

this limit by losing the phase of the signal completely.
306 CRYOELECTRONICS

7.3. SEMICONDUCTOR DEVICES

7.3.1. The Parametric Amplifier

The semiconductor diode parametric amplifier [9] has given the

sohd-state maser some of its stifFest competition for appHcation to field problems. A major reason for this is that, unlike the maser, the para- metric amplifier can operate over a temperature range from near absolute

zero to ambient. The noise is of course worse at higher temperatures,

but the fact that it will operate at all should the refrigeration fail is a great practical advantage.

Any amplifier requires a source of power that can be used to in- crease the signal power. The maser uses microwave power to store energy in the excited levels of a paramagnetic crystal, and the para- metric amplifier is similar in that it too is driven by microwave power

at a frequency higher than that of the signal. There are other similarities but attempts to describe both devices with a common fundamental

theory have been unsuccessful.

In the parametric amplifier (or paramp) the power source is used to vary (pump) one of the circuit parameters of a combination of resonant

circuits [10]. Such a variation of any component, such as a capacitor or inductor, which is capable of storing energy, can result in amplification

of a signal of a suitable frequency.

To understand paramp action in its simplest form let us consider

the resonant circuit shown in figure 7.4, in which the capacitor is variable.

As the circuit resonates the capacitor is alternately charged and dis- charged through the inductor, twice per cycle with alternating polarities.

The simplest way to decrease the capacitance is to increase the separation of the capacitor plates, and when they are charged this requires

mechanical work against their electrostatic attraction, which increases

the stored energy of the system. By separating the plates in those parts

of the cycle when they are charged and restoring their original positions

when the are discharged, it is possible to supply energy continuously

to the system and sustain the oscillation of the circuit. The same principle is used by a child pumping a swing'.

This simplest example, with the pump frequency double the signal

frequency, is known as a degenerate paramp. A more complex arrange- ment could have a circuit resonant at the signal frequency fs coupled to another circuit resonant at an idler frequency fi by a coupling which

varies at the pump frequency fp. l{fp=fs -\-fi this arrangement also sup-
PARAMETRIC AMPLIFIER 307

Figure 7.4. The action of a degenerate parametric amplifier, pumped by moving the capacitor plates at twice the resonant frequency of the circuit. plies energy to the signal. It has the advantage that the presence of the

idler circuit removes the need for fixing the phase of the pump relative

to that of the signal. This is the arrangement normally used for micro- wave work. In order to achieve stable amplification, rather than oscilla- tion, it is necessary to separate the input and output signals, which can

be accomphshed by means of isolators and circulators. Another possibility is the so-called "up-converter" mode, in which the output is taken

at the idler frequency.

The variable reactance commonly used in paramps is the varactor

diode. A junction between p and ^-type semiconductors normally contains a depletion region, with very few charge carriers, which increases

in width when a reverse bias voltage is appHed. The material in this region has a high resistivity and therefore acts as the dielectric in a capacitor whose "plates" are the conducting p and M-type regions on

either side. The variation of capacitance of this arrangement with re- verse bias voltage is used to drive the paramp.

There are two main sources of noise in the paramp. The first is the varactor diode itself, whose leakage resistance generates Johnson noise.

The shot noise from the leakage current is usually neghgible. Reducing

the temperature reduces the Johnson noise provided that the diode

resistance does not increase too much. The second source of noise is
308 CRYOELECTRONICS

the idler circuit, which receives noise generated elsewhere and also generates some of its own in lossy components. The amount of this noise coupled into the signal circuit depends upon the ratio of idler to signal frequencies. It can be reduced by making this ratio large and by

cooling the idler circuit. The ratio of idler to signal frequencies is limited by diode loss, which

becomes more serious at higher frequencies. In practice there is there- fore an optimum idler frequency.

In general cooling a semiconductor to a low temperature raises its resistivity by "freezing out" the carriers. However, the diodes used for varactors are usually so heavily doped that this problem does not arise,

and they work quite well at low temperatures. Diodes made of GaAs

and InSb have been used successfully in the liquid helium temperature

range to make very low noise amplifiers.

There are other variations of the paramp principle and other possibilities for the active element besides the varactor diode. In particular,

several up-converters have been made with the incoming signal at zero frequency, or direct current. They can be used as preamplifiers in ex- tremely sensitive systems for measuring small steady voltages. One

device of this type was developed by Hille, Reid, and Beer [11]. The

active element consists of a pair of coils separated by a sheet of bismuth.

At low temperatures the skin depth for alternating currents in the bis- muth varies with an apphed magnetic field because of magnetoresistance.

The resulting variation in the mutual inductance of the coils is sufficient for efficient paramp action. Another way to modulate the mutual induct- ance of two sets of coils is to move them mechanically. This principle was

used by Ries and Satterthwaite [12] to create a device with a sensitivity

of 4 X 10~^^ V with a time constant of 1 s. The coils were made of super- conducting wire to reduce the input impedance. One set was mounted

on an electrically driven tuning fork to provide the motion.

We will describe other parametric amplifiers in section 7.4.

7.3.2. Radiation Detectors

In general, perfectly adequate detectors for fight of visible or shorter

wavelengths can be made to operate at ambient temperature [13]. The

fundamental reason for this is that the quanta of radiation are large compared to kT, when T is the ambient temperature, so it is relatively easy

to find processes which can be driven by the radiation but not by avail- able thermal energy. In the far infrared, however, the situation is dif-
RADIATION DETECTORS 309

ferent, and the best available detectors for this region of the spectrum

all operate at a low temperature [13, 14]. In most infrared detectors, the radiation causes a change in the electrical resistance of the active element, which is detected by maintaining

a steady current and observing the change in voltage. It is usual to chop

the incident radiation with a rotating shutter, and detect variations of voltage at the chopping frequency, in order to discriminate against background radiation. Because of the low conversion efficiency of presentday transducers, this requires a very low-noise, low frequency preamplifier. The overall sensitivity of the complete detector could be improved

either by a radical improvement in transducers, or by using a preamplifier

with lower noise [15]. One of the future contributions of cryoelectronics

may be directed to this problem.

There are several mechanisms by which absorption of infrared radi- ation can affect the resistivity of a semiconductor. The simplest is used

in the bolometer, where the incident radiation raises the temperature of the semiconductor and affects the resistivity through its temperature de- pendence. This simple principle has been developed into a very sensitive instrument by Low [16]. Putley [17] has developed a refinement of the principle in which only the energy distribution (and hence temperature) of the conduction electrons is affected by the radiation. The temperature of the crystal lattice does not need to change, so the effective

heat capacity of the device is very small. In photoconductive detectors

[13, 18] the radiation excites current carriers out of bound states and

thereby increases the conductivity of the semiconductor. For long wavelengths the bound states must be very close to the conduction band in energy, so a low temperature is necessary to avoid masking the photoconductive effect with thermal excitations. Arndt and Hartwig [19]

have developed a variation of this detector in which the semiconductor

is placed in a resonant microwave cavity. The carriers which are excited

by the incident radiation affect the real part of the dielectric permittivity and hence the resonant frequency of the cavity. The resulting

phase shift in microwave transmission or reflection by the cavity can

be detected with a microwave bridge. The sensitivity of this system can

be increased greatly using a superconducting microwave cavity.

We will mention some superconducting infrared detectors in section 7.4. 7.3.3. The Cryosar

This device was developed for use as a computer memory [20],

Its name is an acronym for low temperature switching by avalanche
310 CRYOELECTRONICS

breakdown and recombination. In compensated crystals of germanium

at temperatures below 10 K, impact ionization of the impurities gives

rise to a negative resistance characteristic in the bulk material, which

can be used to make a simple bistable device. It is switched by voltage

pulses of appropriate size, applied through ohmic contacts. Although this device promises to be fast, compact, and simple to fabricate, it has not yet been used in a practical computer.

7.4. SUPERCONDUCTOR DEVICES

7.4.1. Some Useful Properties of Superconductors

The most striking property of superconductors, for which they are named, is the absence of electrical resistance at all temperatures below

the characteristic critical temperature, provided that they are not ex- posed to an excessive current or magnetic field. Small magnetic fields and currents do not permeate the bulk of a superconductor. They are confined to a surface region whose thickness, known as the penetration

depth, is of the order of 1000 A. This penetration depth is sensitive to magnetic fields and may be altered significantly in a thin film, which

can retain its superconductivity in the presence of magnetic fields strong enough to drive a thick specimen of the same metal into the

normal state. This results in a nonlinear response to electromagnetic

radiation. In addition to these properties of isolated superconductors, there

are some fascinating phenomena which are observed when a weak link is formed between two superconductors. In practice this can be achieved

quite simply by touching lightly together two pieces of superconducting

metal which have been exposed to normal atmospheric corrosion. The

properties of such a junction depend strongly on the contact pressure and

are usually improved by shaping one of the superconductors to a fine point. Another technique is to form a thin uniform layer of oxide or other

insulator on the smooth surface of a thin film of one superconductor and

then to deposit a thin film of the second superconductor on top. There are various mechanisms for conduction of electric current across these

structures, but the most striking effects are due to the quantum mechanical tunnel effect [21, 22], and other contributions to the current tend to have nuisance value only and must be suppressed as much as possible in a well-fabricated junction.
USEFUL PROPERTIES OF SUPERCONDUCTORS 311

Figure 7.5. Current versus voltage

characteristic of a tunnel junction

betiveen two superconductors sepa- rated by an insulating barrier too thick to pass a supercurrent.

VOLTAGE

If the insulating barrier in a tunnel junction between two identical

superconductors is rather thick, the current versus voltage characteristic

typically looks Hke the curve shown in figure 7.5. For small voltage, the current is very small and temperature dependent, tending to zero at zero temperature. As the voltage is increased, there is a well defined value of the order of a milHvolt at which the current increases sharply, and thereafter the junction has an ohmic response. This characteristic curve is controlled mainly by the influence of the energy gap for excitations in superconductors, and may be modified by the presence of electromagnetic radiation [23] or ultrasound [24, 25] of sufficiently high frequency.

A very thin insulating barrier permits the two superconductors to influence each other enough to create various effects due to quantum

mechanical phase coherence, first predicted by B. D. Josephson [26] and subsequently verified when the experimental difficulties were over- come. Briefly, these eff"ects consist of the passage of electric current

across the junction without dissipation: a direct current can flow at zero

voltage; when the junction is biased to a finite voltage, the energy gained

by each electron pair as the current flows is radiated away as a quantum

of electromagnetic radiation, the frequency / being given by /i/— 2eV

where h is Planck's constant, e is the electron charge, and V is the bias

voltage. The d-c effect at zero bias may be distinguished from a simple

metallic short in the junction by its extreme sensitivity to a magnetic

field, which disturbs the phase coherence. The maximum current Ic which can flow at zero bias through a typical thin film junction, covering

about 1 mm square, is plotted against the strength of a magnetic field.

306-922 0-68—21
312 CRYOELECTRONICS

FLUX DENSITY , microtesia Figure 7.6. Maximum supercurrent Ic plotted against applied magnetic

field for an ideal rectangular Josephson junction.

The positions of the nodes depend upon the dimensions of the junction.

applied in the plane of the junction, in figure 7.6. The Josephson effect

is clearly a rich source of potential devices.

7.4.2. Superconducting Switches

The earliest and, at the time of writing, most developed superconducting switch is the Cryotron [27]. In its present form [28] it consists of two thin overlapping superconducting films, on a suitable substrate,

which are separated by an insulating layer too thick for leakage by elec- tron tunneling. These two films are in the form of narrow strips and are known as the "gate" and the "control." The gate can be switched be- tween the superconductive and resistive states by the magnetic field due to a current in the control strip, which is designed to remain always

in the superconductive state. The gate could also be switched by its own current if it were large enough, and it is meaningful to talk of a current gain defined as the ratio of gate current to control current re- quired to switch the gate into the resistive state. If this gain exceeds

unity it is possible to devise many computer circuits [29], both for logic

and for storage, in which the paths taken by electric current in a super- conducting network are controlled by cryotrons.

Although cryotrons can be combined to form bistable circuits suit- able for computer memories, more compact devices for this purpose

have been devised. One of the most promising of these is the continuous

film memory described by Burns [30] . This consists of a sheet of super- conducting tin covered on one side by a square grid of insulated strips
SUPERCONDUCTING SWITCHES

Lead Driving Lines

313

0

Lead Sense Line Figure 7.7. Single cell of a continuous film computer memory.

of lead. Each intersection of this grid forms one storage cell, and there

is an insulated sense line on the other side of the sheet consisting of a

strip of lead which passes over every cell on the sheet. A single cell of

this arrangement is shown in figure 7.7. The memory is operated by

pulses of current in the X and Y strips of the grid. The ampHtude of these

pulses must be adjusted so that only the combined current in both X and

Y strips of a cell creates sufficient magnetic field to switch the cell by

driving the nearby part of the tin sheet into the normal state and thereby

trapping a loop of magnetic flux. Once this metastable loop of flux is established, a current pulse in either the X or the Y strip can drive the

tin normal again, and induce a voltage pulse in the sense line, only if

its direction is such that its magnetic field adds to the flux loop. A single

pulse in the opposite direction does not communicate with the sense

line. Units with 16,384 cells together with cryotrons to drive the switching currents have been made on a single sheet in a very compact ar- rangement [30] . Burns argues that the chief advantage of this type of memory is that the price per bit of storage falls rapidly as the size of

the memory increases. He estimates that a memory of about 13 million

bits arranged on 50 sheets would cost about one cent per bit, a very
314 CRYOELECTRONICS

LlI

tt: 3

O

Figure 7.8. Current versus voltage

The operating points of interest for the tunneling

cryotron are marked A and B.

characteristic of a Josephson junction between superconducting thin films.

0

VOLTAGE

competitive price. The speed of these memories would be hmited by

the cryotron circuits that drive them.

The speed of a cryotron circuit is Hmited by the resistance and in- ductance of the gate in the normal state and by the time required for the

superconducting to normal state transition. Matisoo [31] has reported a new type of cryotron which promises to be much faster than the device we described earlier in this section. The improved performance is obtained by replacing the gate film with a Josephson junction. The control strip is retained. The current versus voltage characteristic of the

gate with zero control current is shown in figure 7.8. A gate current less than Ic (see fig. 7.8) would normally be passed with zero voltage at an

operating point such as A. The effect of a control current of the right sense is to reduce /c, switching the cryotron into the finite voltage operating point B. If there is a second cryotron connected in parallel the current would then be transferred to it in a time which can be as small

as a few hundred picoseconds. After the transfer of current the first cryotron would be free to switch back into the zero voltage mode, but

if all the connections are superconducting the current would remain in

its new path until it is disturbed again by the introduction of resistance

into the circuit. Thus the combination of two such cryotrons forms a

bistable device with a very fast switching time.

One disadvantage of this device is that Josephson junctions are dif- ficult to fabricate with a high success rate. However, a development

effort equivalent to a small fraction of that which transistors received

should solve this problem.
SUPERCONDUCTING AMPLIFIERS 315

7.4.3. Superconducting Amplifiers

Because of their nonlinear response to currents and magnetic

fields, superconductors lend themselves very readily to parametric

amplification. Clorfeine [32] and Bura [33] have both reported parametric

microwave ampHfiers in which the mixing of signal and pump frequencies

is achieved by nonlinear transmission through superconducting thin

films. Point contact Josephson junctions have also been used for micro- wave mixing [34]. One striking achievement was reported by Zimmerman and Sil- ver [35] , who used a point contact Josephson junction as the active

element in a parametric up-converter. The input signal was the oscilla- tion of the junction itself under the influence of a very small bias voltage.

They were able to convert a detectable signal up to 30 MHz, with an

input at 1 Hz. The amphtude of the input could not exceed the bias

voltage of 2 X 10~^^ V, which could be measured by measuring the

frequency of oscillation. Such a sensitivity can only be achieved in a superconducting circuit. Another superconducting device which can function as the active

element in a parametric amplifier is the Ryotron [36, 37]. This device

is superficially similar to the cryotron, in that it consists of a pair of superconducting strips, the "gate" and the "control," in close proximity.

However, the function of these strips is different. The gate always re- mains in the superconducting state during normal operation, and its important property is its self inductance, which is influenced by the

tendency of the control strip to screen the magnetic field due to the gate

current. The device operates when a large current is passed through the

control strip, driving it i'nto the normal state and allowing magnetic

flux to penetrate it. By careful attention to geometry it is possible to cause the inductance of the gate to change by over two orders of magnitude when the control switches into the normal state. This switching can be accomphshed in a time of the order of a nanosecond, giving a broad bandwidth for parametric applications. It is also possible to use

the ryotron to steer currents in a superconducting network just like a

cryotron, and it has the same potential for logic circuits. In contrast to these parametric amplifiers, Newhouse and Edwards

[38] have developed a broadband, linear amplifier based on the cryotron. The switching of the gate of a cryotron from the superconducting

to the resistive state does not happen instantaneously, but occupies a

finite range of control current. If the cryotron is biased into the center

of this range, and kept there by a feedback circuit incorporating a low-
316 CRYOELECTRONICS

pass filter, then the gate resistance is a very sensitive function of the

control current for frequencies above the cutoff of the filter. The gain

of the cryotron used in this mode is of course much greater than the "current gain" commonly used to describe the switching mode.

These amplifiers usually have a low input impedance and excellent

noise characteristics. Newhouse and ^dwards [38] report a minimum

detectable signal of 1.3 X 10"^^ V for an amplifier operating at 100 Hz

with a bandwidth of 10 Hz. The source resistance was less than 3 X 10~^

n and the operating temperature was 3.6 K.

7.4.4. Superconducting D-C Instruments

We have already mentioned two excellent instruments for measuring very small, steady voltages: namely, Ries and Satterthwaite's para- metric amplifier [12] with vibrating coils (sec. 7.3.1) and Zimmerman and

Silver's detector [35] using parametric up-conversion of oscillations due to the Josephson effect (sec. 7.4.3). The performance of these instru- ments is unsurpassed at the time of writing.

An interesting low-level null detector has been developed by

Clarke [39]. In essence, the input circuit is a short piece of superconducting niobium wire, with negligible resistance and a small inductance.

Any current which flows in this wire is sensed by the effect its accompanying magnetic field has on the critical current of a Josephson-hke

junction between the wire and another superconductor. Clarke formed

these junctions by freezing a blob of lead-tin solder onto the niobium

wire. The solder does not wet the niobium but instead forms an electrical connection which behaves like a multiple point-contact Josephson junction. It is possible to form junctions of this type whose critical current,

or resistance for currents slightly over critical, is a well defined periodic

function of the current in the niobium wire. If a voltage is applied to the niobium wire a current builds up linearly with time until it becomes

large enough to have a detectable effect on the junction. Allowing a time

of one second, Clarke found he could detect 10~^^ V.

In addition to these very sensitive low-level instruments, the

Josephson effect may make a significant contribution to precise measure- ments of fundamental quantities. Parker, Taylor, and Langenberg [40]

have measured the relationship between frequency and voltage,

hf~2 eV, to an accuracy of a few parts per million, deriving a new value

for the constant lejh from their results. Assuming that there is no un- known electrodynamic correction to be applied to this quantity, they

used it to calculate a new value for the fine structure constant, a = e^lhc.
SUPERCONDUCTING RADIATION DETECTORS 317

which resolved a long-standing discrepancy between the theoretical

and experimental values for the hyperfine splitting of the ground state

I of the hydrogen atom.

I Their experience with this measurement suggested to them that the

Josephson effect might be used with advantage to maintain a standard

: volt by frequency to voltage conversion [41]. Among the advantages they

listed were the simphcity of the equipment, the lack of stringent requirements on the preparation of the junction and experimental conditions,

jj and the ready availability of very accurately known frequencies.

A companion to this idea was Meservey's [42] suggestion that the

quantum of magnetic flux, /i/2e, should be used as a fundamental elec- tromagnetic unit. The volt could then be derived from it by reversing

the emphasis of Parker, Taylor, and Langenberg's measurement [40]

,

and the ampere could be derived from it by a direct measurement of the flux density generated by a precisely made solenoid. This would be

accomplished by using the Josephson effect to count the flux quanta

Unking a superconducting circuit of known area.

7.4.5. Superconducting Radiation Detectors

As we mentioned in the introduction to this chapter, it is only necessary to use a cold detector for radiation of long wavelength. Both the

detectors we describe in this section were designed for the far infrared.

The older of the pair is the superconducting bolometer, which was

developed into a working instrument by Martin and Bloor [43]. It consists of a thin superconducting film, on a very light substrate, connected

; to the helium bath through a thermal resistance. The transition from the

! superconductive to the normal resistive state occupies a finite range of

j temperature, and in this range the resistance of the film is a very sensi-

' tive function of temperature. Martin and Bloor found that the minimum

incident power detectable by its heating effect on this structure was

10-12 ^ (noise equivalent power), with a time constant of 1.25 s. In

I order to achieve its sensitivity this detector suffers from the rather

slow response time of 15 ms, which can be a disadvantage for some

applications.

Grimes, Richards, and Shapiro [44] showed that this situation can

be improved, as always, by making use of the Josephson effect. The

junctions they used were point contacts with current versus voltage

characteristics as shown in figure 7.9. (Note the difference from the thin

film junctions illustrated in fig. 7.8.) They found that the maximum

current at zero voltage, as well as the nearby part of the curve, were
318 CRYOELECTRONICS

0

VOLTAGE

Figure 7.9. Current versus voltage characteristic of a point contact Josephson

junction {solid curve). Long wavelength radiation causes the characteristic

to follow the dotted curve.

sensitive to irradiation with far infrared, which would cause the curve

to shift as shown in figure 7.9. By biasing the junction to a suitable operating point, they were able to detect 5 X 10~^^ W (noise equivalent

power) with a time constant of one second. The response time was less than 10 ns and the detector was sensitive to radiation of any wavelength

shorter than 300 fxin (corresponding to a quantum of about twice the

energy gap of the superconductor they used). The difficulty of concentrating a reasonably large portion of the incident radiation onto the very

small junction was overcome with a metal light pipe.

7.4.6. Passive Superconducting Circuit Elements

The obvious benefit of superconductivity to passive circuit elements

is to reduce loss and hence reduce power consumption and permit tuned

circuits of narrower bandwidth. In this connection we should remark

that superconductors show zero resistance only for direct current. Alter- nating currents suffer loss through hysteresis effects and, at higher

frequencies, through the presence of thermally excited carriers which

behave as if they were in the normal state. This latter mechanism be- comes stronger as the frequency is increased, but still allows super- conductors to show superior performance to normal metals for wavelengths through the millimeter region.
SUPERCONDUCTING MAGNETOMETERS AND GRAVIMETERS 319

The most ambitious use of superconducting passive circuits to date

is in the hnear accelerator, described in chapter 8. Its performance

will be improved by several orders of magnitude as a direct result of using superconductmg microwave cavities. There has been some work on superconducting radiofrequency

filters. A filter based on circuits with a (? of 500,000 at 27 MHz has been

reported [45] . It was made of superconducting NbaSn and had a working temperature of 15 K. Superconducting pulse delay lines seem to be very promising. The

reduction in loss in the line yields a wider bandwidth and hence less dis- tortion of sharp pulses [46]. Quite a crudely made coaxial line, with a niobium inner conductor, teflon dielectric and lead outer conductor,

operating at 4 K, was found to permit a pulse rise time of 500 ps in a length sufficient to give a delay of 2 /jls. Large coils carrying heavy currents store a great deal of energy

inductively. This is a source of embarrassment to the designers of stable superconducting magnets, but it is the cheapest way to store energy in megajoule quantities for release in short pulses. If the cycle time is short

it is unnecessary to use a superconducting coil, and there may be some

advantage in using a coil of normal metal, cooled to reduce its resistivity

somewhat [47].

Another property which superconductors have is that in the form of

thin filaments they can pass large currents with very few carriers.

Little [48] has shown that when they do so the energy stored as kinetic

energy of the electrons contributes a significant amount to the self- inductance of the circuit. This offers a solution to a problem which has

troubled the designers of miniature integrated circuits: that of making

inductors of very small physical size.

7.4.7. Superconducting Magnetometers and Gravimeters

The superconducting magnetometers which have been proposed

or built fall into two classes. Either they depend upon sensing the persistent current induced in a superconducting loop or they depend upon

counting flux quanta in a quantum interference device. Both kinds really measure changes in field and need precautions to set the zero for ab- solute measurements.

Since a superconducting loop has zero resistance, it follows from

Lenz's law that if it is placed in a magnetic field a persistent electric current must flow so as to keep the magnetic flux linking the loop

constant. Regardless of the geometry of the loop, this current is linearly
320 CRYOELECTRONICS

related to the change of field, provided the superconductor remains

always in the superconducting state. Several schemes have been used

to detect this persistent current nondestructively. If the self inductance

L of part of the circuit carrying a current / is modulated an emf equal

to didt {LI) appears, which can be coupled out of the circuit by means

of a transformer. The modulation can be accomplished by moving a superconducting ground plane mounted on a vibrating quartz crystal

close to the loop [49], or by driving a superconducting core in a small

coil repeatedly through its transition temperature [49, 50]. Some

advantage can be gained by tuning the circuit to resonate at the modulation frequency. With a wire loop of radius 1 cm a calculated sensitivity [49] of 10~^4 T should be achievable with a bandwidth of 1 Hz. Another sensitive detector of the induced persistent current, which

does not require modulation, is the Clarke [39] null-detector which we

mentioned in section 7.4.4. By using a two-stage detector it might be

possible to achieve even greater sensitivity than is feasible by the

modulation technique.

The classic flux-quantum counter [51] depends upon the fact that the Josephson current which crosses an insulating barrier is proportional to sin where (f is the local difference in quantum mechanical

phase across the barrier. If two such barriers are connected together in a superconducting ring, as shown in figure 7.10, then the total quantum

mechanical phase difference integrated round the ring is equal to the

number of flux quanta {hl2e) enclosed multiphed by 277. Most of this phase difference appears across the barriers, so it follows that if they are

identical and if the ring encloses an integral number of flux quanta the

Josephson currents through the two barriers must be equal and opposite.

For a half-integral number of flux quanta the two currents will flow in the same sense. Thus the maximum current that the two-branch struc- ture shown in figure 7.10 can pass with zero voltage is a periodic function

of the magnetic flux finking the ring, the period being one flux quantum.

Recently Mercereau [52] has shown that it is possible to count flux quanta entering a weak superconducting ring without constructing such

an elaborate device as the double junction ring we have described. The

achievable sensitivity appears to be comparable to that of the other

superconducting magnetometers we have described.

Closely related to these magnetometers is Prothero and Goodkind's

superconducting gravimeter [53]. This is a device for detecting changes

in the acceleration g due to gravity over a long period of time. It consists of a superconducting ball which is levitated by a magnetic field.
SUPERCONDUCTING MAGNETOMETERS AND GRAVIMETERS 321

t

Figure 7.10. Quantum interferometers for counting quanta of magnetic flux.

(a) Ideal, with point contacts

(b) Practical version with thin film junctions described in ref. [51].

The field gradient required to keep the ball in position is proportional

j to g and is maintained by a sensitive servo loop. The position of the ball

I is sensed both by measuring the capacitance between it and some fixed

plates and by using a sensitive magnetometer to detect the distortion of

' magnetic field it causes. The advantage they hope to obtain over the more usual spring balance is in long-term stability. In principle, this device is perfectly stable and it is designed for a sensitivity of 10~^^ g.
322 CRYOELECTRONICS

7.5. WORKING DEVICES

In this section we describe a few examples of cryogenic devices

which have been used in the field because their performance warranted

the inconvenience of keeping them cold. This inconvenience will hopefully be reduced by the development of compact, rehable refrigerators

to handle the rather modest heat loads from these devices.

7.5.1. Working Masers and Paramps

An early application of the soUd-state maser followed Bloembergen's

suggestion [1] that it would be useful in radio telescopes. The residual

background temperature of the universe appears to be only about 3 K

when measured by microwave radiometry, a discovery which owes an

obvious debt to cryoelectronics [54, 55]. Although the background temperature seen by most radio telescopes exceeds this, it is low enough so

that faint sources may be picked out by a sensitive instrument without

being swamped.

A UHF solid-state maser [56] was used with the Millstone radar at the Lincoln Laboratory in an attempt to make a radar observation of the

planet Venus. Operating at a temperature of L6 K, it gave a gain of 10 dB

with a bandwidth of 100 kHz.

A self-contained package including a ruby maser with a closed

cycle refrigerator has been developed at the Jet Propulsion Laboratory

[57]. It gives a gain of 37 dB with a bandwidth of 18 MHz centered on 2295 MHz. The operating temperature is 4.5 K and the noise temperature

is 15 K. The whole package weighs 200 kg.

Masers were used in the Telstar sateUite communication experiments [58]. They gave gains near 30 dB at 4170 MHz with a bandwidth

of 25 MHz. The noise temperature was 3.5 K at an operating temperature

of 4.2 K, indicating very efficient design.

An experimental transportable ground terminal for space communications at the Lincoln Laboratory incorporates a parametric pre- amphfier [59] with two channels, in the band between 7.7 and 8.4 GHz.

Each channel consists of a two-stage paramp using gallium arsenide

varactor diodes. It has an overall noise temperature of 55 K. Helium-cooled masers and paramps have even been used in missile

tracking and defense radars, for which reliability is at a premium.
WORKING RADIATION DETECTORS 323

7.5.2. Working Radiation Detectors

Many of the infrared detectors we described earlier in this chapter

have seen considerable service in spectroscopy, and some of them are

available as commercial instruments.

The superconducting bolometer [43], for example, is the workhorse of a productive group of infrared spectroscopists at Queen Mary

College in London [60]. Low's germanium bolometer [16] is available

as an optional detector with several commercial spectrometers. The

performance and commercial availability of many photoconductive

detectors have been summarized by Putley [61]

.

Infrared detectors are also used occasionally in astronomy. Westphal, et al. [62] . have used a crystal of germanium doped with mercury

for this purpose. It was cooled to 20 K with hquid hydrogen and used in the wavelength range from 8 to 14 ^tm. The reason for this choice was

to take advantage of a "window" in the atmospheric absorption spectrum

in order to make observations on the planets which would help to reveal

the compositions of their atmospheres.

7.5.3. Conclusion

In looking over the hst of ingenious devices we have mentioned in

this chapter, it seems that rather a small proportion have won presentday acceptance in pragmatic engineering. Many of the others offer remarkable performance which we do not know how to use yet. No

doubt the all-round advance in technology will one day create problems

requiring these elegant solutions.
324 CRYOELECTRONICS

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[3] J. Weber, Masers, Rev. Modern Phys. 31, 681 (1959).

[4] N. Bloembergen, Solid state masers, Progress in Low Temperature Physics (ed. C. J. Gorter, North Holland Publ. Co., Amsterdam, 1961), Vol. Ill, p. 396.

[5] A. E. Siegman, Microwave Solid-State Masers, (McGraw-Hill Book Co., Inc., New

York, N.Y., 1964).

[6] C. T. Stelzried, Temperature calibration of microwave thermal noise sources, IEEE

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[7] C. Trembath, D. F. Wait, G. Engen, and W. foote. The NBS Helium-Cooled Micro- wave Noise Standard, to be published.

[8] H. Heffner, The physical aspects of low noise electronics. Low Noise Electronics

(ed. K. Endresen, Pergamon Press, Ltd., London, 1962), p. 3.

[9] W. W. Mumford, Some notes on the history of parametric transducers, Proc. IRE

48, 848 (1960).

[10] W. H. Louisell, Coupled Mode and Parametric Electronics (John Wiley & Sons, Inc., New York, N.Y., 1960).

[11] P. F. Hille, F. J. Reid, and A. C. Beer, Amplification by magnetic control of skin depth.

Solid State Electronics 9, 453 (1966).

[12] R. P. Ries and C. B. Satterthwaite, A superconducting parametric amphfier for the measurement of small DC voltages. Rev. Sci. Instr. 38, 1203 (1967).

[13] R. A. Smith, Detectors for ultraviolet, visible, and infrared radiation, Appl. Opt. 4,

631 (1965).

[14] P. Bratt, W. Engeler, H. Levinstein, A. MacRae, and J. Pehek, A status report on

infrared detectors. Infrared Phys. 1, 27 (1961).

[15] E. H. Putley, The ultimate sensitivity of sub-mm detectors, Infrared Phys. 4, 1 (1964).

[16] F. J. Low, Low-temperature germanium bolometer, J. Opt. Soc. Am. 51, 1300 (1961).

[17] E. H. Putley, Indium antimonide submillimeter photoconductive detectors, Appl. Opt. 4, 649 (1965).

[18] H. Levinstein, Extrinsic detectors, Appl. Opt. 4, 639 (1965).

[19] G. D. Arndt, W. H. Hartwig and J. L. Stone, Photodielectric detector using a super- conducting cavity, J. Appl. Phys. 39, 2653 (1968).

[20] A. L. McWhorter and R. H. Rediker, The cryosar — A new low-temperature computer

component, Proc. IRE 47, 1207 (1959).

[21] I. Giaever, Energy gap in superconductors measured by electron tunneling, Phys. Rev. Letters 5, 147 (1960).

[22] B. D. Josephson, Supercurrents through barriers. Advances in Physics 14, 419 (1965).

[23] A. H. Dayem and R. J. Martin, Quantum interaction of microwave radiation with tun- nehrig between superconductors, Phys. Rev. Letters 8, 246 (1962).

[24] Y. Goldstein, B. Abeles, and R. W. Cohen, Tunneling induced by longitudinal micro- wave phonons in Al-Pb, Al-Sn, and Pb-Pb superconducting diodes, Phys. Rev. 151,

349 (1966).

[25] W. Eisenmenger and A. H. Dayem, Quantum generation and detection of incoherent phonons in superconductors, Phys. Rev. Letters 18, 125 (1967).

[26] B. D. Josephson, Possible new effects in superconductive tunnehng. Physics Letters

1, 251(1962).
REFERENCES 325

[27] D. A. Buck, The cryotron — a superconductive computer component, Proc. IRE 44,

482 (1956).

[28] J. M. Lock, Superconductive switching, Reports on Progress in Physics 25, 37 (1962).

[29] J. D. Barnard, F. A. Behnke, A. B. Lindquist, and R. R. Seeber, Structure of a cryo- genic associative processor, Proc. IEEE 52, 1182 (1964).

[30] L. L. Burns, Cryoelectric memories, Proc. IEEE 52, 1164 (1964). See also: J.D. Barnard, R. H. Blumberg. and H. L. CasweU, Operation of the cryogenic continuous

film memory ceU, Proc. IEEE 52, 1177 (1964).

[31] J. Matisoo, The tunnehng cryotron — a superconductive logic element based on elec- tron tunneling, Proc. IEEE 55, 172 (1967).

[32] A. S. Clorfeine, Microwave amplification with superconductors, Proc. IEEE 52 , 844

(1964). See also: A. S. Clorfeine, NonUnear reactance and frequency conversion in superconducting films at millimeter wavelengths, Appl. Phys. Letters 4, 131 (1964).

[33] P. Bura, Parametric amplification with superconducting films, Proc. IEEE 54, 687

(1966) . [34] G. K. Gaule, R. L. Ross, and K. Schwidtal, Microwave mixing with weakly coupled

superconductors, Symp. Physics of Superconducting Devices, Charlottesville, Va.

(1967) . Available as an ONR report.

[35] J. E. Zimmerman and A. H. Silver, High sensitivity superconducting detector, Appl.

Superconductivity Conf., Austin, Tex. (1967). To be published. See also: A. H.

Silver, J. E. Zimmerman, and R. A. Kamper, Contribution of thermed noise to the

line-width of Josephson radiation from superconducting point contacts, Appl. Phys.

Letters 11, 209 (1967).

[36] R. A. Gauge, The cryotron — a new cryogenic device, Proc. IEEE 52, 1216 (1964).

[37] J. C. Miller, C. M. Wine, and L. S. Cosentino, The ryotron — a variable inductance

cryogenic device, Proc. IEEE 52, 1223 (1964).

[38] V. L. Newhouse and H. H. Edwards, An ultrasensitive linear cryotron amplifier, Proc. IEEE 52, 1191 (1964).

[39] J. Clarke, A superconducting galvanometer employing Josephson tunnehng, Phil. Mag.

13, 115 (1966). See also: J. Clarke, The measurement of small voltages using a quantum interference device, Symp. Physics of Superconducting Devices, Charlottesville, Va. (1967). Available as an ONR report.

[40] W. H. Parker, B. N. Taylor, and D. N. Langenberg, Measurement of 2e//i using the AC Josephson effect and its implications for quantum electrodynamics, Phys. Rev.

Letters 18, 287 (1967).

[41] B. N. Taylor, W. H. Parker, D. N. Langenberg, and A. Denenstein, On the use of the AC Josephson effect to maintain standards of electromotive force, Metrologia 3, 89 (1967).

[42] R. Meservey, Proposed absolute ammeter using flux quantization, J. Appl. Phys. 39,

2598 (1968K

[43] D. H. Martin and D. Bloor, The application of superconductivity to the detection of radiant energy. Cryogenics 1, 159 (1961).

[44] C. C. Grimes, P. L. Richards, and S. Shapiro, Far infrared response of point-contact

Josephson junctions, Phys. Rev. Letters 17, 431 (1966). See also: S. Shapiro,

Josephson effect far infrared detector, Symp. Physics of Superconducting Devices,

Charlottesville, Va. (1967). Available as an ONR report.

[45] K. Siegel, R. Domchick, and F. R. Arams, High circuit ^'s with niobium stannide at radio frequencies, Symp. Physics of Superconducting Devices, Charlottesville, Va.

(1967). Available as an ONR report.
326 CRYOELECTRONICS

[46] R. J. Allen and N. S. Nahman, Analysis and performance of superconductive coaxicd transmission lines, Proc. IEEE 52, 1147 (1964).

[47] V. Arp, Cryogenic coil for megajoule energy storage, Proc. Intern. Symp. Magnet

Technology, Stanford, Calif. (1965).

[48] W. A. Little, Device application of super-inductors, Symp. Physics of Superconducting Devices, Charlottesville, Va. (1967). Available as an ONR report.

[49] J. M. Pierce, A persistent current magnetometer witjjj novel applications, Symp.

Physics of Superconducting Devices, Charlottesville, Va. (1967). Available as an ONR report.

[50] B. S. Deaver and W. S. Goree, Some techniques for sensitive magnetic measurements

using superconducting circuits and magnetic shields. Rev. Sci. Instr. 38,311 (1967).

[51] R. C. Jaklevic, J. Lambe, J. E. Mercereau, and A. H. Silver, Macroscopic quantum

interference in superconductors, Phys. Rev. 140, A1628 (1965).

[52] J. E. Mercereau, Modulated flux flow in superconducting films, Symp. Physics of Superconducting Devices, Charlottesville, Va. (1967). Available as an ONR report.

[53] W. M. Prothero, Jr., and J. M. Goodkind, A superconducting gravimeter, Symp.

Physics of Superconducting Devices, CharlottesviUe, Va. (1967). Available as an ONR report.

[54] R. A. Stokes, R. B. Partridge, and D. T. Wilkinson, New measurements of the cosmic microwave background at \ = 3.2 cm and \=1.58 cm — Evidence in support of a black body spectrum, Phys. Rev. Letters 19, 1199 (1967). This paper contains ref- erences to the earlier work.

[55] R. H. Dicke, Gravitation and the universe. Science J. 2, 95 (Oct. 1966).

[56] R. H. Kingston, A UHF soHd state maser, Proc. IRE 46, 916 (1958).

[57] R. C. Clauss, W. Higa, C. Stelzried, and E. Wiebe, Total system noise temperature:

15 K, IEEE Trans. MTT-12, 619 (1964).

[58] W. J. Tabor and J. T. Sibilia, Masers for the Telstar Satellite communications experi- ment. Bell Syst. Tech. J. 42, 1863 (1963).

[59] L. W. Bowles, Parametric amplifiers in the Lincoln experimental terminal, Nerem

Record 7, 212 (1965) (Published by E. E. Witschi, Jr., Newton, Mass.) See also: C. Blake, L. W. Bowles and E. P. McCurley, Helium-cooled parametric amplifier, Appl. Phys. Letters 2, 17 (1963).

[60] D. Bloor, J. J. Dean, G. O. Jones, D. H. Martin, P. A. Mawer, and C. H. Perry, Spectroscopy at extreme infrared wavelengths, Proc. Roy. Soc. A260, 510 (1961).

[61] E. H. Putley, SoHd state devices for infrared detection, J. Sci. Instr. 43 , 857 (1966).

[62] J. A. Westphal, B. C. Murray and D. E. Martz, An 8-14 micron infrared astronomical photometer, Appl. Opt. 2, 749 (1963).
CHAPTER 8

Applications

Robert A. Kamper, R. H. Kropschot, and A. F. Schmidt ^

Page 8.1. Magnets 329

8.1.1. Introduction 329

8.1.2. Superconductors 329

8.1.3. Non-superconductors 335

8.2. Linear Accelerators 337

8.3. Bubble Chambers 339

8.3.1. Introduction 339

8.3.2. Helium chambers 339

8.4. Cryopumping 343

8.4.1. Introduction 343

8.4.2. Space simulation chambers 343

8.4.3. Low-density wind tunnels 347

8.4.4. High-altitude air-sampling cryoprobe 350

8.5. Non-Condensable Gas Injection 351

8.5.1. Introduction 351

8.5.2. Injection cooling 352

8.5.3. Thermal and helium lift pumping 356

8.6. Other Apphcations in Missile and Space Systems 357

• Cryogenics Division, NBS Institute for Basic Standards, Boulder, Colo. 80302.

306-922 0-68—22 327
328 APPLICATIONS

Low-temperature research utilizing liquid helium has revealed

many unique properties of matter which are currently being exploited

for technological applications. These applications have stimulated much

of the work discussed in the earlier chapters. Thus it appears appropriate to present a brief description of some of these uses and to indicate

how the field may develop in the future. The reader is requested to bear

in mind that the material presented here is not meant to provide a comprehensive treatment of the subject and therefore selected references

are provided as a guide for further study. The intention is to illustrate the use of liquid or cooled gaseous helium for specific applications.

The chapter is divided into six major subdivisions where the in- dividual apphcations are discussed. Perhaps the most exciting and far- reaching low-temperature development is in the area of high-field super- conducting magnets. These and non-superconducting cryogenic magnets

are described in the first section. The second section describes how

superconducting resonance cavities can be utiUzed to construct highenergy Unear particle accelerators. The third section describes the use

of liquid helium in bubble chambers as a detection device for high-energy

physics investigations. The fourth section explains the use of helium

cooled surfaces for the production of high and ultra high vacuum by

cryopumping. The last two sections deal with non-condensable gas in- jection and some other aoDlications in missile and space systems.
MAGNETS 329

8.1. MAGNETS

8.1.1. Introduction

The reason for operating an electromagnet at a low temperature is to avoid the cost and inconvenience of supplying and dissipating a

large quantity of power. This can be achieved by making the windings

of either a superconductor or a very pure metal whose resistivity falls to a very small, but nonzero, value at low temperatures. Both alternatives have been used successfully to achieve flux densities over 10 tesla (100 kilogauss) [1, 2].^ They offer great promise for applications such as magnetohydrodynamic generators, plasma containers, satellite radiation

shields, particle accelerators, and research tools, where the cost, power

consumption or weight of the power supply for a conventional magnet,

working at or above ambient temperature, would be unacceptable. We

wiU discuss these two alternatives separately, concentrating mainly on

the hmitations set by the properties of materials.

8.1.2. Superconductors

Superconductors may be classified into several types according to their behavior in a magnetic field. In general, persistent currents flow

which partly screen a steady appKed field from the interior of a thick

superconducting body. Figure 8.1 shows how the resulting magnetic

moment, M, per unit volume of a specimen with zero demagnetizing

coefficient varies with the applied field H for the different types.

Figures 8.1(a) and 8.1(b) represent two ideal types to which very

carefully prepared specimens of real metals may approximate. They dif- fer in that the surface energy at a phase boundary, between superconducting and normal regions of the same metal, is positive in (a) and

negative in (b). Abrikosov [3] has called them ideal superconductors of type I and type II respectively. He has shown that the main features of

their behavior are described very weU by a phenomenological theory of superconductivity due to Ginzburg and Landau [4]. According to this theory the difference between type I and type II behavior depends upon

the value of a dimensionless parameter k which is equal to the ratio of two lengths, X/^, where \ is the depth of penetration of a weak magnetic

field into the surface of the superconductor and ^ is the characteristic

distance within which the wave functions of the electrons change, from

^ Figures in brackets refer to the literature references at the end of this chapter.
APPLICATIONS

Figure 8,1. Curves of volume magnetization M versus field H, illustrating the

difference between the types of superconductors: (a) ideal type I, (b) ideal type II, and (c) hard.
SUPERCONDUCTORS 331

the superconducting to the normal state, at a phase boundary. A type I superconductor has k < 1/V2 while a type II has k > I/V2.

In ideal type I superconductors the screening currents are confined to a surface layer of depth \ (typically 10~^ to 10~^ cm) and totally

exclude the field from the interior. When the appHed field exceeds the

critical field Hcb the whole superconductor makes an abrupt transition

to the normal state. Good specimens of most pure superconducting ele- ments (e.g., tin, indium, lead, etc.) fall into this class. The maximum current that can be carried without resistance by a type I superconducting wire is just that current which brings the total

field at any point on the surface of the wire up to the critical value Hcb

(Silsbee rule). This critical current decreases in an apphed field and

falls to zero at ^Hcb (the factor of ^ is the demagnetizing factor) for a circular wire in a transverse field. Since Hcb is typically a few hundreths

of a tesla, this sets a severe Hmit to the usefulness of this type of super- conductor. There has, however, been a proposal to use type I superconductors for power transmission [5], the equipment being designed so as

to hmit the maximum field to less than HcbThe above remarks refer to thick specimens. If one dimension of a superconductor, perpendicular to the field direction, is of the order of the penetration depth A. or less, the critical field is increased. For ex- ample, the Ginzberg-Landau theory predicts that a film of thickness

2did < k) with a smaU value of k would have a critical tangential field He where

Hc=V6 {kid) Hcb- (8.1) The Silsbee rule for critical current no longer applies. Instead, the

critical current density Jc tends to the value

J ^

4x10-8 Hcb

3 V6 ^ ©]3/2 MKS units (8.2)

in the Hmit of small d. This is typically of the order of W A/m^ when

H <^ He- There have been proposals [6] to make superconducting magnets with thin film windings. A successful thin film magnet would

require very careful attention to geometry, because the critical per- pendicular field [7] is just V2 k Hcb, which is less than Hcb for a type I superconductor.
332 APPLICATIONS

Type II superconductors behave like type I in small magnetic

fields. When the applied field exceeds a lower critical field Hc\{< Hcb),

however, they assume a mixed state in which the metal behaves as if it were finely divided into superconducting and normal phases, allowing

partial penetration of the magnetic field into the interior. This mixed

state persists up to a second critical field Hc2 = \^KHcb at which the

interior of a thick specimen is essentially in the normal state, but super- conductivity persists as a sheath on surfaces parallel to the field [8] up

to an upper critical field Hc3 = 1-69 Hc2- Good specimens of niobium and

many well annealed, single-phase alloys approximate to this type.

Below Hci it is expected that type II superconductors carry trans- port currents on the surface, like a type I superconductor, but in the

mixed state above Hci current paths in the interior would be mobile and

unstable against Lorentz forces, at least in transverse fields [9]. There

is experimental evidence that defect free type II superconductors do

indeed obey the Silsbee rule in transverse fields [10,11], with Hci as the

limiting surface field, while a longitudinal field can enhance the critical current by creating a force-free current path [12]. The behavior represented by figures 8.1(a) and 8.1(b) corresponds

to true equilibrium, so the same magnetization curves are obtained in both increasing and decreasing fields. The so-called "hard" superconductors are characterized by in- homogeneous structures due to the presence of several metallurgical

phases, lattice defects, or variations in composition. They usually behave

as shown in figure 8.1(c). They may be regarded as extreme type II super- conductors with the mixed state stabilized against Lorentz forces by

the irregularities in their structure, resulting in both magnetic hysteresis

and the ability to carry substantial transport currents distributed over

the entire cross section of a wire [13, 14]. These irregularities also have

the effect of decreasing the effective value of f and hence increasing

Hc2' There is, however, an upper limit to this effect which is imposed

by the paramagnetism of the conduction electrons in the normal state [15, 16^ 17]. It is unlikely that any superconductor will be found to withstand flux densities in excess of 50 T (500 kG).

Typical examples of hard superconductors which have been suc- cessfully used in electromagnets are cold-worked alloys of niobium with

zirconium or titanium and the intermetallic compound NbsSn. They

have all been observed to carry distributed current densities of the order

of 10 A/m2 at 4 K without resistance. The Nb-Zr alloys will carry sig-
SUPERCONDUCTORS 333

nificant current densities in flux densities up to about 7 T (70 kG) while

Nb-Ti and NbgSn can be used over 10 T (100 kG) [18, 19]. There is now a considerable quantity of literature on the design of superconducting magnets [18-22, 2]. The choice of superconductor

depends upon convenience and the performance required. The Nb-Zr

and Nb-Ti alloys are both ductile and can be wound into coils very easily.

NbaSn, which can show a better performance, is very brittle. It is nec- essary either to use a very thin layer deposited on a flexible substrate

or to wind the coil with the mixed but unreacted elements and form the

compound afterwards by heat treatment.

H

Figure 8.2. Typical variation of critical current

Ic with transverse field H for a short sample of a hard superconducting wire.

The design of the magnet then requires a knowledge of the critical current of the chosen wire as a function of transverse field at the working

temperature. Figure 8.2 represents a typical curve measured with a short

sample. Unfortunately, unless special precautions are taken, a long
334 APPLICATIONS

length of the same wire wound into a coil would probably not perform

as well as figure 8.2 would suggest, especially in the low field, high cur- rent range. Various explanations have been proposed for this degradation effect [19], but the one currently in favor was suggested by Riermersma, Hulm, and Chandrasekhar [23, 18]. The magnetization curve

for a hard superconductor is not usually observed as the smooth curve represented in figure 8.1(c), especially in the low field region. Instead,

magnetic flux is observed to enter and leave the superconductor in a

series of irregular steps known as "flux jumps." These are irreversible

and release pulses of heat. In a coil, a flux jump at one point may precipitate flux jumps in neighboring windings and nucleate a region in the

normal state. If the coil were carrying a sufficiently large current, the

resulting Joule heat could then spread this region throughout the windings, driving them into the normal state without exceeding the critical steady current (at the working temperature) at any point.

Stekly [24] suggested that a magnet could be stabilized against this effect by having the superconducting wire in electrical contact with a copper strip throughout its length, to share the current during flux jumps. His original criterion for the thickness of the copper was that it should be capable of carrying the full current while remaining below the zero current transition temperature of the superconductor, albeit with

a vigorous temporary boiloff of helium. However it was subsequently

found that considerably less copper is required [25, 26], resulting in a

big improvement in the winding packing factor. Using this principle it is possible to construct a magnet with a technologically useful overall current density which is free of the degradation effect. Another aspect of the transition to the normal state which requires

attention is the protection of the magnet from damage. Electrical transi- ents in a large, unprotected coil are capable of causing breakdown of the

insulation of the wire and local heating sufficient to spoil the heat treat- ment of the metal. Smith [27] has discussed several protection schemes.

Stekly's stabilized coils are also completely protected and offer a solu- tion to the problem which should be acceptable for large magnets, which

do not require high current densities.

The use of hard superconducting coils with alternating current is severely Hmited by hysteresis loss [14, 28-30]. This could conceivably

create a problem when running up the current in a large superconducting

magnet. A certain amount of heat will be hberated by the irreversible

process of attaining the final field and current distribution within the
NON-SUPERCONDUCTORS 335

superconducting metal, and provision must be made to absorb it. The

quantity involved is of the order of Q where [14]

Q^tJBJ/m^ (8.3) where t is the thickness of the wire in meters, J is the maximum current

density in A/m^, and B is the final flux density in tesla. The prospects and achievements in superconducting magnet

technology at the time of writing have been reviewed by Laverick [31].

8.1.3. Non-superconductors

We turn now to magnets using non-superconducting metals at low

temperature. A general discussion of these has been given by Taylor and

Post [32]. Corruccini [33] has also discussed the choice of metal. The

success of this type of magnet depends upon the reduction in the electric resistance of the windings, with a decrease in the working temperature,

overcompensating for the work required to extract a given quantity of

joule heat. The three major contributions to the resistivity of a metal are:

(a) scattering of electrons by lattice vibrations;

(b) scattering by impurities, dislocations, grain boundaries, etc;

(c) magnetoresistance.

To a first approximation, these three contributions may be simply added

together.

For most pure metals in attainable magnetic fields, ^a) dominates

the other two at room temperature. It may be reduced to an arbitrarily

small amount by reducing the temperature.

The residual resistivity to which a metal tends at very low temperature and zero magnetic field is a measure of (b). It can be reduced by

purifying and annealing. For an optimum cryogenic magnet material the

residual resistivity must be negligible compared to the other two contri- butions under working conditions, but in practice the cost of refining

may force a compromise which falls short of this. The relative change in resistivity Ap/p due to magnetoresistance in a

field H usually scales approximately as a function of H/p when p is changed either by changing the temperature or adding impurities,

although there are exceptions to this rule especially for very pure metals

at low temperature. The exact form of this function is a property of each
336 APPLICATIONS

individual metal and depends on the relative orientation of the field, current and crystal axes. For many metals the magnetoresistance is proportional to {HjpY at low fields, and it may or may not saturate at a constant value Ap/p ~ 2 or 3 at high fields depending on the electronic

structure. Thus the magnetoresistance in, a given field in general in- creases as the contribution of scattering |:o the resistivity is reduced,

and it imposes the ultimate limit on the attaihable efficiency of a magnet.

The choice of metal for a cryogenic magnet must depend upon the use for which it is intended. If it is only to be used intermittently, its chief advantage over a room-temperature magnet would be the lower

capital cost of the power supply, and the chief consideration in choosing

the winding metal would be a balance of its resistivity under working

conditions against its cost. For a more or less continuously running

magnet, however, the cost and power consumption of refrigerating plant

become more important. The total power consumption at a given cur- rent density would then become a more important consideration. This is the sum of the power delivered by the power supply and the power re- quired by the refrigerator (or liquefier) to remove it in the form of joule

heat. There are only six metals with which low-temperature operation is known to offer any advantage over copper at room temperature in this respect [33], assuming a magnet sustaining field of the order of 10 T. They are: aluminum, sodium, indium, hthium, copper, and beryllium

[34]. Of these, aluminum and sodium are distinctly better than the

others. For very pure samples, the optimum working temperatures [32] are about 17 K for aluminum and 10 K for sodium. At these temperatures

they both require a total power consumption as low as 10 percent of that required to sustain the same current density in copper at room temperature. It should be emphasized, however, that this does not represent a

direct comparison of the performance of actual magnets, since there is no reason why an optimum room-temperature magnet should work at the same current density as a cryogenic magnet for the same purpose.

The very pure metals used in cryogenic magnets are usually very

weak mechanically, especially as they must be used in an annealed state. The problem of supporting the windings against Lorentz forces while

maintaining cooling channels for efficient heat transfer may become

acute. It is usually necessary to interleave the windings with a stronger

metal.

Designs and working cryogenic magnets have been reported in the

literature by Taylor and Post [32], Purcell and Payne [35], Laurence and

Coles [2], and Laquer [21].
LINEAR ACCELERATORS 337

8.2. LINEAR ACCELERATORS

Liquid helium, below about 2.18 K, has many strange characteristics which are not possessed by any other material and therefore it has

been given the name "superfluid." It has been suggested [36, 37] that the use of the nearly zero electrical resistance of superconductors (exactly zero for direct current only) and the superfluid helium properties will revolutionize the construction of high energy particle accelerators. The

Stanford University group has constructed a 5 ft (1.5 m) long accelerator

of the type which will accelerate electrons to about 6 Mev (and probably

to 10 Mev with improved injection) and are constructing a 500 ft (150 m)

long machine.

The schematic of a bi-periodic 7r/2 mode accelerator is shown in figure 8.3. It consists of a series of microwave cavities connected to- gether so the entire system resonates as a single unit. The electric field in the short cavities is small while the field in the other cavities is along

the axis of the accelerator and alternates periodically in time. The dis- tance from one accelerating cavity to the next is adjusted such that the

electrons will obtain maximum acceleration.

Figure 8,3. Standing wave bi-periodic 77/2 mode accelerator.

The major advantage of using a superconducting cavity is its low

electrical surface resistance R (zero only for direct current) or high Q.

The ^ of a cavity is proportional to l/R and is an exponential function

of temperature [38] as shown in figure 8.4. Using very careful preparation

techniques, McAshan has been able* to obtain residual ^'s of about

18 X W for lead and 15 X W for niobium.
338 APPLICATIONS

MEASURED

THEORETICAL

TEMPERATURE Figure 8.4. Experimental and theoretical Q

of a lead superconducting cavity as a function of temperature.

The lower the temperature the smaller the power dissipation and

therefore superfluid helium will be used to cool the accelerator. Helium

below the superfluid transition or A. point has a very high thermal conductivity and estimates indicate that refrigeration at the center of the 500

ft (150 m) accelerator will be practical using a hehum thermal bus bar. A 9-in (23 cm)-diam tube of helium at 1.8 K will keep the accelerator to within ±0.03 K over its entire length. It would require a 100 ft (30 m)

diameter copper bar to do an equivalent job [37]. Superfluid helium, which has fascinated and entertained low temperature physicists for about 40 years is being put to work. Professor

Fairbanks and his group are utilizing the high thermal conductivity and

specific heat to do a job which would be extremely difficult in other ways.
339

8.3. BUBBLE CHAMBERS

8.3.1. Introduction

Helium bubble chambers are used as targets and detectors of high

energy particles for selected nuclear physics experiments. The bubble

chamber, as first suggested by Glaser [39], consists of a container of

liquid under sufficient pressure to be sub-cooled. When the pressure is reduced to below the normal boihng condition (superheated), the bubbles

nucleate around charged particles [40]. Bubble tracks define the trajectory of charged particles traversing the liquid. The chamber is con- structed with one or more windows to permit photographing the events and the entire unit is placed in a magnetic field so that the charged particles will have curved paths which in turn determine their momenta

[41]. Bubble chambers combine many of the advantages of cloud chambers and nuclear emulsions to give a high-density detection device.

Hydrogen and deuterium chambers have been used extensively as universal detectors and are the backbone of the high-energy nuclear

physics detection and analysis system [42, 43].

8.3.2. Helium Chambers

Liquid helium chambers are of specialized interest in the study of

particular events which occur in high energy physics reactions. For

example, when a beam of K~ particles impinges upon the helium chamber, some of the interesting events include the measurement of the

hfetime of aH^ and aH^ and the binding energies of AHe"* and aH^, where

the A indicates unstable excited states of the atoms known as hypernuclei. Also of particular interest is the study of the decay modes of these hyper-nuclei such as K~ + He^ aH^ + n-\-7T^ and the subsequent

decay of aH^.

Figure 8.5 shows a cut-away view of the Argonne-Carnegie Tech

10-in. helium bubble chamber. Table 8.1 gives the characteristics of

this and other existing helium chambers.

The operating conditions for good quality bubble tracks is a fluid temperature of about 3.5 K and an operating pressure which varies

between about 200 and 500 torr. As a high-energy charged particle passes

through the hquid it ionizes some of the atoms giving rise to free electrons. These electrons are nucleation sites for boiling in a superheated
340 APPLICATIONS

Expansion ^ Bellows-. E

Light Source

Refrigeration Source

Camera

Superconducting

Magnet

Figure 8.5. Schematic of the Argonne-Carnegie Tech 10-in helium

bubble chamber.

Particles enter at right angles to the optical system.

Table 8.1.

Source Vol /* Expansion

system

Field Refrig Reference

3 Piston He reservoir [44] 4 Piston He reservoir [45] Oxford/NIRNS 150 Bellows 20Kg 80 W refrig [46] Argonne 20 Bellows 45Kg He reservoir [47] Northwestern 45 Bellows 28Kg He reservoir [48]

*These values are approximately the volume available for photographing. The value

for thermodynamic calculations will be larger.
HELIUM CHAMBERS 341

liquid and bubbles will grow if their radius is greater than the critical value Tc where

8P'

cr is the surface tension, 8P = Pi — Pv, Pi is the pressure of the liquid

and Pi is the saturated vapor pressure in the bubble. The critical radius

for helium is approximately 1 X 10~^ cm.

Figure 8.6(a) shows a schematic of an ideal pressure-time cycle

(solid line) and an actual (dotted line) pressure-time cycle. The expansion is not an equilibrium process and therefore cannot be represented on a conventional pressure-volume or temperature-entropy diagram.

However, it is useful to use such a diagram for estimating the amount of

refrigeration which will be required. Figure 8.6(b) is a schematic of a temperature-entropy diagram for hquid expansion. The area inside the curve must be estimated in order to adequately size the external

refrigerator.

The external refrigerator must be sized such that it will cool the

entire chamber to operating temperature in a reasonable time, remove

heat transferred to the chamber from external sources, and remove the

work of compression during operation. The work of compression for existing chambers is of the order of 0.05 to 0.2 joules/pulse/liter. The

specific value is dependent unon the irreversibility of the expansion cycle

and conceivably could be made much less than existing values by careful

cryogenic design. Although gas expansion has been frequently used for hydrogen, it has not been used in helium chambers because of helium's

low heat of vaporization and the large refrigeration load it would cause. Small work of compression is desirable primarily because of the greater ease of maintaining temperature stability as well as the high cost of

refrigeration. Heat from external sources can be reduced greatly by

use of nitrogen-cooled radiation shields, but in any case it can be esti- mated accurately from knowledge of the physical properties of the

materials of construction.

A uniform temperature within the chamber yields uniform track

densities and thus allows particle energy determination using bubble

densities. Whereas, small chambers can operate using a reservoir of

liquid and only the latent heat of vaporization, a large refrigeration load

or an extended time of operation will very easily justify the capital

investment for integral refrigeration which utilizes the enthalpy of the

cold gas in the cycle [49].
342 APPLICATIONS

v>

a>Q.Eoo

Accelerator

/^Pulse Photo

1

7—^17

Bubble J

G.'owth I TIME

Figure 8.6(a). Schematic of pressure-time cycle for bubble chamber operation.

Figure 8.6(b). Temperature-entropy

diagram for liquid expansion.

Process is non-equilibrium and therefore cannot be

traced point by point.
CRYOPUMPING 343

8.4. CRYOPUMPING

8.4.1. Introduction

The production of high- and ukra-high vacuum can be enhanced

through the utiHzation of cryogenic cooUng, where the gas is condensed,

sorbed or trapped on a low temperature surface [50-57]. Alternate

schemes for producing high vacuums include ion, subhmation, and

sputter type pumps which operate on the similar principle of capture

or retention of mass from the gas phase, as well as the more conventional momentum transfer, mass removal pumping techniques. The

principal advantages of the cryopump over these other pumping schemes

are: simplicity, high pumping efficiency, low ultimate pressures, no moving parts, no premature shut-downs due to burn-out of filament

needed for ionizing current, no heavy magnets or high voltages, etc.

It is a technique that provides a clean, oil-free system in which the

pumping capabilities for atmospheric gases closely match their respective partial pressures in the atmosphere. Limitations include: the

necessary size requirements of cryosurface areas, cryosurface capacity

(operation time between periods of surface regeneration), temperature

rise across the deposit due to the heat of condensation, unfavorable

thermal properties of the deposit, and the limited ability to pump

certain low boiling point gases. Even the removal of gases which are normally termed non-condensable can be augmented by cryogenic

cooling. The cryogenic cooling of a plane surface can increase the

pumping capacity of that surface; however, more effective utilization of this scheme is realized by providing extended surfaces such as molecular sieves, siHca gel, charcoal or the soHd particles of a condensed gas. Three major research areas where cryogenic pumping is beneficial are space chambers, low density wind tunnels, and high altitude air samplers.

8.4.2. Space Simulation Chambers

It has not proven feasible, nor is it possible at this time, to simulate

either individually or collectively all of the various aspects of a true space

environment. However, during the past half-decade or so, considerable

strides have been made in the development of chambers capable of accommodating and producing many of the extreme conditions of space

[58-70]. The high cost of space ffight coupled with the consequences of

insufficient reliabihty have dictated the necessity for simulation testing on the ground.

306-922 0-68—23
344 APPLICATIONS

Inasmuch as this discussion is primarily concerned with the several

effects of refrigeration on simulations as afforded by Hquid hehum or cold

helium gas, the principal environmental conditions of interest are asso- ciated with the low temperature and pressure of space. The infinite heat sink of space is known to have an effective temperature as low as that of liquid helium. This, in itself, is not a criterion

which requires that a simulated space radiation heat sink be at this temperature, since the fourth power Stefan-Boltzmann radiation law

indicates that only small error can be derived by substituting sink temperatures on the order of 80-100 K (providing that the radiating test body is maintained no lower than 300 K). Of greater importance is the

fact that surfaces cooled to a temperature only slightly greater than

absolute zero are able to provide increased molecular retention capability, thereby reducing the reflection or rebounding of gas molecules

back to the test piece from the chamber walls after originally departing

from that surface. The value of an extremely low temperature is in the

establishment of a molecular sink more closely approximating the infinity

of space. In certain small systems or experiments where true temperature

and pressure simulation is deemed necessary, liquid helium has been

used as the cryopanel refrigerant in large quantities and with attendant

cost. This is possible at the few facilities fortunate enough to possess

sizeable helium liquefaction capability, but even then it is expensive

and requires good justification. In order to achieve the extreme high vacuum (or low pressure) of

space, it is generally recognized that various cryopumping techniques

are invaluable, if not indispensable. With the exception of neon, hydrogen, and helium, all of the constituents of air have vapor pressures

lees than 10"^^ torr at a temperature of 20 K, which is the normal boihng

point of hydrogen. Assuming that the process of cryogenic condensation

will be used to take out the primary gas load in a typical cryopumped

simulator, and assuming that cryo-sorption, -trapping, -gettering, or some other method is used to remove the noncondensables, it is seen

then that the temperature of liquid hydrogen may be considered an upper limit to many test requirements. However, although liquid hydrogen is now in plentiful supply and reasonably inexpensive, its use constitutes a safety hazard which is not warranted for such an appfication.

Instead, dense hehum gas refrigeration usually is adopted at a working

temperature level of 10-20 K. In order to minimize the heat load and

hence the appreciable cost of low-temperature refrigeration to the

hehum cooled cryopanel, a hquid nitrogen-cooled thermal shroud or
SPACE SIMULATION CHAMBERS 345

Figure 8.7. Typical shielded cryo- pump arrays.

In each geometrical arrangement illustrated here, the liquid nitrogen cooled radiation shielding

is positioned to maximize pumping speed of the hdium cooled cryopump condenser surface while minimizing the high temperature thermal radiation load.

Test Object

, ^«««««««C — .K sy

20 "K-^

Test Object

20 °K (-77

fbj

Test Object

Test Object 77 °K 20 "Kshield is provided to intercept incident radiation. Typical shielded

cryopump arrays are shown in figure 8.7. As one example, each of the

39-ft spherical General Electric Valley Forge space simulators [71, 72]

shown in figure 8.8 consist of a 1-kW (below 20 K) helium gas cryopanel

surrounded by a 150-kW liquid nitrogen shroud in two of the chambers

and a 300-kW nitrogen shroud in the third. The origin of gases to be evacuated from a space chamber generally consists of (a) an initial or residual volume of air in the chamber,

(b) desorption from the chamber walls and test body, and (c) leaks. However, a somewhat different category of near-space propulsion test- ing requires also that rocket-exhaust products be collected and con- densed on an appropriate cryogenically-cooled surface. The AEDC

Engine Test CeU J-2(A) [73] (figs. 8.9 and 8.10) consists of an 18-ft (5.5 m) diam by 32-ft (9.8 m) long nitrogen-cooled vacuum chamber

equipped with a gaseous helium refrigerated system of cryoplates capable of absorbing a 1-kW heat load at a maximum refrigerant return temperature of 18 K.
346 APPLICATIONS

Figure 8.8. 39-ft diam General Electric space test chambers (Valley Forge, Pa.).

Figure 8.9. AEDC Propulsion Engine Test Cell (J-^A).
LOW-DENSITY WIND TUNNELS 347

Figure 8.10. AEDC Propulsion Engine Test Cell {J-2A).

8.4.3. Low-Density Wind Tunnels

As a result of intensified interest in the exploration of space and the

upper atmosphere, it has become increasingly desirable to develop large

low-density wind tunnels [74] capable of operating over extensive

periods of time at hypersonic velocities in the transitional flow regime,

i.e., altitude range between 200,000 ft (60 km) and 500,000 ft (150 km).

In this region between continuum and free-molecule flow, neither of the bounding flow regime concepts have been sufficient to provide

quantitative design information or to explain certain aerodynamic

phenomena.

The first large-scale apphcation of cryopumping occurred in the

spring of 1959 when the University of Southern California low-density

wind tunnel [75] became fully operational. A helium refrigerator, originally constructed as part of a refrigerated liquid hydrogen transport

dewar, was used to cool the condenser. The information obtained on

this system has been invaluable in designing more recent cryotunnels.
348 APPLICATIONS

The choice of refrigerants for such appHcation is usually limited

to neon, hydrogen, or helium using either an open or closed refrigeration cycle. Closed cycle systems are more economical over the long run whereas an open cycle would require only moderate capital investment

but high operating cost. In view of the potential safety hazard of hydrogen, and the low-temperature limitation and cost of neon, hehum

refrigeration appears to offer the best solution for a large cryopumo.

In addition to the primary cryosurface considerations, tunnel design

criteria require that liquid nitrogen-cooled precoolers be installed up- stream of the final condenser in order to decrease the refrigeration load

(by reducing the heat content of the flowing gas) on this condenser; a

liquid nitrogen-cooled shroud is also furnished to protect the primary

element from thermal radiation. The importance of this is illustrated by

the fact that approximately 70 percent of the available refrigerator

power in the early USC tunnel was dissipated in radiation losses, leaving

only the remaining 30 percent to useful condensation. To alleviate the

problem of non-condensables in the tunnel working fluid stream, it has

been found desirable to substitute high-purity nitrogen for the atmospheric air. For testing purposes, its properties are similar to air, and it is readily condensed on the cryosurface. Also, to increase the useful

operating time of a cryotunnel, means have been considered to remove

Figure 8.11. AVCO precooler and cryopump condenser.
LOW-DENSITY WIND TUNNELS 349

the built-up condensate on the cold surface(s). Periodic scraping, or a

flexible surface, would permit continuous operation if desired. The

AVCO tunnel design [76] suggests radiative heating of the cryodeposit

in such a manner that a minimum amount of heat is transferred to the

cooHng fins. This causes the condensate temperature to be raised above

its triple point, permitting it to drip off and be removed in the Hquid

phase. A schematic layout of the AVCO hardware is presented in figures 8.11 and 8.12.

FIN CONSTRUCTION DETAIL FOR LOW MASS FLOWS

Figure 8.12. AVCO condenser fin arrangement and construction detail.
350 APPLICATIONS

8.4.4. High-Altitude Air-Sampling Cryoprobe

Included in a variety of devices which have been, or are being,

developed in the interest of examining the constituents of the upper

atmosphere are several which make use of cryogenic collection tech- niques. One of these, the Nesco cryoprobe [77], has been designed to collect a whole-air sample volume (including particulate matter) of up

to 800,000 ft^ (20,000 m^) within an altitude range of 300,000 to 400,000

ft (90 to 120 km). As part of the USAF's air-launched, air-recoverable

rocket (ALARR) program, the payload is fitted to a Genie rocket booster

and launched by supersonic aircraft; recovery is accomplished using a conventional air-snatch technique, with the sampler then flown to the

laboratory for sample evaluation and analysis [78].

Figure 8.13. Nesco cryoprobe paddle deployment.
NON-CONDENSABLE GAS INJECTION 351

In-flight cryopumping is achieved by a large hehum-cooled parasol

(development of which is shown in fig. 8.13) unfurled by centrifugal

force to sweep the prescribed flight path in 30 to 60 s. After sampling,

the collector surface is refolded and stored in its original nosepiece.

Approximately 20 lb (9.1 kg) of indium — a metal having the highest

specific heat capacity of any useful to the program — are fabricated

into 0.05 in (1.3 mm) blades which then are sandwiched between 0.015 in (0.38 mm) gold-plated aluminum face sheets; the plating improves

bonding to the indium and provides a highly reflective surface for solar heat rejection. On the basis of both theoretical and experimental analysis, it has been determined that no shock wave forms at the face of the blades as long as the temperature does not exceed 32 to 38 K in the applicable sampling altitude range. Therefore the temperature of 4 K when deployed can rise to something less than 32 K with small

decrease in efficiency. The cryoprobe functions first as a vacuuminsulated dewar, retaining liquid hehum about the sampling paddles

for several hours prior to the actual collection period. Following col- lection, the condensate turns to gas with rising temperature and the

container becomes a low-pressure vessel.

8.5. NON-CONDENSABLE GAS INJECTION

8.5.1. Introduction

The involvement of cryogenic propellants in missiles and space

vehicles has provided the aerospace industry with a multitude of problems, and solutions for several of these have relied upon the avail- abihty and unique characteristics of hehum. As a result of its low

boiling point or non-condensabihty, low solubility, low molecular weight,

and, because it is inert, helium has many cryopropellant applications. In this section, examination is made of the use of low-temperature

helium gas injection as a cryogenic fluid conditioning technique. If a non-condensable gas is bubbled through a volume of liquid, evaporative

cooling — the vaporization of some of the liquid into the gas bubbles —

causes the temperature of the hquid to be lowered. When properly

applied to a tank or pipehne of hquid hydrogen or hquid oxygen, for example, injected hehum is able to refrigerate and thereby reduce the

temperature of the hquid [79-82], suppress geysering [83], and create

or maintain solid or solid-liquid propellant mixtures [84, 85]. Various
352 APPLICATIONS

cryogenic propellant stages of the Saturn family of space vehicles make

use of helium injection as a means of conditioning propellant in the

suction lines of the pumps. Several gas injection-based methods are described below.

8.5.2. Injection Cooling

The liquid duct connecting the bottom of the cryopropellant tank

to the rocket engine-feeding turbopump furnishes the physical system

for a variety of undesirable phenomena. The combined effect of suction

line length, diameter, and environmental heating is able to cause an

excessive temperature rise in the contained fluid v^hich either can be

objectionable or impossible to accommodate by the vehicle system.

Should insufficient net positive suction head (pump inlet total pressure minus vapor pressure) be available to the pump, cavitation

may occur and the mission will be jeopardized; since vapor pressure

is a function of temperature, this latter quantity must be controlled in some way to maintain the liquid in a sufficiently subcooled condition.

If, due to ambient heating at anytime prior to launch, the column

of liquid becomes superheated, geysering can result from a violent

generation of vapor caused by the release of superheat, which forces

liquid from the line upward into the propellant tank. Cooling from the

vaporization process, and colder liquid falHng back through the residual vapor, sets up a condition wherein the momentum of dropping liquid

must be absorbed by the bottom of the fluid column. This results in a large pressure spike — at times a factor of ten or more greater than the

initial pressure — which can be extremely destructive to the confining

hardware. A mechanism to refrigerate (or somehow adjust the temperature of) this pipe of liquid therefore becomes necessary to preserve the

integrity of the fluid system. Injection cooHng can be used to reduce or eliminate both cavitation and geysering.

A more recent development of this technique concerns the formation

and maintenance of solid-liquid (slush) mixtures of cryogen. The ad- vantages of creating and using such mixtures lie in their ability to absorb more heat before any evaporation occurs — i.e., to provide an

increased heat sink — and to require a smaller container volume due to greater density of the solid relative to the Hquid. Injection coohng is not considered to be an economically justifiable large-scale production

method, but it may well be an excellent way to upgrade the mixture

within a thin-walled space vehicle propellant compartment. Upgrading

is necessary because of the uncertain, or long, periods of ground-hold
INJECTION COOLING 353

or standby time, imperfect tank insulation characteristics, and a possible requirement to increase the soUd percentage above a flowable

mixture.

Gas injection cooUng is accomplished in the following manner

[81. 82]. Assuming that the non-condensable hehum gas being injected

into a Uquid cryogen is pure, then the partial pressure of the cryogen

within any given helium bubble is zero. A portion of the hquid is required

to evaporate into the gas bubble since a difference has been established

between the vapor pressure of the cryogenic fluid surrounding the bubble

and the partial pressure of the cryogen within the bubble. The natural

result of vaporization is a cooHng effect whereby heat is extracted from

the Hquid and its temperature decreased accordingly.

A heat balance of the given control volume of figure 8.14 yields

q =qi -\- Qo — qs where q is the total system heat flux, qi the environmental heat flux, q-z the injected hehum heat flux, and qs the evaporative cooling rate. Using

the simplifying assumptions of instantaneous cool-down of helium gas to the temperature of Hquid cryogen and instantaneous diffusion of cryogen

vapor into the heHum gas (the model does not apply for a system where

Figure 8.14. Non-condensable gas injection model.
354 APPLICATIONS

TIME, minutes

TIME, minutes Figure 8.15. Comparison between calculated and experimental temperature his- tories for different injection gas flow rates.

Solid line experimental, dotted theoretical. (20 scfm = 0.57 m^/min.)

significant diffusion of gas to liquid occurs), a comparison of experiment

and theory is shown in figure 8.15. The correlation is quite good for two

rather extreme values of injection flow rate -0.1 scfm (0.0028 m^/min)

and 20 scfm (0.57 m^/min) of hehum into Uquid hydrogen.

Figure 8.16 is a graphical representation of the analytical model

for a specific helium injection flow rate of 10 scfm (0.28 m^/min). An

environmental heating load qi of 2.7 W/m^ of wetted column wall area

is used as an example but substitution may be made for any value. Also,

since the injected hefium heat flux qz and the evaporative cooHng rate

qs vary directly as the injected gas flow rate, these curves may be altered

by simple ratio, e.g., (^2)20 scfm = (2^2)10 scfm, and (^3)20 scfm = (2^3)10 scfm- The

curves are useful in predicting the minimum theoretical temperature

attainable for a given set of conditions. That is, for a specified gas flow

rate and a known system heat load, the minimum temperature may be

found simply by setting dT/dO (and therefore q) equal to zero, from which

91 + ^2 — <73 = 0; graphically (fig. 8.16 or equivalent), the temperature at qi~^q2 = qz is the desired temperature.
INJECTION COOLING 355

Figure 8.16. Graphical representation of the analytical model at 10 scfm

{0.28 m^lmin).
356 APPLICATIONS

8.5.3. Thermal and Helium Lift Pumping

Where more than one column of liquid may be considered usable

in a fluid conditioning system — many space vehicles are now equipped

with multiple propellant lines due to the parallel clustering of both

engines and propellant tanks — several liquid recirculation schemes

have been developed where in a given set of two lines each is interconnected near the bottom. Whether one tank, or several manifolded tanks,

are used to supply these line combinations makes little difference as long as a continuous circulation path is provided (see fig. 8.17) from

tank(s) to line A through connector to line B to tank(s), etc. Since we

are concerned with vehicle standby and engine prestart suction line

fluid conditioning (i.e., dead-ended lines with nonoperative pumps),

hne A will be designated "feed" and line B "return". Assume now that the two ducts are filled with liquid at diff"erent temperatures and that they open into a common tank filled with hquid at the colder temperature. Density differentials in the lines due to this temperature difference then cause a natural convection, or thermal

pumping, process to be established which moves fluid from the cold

column into the warm column, warm fluid into the tank, and cold tank

Figure 8.17. Thermal and helium lift pumping model.
OTHER APPLICATIONS IN MISSILE AND SPACE SYSTEMS 357

liquid down into the colder line as replenishment. This continues as long as the feed and return line temperatures are different, which is a function of the normal ambient heat flux to the columns and a continuing

supply of cold Hquid from the propellant tanks.

If it is found that the natural circulation process outlined above is not quite sufficient to do the job, and that it is undesirable or impossible

to insulate the cold line and/or adjust line sizes, restrictions, etc., then

it becomes necessary to implement the flow by some other means. By

injecting helium gas into the bottom of the return line, the effective mean

density of the two-component fluid system within this line is decreased,

with subsequent increase in density differential between lines and con- comitant rise in flow driving force. From the previous section on injection

cooling, it will be recalled that the evaporative cooling effect of non- condensable gas introduced here eliminates geysering by temperature

reduction of the affected (return) flne liquid. Also, it has been verified experimentally that flow oscillations in such a system at low (filling or

draining) propellant tank head are stabilized due to the accelerated flow

rate and subcooled liquid resulting from injection.

Analysis of this so-called helium lift pumping process is complicated because of the interdependency of several variables — two-phase

flow, liquid-gas interfacial mass transfer, etc. However, computer pro- grams have been written by a group at the NASA-Marshall Space Flight

Center which predict steady state recirculation flowrates and temperature distributions for both helium lift and thermal pumping systems.

Extensive analytical and experimental investigations of these systems

at Marshall have proven the feasibility for space vehicle application [86].

8.6. OTHER APPLICATIONS IN MISSILE AND

SPACE SYSTEMS

Because of the many desirable properties of helium which were

listed in a previous section, the fluid enjoys multi-faceted utilization in space vehicle applications [87, 88]. As propeUant pressurant in a pump-fed engine system, the gas is used to displace the liquid fuel or oxidizer and to guarantee that pro- pellant reaches the pump at a required pressure and usable fluid condition. The fact that helium is an inert gas, non-condensable and only

very slightly soluble (therefore non-contaminating) in these propellants,

and is of low molecular weight, identifies several good reasons for its
358 APPLICATIONS

choice of suitability as a pressurization gas. The helium is carried on- board in high pressure vessels, either external to, or within, the propellant tanks. For an internal configuration the gas loading is maximized

since the fluid density is increased by a low temperature environment;

however, subsequent to use, the temperature of the gas should be in- creased for efficient use at low density. Heat exchange is accompHshed

generally at some point in the connecting fine between high pressure

gas storage and tank ullage where the heat sink of the cold helium gas

can, perhaps, be used to some advantage.

In the liquid hydrogen-fueled Centaur upper stage, liquid helium

is employed as a pre-launch engine coolant [89]. The requirement here

is based upon a need for near-operating temperatures of engine hardware

prior to ignition to forestall the possibility of turbopump cavitation and

overspeeding. A previous solution was to bleed liquid hydrogen from the

fuel tank in flight after booster separation, but this procedure involved a valuable time period of roughly one-half minute and additional propellant

to be carried solely for its refrigeration capacity. Results of the decision

to pre-chill the engine with liquid helium on the ground before launching

shortened the in-flight chilldown time by a factor of four and permitted a payload increase of 50 lb. At the expense of additional development

time, a refrigerated helium gas system could have been devised to accomplish the same purpose.

Because of its low condensing point temperature, hehum gas is also of value in certain extensive space travel pneumatic control systems

where any long term exposure to a low temperature environment might

freeze the more conventional actuating fluids. Cryogenic gas bearings

are often lubricated with helium gas if the rotating device is required to operate at very low temperatures. The characteristics of helium pre- viously mentioned as being beneficial for a pressurant are equally advantageous in purging operations which ultimately involve the handling of cryogens. Helium purge gas does not freeze, contaminate, or react with

subsequently introduced low boiling point fluids.
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360 APPLICATIONS

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[25] Z. J. J. Stekly and J. L. Zar, Stable Superconducting Coils, IEEE Trans. Nucl. Sci. NS-12, 367 (1965).

[26] C. Laverick, G. M. Lobell, J. R. Purcell, and J. M. Brooks, Use of aluminum in super- conducting cables, Rev. Sci. Instr. 37, 806 (1966).

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[29] C. H. Jones and H. L. Schenk, A.C. losses in hard superconductors. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1963), Vol. 8, p. 579.

[30] T. A. Buchhold, The nature of the surface losses of superconductors at low frequencies, Cryogenics 3, 141 (1963).

[31] C. Laverick, The future of superconducting magnets, IEEE Trans. Nucl. Sci. NS-14,

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[32] R. F. Post and C. E. Taylor, Air core cryogenic magnet coils for fusion research and

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NBS Tech. Note 218 (1964).

[34] P. Burnier, Cryogenics and aluminum in electrical manufacturing. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1966), Vol. 11,

p. 668.

[35] J. R. Purcell and E. G. Payne, High-field liquid Ha-cooled aluminum-wound magnet.

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[36] H. A. Schwettman, J. P. Turneaure, W. M. Fairbanks, T. I. Smith, M. S. McAshan,

P. B. Wilson, and E. E. Chambers, The 1967 U.S. National Particle Accelerator Conference (to be published).

[37] W. M. Fairbanks, A superconducting linear accelerator and the use of superconductivity in some fundamental experiments in physics, Proc. Symp. Physics of Superconducting Devices (ed. B. S. Deaver, Jr., Univ. of Virginia, April 28-29, 1967),

p. A-1.

[38] M, S. McAshan, The application of superconductors in the construction of high-^

Microwave cavities, Proc. Symp. Physics of Superconducting Devices (ed. B. S. Deaver, Jr., Univ. of Virginia, April 28-29, 1967), p. C-1.

[39] Donald A. Glaser, The bubble chamber, Handbuch der Physik (ed. S. Flugge, Springer,

Berhn, 1958), Vol. 45, p. 314.

[40] F. Seitz, On the theory of the bubble chamber, Phys. Fluids 1, 2 (1958).

[41] D. F. Shaw, Hydrogen and heUum bubble chambers. Cryogenics 4, 193 (1964).

[42] R. P. Shutt, Recent advances in the bubble chamber technique. Nuclear Instr. Methods 20, 71 (1963).

[43] N. C. Barford, Low temperature bubble chambers. Progress in Cryogenics 2, 88 (1960).

[44] M. M. Block, W. M. Fairbanks, E. M. Harth, T. Kikuchi, C. Mehzer, and J. Leitner, Proc. Intern. Conf. High-Energy Accelerators and Instrumentation (C.E.R.N., Geneva, 1959), p. 461.

[45] E. Di Capua, U. Dore, G. C. Gialanella, P. Guidoni and I. Laakso, Nuclear Instr. and Methods 15, 273 (1962).
REFERENCES 361

[46] J. Moffatt, Nuclear Physics Laboratory, Oxford (1960) (unpublished report) (see ref. 41). [47] E. G. Pewitt, M. Derrick, T. H. Fields, L. Hyman, C. Laverick, K. B. Martin, J. G. Fetkovich, and J. McKenzie, The 19th Intern. Conference on High Energy Physics,

Dubna, U.S.S.R. (1964).

[48] M. M. Block, Northwestern University (unpublished data).

[49] D. B. Mann, Chapter 2. [50] H. Adam. The significance of cryogenic techniques for the production of low pressures. Rev. Soc. Roy. Beige Ingrs. Ind. 11 457 (1963).

[51] J. P. Dawson and J. D. Haygood, Cryopumping, Cryogenics 5, 57 (1965).

[52] P. J. Gareis and G. F. Hagenbach, Cryosorption, Ind. Eng. Chem. 57, 27 (1965).

[53] E. L. Garwin, Cryogenic pumping and space simulation, Advances in Cryogenic

Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1963), Vol. 8, p. 37.

[54] E. L. Garwin, Cryogenic pumping and space simulation. Cryogenic Technology (ed. R. W. Vance, John Wiley & Sons, Inc., New York, 1963), p. 332.

[55] G. F. Hagenbach, Cryogenic vacuum pumps, Ind. Res. 7, 67-68, 70-72, 74 (1965).

[56] P. B. Henault, P. J. Fennema, and B. A. Buffham, Cryogenic pumping, J. Environ- mental Sci. 6, 15(1963).

[57] L. O. Mullen and M. J. Hiza, The role of cryogenics in the production of high and

ultra-high vacuum. Cryogenics 4, 387 (1964).

[58] J. L. Aberle and A. J. Westbrock, Liquid helium and nitrogen supply systems for space

simulators. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum

Press, New York, 1963), Vol. 8, p. 190.

[59] C. W. Alstrom, Design of extreme high vacuum facilities. Test Eng. 11, 28-30, 34, 41 (1964).

[60] N. C. Breddy and H. J. Smith, Some problems in space simulation chamber design.

Vacuum 13,449(1963).

[61] E. E. Callaghan, Hard vacuums and cryogenics. Machine Design 34, 162 (1962).

[62] I. Farkass and G. W. Horn, Cryogenic pumping in space simulators, A.I.Ch.E., 56th Annual Meeting, Symp. Molecular Properties at Cryogenic Temperatures, Houston,

Tex. (1963).

[63] S. Giles, Cryogenic pumping systems. Test Eng. 12, 38 (1964).

[64] B. H. Goethert and H. M. Cook, High ground for aerospace simulation, Astronaut. Aeronaut. 2,96(1964).

[65] R. A. Hindle, Space simulation, ASME Paper No. 62-AV-34 (1962).

[66] R. T. HoUingsworth, A survey of large space chambers, NASA Tech. Note D-1673

(1963) . [67] H. Mark and R. D. Sommers, The combined use of liquid and gaseous helium to pro- vide near actual space environment, Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1963), Vol. 8, p. 93.

[68] V. Outmann and J. C. McLane, Jr., Testing in the simulated space environment —

an assessment, AIAA Paper No. 65-474 (1965).

[69] J. C. Simons, Jr., and M. P. Hnihcka, Future trends in space simulation — the role of vacuum and cryotechniques, Inst. Environ. Sci. Proc. (1963), p. 363.

[70] K. D. Timmerhaus, Cryogenics in space simulation. Test Eng. 12, 26, 28-29, 34, 42

(1964) . [71] J. Richman, Space simulation facility. Test Eng. 11, 16-18, 22(1964).

[72] J. Richman and C. B. Hood, Report on three 32,000 cu. ft. space simulation systems.

Trans. 9th Natl. Vacuum Symp. (ed. G. H. Bancroft, MacMillan Co., New York,

1962), p. 282.
362 APPLICATIONS

[73] J. R. Reeves, Jr., General description and performance of the propulsion engine test

cell (J-2A), Arnold Engineering Development Center, Arnold AF Station, Tennessee, Report No. AEDC-TDR-64-138 (1964).

[74] W. A. Clayden and R. B. Reynolds, Cryopumped low density wind tunnels. Vacuum

13,461(1963).

[75] J. G. Everton, Operational characteristics of a cryopump used in a low density wind

tunnel. Trans. 9th Natl. Vacuum Symp. (ed. G. H. Bancroft, MacMillan Co., New

York, 1962), p. 227.

[76] B. D. Henshall and E. M. Brower, A cryogenic hypersonic low-density wind tunnel. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New

York, 1962), Vol. 7, p. 73.

[77] National Engineering Science Company, Cryogenic air sampler for high altitude appHcation, NESCO Rept. SN lOOA-1 (1964).

[78] C. M. Plattner, Super-cooled paddles used in high altitude air sampler. Aviation Week and Space Technology 82, 44-45, 47 (1965).

[79] P. S. Larsen, J. A. Clark, W. O. Randolph, and J. L. Vaniman, CooHng of cryogenic

liquids by gas injection. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus,

Plenum Press, New York, 1963), Vol. 8, p. 507.

[80] F. W. Lytle and J. T. Stoner, Cryogenic cooling by noncondensible-gas injection. Science 148,3678, 1721 (1965).

[81] W. O. Randolph and J. L. Vaniman, Sub-coohng of cryogenic liquids by injection of non-condensing gas, NASA-Marshall Space FHght Center, MTP-S&M-P-61-19

(1961).

[82] A. F. Schmidt, Experimental investigation of liquid-hydrogen cooling by helium gas

injection. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1963), Vol. 8, p. 521.

[83] S. K. Morgan and H. F. Brady, EHmination of the geysering effect in missiles. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York,

1962), Vol. 7, p. 206.

[84] G. J. Caras, Slush hydrogen, Redstone Scientific Information Center, Redstone

Arsenal, Alabama, Report No. RSIC-288 (1964).

[85] R. R. Carney, R. F. Dwyer, H. M. Long, and L. R. Niendorf, Theoretical, experimental, and analytical examination of sub-cooled and solid hydrogen, Air Force Tech.

Doc. Rept. No. APL-TDR 64-22 (1964).

[86] H. F. Trucks and W. O. Randolph, Analytical and experimental investigation of thermal and helium lift — pumping recirculation systems, Advances in Cryogenic

Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1965), Vol. 10, p. 341.

[87] W. M. Deaton, Helium for missiles and space vehicles. Missile Liquid Propellants

Symp., Hershey, Pa. (1964).

[88] Rocket Propellant and Pressurization Systems (ed. ElHot Ring, Prentice-Hall, Inc., Englewood CUffs, N.J., 1964).

[89] J. J. Gilbeau, D. L. Martindale, and K. R. Burton, A large-scale storage and transfer system for distribution of liquid helium as a refrigerant medium. Advances in Cryogenic Engineering (ed. K. D. Timmerhaus, Plenum Press, New York, 1965), Vol. 10, p. 428.
Author Index

A

Abeles, B., 311, 324

Aberle, J. L., 70, 76, 343, 361

Abrikosov, A. A., 329, 359 Adam, H., 343, 361 Akin, S. W., 35, 53

Allen, R. J., 319, 326

Alstrom, C. W., 343, 361 Anderson, C. C, 3, 32 Arams, F. R., 319, 325

Arndt, G. D., 309, 324

Arp, V. D., 319, 326, 331, 359 B

Bannister, J. D., 194, 195, 251

Barford, N. C, 339, 360

Barnard, J. D., 312, 313, 325

Bean, C. P., 332, 359

Beanie, J. A., 40, 53 Beck, D. R., 185, 250

Beer, A. C, 308, 324 Behnke, F. A., 312, 325

Bell, J. H., Jr., 196, 251

Benedict, M., 36, 53

Berlincourt, T. G., 333, 334, 359 Berman, R., 35, 37, 53

Biel, R. G., 210, 252

Bijl, A., 52, 53 Birmingham, B. W., 167, 169, 249

Bitter, F., 333, 359

Black, I. A., 187, 189, 191, 250, 251

Blake, C., 322, 326

Block, M. M., 340, 360, 361 Bloembergen, N., 301, 304, 322, 324

Bloor, D., 317, 323, 325, 326 Blumberg, R. H., 312, 313, 325

Boone, W. J., 3, 32

Bowles, L. W., 322, 326

Brady, H. F., 351, 362

Brandt, L. W., 24, 32

Bratt, H. F., 309, 324

Breddy, N. C., 343, 361

Brentari, E. G., 285, 298

Brooks, J. M., 334, 360

Brower, E. M., 349, 362 Brown, R. E., 155, 249 Brown, E., 169, 249 Buchhold, T. A., 334, 360 Buck, D. A., 312, 325 Buffham, B. A., 343, 361

Bura, P., 315, 325

Burley, R. M., 177, 249

Burnier, P., 336, 360

Burns, L. L., 312, 313, 325

Burton, K. R., 358, 362C

Callaghan, E. E., 343, 361

Calvert, A. P., 155, 249

Canty, J. M., 187, 226, 227, 250, 252

Caras, G. J., 352, 362

Caren, R. P., 179, 249 Carney, R. R., 352, 362

Caswell, H. L., 312, 313, 325 Chambers, E. E., 337, 360 Chandrasekhar, B. S., 332, 334, 359

Chelton, D. B., 142, 151, 218, 252

Christiansen, R. M., 185, 250

Churchill, S. W., 186, 250

Clark, J. A., 351, 362

Clarke, J., 316, 320, 325

Clauss, R. C, 322, 326

Clay, J. P., 187, 250

Clayden, W. A., 347, 362 Clement, J. R., 36, 53, 193, 251

Clogston, A. M., 332, 359

Clorfeine, A. S., 315, 325

Cockett, A. H., 279, 298 Cohen, R. W., 311, 324

CofiFey, H. T., 333, 334, 359

Coles, W. D., 329, 333, 336, 359

Collier, R. S., 332, 359
364 AUTHOR INDEX

Collins, S. C, 55, 107, 150

Colyer, D. B., 130, 138, 147, 150, 151 Cook, H. M., 343, 361 Cooke, J. P., 210, 252

Corruccini, R. J., 164, 177, 181, 182, 249, 250, 335, 336, 360

Cosentino, L. S., 315, 325

Cost, J. S., 258, 297

Coston, R. M., 187, 250

Creyghton, J. W., 279, 298

Croft, A. J., 259, 280, 297

Currie, R. B., 74, 75, 76, 136, 141, 151 D

Dana, L. I., 185, 250

Darrel, B., 233, 252 Dawson, J. P., 343, 361 Dayem, A. H.., 311, 324 Dean, J. J., 323, 326 Dean, J. W., 89, 92, 99, 111, 150 Deaton, W. M., 1, 24, 32, 357, 362

Deaver, B. W., 320, 326

de Gennes, P. G., 332, 359

Denenstein, A., 317, 325

Derrick, M., 340, 361

Di Capua, E., 340, 360

Dicke, R. H., 322, 326

Dickson, P. F., 177, 180, 249

Din, F., 279, 298

Doherty, P. R., 189, 191, 250 Domb, C, 35, 53 Domchick, R., 319, 325

Dore, U., 340, 360 Dubs, M. A., 185, 250

Dugdale, J. S., 35, 53 duPont de Nemours, E. I., 15, 32 Dushman, S., 183, 250 Dwyer, R. F., 352, 362E

Edeskuty, F. J., 35, 37, 53 Edwards, H. H., 315, 316, 325 Eisenmenger, W., 311, 324

Elgin, D. R., 187, 250

Engeler, W., 309, 324 Engen, G., 303, 324

Epstein, M.', 227, 252

Ergene, S., 241, 253

Everton, J. G., 347, 362F

Fairbanks, W. M., 337, 338, 340, 360

Farkass, I., 343, 361

Fawcett, H. H., 260, 297 Fenema, P. J., 343, 361

Fetkovich, J. G., 340, 361

Fields, T. H., 340, 361

Firth, M., 142, 151 Fleming, R. B., 132, 151

Foote, W., 303, 324

Fox, D. K., 333, 334, 359

Frainier, R. J., 270, 297

Frazier, D., 15, 32

Fulk, M. M., 177, 185, 189, 249, 250, 251

Fuller, P. D., 215, 252G

Gaffney, J., 36, 53, 193, 251

Gallet, R. M., 37, 53 Gauge, R. A., 315, 325

Gardner, J. B., 142, 151

Gareis, P. J., 15, 32, 343, 361 Garwin, E. L., 343, 361

Gaule, G. K., 315, 325

Gessner, R. L., 130, 138, 150

Getty, R. C, 187, 250

Giaever, I., 310, 324

Gialanella, G. C, 340, 360

Gibbons, H. P., 196, 251, 281, 298

Gifford, W. E., 70, 76

Gilbeau, J. J., 358, 362

Giles, S., 343, 361

Ginzburg, V. L., 329, 359

Glaser, D. A., 339, 360

Glaser, P. E., 186, 187, 189, 191, 250, 251

Glassford, A. P. M., 52, 53

Goethert, B. H., 343, 361 Goldsmith, A., 177, 249

Goldstein, Y., 311, 324 Good, S. H., 196, 251 Goodkind, J. M., 320, 326

Goree, W. S., 320, 326

Gorter, C. J., 332, 359 Gottzmann, C. F., 233, 252 Grimes, C. C, 317, 325

Gubareff, G. G., 177, 180, 249

Guidoni, P., 340, 360H

Haarvis, G. J., 119, 150 Habgood, H. W., 15, 32 Hagenbach, G. F., 343, 361 Haisma, J., 119, 150 Hansen, O. A., 262, 297

Harth, E. M., 340, 360

Hartwig, W. H., 309, 324
AUTHOR INDEX 365

Hawthorne. R. C, 231, 252 Haygood. J. D., 343, 361 Heaton, J. W., 332, 359

Heffner. H.. 303. 305. 324

Henault, P. B.. 343, 361 HenshaU, B. D.. 349. 362 Hersch, P. A.. 248, 253

Higa, W., 322, 326

HiU, F. C, 258, 297

HiU, R. W., 37, 41, 42, 50, 53 HiUe, P. F., 308, 324

Hindle, R. A., 343, 361 Hinson, H. H., 3, 32 Hirschhorn, H. J., 177, 249

Hiza, M. J.. 343, 361

Hnilicka, M. P., 343, 361 Hoare. F. E., 279, 298 Hogan, W. H., 122, 150

Hollingsworth, M., Jr., 185, 250

Hollingsworth, R. T., 343, 361 Hood. C. B., 76, 136, 151, 345, 361 Horn, G. W., 343, 361 Hover, H., 240. 252

Hover, H. K., 261, 297 Huffman, D. R., 15, 32 Hulm, J. K.. 333, 334, 359

Hunter, B. J., 189, 251 Hyman, L., 340, 361

I

Ishaghoff, I., 187, 250J

Jackson, L. C, 279, 298

Jacobs, R. B., 222, 252

Jaklevic, R. C, 320, 326

Jakob, M., 132, 151 Janssen, J. E., 177, 180, 249

Johnson, V. J., 281, 298

Jones, C. H., 334, 360

Jones, G. O., 323, 326

Jones, M. C., 177, 180, 249

Jones, W. B., 37, 53 Josephson, B. D., 310, 311, 324 K

Kamper, R. A., 299, 315, 316, 325, 327, 332,

334, 359, 360

Kantrowitz, A. R., 334, 360

Kays, W. M., 132, 151 Keesom, W. H., 35, 37, 51, 52, 53

Keidel, F. A., 23, 25, 32, 248, 253

Keys, R. D., 196, 251

Kiefer, T. F., 196, 251

Kikuchi, T., 340, 360

Kingston, R. H., 322, 326

Kinney, G. D., 15, 32

Kintz, G. M., 258, 297

Kirk, B. S., 153, 181, 249

Kirkland, C. G., 24, 32

Kleinbaut, J., 237, 243, 252 Kolm, H., 333, 359

Kreith. F., 185, 250

Kremzier, E. J., 187, 250

Kropschot, R. H., 185, 189, 196, 250, 251,

327, 331, 359

Kurti, N., 279, 298 L

Laakso, I., 340, 360 Lamb, H., 192, 251 Lambe, J., 320, 326

Lancaster, J. F., 280, 298 Landau, L. D., 329, 359 Langenberg, D. N., 316, 317, 325

Larkin, B. E., 186, 250

Larsen, P. S., 351, 362

Laquer, H., 336

Laurence, J. C., 329, 333, 336, 359

Laverick, C., 333, 334, 335, 359, 360

Lax, B., 333, 359

Leitner, J., 340, 360

Leonhard, K. E., 187, 250

Levenstein, H., 309, 324

Lindquist, A. B., 312, 325

Little, A. D.. 192, 251

Little, W. A., 319, 326 LobeU, G. M., 334, 360

Lock, J. M., 312, 325 Logan, J. K., 36, 53 London, A. L., 132, 151 London, H., 332, 334, 335, 359

Long, H. M., 239, 252, 255, 352, 362

Loser, J. B., 189, 251

LouiseU, W. H., 306, 324 Lounasmaa, O. V., 35, 37, 41, 42, 50, 52, 53 Loveday, P. E., 255, 266, 297 Low, F. J., 309, 323, 324

Lytle, F. W., 351, 362M

MacKinnon, L., 279, 298 MacRae, A., 309, 324 Manko, H. H., 280, 298
366 AUTHOR INDEX

Mann, D. B., 33, 35, 40, 42, 52, 53, 89, 92, 99, 111, 150, 155, 167, 169, 249, 341, 361 Mark, H., 343, 361 Marsh, N. M., Jr., 185, 250

Marshall, J. W., 233, 252

Martin, D. H., 317, 323, 325, 326

Martin, K. B., 340, 361

Martin, R. J., 311, 324

Martindale, D. L., 358, 362

Martz, D. E., 323, 326 Mate, C. F., 35, 37, 53

Matisoo, J., 314, 325 Mawer, D. A., 323, 326 McAdams, W. H., 132, 151 McAfee, K. B., Jr., 15, 32 McAshan, M. S., 337, 360 McCarty, R. D., 52, 53 McClintock, R. M., 196, 251, 281, 298 McCurley, E. P., 322, 326 McFee, R., 331, 359 McKenzie, J., 340, 361 McLagan, J. N., 215, 252 McLane, J. C, Jr., 343, 361 McMahon, H. O., 70, 76 McWhorter, A. L., 309, 324

Meier, H., 125, 150 Melnyk, N., 15, 32 Mekzer, C, 340, 360 Mercereau, J. E., 320, 326 Meservey, R., 317. 325 Meyer, J. W., 299

Miller, J. C, 315, 325

Miller, R. D., 3, 32

Mills, R. G., 333, 359

Moeller, C. E., 189, 251

Moffatt, J., 340, 361 Mohr, P. H., 15, 32 Monte, L. A. J., 52, 53 Montgomery, D. B., 333, 359 Moore, B. J., 3, 32 Morgan, S. K., 351, 362 Muhlenhaupt, R. C, 99, 108, 116, 150 Mullen, L. O., 343, 361

Mullins, J. C, 181, 249 Mulhns, P. v., 1 Mumford, W. W., 306, 324 Munnerlyn, R. D., 3, 32 Murray, B. C, 323, 326N

Nahman, N. S., 319, 326

Navikas, C. R., 191, 251 Neary, R. M., 204, 252 Newhouse, V. L., 315, 316, 325

Niendorf, L. R., 352, 362

Norrell, G. P., 3, 32

'Norton, M. T., 194, 251 Nubel, G. C., 153, 240, 252, 261, 297 O

Ohori, Y., 332, 359 Onnes, H. K., 78, 150, 155, 249 Outmann, V., 343, 361P

Paivanas, J. A., 161, 249, 266, 297

Parker, W. H., 316, 317, 325 Parmley, R. T., 187, 250

Partridge, R. B., 322, 326

Payne, E. G., 336, 360 Pehek, J., 309, 324

Perry, C. H., 323, 326

Pewitt, E. G., 340, 361

Pickard, G. L., 87, 150

Pierce, J. M., 320, 326

Pieters, H. A. J., 279, 298

Plattner, C. M., 350, 362

Post, A., 70

Post, R. F., 335, 336, 360

Prothero, W. M., Jr., 320, 326

Purcell, J. R., 334, 336, 360

Putley, E. H., 309, 323, 324, 326

Pylant, H. S., 13, 32 R

Randolph, W. O., 351, 353, 357, 362

Rediker, R. H., 309, 324

Reeves, J. R., Jr., 345, 362

Reid, F. J., 308, 324

Reynolds, M. M., 177, 249

Reynolds, R. B., 347, 362

Reynolds, W. T., 333, 334, 359

Richards, P. L., 317, 325

Richards, R. J., 199, 251 Richman, J., 345, 361

Riede, P. Mf, 185, 250; Riemersma, H., 334, 3$9

Ries, R. P., 308, 316, 324

Rietdijk, J. A., 118, 150 Robbins, R. F., 167, 169, 249

Robinson, R. J., 15, 32

Rogers, A. D., 122, 150 Roozendal, K., 119, 150

Rose-Innes, A. C., 279, 282, 298, 332, 359

Ross, R. L., 315, 325

Rubin, L. C., 36, 53
AUTHOR INDEX 367

S

Saint-James, D., 332, 359

Satterthwaite, C. B., 308, 316, 324

SchaUa, C. A., 232, 252 Schenk, H. L., 334, 360

Schiffhauser, J. A., 268, 297 Schmidt, A. F., 327, 351, 353, 362 Schock, K., 233, 252

Schrader, E. R., 329, 359

Schrodt, J. E., 189, 251 Schwartzberg, F. R., 196, 251 Schwettman, H. A., 337, 360

Schwidtal, K., 315, 325

Scott, L. E., 87, 150, 153, 155, 160, 167, 169, 203, 204, 234, 237, 243, 249, 252, 266, 268, 297

Scott, R. B., 52, 53, 160, 177, 183, 185, 193, 215, 220, 249, 279, 298

Seeber, R. R., 312, 325

Seitz, F., 339, 360

Shal'nikov, A. I., 278, 298

Shapiro, S., 317, 325 Shaw, D. F., 339, 360

Short, A. J., 278, 298 Sherman, R. H., 35, 37, 53 Shrewsbury, R. D., 3, 32

Shutt, R. P., 339, 360

Sibiha, J. T., 322, 326

Siegel, K., 319, 325 Siegman, A. E., 301, 324

Silver, A. H., 315, 316, 320, 325, 326 Simon, F., 87, 150 Simons, J. C, Jr., 343, 361

Singer, J. R., 301, 324

Sinclair, T. F., 15, 32

Sixsmith, H., 137, 151 Smith, H. J., 343, 361

Smith, J. L., Jr., 52, 53

Smith, P. F., 334, 360

Smith, P. K., 13, 32 Smith, R. A., 308, 309, 324

Smith, R. v., 285, 298

Smith, T. I., 337, 360 Snow, W. A., 74, 75, 76, 85, 136, 141, 150, 151 Sommers, R. D., 343, 361

Sonntag, R. E., 79, 80, 81, 150 Span, R. E., 333, 334, 359

Steiner, W. A., 15, 32

Stekly, 2. J. J., 334, 360

Stelzried, C, 322, 326

Stelzried, C. T., 303, 324

Stern, S. A., 15, 32

Stewart, R. B., 35, 40, 52, 53

Stoecker, L. R., 185, 250

Stokes, R. A., 322, 326

Stoner, J. T., 351, 362

Streeter, M. H., 107, 150

Strobridge, T. R., 36, 52, 53, 77, 99, 108, 116,

142, 150, 151

Strong, J., 280, 298

Stuart, R. W., 107, 122, 150 T

'Tabor, W. J., 322, 326

Taylor, B. N., 316, 325

Taylor, C. E., 87, 150, 335, 336, 360 Thompson, M. B., 189, 251 Thompson, P. A., 329, 359 Timmerhaus, K. D., 343, 361 Tinkham, M., 331, 359

Torborg, R. H., 177, 180, 249 Trembath, C, 303, 324

Trepp, Ch., 66, 241, 253

Trucks, H. R., 357, 362 Turneaure, J. P., 337, 360 V

Vahldieck, N. P., 15, 32 Vaniman, J. L., 351, 353, 362 Van Meerbeke, R. C, 153, 192, 193, 200, 251,

273, 287, 297 Van Wylen, G. J., 79, 80, 81, 150 Von Helms, H. C, 231, 252 W

Wait, D. F., 303, 324 Wang, D. I-J., 185, 250 Wapato, P. C, 129, 138, 150 Ward, D. E., 142, 151 Waterman, T. E., 177, 249 Webb, G. B., 36, 53 Weber, J., 301, 304, 324

Weller, S., 15, 32 Westbrock, A. J., 70, 76, 343, 361 Westphal, J. A., 323, 326 Wexler, A., 158, 160, 193, 226, 249, 261, 297

White, G. K., 279, 283, 298 Wiebe, E., 322, 326

Wilkinson, D. T., 322, 326

Wilson, P. B., 337, 360

Wilson, W. A., 137, 151 Wine, C. M., 315, 325

Winters, A. R., 74, 75, 76, 85, 136, 141,

150, 151
368 AUTHOR INDEX

Witter, C. E., 141, 151 Wood, W., 260, 297

Zabetakis, M. G., 256, 297

Zar, J. L., 334, 360 Zelmanov, J., 52, 53 Zenner, G. H., 262, 297

Ziegler, W. T., 181, 249 Zimmerman, J. E., 315, 316, 325

Zurcher, A., 125, 150
Subject Index

A

Absolute electromagnetic

measurements, 316 Accommodation coefficient, 182 Adiabatic expansion, 86

Adsorption, oil removal, 123

Analysis methods, 3, 22 Automotive semitrailers, 28, 239 B

Boiling point, normal, 35, 36 Brayton cycle, 99 Bubble chambers, 339C

Camot efficiency, 82 Cascade cycle, 99 Claude cycle, 106 Compressed Gas Association, 198 Compressor. 122, 246

gas lubricated, 127

isothermal efficiency, 92, 130

oil free, 119, 123

reciprocating, 123

turbine, 127 Conservation

extraction plants, 11 federal program, 10

future supply, 11 storage field, 11 Container support systems, 160 Continuous film memory, 313 Cooldown, 81. 90

Cost of gaseous helium, 31 Costs of transportation, 29, 30

Couplings, transfer line, 215

Cryoelectronics, 229 Cryopumping

high-altitude air sampler, 350

space chambers, 343 wind tunnels, 347

D

Density, pressure-density diagram, 38

Depressurization losses, 233, 235 Design of equipment

cryostat safety, 279, 282

liquefier safety, 258

materials and methods, 262

relief devices, 207, 266

effect of altitude, 273

venting of cryostat, 284

transport and storage safety, 261 E

Emittances, radiant thermal, 175 Enthalpy, 40, 41, 43, 46, 47, 50

Entropy, 36, 40, 41, 42, 43, 44, 45, 46 Evacuation of insulations, 171, 191 Evaporation rate, 84 Expansion engine, 65, 72, 132

adiabatic efficiency, 104

gas lubricated, 137, 140

reciprocating, 136

turbine, 136, 140 Expansion ejector, 118 Expansion valves, 107, 110, 142 F

Flash losses, 229 G

Gas conduction heat transfer, 180

G-loads on containers, 200 Gifford-McMahon cycle, 119, 138 H

Heat exchanger, 130

cooling curve, 93, 109

finned, 98 369
370 SUBJECT INDEX

Heat exchanger — Continued

Hampson, 132

invalid temperature profile, 93, 96, 110

laminated, 119

plate fin, 132 pressure drop, 114 Heat exchangers, 63, 64, 72 Heat leak, determination of, 85, 214 Heat of vaporization, 41, 45, 84 Helium production Federal Government, 6, 12, 16

foreign, 7, 15

history, 6

other processes, 15

private industry processes, 7, 13, 14 Helium resources

geographical locations, 4, 5 occurrence, 3

quantity, 4

survey, 3 Helium, thermophysical properties, 33 Helium uses and markets, 8 Helium vaporization, 163

I Impact tests of materials, 198

Injection cooling, 352

Insulation, multilayer, 170

Insulation, porous, 172

Insulation, simple high vacuum, 170,

171, 173

Inversion curve, 89

Isothermal compression, 92

J Josephson Effect parametric amplifier, 315

radiation detectors, 308, 310, 323 Joule-Thomson effect, 56, 89

process, 88 L

Lift pumping, 356

Linear accelerators, 337

Liquefaction cycles, 63, 66, 69

Liquefaction facilities Air Products and Chemicals, 74 500 Incorporated, 69 CVI Corporation, 76

Otis, Kansas, 66 Union Carbide, Linde Division, 70

Liquid containers,

cost, 157

load rates, 159 Liquid production, 155, 241 Liquid properties, 156 Liquid semi-trailer, 166, 167, 169, 241, 244 M

Magnetoresistive parametric amplifier, 308 Magnets, cryogenic, 335 Magnets, superconducting, 329 Maser, 301 Measurement of container loss rate, 163 N

Noise standard, 304

Noise temperature, 303, 304 O

Oil detection, 124 P

Parametric amplifier (paramp), 306, 308,

14, 315, 322 Photon counter, 304

Pipelines, 29

Pressure drop, 214, 230

Purification

.-CO2 removal, 17 hydrogen removal, 19 water vapor, 17 Purity of helium, 21, 22, 248 R

Radiation heat transfer, 160, 173 Railway tank cars, 28

Refrigeration, 77, 138 Brayton cycle, 99 Cascade cycle, 99 Claude cycle, 106, 139, 141 continuous vs. bulk Uquid, 85

efficiency, 142 evaporating liquid hehum, 84 Gifford-McMahon cycle, 119, 138 Joule-Thomson process, 56, 88

relationship to liquefaction, 83 Simon process, 61, 86

Stirhng cycle, 59, 116 volume, 147

weight, 147
SUBJECT INDEX

Regulations governing Federal sales, 31

Relief valves, 208, 212

Ryotron, 315 S

Safety, 255 codes and regulations, 261

cryostat design, 215

general precaution, 257

liquefier design, 258

liquefier operation, 260 neck plug hazards, 265

relief devices, 221, 266 Semiconductor parametric amplifier, 307 Shipment, aircraft, 207, 210 Simon process, 61, 86 Small cylinders, 28 Space systems

coolant, 358

pressurization, 357

Specific heat, 35, 49, 50

Stirhng cycle, 59, 116 Storage of gaseous helium above ground, 26, 28 underground, 4 Superconducting

amphfier, 315

flux-quantum counter, 320

gravimeters, 320

inductors, 319 magnetometers, 319

null detectors, 308, 316

parametric amplifier, 16, 308

passive circuit elements, 318

Superconducting — Continued

pulse delay lines, 319

radiation detectors, 310

cryosar, 309

tunnel diode, 310T

Thermal conductivity, 51 Thermal oscillation, 193, 277 Thermal stratification, 168, 169 Thermodynamics, 79

control volume, 79

first law, 79 second law, 79 system, 79 Throtthng losses, 228

Transfer efficiency, 227, 237

Transfer of fiquid, 199, 213, 257, 274

precooling and purge, 274

transfer to cryostat, 284

Transistor, field effect, 300

Transportation of hquid, 157, 239, 271 Transportation of gas, 26 Tunnel diode, 300

Tunneling cryotron, 314

Tunneling cryotron amplifier, 315 V

Valves, 215 Vapor shielding, 161 Vapor superheat, 163

Viscosity, 51

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