NISTIR 7352

NIST Liquid Hydrogen

 Cold Source

P. Kopetka

 R. E. Williams

J. M. Rowe

NISTIR 7352

NIST Liquid Hydrogen

 Cold Source

P. Kopetka

 R. E. Williams

J. M. Rowe

Materials Science and Engineering Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899-8560

September 2006

U.S. Department of Commerce

Carlos M. Gutierrez, Secretary

Technology Administration

Robert Cresanti, Under Secretary of Commerce for Technology

National Institute of Standards and Technology

 William Jeffrey, Director

Certain commercial entities, equipment, or materials may be identified in 

this document in order to describe an experimental procedure or concept 

adequately. Such identification is not intended to imply recommendation 

or endorsement by the National Institute of Standards and Technology 

(NIST), nor is it intended to imply that the entities, materials, or 

equipment are necessarily the best available for the purpose.

Disclaimer
ABSTRACT

Nearly two-thirds of the experiments performed at the NIST Center for Neutron Research (NCNR) utilize 

cold neutrons with wavelengths greater than 4 Angstroms. This report documents the development of the 

liquid hydrogen cold neutron source in the NIST research reactor. The source was designed to optimize 

the flux of cold neutrons transported to the scattering instruments in the guide hall. It was also designed 

to be passively safe, and operate simply and reliably. All hydrogen system components are surrounded with 

monitored helium containments to ensure that there are at least two barriers between the hydrogen and the 

atmosphere. Monte Carlo simulations were used to calculate the cold source performance and estimate the 

nuclear heat load at full reactor power. Thermal-hydraulic tests in a full-scale mockup at NIST Boulder 

confirmed that a naturally circulating thermosiphon driven by the 2 meter height of the condenser could easily 

supply the moderator vessel with liquid hydrogen while removing over 2000 watts. The cryostat assembly was 

designed to withstand any high pressure generated in a credible accident. It was fabricated to rigorous quality 

assurance standards, resulting in over 10 years of leak-free operation. 
Contents

1.0 Introduction

2.0 Design Basis

2.1 Requirements of 10 CFR 50.59

2.2 Extraction of Beams

2.3 Simple, Reliable Operation

2.4 Maximize Intensity of Neutrons Below 5 meV

3.0 Cold Neutron Performance and Nuclear Heat Load Calculations

3.1 NBSR Modeling with MCNP

3.2 Neutron Performance Calculations

3.3 Nuclear Heat Load Calculations

3.4 Evolution of the Liquid Hydrogen Sources

3.5 Source Performance

4.0 Thermal Hydraulic Design

4.1 Design Criteria

4.2 Initial Design

4.3 Thermal Hydraulic Tests at NIST Boulder

4.4 Other Components (Condenser, Ballast Tank, Hydride Storage)

4.5 Helium Refrigerator

5.0 Mechanical Design

5.1 Hydrogen Cryostat Assembly

5.2 Outer Plug Assembly

5.3 Cold Hydrogen Transfer Line Assembly

5.4 Hydrogen Condenser Assembly

5.5 Hydrogen Gas Management Assembly
5.6 Insulating Vacuum Pump Assembly

6.0 Instrumentation and Control

6.1 Programmable Logic Controller (PLC)

6.2 Control Program

6.3 User Interface

6.4 Rundowns and Trouble Alarms

7.0 Startup and Operation of the Liquid Hydrogen Cold Sources

7.1 Authorization

7.2 Startup and Operation of Unit 1

7.3 The Advanced Liquid Hydrogen Cold Source, Unit 2

7.4 Heat Load Measurements

8.0 Accident Analysis

8.1 Reactor Power Change

8.2 Refrigerator Failure

8.3 Sudden Loss of Electrical Power

8.4 Loss of Cooling Water Flow to Cryostat and Plug Assembly

8.5 Loss of Cryogenic Insulating Vacuum

8.6 Sudden Release of Liquid Hydrogen to Vacuum Space

8.7 Slow leak of Air into Vacuum Vessel

8.8 Hydrogen Release to Confinement Building

8.9 Stoichiometric Mixture of Hydrogen and Air in Moderator Chamber

8.10 Maximum Hypothetical Accident

9.0 Acknowledgements

9.1 References

9.2 Appendix
9

1.0 Introduction

The neutrons that are used for neutron scattering and 

other research applications at research reactors are 

those that have been slowed down from the range of 

MeV, at which they are created, to tens of meV by 

scattering in the moderator and reflector. In most 

reactors, the neutron energy spectrum is essentially 

Maxwellian, with a characteristic temperature 

somewhat higher than the actual temperature of the 

moderator (usually approximately room temperature). 

For such a spectrum, the mean neutron wavelength, 

energy and speed are 1.8 Å, 25 meV and 2200 ms-1, 

values well suited to study the properties of condensed 

matter at the atomic scale, where interatomic spacings 

are approximately 1 Å, and characteristic vibrational 

energies are of order meV.

Over the past two-three decades, structures with 

characteristic lengths of 100 Å and correspondingly 

smaller vibrational energies, have become increasingly 

important for both science and technology. For 

example, polymers are being developed with 

totally new properties as substitutes for other more 

traditional metal materials for weight savings, nanostructures are being developed with novel properties 

for processing drug delivery plus other applications 

and biotechnology applications are increasing in 

importance. All of these structures have characteristic 

dimensions that are best matched by neutrons of 

longer wavelength, which also allow better resolution 

in energy to match the slower relaxation and vibration 

of the more massive objects. However, the fraction of 

neutrons with energies less than 5 meV in a normal 

moderator spectrum is less than 2 % of the total, 

which makes their use impractical, except in special 

cases. In order to address this problem, several groups 

around the world1 have developed systems that place 

cryogenically cooled moderators called “cold sources” 

in the reflector of neutron sources. These cold sources 

shift the spectrum down in energy to a temperature 

somewhat above the moderator temperature, providing 

large gains in intensity for low energy neutrons.

The NIST research reactor, NBSR (National Bureau of 

Standards Reactor), was designed from the beginning 

with a provision to add a cryogenically cooled 

moderator system, or cold neutron source. The initial 

design, as discussed in NBSR-92 and NBSR-133, was 

a block of D2O ice, cooled to approximately 25 K 

by circulating helium gas. A water cooled lead and 

bismuth gamma ray shield was included to reduce 

heating of the ice. The source was installed in the 

reactor in 1987. Its successful operation was a key to 

the development of the NIST Cold Neutron Research 

Facility, a large new experimental area constructed to 

take advantage of the longer wavelength neutrons 

produced by this source. The success of this facility, 

and the high national demand for cold neutrons, 

led to a design effort to develop a more intense cold 

neutron source, based on liquid hydrogen. The first 

version of this source was installed in the NBSR in 

1994, and resulted in a gain in cold neutron intensity 

of a factor of seven. Further refinements of the design 

resulted in an additional gain of a factor of two for the 

unit installed in 2002.

This report documents the design and analysis of the 

latest version of the NIST liquid hydrogen source.
10

2.0 Design Basis

The design basis for the liquid hydrogen cold source 

is driven by the following requirements, listed in 

priority order:

The finished cold source must satisfy the safety requirements of 

10 CFR 50.59.

The source must allow extraction of beams from all existing cold 

source beam ports (CT-E and CT-W), as well as from the insertion 

port, with full guide illumination to 10 Å.

Operation should be simple and reliable, with minimal impact on 

reactor operation.

The intensity of neutrons with energies less than 5 meV should be 

maximized.

Each of these requirements has several implications 

for the design of the liquid hydrogen cold source, 

which will be addressed in turn.

2.1 Requirements of 10 CFR 50.59

At the time of cold source design and construction, 

the requirements of 10 CFR 50.59 were:

(a)(1) The holder of a license authorizing operation of a 

production facility may (i) make changes in the facility as 

described in the safety analysis report, (ii) make changes in the 

procedures as described in the safety analysis report, and (iii) 

conduct tests or experiments not described in the safety analysis 

report, without prior Commission approval, unless the proposed 

change, test or experiment involves a change in the technical 

specifications incorporated in the license or an unreviewed safety 

question. 

(2) A proposed change, test, or experiment shall be deemed to 

involve an unreviewed safety question (i) if the probability of 

occurrence or the consequences of an accident or malfunction 

of equipment important to safety previously evaluated in the 

safety analysis report may be increased; or (ii) if a possibility for 

an accident or malfunction of a different type than any evaluated 

previously in the safety analysis report may be created; or (iii) 

if the margin of safety as defined in the basis for any technical 

specification is reduced.

In order to comply with this requirement, the 

cold source design emphasizes simplicity, passive 

safety features, complete blanketing of all hydrogen 

containing components with the inert gas helium, and 

a vessel designed to withstand a hypothetical ignition 

of a hydrogen-air mixture without damage to the cold 

source thimble. Based upon experience in Europe, 

these requirements have led to a thermosiphon design, 

with gravity driven flow of liquid hydrogen from a 

condenser into the moderator chamber, removal of 

•

•

•

•

•

•

the heat (generated by the reactor) by boiling of the 

liquid and return of the vapor to the condenser. The 

hydrogen-containing components of this system have 

no moving parts, and the entire hydrogen inventory 

is contained in a closed system with no automatic or 

remotely operated valves or rupture discs. Hydrogen 

vapor returning from the moderator chamber is 

recondensed in a heat exchanger cooled by helium gas 

at 15 K, and then flows back down into the vessel 

under gravity. The hydrogen supply and return lines 

are sized to provide for two phase (liquid and vapor) 

flow in the return from the moderator chamber. This 

provides a very stable operating condition under 

changing conditions of heating, as shown for a source 

designed for the High Flux Reactor (HFR) at the 

Institut Laue Langevin4 (ILL), and verified for the 

NIST geometry in a series of mockup tests5 carried 

out at NIST-Boulder. A schematic view of this system 

is shown in Figure 2.1. 

Note that all hydrogen-containing components, or 

components that could conceivably contain hydrogen, 

are surrounded by helium, including the ballast tank 

"ALLAST4ANK

,IQUID(YDROGEN4HERMOSIPHON

,ITERS	

6ACUUM

0UMP

-ODERATOR

#HAMBER (ELIUM#ONTAINMENT

)NSULATING6ACUUM

(YDROGEN

#ONDENSER

(E/UT (E)N

$//UT

$/)N

(YDROGEN6APOR

(YDROGEN,IQUID

(EAVY7ATER

Figure 2.1 Schematic of NIST Advanced Cold Neutron 

Source, Unit 2.
11

(sized to hold the hydrogen gas at less than 0.5 MPa 

when the system is warm), and the vacuum pumps. 

There are at least two barriers to mixing of hydrogen 

with air, which prevents formation of a flammable 

mixture. In addition, all hydrogen containing 

components, including the lines from the moderator 

chamber to the condenser and ballast tank, are 

protected from damage by being surrounded by steel 

barriers or heavy, non-movable shields. This is the 

basis of the safety case – mixing of hydrogen and air 

is prevented to avoid any possibility of damage to the 

reactor. In addition, one of the vessels surrounding 

the moderator chamber, the helium vessel, is sized 

to withstand a hypothetical reaction of a room 

temperature mixture of air and hydrogen at 0.1 MPa, 

without any damage to the reactor vessel thimble that 

contains the source.

2.2 Extraction of Beams

The arrangement of the cold source thimble and 

beams in the original reactor design is shown in 

Figure 2.2. As can be seen from the figure, two 

horizontal beam ports (CT-E and CT-W) were 

included, separated by an angle of 33°, and crossing 

before the end of the thimble. This design was chosen 

to accommodate a large D2O ice cold source, which 

would have a viewed surface at the point of the beam 

crossing. For a hydrogen source, the viewed surface 

should be much closer to the tip of the thimble, as 

a result of the much higher scattering cross-section 

for hydrogen than deuterium. The net result is that 

the viewed surface must be large, in both height 

(determined by the beam port diameter of 170 mm 

and the guide height) and width (determined by the 

crossing point of the beam port lines-of-sight and the 

guide width). In calculating the required sizes, guide 

coating with 58Ni is assumed, which gives a critical 

angle of 0.02 radians at 10 Å. This adds nearly 80 

mm to the required area in both dimensions, since the 

guides begin approximately 2 m from the end of the 

thimble, and the additional area is required to fully 

illuminate the guides. The required size of the viewed 

surface is approximately 300 mm × 300 mm, a size 

that imposes severe constraints on design. In order 

to make a moderator chamber with such a large, flat 

viewed surface, able to withstand the pressures involved 

(up to 0.5 MPa internal pressure), the material would 

have to be tens of mm thick, which would both 

reduce the intensity due to absorption and scattering 

and increase the heat load beyond reasonable values. 

One solution examined was to use an elliptical shell of 

at least 2:1 axial ratio, while a second was to connect 

a number of vertical tubes arranged to provide an 

almost uniform thickness moderator of appropriate 

dimensions. While either of these solutions would 

have worked, they both have serious defects. For the 

former, the thickness of hydrogen would have been far 

too large for optimum intensity, while for the latter, 

fabrication would have been quite complex, involving 

many welds, and the fluid flow in the thermosiphon 

would have been difficult do in a way that ensured 

safe, effective operation.

2.3 Simple, Reliable Operation

In our research on the causes of incidents and failures, 

it quickly became clear that human error and failure of 

non-passive devices were the most prevalent initiators. 

For this reason, the design studies focused on passive, 

fail-safe operation, with an absolute minimum of 

active components. It was decided that the reactor 

would not operate unless the cold source was 

operational, eliminating any backup cooling systems. 

This was done primarily to simplify the design of the 

in-pile components, as any backup cooling would 

have to involve gas cooling of the moderator chamber, 

implying a high velocity flow of helium. Experience 

with the D2O ice source, which used high velocity 

gas cooling, and at Brookhaven National Laboratory, 

where a directly cooled hydrogen source operated for 

several years, showed that this was a source of problems 

Figure 2.2. Original layout of the cold neutron port.
12

for several reasons. The most critical was that the gas 

would be flowing at high velocity through a region 

of high neutron flux, so that any erosion could lead 

to transport of radioactive material out of the in-pile 

components, presenting both a maintenance problem 

(due to internal contamination of the gas circulation 

system) and a possible health physics problem (if 

a seal material such as indium were to break off, 

become highly activated, and be transported out into 

piping in occupied areas). Another was that the gas 

circulation system would have to be jointed, to allow 

for maintenance, providing opportunities for leaks of 

contaminated gas.

Experience here and at other facilities was an important 

factor in the design if the NIST liquid hydrogen (LH2) 

cold source. The only pipes that go into the in-pile 

area in the chosen design are the hydrogen supply and 

return, and D2O cooling for the heavy walled vessels 

in the cryostat. There are no connections in any of 

these pipes – all joints are welded, and the welds were 

radiographed and helium leak tested at the time of 

installation. The hydrogen thermosiphon is operated 

under saturated conditions, so that the only relevant 

parameter is the vapor pressure, which determines 

the temperature, and the instrumentation is outside 

the high radiation fields. Finally, the thermosiphon 

is operated in a regime with two phase flow in the 

return line, providing stable operation over a wide 

range of heat inputs.

The helium gas used to condense the hydrogen 

is supplied by a closed cycle refrigerator with a 

compressor, coldbox, and expander turbine. The model 

chosen provides up to 3.5 kW of cooling at 14 K, 

which is far more than required, but which allows for 

consideration of a deuterium source at a future date. 

The refrigerator is controlled by a Programmable Logic 

Controller (PLC), with automatic startup, shutdown, 

and warmup functions, and is operated at a fixed 

cooling power. The cooling loop is controlled by the 

same PLC, which controls a valve regulating the flow 

of cold helium to the condenser (the remainder of 

the flow bypasses the condenser) so as to maintain a 

constant hydrogen pressure. This mode of operation 

maximizes refrigerator reliability, and provides rapid 

responses to changes in heat load (such as reactor 

power changes).

2.4 Maximize Intensity of Neutrons Below 5 

meV

This consideration is of course the primary reason 

for installation of a new cold neutron source, but 

the solution is seriously constrained by the preceding 

requirements. In order to optimize this constrained 

design, it was necessary to iterate repeatedly, first 

calculating the neutronics for an idealized source 

geometry, then doing a rough mechanical design 

incorporating constraints, then repeating the neutronics 

calculations. Several concepts were considered before 

settling on the actual design used, which is shown in 

Figure 2.3. The fundamental concept is creation of 

a shell of hydrogen inside a mechanically efficient 

shape – first a sphere, then in the second version of the 

source, an ellipsoid of revolution. The large neutron 

scattering cross-section of hydrogen, along with the 

finite absorption cross-section, leads to the existence 

of an optimum thickness of moderator material. 

There must be some material for the neutrons to 

scatter and thus lose energy, so that the yield of cold 

neutrons increases as the thickness increases from zero. 

However, as the thickness increases further, the yield 

of cold neutrons begins to decrease, as some of the 

neutrons slowed down in the first layers will not be 

able to reach the viewed surface because of scattering 

and absorption. The optimum thickness depends on 

the relevant cross-sections, which introduces another 

complication.

Hydrogen molecules exhibit well-defined states of 

Figure 2.3. Layout of the Advanced Cold Neutron Source and 

guides.

NG-0

NG-1

NG-2

NG-3

NG-4

NG-5

NG-6

NG-7

MODERATOR CHAMBER

INSULATING VACUUM

FUEL ELEMENT

D2O REFLECTOR

D2O COOLING JACKET

HELIUM CONTAINMENT

REACTOR VESSEL
13

angular momentum characterized by rotational 

quantum numbers J=0,1,2..., and at low temperatures, 

in the stable equilibrium state, all of the molecules will 

be in the J=0 state, which has the lowest energy. As a 

result of the spin dependence of scattering, molecules 

with even angular momentum (para-hydrogen) have 

a small scattering cross-section, unless the neutron 

energy is large enough to allow excitation to a state 

of higher angular momentum, while molecules with 

odd J (ortho-hydrogen) have a large scattering crosssection for all neutron energies. As stated above, the 

equilibrium state would be almost all para-hydrogen 

at 20 K, and thus would have a low scattering crosssection for neutron energies less than 15 meV. This 

would imply that the optimum thickness could 

be quite large. However, there is a great deal of 

uncertainty about the actual state, with contradictory 

evidence6.7 and most designers have assumed that 

the actual state is an equal mixture of ortho and para 

hydrogen8. Our studies indicate that the optimum 

thickness is approximately 30 mm, so long as the 

ortho content is greater than 40 % to 50 %, and the 

close agreement of our measured and calculated cold 

neutron intensities and energy dependence seems to 

confirm that the concentration is closer to normal (75 

% ortho) than to para for our source. In order to 

extract the maximum number of neutrons from the 

shell, an entrance hole was included with minimal 

hydrogen in the path, as shown in Figure 2.3. The 

details of the neutronic design are given in Section 3.
14

3.0 Cold Neutron Performance and 

Nuclear Heat Load Calculations

The design of the NIST liquid hydrogen cold 

sources was guided by the results of Monte Carlo 

simulations of neutron and photon transport through 

detailed models of the reactor core and cold source 

thimble. Computer codes and platforms have evolved 

tremendously since the first Cold Neutron Source 

(CNS) calculations in the late 1980’s. No effort will 

be made to detail chronologically all of the steps and 

miss-steps in the process. The first three parts of this 

section describe the currently-used methods for cold 

source development. The concluding sections outline 

the evolution of the source, and the parameters that 

were optimized in its development, with brief notes 

about early results using what are now outdated 

methods.

3.1 NBSR Modeling with MCNP

A three-dimensional, Monte Carlo Neutron Photon 

(MCNP9) model of the cold source thimble was first 

used at the NCNR for the development of a liquid 

hydrogen cold neutron source in 1989. A full core 

model was completed in 1993 accurately predicting 

the performance and nuclear heat load of the first 

LH2 source, Unit 1. MCNP is a Monte Carlo 

neutron and photon transport code developed at Los 

Alamos National Laboratory (LANL) that is used 

for a wide variety of problems including criticality 

simulations. It features generalized surfaces and cells, 

so that complex geometries can be defined, along 

with continuous energy cross-section data. Hundreds 

of cross-section files with gamma-ray production data 

have been formatted for use with the code, including 

thermal neutron scattering kernels for all common 

reactor moderators, and four cold moderators. The 

most recent version, MCNP 5, includes seven kernels 

for ortho- and para-LH2 in 1 K intervals from 19 K to 

25 K (more information about ortho/para-hydrogen 

will be presented in later sections).

The core model has been continually modified and 

expanded, and it was used extensively in the reactor 

physics analysis of the NBSR submitted to the 

Nuclear Regulatory Commission (NRC) in April 

2004 in support of the relicensing effort. All of the 

major features of the NBSR that affect the reactivity 

of the core are represented in great detail in this model 

including:

A hexagonal array of the 30 fuel elements, 6 vertical thimbles, 

and the regulating control rod.

All 1020 fuel plates, their Al cladding and D2O-filled coolant 

channels, positioned in hexahedral repeating structures for the 

upper and lower sections of the core.

Fifteen fuel material specifications representing each step in the 

fuel management pattern for the 7 and 8 cycle fuel elements.

The four shim arms, which can be positioned at any angle 

between their scram and fully withdrawn positions.

Nine radial beam tubes, two tangential beam tubes, the vertical 

beam tube, and the four in-core pneumatic ‘rabbit’ tubes.

The large cryogenic beam port and the liquid hydrogen cold 

source(s).

The reactor vessel, filled with D2O, representing the moderator 

between the fuel elements, and the core reflector regions. 

Layers of lead and iron outside of the vessel, comprising the 

thermal shield, and a layer of concrete, for part of the biological 

shield.

A portion of the D2O tank, providing neutronic coupling with the 

graphite in the thermal column.

Many of the above features are shown in Figure 3.1. 

The lines of the hexagonal array are for computational 

purposes, locating the 30 fuel elements. Concentric 

areas surrounding the core are the D2O reflector, the 

Al vessel, 5 cm of lead and 20 cm of iron comprising 

the thermal shield, and the first 20 cm of the concrete 

biological shield. Four cells between the inner 

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•

•

•

•

•

•

•

Figure 3.1. The MCNP model (in cm) of the NBSR core at its 

mid-plane, in the unfueled region.

-150 -100 -50 0 50 100 150 -150 -100 -50 0 50 100 150
15

most rows of fuel are excluded from the array to 

accommodate the cells of the Cd shim arms. Starting 

at 5 o’clock, and moving counterclockwise, the radial 

beam tubes are BT-1 through BT-9. Between BT5 and BT-6 is the large cryogenic beam port, CT, 

containing the Unit 2 liquid hydrogen CNS. The axis 

of CT is 11.25o west of the y axis of the model, which 

is the north direction. Part of a D2O tank leading to a 

graphite thermal column (not shown) substitutes for 

the thermal shield on the south side.

The MCNP model was provided to the Energy 

Sciences and Technology Department at BNL as a 

basis for their reactor physics and safety analysis study, 

Appendix A of NBSR-14, the updated SAR10. Section 

3 of that appendix contains a thorough description of 

the MCNP model, including the fuel geometry and 

fuel loading scheme, control elements, experimental 

facilities, etc. The 235U inventory in each fuel element 

was determined iteratively with three computer codes, 

MCNP, ORIGEN211, and MONTEBURNS12, with 

MONTEBURNS managing the interactions between 

MCNP and ORIGEN2. The model has been 

carefully benchmarked, and was used to calculate the 

power distribution, reactivity coefficients, integral 

and differential reactivity worths of the shim safety 

arms and the regulating control rod, neutron fluxes in 

the experimental facilities, etc.

3.2 Neutron Performance Calculations

The figure of merit for judging the performance of a 

cold neutron source is the “brightness” of the source 

in the direction of the beams or guides to various 

instruments. Brightness, or d2φ/dλdΩ, has units 

of neutrons/cm2-s-Å-ster, when written in terms of 

wavelength, or d2φ/dEdΩ with units of n/cm2-s-meVster, when written in terms of energy. Specifically, it 

is the brightness of cold neutrons, E < 5 meV or λ > 

4 Å, in the direction of the guides, with angles less 

than the critical angle for reflection from the 58Nicoated surfaces within the guides. The critical angle 

is given by θC = 0.002 λ(Å) rad for 58Ni. An MCNP 

current tally is used to calculate the brightness in a 

two-step process. In the first step, a lengthy criticality 

calculation generates a surface source file for neutrons 

entering the CT region. Much shorter calculations 

starting with the source file can exploit powerful 

variance reduction techniques to optimize the cold 

source geometry (see the MCNP input file, “unit2”, 

in the appendix).

A surface source file containing the location, direction, 

energy and weight of the starting neutrons to be used 

for the MCNP calculation was generated in a KCODE 

calculation of the entire NBSR core including the 

region surrounding the CNS (Figure 3.1). The CT 

axis, y ́, is 11.25° west of the y-axis. A coordinate 

transformation to y ́ is needed for all the surfaces in 

the CNS region. Typically, the surface source is on the 

MCNP surfaces PY ́ = 55.5 cm (a plane perpendicular 

to the y ́ axis, 55.5 cm from the core center) and CY ́ 

= 30.0 cm (a cylinder along the y ́ axis with a 30 cm 

radius). Data for every particle crossing PY ́ in the 

positive, +y, direction, and every particle passing into 

the CY ́ cylinder are recorded in the surface source 

file. Such files may be very large, 500 MB, to obtain 

good statistical results. A track for a given neutron 

may pass into and out of the CT region many times. 

Data are recorded for each entrance.

 A given surface file can be used to study the effects 

of small changes in the CNS geometry, using much 

shorter calculations. A large change, such as a 

modification in the D2O cooling jacket, will change 

the reactivity of the core, and will require a new surface 

Figure 3.2. An MCNP geometry plot (in cm) of the cold 

source region containing the Unit 2 CNS.

-80 -60 -40 -20 0 20 40 60 80 -80 -60 -40 -20 0 20 40 60 80
16

source to be generated. 

Figure 3.2 shows just the CNS region as specified 

in the input file “unit2” (the input file is listed in 

Appendix A of this document). The small circle at 

the top of Figure 3.2 is a DXTRAN sphere, just large 

enough to enclose the current tally surface, which is 

a 6 cm diameter circle perpendicular to the y ́ axis 

and 124 cm from the center of the CNS. At every 

neutron collision, a pseudo particle is generated, and 

the probability that it will be scattered and transported 

to the DXTRAN sphere is calculated. Inside the 

sphere, pseudo particles are transported normally, 

contributing to tallies. Very few cold neutrons will 

be transported to the tally surface in analog MCNP; 

the neutron is killed if it does manage to enter the 

sphere. The DXTRAN feature is a powerful variance 

reduction tool described at length in the MCNP 

User’s Guide. DXTRAN cannot be used in problems 

with repeated structures, preventing its use in the 

original KCODE calculation of the NBSR core. As 

another variance reduction measure in “unit2”, the 

neutron importance is increased from 1 to 27 as 

neutrons move toward the LH2 cells. The probability 

that a fission neutron will be transported to the tally 

surface is about 10-7. A combination of generating 

a surface source and using it, together with MCNP 

variance reduction techniques, provides tallies with 

reliable statistics without generating billions of source 

neutrons. 

The tally surface is located near the entrance of the 

in-pile piece of neutron guides NG-1 through NG4. These tallies provide both a relative measure of 

the performance of various CNS geometries, and 

a way to calculate the absolute brightness of the 

source. Brightness is independent of the location of 

the tally used to compute it. Once the brightness is 

determined, the fluence anywhere in the beam line 

can be determined. The direction cosine bin of the 

current tally should be approximately equal to cos θC, 

where θC is the critical angle of the guides for the 

longest wavelength to be “fully illuminated” by the 

source. Most of the NCNR guides are coated with 

58Ni (or 58Ni equivalent super-mirror for the majority 

of their length) which has a θC = 0.02 rad for 10 Å 

neutrons. (The MCNP user specifies direction cosine 

bins, rather than an angle; the angle is between the 

neutron direction vector and a vector normal to the 

tally surface.)

The size of the cold source with respect to the guides, 

determines whether or not the guides are fully 

illuminated. Consider a guide of width, w, located a 

distance, D, from the source. To be fully illuminated, 

the cold source width must exceed w + 2 × D × θC. 

For example, if guides with w = 6 cm are to be fully 

illuminated to 10 Å by a source at a distance D = 110 

cm, the source width must be WCNS > 6 cm + 2 × 110 

× 0.02 = 10.4 cm. The exit hole of Unit 2 is 15 cm 

wide so the guides are fully illuminated to at least 10 

Å. The tally surface should also be fully illuminated 

to 10 Å, so we chose a cosine bin between cos (0.02) 

= 0.9998 and 1.0000. The width of the cosine bin is 

then dμ = 0.0002. A larger value of dμ is sometimes 

used for the brightness calculations to obtain better 

statistics, but as the full illumination condition is 

relaxed, the cold neutron brightness will appear to 

decrease.

To approximate the beam intensity at the guide 

entrance, the current tallies are divided by the tally 

surface area, 28.27 cm2, and multiplied by the 20 

MW normalization factor of 1.525 x 1018 neutrons/

sec per starting neutron (MCNP starting particles 

in the KCODE calculation). The default MCNP 

normalization is per starting particle, so the tallies are 

independent of the length of the calculation. MCNP 

normalization is preserved in the process of generating a 

surface source and using it in a subsequent problem. The 

user can easily verify this with identical tallies in both 

the source generation calculation and problems using 

the source.

Specific tallies needed to compare competing 

geometries are given in the “unit2” input file in the 

appendix. When using the DXTRAN feature for a 

current tally, the surface must lie completely within 

the DXTRAN sphere. The cosine bins of the tally 

and size of the surface must be chosen to exclude 

illumination of the tally surface by the exterior of 

the cold source or any room temperature part of the 

reactor structure. A 6 cm diameter tally surface was 

chosen, approximately 1.25 m from the source, with 

the smallest cosine bins of either 0.9998 or 0.9996. 

Twenty energy bins of 1 meV or 5 meV were chosen 
17

to provide either the cold neutron spectrum or the 

thermal neutron spectrum, respectively. The same 

tallies are used for all CNS calculations.

Sample Brightness Calculation:

Let the brightness be d2J/dEdΩ = B. Its units are 

n/cm2-s-meV-ster, and its value should not depend 

on the distance of the tally surface from the source, 

if the tally surface is on the beam axis and is fully 

illuminated as mentioned above.

Let F(E,Ω) be the MCNP current tally per unit area 

in energy bin dE and cosine bin dμ times the reactor 

normalization at full power, 1.525 × 1018 at 20 MW.

Then F(E,Ω) = B × dE × dΩ, where dΩ = 2 × π × 

dμ.

So B = F(E,Ω)/[(2π) dE d(cosθ)]. 

For the case of Unit 2, F(E,Ω) = 1.9 × 108 n/s cm2, 

for 0 < E < 1 meV, and 0.9998 < θ < 1.0000. Then 

we have dE = 1 meV, d (cos θ) = dμ = 1 - 0.9998 = 

0.0002, and B = 1.5 × 1011 n/cm2-s-meV-ster.

3.3 Nuclear Heat Load Calculations

MCNP can also be used to estimate the energy 

deposited in the LH2 and the Al vessel with the NBSR 

at its full power of 20 MW. The sources of energy 

are:

Fast neutrons.

Fission gamma rays.

Capture gamma rays.

Fission product gamma rays.

Activation product gamma rays.

Activation beta particles.

The code will only directly tally the prompt 

contributions to the heat load, namely items 1, 2, 

and most of 3 (some cross section files lack capture 

gamma-ray production data). Since the activation 

product responsible for most of items 5 and 6 

happens to be 28Al, these contributions can also be 

determined. MCNP 5 includes an aluminum cross 

section file (13027.62c) with the 1.78 MeV gamma 

ray accompanying the beta decay of 28Al added to 

the prompt capture photon production data. This 

1.

2.

3.

4.

5.

6.

extra photon will always be present in a steady state 

case because 28Al has a 2.2 minute half-life and will 

rapidly reach saturation in the cryostat assembly 

and reactor components. Likewise, the 28Al beta 

particle contribution can be calculated by tallying the 

27Al(n,γ)28Al reaction rate, which will be equal to the 

28Al decay rate, and multiplying by 1.25 MeV, the 

average beta-particle energy. It is assumed that all 

of the beta-particle energy is deposited in the cell in 

which it is created.

Gamma rays from fission products are not included in 

any of the cross section sets provided with MCNP. To 

estimate the contributions of item 4, an ENDF5/B 

235U cross section set (92235.50c) was modified at Oak 

Ridge National Laboratory (ORNL) to add “prompt” 

photon production data (92235.02c) characteristic 

of the fission product gamma rays in equilibrium in 

HFIR (High Flux Isotope Reactor) fuel13. These data 

are used for heat load calculations only. 

The energy deposited in the Unit 2 moderator 

chamber from neutrons, beta particles, and gamma 

rays (prompt and delayed), is summarized in Table 

3.1.

The calculated heat load is 1415 watts, based on MCNP 

results using the modified cross sections with delayed 

gamma rays described above. Indirect measurements 

indicate 1150-1200 W, so the measured heat load is 

81 % to 85 % of the calculated value. The same can 

be said for the Unit 1 model heat load calculation, 

950 W. It over-estimated the measured heat load, 

865 W, by about 10 %. (The measured heat load is 

obtained indirectly by comparing the output of the 

refrigerator return gas heater with and without the 

reactor operating. Since some of the refrigerator 

parameters change during the transition, the method 

Table 3.1. Nuclear heating (in watts) in the moderator 

vessel.

LH2 (324 g) - Watts Al (2840 g) - Watts

Neutrons 104 3

Beta particles - 308

Gamma rays 185 815
18

leads to a wide range of measured values. This is 

discussed further in the Section 7, CNS operations.)

3.4 Evolution of the Liquid Hydrogen Sources

As mentioned in Section 2.2, the requirement of 

full illumination of the CTE and CTW beams 

intersecting outside the CT thimble, dictated that the 

source dimensions would be large. An early candidate 

was a double row of vertical tubes, approximating a 

plane source that was roughly a 30 cm square close 

to the front (reactor side) of the thimble. A second 

possibility was a vertical cylindrical annulus, similar to 

the second generation CNS in the Orphée reactor at 

Saclay14 or a spherical annulus. Ultimately we chose 

a 32 cm OD spherical annulus that differed from the 

Saclay source in that the shell had a 20 cm exit hole 

through the LH2 facing the neutron guides. (The 

source at Saclay illuminates beams on opposite sides.) 

The 32 cm annulus with the exit hole is referred to as 

Unit 1 in this report.

Several parameters were varied in a series of early 

MCNP calculations in an attempt to optimize the 

source performance: diameter, LH2 thickness, exit hole 

size, LH2 density (or void fraction in the boiling liquid) 

and the ortho-LH2 content of the cold moderator. The 

latter two may be easy to vary in calculations, but they 

are not so easy to control or measure in practice. The 

outside diameter was chosen to be as large as possible 

given a number of engineering constraints: 

The annulus needed a 5 cm “bubble-cap” vapor separator above 

the sphere connecting to the LH2 supply and return lines (see 

Section 4.2).

It needed to be as close as possible to the front end of the cryostat 

assembly.

The inside radius of the CT thimble is 27.5 cm.

Surrounding the source there were three additional vessels, 

vacuum, helium containment for a 2 mm to 3 mm layer of He, 

and D2O jacket containing about a 2 cm layer of D2O to cool the 

assembly.

Neutron optics dictated the size of the exit hole. Its 

diameter had to exceed the guide height + 2 × D × θC. 

The exit hole in Unit 1 had a 20 cm diameter; it was 

changed in the elliptical geometry of Unit 2 to be 20 

cm high by 15 cm wide.

The optimum thickness of the LH2 annulus depends 

on both the void fraction and the ortho/para ratio. An 

overall density of 0.63 g/cm3 (90 % of the nominal LH2

•

•

•

•

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 5 10 15 20 25 30 35 40

Thickness (mm)

Relative Brightness, 0 - 5 meV, vs. LH2 Thickness

(10 % voids and 50 % ortho-LH2)

Figure 3.3. Relative brightness versus the thickness of the hydrogen moderator, Unit 1.
19

density), and a 50/50 ortho/para content, were chosen 

for the thickness optimization calculations. Figure 

3.3 is a plot of relative brightness vs. the thickness of 

the LH2 annulus for Unit 1. The LH2 volumes of the 

20 mm, 30 mm, and 40 mm thick annuli are 5.0, 6.9, 

and 8.4 liters, respectively. Although the maximum 

gain occurs at a thickness of 32 mm, we chose 20 

mm as a compromise between neutron performance 

and hydrogen inventory. (A 1000 g inventory at 5 

bar was self-imposed in the initial request for a NRC 

preliminary review). Increasing the thickness from 20 

mm to 30 mm, for example, increases the inventory 

by nearly 40 %, but only increases the gain by 11 %. 

At the time, the estimated operating pressure of the 

source was between 1 and 2 bar, so the 30 mm vessel 

was rejected because at 2 bar, there would have been 

very little excess LH2.

Later, the maximum thickness of Unit 2 was increased 

to 30 mm without increasing the inventory because 

the moderator chamber is smaller, a 32 cm by 24 cm 

ellipsoidal annulus. In either case, if the hydrogen 

converted entirely to para-LH2, the thickness would 

have to be considerably greater, 50 mm to 80 mm.

As mentioned in Section 2.4, the scattering cross 

section for ortho-H2 is much greater than that of 

para-H2 at neutron energies less than about 15 meV. 

It is not surprising, therefore, that the gain and the 

cold neutron energy spectrum from the source vary 

considerably over the range of possible para-H2

concentrations, 25 % for normal (room temperature) 

hydrogen to 100 % at 20 K. In the absence of ionizing 

radiation, LH2 molecules would convert slowly to 

over 99 % para, because it has the lowest energy 

level. If, however, the molecules are dissociating and 

recombining at high rate, particularly in a two-phase 

system, the ortho content may be as large as 50 % 

to 75 %. Figures 3.4 and 3.5 show the effects of 

the presence of para-LH2 in the Unit 1 source. The 

conversion of the moderator to 100 % para causes 

the brightness to drop 40 %, and a noticeable shift in 

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 10 20 30 40 50 60 70 80 90 100

Relative Brightness, 0 - 5 meV, vs. Para-LH2 Fraction

(10% voids, 20-mm thick annulus)

Percent Para-LH2

R e l a t i v

e

B r i g

h t n

e

s

s

Figure 3.4. Brightness of the Unit 1 CNS as a function of the para LH2 content.
20

the neutron energy spectrum to an under-moderated 

state. Because the scattering cross section of ortho is 

so much greater than para, just 15 % ortho results in 

82 % of the normal LH2 brightness. From the shape 

of the measured time-of-flight spectrum in Figure 

3.5, it is clear that the ortho content is sufficient to 

dominate the scattering in the moderator. Additional 

evidence of the large ortho content will be presented 

in the section on operating characteristics.

The void fraction due to boiling in the liquid hydrogen 

also affects the brightness, as shown in Figure 3.6. 

To verify that the annular vessel would not result 

in an excessive void fraction or thermal hydraulic 

instabilities, a full scale mockup was built at the NIST 

campus in Boulder for series of tests (see Section 4.3). 

Measurements of the mass of LH2 vs. heater power 

indicated that the average void fraction would be 

about 10 % at the anticipated 800 watt heat load. 

As seen in Figure 3.6, there is only an 8 % decrease 

in brightness as the void fraction 

increases from 0 to 20 %. 

In the first weeks of operating Unit 

1, it was observed that reducing 

the system pressure resulted in a 

brightness increase of up to 5 % 

at the longest wavelengths. This 

result seems to conflict with the 

density dependence (Figure 3.7) 

above, because a pressure increase 

will reduce the bubble size in the 

boiling liquid. The flux increase, 

however, is due to decreased 

scattering of cold neutrons out 

of the beam directed at the 

guides, as the molecular density 

in the H2 vapor decreased. The 

phenomenon is further evidence 

that the LH2 and vapor are largely 

composed of ortho-hydrogen. 

Para-hydrogen, with its very small 

scattering cross section at low 

energies, is nearly transparent 

to cold neutrons. MCNP 

calculations show that brightness 

increases with increasing pressure 

if the vapor is 100 % para, as shown in Figure 3.7. 

A 50-65 % ortho-LH2 concentration in the vapor, 

however, is shown to be consistent with the observed 

behavior.

Fixed-Source vs. Surface-Source Calculations:

Most of the early optimization work on Unit 1 was 

done using a fixed source of neutrons rather than 

the surface-source method described above. A source 

subroutine was added to the MCNP code to start 

thermal and epithermal neutrons from the outside 

surfaces of the CNS region. The spatial distribution 

of these neutrons was dictated by neutron flux profiles 

taken from the original SAR of the reactor, NBSR92

. Starting neutrons were distributed uniformly over 

the base of the cylinder defining the region (55.5 cm 

from the core center), and distributed with a linearly 

decreasing probability along the cylindrical surface, as 

the distance from the core increased. In the absence 

of a model of the reactor core, this source subroutine 

Figure 3.5. Measured TOF spectrum (April 1996) and MCNP calculations for 35% 

and 100% para-hydrogen in Unit 1. The TOF data have been renormalized showing 

that the shape of the measured spectrum is very simialr to the MCNP prediction 

for the 65% ortho-LH2 case.

0 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30

E (meV)

5.0e+10

1.0e+11

1.5e+11

2.0e+11

2.5e+11

3.0e+11

d2

φ/dEd

Ω (cm-2

s

-1

me

V-1ster-1)

Peak brightness (MCNP) 65% ortho:35% para

Peak brightness (MCNP) 100% para

Relative brightness (from NG-1 TOF measurements)

NIST CNS Unit 1 Brightness
21

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0 10 20 30 40 50 60 70 80 90 100

Percent Void

Relative Brightness, 0 - 5 meV, vs. LH2 Void Fraction

(20 mm thick annulus, 50 % ortho-LH2

Figure 3.6. Brightness versus void fraction for the Unit 1 source.

Unit 1 Relative Brightness, 0 - 5 meV vs. Pressure

 and Vapor Para-Hydrogen Content 

2.0

2.2

2.4

2.6

2.8

3.0

3.2

3.4

3.6

3.8

0 50 100 150 200 250

Pressure (kPa)

100% para vapor

50% para vapor

Figure 3.7. Brightness versus pressure for 100% and 50% para-hydrogen vapor (the liquid density held constant at 90%).
22

mimicked the flux expected around the CT thimble. 

There was no effort to normalize the subroutine to 

reactor power.

Although the fixed source calculations could not 

provide absolute values for the source performance or 

the nuclear heat load, they did allow us to compare 

the relative behavior of the candidate geometries, and 

calculate the effects of thickness, density and orthopara ratio. The geometry of Unit 1 was determined 

by the fixed source calculations. 

The first model of the core was used to estimate the 

nuclear heat load in Unit 1 in 1992. By that time, 

the thermal hydraulic tests were well underway in 

Boulder (see Section 4.3), and the 3.5 kW refrigerator 

was installed. A heat load of 800 W to 900 W was 

calculated, assuming that the fission product gamma 

ray heating was equal to the prompt fission gamma 

ray contribution.

The whole-core model confirmed the expected cold 

neutron gain of 22 for Unit 1 with respect to an 

empty cryostat, but predicted a 25 % lower gain 

with respect to the existing D2O source than did 

the fixed source calculations (6.7 rather than 8.8). 

The discrepancy arises because the fixed source is 

insensitive to the impact of the cold source itself on 

the neutron flux in the adjacent fuel elements in the 

core. Unit 1 introduced a large void in the reactor 

reflector compared to the D2O source, lowering the 

thermal neutron flux in the northwest corner of the 

core. In fact, the whole-core model revealed that 

there was a one-time loss of about 0.6 % reactivity in 

the core when Unit 1 replaced the D2O source, worth 

about 2 days of full power operation. As the core 

model was improved, preliminary design work began 

on a second generation LH2 source even as Unit 1 was 

being installed.

The Advanced Liquid Hydrogen Cold Source:

Modifications incorporated in Unit 2 resulted from 

improved computational models and experience 

gained in the operation of Unit 1. Calculations 

demonstrated that filling a portion of the insulating 

vacuum space of Unit 1 with D2O, so the source was 

partially surrounded, would increase the neutron flux 

throughout the cold source region and increase the 

brightness by over 40 %. Experimenting with the 

operating pressure of Unit 1 showed that the vapor 

within the inner sphere was scattering at least 10 % 

of the cold neutrons from the direction toward the 

guides. 

A lengthy series of calculations was needed to study 

and optimize the neutronic coupling of the core and 

the source. An ideal limiting case, although totally 

impractical, is a source like Unit 1, with no vapor in the 

center, nearly surrounded with D2O, and a very small 

exit hole. Such a source could provide one small beam 

with a brightness of 3.4 times that of Unit 1. Since the 

Figure 3.8. Comparison of the Unit 1 assembly (left) and that of the Advanced Cold Neutron Source (Unit 2).
23

cold neutron beam ports through the biological shield 

of the NIST reactor span a range of 16.5o on either 

side of the axis of the center cryogenic port, we needed 

to balance the conflicting goals of surrounding the 

source with D2O and fully illuminating the existing 

guides. MCNP is ideally suited for this task, but the 

process required many, lengthy criticality calculations 

because the addition or subtraction of a few wellpositioned liters of D2

O changed the reactivity of the 

core (keff), the thermal neutron flux in the region, and 

even the fuel utilization.

Engineering constraints must also be considered in 

the MCNP calculations. The moderator chamber is 

surrounded by a vacuum vessel, that is surrounded 

by a helium containment vessel, strong enough to 

withstand the design basis accidental detonation of 

liquid hydrogen and solid oxygen. The helium vessel 

determines the extent of the D2O volume. Thus, 

through an iterative process, the conceptual design 

had to be modified as the complex mechanical design 

was finalized.

The geometry of the advanced source differs from 

Unit 1 in many key respects, as shown in Figures 

3.8 and 3.9. The most important change is the 

addition of 60 liters of D2O to the cooling jacket, 

partially surrounding the moderator chamber. This 

substantially reduced the area of the void through 

the reactor reflector, increasing the thermal neutron 

flux in the cryostat region about 40 %. Unit 2 is an 

ellipsoidal, rather than spherical, annulus, with major 

axes of 320 mm, and a 240 mm minor axis in the 

transverse horizontal direction. This 5 liter annulus 

is between two nearly concentric Al-6061 ellipsoids; 

its average thickness is 25 mm. The center of the 

inner ellipsoid is offset 5 mm, however, so that the 

annulus is 30 mm thick near the core, and 20 mm 

Figure 3.9. Plan view of the Advanced Cold Neutron Source. The cryostat assembly was inserted horizontally into the cryogenic 

beam port. The cold neutron beam ports intersect at a point near the exit of the thimble, about 20 cm from the center of the 

LH2 vessel.
24

thick at the exit hole. The inner ellipsoid is evacuated 

through a small vacuum port. For structural reasons, 

hydrogen vapor fills the elliptical exit hole (200 mm 

by 150 mm), but in Unit 2, the cold neutron beam 

will pass through only 20 mm of vapor, rather than 

300 mm. These changes in the moderator chamber 

added another 20 % to 25 % gain. 

An ellipsoidal annulus provides three advantages. 

Because it has a smaller volume, more D2O can be 

introduced in the cryostat assembly. It is also possible 

to increase the LH2 thickness while maintaining the 

same volume as Unit 1. An elliptical shape is also 

desirable from a neutron optics standpoint; the 

neutron guides at NIST are all rectangular, most are 

60 mm wide and 150 mm tall, so they are still fully 

illuminated with the smaller cold source volume. 

A disadvantage was that the mass of the moderator 

chamber increased over that of Unit 1, from 2.1 kg 

to 2.8 kg, because the inner vessel had become part of 

the hydrogen boundary.

A separate calculation, presented in Section 3.3, was 

required to determine the nuclear heat load in the 

chamber when the reactor is operating at 20 MW. 

For the new moderator chamber described above, 

the calculated heat load is 1400 W. It increased as 

a result of the additional mass of aluminum and the 

higher neutron flux. From our previous attempts 

at a benchmark for Unit 1, however, the MCNP 

result overestimated the heating by 10 % to 15 %, 

so we expected that 1200 W to 1300 W would be 

deposited in Unit 2. Based on our tests at Boulder, 

the increased heat load over Unit 1 would be easily 

removed by exploiting our excess refrigerator capacity. 

The measured heat load was 1150 W to 1200 W.

Unit 2 was first operated in March, 2002. Only one 

aspect of its behavior was unexpected. It was thought 

that with most of the H2 vapor removed, there would 

be an increase in brightness as the system pressure 

increased, owing to the reduced void fraction in the 

LH2. Instead, just like Unit 1, the flux decreased as 

the pressure was raised. Even though there is only 20 

mm of vapor in the exit hole, the loss from the vapor 

scattering still outweighs the gain due to the lower 

void fraction. Therefore, the source is operated at 

nearly the same pressure as Unit 1 (90 to 100) kPa. 

0

5

10

15

20

25

30

35

40

45

0 5 10 15 20 25

Wavelength (A)

Full/Empty

H2/D2O

Figure 3.10. Neutron gain versus wavelength for Unit 1. The bottom curve is the gain with respect to the old D2O source, and 

the top curve is the gain with respect to the source while warm.
25

3.5 Source Performance

The neutron performance of the source has been 

evaluated in many ways, including gain over the 

previous D2O ice source, gain full to empty, absolute 

flux, and cold neutron spectrum. The measured gains 

for Unit 1 are shown in Figure 3.10 as a function of 

neutron wavelength. Agreement with calculation was 

highly satisfactory, and the cold neutron gains (full/

empty) are as high as those seen for any hydrogen 

source. The gains measured at all of the instruments 

were fully consistent with those shown here, which 

are those measured at the NG-7 30 m SANS. The 

intensities at approximately 2.5 Å for the SPINS triple 

axis spectrometer were equal to or better than those 

at the best spectrometers at thermal beam ports at 

similar resolution (note that the NIST guides are not 

curved, providing excellent continuous measurement 

capability from 2.5 Å to 6 Å)

Another benchmark for source performance is the 

total capture neutron flux in the guides, an integral 

measurement that is most sensitive to the longest 

wavelengths (e.g. for a 45 K Maxwell-Boltzmann 

spectrum, over half of the capture flux comes from 

neutrons with wavelengths longer than 10 Å). 

Hydrogen sources are certainly inferior to deuterium 

sources at the longest wavelengths because of 

absorption and, as stated earlier, the primary region 

of interest is below 10 Å. For the Unit 1 source, the 

Table 3.2. Capture flux measurements from Unit 1 at various 

guide locations.

Guide - location φc

meas/cm-2s-1

NG-0 (NDP in C-100) 2. 6 x 109

NG-1 (Reflectometer) 3.0 x 109

NG-2 (Before filter) 3.4 x 109

NG-3 (Before SANS filter) 1.7 x 109

NG-4 (Shutter position) 2.7 x 109

NG-5 (Entrance guide hall) 2.3 x 109

NG-6 (Entrance guide hall) 2.3 x 109

NG-7 (Reflectometer) 1.9 x 109

0

0.5

1

1.5

2

2.5

0 5 10 15 20

Wavelength (A)

MCNP

NG-7

NG-4

NG-1

Figure 3.11. Measured and calculated gains of the Advanced Cold Neutron Source. Some gains exceeded predicted values 

because improvements were made to the neutron guide system.
26

capture flux measured inside the neutron guide hall 

(35 m from the source) was 2-3 x 109 n/cm2/s, see 

Table 3.2. Allowing for guide losses and for the high 

ortho/para ratio in the gas as discussed above, this was 

in satisfactory agreement with the calculated value of 

4 x 109 n/cm2/s, assuming no losses.

A time-of-flight measurement of the cold neutron 

spectrum was obtained at the end of NG-1, and 

plotted against the MCNP estimates in Figure 3.5. 

The TOF data has been corrected for solid angle, 

guide acceptance and reflectivity, and absorption 

in aluminum windows, to determine the source 

brightness. Both the shape and intensity of the 

measured spectrum agree very well with MCNP 

calculations in which the ortho content of the LH2

was 65 % (best fit). Simulations with 100 % parahydrogen predict intensities above 4 Å to drop by a 

factor of two. All of the performance measurements 

are consistent with our conclusion that the liquid 

hydrogen is maintained with at least 50 % ortho.

The neutronic performance of the new source was 

benchmarked by duplicating flux measurements made 

with the Unit 1 cold source at several Guide Hall 

instruments. Figure 3.11 is a plot of the cold source 

gain, defined as the ratio of intensities between Unit 

2 and Unit 1, as a function of neutron wavelength. 

The figure includes the gain factors measured at three 

spectrometers and the gains calculated using MCNP. 

While the agreement is excellent, the measured gains 

were actually somewhat greater than predicted, 

especially at long wavelengths. This additional gain is 

likely due to the new guide sections replaced during 

the cold source installation, an effect not included 

in the MCNP models. The new cryostat design also 

substantially reduces the number of fast neutrons 

from the cold source resulting in improvements in 

instrument signal-to-noise levels.
27

4. Thermal Hydraulic Design

4.1 Design Criteria

The first and over-riding design requirement was 

simple, robust operation under varying conditions 

of heat load (e.g. reactor on/off), which immediately 

focused the design on a liquid hydrogen thermosiphon. 

In this design, which has been used at many facilities, 

heat is removed by allowing cold hydrogen to flow 

into the moderator chamber under gravity, forcing 

warmer hydrogen (vapor, liquid, or two phase) 

out of the chamber to a secondary source of heat 

removal (condenser, heat exchanger). The first step 

in the design was to establish the likely minimum and 

maximum heat loads, as described in the previous 

section. Rather quickly, the neutronics design led to 

results that established the likely mass of hydrogen 

and cryostat metal that would be in the high flux 

area, enabling initial studies of the thermal hydraulic 

design.

4.2 Initial Design

Originally, the maximum expected heat load of Unit 

1 was determined to be 1500 W, with a probable 

value of less than 1000 W. An operating pressure 

of approximately 100 kPa to 200 kPa was chosen 

as a design goal, as a compromise between higher 

pressures which tend to reduce the void fraction 

and lower pressures which tend to reduce scattering 

losses from the vapor in the exit path. This choice 

also yields a reasonably sized storage vessel, and allows 

the use of containment helium pressures only slightly 

above barospheric. The latter condition ensures that 

any leak is from the helium space into the hydrogen 

volume, and that monitoring of the helium pressure 

will give indications of a leak, while minimizing the 

design pressure for the containment system.

Since the cold source was to be installed into an 

existing horizontal beam port, the next step was to 

calculate the length of horizontal piping that would 

be required, and then to size the piping using the 

physical constraints on the location of the condenser, 

which in turn sets the available driving head for the 

thermosiphon. At this stage, the configuration of the 

lines leading into the source also had to be specified 

(to obtain a length, and to ensure that adequate 

radiation shielding would be in place). The final 

piping design, which differs only in detail from the 

original conceptual design, is shown schematically 

in Figure 2.1. For reasons of space constraints, it 

was decided to go with a design in which the liquid 

flowing to the source passed through an inner tube, 

while the return (gas or two phase mixture) passed 

through a concentric annulus surrounding the supply 

line. In this design, the supply and return lines act as 

a relatively efficient counter-flow heat exchanger. This 

design has several advantages:

Compact design to meet space constraint.

Good thermal isolation of the supply line.

A nearly isothermal system, so that all thermodynamic variables 

are determined by pressure alone (allowing measurement of only 

one variable, outside any high radiation area).

Once flow is established, the operating point is stabilized by the 

properties of the supply and return lines acting as a counter flow 

heat exchanger.

And it has some disadvantages (some of which were 

only discovered in later tests):

Fabrication and assembly is complex.

Establishment of initial flow and re-establishment of flow after 

an upset is made difficult by the heat exchanger effects.

For design purposes, a total length of 5 m, of which 

1.5 m is vertical, was assumed. Using these values, a 

preliminary design consisting of a 1⁄2 inch (1.27 cm) 

diameter inner tube, surrounded by a 11⁄4 inch (3.175 

cm) diameter tube was chosen. This provided flow 

areas of approximately 1 cm2 for the liquid supply 

line and 4 cm2 for the return line. Using single phase 

estimates of pressure drop, this design would allow for 

a maximum heat load of more than 2500 W at full 

flow, limited only by the tubing pressure drops. This 

model was used in the tests described below. More 

recent calculations, using the as-built drawings, and a 

model for two phase flow developed by Dukler15, give 

the results shown in Figure 4.1. In these calculations, 

the pressure drop within the moderator chamber itself 

is neglected, since the flow area is approximately 200 

cm2. Note that the total flow through the system 

consists of the liquid going to the moderator, and 

the two phase return, driven by the head arising from 

the difference in density of the liquid and two phase 

fluid. At approximately 3000 W, the fluid leaving the 

moderator chamber is all vapor, and a further increase 

•

•

•

•

•

•
28

in heat would empty the chamber and flow would 

cease.

A key parameter in the neutronics performance of this 

system is the void fraction present in the moderator 

chamber as a result of vapor bubbles rising through 

the liquid. The vapor lowers the effective density of 

the hydrogen, and reduces the effective moderator 

thickness. Although this can be partially rectified 

by making the moderator annulus thicker, the 

intensity in the NIST reactor drops monotonically 

with distance from the center of the reactor, so that 

this strategy is not fully effective, and at some point, 

will result in reduced intensity. The geometry of the 

concentric spheres provides a constant flow area for 

the vapor over the height of the inner sphere; however, 

above this level, the area rapidly decreases, and is at a 

minimum where the vapor return line attaches to the 

outer sphere. This causes a large acceleration of the 

vapor, which tends to carry along extra liquid into the 

return. This could lead to a relatively large volume 

at the top of the source that is almost entirely vapor, 

decreasing source performance. In an attempt to 

improve performance, a “bubble cap” (phase separator) 

was added at the top of the moderator, as shown in 

Figure 4.2. This addition provides a space where the 

area again increases, allowing the vapor to expand, 

and liquid to strike the top of the cap, and drain to 

the bottom of the cap, where holes allow the liquid to 

return to the moderator, in counter-flow. The supply 

line ended in the exit of the vapor line from the outer 

sphere, inside the cap as shown in Figure 4.2.

In order to choose a nominal thickness, an estimate 

of the expected void fraction is needed. One estimate 

can be obtained by using the prediction for the 

terminal velocity of bubbles rising in a swarm of other 

bubbles16:

U = (1-α)

1/2 1.53 [σg(ρL

-ρG)/ρL

2

] 1/4

where:

α = the void fraction

σ = the surface tension

ρL,G = the density of the liquid, gas

The void fraction α is simply given by the superficial 

velocity UGS of the vapor (volume of vapor generated/s 

over the flow area) divided by the terminal velocity U. 

For the estimated heat load of 1 kW (which gives the 

total gas volume generated over the entire height) and 

an approximate flow area of 200 cm2, the average (i.e., 

the superficial velocity at mid-plane, corresponding 

the one-half the heat load) superficial velocity UGS = 

4 cm/s at 0.1 MPa.

Figure 4.1. Calculated performance of the thermosiphon for 

the NIST liquid hydrogen cold source.

Figure 4.2. Original design of the first hydrogen cold source 

moderator chamber (spherical annulus).
29

Therefore,

α = UGS/U = UGS/(1-α)

1/2 1.53 [σg(ρL

-ρG)/ρL

2

] 1/4

This leads to a predicted average void fraction of 28 

%, which could be reduced to 20 % by increasing the 

pressure to 0.15 MPa. This was the preliminary design 

– to operate the source at approximately 0.15 MPa, 

and to optimize the pressure by observing the neutron 

intensity as a function of pressure. The initial thermal 

hydraulic design also assumed that the thermosiphon 

would be operated with single phase flow in both the 

supply (liquid) and return (vapor) leg by controlling 

the hydrogen inventory and pressure.

4.3 Thermal Hydraulic Tests at NIST Boulder

As a result of the uncertainties involved in the novel 

shape of the moderator chamber volume, a full scale 

mockup of the cryostat and transfer lines was built and 

tested at the Boulder site of NIST. The details of the 

mockup, tests and calculations have been published as 

a NIST internal report5. The mockup (of the original 

spherical shell system) was constructed from two 

standard, commercially available glass spherical flasks 

of volume 12 liters and 22 liters. This gave an annulus 

of almost 2 cm, as in the planned source, but a larger 

total volume of 6.755 liters rather than the 5 liters 

planned for the actual source. The nuclear heat load 

was simulated by a resistance heater consisting of a flat 

nichrome ribbon (with dimensions chosen such that 

the heat flux at 2 kW input was less than one-forth 

of the critical heat flux for departure from nucleate 

boiling) wound loosely around the inner glass sphere, 

as shown in Figure 4.3. As a result of the slightly 

larger diameter of the test system as compared to the 

planned source, the flow area between the two spheres 

is 1.137 times as large as that for the actual system, 

and the volume is 1.35 times as large. The very close 

match of the annular gap gives high confidence that 

the results obtained will be representative of the actual 

system. As shown in Figure 4.4, the entire moderator 

chamber was mounted on a balance scale, so that 

the actual mass of hydrogen in the system could be 

measured under operating conditions.

After some initial difficulties achieving an adequate 

insulating vacuum preventing liquid from flowing to 

the vessel, the vessel was successfully filled with LH2

and void fraction measurements as a function of heater 

power were made. A second test was made to assess 

the ability of the thermosiphon to restart following 

a temporary interruption of the LH2 supply. Even 

when the heater was turned off, however, with the 

supply line configured as shown in Figure 4.3, the flow 

of LH2 to the vessel could NOT be restarted until all the 

liquid in the vessel had evaporated. It was not possible 

to overcome the backward flow of vapor up the supply 

line. This failure led directly to an important design 

change, namely, to extend the LH2 supply tube to the 

bottom of the vessel.

A smooth tube was added to extend the liquid fill 

line to the bottom of the annulus, and two complete 

runs were obtained in which the input power could 

be varied and other parameters tested. In the course 

of arriving at this test geometry, a corrugated tube 

was used initially to extend the fill line, and filling 

could not be achieved. This result clarified one of the 

primary issues with this design – the effect of the coaxial 

supply and return lines. In this design, it is difficult to 

change the operating point because the transfer line 

acts as a counter-flow heat exchanger, making changes 

difficult (such as getting the entire system cold). 

During the successful filling of the system, the vessel 

was cooled by individual liquid drops “dancing” down 

the fill line, in a manner analogous to liquid nitrogen 

on a floor. This provided the necessary imbalance in 

the system to allow cooling, and showed clearly that 

corrugated lines (i.e. bellows) would totally inhibit 

filling. All bellows used in the system are lined with Figure 4.3. NIST-Boulder moderator test vessel.
30

a smooth insert to reduce friction, and there are no 

bellows in the liquid supply line. Although attempts 

were made to operate with limited hydrogen flow, this 

proved impossible. However, by keeping the supply 

line full (by constantly overflowing the constant 

head cap), stable operation was routine, under low 

to high (2200 W) heat load. Following a heat input 

change, the system immediately stabilized at a new 

flow condition, with no instability of any kind. This 

was true even for almost instantaneous large power 

changes.

During each of the tests a video record of the boiling 

behavior was recorded for later analysis. This record 

was especially illuminating regarding the modes 

of boiling and bubble rise, which are notoriously 

difficult to model properly. Among the more salient 

observations from these video records are the 

following:

The bubbles rise in a spiral path around the vertical axis, 

completing 1/8 to 1/4 of a revolution during their rise.

The total rise time of a bubble from the bottom of the sphere 

to the top is of order 1s, implying an approximately constant 

velocity of approximately 35 cm/s.

The bubbles seem to be of uniform size, approximately 1 cm in 

diameter.

Bubbles tend to follow each other during their rise.

The areal distribution of bubbles seems to be uniform, and 

relatively stable at any height.

Up to 2200W, there was no liquid-vapor interface in the spherical 

annulus.

The bubble cap on the top of the sphere did seem to accomplish 

some phase separation.

With the constant head in this test, the return line flow was 

always two phase up to 2200W input power.

The void fraction as a function of input power 

at 85 kPa (barospheric pressure in Boulder) was 

measured by observing the weight of hydrogen in 

•

•

•

•

•

•

•

•

Figure 4.4. Diagram of the NIST-Boulder test configuration.
31

the chamber, and determined to be almost linear, 

reaching approximately 20 % at 1800 W. The effect 

of pressure on mass in the chamber at an input power 

of 860 W was measured over a limited range up to 

125 kPa, and the mass was found to increase almost 

linearly. Experimental difficulties prevented a detailed 

examination of void fraction as a function of pressure, 

but a model was proposed that indicated a maximum 

in mass, in the chamber, should occur at 150 kPa.

Based upon this model, and earlier experience in 

the design of a cold neutron source for the Institut 

Laue Langevin4, it was decided to operate the NIST 

source with a sufficient inventory of liquid hydrogen 

that flow would be self-limiting, i.e. by providing the 

equivalent of the constant head cup at the top of the 

hydrogen supply line, as shown in Figure 4.5. This 

required a design for the condenser that provided 

room for excess liquid, so that there would be no need 

to control the inventory too closely. Further, the inlet 

liquid line was extended to the bottom of the annular 

space, allowed restarting of flow after an interruption 

caused, for example, by a refrigerator upset. Finally, 

the decision was made to plan on a hydrogen pressure 

near 150 kPa, but to check for the value that gave the 

best intensity with neutrons.

4.4 Other Components (Condenser, Ballast 

Tank, Hydride Storage)

While sizing all of the non-cryostat components, the 

decision was made to allow for future changes, such 

as development of a liquid deuterium source, which 

was estimated to require 3.0 kW to 3.5 kW of cooling 

capacity. Therefore, a design effort was begun to 

develop a condenser capable of at least that amount 

of heat transfer from cold helium gas to the hydrogen 

vapor. During the commissioning of the refrigerator, 

the plate-type heat exchangers used for that system 

were being examined, and the engineer from the 

manufacturer suggested that the same company could 

make a condenser to our specifications. The actual 

condenser is a hydrogen reflux condenser, consisting 

of multiple welded aluminum plates in counter-flow 

geometry, providing 5.042 m2 of heat transfer area for 

the hydrogen side and 3.526 m2 for the helium side. 

This provides 848 W/K at 5.16 K logarithmic mean 

temperature difference, or 4375 W, thus yielding a 

25 % margin over the 3.5 kW requirement. The 

effective length of each plate is 291 mm. The entire 

assembly is contained in a stainless steel vacuum 

jacket, with connections for helium transfer lines, a 

reflux line for hydrogen, and two junctions to allow 

connection of the condenser to a remote storage tank 

and gas handling system. As shown in Figure 4.5, the 

hydrogen line from the condenser comprises an inner 

supply line along with a concentric, annular return 

line, that attaches to the condenser by means of a 

specially designed fitting incorporating a weir over 

which the liquid must flow. The height of the weir 

is chosen such that any oxygen impurities would be 

trapped there, rather than going into the moderator 

chamber and exposed to high radiation.

The amount of liquid hydrogen present on the plates 

is of interest, since that provides a reservoir for liquid 

during operation, and affects the inventory required. 

To calculate this volume, it is necessary to estimate 

the thickness of the film formed on the heat transfer 

plates. From standard texts17, the film thickness δ is 

given by:

δ = [4 μky(TG-TW)/(ghfgρ(ρ-ρv))] 1/4

Figure 4.5. Arrangement of hydrogen supply and return lines 

at the bottom of the hydrogen condenser.
32

where:

μ = viscosity of liquid = 1.32x10-5 kg-m-1s

-1 at 100 kPa and 20.2 K

k = thermal conductivity = 0.0989 W-K-1m-1

y = distance from top of plate (m)

T

G,W = gas, wall temperature

hfg = enthalpy of vaporization = 445.5 kJ/kg

ρ = liquid density = 70.8 kg/m3

ρv = vapor density = 1.33 kg/m3

This yields an average thickness of 0.8 × 0.00018 × 

y1/4, or 10-4 m for the plate length used, which, when 

combined with the total plate area of 5 m2, gives a 

volume of 0.5 liters of liquid on the plates during 

operation.

The NIST source is designed so that there are no active 

components required in the hydrogen system during 

the entire cycle, from all gas at room temperature 

at a moderator chamber full of liquid. This requires 

provision of a “ballast tank” that can contain all of the 

hydrogen at room temperature and a modest pressure 

(0.5 MPa). The tank is sized at 2 m3, and is filled to 

a warm pressure such that when hydrogen is liquefied 

to fill the moderator, the supply and return lines, 

and the reservoir at the bottom of the condenser, the 

pressure drops to approximately 0.1 MPa. As is true 

for all parts of the design, the entire system, (including 

piping and valves) is surrounded by a blanket of inert 

helium gas. In addition, all parts of the gas handling 

system are protected from accidental damage by steel 

panels, heavy shielding blocks, routing below the 

floor surface, or other means. The ballast tank itself 

is located outside the radius of the reactor crane, and 

protected from damage by steel plates.

4.5 Helium Refrigerator

A helium gas refrigerator was purchased from CVI 

Incorporated in 1990, and installed at NIST in the 

early 1990’s. It was designed for a capacity of 3,500 

watts at a supply temperature of 14 K, before the LH2

cold source was designed. A PLC controller made it 

a fully automated system. In 1995, it was upgraded 

and expanded to monitor and control cold source 

operation (see Section 6) and in 1998 a second, 

backup compressor was added.

System Components:

Originally, the refrigeration system was made up of 

three skid-mounted modules, plus a gas charging and 

storage system, and various ancillary components. 

The skid-mounted modules are respectively the 

Compressor modules (now two modules), the Final 

Oil Removal/Gas Management module and the Cold 

Box module. The gas charging/storage system consists 

of a low-pressure storage tank and a connection to 

high-pressure He cylinders maintained by the reactor 

operations group. Ancillary equipment (major 

components are shown in Figure 5.1) includes the 

following items:

Interconnecting piping for the skids.

A 22712 liter (6000 gal) LN2 tank, and a vacuum jacketed pipe to 

the Cold Box.

A free standing motor starter enclosure.

Uninterruptible power supply (UPS) for the compressor motor 

controls.

Two identical compressor modules consist of 

Sullair Model C20LA704-26 single stage, positive 

displacement, oil flooded (lubricated and cooled) rotary 

screw compressors, bulk oil separators, helium ‘after’ 

coolers, oil coolers, two oil pumps per compressor, 

filters, valves, piping, instrumentation and controls. 

Only one compressor at a time can be operated, but, 

in the event of a failure, a manual switch to the second 

one can be accomplished in a few minutes. Water to 

cool the compressor oil is supplied from a separate 

pump module can provide up to 757 lpm (200 gpm). 

The compressors are ultimately cooled by the reactor 

cooling tower using an isolated pump system, Helium 

Compressor Secondary Cooling (HCSC), served by 

a plate-type heat exchanger and the reactor secondary 

coolant.

The Final Oil Removal/Gas Management module is 

also installed in the Compressor Building. It consists of 

three coalescer vessels with internal elements, a carbon 

adsorption bed with internal particulate filters, three 

control valves, piping, instrumentation and controls. 

Control valves are needed for mass inventory control 

and bypass capability (discharge pressure control).

The vacuum insulated coldbox contains the heat 

exchangers, valves, piping, a purifier and a turbine 

expander. The coldbox is horizontally oriented with 

turbine and valve penetrations in the vacuum vessel. 

A single turbine is accessible for operation and 

maintenance without breaking the vacuum in the 

•

•

•

•
33

coldbox. Helium purification is accomplished with an 

80 K purifier, which is regenerated without removal 

from the coldbox. The turbo molecular pump and the 

rough vacuum pump located on the coldbox module 

are used to maintain a vacuum of order 10-5 Pa within 

the coldbox. A 10.8 kW heater in the return line is 

programmed to maintain 18 K at the return gas inlet 

of heat exchanger HX-3, and a load bypass line from 

the turbine to the heater allows variable flow rates to 

the condenser to follow the CNS heat load. 

A 16990 liter (600 ft3) low-pressure helium storage 

tank is used as a reservoir to store helium when the 

compressor is shut off. The high-pressure cylinders 

assure a nearly inexhaustible supply of “zero” grade 

He, also required to operate the NBSR.

The Helium Refrigerator layout is shown in Figure 5.1. 

The coldbox module is located in room C-100 of the 

reactor building. Nearly 100 m of vacuum jacketed 

piping connects it with the hydrogen condenser, 

also in C-100, on the north face of the reactor. The 

compressors and final oil removal modules are located 

in the compressor building with all controls for 

operation at ground level. The motor control centers 

are also located in the compressor room. The lowpressure gas storage tank is situated adjacent to the 

Compressor Building on a concrete pad on the north 

side, and the high-pressure He cylinders are likewise 

adjacent to the building but on the south side.

Process Description:

High purity helium gas from the high-pressure 

cylinders is supplied to the low-pressure storage 

tank and regulated at a pressure somewhat higher 

than the compressor suction pressure. Helium 

enters the refrigerator cycle, on the suction side 

of the compressor, through a controller-operated 

valve. Helium gas is compressed by the rotation of 

two intermeshing helical rotors, one lobed (the male 

rotor) the other fluted (the female rotor). As the 

rotors turn their point of contact progresses along 

their axis, causing an axial movement of the helium 

gas introduced at the suction pressure and trapped 

in the inter-lobe volume. The inter-lobe volume 

progressively decreases, increasing the gas pressure, as 

it advances down toward the discharge port. Oil is 

injected along the periphery of the rotors to seal, cool 

and lubricate. The screw compressor also is equipped 

with a slide valve for capacity control. When the slide 

valve is closed, all of the gas introduced into the rotors 

is compressed, and when partially open, a fraction of 

the helium is bypassed to the suction side.

On exiting the compressor, the oil laden helium gas 

enters the oil separator. Here it impinges against 

a baffle that separates the bulk of the oil by gravity 

into a sump. The oil is removed from the sump and 

directed through a water-cooled heat exchanger and 

particulate filters to the oil manifold where it is reinjected into the compressor. The gas over the sump 

then passes through a first stage coalescer, reducing 

the oil content to the order of 100 μg/g. This stream 

is then water-cooled (in the after-cooler) before 

going to the final oil removal system. The final oil 

removal system consists of three coalescers in series, 

that remove the oil aerosol to approximately .001 

μg/g followed by an activated carbon adsorber for the 

removal of any vapors (decomposition products and 

trace contaminants).

Following oil removal, the high-pressure helium 

flow (approximately 165 grams/sec for 3500 W) 

enters the coldbox, which thermally insulates the 

cold components of the refrigerator. Helium enters 

the coldbox at a pressure of 19.0 barospheres and 

is cooled to near LN2 temperature in the first heat 

exchanger (HX-1) by receiving refrigeration from the 

low pressure return stream as well as the cold GN2 boiloff from HX-2 (see Figure 4.6). The stream is further 

cooled to 79 K in HX-2 by heat exchange with boiling 

LN2. The helium then passes through an activated 

carbon adsorber (purifier) for removal of impurities 

before being further cooled to approximately 20.7 

K in HX-3 by counter-flow exchange with the low 

pressure return stream. The process gas enters the 

turbine expander at approximately 18.8 bar and 

20.7 K and is expanded to approximately 5.42 bar 

and 14 K, removing approximately 4760 watts from 

the process gas and thus producing the required 

refrigeration. The expander discharge is split, going 

to the load as well as bypassing the load and entering 

the inlet of the return gas heater. The return gas heater 

supplies additional heat to bring the total return gas 

temperature to the control setpoint of 18 K. The 
34

stream then returns through heat exchangers HX-3 

and HX-1 to the compressor suction and the cycle is 

complete.

Generally, the compressors are not operated fully 

loaded because the Unit 2 heat load is only one third 

of the refrigerator capacity. A refrigerator power of 

about 2500 W is more than adequate to cool the 

condenser, make up the 500 W to 700 W heat leak 

in the coldbox end, He transfer lines and condenser 

and still require a few hundred watts from the return 

gas heater. The interaction between the CNS and the 

refrigerator is presented in Section 6.

HX-1

HX-2

HX-3

PURIFIER

EXPANDER

LN2

GN2

COMPRESSOR

LOAD

HEATER

Figure 4.6. Simple flow schematic for the 3.5 kW helium 

refrigerator. See Figure 6.2 for typical operating temperatures 

and pressures.
35

5.0 Mechanical Design

The advanced hydrogen cold source facility was 

designed to be reliable, maintainable and above all 

safe. The success of the first cold source, D2O ice, 

installed in the mid 1980’s, provided the momentum 

for the development of the Cold Neutron Research 

Facility and for the hydrogen cold sources. The D2O 

ice source operated successfully for several years and 

provided valuable experience, insight and knowledge 

in the design and operation of a cold source facility. 

The D2O cold source consisted of a complex maze of 

tubes, valves, pumps and gas systems presented to the 

operator through a variety of interfaces. An integral 

part of the facility was a 1 kW helium refrigerator 

that provided the means to cryogenically cool the 

block of ice. Included with the refrigeration system 

and coupled to it was a liquid nitrogen cooled helium 

circulator (small refrigerator) that was used to keep 

the ice frozen should the main refrigerator fail to 

operate.

Due to the complexity of the D2O ice cold source 

system considerable time and effort was required to 

start, maintain, operate and even shutdown the facility. 

Preparing the cold source for operation required 

making ice by transforming a fixed inventory of heavy 

and light water, cooled with liquid nitrogen, into a 

layered block of frozen moderator material. The icemaking process was complicated and time consuming 

and required meticulous attention to procedure to 

ensure success. Failure to adhere to the prescribed 

procedure could abort the ice-making session and 

possibly damage the cryostat beyond repair. The 

old helium refrigerator, designed and built more 

as an experimental model than a production unit, 

also required a great deal of attention to start, cool 

down and bring the block of D2O ice to an operating 

temperature of 25 K. The refrigerator system was 

manually operated, as were all the systems of the 

cold source facility, and therefore required constant 

attention for a considerable period of time until 

the system and cold source reached their operating 

temperatures. After the systems reached equilibrium, 

refrigerator operation required only inspection and 

data collection at regular intervals. Automatic data 

acquisition was not available for the refrigerator or 

cold source and therefore process information was 

recorded manually by reactor operations staff.

Normal operation of the cold source also required a 

cold source operator to manually perform a slight 

warm up of the source to force a controlled release 

of stored energy from the ice. This operation was 

performed every forty-eight hours to prevent a 

spontaneous release of energy from occurring that 

could adversely affect the operation of the refrigerator 

causing a shutdown and requiring even further 

operator intervention. Shutdown of the cold source 

involved slowly warming the refrigerator to melt 

the ice, warming the water and then subsequently 

draining the water into a radiation shielded holding 

tank. After several days the water could be safely 

removed from the holding tank for proper disposal. 

A fresh mixture of light and heavy water would then 

be prepared for the next freeze cycle. Other significant 

maintenance tasks for the refrigerator and cold source 

included extensive oil vapor removal procedures from 

the highly inadequate oil coalescers downstream from 

the refrigerator compressor and regeneration of the 

cover gas sweep hydrogen system getters.

The complexity of the D2O ice cold source coupled 

with the poor reliability of the 1 kW helium 

refrigerator and associated components plus the quest 

for improved cold neutron performance drove the 

development of a hydrogen cold source system that 

had to be functionally simple. The entire system also 

had to be dependable, easy to operate and maintain 

and above all safe. 

The 3.5 kW refrigerator purchase specification 

required a rotary screw type compressor for efficiency 

and reliability, extensive oil removal capability to 

ensure a clean system, a gas bearing turbine for 

reliable operation and zero maintenance, operator 

initiated automatic operation for easy cool-down, 

warmup and recovery from one central location, fast 

cool-down and warmup, corrosion resistant materials 

to minimize or eliminate required part replacement, 

extensive instrumentation and automated data 

collection to help anticipate problems before they 

develop into failures.

For the hydrogen cold source, simplicity in design was 

achieved principally by changing moderators from 
36

ice to liquid hydrogen. Utilizing a thermosiphon to 

circulate the liquid hydrogen between the moderator 

chamber and the condenser greatly simplified the 

system and allowed for the a completely passive 

design. The vast matrix of valves that existed in the 

earlier D2O source was replaced by just a few that 

are used only for initial charging of the system with 

hydrogen, isolating the cryostat and buffer tank 

during major system upgrades or decommissioning 

the system and subsequent hydrogen removal. In 

addition, there are no pressure relief devices of any 

kind; the hydrogen enclosure is hermetically sealed 

to assure the integrity of the system. During normal 

operation the valve handles are removed to prevent 

accidental operation (the valves have only been used a 

few times in the past ten years). Since the cold source 

is a passive design operator intervention is never 

required to maintain the safety of the system. Should 

cooling fail (refrigerator failure) the hydrogen will 

boil off and safely return to the buffer tank. 

To ensure a reliable and safe cold source, quality 

assurance played a crucial role from the very 

beginning through a judicious approach to system 

operation and the meticulous selection of materials, 

equipment, services and processes. This methodology 

has provided many years of safe and reliable cold 

source operation. A quality assurance plan was 

developed that provided a comprehensive approach 

to quality management and yet provided methods 

that were by design simple, easy to implement and 

execute. Some of the basic elements of the plan called 

for certified material and processes, comprehensive 

tests to ensure reliability and rigorous standards for 

acceptance.

All materials procured for the fabrication of the 

cold source were certified in their composition and 

mechanical properties (e.g., strength and elongation). 

Equipment was chosen based on reliability, simplicity 

of operation, availability and low maintenance. 

Heat trebarents and other processes performed on 

Reactor

Refrigerator 

Cold Box 

Module

Refrigerator Compressor Module

Refrigerator 

Compressor 

Module

Refrigerator Oil 

Removal Module 

Refrigerator 

Low Pressure 

Helium

Hydrogen Gas 

Management

Vacuum 

Insulating 

System

Cryogenic Port Shutters

Outer Plug

Floor Trench

Hydrogen Condenser

Cryostat

First Floor -Room C100

Compressor 

Building

Figure 5.1. The NCNR cold source facility.
37

cryostat components by outside vendors also had to 

be certified.

Non-destructive testing of system components and 

assemblies focused mainly on leak testing using a 

helium mass spectrometer to ensure system integrity. 

Additional tests included x-ray radiography on all 

in-pile welds, cold and warm thermal soaks on all 

cryogenic components and assemblies, and when 

needed, dye penetrant and ultrasonic tests. Pneumatic 

proof pressure tests were also performed on the 

completed assemblies. In addition, prototypes of the 

hydrogen moderator chamber and helium containment 

vessels were subjected to high pressure (hydraulic) 

tests designed to measure their burst points. The 

moderator chamber ruptured at a pressure of 27 bar 

while the helium vessel survived a pressure of 88 bar 

and remained intact with only slight deformation.

Extremely high standards for accepting components 

and assemblies for the hydrogen cold source increased 

confidence in the reliability of the finished product. 

As an example, the NCNR leak tests provided for a 

nominal acceptance helium leak rate of 10-9 std-cc-s-1. 

Although quite stringent, in reality, any leak that was 

encountered during testing was repaired to a zero leak 

rate (no leak detected). The result was a cold source 

assembly free of known leaks.

 The cold source facility, shown in Figure 5.1, consists 

of several major component assemblies:

In-pile cryostat assembly.

Cold hydrogen transfer line.

Condenser.

Hydrogen gas management system.

Vacuum insulating system.

3.5kW helium refrigerator.

The in-pile assembly, installed in the cold port of the 

NCNR reactor, is comprised of the hydrogen cryostat 

mated to a stainless steel shield and support, the 

outerplug. The condenser and cold hydrogen transfer 

line are mounted to the face of the reactor immediately 

above the cold port. Located away from the reactor in 

the northwest corner of room C-100 is the hydrogen 

gas management system. A warm hydrogen transfer 

line, protected in a trench in the C-100 floor, connects 

the condenser to the gas management system. A dual 

pump vacuum system sits atop the guide shields (guide 

•

•

•

•

•

•

shields not shown in figure) and provides the thermal 

insulation for the cryogenic system components, the 

condenser and the hydrogen cryostat. The coldbox 

module of the 3.5 kW helium refrigerator resides 

along the west wall of C-100 while the compressors 

and ancillary refrigerator equipment are located in a 

separate building immediately adjacent to the west 

wall of the reactor confinement building.

5.1 Hydrogen Cryostat Assembly

The advanced hydrogen cryostat is an all-welded 

assembly consisting of a moderator chamber 

surrounded by a closely fitting insulating vacuum 

jacket, a similarly shaped helium containment jacket 

and a large water-cooling jacket.

The in-pile cryostat assembly was fabricated almost 

entirely of aluminum alloy 6061. This alloy was 

chosen for the following reasons:

Superior cryogenic properties; ductile at low temperatures, high 

thermal conductivity.

Resistant to hydrogen embrittlement.

Radiation tested; years of experience with this alloy in high 

radiation fields.

Neutronic properties; relatively low capture cross-section.

Ease of fabrication; high speed machining possible, weldable.

To enhance safety and reliability, the use of mechanical 

joints in the hydrogen cold source system was greatly 

minimized. The in-pile cryostat assembly is devoid of 

all mechanical joints. The hydrogen system from the 

cryostat to the ballast tank, including the valves, is 

joined entirely with welds with the exception of a few 

high vacuum/high pressure mechanical joints.

Section VIII, Rules for Construction of Pressure 

Vessels, of the ASME Boiler and Pressure Vessel 

Code provided a design methodology for preliminary 

sizing, vessel wall thickness and allowable pressure 

of the vessel shells. This also provided for the initial 

design feedback (geometry, mass) to the neutronic 

calculations. In addition, the ASME code also 

furnished material requirements such as maximum 

allowable stress and operating temperature envelope. 

Since the unusual shape of these vessels did not readily 

lend itself to easy structural analysis, finite element 

techniques utilizing solid and surface models, were 

used to analyze the pressure vessels under a variety 

•

•

•

•

•
38

of internal and external load conditions. Figure 5.2 

shows typical von Mises stress results from finite 

element analysis (FEA) for the moderator chamber, 

vacuum vessel and helium containment vessel.

The advanced solids based computer aided design 

(CAD) program that was used to generate a variety of 

models for FEA was also used to provide solids based 

models for computer aided machining (CAM). All of 

the cryostat vessel parts were fabricated on computer 

numerical control (CNC) machines programmed 

directly from the 3-D design model files. Large 

pieces of aluminum alloy 6061 plate stock were 

whittled down to form the thin shell components 

that were subsequently welded together to form the 

individual pressure vessels in the cryostat assembly. 

Heat trebarent was used between machining steps, 

as necessary, to relieve residual stresses and therefore 

reduce part warpage. After machining was complete, 

the parts received a final heat trebarent that brought 

the finished pieces to a T4 material condition before 

they were welded together.

The right image of Figure 3.8 is a cross-sectional plan 

view of the cryostat tip showing the inner and outer 

vessels of the moderator chamber, the insulating 

vacuum, helium containment and the water-cooling 

jackets. Both the vacuum and helium vessels are 

extended beyond the moderator chamber along the 

neutron beam to form the neutron beam window.

Considerable effort went into the design and analysis 

stage of the cryostat to reduce the mass of the 

moderator chamber, and thereby minimizing the heat 

load on the hydrogen system and eliminating any 

adverse effects that might occur on the performance 

of the thermosiphon. The liquid hydrogen moderator 

is arranged in the shape of an annulus formed by the 

inner and outer thin-walled shells. The annulus has 

a non-uniform thickness ranging from 2.0 cm to 

3.0 cm and an approximate volume of 5000 cm3 (5 

liters). The inner vessel of Unit 1 cold source was open 

to the annulus and filled with hydrogen vapor during 

operation. In the advanced design the inner vessel is 

open to the insulating vacuum space at the rear of 

the vessel. Evacuating the inner vessel eliminated the 

hydrogen vapor and increased the cold neutron flux. 

The liquid hydrogen supply line and vapor return line 

enter at the top of the chamber. The inner vessel has 

a small opening in the rear and is held in place by the 

outer vessel end cap. The closed space formed between 

the downstream inner vessel shell and the end cap is 

open to the annulus and filled with hydrogen vapor 

during normal operation. (The end cap design was 

later modified for the backup cryostat by eliminating 

the hole that vented the space between the inner vessel 

and the cap to the hydrogen annulus. Instead, this 

small space is also opened to the insulating vacuum 

jacket.)

The insulating vacuum vessel closely follows the 

shape of the hydrogen moderator chamber with a 

nominal operating thermal insulating gap of 0.25 

in (6.4 mm). When warm, the front or upstream 

end of the moderator chamber will bump or barely 

touch the vacuum vessel. As the system cools, the 

hydrogen supply and return tubes contract and pull 

the moderator chamber away from the vacuum vessel 

providing a uniform clearance gap all around the 

chamber.

Moderator Chamber Vacuum Vessel Helium Containment Vessel

Figure 5.2. FEA results for the moderator chamber, vacuum and helium vessels.
39

The rear section of the vacuum vessel is extended 

along the neutron beam line to the end of the cryostat 

tip where it forms a window for the neutrons to exit. 

A tube exiting horizontally from the top of the vessel 

extends the insulating vacuum along the path of the 

concentric hydrogen lines. It also serves as the sole 

mechanical support for the moderator chamber and 

as a pump-out line for the insulating vacuum system. 

The support system is designed to restrict movement 

of the moderator chamber to along the axis of the 

hydrogen line as the temperature of the hydrogen 

system changes.

The insulating vacuum jacket is completely surrounded 

by a helium containment jacket that serves to prevent 

the possibility of undetected air from entering the 

hydrogen system and also serves as a containment for 

any adverse hydrogen-oxygen event. 

From an engineering standpoint, the most significant 

task in designing a cold source with a helium jacket 

is that it leaves one shell, in this case, the vacuum 

vessel, without direct cooling to remove the energy 

deposited from neutron capture and gamma 

radiation. The moderator chamber shells are cooled 

by the liquid hydrogen, the helium shell is cooled by 

the circulating heavy water, but the vacuum jacket 

has no direct thermal cooling and has to rely on 

thermal radiation to the moderator chamber shell or 

conductive cooling to the helium vessel. If the energy 

deposited in the vacuum vessel could not be removed 

effectively then the shell temperature would exceed 

the allowable design operating condition (100  ̊C) and 

compromise the integrity of the system. The amount 

of energy lost to the moderator chamber through 

radiation is insignificant, and furthermore not the 

preferred mode of heat removal for the vacuum shell. 

Fortunately, dimensional constraints from the size 

of cold port thimble (there is only so much room to 

work in) and the desire for a significant heavy water 

presence outside the helium vessel, relegated the gap 

between the vacuum and helium vessels to a small 

one. Since the gap was too small to take credit for 

reliable thermal convection, conduction through the 

helium layer is the only viable method of removing 

the heat from the vacuum shell. The gap was made as 

small as possible to enhance the thermal conduction. 

This allowed for acceptable heat removal and vacuum 

vessel wall temperature and yet still permitted the 

unit to be assembled without too much difficulty. 

Without taking credit for any heat transfer from free 

convection or thermal radiation, a helium gap of 

0.25 cm is sufficient to remove the heat generated in 

the vacuum vessel without exceeding the maximum 

allowable working temperature. To further improve 

the margin of safety, the effective thermal conductivity 

through the helium filled gap was increased by filling 

the space between the vacuum and helium vessels with 

multiple layers of expanded aluminum. This material 

is commonly used to make air filters for HVAC and 

electronic cooling systems. The material is very soft 

and conforms easily to complex shapes such as the 

vacuum and helium vessel assemblies.

The helium containment jacket consists of a thick 

vessel and tubes that completely surround the vacuum 

vessel and pump-out line. The helium jacket, like the 

insulating vacuum jacket, consists of an ellipticallyshaped vessel with a large extended portion along the 

beam line. The entire helium vessel with the vacuum 

and moderator chamber inside is cantilevered off 

the helium vessel rear support plate that forms the 

downstream end of the cryostat. Since this assembly is 

also the primary protection or safety barrier between 

the moderator chamber and the reactor beam port, 

the helium vessel was designed to withstand an 

internal static pressure of 8270 kPa (1200 psi) before 

rupturing.

The cryostat water-cooling system is designed to 

remove the nuclear heat generated in the vacuum, 

helium and water vessels. Water for cooling the 

cryostat assembly is supplied by the reactor auxiliary 

D2O system and delivered to the water vessel through 

the helium vessel rear support with a 2.67 cm (3⁄4 

inch) Schedule 40 pipe and distributed inside the 

jacket through a manifold and four 1.27cm (1⁄2 inch) 

tubes.

Before the exterior hydrogen, drain, water-cooling 

and thermowell lines were added, the cryostat tip 

assembly was placed in a large laboratory oven for 

heat-trebarent (artificial aging). This procedure was 

designed to bring the T4 temper aluminum to a T6 

temper and strengthen the parent material in the weld 

heat-affected zones. 
40

5.2 Outer Plug Assembly

The cryostat assembly is attached to a reusable stainless 

steel plug and radiation shield, shown in Figure 5.1, 

This support also contains several openings for the 

neutron beams as well as the new reactor cold port 

shutters. The plug was designed at the NCNR and 

fabricated by an outside contractor. Two welded 

subassemblies and two large machined pieces are the 

principle components of the plug assembly. Stainless 

steel plates were welded together to form the top 

and bottom sections of the plug. After welding, 

these subassemblies were machined to the required 

dimensions. The two side pieces were machined 

from solid stainless steel plate. Final machining was 

accomplished with these four main pieces bolted 

together. Heavy concrete was used to fill all the hollow 

sections of the top and bottom subassemblies.

The plug also has four adjustable wheels to facilitate 

easy installation and removal. Four long connecting 

rods are used to secure the cryostat assembly to the 

front of the plug. Attachment is made from the 

downstream or “cool” end of the plug. This allows 

the plug to be reused a number of times for different 

cyrostat tips therefore saving considerable hot plug 

storage space.

5.3 Cold Hydrogen Transfer Line Assembly

The cold hydrogen transfer line comprises the section 

of super-insulated vacuum-jacketed concentric 

annular tubes along the face of the reactor from the 

condenser assembly to the in-pile cryostat assembly, 

see Figure 5.1. These tubes are also surrounded by a 

layer of helium. Stainless steel alloy 304 is used for 

all components except for the hydrogen vapor line 

components which used alloy 316, an alloy slightly 

more resistant to hydrogen embrittlement. The 

transfer line has several special sections that provide 

for the thermal expansion and contraction of the 

hydrogen liquid and vapor tubes. Pump-out ports 

for connection to the insulating vacuum system are 

available through this assembly for the cryostat and 

condenser insulating vacuum jackets. The transfer line 

provides an over-lapping vacuum break that separates 

the two vacuum spaces yet maintains the integrity of 

the vacuum and insulation without thermal shorts. In 

a similar fashion, the helium containment systems of 

the cryostat and condenser are split and are so designed 

to eliminate all direct paths from the barosphere to 

the vacuum jacket, an important safety feature.

5.4 Hydrogen Condenser Assembly

The hydrogen condenser assembly utilizes a cold gas 

stream from the helium refrigerator to liquefy the 

hydrogen used to moderate neutrons and cool the cold 

source moderator chamber. Under normal operating 

conditions, cold helium gas enters the heat exchanger 

at approximately 15 K and at a pressure of 3 bar. 

There it is warmed by condensation of hydrogen and 

returns to the refrigerator coldbox at approximately 20 

K. Liquid hydrogen drains to the bottom of the heat 

exchanger into the phase separator region, allowing 

the incoming vapor stream to expand and any liquid 

carried over to be separated and returned to the 

cryostat through the liquid supply line. The operating 

point for the condenser is controlled by the incoming 

helium temperature and mass flow, which in turn are 

regulated by the system hydrogen pressure.

The condenser, like the moderator chamber, is 

surrounded by an insulating vacuum jacket, which in 

turn, is surrounded by a helium containment jacket. 

The entire assembly is mounted to the reactor face 

above the cold port, Figure 5.1, and surrounded by a 

protective shield, also mounted to the reactor face.

The insulating vacuum and helium containment 

jackets each contain pressure vessels that were designed 

to withstand any hydrogen-oxygen event. Each vessel 

was designed using Section VIII of the ASME Boiler 

and Pressure Vessel Code.

The hydrogen condenser is a brazed plate-fin heat 

exchanger made entirely of aluminum. With working 

pressures for the helium and hydrogen sides of 2070 

kPa and 986 kPa respectively, and a 3.5 kW heat 

transfer capability, the condenser can easily handle 

the heat load required for successful operation of the 

NCNR hydrogen cold source. Pressure drop across 

the condenser is less than 7 kPa at a helium flow of 

162 g/s (3500 watts). As shown in Figure 5.1, the inlet 

and outlet pipes for the refrigerated helium stream are 

on either side of the heat exchanger core. Two inch 
41

stainless steel pipe ends are attached to the aluminum 

pipes on the condenser through Bi-Braze transition 

joints. A 3.81 cm diameter drain tube for collection of 

the condensed hydrogen and delivery to the cryostat 

is located at the bottom of the condenser.

A hydrogen phase separator is installed in the drain 

tube at the bottom of the condenser to separate the 

return vapor coming back from the cryostat from the 

liquid supply to the cryostat. A partial cross-sectional 

view of the separator, shown in Fig 4.5, illustrates 

how the vapor return flow from the cryostat is split 

away from the fill tube into two chimneys. Liquid 

hydrogen pools at the bottom of the condenser as 

shown in the figure and flows between the two vapor 

chimneys and into the fill line to the cryostat. An 

umbrella is installed on top of the separator assembly 

to prevent condensate, either impurities or hydrogen, 

from dripping from the condenser core directly into 

the supply or return lines.

5.5 Hydrogen Gas Management Assembly

The hydrogen gas management assembly and the 

warm hydrogen transfer line connecting it to the 

rest of the system are located on the first floor of the 

confinement building, room C-100 and are shown 

in Figure 5.1. The system consists of four principle 

assemblies: the ballast tank, the valve tank, the warm 

hydrogen transfer line and the protective frame 

assemblies. The ballast tank assembly contains a large 

double-walled gas reservoir to safely hold the entire 

hydrogen inventory. The valve tank assembly provides 

the means for isolating the ballast from the cryostat 

and connection ports for charging and removing the 

gas from the system. The protective cage assembly 

consists of steel plates and frame surrounding 

the hydrogen containing components to protect 

them from accidental damage. Also included is the 

necessary instrumentation to assure that hydrogen fill 

and removal operations are performed safely.

All sections containing hydrogen are completely 

surrounded by a helium blanket. This is designed 

to prevent undetected air from entering hydrogen 

system. The plumbing design allows the hydrogen 

system to be separated into sections when the system 

is not operating (i.e. during maintenance) to limit 

the volume of gas available when the system is most 

vulnerable to accidental rupture, since shields may be 

removed.

The warm hydrogen transfer line piping from the 

ballast tank to the reactor biological shield lies inside 

the existing floor trench, where it is protected from 

any possibility of accidental rupture. At the face of 

the reactor biological shield, the line passes through 

the massive radiation shielding to the condenser, that 

is also surrounded by a protective steel enclosure. 

Since the cold source system is not operated unless 

the shielding is in place, the entire hydrogen system 

is enclosed by protective shielding that will prevent 

the possibility of damage as a result of any external 

occurrence while the system is cold, and the inventory 

of hydrogen available for release is large. When the 

system is warm, 98 % of the hydrogen is in the ballast 

tank. Before any shielding is removed, the hydrogen 

is protected from release by closing the pair of manual 

isolation valves shown schematically in Figure 5.3. This 

design protects against a massive spill of the inventory 

to the confinement building, with the possibility of 

delayed ignition and detonation. Furthermore, any 

remaining hydrogen in the warm hydrogen line, 

condenser and cryostat is easily removed with portable 

metal hydride storage units.

The ballast tank assembly consists of two pressure 

vessels, a hydrogen vessel within a helium vessel. 

The vessel assembly was designed and constructed 

by an outside contractor to meet NCNR cold source 

specifications and the requirements of Section VIII 

of the ASME Boiler and Pressure Vessel code. The 

hydrogen vessel is a large, 2 m3, all-welded cylindrical 

pressure vessel, completely contained and supported 

inside the slightly larger helium vessel. Each vessel is 

fabricated from a rolled cylinder with a 2:1 elliptical 

head welded at each end. Each vessel has a wall 

thickness of 0.95 cm. Stainless steel alloy 316 was used 

for the components of the hydrogen vessel and alloy 

304 used for the helium vessel. Six spacers center and 

support the hydrogen tank inside the helium vessel. 

Hydrogen enters and exits the inner vessel through a 

single 6.03 cm (2 inch) pipe located in the center of 

one of the elliptical heads. Helium flows into the outer 

vessel through a 16.83 cm (6 inch) pipe concentric to 
42

the hydrogen fill pipe.

The assembly rests on two supports welded to the 

cylindrical section of the helium vessel, which in turn 

are welded to cross members of the protective steel 

frame.

The valve tank assembly is a flanged pressure vessel 

containing the system control valves and sensors to 

monitor the system pressures. 

The pressure vessel is essentially a stainless steel tank 

composed of a flanged center cylindrical section, 

fabricated from 304 stainless steel, with flanged 

elliptical stainless steel heads bolted to each end of the 

tank body. There are numerous penetrations in the tank 

body to accommodate the mechanical feedthroughs 

to operate the variety of control valves and electrical 

feedthroughs to handle the signal requirements of the 

electrical equipment and instrumentation sensors. 

There is a large opening in the rear of the tank body 

for a pipe connection to the hydrogen ballast tank 

assembly. In addition, a penetration at the bottom of 

the tank body receives the warm hydrogen transfer 

line rising from the floor trench beneath the gas 

management assembly.

5.6 Insulating Vacuum Pump Assembly

The insulating pump assembly consists of two 

identical vacuum pump tank assemblies, one valve 

tank assembly, a support frame with plumbing, and 

several flexible hose assemblies. The vacuum pumps 

and vacuum valves are located inside vessels filled 

with helium. Each helium containment is physically 

separated and monitored. The assembly can pump 

on either of the two insulating vacuum systems or 

both simultaneously and is accomplished using the 

valves located in the valve tank assembly. An external 

manifold and plumbing assembly is used to evacuate 

and fill the helium containments surrounding the 

cryostat, condenser and vacuum pumps. Since the 

Figure 5.3 Advanced hydrogen cold source plumbing diagram.
43

pumps operate in sealed helium containments, 

each tank environment is cooled with a small heat 

exchanger supplied with plant chilled water to prevent 

the pumps from overheating.
44

6.0 Instrumentation and Control

A console on the coldbox module houses the main 

PLC controller (Allen-Bradley PLC-5/40) monitoring 

the refrigerator and cold source instrumentation. 

Two PCs are also located at the console providing 

an interface with the entire system (RSView). Most 

alarms regarding abnormal conditions are generated 

by the PLC, but there are two reactor rundown signals 

that are directly wired into the reactor console. The 

PLC is programmed to automatically step through the 

procedures to cool the system to operating conditions, 

warm it to ambient temperatures, and shut off the 

compressor. All components of the Instrumentation 

and Control system receive AC power from an 

Uninterruptible Power Supply (UPS). 

6.1 Programmable Logic Controller (PLC)

The main PLC was part of the refrigerator operating 

system provided by CVI. It is located in the console 

next to the coldbox, the area where the CNS team 

operates the system. The PLC is mounted in one of 

two racks of instrumentation modules used for analog 

and digital input and output for the refrigerator and 

Compressor 1. The refrigerator was installed long 

before the cold source, and was tested and operated as 

an independent system for over two years. (A 1960’s 

vintage refrigerator was still being used to cool the 

D2O cold source; that system was removed in 1994.) 

A second PLC was included with the refrigerator 

turbine package to monitor its inlet and exhaust 

pressures, bearing, brake, and seal gas pressures, 

exhaust temperature, and speed. This turbine PLC 

determines the run/alarm/trip status of the turbine, 

conveyed as digital data to the main PLC.

The PLC is programmed in ladder-logic code in which 

most steps (called rungs) consist of a conditional 

inquiry followed by an action if the condition is true. 

A printed page of the software has the appearance of a 

ladder. Many subroutines are used to perform specific 

tasks such as startup and shutdown, block transfers 

of data, etc. The original software and database were 

created by CVI, but subroutines were added in 1995 

to include CNS data monitoring and control, and 

again in 1998, to add Compressor 2 instrumentation 

and logic. Regardless of the operating status of the 

refrigerator, the program is always being executed at 

a rate of about 50 times per second. Data memory 

locations are not necessarily updated at that rate, 

however, as block transfers of analog I/O are performed 

asynchronously through memory buffers; most are 

updated at least every 5 seconds. The program can be 

modified as it is running, with due caution, of course, 

so new logic may be tested instantly. The modified 

program is then downloaded to a PC where the most 

recent versions are stored.

As the system was expanded, remote PLC racks 

were added to the data highway, shown in Figure 

6.1. Rack 4, located on the condenser enclosure, 

close to the insulating vacuum system, contains data 

modules for the vacuums, condenser temperatures, 

helium containments, thermocouples in the cryostat 

assembly, as well as AC outputs for controlling the 

vacuum valves. Rack 5, located in the pump house, 

receives compressor cooling water flow rates and 

temperatures. Rack 6, in the compressor building, 

was added when Compressor 2 was installed. At the 

same time, the compressor motor control centers 

were each upgraded with a smart motor controller 

(SMC), communicating power, current, line voltage, 

and faults to the main PLC as Racks 7 and 10.

6.2 Control Program

As mentioned above, the ladder-logic program consists 

of a main program and many (34) subroutines. A list 

of the source code is well over 100 pages. Many of 

the subroutines simply execute data block transfers, 

and floating point data manipulations to express 

parameters in familiar units. Other subroutines are 

PID controls that compare a process value to its 

set point, and adjust the output of a component to 

maintain the variable at its set point. (PID is the 

acronym for Proportional-Integral-Derivative, in 

which the strength of the feedback is calculated using 

any or all of the three algorithms. Proportional is the 

simplest as the feedback is directly proportional to 

the difference between the variable and its set point.) 

Still other subroutines contain the sequence of steps 

necessary for the refrigerator to perform its functions, 

cool the refrigerator, recover from trips, generate 

alarms, etc. A few of these subroutines are described 

below to present an overview of the program.
45

Main Program: 

Most of the rungs in the main program are jumps 

or conditional jumps to various subroutines for 

data transfer. Since block transfers take seconds to 

complete, and the program cycles in milliseconds, 

a block transfer command is not executed until the 

previous transfer command has been completed and 

has set its ‘done’ bit. A conditional jump is one which 

occurs only if the user has selected it, such as a jump 

to the cool-down sequence. 

Auto-Cool:

This subroutine executes a sequence of numbered 

steps required to bring the refrigerator to its normal 

operating conditions from a shutdown and warm 

state, its presumed starting point. It specifies suction 

and discharge set points, opens bypass valves, starts the 

flow of liquid nitrogen, starts and loads the selected 

compressor, realigns bypass valves as temperatures 

drop, starts the turbine when it is cold enough, closes 

the turbine bypass valve, and cools the load to 20 K, 

finally changing status from the Cool-Down mode to 

the Auto-Operate mode. Along the way, most of the 

hydrogen is liquefied and eventually flows into the 

moderator chamber as the thermosiphon starts. The 

process takes about fours hours and usually requires 

no human intervention.

Auto-Warm:

 The Auto-Warm subroutine was included to be able 

to warm the turbine to ambient temperature in three 

hours. It simultaneously opens the turbine bypass 

valve and closes the turbine inlet valve, increases the 

return heater set point to 300 K, and jumps into the 

compressor shutdown routine when the heater inlet 

temperature reaches 290 K. This sequence leaves 

HX-1 and HX-2 quite cold, however, so to warm the 

entire coldbox, the sequence is cancelled and valves 

are realigned manually; this process takes (6 to 7) 

hours to complete.

Shutdown:

The compressor is shutdown ‘softly’ by this subroutine 

in (6 to 8) minutes. The discharge pressure set point 

Figure 6.1. Data Highway configuration for the helium refrigerator and cold source.
46

is reduced slowly to 1240 kPa psia before unloading 

the compressor slide valve. It cools by operating at 

low power briefly before the motor is turned off.

Helium Mass Inventory:

Three proportional-integral-derivative (PID) 

controllers maintain the suction and discharge 

pressure set points, by operating the Mass-In, MassOut and Bypass valves. If the suction pressure is lower 

than the set point, the Mass-In valve opens allowing 

additional He to enter the suction line from the lowpressure storage tank. If it is too high, the Mass-Out 

valve opens, and He flows from the discharge line 

into the tank. The Bypass valve allows flow from 

the discharge line to the suction line if the discharge 

pressure is above its set point.

Cold Source:

The interaction between the refrigerator and the 

CNS is greatly simplified by the facts that both 

are controlled by the same PLC, and that a single 

variable, hydrogen pressure, is used to control the cold 

helium flow through the load lines to the condenser. 

This PID uses only proportional feedback, and the 

subroutine has only eight lines. A built-in load bypass 

line in the coldbox, and the return gas heater, already 

programmed by CVI, are also vital components in this 

scheme. The bypass valve is always left partially open, 

so that most of the flow can be accommodated when 

the condenser load is small, that is, when the reactor is 

shutdown. As reactor power increases, the increasing 

hydrogen pressure drives the PID to incrementally 

open the load line valve, and more cold helium flows 

to the condenser, where it is warmed a few Kelvin. 

The bypass and load flow streams mix before entering 

the return gas heater. The heater maintains the return 

helium temperature at 18 K, and it substitutes for the 

absent load when the reactor is shutdown. In this 

way, the refrigerator cooling power is nearly constant 

as the system follows the condenser load. No human 

intervention is required for the reactor startup or 

shutdown, or during a reactor scram.

Trouble Alarms:

Section 6.4 contains a thorough description of the 

thirty trouble alarms generated by the PLC when 

parameters are outside their normal range. All of 

the alarms are generated in one subroutine added 

to the program in 1995. The bits of three words of 

memory are used for the individual alarm conditions. 

The subroutine also utilizes two other pairs of 

memory words to bypass cold source alarms and to 

acquire “handshakes” from the RSView interface 

acknowledging that the alarms were recorded.

Restart:

A brief power interruption can cause the compressor 

motor to trip. When the power is restored, however, 

it is desirable to restart the refrigerator and resume 

CNS operation. If the refrigerator is in the AutoOperate mode, but neither compressor motor is 

operating, a 20 second timer starts when the high 

voltage is restored. If the power stays on, the AutoOperate mode is cancelled and the Auto-Cool mode 

is initiated (above). The compressor is restarted and 

the turbine starts as soon as the compressor is loaded 

because it is already cold. It takes 15-20 minutes 

to re-condense the hydrogen, so the reactor may be 

restarted (see Section 7 for a detailed account of the 

recovery process).

Compressor 2:

The last subroutine added to the program includes 

several data block transfers from Racks 6, 7 and 10 

needed to integrate the new compressor and SMC’s 

into the system. Compressor 1 transducers are directly 

wired to the console. Data from both are constantly 

read, so the subroutine must also identify which one is 

supposed to be operating. A compressor is “selected” 

when it receives high voltage; that bit is latched until 

the other compressor has high voltage, indicating a 

manual switch was performed.

6.3 User Interface

Two local PCs and several remote PCs provide the 

graphical user interface (GUI) with the main PLC. 

The windows-based software compatible with a 

database in the Allen-Bradley line of PLCs is RSView. 

RSView allows the user to create elaborate graphical 

representations of system components displaying 
47

component screens, a startup screen, and a screen to 

display previously defined trends. The component 

screens and some of their functions are:

Overview – General layout of components; no control functions

Compressor – Suction and discharge gas temperatures and 

pressures; motor power, line currents and voltage; cooling system 

temperatures and flows.

Oil Removal/Gas Management – Suction and discharge process 

variables; sub-screens to change set points or manually control 

gas management valves.

Coldbox – Temperatures and pressures at many points such as 

turbine inlet and exhaust, etc.; status of valves and LN2 system; 

sub-screens to change set points for hydrogen pressure, return 

gas temperature and purifier temperature; ability to manually 

operate bypass valves. See Figure 6.2.

Turbine – Speed, exhaust temperature, and local pressures from 

the turbine PLC; turbine alarm/trip status; no control functions.

Vacuum – CNS insulating vacuums and most He containment 

pressures; control of vacuum manifold valves.

Hydrogen System – Hydrogen pressure, condenser and moderator 

chamber temperatures; cryostat assembly temperature and D2O 

coolant flow; reactor power; control of the hydrogen set point or 

•

•

•

•

•

•

•

temperatures, pressures, the status of valves and 

motors, etc. Control “buttons” can be created such 

that a mouse click can initiate or terminate actions, or 

enable or disable automatic control of a component. 

Data entry fields can be called prompting the user 

to enter a new set point for a parameter subject 

to automatic control, or a new output level for a 

component in manual control. Messages regarding 

alarms and trips are displayed at the bottom of each 

screen.

RSView also performs data logging, providing the 

ability to display trends in past or current data. Any 

of the 420 I/O tags in the RSView database can be 

logged with frequencies ranging from seconds to days. 

Data files can be exported to spreadsheets for archival 

documentation. Section 7 contains plots of logged 

data from startup tests. 

Information is displayed on seven major system 

Figure 6.2. Refrigerator Coldbox screen.
48

Runtime is used only to monitor and control the 

system. It is installed in the reactor control room, and 

on the office computers of the CNS team members. 

Remote PCs connect to the data highway via the 

Building 235 intranet. Very remote PCs, in the homes 

of Cold Source Contacts, connect to one of the two 

local PCs using secure software that allows the home 

PC control of the local PC, at the refrigerator console. 

The contact can then monitor and control the system 

from home. Remote access is needed to respond to 

trouble alarms reported by the reactor staff after office 

hours.

6.4 Rundowns and Trouble Alarms

In the event of a refrigerator or auxiliary D2O pump 

failure, the reactor power must be reduced to avoid 

overheating the moderator chamber or the cryostat 

assembly. Two rundown signals exist, therefore, to 

protect the cold source. A compressor or turbine 

manual operation of load line valve. See Figure 6.3.

The Startup screen gives access to compressor and 

refrigerator controls. A user can initiate or cancel the 

Auto-Cool, Auto-Warm and Shutdown subroutines 

in the PLC program, or manually start, stop, load, 

or unload the compressor. One can reset refrigerator 

or cold source trouble alarms, and access still another 

screen that can be used to bypass any of the cold source 

alarms. The activities launched from the Startup 

screen are critical for proper operation of the cold 

source (and the reactor!), so great care has been taken 

to ensure that the CNS operator will not mistakenly 

upset the system with an inadvertent keystroke or 

mouse click. Three or four mouse clicks are required 

to reach a control screen and initiate a critical action.

There are two versions of RSView. RSView-Works 

is used to establish the database, create the display 

screens, create data logs and trends, and monitor and 

control the cold source and refrigerator. RSViewFigure 6.3. Hydrogen System screen.
49

trip will stop the flow of LH2 to the moderator 

chamber, resulting in a rapid increase in pressure. 

Two hydrogen pressure transducers are connected to 

the reactor protection system to generate redundant 

rundown signals when the pressure reaches roughly 

135 kPa (5 psig). Likewise, two differential pressure 

signals from the D2O orifice flow meter provide 

redundant rundowns if the flow drops below 18.9 

lpm. Rundowns are appropriate, rather than reactor 

scrams, because the temperature increases are not 

rapid in either case. To keep the reactor critical 

and speed the return to full power, the reactor shift 

supervisor may bypass either rundown signal when 

the reactor power has dropped to 200 kW. A return 

to normal operation is permitted when the failure has 

been corrected. Further, following a refrigerator trip 

and restart, the reactor power can be increased to 2 

MW if the hydrogen pressure has started to decrease 

and it continues to decrease. The return to full power 

must wait until all cold source trouble alarms have 

cleared.

In addition to the two rundown signals, there are 

thirty abnormal cold source or refrigerator conditions 

that generate alarms through the ladder-logic code 

executed by the PLC. These alarms annunciate in the 

reactor control room, indicating that some physical 

parameter is outside its normal range, but the signals 

do not directly affect reactor operation. The reactor 

operators are instructed to phone the Cold Source 

Contact, who will decide if the problem requires a 

shutdown, immediate repairs, or if the problem can 

wait until the next work day or next reactor shutdown 

for a fix. Tables 6.1 and 6.2 list all the alarms; they are 

briefly described below.

Hydrogen Pressure:

There are redundant alarms if the hydrogen pressure is 

above or below its normal range with the refrigerator 

in its Auto-Operate mode. These alarms could be due 

to a refrigerator failure, in which case there will be 

other alarms, or could indicate the control system is 

malfunctioning. Three additional alarms are generated 

if there is low hydrogen pressure when the condenser 

temperature is far above the hydrogen boiling point 

(normally when the system is shutdown); these would 

alarm if there was a large hydrogen leak.

Insulating Vacuum:

There are two levels of vacuum alarms, ‘poor’ vacuum 

and ‘lost’ vacuum, with set points of .133 Pa and 

13.3 kPa, respectively. A poor vacuum may be due 

Table 6.1 Cold Source Alarms from the PLC.

Cold Source Trouble Alarms

1 VG > 13.3 Pa -- Gross Lost Vacuum

2 HD-1 or HD-2 > 20 % LFL -- Hydrogen leak in C-100

3 VG-1 and VG-2 < 10-7 Pa -- No AC Power to Vacuum Gauge

4 Flow - A < 18.9 lpm -- Loss of D2O Cryostat Cooling

5 Flow - B < 18.9 lpm -- Loss of D2O Cryostat Cooling

6 105 < P-5 < 115 kPa -- Ballast Tank Containment Leak

7 P-8 < 105 kPa -- Pump ‘A’ Containment Leak

8 P-9 < 105 kPa -- Pump ‘B’ Containment Leak

9 P-11 < 105 kPa -- Vacuum Valve Containment Leak

10 P-12 < 105 kPa -- Moderator Containment Leak

11 P-13 < 105 kPa -- Condenser Containment Leak

12 VG-1 > .133 Pa -- Poor Vacuum

13 VG-2 > .133 Pa -- Poor Vacuum

14 P1-A not 65 kPa to 125 kPa -- Abnormal Hydrogen Pressure

15 P1-B not 65 kPa to 125 kPa -- Abnormal Hydrogen Pressure

16 P1-A < 375 kPa, Condenser > 250 K -- Hydrogen Leak

17 P1-B < 375 kPa, Condenser > 250 K -- Hydrogen Leak

18 PLC Rack 2 or Rack 4 Fault -- Alarm Disabled

19 P-3 < 375 kPa, Condenser > 250 K -- Hydrogen Leak

20 P-1a Data Transfer Error -- Alarm Disabled

21 P-1b Data Transfer Error -- Alarm Disabled

22 VG Data Transfer Error -- Alarm Disabled 

23 He Data Transfer Error -- Alarm Disabled 

24 D2O Tin > 50 oC or Tout > 60 oC, Reactor > 1 MW -- Loss of Cooling

25 Pump ‘A’ or ‘B’ Containment Temp > 27 oC -- Loss of Cooling

Table 6.2 Refrigerator Alarms from the PLC.

Refrigerator Trouble Alarms

1 Turbine or Compressor Trip in Auto-Operate Mode

2 Compressor Trip after Step 40 in Auto-Cool Mode

3 Compressor Trip in Auto-Warm Mode

4 Compressor Trip during Auto-Shutdown 

5 Compressor Oil Temp > 50 oC

6 HCSC Flow < 189 lpm with Compressor Operating
50

to a malfunctioning pump or very small leak, which 

might be rectified by using the standby pump. A lost 

vacuum is insufficient insulation, and could indicate 

a hydrogen leak. The lost vacuum alarm is the only 

alarm that causes an active response; it seals the vacuum 

vessels by closing all the vacuum valves, and it stops 

cooling the condenser by closing the refrigerator load 

line valve. (A rundown signal would quickly shutdown 

the reactor, also, if it was operating.) Two additional 

alarms could indicate vacuum system problems. If 

both gauges indicate vacuum less than 10-6 Pa, an 

alarm indicates a loss of AC power. Finally, if either 

pump containment is too warm, an alarm indicates a 

problem with the chilled water system.

Helium Containment Pressures:

There are low-pressure alarms for all of the He 

containments that would indicate leaks to the 

barosphere or into the vacuums. In addition, there is 

a high-pressure alarm on the ballast tank containment, 

He-III, which could indicate a direct hydrogen 

leak. (Since He-III has the reference pressure for 

two hydrogen differential pressure transducers, the 

allowable pressure range in He-III is just 10 kPa.)

Cryostat Assembly Cooling:

In addition to the rundown, there are redundant low 

D2O flow alarms. There are also high temperature 

alarms for the D2O coolant to indicate insufficient 

cooling of the cryostat.

PLC Rack Faults:

If the PLC is not being updated because block transfers 

of data are failing, the PLC may continue to use the 

most recent value of the parameters it received, which 

would disable alarms using these data. Fortunately, 

the block transfer control software sets a fault bit 

if the transfer is not successfully completed. Four 

alarms were included to monitor the transfers of data 

from input modules for the vacuum gauges, the He 

containment pressures, and two separate modules 

with hydrogen pressure inputs. In addition, another 

alarm is generated if the PLC fails to communicate 

with either Rack 2 or Rack 4.

Hydrogen Detectors:

There are two hydrogen detectors monitoring room 

C-100, one just above the ballast tank, and one on 

the ceiling over the condenser. If either indicates the 

presence of hydrogen at a concentration of 0.8 %, 20 

% of the Lower Flammable Limit (LFL), there are local 

audible and visual alarms, and there is a cold source 

trouble alarm annunciated in the control room. The 

output of the detectors is then automatically displayed 

on the PCs. If the H2 level continues to increase, the 

operators will have C-100 evacuated and shutdown 

the reactor.

Refrigerator Trouble:

There are six refrigerator alarms, four of which are 

generated by a turbine or compressor trip in the four 

modes of automatic operation. The two remaining 

alarms indicate problems with the compressor cooling 

system. Since the reactor secondary and cooling 

tower comprise the ultimate heat sink, an alarm on 

the secondary flow to the heat exchanger indicates 

low flow. There is also an alarm if the compressor 

oil temperature exceeds 50  ̊C. (Each compressor is 

internally protected, as well, against high discharge 

temperature or pressure, low suction pressure, low 

oil pressure, and high oil temperature. The SMCs 

protect the compressor motors from operating with 

low or high line voltage or high current.)

The above PLC alarms are connected to the annunciator 

panel in the control room in a fail-safe manner. That 

is, the alarm is actually generated by the absence of an 

“All Clear” signal from the PLC. Alarm conditions set 

bits in words that give these memory locations nonzero values, and cause the PLC to stop generating the 

all clear signal. In this way, the alarms are generated if 

the PLC is disabled or loses AC power.
51

7.0 Startup and Operation of the Liquid 

Hydrogen Cold Sources

The startup of the first LH2 CNS occurred in 1995, 

and Unit 2 started in 2002. This section contains 

summaries of the initial startup of each unit and 

their operating records. It is a record of safety and 

reliability. There have been no accidents or hydrogen 

leaks of any kind. Over its 38 reactor cycles Unit 1 

was unavailable for only 14 of the days the reactor 

was scheduled to operate (1440 days); that is better 

than 99 % reliable. Unit 2 has had 100 % reliability. 

A chronological list of cold source milestones is 

presented in Table 7.1.

7.1 Authorization

As design work on the Unit 1 LH2 source was nearing 

completion in 1992, a description of the system then 

envisioned was forwarded to the Non-Power Reactors 

Project Directorate in the Office of Nuclear Reactor 

Regulation of NRC, for an opinion as to whether the 

installation of the source could proceed under the 

provisions of 10 CFR 50.59 (see Section 2.1). Our 

proposal was sent to a group in INEL for review, and 

after a round of questions, and a small modification 

in the piping, we received a favorable reply from NRC 

on May 17, 1993. In the meantime, a separate outside 

review panel of experts in cryogenics, reactor safety, 

and nuclear radiation transport calculations, came 

to NIST in March, 1993 for a second independent 

appraisal of the proposed source. That panel strongly 

recommended the installation of the source to the 

Chief of the Reactor Radiation Division (now NIST 

Center for Neutron Research) and to the NBSR Safety 

Evaluation Committee (SEC). The SEC reaffirmed 

its support for the proposed source in Meeting 318 

on May 6, 1993.

An Engineering Change Notice (ECN #413) was 

prepared for the committee, and it was approved 

on October 3, 1994. Finally, in early September of 

1995, the SEC reviewed the startup and operating 

procedures, and approved the loading of up to 5 bar 

of hydrogen into the system and operation with liquid 

hydrogen in the NBSR. The ECN required startup 

tests in the first reactor cycle to measure:

Table 7.1 Liquid Hydrogen CNS Milestones.

Date Milestone

26-Oct-1994 Unit 1 installed in CT beam port

8-Dec-1994 Condenser mounted on reactor face

26-Jun-1995 Ballast tank and hydrogen manifold complete

3-Aug-1995 Refrigerator connected to CNS

8-Sep-1995 Insulating vacuum system installed

12-Sep-1995 725 grams of hydrogen loaded into system

First filling with LH2

25-Sep-1995 Reactor startup with refrigerator OFF; 10 MW limit

27-Sep-1995 First reactor cycle with LH2 source

5-Oct-1995 Power reaches 20 MW

6-Oct-1995 Thermosiphon restart tests; failed at 250 kPa

20-Oct-1995 Test of ortho/para catalyst pump

23-Oct-1995 Reduced operating pressure to 105 kPa

6-Nov-1995 End of first cycle

4-Apr-1996 Turbine inlet valve fuse blown - 1 lost day

19-Nov-1996 Late cycle compressor trip - 2 lost days

3-Mar-1997 Approx 18 g of hydrogen removed to hydride unit

26-Mar-1997 Add hydrogen to return to 725 g

23-Nov-1997 Turbine speed circuit failed - 2 lost days

25-Feb-1999 Start first cycle with Compressor 2

26-Feb-1999 Comp 2 and Comp 1 fail - 1 lost day

10-Aug-2000 600-Amp breaker fails - 1.5 lost days

25-Aug-2001 EOC #38; last day of operation of Unit 1

28-Aug-2001 18 g of hydrogen removed from source & condenser

2-Nov-2001 Unit 1 removed; stored in high-radiation cavity

23-Nov-2001 Unit 2 installed

29-Jan-2002 Ortho/para catalyst and pump removed

16-Feb-2002 Unit 2 ready; add hydrogen to 500 kPa

19-Feb-2002 First fill of Unit 2 with liquid hydrogen

6-Mar-2002 Start 1st cycle on Unit 2; 20 MW; P = 134 kPa

Test of thermosiphon restart

14-Apr-2002 Lower pressure to P = 115 kPa

24-Apr-2002 100 grams of H-2 removed to hydride; P = 425 kPa

2-May-2002 Start 2nd cycle at P=100 kPa

24-Aug-2004 14 g of hydrogen removed; not replaced
52

The moderator chamber temperature with the refrigerator 

shutdown and its insulating vacuum backfilled with helium.

The power deposited in the cryostat assembly based on D2O flow 

and ΔT.

The rate of increase of the assembly temperature with no D2O 

flow.

The response of the system to a reactor startup, power increase, 

rundown and scram.

The optimum steady-state temperatures and pressures of all 

components at 20 MW.

The recovery rate of the system following a refrigerator trip.

7.2 Startup and Operation of Unit 1

By September 12, 1995, installation of all of the 

components of the cold source was complete. One 

minor deviation from the system described in Section 

5 was made; the Vacuum Valve containment vessel, 

He-VI, was filled with nitrogen (also inert) rather 

than He because the cold cathode vacuum gauges 

failed in a helium environment. (The gauges were 

modified six months later, and He has been used in all 

of the containments since then.) The volume of the 

hydrogen system was measured earlier by loading the 

system with known quantities of helium; precise mass 

measurements were made on three He gas cylinders 

before and after filling the ballast tank with each. The 

volume of the ballast tank is 2009 liters (± 0.25 %), 

the volume outside of the tank (moderator chamber, 

condenser and piping) was 48 liters (± 4 %), and 

therefore the total hydrogen volume of Unit 1 was 

2057 liters.

After checking the operability of the area hydrogen 

monitors, about 725 grams of hydrogen was loaded 

into the system, resulting in a pressure of just over 

425 kPa at room temperature. With the reactor 

shutdown the refrigerator was started and the source 

was filled with LH2 about four hours later that same 

day, September 12. Since the hydrogen pressure 

PID was still unfinished, there was considerable 

experimentation with the load and bypass valves. The 

vessel filled rapidly with LH2 once the thermosiphon 

started following a pressure spike as liquid rushed 

into the warm vessel (see Section 7.3, also). The 

refrigerator was then shutdown and the LH2 slowly 

boiled away overnight, the system returning to its 

starting pressure about twelve hours later. From the 

rate of pressure increase, the thermal radiation heat 

load was determined to be only 9 W.

1.

2.

3.

4.

5.

6.

Startup Tests:

Startup tests with the refrigerator shutdown began 

on September 25, 1995. The vacuum surrounding 

the moderator chamber was backfilled with He to 

allow some heat transfer through the vacuum and He 

jackets to the D2O coolant. The moderator chamber 

temperature was measured indirectly by isolating the 

chamber, condenser, and connecting lines (48 liters) 

from the ballast tank, and measuring the hydrogen 

pressure in this part of the system as the chamber (17 

liters) was warmed by the reactor. At 1 MW, 5 MW, 

and 10 MW, the measured temperatures were 40  ̊C, 

67  ̊C, and 114  ̊C, respectively. Test 1 determined 

the conditions under which the reactor could have 

operated with the Unit 1 CNS and refrigerator 

shutdown:

Moderator chamber filled with hydrogen (or helium) to at 

least 60 psia.

Insulating vacuum and containment helium-filled to at least 45 

psia.

Normal auxiliary D2O cooling maintained.

Reactor power not to exceed 10 MW.

This was the only time ever that the reactor has been 

operated above 200 kW with the CNS warm. 

Test 3 was also performed that day. At 10 MW the 

D2O flow was secured and the cryostat assembly 

temperature rose 5  ̊C in 45 seconds. Extrapolating 

to 20 MW, the heating rate at full power would be 

13.3  ̊C/min. Since the average D2O temperature is 

less than 35  ̊C at 20 MW, at least 5 minutes would 

be needed to reach the boiling point. Test 3 verified 

that a low-flow (18.9 lpm) reactor rundown provides 

more than adequate protection against overheating. 

The cryostat heat load at 20 MW, Test 2, was later 

determined to be 27 kW to 30 kW; it is constantly 

monitored and displayed.

Following these tests, the insulating vacuum was 

reestablished and the refrigerator was started the 

following day. On September 27, the first reactor cycle 

with the LH2 source began. Since it was also the first 

cycle with new plate-type primary heat exchangers, 

the power was limited initially to 10 MW. The 

source performed as expected and preliminary gain 

measurements indicated that instruments in the guide 

hall were receiving up to 40 times the beam intensity 

•

•

•

•
53

measured two days before, with the source warm. 

A week later, with the reactor at 20 MW, the response 

of the cold source to a refrigerator failure (Test 6) 

was simulated by twice interrupting the flow of cold 

helium to the hydrogen condenser. Before the start 

of this test, the hydrogen pressure was maintained 

at 134 kPa through the action of a controller on the 

He load line return valve. The cold neutron flux 

was monitored at the NG-7 SANS. In the first case, 

the valve was closed and the hydrogen pressure was 

allowed to increase to 200 kPa before the valve was 

restored to its automatic mode, 47 seconds later. 

The pressure decreased immediately, returning to its 

initial value in about 3 minutes, and the flux dropped 

about 2 %, before returning to its initial value. We 

repeated the test, but allowed the pressure to rise to 

250 kPa before returning the valve to its automatic 

mode (87 seconds). Again the pressure immediately 

started to decrease, but the flux had dropped to only 

10 % of its initial value, prompting a verbal request 

that the reactor be shutdown. The sharp decrease 

in neutron flux, even as the hydrogen pressure was 

decreasing, indicated that the liquid was not refilling 

the moderator chamber. 

As a result of these tests, a maximum reactor rundown 

set point was established at 200 kPa. In addition, 

the reactor is not restarted after a hydrogen pressure 

rundown, until either all of the alarms have cleared, 

or the operators have discussed the situation with the 

Cold Source Contact. It became clear in the first 

cycle, that a turbine or refrigerator trip would always 

lead to a reactor rundown. Later, the set point was 

lowered to 130 % of the operating pressure to speed 

the recovery from such a trip.

Optimization:

The hydrogen pressure was varied to optimize 

performance. Increasing the hydrogen pressure caused 

the long-wavelength intensity to drop, so the pressure 

was lowered to about 105 kPa, where it remained 

for the life of the Unit 1 source. The hydrogen 

recirculation pump, designed into the system to 

counteract the feared conversion of ortho-LH2 to 

para-LH2, was twice tested for many hours during 

the first cycle. Its operation had no measurable effect 

on the beam intensity, and the system was never used 

subsequently. The pump and catalyst were removed 

in 2002 before the startup of Unit 2. 

Both of the above phenomena support the conclusion 

that the LH2 and vapor are NOT completely converted 

to para-hydrogen in the source (see Section 3 and 

Figure 3.5). In addition, if the LH2 were to gradually 

convert to para, we would expect to observe the cold 

neutron flux decrease at the instruments during the 

first few days of a reactor cycle. No changes are 

observed over the 38 days. Finally, if the LH2 were 

to convert to para quickly, in a few hours, there 

might not be any drop to be observed during a cycle. 

However, a refrigerator trip and recovery would mix 

the presumed para vapor from the source with the 

normal hydrogen remaining in the ballast tank. The 

cold neutron intensity after a trip would be greater 

than it was previously. This has never been observed, 

either. One can only conclude that the ortho content 

of the LH2 is always maintained at a fraction high 

enough to cause the scattering to be dominated by 

ortho.

Reliability:

A review of the operating logs for Unit 1 shows that 

there were 46 refrigerator trips during its 38 cycles. 

Two of these were turbine trips only (failed inlet valve 

fuse and speed circuit) and the remaining 44 were 

compressor trips, usually caused by storm-related AC 

power interruptions. Only nine of these trips caused 

135Xe poison shutdowns (30-40 hours), resulting 

in about 14 days lost (though some days were just 

postponed); thus the reliability of the Unit 1 CNS 

was 99 %. A few trips coincided with other reactor 

system failures that prevented restarts.

Of the 44 compressor trips, the system restarted 

itself eighteen times, and was successfully restarted 

manually fifteen times. The manual restarts involved 

either re-initiating the Auto-Cool program, or, on 

eight occasions starting in January 1999, switching 

compressors and restarting Auto-Cool. It takes less 

than 5 minutes to make the switch, so unless the trip 

comes in the last few days of a reactor cycle, there is 

ample time to change to the other compressor and 

start it, rather than diagnose the failure of the first 
54

compressor to restart. The cold source can recover 

completely 15 to 20 minutes after the refrigerator 

restarts.

7.3 The Advanced Liquid Hydrogen Cold 

Source, Unit 2

Unit 1 was replaced during a six-month shutdown in 

2001-02. It had not reached the end of its life, nor 

was there anything wrong with Unit 1. The Advanced 

LH2 Cold Source was installed because it was expected 

to increase the brightness by 70-80 %, as described in 

Section 3. A new cask and transfer system were built 

for the removal of Unit 1, and the entire job of pulling 

it from the reactor, transporting it to its storage vault, 

and securing it safely inside the vault was accomplished 

in 5 hours on November 2, 2001. Only the cryostat 

assembly was replaced; the condenser, ballast tank, 

refrigerator, vacuum system etc. are still in service. 

We needed only to sever the D2O and LH2 supply 

lines to free Unit 1 from the reactor, and rebuild these 

lines to connect Unit 2 to the existing components. 

Authorization for the change was sought and received 

in the form of an Addendum to ECN #413, as there 

were no unreviewed safety questions.

Thermosiphon:

Having removed the large volume of H2 vapor from 

the inner vessel, it was anticipated that the optimum 

operating pressure for Unit 2 would be 150-200 

kPa. Accordingly, once the system was completely 

reassembled, over 100 more grams of hydrogen were 

added, increasing the pressure at STP to 500 kPa, and 

the total inventory to 825 g. The new source was 

first filled with LH2 on February 19, 2002. A plot 

of the hydrogen pressure and moderator chamber 

temperature as the thermosiphon started is shown 

in Figure 7.1. Unit 2 behaved very much like Unit 

1, although the pressure was a little higher, and, this 

time, the thermocouples on the LH2 vessel survived 

the installation welding. As the condenser cooled 

toward 20 K, LH2 collected beneath and between the 

plates, prevented from filling the vessel by the back 

pressure created as droplets evaporate in the lines and 

vessel. An unstable equilibrium existed 25-30 minutes 

until a larger quantity of LH2 reached the vessel 

where it boiled rapidly, producing the 60-kPa (10-

psi) pressure spike. Sometimes there are two or three 

spikes, and sometimes a spike is barely noticeable, but 

the start of the thermosiphon is unmistakable because 

F irst L iquid Hydrogen F ill of Unit 2, Warmup

0

50

100

150

200

250

300

350

400

0 50 100 150 200

T ime (min)

H ydrogen P r e s s ur e (k P a )

V e s s e l T em pe r a tur e (K )

Figure 7.1. Hydrogen pressure and Unit 2 moderator chamber temperature as the thermosiphon starts for the first time.
55

the pressure drops very quickly to its set point. The 

temperature of the vessel also drops rapidly to 20 K 

(the thermocouples, however, indicate 40 K).

Immediately before the thermosiphon started, when 

the pressure was just under 200 kPa, about 60 % of 

the inventory had liquefied, 510 g, corresponding to 

7 liters of LH2. Very little of this had reached the 

moderator chamber, which was still quite warm, so 

the condenser plates were completely flooded; the 

hydrogen side of the condenser has a volume of 6.4 

liters, and there is only a 0.5 liter plenum below it. 

Once the thermosiphon was operating, the pressure 

dropped to 95 kPa, indicating that there were nearly 

9.5 liters of LH2 in the system. Given that the vessel 

holds 5 liters and there may be another 2 liters available 

under the condenser and in the transfer tubes, the 

condenser plates were still nearly half flooded. We 

operated the first cycle at 135 kPa, which would have 

left the plates one-third covered with LH2. (Later, we 

reduced the inventory to 695 g, but it is very likely 

that we still operate with 20-25 % of the condenser 

flooded.)

Still referring to Figure 7.1, the refrigerator was 

shutdown 20 minutes after the thermosiphon started, 

and the hydrogen slowly boiled away. The resulting 

pressure rise appears to occur in three distinct regimes. 

For the first 12-14 min, until the pressure reached 200 

kPa, the thermal radiation heating rate was about 70 

W. At this point, the condenser plates were no longer 

flooded, but the transfer tubes and vessel were full of 

liquid. For the next hour, the heating rate decreased 

from 40 to 20 W, until the pressure was 300 kPa. 

In the final regime, only the vessel contained LH2, 

and the thermal radiation into the vessel alone was 7 

watts.

Startup Tests:

The reactor was first operated with the Unit 2 CNS 

on March 6, 2002. Since the NBSR was never 

operated with Unit 1 warm, it was decided to skip the 

measurement of the moderator chamber temperature 

vs. reactor power with the refrigerator shutdown and 

the vacuum backfilled. Instead, there was a routine 

reactor startup, progressing through steps to 20 MW. 

The tests included verification of stable operation, 

measurement of routine operating parameters, and 

determination of limits governing reactor restart 

following a refrigerator failure. Since we have not 

yet tested our ability to operate the reactor with the 

refrigerator shutdown, the reactor must never be 

operated at a power greater than 200 kW with the 

cold source shutdown.

Routine Reactor Startup:

The first test of the cold source was to verify its stable 

operation as heat deposited in the moderator chamber 

and liquid hydrogen increased from its thermal 

radiation ‘background’ to as much as 1360 watts, as 

predicted by the MCNP results. We had successfully 

demonstrated the operation of the thermosiphon 

with no heat load, and our ability to restart the 

thermosiphon after a refrigerator trip, during two 

weeks of tests before the reactor startup.

To assure that the moderator chamber remained filled 

with liquid hydrogen, and to report initial estimates of 

the gain over Unit 1, we were assisted by instrument 

scientists at NG-0, NG-4 and NG-7. At least one 

of those instrument monitors was recording the flux 

each time the reactor power was increased.

As was the case with Unit 1, the surest indication that 

the thermosiphon was operating properly was the 

decrease in the output of the return gas heater in the 

refrigerator as the heat deposited in the cold source 

increased. The refrigerator is programmed to adjust 

the flow of cold helium to the condenser as needed to 

maintain a nearly constant hydrogen pressure AND 

to maintain the helium return gas temperature at 

18 K. As the rate of evaporation of LH2 increases 

with reactor power, more hydrogen vapor must be 

condensed to maintain a constant pressure, so the 

helium refrigerant is increasingly warmed flowing 

through the condenser and less heat is required from 

the heater. The difference between the heater power 

with the reactor shutdown and heater power with the 

reactor at 20 MW is an indirect measurement of the 

heat deposited in the source (also see Section 7.4).

Figure 7.2 is a plot of the heater output and reactor 

power during the March 6, 2002 startup. The reactor 

power has been renormalized and multiplied by 100 
56

for display purposes. During the startup, the reactor 

power was increased in steps to 100 kW, 1 MW, 5 

MW, 10 MW (for at least one hour), and finally 

20 MW. The actual power somewhat exceeded 

the ‘nominal’ power for the steps prior to 10 MW, 

at which time the reactor thermal power could be 

determined. With every increase in reactor power 

there was a proportional decrease in heater power, 

an indirect measure of the nuclear heat load. (The 

average of dozens of these indirect measurements of 

the heat deposited in the cold source at 20 MW is 

1220 ± 60 watts, about 10 % less than the 1360 watt 

estimate using MCNP.)

Refrigerator Trip and Recovery:

From our operating experience with Unit 1, we know 

we can expect several interruptions of AC power each 

year that will trip the compressor and may or may not 

cause a reactor scram. If there is a refrigerator trip 

without a reactor scram, there must be an automatic 

reactor rundown to protect the moderator chamber 

from over heating. The rundown is generated by one 

of two hydrogen differential pressure transducers, 

comparing the H2 pressure with the helium pressure 

in the ballast tank containment. To verify that the 

thermosiphon will restart, and continue to refill the 

vessel even with the reactor at 2 MW, we deliberately 

stopped the compressor and restarted the Auto-Cool 

sequence 55 seconds later, simulating an AC power 

interruption. Since this test caused a rundown and 

required a quick return to power, we first obtained 

approval from the Chief of Operations, and then 

established close coordination with the reactor 

operators. The above-mentioned flux monitors were 

ready to verify that the vessel remained filled with 

liquid hydrogen.

At the start of the test, the reactor had been at 20 

MW for nearly an hour, so the heat load in the LH2

Unit 2: Initial Reactor Startup

0

500

1000

1500

2000

2500

0 50 100 150 200 250

Time (min)

Power (MW or W)

Heater Output (W)

Reactor Power (MW) x 100

Figure 7.2. Response of the refrigerator return gas heater as reactor power is increased from 0 to 20 MW. As the heat load 

from the cold source is met by the hydrogen condenser, the heater power needed to keep the return gas at 18 K decreases. The 

difference in the heater output provides an indirect measure of the CNS heat load.
57

vessel was at equilibrium (28Al, with a half-life of 2.25 

minutes, reaches saturation and contributes over 20 

% of the heat load). The hydrogen pressure was 134 

kPa, and reactor rundown had been set at 200 kPa. A 

brief chronology of events, as recorded every 5 seconds 

by the data logger, is as follows:

17:46:37 STOP Compressor 1.

17:47:12 Reactor rundown.

17:47:32 Start the Auto-Cool sequence, successfully 

 restarting the compressor.

17:49 Rundown bypassed (by reactor operators) at 

 approximately 200 kW.

17:50:12 Refrigerator turbine starts OK.

17:52:32 Maximum H2 pressure recorded, P-3 = 355 kPa.

17:53 Start increasing reactor power to 2 MW.

17:55 Reactor at 2 MW; P-3 = 195 kPa and CNS alarms 

 clear.

17:59 Increasing reactor power.

18:05 20 MW.

Figure 7.3 shows the hydrogen pressure and reactor 

power during the test. After the refrigerator has 

cooled the condenser below about 25 K, the H2

pressure dropped rapidly. The thermosiphon restarts 

•

•

•

•

•

•

•

•

•

•

immediately, supplying the vessel with more LH2. 

The vessel remained filled based on the behavior of 

the flux monitor at NG-7; the monitor count rate 

had exactly the same time-dependence as the reactor 

power curve in Figure 7.3. Increasing the reactor 

power to 2 MW slows only slightly the rate of the H2

pressure decrease. Thus it is safe to raise the reactor 

power from 200 kW to 2 MW, holding at 2 MW 

as long as the pressure continues to drop. After all 

alarms have cleared, the power can be raised to 20 

MW. The increase to 2 MW is at the discretion of 

the reactor supervisor, of course, but from the data, it 

could speed the recovery by about 3 minutes.

The ultimate confirmation of the proper operation of 

the thermosiphon, in the absence of a flux monitor, 

is the feedback through the return gas heater, as 

discussed in the reactor startup monitoring above. 

Figure 7.3 also shows reactor power and the heater 

output during the trip test. The heater output is slow 

to return to its initial value because the refrigerator 

must return to equilibrium after the trip. After 

Unit 2: Refrigerator Trip and Recovery

0

500

1000

1500

2000

2500

3000

3500

4000

0 50 100 150 200 250 300 350 400

Time (5-sec intervals)

Power (W or MW*100) or Pressure (kPa*10)

Reactor Power (MW x 100)

Hydrogen Pressure (kPa x 10)

Return Gas Heater (W)

Figure 7.3. Unit 2 startup test of the ability to recover from a refrigerator trip.
58

passing 10 MW, however, the heater power was 

decreasing while the reactor power was increasing, 

proof that the thermosiphon had restarted. In this 

case, the heater power was about 620 watts before 

the trip, and returned to a new equilibrium of about 

750 watts, as the reactor returned to 20 MW. When 

the reactor was shutdown, the heater supplied about 

1900 watts in the absence of a load on the condenser. 

It is this roughly 1300 watt difference that proves the 

thermosiphon is supplying LH2 to the vessel.

Typical values of operating parameters for Unit 2 and 

the refrigerator as of November 2005 are listed in 

Table 7.2.

Reliability:

Since the installation of Unit 2, the CNS reliability 

has been 100 %. “Lost” time due to the cold source 

is measured in minutes because it has not caused a 

single 135Xe poison shutdown in 22 reactor cycles. The 

improvement is not the result of any design change of 

Unit 2. The restart routine in the PLC was modified 

and refined several times since the SMCs were first 

used in 1999, reflecting our experience with them. 

In addition, the large capacitors installed to balance 

the three phases of the 480 Volt supply in each motor 

control center were disconnected in February 2002 as 

we prepared to resume operation. Since that time, the 

compressors have been less likely to trip because of a 

brief AC power interruption. There have been only 

six compressor trips, and only one required changing 

compressors to recover.

7.4 Heat Load Measurements

Every shutdown and startup of the reactor provides 

an opportunity to record the change in the return 

gas heater power and make an indirect estimate of 

the nuclear heat load. Figure 7.4 shows the results of 

about forty such measurements on Unit 1 (top graph) 

and over sixty values for Unit 2. Unfortunately, the 

very nature of the measurements introduces a large 

uncertainty in the results, because the refrigerator 

parameters vary somewhat as the load-line valve, CV424, opens or closes. As the valve opens, for example, 

from a few percent with the reactor shut down, to 40-

50 % at 20 MW, the turbine exhaust pressure drops a 

few psia, increasing the turbine speed, efficiency and 

the refrigerator cooling capacity. An additional part 

Table 7.2 Typical operating parameters for Unit 2 and 

refrigerator.

Parameter Value Units

Hydrogen Volume 2048 liters

 Ballast Tank 2009 liters

 LH2 Vessel 5.0 liters

 Condenser (H2 side) 6.4 liters

 Piping, etc. 27.6 liters

Pressure (warm) 420 kPa

Density (warm, 300 K) 0.339 kg/m3

Inventory 695 g

Pressure (cold) 97 kPa

 Temperature (saturation) 20.2 K

 LH2 Density 70.8 kg/m3

 GH2 Density (20 K) 1.34 kg/m3

 Heat of Vaporization 446 J/g

LH2 - Mass (total) 530 g

LH2 - Volume (total) 7.6 liters

Moderator Chamber

 LH2 Mass 325 g

 Al Mass 2840 g

Insulating Vacuums 10-5 Pa

He Containments 110-130 kPa

Refrigerator Power 2500 W

Motor Power 280 kW

Helium Flow * 125 g/s

Temperatures:

 Turbine Inlet 20.6 K

 Turbine Exhaust 14.6 K

 Condenser Inlet 14.9 K

 Condenser Outlet 20.2 K

 Heater Inlet 17.0 K

 Heater Outlet (controlled) 18.0 K

CNS Heat Load ** 1200 W

Heater Output ** 700 W

Losses ** 600 W

* 50 - 60% bypasses the load

** ± 100 watts
59

Figure 7.4. . Indirect Heat Load Measurements on Units 1 and 2. The ΔQ values are the changes in the return gas heater 

output before and after a reactor startup or shutdown. It is usually impossible to control all of the refrigerator parameters, 

which results in the large variation in the values of ΔQ.

1000.0

1100.0

1200.0

1300.0

1400.0

1 21 41 61 81 101 121 141 161 181 201 221

Return Gas Heater Power Change (SU or SD) for Unit 2

(Indirect CNS Heat Load Measurements)

Delta -Q (watts)

Measurement Number

600

700

800

900

1000

1 21 41 61 81 101 121 141 161

Measurement Number

Delta -Q (watts)

Return Gas Heater Power Change (SU or SD) for Unit 1

(Indirect CNS Heat Load Measurements)
60

of the uncertainty arises from the fact that the heater 

power may vary 30-50 W for no apparent reason. 

The results for both units seem to divide into two 

groups. Compressor 2 was integrated into the system 

at measurement index number 70, for Unit 1, and the 

average of the results before and after that event are 830 

W and 900 W, respectively. (The slide-valve loading 

of Compressor 1 was increased at that point to match 

the somewhat higher performance of Compressor 2.) 

There is an even more distinct divide for the Unit 

2 results when the load bypass valve, CV-421, was 

closed from 75 % to 70 %, at index number 79; the 

average was 1280 W before the change, and 1160 W 

thereafter. It is not clear, in either case, which of the 

two groups has the lower systematic error. Giving 

equal weight to the groups of measurements we arrive 

at the following values for the nuclear heat loads:

Unit 1 865 ± 40 W.

Unit 2 1220 ± 60 W.

Since there is so much excess cooling capacity, it would 

seem that precise knowledge of the heat load is not so 

important. A future replacement or additional source, 

however, could tax the capacity of the refrigerator. It 

would also be nice to know the origin and magnitude 

of the losses in the load and the transfer lines. Finally, 

the heat load measurements are benchmarks for the 

MCNP computational method. The calculated values 

for Units 1 and 2 are 915 W and 1360 W, overestimates 

of 6 % and 11 %, respectively.

•

•
61

8.0 Accident Analysis

The cold neutron sources have operated very reliably 

for more than ten years. This section presents analysis 

of likely abnormal situations and highly unlikely 

accidents. The first five parts of this section describe 

the CNS response to abnormal incidents that have 

occurred or that are likely to occur over its lifetime. 

Parts 6 through 10 are analyses of unlikely or beyond 

unlikely accidents that could damage the cold source 

but would not pose any unreviewed safety question 

regarding the NBSR. These hypothetical events bound 

any credible accident scenarios.

8.1 Reactor Power Change

The refrigerator and hydrogen system control systems 

have been designed to automatically adapt to changes 

of heat load on the moderator chamber and cryostat. 

This implementation has been extensively tested, and 

it operates completely predictably even under the 

most difficult perturbation of a reactor scram. There 

are no consequences of an increase or decrease in 

reactor power.

Conversely, a change in the quantity of LH2 in the 

moderator chamber has a very small effect on the 

reactivity of the core. Lengthy MCNP calculations 

show that emptying the vessel of LH2 adds 0.1

(+0.13$) or (+0.1 %Δk/k). This reactivity insertion is 

just one-fifth of the limit on “moveable experiments” 

in the core, and is easily offset by a 4-5 inch insertion 

of the regulating control rod. 

Flooding of the cryostat assembly vacuum region with 

D2O would represent a larger reactivity insertion, 

nearly 0.5 %Δk/k, which is the limit for a single 

moveable experiment. It is still a much smaller 

and slower reactivity insertion than the maximum 

reactivity insertion accident, +1.3 %Δk/k in 0.5 sec, 

that was shown not to damage the core in NBSR1410.

8.2 Refrigerator Failure

If the refrigerator fails, there will be no cold helium 

to re-condense the hydrogen vapor, and eventually 

the pressure in the moderator/ballast tank assembly 

will start to rise. In this case, the reactor will receive 

a rundown signal when the pressure reaches 135 

kPa, upon which all shim control arms are inserted 

continuously into the reactor until the condition is 

removed. This will reduce reactor power immediately, 

and reduce the heat load on the moderator chamber. 

If the refrigerator function is not restored, all of the 

liquid will evaporate, and the system will be in a safe 

condition, as the reactor will not be restarted until 

cooling is restored and the pressure is reduced below 

the setpoint. The rundown can be bypassed when 

the reactor power is reduced to 200 kW, at which 

power level the heat load is reduced to an insignificant 

amount. There are no negative consequences to reactor 

safety from such an incident.

8.3 Sudden Loss of Electrical Power

When electrical power fails, the refrigerator will cease 

operation, and will come to a safe condition without 

operator intervention of any kind, even in the 

complete absence of power, as all automatic valves fail 

safe. The liquid in the moderator chamber will boil 

away slowly, and the hydrogen pressure will slowly 

rise to normal shutdown pressure as the vapor fills the 

ballast tank. Again, no operator intervention of any 

kind is required to reach a safe condition. If the power 

failure outage is site-wide, the reactor will also scram, 

so that heating will quickly decay to nominal values; 

if the reactor continues to operate, a rundown signal 

will be generated when the hydrogen pressure exceeds 

135 kPa, and reactor power, and therefore nuclear 

heating, will be reduced to a level safe for the cold 

source. If the power outage is prolonged, the system 

will remain in a safe condition indefinitely, and will 

require no operator intervention. All valves in the 

refrigerator are fail safe, and the hydrogen system is 

entirely passive. No damage to any reactor system 

would occur from such an incident.

8.4 Loss of Cooling Water Flow to Cryostat and 

Plug Assembly

The non-cryogenic components of the cryostat 

assembly are cooled by circulating D2O. The flow is 

monitored by two (for redundancy) flow sensors. If 

the cooling drops below 18.9 lpm, a value providing 

ample cooling at full reactor power, a rundown signal 
62

will be generated and the reactor power will be reduced 

to a safe level until the problem is corrected. There 

are two rundown channels, and the rundown logic is 

one-of-two, so that failure to run down would require 

two failures, and is highly improbable. If both circuits 

were to fail, there are additional alarms generated by 

the PLC that would alert the reactor operators before 

there is any danger to the structural integrity of the 

cryostat assembly. If the reactor continued to operate 

for an extended period in spite of this signal, the 

cryostat assembly could be damaged, but there would 

be no damage to the reactor or its systems.

8.5 Loss of Cryogenic Insulating Vacuum

The only reasonable scenarios leading to a total loss 

of insulating vacuum are: failure of a vacuum pump 

and of the protective circuit which would isolate the 

system; a sudden leak in the helium containment 

surrounding the vacuum space; or a failure of 

the moderator chamber itself. The last of these is 

analyzed in Section 8.6 for the worst case of a sudden 

leak, and presents no hazard. For either of the other 

two possible scenarios, helium would leak into the 

vacuum, spoiling the cryogenic insulation. The heat 

load would exceed the capacity of the refrigerator, 

and the liquid would all vaporize (relatively slowly). 

The reactor would then be reduced in power by the 

rundown signal generated by high hydrogen pressure, 

and the incident would be over. No damage to any 

reactor component or system would occur.

8.6 Sudden Release of Liquid Hydrogen to 

Vacuum Space

Although the normal operating condition of the cold 

source with liquid in the moderator chamber is at 

approximately 100 kPa, and the moderator chamber 

is therefore at low stress (design working stress 500 

kPa), a sudden large leak that releases all of the 

liquid to the vacuum space has been analyzed. The 

liquid hydrogen in the moderator chamber during 

operation is under its saturated vapor pressure, and 

is approximately 5 liters at 20.27 K and 101.3 kPa 

(1 bar), which corresponds to 354 g (see table 8.118). 

The volume contained by the helium jacket, which is 

the pressure containing barrier, is 47 liters, neglecting 

all connecting tubing.

In the event of a sudden leak in the moderator, the 

354 g of liquid would first cool down as a result of 

the expansion into a larger volume, but would then 

increase rapidly as a result of vaporization. Some of 

the excess pressure would be vented to the ballast tank 

and other components outside the cryostat. For a 

conservative bounding estimate, however, this venting 

is neglected, and it is assumed that all is converted 

instantaneously to vapor at 28 K (lowest temperature 

at which entire inventory will be vapor). From Table 

8.1, the vapor density at 28 K is 7.53 g/liter, so that 

the 354 g will fill the available volume of 47 liters 

at just over 600 kPa (87 psi). This is well below 

the design working pressure, let alone the rupture 

pressure (>8270 kPa) of the helium jacket. Therefore, 

there would be no damage to any reactor component, 

nor would there be any release of hydrogen to the 

confinement building. Note that venting will occur, 

and the actual pressures would be much less than 600 

kPa in an actual incident.

8.7 Slow leak of Air into Vacuum Vessel

The design of the NIST liquid hydrogen cold neutron 

source is focused on prevention of mixing of hydrogen 

and air. To this end, all hydrogen containing structures 

are completely surrounded by inert gas (usually 

helium) contained in an additional containment 

vessel. These containments are filled to a pressure of 

approximately 125 kPa with helium that is analyzed 

for oxygen impurities immediately before filling, so 

that there is no possibility of air leaking into the space 

immediately surrounding the hydrogen moderator 

chamber. The pressure in each containment volume is 

continuously monitored, and an alarm is annunciated 

when the pressure reaches 100 kPa. In spite of these 

precautions, the consequences of a slow leak from the 

Table 8.1 Selected cryogenic properties of hydrogen.

Temperature Form Pressure 

(kPa)

Density 

(kg/m3)

Enthalpy 

(kJ/kg)

20.27 Liquid 101.3 70.786 -256.2

20.27 Vapor 101.3 1.3380 189.3

28 Vapor 600. 7.53 189.3.
63

containment into the vacuum surrounding the cold 

moderator chamber is calculated, assuming that the 

containment is filled with 100 % air. For this to happen, 

a helium supply bottle would have to be filled with air, 

and not have been analyzed, in direct contradiction 

of all procedures. It is further assumed that the leak 

is not noted until the pressure drops to the alarm 

point of 100 kPa. The volume of the containment 

surrounding the cold zone of the moderator chamber 

is approximately 14 liters; therefore, this would imply 

a total volume of 14x25/125 = 2.8 STP liters (3.3 g) 

of air that would be cryo-trapped on the moderator 

chamber. The oxygen in this air would be converted to 

ozone by the radiation field, which can spontaneously 

recombine with reasonably large energy release. 

Assuming that this energy both ruptures the vessel 

and ignites a detonation, the magnitude of the peak 

pressure can be estimated. Measurements have been 

made for the more challenging conditions of 300 cc of 

solid air (528 g) combining with liquid hydrogen in a 

closed vessel19 of total volume 32.6 liters. In this case, 

the peak pressure reached only 1050 psi, a value that 

would be greatly reduced for the actual conditions of 

a 47 liters volume and 3.3 g of air. These values imply 

a pressure less than 883 kPa, which is much less than 

the allowed working pressure for the helium jacket. 

Therefore, there would be no safety consequences, 

either from damage to the reactor thimble or from 

any release to the confinement building from such a 

hypothetical accident.

8.8 Hydrogen Release to Confinement Building

The cold source system is designed to prevent such an 

occurrence during normal operation, both by passive 

features and by procedures. All hydrogen containing 

systems are protected from accidental breakage by being 

buried in shielding, in the floor, or being surrounded 

by steel protective structures. The 133 kN (15 ton) 

annular crane is prevented from entering the zone 

where hydrogen system components are contained 

by electrical interlocks, that can only be bypassed by 

reactor management (reactor supervisor or above). 

During reactor operation, additional personnel would 

be detailed to observe and control any crane use in this 

area, to ensure that no shields that protect hydrogen 

lines were removed. During maintenance procedures, 

the ballast tank is isolated from the moderator, so that 

only the hydrogen present in the moderator chamber, 

heat exchanger and piping could be released. The 

volume of the moderator, heat exchanger and piping 

were measured for the first source, and determined to 

be a total of 47 liters; the volume of the moderator 

chamber alone was 17 liters. For the second generation 

source, the moderator chamber is just 5 liters, so that 

37 liters could be available for release when the ballast 

tank is isolated. Maintenance procedures require 

emptying this volume by use of a hydrogen absorbing 

system, and backfilling with helium to obtain a nonflammable mixture before opening or working on the 

hydrogen system.

Although any release would require multiple system 

and procedural failures, the consequences of a release 

have been analyzed. The design allows a warm pressure 

of 500 kPa, which would allow release of approximately 

13 g (the mass release required to reduce the Unit 2 (17 

g for Unit 1) pressure to barospheric). It is extremely 

difficult to predict actual gas behavior during any 

accidental release, but it is useful to look at some limit 

cases. The volume of the first floor of the confinement 

building, which is the area to which hydrogen would 

be released, is approximately 4000 m3, and a complete 

release as described would add 0.15 m3 of hydrogen. 

When completely mixed, this is considerably below 

the flammable limit of 4 % by volume. Thus, the 

possibility of deflagration or detonation exists only 

during that period when the gas is being released, and 

is mixing with the air. The ventilation system supplies 

conditioned air from a supply duct on the wall behind 

the area of hydrogen containing equipment at a 

measured rate of approximately 240 m3 per minute, 

divided between three equal registers, one of which 

discharges directly into the area containing hydrogen. 

For estimation purposes, the local volume for hydrogen 

release and mixing can be specified as the area above 

the guide shielding, a volume of approximately 

200 m3, which connects directly with no obstructions 

to the rest of C-100. Further progress in estimating 

consequences requires specification of the nature of 

the release. For a “slow” release of less than 1 l/s for 

140 s or more, there is little or no danger of a large 

energy release. The quantity of hydrogen-air mixture 

within flammable limits at any time will be small. (A 

recent simulation for a garage of dimensions similar 
64

to the hydrogen release volume as part of a study of 

hydrogen use for transportation with a much faster 

release rate indicates little risk. This simulation did 

not incorporate active ventilation.) 20. The hydrogen 

will rise into the local volume and rapidly be mixed 

to well below flammable limits by the action of 

convection and the forced ventilation. Any reaction 

would be a slow deflagration, which could generate 

hot flames, but no destructive shock waves.

A rapid release, however, that might occur if a 

guillotine break occurred as a result of misuse of a 

crane, tearing the heat exchanger from the wall, 

would have more serious effects. In this case, the 

release would proceed at sonic velocities, so that the 

entire 13 g was released in less than 1-2 s. Such a 

release would create a well mixed volume as part of 

the release – less than the full 13 g, but certainly 

a substantial fraction of 13 g. In order to estimate 

the consequences of such a release, the equivalent 

weight of TNT must be determined. This is not as 

straightforward as simply converting at the usual 

rate of 24 g TNT per g hydrogen, since allowance 

must be made for the details of mixing (not all of 

the gas will be at stoichiometric conditions), and for 

reflections of the blast from adjacent walls, ceilings, 

shields, and reactor shielding. Given the arrangement 

at the assumed point of release (near the reactor face 

over the guide shields), a reasonable assumption is 

that we should double the charge to account for 

reflection from the ceiling, while halving the charge 

to account for the actual mixture conditions. The net 

result is that the effective TNT weight, W, is taken 

as 0.3 kg for use in the scaling for hydrogen blast 

effects. Although a detonation is highly unlikely (no 

detonations are known for unconfined spills of less 

than 100 kg of hydrogen, and this assumed spill is 

only weakly confined), that is what is assumed by 

the use of TNT equivalent, and so the assumption 

is highly conservative. The distance, Z, of the release 

and detonation from the confinement building walls 

is approximately 10 m, so that the scaled distance (Z/

W1/3) is 15 m/kg1/3. Using standard blast data21, this 

gives rise to a peak reflected overpressure of 6000 Pa 

(0.7 psi) and a peak reflected impulse of 25.4 Pa-s at 

the wall. These values are well below any possibility 

of serious structural damage, and will result in no 

harm to the reactor or confinement. However, the 

consequences for equipment and personnel closer 

to such a blast would be very serious. Once again, 

all of the design and procedures established for the 

maintenance and operation of the source are explicitly 

for the purpose of preventing such an event, and two 

or more independent failures are required to even 

make it possible.

8.9 Stoichiometric Mixture of Hydrogen and Air 

in Moderator Chamber

Although the entire design philosophy is to prevent 

air from mixing with the hydrogen in the moderator, 

and despite the absence of any scenario to accomplish 

the hypothetical accident, it is assumed that a 

stoichiometric mixture of hydrogen and air is present 

in the moderator chamber at a pressure of 100 kPa. If 

a normal deflagration occurs, the peak pressure would 

be 827 kPa (120 psi), which is much less than the 

calculated 2.6 MPa (375 psi) or measured breaking 

pressure 3.3 MPa (475 psi) for the moderator chamber. 

The temperature of the gas produced after the event 

would be approximately 3000 K, which would be 

given up to the aluminum vessel. Since the relative 

mass of the hot gas to aluminum is of order 0.4 %, 

the temperature of the vessel will not increase by more 

than 8 K, which will not affect the properties of the 

vessel. Since everything will be contained within the 

moderator, there will be no effect on the reactor or the 

confinement building.

8.10 Maximum Hypothetical Accident

The maximum hypothetical accident (MHA) for 

the NBSR liquid hydrogen source begins as a 

sudden rupture of the vacuum line to the cryostat 

simultaneously with its helium jacket. This in itself 

has no serious consequences, but the MHA goes on to 

postulate that this rupture is followed by a rupture in 

the moderator shell, followed by complete mixing of 

the solid air cryo-pumped on the moderator chamber 

shell, followed by a strong ignition source. This scenario 

is not credible, so that the consequences of this analysis 

bound all possible accidents of this type. At each stage 

of the analysis, the extraordinarily conservative nature 

of the assumptions will be elucidated.
65

The first stage of the scenario, a rupture of the vacuum 

line and its helium shield line, is improbable, but not 

incredible. The only identified scenario assumes that 

the 133 kN (15 ton) crane (or a load on it) snags 

the module containing the vacuum pumps, pulling 

it over, and ripping the vacuum line. However, the 

crane is interlocked to prevent operation in the area 

of the vacuum module, so that this assumes a failure 

of the interlock. Further, it is unlikely that simply 

tipping the module over would rupture both lines 

– either no damage to the vacuum line or a slow leak 

of helium is more probable. However, the assumption 

is that there is a “guillotine” break in the vacuum line, 

which admits air into the vacuum space around the 

moderator chamber. The source is assumed to be full 

of liquid hydrogen (a volume of 5 liters), with the 

refrigerator operating. Under these conditions, air 

will rush into the vacuum space, and will be cryopumped onto the surface of the moderator, which will 

be at 21 K, well below the solidification temperature 

of both nitrogen and oxygen. At this point, the gross 

vacuum alarm on the hydrogen system will cause 

an interruption in the flow of cold helium into the 

condenser, and the liquid hydrogen will begin to boil 

away. It is conservative to assume that the reactor is 

shut down, as this will reduce the heat load, leaving 

more hydrogen to be boiled off by removing the heat 

evolved in the condensation and solidification of air.

As the room temperature air enters the moderator 

chamber vacuum space, part of it will be cooled down 

by contact with the moderator chamber surface, and 

solidified. The heat evolved in each of these steps is 

computed from the data in Table 8.222,23.

This heat must be removed from the system by boiling 

of liquid hydrogen, which requires an average heat 

input between 100 and 400 kPa of approximately 419 

J/g. Thus, for every gram of air that is cryopumped 

onto the moderator chamber surface, there will be 

555/419 = 1.32 g of hydrogen boiled off from the 

moderator. The limiting case of hydrogen available 

comes with the reactor off, so that very little vapor 

is being generated as a result of nuclear heating. 

This condition also maximizes the amount of air 

cryopumped, again because the heat load is minimal. 

The volume of the moderator chamber is 5 liters, 

and at 100 kPa pressure, this corresponds to 354 g 

of liquid hydrogen. This liquid will all be boiled 

off by condensation of 354/1.32 = 268 g of air, or 

approximately 240 cm3. As soon as the vacuum 

ruptures, the refrigerator will automatically stop 

providing cooling to the condenser, and the liquid 

hydrogen will decrease in volume in the moderator 

chamber until it is all gone. Note that the vacuum line 

surrounds the supply and return lines for hydrogen 

between condenser and moderator chamber, so that 

no new hydrogen will flow into the chamber after the 

rupture. In fact, it is probable that most of the liquid 

would be immediately blown from the moderator 

chamber by the sudden increase in heat load; 

however, this possibility is ignored for conservatism. 

Clearly, the amount of hydrogen is considerably 

above stoichiometry at the beginning of the incident, 

and the amount of oxygen available for a reaction 

will be maximum when all of the liquid is gone, at 

which point there will be 25 g of cold hydrogen vapor 

available. The amount of oxygen available at this time 

(from above estimate) will be 268*0.23 = 62 g, which 

can react with 7.8 g of hydrogen.

The MHA now assumes that the moderator vessel 

now ruptures (with no reasonable scenario), and that 

a large energy ignition source then fires (again with no 

credible scenario). The results of Ward19 et al showed 

that reaction of 300 cm3 of solid oxygen with 21 g of 

hydrogen in a closed vessel of 33 liters resulted in a 

maximum pressure of 6.8 MPa. The present vacuum 

Table 8.2 Selected properties of nitrogen and oxygen.

Property N2 O2 Air(J/g)

Volume Fraction Used 0.79 0.21 1.0

Mass fractions .77 .23 1.0

Enthalpy (300 K, 100 kPa) 272.71 (J/g) 311.20 (J/g) 282.6 (J/g)

Enthalpy (MP1, 100kPa) -193.53 -150.78 -183.7

Heat of Fusion 26.0 13.9 23.2

Solid cooling to 25 K 

(including phase transitions) 67.5 60.1 65.8

Total Heat Evolved J/g 559.7 536.0 555.3

1

 MP=Melting Point; Air 59.8 K (liquidus), N2 63.16 K, O2 54.37 K •
66

vessel has a volume of 47 l, and Ward et al showed 

that yield was inversely proportional to volume, 

implying a maximum pressure of 4.8 MPa, which is 

significantly below the calculate breaking pressure of 

> 8270 kPa (the test vessel did not break at 8960 kPa). 

Therefore, the MHA will be entirely contained within 

the vacuum vessel, and will cause no damage to the 

reactor or confinement building.
9.0 Acknowledgements

The authors acknowledge the help and assistance of many different people at the NCNR over the past 10-15 

years, in the design, fabrication, installation, operation and maintenance of the hydrogen sources. In particular, 

we thank Jim Moyer and Lew Robeson for their engineering design and drafting; Don Pierce for leading the 

installation of Unit 2; George Baltic and his technical team for their work on installation; Bob Williams 

(Big Bob) for his removal and installation work on the first source; George Baltic and Scott Slifer for their 

exceptional welding work; the reactor electronics staff for their work on the interface with reactor systems; Don 

Fravel, Tom Green and Dennis Nester for their operation and maintenance of the source; Jeremy Cook for the 

TOF analysis; the reactor operations staff for their help in many areas, and the Health Physics group, especially 

Les Slaback for helping us to do it safely. Thanks also to Jim Siewarth and Doug Olson for creating the mockup 

and conducting the thermal hydraulic tests in Boulder. We have also benefited from many discussions with 

almost everyone in the organization.
9.1 References

1. S. J. Cocking and F. J. Webb, Section 4 in Thermal Neutron Scattering, edited by P. A. Egelstaff, Academic 

Press, London and New York (1965); F. J. Webb, Reactor Science and Technology (Journal of Nuclear 

Energy Parts A/B), 187-215 (1963).

2. National Bureau of Standards, FSAR on the National Bureau of Standards Reactor, NBSR-9, U. S. 

Department of Commerce, (1966).

3. Robert S. Carter, Safety Analysis Report on the D2O Cold Neutron Source for the National Bureau of 

Standards, NBSR-13, National Bureau of Standards, U. S. Department of Commerce, (1984).

4. H. Hoffman, “Experimental Investigations of Heat Transfer Behavior of a Horizontally Arranged 

Cylindrical Cold Neutron Source in a Vaporizing Deuterium Flow,” Proc. Cryo. Eng. Conf., 390-303, 

Helsinki (1984); H. Hoffman and J. M. Astruc, “Experimental Investigations of the Fluid Dynamics 

of a Thermal Siphon with Vaporizing Deuterium for a Horizontally Arranged Cylindrical Cold Source,” 

462-465, ibid; “Natural Convection Cooling of a Cold neutron Source with Vaporizing Deuterium at 

Temperatures of 25 K”, in Natural Convection: Fundamentals and Applications, S. Kakac, W. Aung, R. 

Viskanta editors, Hemisphere Publishing Corp., New York, 1140-51 (1985).

5. J. D. Siegwarth, D. A. Olson, M. A. Lewis, J. M. Rowe, R. E. Williams and P. Kopetka, “Thermal 

Hydraulics Tests of a Liquid Hydrogen Cold Neutron Source,” National Institute of Standards and 

Technology, NISTIR 5026 (1994).

6. Paul Ageron, “Cold Neutron Sources at ILL,” Nuc. Inst. Meth., A284, 197-199 (1989); P. Ageron, 

“Neutronic Design of the ILL Cold Sources – An Historical Perspective” in Proceedings of the 

International Workshop on Cold Neutron Sources, March 5-8, 1990, Los Alamos, N. M., LA-12146-

C, 1-19 (1991).

7. T. E. Fessler and J. W. Blue, “Radiation-Induced Conversion of Liquid Hydrogen,” Phys. Rev. Let. 14, 

811 (1965).

8. Nelms, L. W. and Carter, H. G., “NERVA Irradiation Program: GTR Test 21, vol. 2, Nuclear RadiationInduced Conversion of Para Hydrogen to Ortho Hydrogen,” U. S. Atomic Energy Commission, FZK351-2 (1968).

9. J. F. Briesmeister, ed., MCNP – A General Monte Carlo N-Particle Transport Code, Version 4B, Los Alamos 

National Laboratory, LA-12625-M, Los Alamos, New Mexico (March 1997).

10. NBSR-14: Safety Analysis Report (SAR) for License Renewal of the National Institute of Standards and 

Technology Reactor – NBSR, NIST Internal Report NISTIR 7102 (2004).

11. Croff, A. G., “A User Manual for the ORIGEN2 Computer Code,” ORNL/TM-7175, Oak Ridge 

National Laboratory (July 1980).

12. Trellue, H.R., “Development of MONTEBURNS: A Code that Links MCNP and ORIGEN in an 

Automated Fashion for Burnup Calculations,” LA-13514-T, Los Alamos National Laboratory (December 

1998).
13. Douglas E. Peplow, “A Computational Model of the High Flux Isotope Reactor for the Calculation 

of Cold Source, Beam Tube, and Guide Hall Nuclear Parameters,” Oak Ridge National Laboratory, 

ORNLTM-2004/237 (2004).

14. B. Farnoux and P. Breant, “Upgrade of the Experimental Facilities of the Orphee Reactor,” Proceedings 

of the Third Meeting of the International Group on Research Reactors (IGORR-III), September 30-

October 1, 1993, Naka, Japan.

15. A.E. Dukler, Moye Wicks III and R. G. Cleveland, “Frictional Pressure Drop in Two-Phase Flow: A. 

A Comparison of Existing Correlations for Pressure Loss and Holdup and B. An Approach through 

Similarity Analysis,” AIChE J 10, 38-51 (1964).

16. L. S. Tong and Y. S. Tang, Boiling Heat Transfer and Two-Phase Flow, 2nd Edition, Taylor and Francis, 

Washington, DC, 188 (1997); A. E. Dukler and A. Taitel, Two-Phase Gas-Liquid Flow: A Short Course on 

Principles of Modeling Gas-Liquid Flow and on Modern Measuring Methods, Univ. of Houston, Houston, 

Texas (1991).

17. J. P. Iman, Heat Transfer, McGraw-Hill Inc. (1981).

18. R. D. McCarty, J. Hord and H. M. Roder, Selected Properties of Hydrogen, NBS Monograph 168, 

(1981).

19. D. L. Ward, D. G. Pearce and D. J. Merrett, “Liquid Hydrogen Explosions in Closed Vessels,” Adv. 

Cryo. Engineering, 9, 390-400 (1964).

20. W. Breitung, G. Necker, B. Kaup and A. Vesser, “Numerical Simulation of Hydrogen Release in a Private 

Garage”, Proceedings of Hypothesis IV, Strahlsund, Germany, 268 (2001).

21. W. E. Baker, P. A. Cox, P. S. Westine, J. J. Kulesz and R. A. Strehlow, Explosion Hazards and Evaluation, Vol. 

5 of Fundamental Studies in Engineering, Elsevier Scientific Publishing Company, New York (1983).

22. R. T. Jacobsen, S. G. Penoncello, and E. W. Lemmon, Thermodynamic Properties of Cryogenic Fluids, 

Plenum Press, New York and London, (1997).

23. R. B. Scott, Cryogenic Engineering, D. van Nostrand Co., Inc., Princeton, NJ (1959).
9.2 Appendix

unit2: advanced LH2 cold source for mcnp5, As built 24x32-cm 

c

c with vapor in exit hole 3.0 cm LH2

c Inner ellipsoid at 76.5, outer centered at 76 cm

c .09 in walls

c

c LH2 cold source ad00 (11/00):

701 4 -.063 201 -202 -210 vol=726 imp:n,p=54 $lh2

702 4 -.063 201 -202 -227 210 imp:n,p=54 $lh2

703 4 -.063 201 -202 -225 227 imp:n,p=54 $lh2

704 4 -.063 201 -202 225 203 imp:n,p=54 $lh2

705 6 -.035 -211 207 imp:n,p=27 $ lh2 separator

706 0 -208 imp:n,p=27

707 6 -.0013 -209 201 225 -202 vol=600 imp:n,p=27 $ vapor @ 1.1 bar

709 0 -218 207 #(-212) imp:n,p=27

710 0 -215 212 218 imp:n,p=27 

711 2 -2.7 -201 208 203 227 imp:n,p=27 $ Al wall, inner sphere

712 2 -2.7 -201 208 -227 imp:n,p=27 $ “ “ “

713 2 -2.7 -212 211 207 imp:n,p=27 $ Dome

714 2 -2.7 -203 209 -207 208 227 imp:n,p=27 $ Al support ring

715 2 -2.7 -207 202 -227 imp:n,p=27 $Al: Chamber

716 2 -2.7 -207 202 227 #760 203 imp:n,p=27 $Al: “

717 2 -2.7 -201 208 -209 225 imp:n,p=27 $ Al - exit hole

718 2 -2.7 -207 202 -209 225 imp:n,p=27 $ Al “

c

720 2 -2.7 -217 218 -225 #(-216) imp:n,p=27 $Al: He jacket

721 2 -2.7 -217 218 225 #(-216) #761 #763

 #751 #752 imp:n,p=27 $Al: “

722 2 -2.7 -216 215 218 #760 imp:n,p=27 $ remove He jackt section

723 3 -1.1 -205 204 -273 232 #(-242) #(-244) imp:n,p=9

724 3 -1.1 -227 -270 217 #(-216) #(-229) imp:n,p=9 $d2o: cryostat

725 3 -1.1 227 217 -273 -225 216 229 imp:n,p=9 $d2o: cryostat

726 3 -1.1 -273 -204 225 217 #751 #752 #(-229)

 #(-242) #(-244) #756 #757 #(-216) imp:n,p=9

727 3 -1.1 205 -273 -127 #741 #742 #743 #(-242) #(-244)

 #744 #745 #746 #747 imp:n,p=9 $source reflector,coolant

728 2 -2.7 227 273 -219 -127 #746 #747 imp:n,p=3 $Al: d2o jacket

729 2 -2.7 -227 -220 270 imp:n,p=3 $Al: “

730 2 -2.7 -205 204 -232 231 imp:n,p=9 $ ‘Al collar’

c

740 0 -205 204 -231 imp:n,p=27

741 0 205 -292 -295 -127 imp:n,p=27 $ void through D2O

742 0 205 -293 296 -127 imp:n,p=27

743 0 205 -298 297 -296 295 -127 imp:n,p=27

744 2 -2.7 -297 287 -127 205 -296 295 imp:n,p=9

745 2 -2.7 -288 298 -127 205 -296 295 imp:n,p=9

746 2 -2.7 -284 292 -127 205 -295 imp:n,p=9

747 2 -2.7 -285 293 -127 205 296 imp:n,p=9

748 3 -1.1 -228 217 imp:n,p=9

749 2 -2.7 -229 228 217 imp:n,p=9

c

751 0 -293 -204 218 299 213 imp:n,p=9

752 0 -292 -204 218 -299 213 imp:n,p=9

756 2 -2.7 -284 292 -204 217 -299 213 imp:n,p=9

757 2 -2.7 -285 293 -204 217 299 213 imp:n,p=9

c

760 0 -241 -127 225 215 imp:n,p=3

761 2 -2.7 -242 241 -127 225 216 imp:n,p=3

762 3 -1.1 -243 -127 225 229 imp:n,p=3

763 2 -2.7 -244 243 -127 225 229 imp:n,p=3

c

731 2 -2.7 -283 219 -127 227 #746 #747 imp:n,p=3 $Al: reactor vessel

732 2 -2.7 -227 -280 220 223 imp:n,p=3 $Al: “

733 3 -1.1 -222 280 -221 223 -224 imp:n=1 imp:p=1 $d2o: reflector

734 3 -1.1 283 227 -224 -127 imp:n=1 imp:p=1 $d2o: “

735 3 -1.1 -224 -227 280 222 imp:n=1 imp:p=1 $d2o: “

736 3 -1.1 223 221 -224 -222 imp:n=1 imp:p=1 $d2o

c

737 3 -1.1 -223 -224 8 imp:n=1 imp:p=1 $(ec=737)

738 3 -1.1 -127 8 -9 224 imp:n=1 imp:p=1 $(sc=738)

c

200 0 -1 2 -3 imp:n=27 imp:p=1 $ beam to DXTRAN

201 0 8 -2 -3 127 imp:n=27 imp:p=1

202 0 8 127 -10 3 -1 imp:n,p=27

998 0 -127 8 9 -10 imp:n,p 0

999 0 1: 10: -8 imp:n=0 imp:p=0 $ outside

c SURFACES:

1 1 py 200.

2 1 py 199.5

3 1 cy 3.

8 1 py 55.5

9 1 cy 30.

10 1 cy 40.

127 1 cz 150.

c these are the CNS surfaces for Unit 2:

 201 1 sq 2.0194 1 1 0 2r -182.25 0 76.5 0

 202 1 sq 1.7778 1 1 0 2r -256 0 76 0

 204 1 py 96.0

 203 1 sq 1.7331 0 1 0 2r -108.16 0 0 0 $ .2 cm thick support ring

 205 1 py 97.0

 207 1 sq 1.7657 1 1 0 2r -261.3072 0 76 0 $ .065-in outer wall
 208 1 sq 2.0519 1 1 0 2r -175.5625 0 76.5 0 $ .090-in inner wall

 209 1 sq 1.7778 0 1 0 2r -100 0 0 0 $ exit hole

 210 1 py 65.0

 211 1 s 0 76 16 5.0

 212 1 s 0 76 16 5.25

 213 1 py 82.

 215 1 s 0 76 16 6.5

 216 1 s 0 76 16 7.5

 217 1 sq 1.6402 1 1 0 2r -333.0625 0 76 0

 218 1 sq 1.7101 1 1 0 2r -289.0000 0 76 0

 219 1 cy 28.0

 220 1 sq 1.000 4.500 1.000 0.0 0.0 0.0 -784.00 0. 69.7 0.

 221 1 sq 1. 5.3389 1. 0. 0. 0. -930.25 0. 67.2 0.

 222 1 py 67.2

 223 1 py 55.52

 224 1 cy 29.98

 227 1 py 69.7

 225 1 py 76.

 228 1 s 0 76 -16 6.2

 229 1 s 0 76 -16 7.2

 231 1 cy 11.5

 232 1 cy 13.

 241 1 c/y 0 19.7 2.8

 242 1 c/y 0 19.7 3.3

 243 1 c/y 0 -19.7 2.8

 244 1 c/y 0 -19.7 3.3

 270 1 sq 1. 4.7452 1. 0. 0. 0. -748.5696 0. 69.7 0.

 273 1 cy 27.36 

 280 1 sq 1. 4.08 1. 0. 0. 0. -858.49 0. 69.7 0.

 283 1 cy 29.3

 284 5 cy 13.

 285 6 cy 13.

 287 1 pz -13.

 288 1 pz 13.

 292 5 cy 11.

 293 6 cy 11.

 295 5 px 0.

 296 6 px 0.

 297 1 pz -11.

 298 1 pz 11.

 299 1 px 0.

m2 13027.62c 1 $Al

m3 1002.66c .66334 1001.62c .00333 8016.62c .33333 $99.5% pure d2o

mt3 hwtr.60t lwtr.60t

c

c MODIFIED ORTHO-PARA MIXING --- ORIGINAL METHOD:

c 1004.62c and h4para.61t are copies of 1001.62c and hpara.61t

c with ZAID = 1004, needed for mixing ortho and para LH2.

c

m4 1001.62c .5 1004.62c .5

mt4 hortho.61t h4para.61t $ from sab2002

m6 1001.62c .5 1004.62c .5

mt6 hortho.61t h4para.61t $ from sab2002

c

c Change angles to 16.5 degrees for CTE,CTW:

*tr5 -18.83 94.66 0 27.75 62.25 90 117.75 27.75 90 90 90 0

*tr6 -18.83 94.66 0 -5.25 95.25 90 84.75 -5.25 90 90 90 0

*tr1 0 0 0 11.25 78.75 90 101.25 11.25 90 90 90 0

c

c Current Tallies:

fq0 e c

c

c tallies from the kcode calc:

f14:n 701 707

fm14 1.528e18

e14 5-9

fc14 Cold neutron flux in LH2 and exit hole.

c

f61:n 204

fs61 -231

e61 5e-9 

c61 .99 1.0

sd161 415.5 1

fm61 1.525e18

c

f161:n 205

fs161 -209

sd161 235.62 1 $ area

e161 5e-9 

c161 .99 1.0

fm161 1.525e18

c tallies inside dxtran sphere:

f11:n 2

fs11 -3

sd11 28.27 1

fm11 1.525e18

e11 1e-9 18i 2e-8

c11 .9992 .9996 1.0 t

tf11 3j 1 j 3 1 j

c

f21:n 2

fs21 -3

sd21 28.27 1

fm21 1.525e18
e21 5e-9 18i 1e-7 1e-6 1-5 1-4 1-3 1-2 .1 1 20 t

c21 .9992 .9996 1.0 t

c

f31:n 2

fs31 -3

sd31 28.27 1

fm31 1.525e18

e31 1e-9 18i 2e-8

c31 .9996 .9998 1.0 t

c

f41:n 2

fs41 -3

sd41 28.27 1

fm41 1.525e18

e41 5e-9 18i 1-7 3-7 1-6 3-6 1-5 3-5 1-4 3-4

 1-3 3-3 1-2 3-2 .1 .3 1 3 10 20 t

c41 .9996 .9998 1.0 t

c

c

f151:n 2

fs151 -3

sd151 28.27 1

fm151 1.525e18

fc151 energy bins for wavelengths 20.5,19.5,18.5, .. 1.5, and 0.5 A

e151 .1933e-9 .2136-9 .2373-9 .2652-9 .2984-9

 .3381-9 .3863-9 .4457-9 .5199-9 .6142-9

 .7368-9 .9001-9 1.1243-9 1.4441-9 1.9226-9 2.6853-9

 4.0114-9 6.631-9 12.997-9 36.102-9 324.92-9

c151 .9996 .9998 1.0 t

c

c PROB

mode n

prdmp 2j 1 1

tmp 1.7235e-9 16r 2.77e-8 41r

c

c ssw 8 (737) -9 (739) $ write surf source card from kcode prob

c

ssr old 8 9 new 8 9 wgt 1. psc 1.

dxt:n -39.02 196.16 0 3.1 3.1

c