NASA/TP—2010–216435

August 2010

Design of Refractory Metal Life Test 

Heat Pipe and Calorimeter

J.J. Martin, R.S. Reid, and S.M. Bragg-Sitton

Marshall Space Flight Center, Marshall Space Flight Center, Alabama
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i

NASA/TP—2010–216435

Design of Refractory Metal Life Test 

Heat Pipe and Calorimeter

J.J. Martin, R.S. Reid, and S.M. Bragg-Sitton

Marshall Space Flight Center, Marshall Space Flight Center, Alabama

August 2010

National Aeronautics and

Space Administration

Marshall Space Flight Center • MSFC, Alabama 35812
ii

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443–757–5802

This report is also available in electronic form at

<https://www2.sti.nasa.gov>
iii

TABLE OF CONTENTS

1. INTRODUCTION .......................................................................................................................... 1

2. ACCELERATED LIFE TEST HEAT PIPE APPROACH ............................................................. 2

2.1 Proposed Test Matrix .............................................................................................................. 3

3. HEAT PIPE DESCRIPTION ......................................................................................................... 6

3.1 Heat Pipe Sizing and Performance ......................................................................................... 6

3.2 Heat Pipe Engineering Drawings ............................................................................................ 12

3.3 Heat Pipe Procurement Specifications .................................................................................... 15

4. CALORIMETER UNIT ................................................................................................................. 16

4.1 Water Calorimeter Description ............................................................................................... 17

4.2 Heat Pipe/Calorimeter Performance ....................................................................................... 20

4.3 The Generalized Fluid System Simulation Program Model ................................................... 34

5. OTHER CONSIDERATIONS ....................................................................................................... 36

5.1 Sodium Working Fluid Sampling (Vanadium Wire Technique) ............................................. 36

5.2 Heat Pipe Closeout and Support Concept ............................................................................... 38

5.3 Heat Pipe Radio Frequency Inductive Heating Coil ............................................................... 40

5.4 Chamber Gas Selection and Purity ......................................................................................... 41

5.5 Heat Pipe Temperature Measurement ..................................................................................... 43

5.6 Water-Cooling System ............................................................................................................ 44

6. SUMMARY ................................................................................................................................... 46

APPENDIX A—HEAT PIPE LIFE TESTING—ALTERNATE APPROACH .................................. 47

A.1 Corrosion in High-Temperature Heat Pipes ........................................................................... 47

A.2 Elementary Scaling Variables ................................................................................................ 47

A.3 Possible Life Test Configuration ............................................................................................ 48

A.4 Life Test Methods .................................................................................................................. 49

A.5 Baseline Heat Pipe Life Test .................................................................................................. 49

A.6 Alternative Heat Pipe Life Test: Option A ............................................................................. 50
iv

TABLE OF CONTENTS (Continued)

APPENDIX B—HEAT PIPE WICK PROCUREMENT SPECIFICATION ..................................... 54

B.1 MSFC IFMP Procurement Specification for Heat Pipe Wick Assembly Fabrication ............ 54

B.2 Wick Assembly Fabrication Specifications ............................................................................ 54

B.3 General Cleaning Procedure for Mo and Mo-Re Parts .......................................................... 58

B.4 Time-Temperature Relation for Mo-5%Re Material .............................................................. 58

B.5 Sample Process Flow for Heat Pipe Wick Fabrication Simplified Process Flow .................. 59

APPENDIX C—HEAT PIPE MATERIAL PROCUREMENT SPECIFICATION ............................ 61

C.1 IFMP Procurement Specification for Mo-Based Refractory Material ................................... 61

C.2 Material Specification ............................................................................................................ 62

C.3 Material Size and Quantity Specifications ............................................................................. 65

C.4 General Cleaning Procedure for Mo and Mo-Re Parts .......................................................... 66

APPENDIX D—HEAT PIPE FABRICATION/ASSEMBLY 

PROCUREMENT SPECIFICATION ...................................................................... 67

D.1 IFMP Procurement Specification for Fabrication of Mo-Alloy Heat Pipe Units ................... 67

D.2 Heat Pipe Layout Specification .............................................................................................. 67

D.3 General Cleaning Procedure for Mo and Mo-Re Parts .......................................................... 73

D.4 Time-Temperature Relation for Mo-5%Re Material .............................................................. 73

D.5 Sample Process Flow for Heat Pipe Fabrication/Assembly .................................................... 74

APPENDIX E—MATERIAL PROPERTIES USED IN ANALYSIS ................................................ 76

APPENDIX F—POWER EXTRACTION PERFORMANCE WORKSHEET ................................. 87

APPENDIX G—HEAT PIPE/CALORIMETER ENVIRONMENTAL LOSSES ............................. 91

APPENDIX H—THE GENERALIZED FLUID SYSTEM SIMULATION PROGRAM ................. 93

REFERENCES ................................................................................................................................... 95
v

LIST OF FIGURES

 1. Life test matrix assuming a maximum 1-μm nucleation site radius 

for F- and G-test series ....................................................................................................... 5

 2. Proposed life test matrix compared to operating limits at two-nucleation site radii .......... 8

 3. Not-to-exceed time-temperature relation for Mo-5%Re material ...................................... 9

 4. Wick assembly engineering drawing ................................................................................. 13

 5. Getter pack design and assembly engineering drawing ..................................................... 14

 6. Heat pipe envelope and assembly engineering drawing .................................................... 14

 7. Calorimeter conceptual design ........................................................................................... 18

 8. Conceptual heat pipe/calorimeter mounting setup ............................................................. 18

 9. Initial calorimeter concept engineering design layout ....................................................... 19

10. Initial calorimeter concept engineering design parts detail ................................................ 20

11. Water boiling point as a function of pressure ..................................................................... 21

12. Mo-44%Re heat pipe diameter growth as a function of temperature ................................ 22

13. Static gap and calorimeter layout ....................................................................................... 23

14. Sensitivity of the static conduction gap with gas composition .......................................... 24

15. Sensitivity of the static gas gap with calorimeter wall temperature ................................... 25

16. Calorimeter power and pressure drop variation with flow rate .......................................... 27

17. Calorimeter temperatures and film annulus sensitivities (∆T=20 K) ................................. 29

18. Calorimeter temperatures and film annulus sensitivities (∆T=60 K) ................................. 29

19. Heat pipe/calorimeter environmental heat loss .................................................................. 31

20. GFSSP model of heat pipe test setup ................................................................................. 35
vi

LIST OF FIGURES (Continued)

21. Concept to measure working fluid oxygen concentration in Na-filled heat pipes ............. 37

22. Estimated time required for equilibrium oxygen concentrations in static 

Na immersed for two V wire diameters ............................................................................. 37

23. Method 1 to close out the heat pipe fill stem and provide mounting support .................... 38

24. Method 2 to close out the heat pipe fill stem and provide mounting support .................... 39

25. Induction coil cluster mounted on five-position bus bar (end view) .................................. 40

26. Oxygen concentration in the heat pipe test chambers after successive dilutions 

of (a) UHP He and (b) UHP Ar fill gas ............................................................................... 41

27. Initial concept for test chamber gas mixture system .......................................................... 42

28. Final oxygen concentration in the heat pipe test chamber with (a) UHP He fill gas

and (b) UHP Ar fill gas purified using an inline SAES MicroTorr ambient gas purifier ... 43

29. Initial concept for water-cooling system layout ................................................................. 45

30. Baseline life test matrix assuming a maximum 1-μm nucleation site radius ..................... 50

31. Baseline life test wire cloth use. Symbols P, N, and I represent performance, 

baseline life, and irradiation tests, respectively .................................................................. 51

32. Option A life test wire cloth use. Symbols P, F, G, and I represent performance, 

Fisher, G68–80, and irradiation tests, respectively ............................................................ 51

33. Option A life test design compared to boiling limits at two nucleation site radii .............. 53

34. Option A life test matrix assuming a maximum 1-μm nucleation site radius .................... 53

35. Wick assembly drawing ..................................................................................................... 56

36. Heat pipe unit layout .......................................................................................................... 68

37. Heat pipe getter pack .......................................................................................................... 69

38. Helium density at 76 torr (1.5 psia) .................................................................................... 77

39. Helium kinematic viscosity at 76 torr (1.5 psia) ................................................................ 77
vii

LIST OF FIGURES (Continued)

40. Helium thermal conductivity at 76 torr (1.5 psia) .............................................................. 78

41. 316 stainless steel thermal conductivity ............................................................................. 79

42. 316 stainless steel specific heat .......................................................................................... 79

43. 316 stainless steel density .................................................................................................. 80

44. 316 stainless steel CTE ....................................................................................................... 80

45. 316 stainless steel emissivity .............................................................................................. 81

46. Mo-44%Re thermal conductivity ....................................................................................... 82

47. Mo-44%Re emissivity ........................................................................................................ 82

48. Mo-44%Re CTE ................................................................................................................. 83

49. Water density at 0.448 MPa (65 psia) ................................................................................ 84

50. Water thermal conductivity at 0.448 MPa (65 psia) ........................................................... 84

51. Water absolute viscosity at 0.448 MPa (65 psia) ............................................................... 85

52. Water specific heat at 0.448 MPa (65 psia) ........................................................................ 85

53. Water boiling point ............................................................................................................. 86

54. Steady-state results from the Fastrac turbopump GFSSP model ....................................... 94

55. Representative transient pressure results from the Fastrac turbopump GFSSP model ...... 94
viii

LIST OF TABLES

 1. Option A life tests, Fisher series (F-series) ........................................................................ 4

 2. Option A life tests, G68–80 series (G-series) ..................................................................... 4

 3. Life heat pipe design and conditions (neutral gravity) ....................................................... 7

 4. Impurity chemical composition (acceptable) final heat Mo-Re alloy 

material (powder metallurgy) ............................................................................................. 10

 5. Time-temperature relationship for Mo-5%Re material ...................................................... 12

 6. Calorimeter parameters for a film gap width of 0.010 in ................................................... 30

 7. Calorimeter parameters for a film gap width of 0.015 in ................................................... 30

 8. Calorimeter parameters for a film gap width of 0.020 in ................................................... 31

 9. Heat pipe and calorimeter thermal loss to environment for pure He (70 torr) ................... 33

10. Heat pipe and calorimeter thermal loss to environment for He-32%Ar (70 torr) .............. 33

11. Initial life test heat pipe arrangement ................................................................................. 40

12. SAES MonoTorr Phase II 3000 performance guarantee for rare gases ............................. 43

13. Baseline life tests ................................................................................................................ 50

14. Option A life tests Fisher series ......................................................................................... 52

15. Option A life tests G68–80 series ...................................................................................... 52

16. Impurity chemical composition (acceptable) final wick assembly .................................... 57

17. Mo-5%Re cumulative time at temperature ........................................................................ 59

18. Impurity chemical composition (acceptable) final heat Mo-Re alloy 

material (PM) ..................................................................................................................... 64

19. Impurity chemical composition (acceptable) wick assembly ............................................ 70
ix

LIST OF TABLES (Continued)

20. Impurity chemical composition (acceptable) final heat Mo-Re alloy 

material (PM) ..................................................................................................................... 71

21. Mo-5%Re cumulative time at temperature ........................................................................ 74

22. Pure He properties at 76 torr .............................................................................................. 76

23. 316 stainless steel properties .............................................................................................. 78

24. Mo-44%Re properties ........................................................................................................ 81

25. Water properties at 0.448 MPa (65 psia) ............................................................................ 83

26. Calorimeter geometry and performance for a cooling film annulus = 0.010 in ................. 88

27. Calorimeter geometry and performance for a cooling film annulus = 0.015 in ................. 89

28. Calorimeter geometry and performance for a cooling film annulus = 0.020 in ................. 90

29. Heat pipe/calorimeter thermal loss spreadsheet for pure He environment ........................ 91

30. Heat pipe/calorimeter thermal loss spreadsheet for He-32%Ar environment .................... 92
x

LIST OF ACRONYMS AND SYMBOLS

Ar argon

ARO after receipt of order

ASTM American Society for Testing and Materials

B boron

C carbon

Ca calcium

CAD computer-aided design

CH4 methane

Co cobalt

CO2 carbon dioxide

EB electron beam

EBR–II Experimental Breeder Reactor–II

ESCA electron spectroscopy for chemical analysis

F Fisher series designee

Fe iron

G mass flux G68–80 series designee

GASP gas properties

GDMS glow discharge mass spectroscopy

GFSSP generalized fluid system simulation program

H hydrogen

HCl hydrochloric acid

He helium
xi

LIST OF ACRONYMS AND SYMBOLS (Continued)

HTPIPE steady-state heat pipe code analysis program developed by LANL

H2O water

ID inside diameter

IFMP Integrated Financial Management Program

ISO International Organization for Standardization

K potassium

LANL Los Alamos National Laboratory

LECO LECO corporation builds combustion analyzers to determine compositions 

in inorganic samples

Mg magnesium

Mn manganese

Mo molybdenum

MSFC Marshall Space Flight Center

N nitrogen

Na sodium

Nb niobium

NDE nondestructive evaluation

Ni nickel

O oxygen

OD outside diameter

Re rhenium

RF radio frequency

SAES SAES Getters Group

Si silicon
xii

LIST OF ACRONYMS AND SYMBOLS (Continued)

Sn tin

SPAR space power advanced reactor

SP–100 Space Power 100 kW (program)

TIG tungsten inert gas

TP Technical Publication

UHP ultrahigh purity

V vanadium

VTASC visual thermofluid analyzer of systems and components

V&V validation and verification

W tungsten

WASP water and steam properties

Zr zirconium
xiii

NOMENCLATURE

A channel dimension; area 

Ac surface area calorimeter shell outer wall diameter 

Af coolant flow area 

Ai surface area inner wall

Ao surface area outer wall 

C impurity concentration

Cp specific heat 

CTE coefficient of thermal expansion

Di initial diameter

d heat pipe diameter 

dci calorimeter channel wall inner diameter 

dco calorimeter channel wall outer diameter

dh hydraulic diameter

di inner diameter 

do outer diameter 

dsi calorimeter shell wall inner diameter 

f friction factor

f

R material fraction of complete recrystallization

g acceleration due to gravity 

h convective heat transfer coefficient 
xiv

NOMENCLATURE (Continued)

hfg latent heat of vaporization 

I irradiation tests

k thermal conductivity 

ke effective thermal conductivity 

L calorimeter full length

La adiabatic length

Lc heat pipe condenser length

Le heat pipe evaporator length

l width 

l

gg gas gap width

M ̋ mass fluence 

MW molecular weight of component 

m mass of component 

m& mass flow rate 

mo stochiometric mass of getter 

N	 baseline life

Nu Nusselt number

n	 nucleation site radius

P pressure

Pr Prandtl number 

Q heat transfer rate
xv

NOMENCLATURE (Continued)

Q

cond conductive heat transfer rate

Q

conv convective heat transfer rate

Q

fluid fluid heat transfer rate

Q

gas cond gas conductive heat transfer rate

Q

rad radiation heat transfer rate

Q

solid cond solid condition heat transfer rate

Q

total total heat transfer rate

q applied evaporator power 

qrad radial heat flux 

R universal gas constant 

Ra Raleigh number

Re Reynolds number

r maximum pore radius

T temperature

T0 design temperature

Ta ambient temperature

Tavg average temperature 

Tcw calorimeter wall temperature 

Tcwi calorimeter channel inner wall temperature 

Tcwo calorimeter channel outer wall temperature 

Tf flow annulus fluid temperature
xvi

NOMENCLATURE (Continued)

Th radiating surface temperature

Thp heat pipe temperature 

Ti inner surface temperature

Tin coolant inlet temperature 

To outer surface temperature

To 1-hr recrystallization temperature

Tout coolant outlet temperature 

T∞ ambient temperature 

t time

V fluid velocity 

Vdisk getter volume 

α thermal diffusivity 

αCTE thermal expansion coefficient 

α(T) normalized Arrhenius diffusion rate

β gas thermal expansion coefficient 

Γ viscosity ratio

∆D change in diameter 

∆E work function

∆H activation energy

∆p delta pressure

∆T delta temperature 
xvii

NOMENCLATURE (Continued)

∆TLM log mean temperature difference

δ gap spacing

ε	 total hemispherical emissivity

εi inner surface emissivity

εo outer surface emissivity

μ absolute viscosity 

μbulk bulk fluid viscosity at average temperature 

μwall fluid viscosity at wall temperature

ν kinematic viscosity

ρ density

σ Stefan-Boltzmann constant

τ time

τi	 time at temperature

τmax maximum time

τo recrystallization time at To

φ flow annulus diameter ratio 

ψ flow annulus gap adjustment parameter
xviii
1

TECHNICAL PUBLICATION

DESIGN OF REFRACTORY METAL LIFE TEST HEAT PIPE 

AND CALORIMETER

1. INTRODUCTION

In the 1980s, a group at Los Alamos National Laboratory (LANL) began to approach heat pipe life 

testing using science-based methods. Tests were conducted on various alkali metal heat pipes for extended 

periods at elevated temperature and power. Material examination and assay were often conducted before 

and after test. Techniques such as x-ray microanalysis, electron microprobe, x-ray fluorescence, and Auger 

electron spectroscopy for chemical analysis (ESCA) were used in characterization. High-temperature corrosion modes were proposed from forensic examination of failed and nonfailed heat pipes. Attempts were 

made to understand complex chemical equilibrium mechanisms using free-energy minimization techniques with data supplied or extrapolated from the literature. One such thermal chemical simulation of the 

niobium (Nb)/potassium (K) system with typical contaminant levels found no life-limiting corrosion after 

7 yr of operation at 875 K.1

During the Space Power 100 kW (SP–100) program, work was also done at LANL to systematically test heat pipes at several temperatures and mass fluences. Sena reported results for eight K-filled 

Nb-1%zirconium (Zr) heat pipes that were tested for 7,000–14,000 hr in the 850–950 K range.2 One of 

the pipes developed a small evaporator leak at 13,000 hr that did not affect the operation of the heat pipe 

and was not detected by the vacuum system monitor. Tests on the other heat pipes concluded with no apparent problems. It is believed that Zr in an Nb-1%Zr centering wire touched the quartz tube during test. A 

wire-quartz contact point on the failed heat pipe surface was discovered after 13,000 hr of test. Zirconium, 

being more stable than quartz partitioned the oxygen, causing the oxygen (O) to diffuse from the quartz 

to the condenser wall. Oxygen from the quartz appears to have migrated to the heat pipe and saturated 

the Zr in the evaporator. Conditions for these eight Nb-1%Zr/K heat pipe tests approximated a 32 factorial configuration with temperature and radial heat flux to the evaporator as the two controlled factors. As 

far as it is known, this test was one of the first attempts to develop a multifactor response surface in a heat 

pipe corrosion test. In this case, results from the partial two-level factorial test series could have identified 

first-order effects and interactions between the two factors.3
2

2. ACCELERATED LIFE TEST HEAT PIPE APPROACH

There are several ways to establish heat pipe material system life without a full-term test. One 

approach is to hasten life-limiting effects. This can be accomplished, for instance, by testing at higher 

than the proposed operating temperature and mass fluence. Acceleration testing is intuitively appealing. 

The impurity corrosion rate is proportional to the accumulation of nonmetallic impurities such as oxygen, 

silicon, and carbon in the heat pipe evaporator. As the working fluid flows into the evaporator it vaporizes, 

concentrating the impurities and making the corrosion rate appear dependent on mass fluence. The radial 

heat flux (q

rad) applied to the evaporator is given by

q&

rad = q& π d L ( e )

–1 , (1)

where q is the applied evaporator power, d is the heat pipe diameter, and Le is the length of the evaporator 

section. Mass flux (G) through the evaporator is a function of the radial heat flux as

G = q&

rad h( fg )

−1 , (2)

where hfg is the latent heat of vaporization. Mass fluence (M ̋) through the evaporator is then given by

M ̋=Gτ , (3)

where τ is the operational time. Mass fluence can easily be increased by applying power along a shortened 

heat pipe evaporator length. Mass fluence can be used as a metric to accelerate heat pipe life or to affordably 

test subscale versions of a flight unit.

Mass diffusion transfers impurities from the heat pipe structure to the working fluid. The Arrhenius 

equation relates impurity diffusion rates to heat pipe temperature. Test conditions can be normalized using 

an expression such as

α(T ) = exp (ΔH /R) T0

−1 −T −1 ( ) ⎡

⎣ ⎤⎦ , (4)

where ∆H is activation energy, R is the universal gas constant, T0 is the design temperature, and T is the 

heat pipe test temperature. It can be argued that since testing at high temperature accelerates the Arrheniusgoverned diffusion rate, this same acceleration applies to heat pipe life. Such simplistic views must be 

approached with caution. The nature of chemical reactions and their kinetics can change dramatically with 

absolute temperature. Conditions that apply at one temperature do not necessarily hold true for others. 

Without deeper understanding of the underlying reaction mechanisms, conclusions drawn from such tests 

may be misleading.

Another approach to heat pipe life prediction involves extrapolation of corrosion effects of a 

reference design from separate tests conducted over different durations. Corrosion or reaction metrics 
3

are taken for each test specimen and the results are plotted versus time. Appropriate metrics include 

changes in fluid composition, wall and wick integrity as determined by three-dimensional x-ray computed 

tomography, species distribution, and grain boundary condition as well as noncondensable gas production 

measured by residual gas analysis. For the corrosion rate, a curve can be generated with a slope that 

normally decreases with time. The long-term corrosion rate is extrapolated from the slope near the longest 

test time. This method is outlined in ASTM G 68–80.4 Conditions of an extrapolated test series can 

closely match those of a proposed reference. Extrapolation tests conducted at prototypic temperatures 

avoid possible problems with elevated temperature acceleration.

Since the extrapolation method, outlined in ASTM G 68–80, is an accepted and defendable test 

practice, serious consideration has been given to following its guidelines for initial screening of life test 

heat pipes for potential material systems. Certainly, if time and budget constraints limit the number of life 

heat pipe tests, extrapolation should be favored over other test methods. 

One or more experiments from an extrapolation test series can be used as the center point conditions 

for a later multifactor star design test series. Life tests that establish corrosion trends as a function of 

temperature, mass fluence, and contamination level are consistent with a test-to-failure design philosophy.

A heat pipe life test matrix is proposed for a molybdenum (Mo)-44.5% rhenium (Re)/sodium (Na) 

material combination as a function of operating temperature and mass fluence. Impurity levels in the heat 

pipe wall and working fluid are to be controlled at constant values. A series of accelerated life heat pipes 

permits compression of life-limiting mechanisms from the mission duration to a period consistent with 

the schedule of a flight program. The target of the proposed test program is to compress 12-yr operational 

life in about 3 yr of actual operation. Condenser calorimetry using forced convection of a cooling fluid 

is necessary to achieve the target power density and to provide realistic boundary conditions of a flight 

system.

Life test interpretation will require the initiation of a science-based program to pursue an 

understanding of mechanisms relevant to heat pipe life. This effort should provide insight into the 

effects that will be observed during periodic removal of life test heat pipes for nondestructive evaluation. 

Destructive evaluation will be planned for preselected heat pipes and for heat pipes that show early 

evidence of corrosion.

2.1 Proposed Test Matrix

The proposed test matrix conserves available material, eases fabrication, provides relevant data, and 

permits the use of compact test fixtures. A series of �16 short heat pipes with thick-walled condensers will 

be constructed. These short heat pipes will have similar condenser-to-evaporator area ratios relative to the 

reference design. The boiling limit for the baseline design is invariant with condenser length, so the short 

heat pipes should perform identically to the longer baseline versions.

A series of accelerated life heat pipes permits extrapolation of corrosion effects of a reference 

design from separate tests conducted over different durations. Destructive examination of the first life test 

heat pipe can begin within 6 mo of test initiation. Testing a series of heat pipes allows identification of 

random manufacturing and processing defects.
4

Tables 1 and 2 show a proposed test series consisting of ≈16 short heat pipes (ASTM G 68–80 

G-series and Fisher multifactor star F-series). One or more experiments from an extrapolation test series 

is used to cross-correlate fluence conditions with a Fisher multifactor star design test series conducted at 

various temperatures and fluences.3 Such life tests can establish and isolate corrosion trends as a function 

of temperature and mass fluence. Figure 1 plots the fluence and temperature relations for the combined 

test series.

Table 1. Option A life tests, Fisher series (F-series).

Heat 

Pipe

Wall/Wick 

Material/

Fluid Lc/Le

Di

(cm)

T

(K) α(T )

q

(W)

Q

rad

(W/cm2)

G

(kg/cm2–s)

t

(hr)

M  ̋

(kg/cm2)

Design Mo-Re/Na 2 1.41 1,250 1.00 5,752 22 0.059 105,120 2,233

F(–4)

F(–3)

F(–2)

F(–1)

F(0)

F(1)

F(2)

F(3)

F(4)

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

3

3

3

3

3

3

3

3

3

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1,273

1,273

1,373

1,173

1,273

1,223

1,223

1,323

1,323

1.30

1.30

3.74

0.38

1.30

0.72

0.72

2.25

2.25

5,000

1,000

3,000

3,000

3,000

2,000

4,000

2,000

4,000

226

45

136

136

136

90

181

90

181

0.599

0.120

0.368

0.351

0.360

0.237

0.474

0.243

0.485

26,280

26,280

26,280

26,280

26,280

26,280

26,280

26,280

26,280

5,669

1,134

3,483

3,321

3,401

2,241

4,481

2,295

4,589

Table 2. Option A life tests, G68–80 series (G-series).

Heat 

Pipe

Wall/Wick 

Material/

Fluid Lc/Le

Di

(cm)

T

(K) α(T )

q

(W)

Q

rad

(W/cm2)

G

(kg/cm2–s)

t

(hr)

M  ̋

(kg/cm2)

Design Mo-Re/Na 2 1.41 1,250 1.0 5,752 22 0.059 105,120 2,233

G–1

G–2

G–3

G–4

G–5

G–6

G–7

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

3

3

3

3

3

3

3

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1,273

1,273

1,273

1,273

1,273

1,273

1,273

1.3

1.3

1.3

1.3

1.3

1.3

1.3

3,000

3,000

3,000

3,000

3,000

3,000

3,000

136

136

136

136

136

136

136

0.36

0.36

0.36

0.36

0.36

0.36

0.36

4,380

8,760

13,140

17,520

21,900

26,280

30,660

567

1,134

1,701

2,267

2,834

3,401

3,968
5

6,000

5,000

4,000

3,000

2,000

1,000

0

Mass Fluence (kg cm–2)

Temperature at Evaporator Exit (K)

F(–4)

F(4)

F(–2)

F(2)

F(–3)

F(1)

F(–1)

F(3)

F(0) G(7)

G(6)

G(5)

G(4)

G(3)

G(2)

G(1)

800 1,000 1,200 1,400

Fisher

G68–80

Design

Figure 1. Life test matrix assuming a maximum 1-μm nucleation 

site radius for F- and G-test series.

These test series attempt to systematically correlate heat pipe corrosion behavior at several 

temperatures and mass fluences. Conditions for the Fisher test (table 1) represent a two-factor 

(22+2× 2 +1) central composite configuration, with temperature and radial heat flux to the evaporator as 

the two controlled factors. Assuming corrosion is measurable, forensics are conducted and a multifactor 

response surface can be developed from this heat pipe corrosion test. In this case, results from the partial 

two-level factorial test series can identify first-order effects and interactions between the two factors, and 

can estimate second-order curvature effects in the data.3

It must be emphasized that the target radial heat fluxes in tables 1 and 2 are quite high, up to 10×

that of a flight heat pipe design. Every attempt will be made to approach these target values by controlling 

surface roughness and noncondensable gas in the heat pipe. These measures should inhibit boiling. An 

initial test will include a performance test of the first pipe to show what radial heat fluxes and temperatures 

can be achieved. It may be necessary at this point to reduce the operating power by some factor to avoid 

a performance limit. Such reductions will necessarily lengthen the test time required to achieve a given 

mass fluence (app. A).
6

3. HEAT PIPE DESCRIPTION

There are several options that will allow a quick start to life tests of a refractory metal heat pipe. In 

the near term, materials availability will dictate the options and schedule of any heat pipe performance test. 

Unfortunately, a supply of 400-mesh Mo-41%Re cloth produced in 1982 for the space power advanced 

reactor (SPAR) reactor project has been depleted. A few remnants of this cloth remain, mostly in small 

pieces suited only for short wicks. However, other materials exist that could make suitable refractory 

metal pipes. For our immediate purposes, a heat pipe will be made at Marshall Space Flight Center (MSFC) 

from readily available refractory metals, such as Mo-44.5%Re, and filled with distilled Na stock produced 

by LANL. This Na stock was used to support a previous test program.5,6

A limited supply of fine wire mesh of Mo-5%Re is available to the program, courtesy of David 

Glass at NASA Langley Research Center (R. Reid, Private Communication, June 2004).There is enough 

400-mesh Mo-5%Re wire cloth on hand to make several heat pipe wicks. Forming wicks from brittle 

Mo-5%Re is not easy. Several techniques have been developed to produce satisfactory porous tubes from 

this material. Making annular or arterial wicks from Mo-5%Re now appears possible and could be completed in approximately 16–22 wk. Wick fabrication will be easier once Mo-41%Re or Mo-44.5%Re screen 

material becomes available that has room temperature ductility after recrystallization. Rhenium has low 

solubility to nonmetallic impurities and alloys with higher Re content should exhibit similar or superior 

corrosion resistance compared to Mo-5%Re. 

Closure plugs can be attached to long Mo porous annular or artery tubes with Zr or palladiumcobalt brazes. Unfortunately, such brazes induce recrystallization and room temperature brittleness. Brittle 

wicks are prone to shipping damage; therefore, they were fabricated and shipped in the ductile condition. Closure plugs will be attached to the wick body at room temperature by mechanical crimping (no 

brazing is planned for the test units). The manufacturer will clean the wick chemically and by firing 

in a dry-hydrogen furnace (hydrogen dewpoint of less than –77 °C) at a temperature that avoids 

recrystallization.

3.1 Heat Pipe Sizing and Performance

Table 3 shows an accelerated life heat pipe design. The heat pipe will be operated horizontally 

to mimic the neutral gravity conditions of space. An annular gap wick is selected for the life heat pipe because 

it is simple, robust, and performs well under ground test. Properly designed short heat pipes tested at relevant temperatures and mass fluence are made geometrically similar (Lc/Le) to the flight heat pipes. Such 

similarity in heat pipe designs form the basis for virtually all long-term heat pipe corrosion testing to date.7

The length of the heated region is 7.5 cm and the cooled region is 25 cm, reflecting approximate geometric similitude with a flight heat pipe. For the case of an annular gap wick, the channel dimension (A) is 

the concentric distance from the wick’s outer and the pipe’s inner surfaces. The value for A is chosen to 

optimize life test heat pipe performance at the reference temperature. A wick’s maximum pore radius (r) 

most directly influences its capillary pumping capacity. A 45-μm effective pore radius has been assumed
7

Table 3. Life heat pipe design and conditions (neutral gravity).

Scale Value

Wick shape

Evaporator length

Adiabatic length

Condenser length

Container inside radius

Channel dimension

Wick pore radius

Nucleation site radius

Solid thermal conductivity

Working fluid

Temperature

Design heat pipe power

Annular gap

Le = 0.075 m

La = ≈0 m

Lc = 0.25 m

R = 0.705 cm

A = ≈0.056 cm

r = 35 μm nominal

n = ≈1 μm

k = 60 W.m–1K–1

f = Na

T = range 1,173 to 1,373 K

Q = 3 kW

for the life heat pipe design. This value reflects bubble-point measurements taken with highly compressed 

400-mesh screen wicks manufactured by Advanced Methods and Materials from February 2002 to the 

present. Radial power densities applied to the evaporator, operating temperature, and nucleation site radii 

are key variables in boiling limit calculations. Nucleation site radii of 1–3 μm are typical for many engineering surfaces and will be assumed for the life heat pipe. For the life test heat pipes designed here, an attempt 

was made to produce surfaces with ≈1-μm nucleation site radius. Section 3.1.1 discusses the nucleation 

site radius sensitivity.

3.1.1 Performance Envelope

A steady state heat pipe analysis computer program, HTPIPE, predicts heat pipe performance 

limits as a function of evaporator exit temperature and calculates axial temperature and pressure profiles.8

The computational routines in HTPIPE have been validated on numerous occasions using alkali metal 

heat pipe test data.9,10 The HTPIPE code uses a pressure balance in the flow direction for both the liquid 

and vapor to define the wicking limit or the viscous limit, depending upon whether the wick pumping 

capacity or the stagnation pressure defines the limiting available pressure. Other heat pipe performance 

limits calculated by the program are the sonic limit, the entrainment limit, and the boiling limit. Figure 2 

superimposes the proposed life test matrix on a performance map calculated with HTPIPE. The heavy red 

line shows the boiling limit assuming a 1-μm nucleation site radius. The lighter line shows the boiling limit 

for a 3-μm radius. This plot emphasizes the importance of producing heat pipes with honed heated regions. 

Failure to achieve a high boiling limit will require that certain set points be rescoped or the possible 

shifting of the entire test matrix.
8

6,000

5,000

4,000

3,000

2,000

1,000

0

Axial Heat Transfer Rate (W)

Temperature at Evaporator Exit (K)

Capillary

Boiling 3 μm 1 μm

–4

3

–1 –2

1

0

2

–3

4

800 1,000 1,200 1,400

Sonic

Figure 2. Proposed life test matrix compared to operating 

limits at two-nucleation site radii.

3.1.2 Wick Assembly Description

During cleaning, the wire cloth material will be vacuum fired and/or hydrogen heat treated, with the 

hydrogen dewpoint less than –77 °C, the nature of which will be discussed with and approved by MSFC. 

During these heat treatments the material will be kept within the following time-temperature envelope:

τ max (T) < τ o f

R exp (– ΔE R) 1 To ( −1 T ) ⎡

⎣ ⎤⎦ , (5)

where f

R is the material fraction recrystallized, τo is the recrystallization time at To (1 hr), R is the universal 

gas constant, and ∆E is the work function. One-hour recrystallization times for pure Mo and activation 

energy data, shown in figure 3, is used to prevent recrystallization of the Mo-5%Re material.11,12 For multiple 

heat-treat steps at different temperatures, the relation 

τ i Ti ( ) τ max Ti ( ) ⎡

⎣ ⎤⎦ < 1

i=1

N

∑ , (6)

applies where τi is the time at temperature Ti for a given step and τmax is the maximum time at temperature 

for a given processing step. Figure 3 shows the not-to-exceed time proposed for wick material as a function 

of temperature at various percent recrystallization.
9

100,000

10,000

1,000

100

10

1

0.1

0.01

Time (min)

Temperature (K)

1,000 1,050 1,100 1,150 1,200 1,250 1,300

0.03

f = 0.01

1

0.3

0.1

Figure 3. Not-to-exceed time-temperature relation 

for Mo-5%Re material.

3.1.3 Getter Pack Description

Corrosion in high-temperature heat pipes can be inhibited with nonsoluble getters that trap impurities 

during operation. Such getters can be distributed to allow wetting by the condensate and can be part of the heat 

pipe container or placed in packs. Assuming a well-distributed getter, the potential for corrosion should relate 

to the total impurity concentration compared with the total stoichiometric impurity capacity in the getter and 

container as follows:

Σ ≡ (impurity capacity of heat pipe getter) × (impurity content of heat pipe)–1 . (7)

Getter material in high-temperature heat pipes should be hyperstoichiometric, compared with nonmetallic impurities from all sources. As a rule of thumb, the impurity capacity to impurity ratio should be 

at least several times stoichiometric proportions, shown as follows:

mZr = ρ ZrVdisks , (8)

mZr

o = mMoMWZr

CO

Mo

MWO

+

CN

Mo

MWN

+

CC

Mo

MWC

+

CSi

Mo

MWSi

+

CB

Mo

MWB

⎡

⎣

⎢

⎢

⎤

⎦

⎥

⎥

+ mNaMWZr

CO

Na

MWO

+

CN

Na

MWN

+

CC

Na

MWC

⎡

⎣

⎢

⎢

⎤

⎦

⎥

⎥

, (9)

and

Σ = mZr

mZr

o . (10)

Using these relations, a getter pack consisting of 1.5 g of Zr foil provides a factor of 6.4 margin for 

a 0.3-kg tube with 20 ppm oxygen (O), 20 ppm carbon (C), 30 ppm nitrogen (N), 30 ppm silicon (Si), and 
10

20 ppm boron (B) filled with 10 g of Na containing 10 ppm oxygen. Hafnium and vanadium (V) getters 

can also be used as in Na heat pipes. Titanium is appropriate at a slightly lower temperature.

3.1.4 Heat Pipe Material/Envelope

The heat pipe envelope will be manufactured from an Mo/Re-based alloy with the two leading 

candidates having blends of 44.5% and 47.5% Re (by weight) to improve low-temperature ductility, supplied in a recrystallized state annealed at 1,775 K. The components (tubing and rod stock) will be drawn/

swaged (or another acceptable technique) from raw material formed by either powder metallurgy or arc 

cast techniques. There have been a series of discussions between Ross Luther (Bechtel Bettis) and MSFC 

personnel regarding composition, material forming, chemical composition, and other metallurgical considerations pertinent to the process. Details are listed in the specification statement submitted for procurement, as provided in section 3.3. Current vendor estimates indicate delivery time of the required Mo-Re 

alloy components is 16–20 wk after receipt of order (ARO). 

A key element of these specifications is the chemical composition of the bulk materials used in 

fabricating hardware components. Test methods to verify the composition can include a LECO Corporation elemental analyzer and glow discharge mass spectroscopy (GDMS) to examine both bulk and surface 

compositions. A list of the maximum and targeted impurity concentrations is provided in table 4.

Table 4. Impurity chemical composition (acceptable) final heat 

Mo-Re alloy material (powder metallurgy).

Element

Maximum

(wppm) 

Target

(wppm) 

O

C

N

H

Ca

Fe

Ni

W

Si

Mg

Mn

Co

Sn

B

20

20

30

25

7

75

50

300

30

10

4

30

20

20

<10

<10

<3

<1

<5

<20

<20

<100

<10

<10

<4

<30

<20

<10

A significant area of concern involves the electron beam (EB) welding process for the Mo-Re 

alloy that will be employed to assemble the heat pipe units. Based on information gathered from Ross 

Luther (Bechtel Bettis), the oxygen and carbon content in the alloy are critical to the success of the welding process. Luther recommends that the maximum oxygen and carbon content concentration should be 

set to 20 ppm for powder metallurgy material, and for arc cast material, the maximum oxygen concentration should be set to 20 ppm with a carbon range of 40 to 80 ppm. During further discussions, it has become 

apparent that some welding tests of appropriate sample coupon materials should be performed to investigate 
11

and document the operations (examine weld ductility, shock resistance, and porosity). Testing could 

potentially be performed at MSFC, LANL, and subcontractor sites such as Advanced Methods and 

Materials. These tests can help resolve any critical surface preparation and welding parameters. Cleaning 

of material can consist of chemical cleaning and vacuum/hydrogen firing. A proposed general chemical 

cleaning procedure for all Mo-Re parts includes the following:

(1) Wash the piece in PF® solvent, an R113 substitute, until all signs of grease have been removed.

(2) Soak the piece for 1–2 min in one part by volume HCl and one part by volume deionized water 

(target of 10 MΩ cm or better) to remove residual iron surface impurities.

(3) Soak the piece in caustic cleaning solution consisting of 11 parts by volume deionized water 

(target of 10 MΩ cm or better), 1 part by volume NaOH, and 1 part by volume H2O2 for 5 min. Remove 

piece from caustic bath. Replenish or replace solution as required.

(4) Wash the part in hot deionized water (target of 10 MΩ cm or better) for at least 5 min.

(5) Repeat steps (3) and (4) three times.

(6) Rinse piece in ethanol inside an ultrasonic cleaner for 5 min.

(7) Vacuum-fire or hydrogen-fire part as appropriate, not exceeding specific time-temperature 

requirements associated with recrystallization of low-Re content Mo materials. Hydrogen used in firing 

will have a dewpoint of less than –77 °C.

The heat pipe wick structure will be produced from existing Mo-5%Re mesh material, a 400×400 

weave with a wire diameter of 0.001 in. This material is in the uncrystallized condition, so it is sufficiently 

ductile for the wick-forming process. During subsequent vendor processing, the cumulative time-at-temperature must be monitored for all expected heat treatments to ensure that the material has not significantly 

recrystallized, becoming brittle at room temperature and increasing the chance of damage during shipping 

and handling. A proposed time-temperature relationship is provided in table 5. For multiple heat-treat 

steps at different temperatures, the relation given in equation (6) will apply.
12

Table 5. Time-temperature relationship for Mo-5%Re material.

Temperature 

(K)

Maximum Time

at Temperature

(min)

Ti τmax(Ti)

1,000

1,010

1,020

1,030

1,040

1,050

1,060

1,070

1,080

1,090

1,100

1,110

1,120

1,130

1,140

1,150

1,160

1,170

1,180

1,190

1,200

1,210

1,220

1,230

1,240

1,250

1,260

1,270

1,280

1,290

1,300

2,049

1,448.9

1,031.6

739.3

533.2

387.0

282.6

207.6

153.3

113.9

85.1

63.9

48.2

36.5

27.9

21.3

16.4

12.7

9.8

7.7

6.0

4.7

3.7

2.9

2.3

1.9

1.5

1.2

1.0

0.8

0.6

3.2 Heat Pipe Engineering Drawings

Required heat pipe performance specifications were translated into an engineering design and drawing generated for the wick structure, getter pack, and envelope. A computer-aided design (CAD) software 

package called Autodesk Inventor Series Version 8 was used to create the engineering drawing set. Three 

drawings were created to describe the complete heat pipe unit. Figure 4 shows the wick assembly with the 

evaporator end plug that provides sealing of the capillary channel. The current approach is to mechanically 

crimp the end plug to the wick for the final test units with a vendor-developed process that eliminates a hightemperature braze resulting in a brittle wick assembly. The sealing integrity of this crimp and of the wick 

material will be checked during bubble-point testing of the assembly. 
13

Figure 4. Wick assembly engineering drawing.

Figure 5 shows details of the getter unit and its assembly. The getter is an integral part of the heat 

pipe evaporator end plug assembly. The getter is comprised of 31-Zr disks stacked and held in place with a 

tack-welded cover plate and support rod. Figure 6 shows details of the heat pipe envelope components and 

final assembly of the unit. Electron beam welding is used exclusively during assembly of the hardware. 

This information, with specifications, is contained in the heat pipe procurement solicitation provided in 

section 3.3.
14

Figure 5. Getter pack design and assembly engineering drawing.

Figure 6. Heat pipe envelope and assembly engineering drawing.
15

3.3 Heat Pipe Procurement Specifications

Procurement specifications have been generated to cover fabrication of the heat pipe units. These 

procurements will be issued by MSFC using full and open competition. The final evaluation of vendor 

quotes will be based on the lowest cost, technically acceptable approach. Three procurement specifications 

were developed covering each of the primary requirement areas as follows:

(1) Wick assembly fabrication for heat pipe units using existing Mo-5%Re 400×400 mesh material 

and an Mo-Re evaporator plug from the material type selected for the heat pipe envelope. Wicks will be 

fabricated by drawing, or another acceptable technique, to produce average 35-μm pores. The completed 

wicks will have documentation that details the process and the results of cleaning, bubble-point testing, 

and material chemical composition testing. Appendix B contains the detailed wick assembly procurement 

specification. The current vendor estimate is 16–22 wk for delivery.

(2) Basic materials for the heat pipe envelope are composed of an Mo-Re alloy with a blend of 

44.5%Re by weight (44.5%Re is a Rhenium Alloys Inc. specific material that has no sigma phase-brittle 

intermetallic) or 47.5%Re by weight. The material quoted can be formed by either powder metallurgy or 

arc cast techniques. The required component shapes include tube, rod, and plate stock of various diameter 

and lengths that have material data sheets detailing chemical composition, ductility, grain size, etc. 

Appendix C contains the detailed material procurement specification. The current vendor estimate is 

16–20 wk for delivery.

(3) Heat pipe fabrication and assembly use the tube, rod, and plate stock with wick assemblies 

detailed under the previous two specifications. This will include machining of component pieces and getter pack, followed by EB welding to assemble the units. Machining and assembly processes will be documented and documentation will be provided, as necessary, for material chemical composition, welding 

technique, leak check, ductility, etc. Appendix D contains the detailed heat pipe fabrication procurement 

specification. The current vendor estimate is 16–20 wk for delivery.

These specifications do not include the engineering requirements needed for final closeout of the 

heat pipe units after Na loading at MSFC. These requirements will be developed once the details pertaining 

to the heat pipe final mounting method and Na-sampling technique have been determined.
16

4. CALORIMETER UNIT

A calorimetry system is required to extract heat from the condenser section of the heat pipe and 

transfer it to a media that can be measured to quantify the power being removed from the heat pipe. 

There are two primary techniques for removing energy from the heat pipe condenser section: (1) Pure 

radiation to the surroundings and (2) conductive/forced convective coupling to a working fluid.13,14 In 

the case of radiation calorimetry, the heat transfer is controlled by the Stefan-Boltzmann relationship, 

Q =εσA(Th

4– Ta4), where Q is the total transferred power, ε is the total hemispherical emissivity, σ is the 

Boltzmann constant, A is the radiation surface area, Th is the radiating surface temperature, and Ta is the 

ambient temperature. This is the simplest configuration, requiring only that the surrounding chamber be kept 

cool and a method employed to monitor the surface temperature of the heat pipe condenser. Radiation heat 

transfer can be high at the expected heat pipe test temperatures; however, for Mo-based alloys, the total 

hemispherical emissivity of the surface is typically very low (≈0.15), resulting in poor overall heat transfer, 

with an approximate range of 200–300 W at proposed test temperatures. The second calorimetric technique 

relies on a conductive/forced convective combination to remove energy from the heat pipe condenser. This 

system is more complex, as it requires additional hardware and a working fluid to absorb and transport the 

energy. The typical approach is a two-step process making use of a static gas conduction gap between the 

heat pipe condenser section and the inner wall of the calorimeter, followed by a flowing fluid contained in 

the calorimeter annulus. The heat transfer rate across the static gas gap is governed by a combination of the 

radiation heat transfer rate, the Stefan-Boltzmann relationship, and the Fourier heat transfer rate relation,

Q =kA(Thp–Tcw)/l, where k is the thermal conductivity, Thp is the heat pipe temperature, Tcw is the calorimeter wall temperature, and l is the gap width with typical gap widths ranging from 0.02 to 0.10 in, depending 

on the magnitude of energy to be transferred. The gas gap can either be sealed, if the calorimeter system 

is physically attached to the heat pipe condenser with a seal weld, or open, if the heat pipe and calorimeter 

are contained in a test chamber with the desired gas constituents. The calorimeter coolant (gas, water, or 

other fluid contained in the flow annulus) absorbs and transports the energy to an external heat exchanger 

where it is removed. Heat transfer into the calorimeter fluid is governed by the convective heat transfer rates,

Q =hA(Tcw–Tf), where h is the convective heat transfer coefficient, Tf

 is the flow annulus fluid temperature, 

and the difference (Tcw–Tf) is commonly evaluated as the log mean temperature difference (∆TLM).

In general, sealing the static gas gap to the heat pipe condenser results in a more complex system 

(calorimeter is designed and fabricated as part of the heat pipe condenser) when compared to the noncontact 

approach with a gas-flooded test chamber. This complexity is compounded when fabricating heat pipe 

units from refractory alloys, requiring the calorimeter to be fabricated from the same material and limiting 

the types of coolant that can be used to maintain compatibility. For Mo-Re, this becomes a very expensive 

endeavor. In addition, there is little experience in fabricating and welding the structural shapes and fittings 

required to make an Mo-Re alloy calorimeter assembly, making it a high-risk undertaking. In contrast, the 

noncontact calorimeter can be manufactured from more common materials, such as stainless steel or copper, 

and supported so the heat pipe and inner diameter of the calorimeter remain nearly concentric. The final 

design will incorporate provisions, such as bumpers or foils, to prevent the heat pipe and calorimeter 

from touching. Such isolation also avoids uncontrolled mass-transfer conditions believed to have been 

experienced in previous tests.2
17

Selection of the coolant flow system to be used in the calorimeter flow annulus can again be 

addressed based on complexity. The most complex arrangement would make use of a gas flow system 

(helium (He) or He/argon (Ar) mixture) requiring a single, high-power gas compressor, or a series of lower 

power units providing sufficient system lift (order of several psia) to account for flow pressure drop with 

a plumbing system capable of operating in the 50–150 psia range. These components are available commercially; however, they are expensive and often require lead times or modifications. A simpler calorimeter 

cooling technique would consist of a pressurized water loop operating at 50–75 psia (to increase the water 

boiling point) with a pump providing up to 20 psi of lift (compensating for system losses). Components 

for this system are commercially available and are typically off-the-shelf items. The downside to a waterbased system is that there is a risk associated with the potential leakage of water into the test chamber 

and its potentially catastrophic interaction with refractory metals at high temperatures. To minimize this 

potential, all water feed lines should be welded/brazed and mechanical connections should be avoided.

Based on these considerations, the selected approach was a noncontact calorimeter design using 

water as the cooling fluid and He/He-Ar gas mixture as the test chamber gas species. The test chamber gas 

pressure would be maintained in the 50–100 torr range to minimize convective losses while providing 

margin above the potential Paschen minimum pressure for ionization.15 The heat pipe power source will be 

a radio frequency (RF) inductive coil that will operate at 100–450 kHz and up to 500 V. 

4.1 Water Calorimeter Description

The selected baseline calorimeter was a smooth tube laminar design lending itself to simple performance estimates supporting initial concept feasibility assessment. The initial configuration layout is illustrated 

in figure 7. It is comprised of three concentric tube shells referred to as: (1) Channel tube, separating the 

gas gap and water film annulus; (2) shell tube, separating the water film annulus and water return annulus; 

and (3) cover tube, separating the water return annulus and test chamber environment. The inlet end of the 

calorimeter makes use of a double wall to create an inlet plenum to distribute water to the film annulus. 

The cooling water passes along this annulus to the end of the calorimeter, where it passes through a series 

of holes/slots into the return path annulus.

The typical film flow annulus width is 0.01–0.02 in and the return path is sized at 0.125-in width; 

the majority of the system pressure drop is expected in the film flow annulus. The calorimeter would be an 

all-welded construction to minimize the potential for leakage. The overall length of the calorimeter body 

could range from 11 to 13 in, depending on the final heat pipe attachment technique. 

Currently, the proposed method for holding the calorimeter and heat pipe unit uses a concentric 

support/mounting bracket, as illustrated in figure 8. The mounting bracket uses a radial slot to capture and 

clamp the calorimeter, while the heat pipe is held in the central bore clamp. Both the calorimeter and heat pipe 

are cantilevered from the mounting bracket, minimizing heat pipe contact and restricting contact location 

to the coolest heat pipe element—the closeout stem. It is believed that the proposed mounting technique, 

that holds radial dimensional tolerance on a single-mounting element, would be the easiest to fabricate and 

would result in the most accurate approach to align both the calorimeter and heat pipe. A Mo or Mo-Re 

bumper foil may still be required, or desired, at the ends of the calorimeter to prevent unanticipated touching 

of the two components. If required, additional support can be provided at the very end of the evaporator plug 

tube extension which is 0.5 in long (0.5 in diameter by 0.03 in wall material).
18

Cover Tube

Seal Welds

Seal Welds

Water Return Path (0.125 in)

Film Flow Gap (0.01 – 0.02 in)

11 to 13 in

Shell Tube Channel Tube

Water Outlet

Water Inlet

Inlet 

Plenum

Flow Slots Between 

Film Flow Gap and 

Return

1 in

Figure 7. Calorimeter conceptual design.

Inductive Heater 

Coil Assembly

Mounting Bracket Uses 

Concentric Support to 

Minimize Misalignment

Heat Pipe Closeout Stem

Gas Gap (0.02 – 0.1 in Depending on Test)

Calorimeter Support/Clamp

Heat Pipe Support/Clamp

Figure 8. Conceptual heat pipe/calorimeter mounting setup.

Material selection for the calorimeter includes either 316 stainless steel or copper alloy. Stainless 

steel is easy to work with during fabrication and assembly (significant shop and laboratory experience), 

although the tight tolerance (±0.002 in on dimensions of 0.01 in over the full length) would prove challenging. On the other hand, stainless steel has several problem areas. First, it is available as tubing in a 

set division of diameters and wall thickness; therefore, any design would have to be tailored to the available sizes and then drawn or machined as necessary, potentially increasing tolerance variability. Second, 

it has a low-thermal conductivity (≈15 W/m-K) resulting in an appreciable temperature gradient across its 

thickness necessitating higher material surface temperatures to achieve required heat transfer rates into the 

calorimeter; however, the gas gap is still the limiting thermal resistance.

Copper is a better alternative for the calorimeter construction. The copper alloy C122 deoxidized 

high-residual phosphorus, commonly used in heat transfer tubing and heat exchangers, is 99.9% copper and 

0.02% phosphorus and has a thermal conductivity of 339 W/m-K at 300 K (value provided by Wolverine 

Tube, Inc.). This is a factor of 20 better than stainless steel and essentially eliminates any material thermal 
19

gradient by placing the hot-side material temperature at the coolant interface where it provides the most benefit. Discussions with Wolverine Tube, Inc. indicate that with their precision tubing process, their dies can be 

adjusted to draw tubing to virtually any diameter from 0.1 to 2 in, with a tolerance of 0.001–0.002 in. In this 

fabrication process, the tube wall thickness would be held constant for all diameters at one of several available wall thicknesses, the most attractive for our purposes being 0.058 in. This simplifies design, fabrication, 

assembly, and performance calculation processes since minimal machining will be required, introducing 

less variability in the final product.

Cost estimates for the calorimeter will be solicited once the final engineering drawings are completed. The initial design is shown in figures 9 and 10. The current material baseline for the calorimeter 

design is copper C122, as suggested by Wolverine Tube, Inc. The following sections provide a detailed 

summary of the calorimeter film flow gap sizing and its sensitivity to parameters such as coolant flow rate, 

pressure drop, calorimeter channel wall temperature, and film flow temperature rise (outlet-inlet).

Figure 9. Initial calorimeter concept engineering design layout.
20

Figure 10. Initial calorimeter concept engineering design parts detail.

4.2 Heat Pipe/Calorimeter Performance

Determination of the size and performance of the calorimeter system to meet the requirements 

for each heat pipe test case was accomplished using a simple spreadsheet-based model. A more detailed 

flow network model is also currently being built using the generalized fluid system simulation program 

(GFSSP); however, it was not completed in time to support the initial design work described in this Technical Publication (TP). The GFSSP model will be used later in the program to reassess the current findings 

and to provide pretest performance predictions.

The spreadsheet model assumes radial heat flow (heat pipe wall to calorimeter) and assumes average properties across the length of the assembly. This is a reasonable approach for steady state operation and provides an established set of operating test conditions. The selection of a pressurized water 

calorimeter, copper construction materials, and the use of the in-house water-cooling system provide 

additional constraints to the analysis. The approach taken was to divide the heat transfer path into two 

component elements including: (1) Heat flux from the heat pipe external wall across the gas gap to the 

calorimeter wall and (2) heat flux through the calorimeter wall into the coolant flow. The common factor coupling these elements, maintaining a balanced heat flux, is the calorimeter wall temperature. This 

temperature can be adjusted to obtain an acceptable solution for flow and pressure drop while meeting the 
21

required power transfer. For all model calculations, temperature-dependent material properties were used 

(evaluated at average conditions). Property information is provided in appendix E.

4.2.1 Static Gas Gap Conduction and Radiation Heat Transfer

The limiting thermal resistance in the heat transfer path between the heat pipe surface and the calorimeter coolant flow is conduction across the static gas gap. This is an important feature allowing the 

inside diameter of the calorimeter channel to be sized for the expected heat pipe operating condition. The 

conductivity across the gap can then be tuned up or down using the gas composition as a knob to provide 

coarse adjustment in the overall heat flux. In addition, the maximum calorimeter channel wall temperature 

is set by the boiling limit of the pressurized water system. Figure 11 shows the variation in the water boiling 

point with pressure. For a cooling system operating pressure of 65 psia, the boiling point of water is ≈420 K. 

To maintain margin in the calorimeter design, a maximum calorimeter surface temperature should be in 

the 375–390 K range. 

Another consideration is the increase in heat pipe diameter due to its thermal growth over 

the expected operating temperature range; this will influence the final width of the gas gap. For the 

Mo-44.5%Re alloy heat pipe, the change in diameter can be assessed using

ΔD = α CTEΔTDinitial , (11)

where ∆D is the change in diameter, αCTE is the coefficient of thermal expansion, ∆T is the temperature 

difference between the two conditions, and Dinitial is the initial material diameter. The final result is an 

increase in heat pipe outer diameter of ≈0.004 in over the temperature range of 300 to 1,400 K, as illustrated 

in figure 12. A curve fit describing the heat pipe unit outer diameter as a function of temperature was used 

in the model to account for heat pipe radial growth at each operating condition. Pressure (Pa)

Temperature (K)

P=9.946×101×T 2 –7.167×104×T + 1.3×107

7×105

6×105

5×105

4×105

3×105

2×105

1×105

0

370 380 390 400 410 420 430 440

Figure 11. Water boiling point as a function of pressure.
22

Delta Diameter (in)

Diameter (in)

Temperature (K)

D = 4.76×10–10×T 2 + 2.891×10–6×T + 6.241× 10–1

0.01

0.001

0.0001

0.6295

0.629

0.6285

0.628

0.6275

0.627

0.6265

0.626

0.6255

0.625

0.6245

0 200 400 600 800 1,000 1,200 1,400 1,600

Diameter

Increase in Diameter

Curvefit

Figure 12. Mo-44%Re heat pipe diameter growth as a function of temperature.

The heat flux balance across the static gas gap is a summation of the radiation and conduction heat 

transfer components. The radiation component can be assessed assuming the operating temperatures of both 

the heat pipe and calorimeter. As expected, radiation accounts for a small portion of the total heat transfer 

(due to the low emissivity, ≈0.15, of the Mo-Re alloy at the operating temperature). The total heat flux rate 

(Q

total) is given by

Q&

total = Q&rad + Q&cond , (12)

where Q

rad is the radiation component rate and Qcond is the conductive heat transfer load rate. Figure 13 

illustrates the static gap and calorimeter layout with relevant parameter identified. The radiative component 

of the heat flux can be calculated using the Stefan-Boltzmann relationship and the emittance and area of 

each of the relevant surfaces as

Q&

rad =

σ Ai Thp

4 − Tcwi4 ( )

1 ε i + Ai Ao ( ) 1 ε o ( − 1) . (13)

With the known conductive heat-flux component and the Fourier conduction heat transfer relationship for 

concentric cylinders,

Q&

cond =

2π Lc

k Thp − T ( cwi)

ln do di ( ) , (14)
23

Heat Pipe 

Centerline

do

cond

rad

cond

conv

Tin Tcwo Ac

Tf

I

gg

Tout

Tcwi o Ao

Thp i Ai

dh/2

di

dci dco dsi

Heat Pipe 

Internals

Heat Pipe Wall

Static Gas Gap

Calorimeter 

Channel Wall

Coolant Annulus

Calorimeter 

Shell Wall

mH2O ∙

ε

ε

Figure 13. Static gap and calorimeter layout.

where do and di are the outer and inner diameters, respectively. The static gas gap width (l

gg) can be solved 

as follows:

l

gg = 1

2

do − di ( ) . (15)

Following this approach, the model was used to calculate the calorimeter sizing parameters for 

all heat pipe test conditions (condenser radial heat flux ranging from 8.5 W/cm2 at 1,000 W throughput 

to 42 W/cm2 at 5,000 W throughput). The results for two static gas gap compositions (pure He and 

He-32%Ar by mass) with an assumed calorimeter wall temperature of 390 K are illustrated in figure 14. As 

expected, the pure He condition requires use of a much larger gas gap width because its thermal conductivity 

is approximately twice that of the He-32%Ar mixture. 

The use of a reduced gas thermal conductivity in the layout provides for some upward margin in 

the design, making up for variability in the final width of the gas gap in the as-built hardware (e.g., more 

or less, Ar can be added to compensate for geometric variations to reach the required heat flux). The most 

difficult case is F(–4) that has the highest heat flux of 5,000 W, requiring an approximate gap width of 

0.014 in. This condition is followed closely by the 4,000-W cases, F(2) and F(4), with gas gap widths of 

approximately 0.0165 in and 0.019 in, respectively. To maintain a design that can be manufactured and 

assembled, a minimum gap width of 0.02 in should be used. Adhering to this gap limit drives the 5,000-W

case Ar concentration to be reduced from 32% to 6%, which still provides some limited upward margin in 

the design. 
24

1

0.1

0.01

Static Gas Gap (in)

Heat Pipe Temperature (K)

900 1,000 1,100 1,200 1,300 1,400 1,500

F(–3)

F(1)

F(3)

F(–4)

F(–1)

F(–2)

F(2) G’sF(0) F(4)

1,000 W He

1,000 W He-32%Ar

2,000 W He

2,000 W He-32%Ar

3,000 W He

3,000 W He-32%Ar

4,000 W He

4,000 W He-32%Ar

5,000 W He

5,000 W He-32%Ar

1,000 W He

2,000 W He

4,000 W He

5,000 W He

3,000 W He-32%Ar

4,000 W He-32%Ar

5,000 W He-32%Ar

3,000 W He

2,000 W He-32%Ar

1,000 W He-32%Ar

Gas Sensitivity to Gas Composition

Calorimeter Wall Temp. = 390 K (H2O Boiling Point = 420 K @ 65 psia)

Calorimeter Copper C122 Emissivity = 0.44

Mo-Re Emissivity = 0.16

Condenser Radial Heat Flux

1,000 W = 8.5 W/cm2

2,000 W = 17 W/cm2

3,000 W = 25.6 W/cm2

4,000 W = 34.1 W/cm2

5,000 W = 42.7 W/cm2

Figure 14. Sensitivity of the static conduction gap with gas composition.

An additional consideration is the sensitivity of the calorimeter channel wall temperature on the 

static gas gap width. As expected, it has as small effect for the F(2) test case at a power setting of 4,000 W, 

as shown in figure 15. The wall temperature was varied over a range from 320 to 420 K, capturing the 

maximum extent of available water temperature. The resulting gap width variation is 0.001 in for the 

He-32%Ar mixture and 0.002 in for the pure He mixture. Therefore, the assumption of a maximum 

calorimeter channel wall temperature of 390 K in the original assessment, used to produce figure 14, 

should be acceptable for the range of actual calorimeter temperatures expected.

The next step in the sequence is to link the calorimeter cooling film flow to the total heat transfer 

though the calorimeter channel wall. This step will require calculation of the convective heat transfer 

coefficient and coupling this to the coolant flow and pressure drop.
25

0.04

0.035

0.03

0.025

0.02

0.015

0.01

0.005

0

Static Gas Gap (in)

Calorimeter Wall Temperature (K)

320 340 360 380 400 420 440

4,000 W He

4,000 W He-32%Ar

Mo-Re Emissivity = 0.16

Copper C122 Emissivity = 0.44

Figure 15. Sensitivity of static gas gap with calorimeter wall temperature.

4.2.2 Calorimeter Coolant Flow Heat Transfer

The power absorbed by the calorimeter is described by the solid conduction heat transfer through the 

calorimeter channel wall thickness and convective heat transfer into the coolant flow. The temperature drop 

across the calorimeter channel wall is assessed using Fourier conduction for concentric cylinders, equation (16), 

where Tcwi and Tcwo are wall temperatures and dci and dco are the inner and outer diameters. Results show 

temperature drops of ≈1 K due to the high thermal conductivity of copper. Heat flux is then absorbed by the 

flowing coolant using convective heat transfer equation (17), where h is the convective heat transfer coefficient 

and Tf is a mean film temperature as follows:

Q&

cond =

2π Lc

k Tcwi − Tcwo ( )

ln dco dci ( ) (16)

and

Q&

conv = hAc Tcwo − T ( f ) = hAcΔTLM . (17)

The temperature delta (Tcwo–Tf) = ∆TLM is the log mean temperature difference and is given by the 

following expression:

ΔTLM = Tout − Tin ( )

ln Tcwo − Tin ( ) Tcwo − Tout ( ) ⎡

⎣ ⎤⎦

, (18)
26

where Tin and Tout are the inlet and outlet temperature of the coolant flow. The power transferred to the 

coolant flow results in a temperature increase (Tout –Tin) as given by

Q&

fluid = m& Cp

Tout − Tin ( ) (19)

and

m& = ρVAf , (20)

where Cp is the specific heat of water, m& is the mass flow rate, ρ is the density, V is the velocity, and Af

is the coolant film flow area. The convective heat transfer is calculated from a laminar form (Reynolds 

number (Re) <2,300) of the Nusselt number (Nu) and the hydraulic diameter (dh) as

hdh

k = Nu = 3.66 + 1.2

dco

dsi

⎛

⎝

⎜

⎞

⎠

⎟

−0.8

Re ≤ 2,300 , (21)

and

dh = dsi − dco . (22)

The flow regime (laminar in all cases for this assessment) is determined by the Reynolds number, which 

is a function of the hydraulic diameter and the average properties of the water coolant flow: 

Re = ρVdh

μ . (23)

The pressure drop (∆p) is also assessed using a laminar Reynolds number correlation to estimate 

the friction factor (f), where ψ is an adjustment factor for annular gap size and Γ is a fluid viscosity 

correction for nonisothermal flow conditions.16 It is assumed that the flow through the film annulus is fully 

developed due to the high aspect (total length-to-width) as follows: 

f = ψ(64 Re)Γ−0.25 Re ≤ 2,300 , (24)

ψ = (1− φ)

2 1+ φ2 + 1− φ2 ( ) ln(φ) , (25)

φ = d

co

dsi , (26)

Γ = μbulk μ wall , (27)

and

Δp = fLρV2

2d

h

. (28)

For the basic calorimeter design, discussed previously, figure 16 shows the power that can be 

removed by the calorimeter (Q

fluid) as a function of flow rate for a series of ∆Ts (where ∆T = Tout – Tin) 

for the coolant film channel.
27

16,000

14,000

12,000

10,000

8,000

6,000

4,000

2,000

0

22

19.25

16.5

13.75

11

8.25

5.5

2.75

0

Power (W) and Reynolds Number

Pressure Drop (psi)

Flow Rate (gal/min)

0 0.2 0.4 0.6 0.8 1 1.2

Power (∆T Film Flow = 20 K)

Power (∆T Film Flow = 30 K)

Power (∆T Film Flow = 40 K)

Power (∆T Film Flow = 50 K)

Power (∆T Film Flow = 60 K)

Reynolds Number

Pressure Drop (Gap 0.010 in)

Pressure Drop (Gap 0.015 in)

Pressure Drop (Gap 0.020 in)

Reynolds 

Number

Gap = 0.020 in

Gap = 0.015 in

Gap = 0.010 in Inlet Water Temp. = 300 K

Avg. Friction Factor: f = 84/Re

Figure 16. Calorimeter power and pressure drop variation with flow rate.

In this figure, both the heat transfer rate (m& Cp∆T) and pressure drop (∆p) are independent of the 

convective coupling of the calorimeter channel wall, as the figure is intended to illustrate the power capability and regime of the flowing water in the film annulus over a wide operating range. Three film flow gap 

sizes are analyzed as a function of water flow rate, showing the pressure loss penalty for smaller gaps. The 

calorimeter pressure drop is a function of fluid properties, Reynolds number, and calorimeter flow path 

geometry and is reduced by a factor of ≈8 by doubling the film gap width.

The desired operating condition for the heat pipe calorimeter is a combination of low-film flow (∆T) 

avoiding boiling point limits and low-pressure drop minimizing water system pumping requirements; however, these conditions must be balanced with the forced convection heat transfer coupling of the calorimeter 

channel wall temperature to water film flow. This coupling is governed by two primary factors. First, the coupling is affected by the convective heat transfer coefficient (h) that is a strong function of the film gap width, 

equations (17) and (21) as the heat transfer coefficient benefits from the smallest gap size possible (the opposite of pressure drop). Second, the coupling is affected by the temperature difference between the calorimeter 

channel wall, high-temperature source, and a mean film flow temperature, the low-temperature sink, used in 

evaluation of the overall convective heat transfer.
28

To examine this sensitivity, figures 17 and 18 illustrate the required calorimeter channel wall 

temperatures for three film gap widths at three fixed cooling film flow ∆Ts. The cooling system in the 

planned heat pipe tests will use pressurized water at 65 psia. At this pressure, the boiling point of water 

is 420 K. Figures 17 and 18 clearly show that for a given film flow rate and power transfer combination, 

the smaller film flow gap provides a lower required calorimeter channel wall temperature, but has a 

corresponding increase in pressure drop. For case F(–4), which is the most demanding of the heat pipe 

tests and requires that 5,000 W be transferred to the fluid, the following is observed:

• For a small film flow gap width of 0.01 in, film flow (∆T) of 20 K and a flow rate of 0.95 gal/min, at a 

pressure drop of 19 psi (fig. 16), the calorimeter wall temperature is 356 K (fig. 17), with a cooling film 

flow outlet temperature of 320 K. This condition requires a large coolant driving pressure.

• Increasing the film flow (∆T) to 60 K with the same 0.01-in film gap width and a corresponding flow rate of 

0.32 gal/min, at a pressure drop of 6.3 psi (fig. 16) the calorimeter wall temperature is 395 K (fig. 18), with 

a cooling film flow outlet temperature of 360 K. This condition yields an acceptable pressure loss for the 

water-cooling system, however, it is approaching the boiling limit.

The tradeoff in selecting the width of the flow gap is an increase in calorimeter pressure drop for 

a reduction in calorimeter channel wall temperature. The final calorimeter geometry that is selected will 

balance the pumping system requirements with sufficient margin on the hot-side temperature to maintain 

a condition that is comfortably below the water boiling point. The effect of increasing the calorimeter film 

annulus reduces the calorimeter pressure drop; opening up the channel improves the ability to machine and 

assemble the unit. However, it also reduces the overall convective heat transfer coefficient (h), which in turn 

requires a higher calorimeter channel-wall temperature. For the case F(–4) with a larger film flow gap, the 

following is observed:

• For an intermediate film flow annulus of 0.015 in, film flow (∆T) of 20 K and flow rate of 0.95 gal/min, 

the required calorimeter wall temperature is 375 K (fig. 17), nearly 20 K higher than with the smaller film 

annulus. This temperature increase is due to the lower convective heat transfer coefficient of 4,174 W/m-K; 

however, the pressure drop falls to 5.6 psi, which is now in an acceptable range.

• Increasing the film flow (∆T) to 60 K for the same 0.015-in film gap (fig. 18) results in a calorimeter 

channel-wall temperature of 416 K, which is unacceptably high and nearly at the boiling point of the 

pressurized water system (420 K). The calorimeter pressure drop, however, falls to 2.1 psi.

• Increasing the film flow annulus to a 0.02-in width with a film flow (∆T) of 20 K (fig. 17) results in a 

calorimeter channel-wall temperature of 390 K and a corresponding pressure drop of 1.9 psi. This is an 

acceptable pressure drop while still allowing some margin on the boiling point.
29

6,000

5,000

4,000

3,000

2,000

1,000

0

420

350

280

210

140

70

0

Power (W)

Temperature (K)

Flow Rate (gal/min)

0 0.2 0.4 0.6 0.8 1 1.2

Power (∆T Film Flow = 20 K)

Cal. Channel Hot-Side Required Temp. (Gap 0.010 in)

Cal. Channel Hot-Side Required Temp. (Gap 0.015 in)

Cal. Channel Hot-Side Required Temp. (Gap 0.020 in)

∆T (Film Flow Outlet to Inlet Temperature)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.010 in)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.015 in)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.020 in)

Inlet Water Temperature = 300 K

Film Flow ∆T = 20 K

Convective Transfer Coeff. = 6,247 W/m-K (Gap 0.010 in)

Convective Transfer Coeff. = 4,174 W/m-K (Gap 0.015 in)

Convective Transfer Coeff. = 3,138 W/m-K (Gap 0.020 in)

Figure 17. Calorimeter temperatures and film annulus sensitivities (∆T=20 K).

18,000

16,000

14,000

12,000

10,000

8,000

6,000

4,000

2,000

0

540

480

420

360

300

240

180

120

60

0

Power (W)

Temperature (K)

Flow Rate (gal/min)

0 0.2 0.4 0.6 0.8 1 1.2

Power (∆T Film Flow = 60 K)

Cal. Channel Hot-Side Required Temp. (Gap 0.010 in)

Cal. Channel Hot-Side Required Temp. (Gap 0.015 in)

Cal. Channel Hot-Side Required Temp. (Gap 0.020 in)

∆T (Film Flow Outlet to Inlet Temperature)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.010 in)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.015 in)

∆T (Cal. Chan. Wall to Film Flow Outlet (Gap 0.020 in) Inlet Water Temperature = 300 K

Film Flow ∆T = 60 K

Convective Transfer Coeff. = 6,247 W/m-K (Gap 0.010 in)

Convective Transfer Coeff. = 4,174 W/m-K (Gap 0.015 in)

Convective Transfer Coeff. = 3,138 W/m-K (Gap 0.020 in)

Figure 18. Calorimeter temperatures and film annulus sensitivities (∆T=60 K).
30

The estimation model applied target calorimeter channel temperatures to provide overall performance given the cooling film flow gap width and water inlet temperature. Specific calorimeter results 

producing feasible designs for the three film flow gap widths of 0.010, 0.015, and 0.020 in are shown in 

tables 6–8. To ease manufacturing and assembly, the 0.015-in gap width is strongly recommended. It is 

noted that there is a workable 0.020-in gap design for all tests cases, with the exception of case F(–4), if 

the Reynolds number is to be maintained below 2,300. The results of the spreadsheet model for each test 

case are provided in appendix F. The calorimeter coolant return annulus is sized with a 0.125-in width to 

minimize pressure loss and to simplify fabrication and assembly; it is not analyzed by this model.

Table 6. Calorimeter parameters for a film gap width of 0.010 in.

Film cooling water inlet temperature

Film gap width

Calorimeter heat extraction length

Calorimeter total length

300 K

0.010 in

9.25 in

13 in

Test Condition Units G Series F(–4) F(–3) F(–2) F(–1) F(1) F(2) F(3) F(4) F(0)

Power

Heat pipe to calorimeter gas 

 gap width

Heat pipe temperature

Calorimeter channel wall temp.

Log mean temp. diff. (wall to fluid)

Cooling film flow ∆T (outlet-Inlet)

Cooling flow outlet temperature

Cal. film flow ID

Cal. film flow OD

Water film flow rate

Heat transfer coefficient

Reynolds No.

Calorimeter pressure drop

W

in

K

K

K

K

K

in

in

gal/min

W/m2-K

–

psi

3,000

0.026

1,273

350

33.1

29.7

329.5

0.796

0.816

0.38

6,071

1,174

7.5

5,000

0.025

1,273

385

54.1

53.5

353.3

0.794

0.814

0.35

6,211

1,345

5.3

1,000

0.104

1,273

328

9.3

26.6

326.4

0.953

0.973

0.14

6,046

354.6

2.6

3,000

0.031

1,373

350

32.7

30.4

330.2

0.807

0.827

0.37

6,075

1,139

7.2

3,000

0.021

1,173

350

33.6

29.1

328.9

0.786

0.806

0.39

6,067

1,205

7.8

2,000

0.037

1,223

335

21.6

23.1

322.9

0.819

0.839

0.33

6,027

916.26

6.9

4,000

0.017

1,223

365

44.9

35.8

335.6

0.778

0.796

0.42

6,110

1,407

7.8

2,000

0.046

1,323

335

21.2

23.7

323.5

0.837

0.857

0.32

6,030

878

6.5

4,000

0.02

1,323

365

44.5

36.5

336.3

0.785

0.805

0.42

6,113

1,378

7.5

3,000

0.026

1,273

350

33.1

29.7

329.5

0.796

0.816

0.38

6,071

1,174

7.5

Table 7. Calorimeter parameters for a film gap width of 0.015 in.

Film cooling water inlet temperature

Film gap width

Calorimeter heat extraction length

Calorimeter total length

300 K

0.015 in

9.25 in

13 in

Test Condition Units G Series F(–4) F(–3) F(–2) F(–1) F(1) F(2) F(3) F(4) F(0)

Power

Heat pipe to calorimeter gas 

 gap width

Heat pipe temperature

Calorimeter channel wall temp.

Log mean temp. diff. (wall to fluid)

Cooling film flow ∆T (outlet-Inlet)

Cooling flow outlet temperature

Cal. film flow ID

Cal. film flow OD

Water film flow rate

Heat transfer coefficient

Reynolds No.

Calorimeter pressure drop

W

in

K

K

K

K

K

in

in

gal/min

W/m2-K

–

psi

3,000

0.025

1,273

360

50.2

18.8

318.6

0.795

0.825

0.6

4,009

1,652

3.7

5,000

0.025

1,273

396

83.3

24.6

324.4

0.794

0.824

0.77

4,035

2,234

4

1,000

0.103

1,273

338

13.8

35.2

335

0.951

0.981

0.11

4,075

290

0.5

3,000

0.031

1,373

360

49.5

20.1

319.9

0.806

0.836

0.56

4,015

1,543

3.3

3,000

0.021

1,173

360

50.9

17.6

317.4

0.786

0.816

0.65

4,004

1,760

4

2,000

0.037

1,223

340

32.7

13.9

313.8

0.818

0.848

0.54

3,986

1,369

3.5

4,000

0.017

1,223

380

68.2

22.8

322.6

0.777

0.807

0.67

4,028

1,936

3.8

2,000

0.046

1,323

340

32

15.2

315

0.836

0.866

0.5

3,991

1,244

3.1

4,000

0.02

1,323

380

67.5

24

323.8

0.785

0.815

0.63

4,033

1,842

3.5

3,000

0.025

1,273

360

50.2

18.8

318.6

0.795

0.825

0.6

4,009

1,652

3.7
31

Table 8. Calorimeter parameters for a film gap width of 0.020 in.

Film cooling water inlet temperature

Film gap width

Calorimeter heat extraction length

Calorimeter total length

300 K

0.020 in

9.25 in

13 in

Test Condition Units G Series F(–4) F(–3) F(–2) F(–1) F(1) F(2) F(3) F(4) F(0)

Power

Heat pipe to calorimeter gas 

 gap width

Heat pipe temperature

Calorimeter channel wall temp.

Log mean temp. diff. (wall to fluid)

Cooling film flow ∆T (outlet-Inlet)

Cooling flow outlet temperature

Cal. film flow ID

Cal. film flow OD

Water film flow rate

Heat transfer coefficient

Reynolds No.

Calorimeter pressure drop

W

in

K

K

K

K

K

in

in

gal/min

W/m2-K

–

psi

3,000

0.025

1,273

375

67.1

15.6

315.4

0.795

0.835

0.73

3,003

1,909

1.8

5,000

0.024

1,273

420

112

16.1

315.9

0.793

0.833

1.17

3,005

3,106

2.6

1,000

0.103

1,273

337

18.4

29.8

329.6

0.951

0.991

0.13

3,046

324

0.3

3,000

0.03

1,373

375

66

17.6

317.4

0.806

0.846

0.65

3,010

1,710

1.6

3,000

0.021

1,173

375

68

13.9

313.7

0.785

0.825

0.82

2,998

2,132

2.1

2,000

0.037

1,223

360

43.6

12.5

312.3

0.818

0.858

0.6

2,992

1,492

1.6

4,000

0.016

1,223

401

91.1

19.6

319.4

0.777

0.817

0.77

3,018

2,166

1.8

2,000

0.046

1,323

350

42.6

14.4

314.2

0.836

0.876

0.53

2,998

1,299

1.4

4,000

0.02

1,323

400

90.3

19.1

319

0.784

0.824

0.79

3,016

2,185

1.8

3,000

0.025

1,273

375

67.1

15.6

315.4

0.795

0.835

0.73

3,003

1,909

1.8

4.2.3 Environmental Losses

To establish good thermal coupling between the heat pipe and calorimeter assembly, the proposed 

test setup requires that the test chamber be flooded with an inert gas (He or He-Ar mixture). The downside 

is that the gas also provides better thermal coupling to the chamber wall, introducing additional losses 

that must be compensated for by the inductive heater system. Figure 19 illustrates the locations of heat 

loss to the environment and support system. At the evaporator section, there is a combination of radiation, 

conduction, and natural convection heat loss to the RF inductive coil and chamber walls. The surface of 

the calorimeter has a natural convection loss component to the test chamber environment (small, due to its 

low temperature, but it must be assessed). The condenser support has a combination of solid conduction 

along the length of the support tube and gas conduction and radiation between the surface of the support 

tube and calorimeter inner wall. 

Qsolid cond Qgas cond

Calorimeter Qrad

4 in 13 in 1 in

3 in

Qrad Qconv

Qgas cond

Qrad

Qconv Inductive Heater 

Coil Assembly

Heat Pipe Closeout Stem

Figure 19. Heat pipe/calorimeter environmental heat loss.
32

A simple spreadsheet model was assembled to estimate the losses; these calculations will be incorporated into the final GFSSP model (sec. 4.3). The approach makes use of constant surface temperature 

for the heat pipe, calorimeter outer surface, mounting bracket, test chamber wall, and RF inductor coil 

(taken from the cases listed in table 7 for a 0.015-in film flow gap). The current assumption is that the test 

chamber wall will be water-cooled to maintain a temperature of ≈300 K. The RF inductor coil will also be 

water-cooled, and for this analysis, the coil temperature is assumed to be 350 K, due to its close proximity 

to the heat pipe evaporator. The heat pipe and calorimeter mounting bracket is assumed to operate at a temperature of 400 K—slightly above the test chamber wall temperature; no cooling of the bracket is currently 

planned. The test chamber is to be held at a pressure of ≈70 torr.

Radiation heat transfer is calculated using the same expression as equation (29) and gas-conduction 

heat transfer, equation (30), with the appropriate values to describe fluid/material properties and boundary 

temperatures. Free convection was assessed as natural convection from a long cylinder using equations 

(29–34), where Ra is the Raleigh number, Pr is the Prandtl number, do is the outer diameter of the component, β is the gas thermal expansion coefficient, α is the gas thermal diffusivity, ρ is the gas density, Cp is 

the gas specific heat, and ν is the gas kinematic viscosity as follows:

Q& = hA T − T ( ∞ ) , (29)

h = k

do

Nu = k

do

0.6 +

0.387 × Ra

0.167

1+ (0.559 Pr) ⎡ 0.562

⎣⎢ ⎤

⎦⎥

0.296

⎡

⎣

⎢

⎢

⎢

⎢

⎤

⎦

⎥

⎥

⎥

⎥

2

(30)

Ra =

gβ T − T ( ∞ ) do

3

vα

10−5 < Ra < 1012 ( ) , (31)

Pr = v

α , (32)

α = k

ρC

p

, (33)

β ≈ 1

T

avg

. (34)

Material and fluid properties are assessed at an average temperature of the hot and cool surfaces 

involved. The results of the heat loss calculations for each of the heat pipe test conditions are provided in 

tables 9 and 10 for pure He and He-32%Ar, respectively. During nominal heat pipe operation, thermal loss 

rates are a function of the surface temperatures; at the maximum temperature of 1,373 K, test case F(–2), 

the heat loss is 890 W with pure He and 560 W with a He-32%Ar mixture. Appendix G shows a complete 

listing of the spreadsheet.
33

Table 9. Heat pipe and calorimeter thermal loss to environment for pure He (70 torr).

Test Condition Units G-Series F(–4) F(–3) F(–2) F(–1) F(1) F(2) F(3) F(4) F(0)

Variables

Heat pipe power level

Heat pipe temperature

Environment temperature

Avg. calorimeter temperature

Avg. RF inductive coil temp.

Support bracket temperature

Gas composition factor (1=pure He)

W

K

K

K

K

K

–

3,000

1,273

300

360

350

400

1

5,000

1,273

300

398

350

400

1

1,000

1,273

300

338

350

400

1

3,000

1,373

300

360

350

400

1

3,000

1,173

300

360

350

400

1

2,000

1,223

300

340

350

400

1

4,000

1,223

300

380

350

400

1

2,000

1,323

300

380

350

400

1

4,000

1,323

300

340

350

400

1

3,000

1,273

300

360

350

400

1

Condenser Support Losses

Solid conduction loss

Radiation loss

Gas conduction loss

W

W

W

12

14

284

12

14

267

12

14

293

14

19

323

10

11

247

11

12

274

11

12

257

13

16

295

13

17

312

12

14

384

Calorimeter Surface Losses

Natural convection W 6 6 5 7 6 3 8 3 8 6

Evaporator Surface Losses

Radiation loss

Gas conduction loss

Natural convection loss

W

W

W

120

233

74

120

233

74

120

233

74

172

270

85

82

199

65

99

216

70

99

216

70

144

252

79

144

252

79

120

233

74

Total Losses to Environment W 745 728 753 890 619 685 673 803 825 745

Table 10. Heat pipe and calorimeter thermal loss to environment for He-32%Ar (70 torr).

Test Condition Units G-Series F(–4) F(–3) F(–2) F(–1) F(1) F(2) F(3) F(4) F(0)

Variables

Heat pipe power level

Heat pipe temperature

Environment temperature

Avg. calorimeter temperature

Avg. RF inductive coil temp.

Support bracket temperature

Gas composition factor 

 (1=pure He)

W

K

K

K

K

K

–

3,000

1,273

300

360

350

400

0.51

5,000

1,273

300

398

350

400

0.51

1,000

1,273

300

338

350

400

0.51

3,000

1,373

300

360

350

400

0.51

3,000

1,173

300

360

350

400

0.51

2,000

1,223

300

340

350

400

0.51

4,000

1,223

300

380

350

400

0.51

2,000

1,323

300

380

350

400

0.51

4,000

1,323

300

340

350

400

0.51

3,000

1,273

300

360

350

400

0.51

Condenser Support Losses

Solid conduction loss

Radiation loss

Gas conduction loss

W

W

W

12

14

145

12

14

136

12

14

150

14

19

165

10

11

126

11

12

140

11

12

131

13

16

150

13

17

159

12

14

145

Calorimeter Surface Losses

Natural convection W 4 3 3 4 4 2 5 2 5 4

Evaporator Surface Losses

Radiation loss

Gas conduction loss

Natural convection loss

W

W

W

120

119

43

120

119

43

120

119

43

172

138

48

82

101

37

99

110

40

99

110

40

144

128

45

144

128

45

120

119

43

Total Losses to Environment W 457 448 461 560 370 414 408 499 511 457
34

The gas gap at both the evaporator/RF coil and condenser closeout stem have a concentric 

arrangement; however, the gap diameter is small, minimizing the potential for natural convection currents. 

For natural convection in an annulus, the total heat transfer can be assessed using an effective thermal 

conductivity (ke) in the following Fourier heat transfer equation:

Q& = 2πke

l

ln do di ( ) Ti − To ( ) . (35)

The ke term can be evaluated based on the annulus inside and outside dimensions and gap width (L) as follows:

ke = k 0.386 Pr

0.861+ Pr

⎛

⎝

⎜ ⎞

⎠

⎟

1/4

Racyl

* ( ) ⎛ 1/4

⎝

⎜

⎜

⎞

⎠

⎟

⎟ , (36)

Racyl

* = ln do di ( ) ⎡

⎣ ⎤⎦

4

δ3 di

−3/5 + do−3/5 ( )

5

Raδ 102 < Racyl

* < 107 ( ) , (37)

Raδ = gβ T − T ( ∞ )δ3

vα , (38)

and

δ = 1

2

do − di ( ) . (39)

For the annular gap (δ) dimensions in the heat pipe calorimeter/RF inductor geometry, the 

calculated value of Rac* is two orders of magnitude below its valid range of 100 < Rac* < 107 and 

therefore not applicable. While equations (35) and (36) are not strictly applicable for the conditions here, 

they do indicate that considering gas conduction alone is an acceptable approach for the small concentric 

components proposed, since the natural convection heat transfer rate yielded by equation (35) produces a 

rate approximately one-fifth that of the gas condition.

4.3 The Generalized Fluid System Simulation Program Model

The GFSSP is a general-purpose computer program for analyzing steady state and time-dependent 

flow rates, pressures, temperatures, and concentrations in a complex flow network.17 The program is 

capable of modeling phase changes, compressibility, mixture thermodynamics, conjugate heat transfer, 

pumps, compressors, and external body forces (e.g., gravity and centrifugal force). Three thermodynamic 

property programs, GASP, WASP, and GASPAKTM are integrated with GFSSP to provide real fluid 

thermodynamic and thermophysical properties for 35 fluids.18–20 Eighteen different resistance and source 

options are available. The flexibility exists to add additional resistance/source options as well as fluid 

properties via user-defined subroutines and lookup tables. The user efficiency of GFSSP is enhanced through 

the incorporation of the visual thermofluid analyzer of systems and components (VTASC) graphical user 

interface. Code validation and verification (V&V) is described in appendix H.
35

A model, currently under development, will provide a tool to simulate the performance of the 

coupled heat pipe and calorimeter configuration. Due to time constraints and the preliminary nature of the 

design phase, the bulk of the analytical work was done via correlations incorporated into spread-sheets. 

The additional fidelity required for design verification and pretest performance predictions is currently 

being incorporated into a GFSSP model. A preliminary nodal layout, representing an earlier version of 

the calorimeter system, is illustrated in figure 20. The existing model treats the heat pipe as a collection of 

ambient nodes at constant temperature. Nodes 41, 1, 4, and 19 are the heat pipe nodes. Node 41 is located in 

the evaporator section, which loses heat to the chamber wall by radiation and natural convection. Nodes 1, 

4, and 19 are located within the calorimeter and transfer heat to the incoming water through an annular 

gap of He and He-Ar mixture and calorimeter channel tube. The gas gap is modeled with two gas nodes 

in the radial direction at five axial stations. The calorimeter water flow path and shell assembly is also 

modeled with a similar series of nodes in both the radial and axial directions. The heat transfer coefficient 

at the surface of the calorimeter was calculated using the Dittus-Boelter equation.21 The thermodynamic 

properties of water were supplied by the WASP computer program, while He properties were supplied by 

the GASPAK computer program. Properties for the calorimeter material were obtained from Incropera.13

SS

FS

13 14 15 16 17 2K

2K

2K

3 6 21 24 27

2 33 35 37 39

31 32 34 36 38

42 30 5 20 23 26

19 22 25 28

12 11 10 9 8

18

7

29

40

41 1 4

43

HTC-R4243

1213

1314

4142

SA

HTC-R

313 FS

K

23 SS

K 336

532

130 45

3233

3521

3435

2034

1920

1922

2115

1516 1617 1718

1112 1011

2829

910 89 78

K

SS

K

SS

K

SS

K

SS

K

SS

K

SS

K

SS

K

SS

K

SS

2336

2223

2225 2526

2638

3839

3927

2717

3637

3724

2416

K

SS

K

SS

K

SS

K

SS

K

SS

K

SS

HTC-R

SA

HTC-R

SA

HTC-R

SA HTC-R

SA

HTC-R

SA

HTC-R

SA

K

312 SS

K

3031 SS

HTC-R

614

HTC-R

FS

HTC-R

FS

HTC-R

SA

HTC-R

FS

Figure 20. GFSSP model of heat pipe test setup.

As is typical with most analytical efforts, a series of iterations will be required to refine the model 

for efficient and accurate use in assessing the performance of the heat pipe and calorimeter assembly. Once 

the GFSSP model is completed and the output is checked against analytical and spreadsheet estimates, 

it should prove to be a useful tool in performing quick turnaround parametric computational analyses 

to support final system design and pretest operations. The model will be updated once test data become 

available.
36

5. OTHER CONSIDERATIONS

5.1 Sodium Working Fluid Sampling (Vanadium Wire Technique)

Interest has been expressed in sampling the Na working fluid after it has been transferred into the 

refractory metal heat pipe unit. This provides insight into the initial oxygen concentration within the unit—the 

sum of impurities in the dispensed Na and what is picked up from the heat pipe materials. A potential 

method to assess the oxygen concentration is based on a V wire equilibration technique (ASTM 997–83).22

Details of this technique are covered in NASA/TM—2005–213902.

23 Prior to incorporating this technique 

into the heat pipe processing sequence, laboratory testing must be performed to confirm the accuracy 

of the method. Sample testing can be performed in the MSFC alkali metal handling glove box with the 

current Na supply that will be used to fill the heat pipes discussed in this TP. To validate the V wire technique, 

detailed engineering will be required to accommodate the sampling hardware in the heat pipe fill stem. The 

key in this design will be to devise an operation in which the V wire sample can be removed from the heat 

pipe without introducing impurities. Additional support hardware required for these operations, such as 

the equilibration furnace and valve cooling, will have to be incorporated into the glove box system. 

The V wire technique was successfully applied in Na loops in the EBR–II Program.24 A V wire, 

present in the heat pipe during processing, appears suitable to characterize oxygen levels in fully assembled 

and filled heat pipe modules. A method has been devised to remove the wire from the Na during closeout 

with negligible impurity introduction. The validity of this technique can be independently confirmed 

by analysis of sample heat pipe modules. 

Figure 21 depicts one possible sequence of steps. Sodium is introduced into the heat pipe through 

a fill stem containing a 0.25-mm-diameter by 3-cm-long V wire (step A). The heat pipe is then closed and 

wetted over a 48-hr period at the core design temperature.

The heat pipe is then inverted and heated to immerse the V wire in the Na (step B). The assembly 

is then brought to temperature for a time sufficient to reach equilibrium oxygen concentration between 

the Na and the wire. Figure 22 shows a time-temperature relation for two wire diameters, assuming that 

oxygen diffusion through the V is the rate-limiting mechanism.

The heat pipe is cooled to room temperature and the fill stem containing the V wire is severed 

from the heat pipe inside an Ar dry box (step C). The wire and an Na sample are extracted from the fill 

stem for analysis. A cap is attached to the heat pipe fill stem using an EB welder connected to the dry box 

(step D). Since the V wire further purifies the Na during equilibration, some correction will be necessary 

to establish the true oxygen concentration in the working fluid before and after equilibration.
37

Heat Pipe

V Wire

Na Fill

A

Na

B C D

Heat Pipe

Sample Cap

Heat Pipe

Figure 21. Concept to measure working fluid oxygen concentration 

in Na-filled heat pipes.

1,000

100

10

1

t (hr)

0.85 0.9 0.95 1 1.05 1.1 1.15

900 800 700 600

T–1 × 10–3 (K–1)

T (°C)

0.5-mm Diam.

0.25-mm Diam.

Figure 22. Estimated time required for equilibrium oxygen concentrations 

in static Na immersed for two V wire diameters.
38

5.2 Heat Pipe Closeout and Support Concept

Closeout of the heat pipe fill stem is required once Na loading and purity measurements are 

completed; this will require a number of welding operations to seal, protect, and provide a support 

mechanism for mounting the heat pipe to the calorimeter. One possible fill stem closeout concept, shown 

in figure 23, requires that the fill stem be crimped near the condenser plug and then sealed using a tungsten 

inert gas (TIG) welder to make a cut/weld pass across the center of the crimp. The sealed fill-stem stub 

can then be protected by placing a plug and tube cover over it and welding it to the condenser plug using 

an EB welder. The EB technique provides a vacuum in the sealed volume surrounding the crimped fill 

stem. The evacuated closeout tube provides a containment enclosure should the TIG seal weld leak; since 

it is evacuated, no external trapped purge gas would be introduced into the heat pipe volume affecting its 

operation. To support the heat pipe, an additional closeout support tube can be EB welded to the closeout 

plug. The length of the support tube will need to be sized to provide sufficient room for assembly and to 

minimize the heat loss through material conduction. Alignment fixtures will be required to assemble the 

closeout and support components so that straightness can be maintained, a key parameter for integrating the 

heat pipe with the calorimeter, so the two components do not touch. 

Figure 23. Method 1 to close out the heat pipe fill stem and provide mounting support.
39

Another possible method to close out the heat pipe fill stem is illustrated in figure 24. After the 

heat pipe is evacuated, it is heated to a temperature to melt all the Na and then positioned with the fill stem 

down, transferring the Na to fill the stem. The heat pipe is then cooled, solidifying the Na in the fill stem 

and creating a freeze plug. The fill stem is then severed using a cutting tool ≈0.25 in from the condenser 

plug; the freeze plug protects the heat pipe internal vacuum (prevents leakage of glove box inert gas 

into the heat pipe). The fill stub is deburred and a stem cap placed over the stub (tight tolerance so that it 

remains snug to the fill stem). The heat pipe unit is then placed in a welding fixture and bagged to maintain 

an inert environment around the unit while it is transported from the glove box to an EB welder. Welding 

is performed in a vacuum, thus removing any purge gas remaining in the stem cap. The capped fill stem 

stub is protected in a fashion similar to that described previously, making use of an evacuated plug and 

tube cover (welded to the condenser plug) to provide an additional containment enclosure. Mounting is 

accomplished by an additional support tube EB welded to the closeout plug, with a length sized to meet 

heat loss and support setup requirements. 

Figure 24. Method 2 to close out the heat pipe fill stem and provide mounting support.
40

5.3 Heat Pipe Radio Frequency Inductive Heating Coil

Induction heating has been selected to provide a well-characterized mass transfer boundary at the 

heat pipe evaporator. Fifteen of sixteen life test heat pipes are inductively heated in three clusters with 

each cluster containing five positions. Table 11 lists the heat pipes that will initially comprise each cluster. 

A water-cooled copper bus bar, shaped like an inverted patriarch cross, forms a pentagonal pattern around 

which each induction coil is mounted, as shown in figure 25. The cross arrangement minimizes inductive interference among positions. Each induction coil can be affixed to or removed from its bus bar with 

mounting pins or screws in a flexible modular fashion. Assessment of the final system design will require 

determination of the evaporator inductor coil RF coupling to the condenser calorimeter end plate. The 

objective is to minimize potential RF heating of the calorimeter, which would skew the power data. In 

addition, RF coupling to any metallic shielding placed around the evaporator section (serving as isolation 

shield) must be assessed.

Table 11. Initial life test heat pipe arrangement.

Position/Cluster 0 1 2 3

1

2

3

4

5

F5c F1c

F2b

F2d

F4b

F4d

F3a

F3c

F3e

G6

G7

G1

G2

G3

G4

G5

Figure 25. Induction coil cluster mounted on five-position bus bar (end view).

Clusters 0 and 1 hold only F-series heat pipes that operate for the duration of the test. The heat pipe 

on cluster 0 operates at 5 kW and is inductively heated on a separate electrical circuit. This single-position 

circuit also serves to determine the performance limits of individual heat pipes before the life test begins. 

Cluster 1 holds heat pipes that require power to be supplied at 1, 2, and 4 kW each (powers of two).

Cluster 2 initially holds F-series and long-duration G-series heat pipes. Cluster 3 initially holds the 

G-series heat pipes that will be first removed for destructive test. Grouping all three kilowatt heat pipes on 

clusters 2 and 3 allows a uniform electrical potential to be applied to both clusters. As G-series heat pipes 

are removed, the configuration of clusters 2 and 3 will be rearranged to help balance the electrical load.
41

5.4 Chamber Gas Selection and Purity

For operation of refractory metal systems at increased temperature for a relatively long time, the 

concentration of oxygen in the ambient environment is critical to the life of the refractory system. For 

the proposed testing, the refractory metal components will be operated in a low-pressure environment of 

ultrahigh purity (UHP) Ar or He so that the thermal coupling between the heat pipe and calorimeter can 

be controlled. To achieve the desired gas purities, successive dilutions and pumpdown of the system are 

required. The pumpdown procedure will also include an initial bakeout of the system at ≈525 K to drive 

off water vapor (and other volatiles) from the gas lines, vacuum chamber, and test components. 

A series of rough calculations was performed to assess the oxygen concentration in the heat pipe 

test chambers over successive dilutions with UHP fill gas, assuming an initial oxygen concentration of 

209,500 ppm in air filling the test chambers (20.95% oxygen by volume). Calculations were performed 

assuming three cylindrical test chambers having approximate dimensions of 24 in diameter by 24 in length. 

The UHP He and Ar, purchased from Sexton Supply, have a guaranteed minimum purity of 99.999%. The 

UHP He has a maximum oxygen content of 3 ppm by volume, and UHP Ar has a maximum oxygen content of 1 ppm by volume. Figure 26 shows the concentration of oxygen in the test chambers in ppm (by 

volume) at a pressure of 70 torr (≈0.1 atm) after multiple dilutions of He or Ar. 

1×105

1×104

1×103

1×102

1×101

1×10

1×10–1

1×10–2

1×10 O

–3 2 Concentration (ppmv) at 70 torr

0 1 2 3 4 5 6 7 8 9 10 11 12

Number of Dilutions

1×105

1×104

1×103

1×102

1×101

1×10

1×10–1

1×10–2

1×10 O

–3 2 Concentration (ppmv) at 70 torr

0 1 2 3 4 5 6 7 8 9 10 11 12

(a) (b) Number of Dilutions

Figure 26. Oxygen concentration in the heat pipe test chambers after successive 

dilutions of (a) UHP He and (b) UHP Ar fill gas. 

For this analysis, the chambers are pumped down from atmospheric pressure (760 torr) to 70 torr 

and then backfilled with the UHP gas to atmospheric pressure. Each subsequent dilution cycle reduces the 

oxygen concentration; pumping to a lower pressure in each cycle would further reduce the oxygen content 

over the same number of dilution cycles. After multiple dilutions (approximately six, based on fig. 26) 

the baseline oxygen concentration, established by the purity of the fill gas, is achieved; these values are 

300 ppb for He and 100 ppb for Ar. Note that while these calculations ignore any additional contamination from impurities in the lines and test chamber, they do provide an ultimate baseline for the minimum 

achievable oxygen concentration for a given fill gas and operating pressure without additional gas purification. Initial bakeout of the system (under vacuum) will assist in driving out volatile impurities (primarily 
42

water if the system is clean and degreased) from the test components and fill lines to reduce additional 

impurities in the system. 

For comparison, an acceptable vacuum level for testing Mo-Re alloys is in the 10–6 torr range; 

at a pressure (1×10–6 torr) the oxygen concentration is 0.28 ppb in the vacuum test environment. The 

oxygen concentration provided by the UHP gases alone is insufficient to achieve the required level; therefore, additional purification is required. Two methods are employed to purify the test chamber gases: 

(1) An SAES ambient temperature MicroTorr gas purifier to clean the incoming gases and (2) an SAES 

MonoTorr point-of-use purifier incorporated in the test chamber recirculation system. Figure 27 illustrates 

a conceptual layout of the of the test chamber gas purification system.

Set 16 psia

Set 10 psig

Set 20 psig

Bottle 

Regulators

Gas Mixture 

Bottle

Evacuation 

Port

0–20 psi

Bleed Valves

Set 20 psig

0–20 psi 0–4,000 psi

0–20 psi 0–4,000 psi

Set 16 psia Vac. Gauge 

0–1,000 Torr

Gas 

Pump

GP

Turbo Vacuum Pump 

and ISO Valve

Test Chamber With 

Heat Pipes

MonoTorr 

Purifier

MonoTorr 

Purifier

He Supply 

K-Bottle

Ar Supply 

K-Bottle Bypass 

Valve

FM TC PG

Figure 27. Initial concept for test chamber gas mixture system.

The MicroTorr unit has an advertised performance for purifying both He and Ar to a final oxygen 

concentration of 1 ppb. The gas mixture bottle can therefore be charged with an appropriate He/Ar mixture 

(purified to 1-ppb oxygen concentration using the inline MicroTorr unit) and, in turn, used to perform 

multiple dilution cycles on the test chamber. Using this new value of oxygen concentration in both He and 

Ar produces the trends shown in figure 28 over multiple dilution cycles. After nine dilution cycles, the final 

oxygen impurity concentration for both the He and Ar fill are 0.1 ppb at a 70-torr test chamber pressure. 

This condition meets the acceptable oxygen concentration for testing Mo-Re alloys, ≈0.28 ppb, that was 

discussed previously corresponding to the oxygen concentration in a 1×10–6 torr vacuum.
43

1×105

1×104

1×103

1×102

1×101

1×10

1×10–1

1×10–2

1×10–3

1×10–4

1×10–5

1×105

1×104

1×103

1×102

1×101

1×10

1×10–1

1×10–2

1×10–3

1×10–4

1×10 O

–5 2 Concentration (ppmv) at 70 torr

0 1 2 3 4 5 6 7 8 9 10 11 12

Number of Dilutions

O2 Concentration (ppmv) at 70 torr

0 1 2 3 4 5 6 7 8 9 10 11 12

(a) (b) Number of Dilutions

Figure 28. Final oxygen concentration in the heat pipe test chamber with (a) UHP He 

fill gas and (b) UHP Ar fill gas purified using an inline SAES MicroTorr 

ambient gas purifier. 

To maintain the purity level of the test chamber environment during test operations, an SAES 

MonoTorr point-of-use purifier with a recirculation pump will also be incorporated in the system. Within 

the MonoTorr purifier, getter materials are used to irreversibly trap gaseous impurity molecules; these 

impurities are captured on the surface of the materials, and upon heating, diffuse into the bulk of the 

getter. When used with Ar or He, the MonoTorr purifier can be used to remove molecules of H2O, O2, H2, 

CO, CO2, N2, and hydrocarbons. The performance of the purifier is dependent on the pumping speed at 

which it is operated. Table 12 provides a summary of the performance guarantee for the SAES MonoTorr 

Phase II 3000 for rare gases. 

Table 12. SAES MonoTorr Phase II 3000 performance 

guarantee for rare gases.

Impurity

Pumping Rate

0–20 slpm 20–50 slpm

O2

H2O

CO

CO2

N2

H2

CH4

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<1 ppb

<10 ppb

<10 ppb

<10 ppb

5.5 Heat Pipe Temperature Measurement

Temperature measurements are required to monitor overall operation and track performance of 

the heat pipe unit. The majority of the heat pipe unit is inaccessible due to the calorimeter and inductive 

heater assemblies. However, the primary location for a performance measurement is at the heat pipe 

evaporator exit, which is accessible through the small gap between the inductive heater and calorimeter. 
44

Two potential techniques have been identified that are capable of taking this measurement and include 

type-C thermocouples and two-band optical pyrometers. 

Thermocouples are an inexpensive method, but have a number of negative issues related to their 

usage. First, they are a contact measurement technique requiring direct attachment to the surface of the 

heat pipe. This can be troublesome since they frequently debond from the surface to which they are 

attached. In addition, to achieve a successful bond, especially in the case of refractory metals, an intermediate material such as nickel foil is typically required. The foil is bonded to the refractory metal surface 

and the thermocouple then bonded to the foil. This intermediate material is not the best option for longterm testing since nickel will diffuse into the heat pipe alloy, potentially resulting in undesirable and/or 

unknown effects. There is a possibility that the type-C thermocouples, which are available with a 26%Re 

content (the balance being tungsten), may have sufficient Re to allow them to be spot welded directly to 

the surface (testing would be required). In general, the use of spot welding is not an attractive idea since it 

can result in pitting of the heat pipe surface and potentially embed or trap impurities. Over the duration of 

the program, many spot welding cycles would be required to account for nondestructive evaluation (NDE) 

cycles, estimated at 6-mo intervals, and thermocouple repair. In addition, the thermocouple leads act as 

fins, locally cooling the location where the measurement was taken, and the long-term degradation of the 

thermocouple wires over time at high temperature introduces unknowns. If thermocouples were used, they 

would be paired at a minimum to provide backup. 

The two-band optical pyrometer is a remote mounted, noncontact measuring technique that makes 

use of a lens to focus the field of view on the location where the temperature measurement is required. The 

two-color, infrared sensing uses a ratio technique between the radiosity at two wavelengths to determine 

temperature rather than the radiosity at a single wavelength. The result is a temperature measurement that 

is independent of emissivity, target size/shape, and dust/contamination on the windows in the optical path. 

The noncontact technique simplifies setup (no internal thermocouple wires) and eases operations involving 

loading and unloading of the heat pipe units. A typical two-band system, the M77S by Mikron, retails for 

$3,000–$4,000, depending on features, and has focusing capability allowing for a spot size of 0.16 in at a 

14-in standoff.

The current recommendation is to make use of two-band optical pyrometers for all heat pipe 

measurements (total of 16 units required). This will significantly reduce the overall effort involved in 

setting up, attaching, and maintaining thermocouples inside the vacuum system. In addition, the two-band 

systems are completely external to the chamber, allowing for rapid replacement or troubleshooting should 

it be required. Thermocouples will still be used extensively on all other parts of the system such as the 

coolant flow, test chamber, etc. where they are easily accessible and operate in less severe conditions.

5.6 Water-Cooling System

Water-cooling has been baselined in the current calorimeter design and is also used to cool the 

test chamber walls and RF induction coils. The current conceptual layout for providing coolant water to 

a calorimeter cluster (a cluster is composed of five calorimeter units arranged in a pentagon) is illustrated 

in figure 29. There are two options to cool additional calorimeter clusters: (1) They can be attached to the 

system inlet and outlet manifolds if sufficient pumping/cooling power is available or (2) a duplicate of 

this system can be fabricated if necessary. The coolant system is fabricated from stainless steel and loaded
45

Facility 

Water 

System

Flow Meter

Thermocouple

Pressure Gauge/Sensor

Relief Valve

Hand Valve

Flow Switch

Temperature Switch

Heat Pipe 

Outlet Flow 

Instrumentation

Heat Pipe 

Flow Control 

Valves

Heat Exchanger Test 

Setup to Facility

Inlet Manifold

Bypass Valve

Bottle 

Regulator

Ar Supply 

K-Bottle

4655 Int 

Return

4655 Int 

Supply

FS

TC TC

PG

0–200 psi

0–200 psi 0–4,000 psi

Set 60 psig

Outlet Manifold

Bypass Valve

Inlet Isolation 

Valve

Water 

Pump

Filter

FS PG TC FM

FS TC FM

Water 

Reservoir and 

Expansion 

Tank

L

P

Test Chamber With 

Heat Pipes

FS TC FM

FS TC FM

FS TC FM

FS TC FM

FM

TC

PG

FS

Figure 29. Initial concept for water-cooling system layout.

with deionized water (target of 10 MΩ-cm or better) and an inline filter to minimize conductivity and 

the potential for buildup of scale/rust etc. The water flow to each heat pipe calorimeter is regulated by a 

hand valve, located at the calorimeter exit. Temperature and flow rate measurements are also taken at this 

location so that calorimeter power can be determined. A facility water-coolant system is used to absorb 

the estimated 60,000 W of power from the heat pipe cooling loop through the use of a heat exchanger. 

The water pump selected (if a single pump is to supply all calorimeters) will be required to provide a flow 

rate of 12–15 gpm, operate at an inlet suction pressure of 65 psia, and provide a pressure rise of at least 

20 psi to conservatively overcome plumbing losses. The facility and heat pipe cooling loops are equipped 

with flow and temperature switches to interrupt the RF generator power in the event of coolant circulation 

failure or an over-temperature condition. The system will be configured so it can be degassed (equipped 

with heating and evacuation capability) to reduce the potential for subcooled boiling. The final system 

configuration will need sufficient flow capability and capacity to remove the thermal capacity of the heat 

pipes (to a safe level) without system damage should an abrupt power-out condition be encountered. 

To control the growth of organic material in the main flow circuit (potential to plug the filter and 

small flow paths), it is recommended that an inline, ultraviolet purifier be added. This setup should allow 

for low-maintenance operation while initial degassing of the water system and the use of inert cover gas 

would reduce free oxygen, limiting the potential for ozone production. An alternate technique would 

be the use of chemicals; however, concentrations need to be monitored and chemicals replenished on a 

regular basis. Depending on chemical selection, there is the risk of interaction with the copper calorimeter 

and potential breakdown of the chemicals at high temperature. In addition, to maintain the water quality 

(10–18 MΩ-cm water resistivity), it is possible to add a deionizing (polishing) loop to the main circuit 

with a throughput of 3%–5% of the main flow.
46

6. SUMMARY

A Mo-Re alloy refractory metal heat pipe design and test matrix have been proposed that attempt 

to systematically address the issues related to corrosion rates at operating temperatures anticipated for 

possible space nuclear operations. A series of test conditions have been proposed that vary temperature 

mass fluence and power to provide for acceleration of impurity transport and corrosion mechanisms. 

Power throughput covers a range from 1,000 to 5,000 W with planned operating times up to 3 yr and 

temperature ranges from 1,173 to 1,373 K. Performance testing of the initial heat pipes will be necessary 

to determine if the high-power/temperature conditions in the test matrix can be achieved. If boiling limits 

are exceeded, the test matrix will have to be adjusted to appropriately lower power/temperature conditions.

To absorb energy from the heat pipe condenser, a gas gap/laminar flow water calorimeter concept 

appears feasible. For all power levels, the gas gap width between the heat pipe surface and calorimeter can 

be sized at 0.020 in or larger, while still retaining performance margin by adjusting the gas mixture ratio 

of He to Ar in the test chamber. Likewise, geometric solutions exist for satisfying the calorimeter film 

flow coolant annulus thermal performance. It is noted that for the higher heat pipe power cases, 4,000 and 

5,000 W, a higher calorimeter channel-wall temperature is required which reduces design margin. For all 

design cases examined, the smaller film flow annulus is desirable due to higher heat transfer coefficient; 

however, it carries a penalty of higher calorimeter pressure drop. Also, the smaller flow annulus width 

requires tight tolerances, taxing the fabrication and assembly process. For the large annulus gap examined, 

0.020 in, a workable set of operating parameters was established with low-pressure loss for all but 

the high-power, 5,000-W case. This case exceeded laminar flow water limits, requiring a calorimeter 

temperature above the local water boiling point. The recommended coolant film gap width, based on the 

current assessment, is 0.015 in. It should be pointed out that there is a potential concern in using water as 

the calorimeter coolant fluid. A water-based system presents a risk for water leakage into the test chamber 

and the potential catastrophic interaction of water with refractory metals at high temperatures. To minimize 

this risk, the calorimeter and feed lines will be an all-welded design and He leak checked; however, there is 

always the potential for a leak.

In addition, initial analysis indicates that the oxygen level (ppb) within the test chamber can be 

reduced to a level consistent with that achieved under normal vacuum testing conditions (10–6 torr range) 

typically used for Mo-Re alloys. The use of commercial purifier systems will be implemented to further 

reduce the impurity level below that achieved by successive dilutions. The test chamber environment can 

be monitored with a residual gas analyzer. Also, the use of two-band, noncontact optical pyrometers to 

measure heat pipe evaporator exit temperature is highly recommended. These units significantly simplify 

setup/maintenance operations and provide better overall data. Additional work is underway to further 

develop hardware requirements and detailed layouts for all supporting systems, including the RF heating 

system and calorimeter/test chamber water-cooling systems.
47

APPENDIX A—HEAT PIPE LIFE TESTING—ALTERNATE APPROACH

A.1 Corrosion in High-Temperature Heat Pipes

An ideal heat pipe is thermodynamically closed with energy crossing the system (container) 

boundary at the evaporator and condenser. By definition, mass does not cross the boundary of a closed 

thermodynamic system. Impurities may diffuse across the container boundary of high-temperature heat 

pipes making the system, in part, thermodynamically open. The second law of thermodynamics predicts 

the conditions under and the extent to which a chemical reaction or physical process occurs. For a given set 

of conditions, such processes proceed spontaneously and unidirectionally toward an equilibrium state of 

minimum free energy. Reaction kinetics at high temperature can overcome many reaction energy barriers 

and hasten an alkali metal heat pipe’s adjustment toward thermodynamic equilibrium.

Distillation of working fluid from the hot zone combined with diffusion and reactions among 

chemical species in the fluid and in the container accounts for impurity movement within a heat pipe. 

Solubility-induced corrosion may occur from interaction between the container and the relatively pure 

condensate in the cold zone. Heat pipe containers, selected to be insoluble in the pure working fluid to well 

under a part per million, usually are not susceptible to direct solubility corrosion. 

Nonmetallic impurities in the cold zone structure can diffuse into the pure condensate. These 

impurities react with or are dissolved in the condensate that flows toward the evaporator. Such impurities 

accumulate in the evaporator and may precipitate, plugging the wick, or may form eutectic mixtures 

with the container. Ternary compounds may also form among the impurities, condensate, and container. Eutectics and ternary compounds with low melting-point temperatures can erode the container 

or contribute to dewetting.

A.2 Elementary Scaling Variables

Corrosion in alkali metal heat pipes is related to the initial concentration of nonmetallic impurities 

in the container and fluid plus any impurity movement into the system from external sources. Pure condensate forms a concentration gradient with respect to the container that tends to transfer impurities by 

condensate flow from the cold zone toward the hot zone. The total impurity content in the condenser and 

attached structures is then a key variable influencing corrosion. The initial mass of impurities in the condenser over the mass of impurities in the evaporator is a scaling variable affecting corrosion similarity 

between heat pipe structures. For heat pipe materials with uniform impurity concentration, the volume of 

condenser to evaporator structures may serve as a proxy for the impurity content ratio as follows:

Ξ ≡ (impurity content in condenser) × (impurity content in evaporator)

–1 ≈ V

c

V

e ( )

−1 . (40)

Corrosion in high-temperature heat pipes can be inhibited with nonsoluble getters that trap 

impurities during operation. Such getters can be distributed to allow wetting by the condensate and can 
48

be part of the heat pipe container or placed in packs. Assuming a well-distributed getter, the potential for 

corrosion should relate to the total impurity concentration compared with the total stoichiometric impurity 

capacity in the getter and container as, 

Σ ≡ (impurity capacity of heat pipe getter) × (impurity content of heat pipe)–1 . (41)

Getter material in high-temperature heat pipes should be hyperstoichiometric compared with 

nonmetallic impurities from all sources. As a rule of thumb, the impurity capacity to impurity ratio should 

be at least 10:1 above stoichiometric proportions.

Most proposed mission lifetimes are greater than the time available to life test any reactor design. 

This constraint suggests the following natural time variable:

τ ≡ (duration of life test) × (design life of heat pipe)

–1 . (42)

Corrosion rates are influenced by temperature because of its exponential effect on reaction kinetics. 

Arrhenius-governed processes govern corrosion rates setting heat pipe lifetime for a set of reactions and 

temperatures. This suggests that test conditions can be normalized using an expression like,

α(T ) = exp ΔH(k)

−1 ( ) T0

−1 − T ⎡ −1

⎣⎢ ⎤ { ⎦⎥} , (43)

where ∆H is activation energy, k is Boltzmann’s constant, T0 is the design temperature, and T is the heat 

pipe test temperature. Note the numerical results presented here use generic activation energies that are 

not necessarily representative of the material systems considered. It can be argued that since testing at 

high temperature accelerates Arrhenius-governed kinetics, this same acceleration applies to heat pipe life. 

Such simplistic views must be approached with caution. The nature of chemical reactions and their 

kinetics can change dramatically with absolute temperature. Conditions that apply at one temperature do 

not necessarily hold for others. Without deeper understanding of the underlying reaction mechanisms, 

conclusions drawn from such tests can be misleading.

As working fluid flows into the evaporator, it vaporizes, concentrating the impurities and possibly 

making the corrosion rate dependent on mass fluence. The radial heat flux applied to the evaporator is 

given by q

rad = q (π d Le)–1; mass flux through the evaporator is a function of the radial heat flux, 

G = q

rad (hfg)–1. The mass fluence through the evaporator is then given by M ̋ = G τ, yielding

M′′ = q& π d Le ( )

−1

h( fg )

−1

τ . (44)

Mass fluence can easily be increased by applying power along a shortened heat pipe evaporator 

length. Variation of the mass fluence can be used to accelerate heat pipe life or to affordably test subscale 

versions of a flight unit. 

A.3 Possible Life Test Configuration

A series of accelerated life heat pipes can be designed to permit compression of life-limiting 

mechanisms from the mission duration to a period consistent with the schedule of a flight program. 
49

Condenser calorimetry, using forced convection of a cooling fluid, is required to achieve the necessary 

power density and provide realistic boundary conditions for a flight system. A life test heat pipe can be 

enclosed in a metallic chamber. Mass transfer can be minimized between the heat pipe and its surround-ings 

by several artifices. A thin-walled tube welded to the cold heat pipe condenser cover mechanically attaches 

and electrically grounds the heat pipe to an Mo holding fixture. This single point of contact creates a 

low-temperature boundary that minimizes solid-state diffusion from external impurity sources. Power is 

applied to the heat pipe evaporator through a noncontact induction heating coil. The chamber is filled with 

a noble gas mixture that slowly circulates through a hot titanium getter. The gas conducts energy across a gap 

between the heat pipe condenser and a gas- or water-cooled calorimeter. Coupling between the condenser 

and the calorimeter is controlled by composition of the gas mixture. The total pressure of the gas mixture is 

kept just high enough to avoid ionization between the induction coil and the heat pipe. Low chamber gas 

pressure tends to reduce heat pipe exposure to residual impurity concentrations in the gas mixture. Heat 

pipe surface temperature is measured with a noncontact, two-band pyrometer.

A.4 Life Test Methods

There are several ways to establish heat pipe material system life without a full-term test. 

Unfortunately, current budgetary limitations and material availability allow comparatively few tests to 

study a few key variables. Under current constraints the relevance of subscale testing must be seriously 

considered. The best available processing procedures must be used for all pipes unless the test scope is 

greatly expanded. Contamination getters should provide margins typical of flight heat pipe designs, at 

least 10:1. A heat pipe life test matrix can be developed for a few key material combinations, operating 

temperatures, and evaporator mass fluences. Boundary conditions should be imposed on the heat pipe 

to avoid random introduction of external impurities. One remotely possible outcome of a realistic, yet 

necessarily limited, life test series might be a null result that conclusively confirms the absence of corrosion 

over the conditions considered. After such an unexpected result, it is a good idea to build a few heat pipes 

and operate them to deliberately induce corrosion or failure. Life tests that compress operating time may 

operate near boiling limit margins. Achievable compression will not be known for sure until the first 

couple of pipes are performance tested.

A.5 Baseline Heat Pipe Life Test

Table 13 shows one possible approach to life testing—a few heat pipes with short evaporators 

and long condensers. A long condenser serves as a proxy for a large-condenser impurity load. A similar 

impurity load can be achieved with thickened or impurity-doped condenser walls. A short evaporator 

increases mass fluence, concentrating impurities, tends to enhance impurity-induced corrosion. The 19:1 

condenser-to-evaporator length ratio is not geometrically similar to the 2:1 ratio of the reference design.

The size of the baseline life test series reflects the limited materials now available for heat pipes 

with comparatively long condensers. The baseline heat pipes in table 13 could be tested over a 3-yr period 

in the configuration outlined in figure 30. Unfortunately, four heat pipes are insufficient to conduct a 

meaningful G68–80 series life test. One possible test method varies mass fluence and temperature above 

the design condition. Both of these variables might change during off-nominal reactor operation, and 

characterization of their functional relationship is important. To exceed the design mass fluence over the 

3-yr period, radial heat flux is concentrated up to 10× the radial heat flux of the reference design. The short
50

Table 13. Baseline life tests.

Heat 

Pipe

Wall/Wick 

Material/

Fluid Lc/Le

Di

(cm)

T

(K) a(T)

q

(W)

Q

rad

(W/cm2)

G

(kg/cm2–s)

t

(hr)

M ′′

(kg/cm2)

Design Mo-Re/Na 2 1.41 1,250 1.00 5,752 22 0.059 105,120 2,220

B(1)

B(2)

B(3)

B(4)

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

19

19

19

19

1.41

1.41

1.41

1.41

1,273

1,273

1,373

1,173

1.30

3.74

1.30

2.25

5,000

3,000

3,000

4,000

226

136

136

181

0.599

0.368

0.36

0.48

26,280

26,280

26,280

26,280

5,669

3,483

3,401

4,589

6,000

5,000

4,000

3,000

2,000

1,000

0

Mass Fluence (kg cm–2)

Temperature at Evaporator Exit (K)

B(1)

B(4)

B(3) B(2)

1,100 1,200 1,300 1,400

Baseline

Design

Figure 30. Baseline life test matrix assuming a maximum 1-μm nucleation site radius.

evaporator and high attendant evaporator radial heat fluxes of these life test heat pipes force operation 

closer to the boiling limit than the reference design. A limited-scope performance test will be required to 

demonstrate whether these high-radial heat fluxes are achievable with existing fabrication and processing 

procedures. Nondestructive testing by x-ray and ultrasound techniques can be conducted on each heat pipe 

at various intervals. At the test conclusion (after 3 yr), appropriate destructive test measurements can include 

changes in fluid composition, species distribution, and grain boundary conditions, as well as noncondensable 

gas production measured by residual gas analysis.

A.6 Alternative Heat Pipe Life Test: Option A

The baseline is by no means the only approach possible to life testing given the current constraints. 

Another approach that attempts to conserve available material, ease fabrication, provide more relevant data, 

and permit the use of more numerous, yet compact, test fixtures is described here. Instead of building four 

heat pipes with long condenser sections of uniform cross section, a series of ≈16 short heat pipes with thickwalled condensers could be constructed. These short heat pipes would have similar condenser-to-evaporator 

area ratios as the reference design. Condenser impurity loading can be imposed by condenser wall thickness. 
51

The boiling limit for the baseline design is invariant with condenser length, so the short heat pipes should 

perform identically to the longer baseline versions.

This approach potentially offers practical and theoretical benefits. Long Mo wicks are Zr brazed to 

end plugs near 2,000 K by induction heating of a vacuum-enclosed furnace susceptor. Such an operation is 

well within the current technology base, but is time consuming to perform. A simpler alternative is to braze 

end plugs to shorter wicks inside a Centorr furnace. Refractory metal tubes and rods (for gun drilling) for short 

heat pipes are easier to procure and process. Precious materials are conserved for future use. Figures 31 

and 32 show wire-mesh screen requirements for the baseline and option A approaches.

80

70

60

50

40

30

20

10

0

Length (in)

Width (in)

0 10 20 30 40 50

P

4N

1N 2N 3N

1I 2I

 

80

70

60

50

40

30

20

10

0

Length (in)

Width (in)

0 10 20 30 40 50

P

1I

2I

3G 4G

3F 4F

6G 7G

3I

1G

2G

5G

–4F –3F –2F

–1F 0F 1F

2F

Figure 31. Baseline life test wire cloth use. Figure 32. Option A life test wire cloth use.

Symbols P, N, and I represent Symbols P, F, G, and I represent

performance, baseline life, and performance, Fisher, G68–80, and

irradiation tests, respectively. irradiation tests, respectively.

Properly designed short heat pipes tested at relevant temperatures and mass fluence can be made 

geometrically similar (Lc/Le) to the reference design heat pipes. Short heat pipes form the basis for virtually 

all long-term heat pipe corrosion testing to date.7 A series of short heat pipes permits extrapolation of 

corrosion effects of a reference design from separate tests conducted over different durations. Relevant 

information becomes available within months, not years. With this approach, destructive examination of the 

first life test heat pipe can begin within 6 mo of test initiation. In addition, testing more pipes allows better 

identification of random manufacturing and processing defects.

Corrosion or reaction metrics can be taken for each test specimen and measurable results can be 

plotted versus time. NDE of wall and wick integrity by three-dimensional, x-ray computed tomography 

could be conducted on the entire series at 6-mo intervals. A plot of the corrosion rate versus time normally 

exhibits a slope that decreases with time; this will be experimentally verified. The long-term corrosion 
52

rate is extrapolated from the slope near the longest test time. This method is outlined in ASTM G68–80.4

Conditions of an extrapolated test series can closely match those of a proposed reference. Extrapolation 

tests conducted at prototypic temperatures avoid possible ambiguity in interpreting elevated temperature 

results.

Since the extrapolation is an accepted and defendable test practice, consideration should be given 

to following the ASTM G68–80 guidelines for initial screening of life test heat pipes for any material 

system seriously considered.4 If time and budget constraints limit the number of life heat pipe tests, 

extrapolation should be favored over other test methods. 

Tables 14 and 15 show a possible test series consisting of ≈16 short heat pipes. One or more 

experiments from an extrapolation test series could be used to cross-correlate fluence conditions with a 

Fisher multifactor star design test series conducted at various temperatures and fluences.3 Such life tests can 

establish and isolate corrosion trends as a function of temperature and mass fluence. Figures 33 and 34 plot 

the heat flux, fluence, and temperature relations for the combined test series.

Table 14. Option A life tests Fisher series.

Heat 

Pipe

Wall/Wick 

Material/

Fluid Lc/Le

Di

(cm)

T

(K) a(T)

q

(W)

Q

rad

(W/cm2)

G

(kg/cm2–s)

t

(hr)

M ′′

(kg/cm2)

Design Mo-Re/Na 2 1.41 1,250 1.00 5,752 22 0.059 105,120 2,233

F(–4)

F(–3)

F(–2)

F(–1)

F(0)

F(1)

F(2)

F(3)

F(4)

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

3

3

3

3

3

3

3

3

3

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1,273

1,273

1,373

1,173

1,273

1,223

1,223

1,323

1,323

1.30

1.30

3.74

0.38

1.30

0.72

0.72

2.25

2.25

5,000

1,000

3,000

3,000

3,000

2,000

4,000

2,000

4,000

226

45

136

136

136

90

181

90

181

0.599

0.120

0.368

0.351

0.360

0.237

0.474

0.243

0.485

26,280

26,280

26,280

26,280

26,280

26,280

26,280

26,280

26,280

5,669

1,134

3,483

3,321

3,401

2,241

4,481

2,295

4,589

Table 15. Option A life tests G68–80 series.

Heat 

Pipe

Wall/Wick 

Material/

Fluid Lc/Le

Di

(cm)

T

(K) a(T)

q

(W)

Q

rad

(W/cm2)

G

(kg/cm2–s)

t

(hr)

M ′′

(kg/cm2)

Design Mo-Re/Na 2 1.41 1,250 1.00 5,752 22 0.059 105,120 2,233

G–1

G–2

G–3

G–4

G–5

G–6

G–7

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

Mo-Re/Na

3

3

3

3

3

3

3

1.41

1.41

1.41

1.41

1.41

1.41

1.41

1,273

1,273

1,273

1,273

1,273

1,273

1,273

1.30

1.30

1.30

1.30

1.30

1.30

1.30

3,000

3,000

3,000

3,000

3,000

3,000

3,000

136

136

136

136

136

136

136

0.360

0.360

0.360

0.360

0.360

0.360

0.360

4,380

8,760

13,140

17,520

21,900

26,280

30,660

567

1,134

1,701

2,267

2,834

3,401

3,968
53

These test series attempt to systematically understand heat pipe corrosion behavior at several 

temperatures and mass fluences. Conditions for the Fisher test (table 14) represent a two factor 

(22+2 × 2+1) central composite configuration, with temperature and radial heat flux to the evaporator 

as the two controlled factors. Assuming corrosion is measurable and adequate and forensics are conducted, a 

multifactor response surface can be developed from this heat pipe corrosion test. In this case, results from 

the partial two-level factorial test series can identify first-order effects and interactions between the two 

factors and estimate second-order curvature effects in the data.3

6,000

5,000

4,000

3,000

2,000

1,000

0

Axial Heat Transfer Rate (W)

Temperature at Evaporator Exit (K)

Capillary

Boiling 3 μm 1 μm

–4

3

–1 –2

1

0

2

–3

4

800 1,000 1,200 1,400

Sonic

 

6,000

5,000

4,000

3,000

2,000

1,000

0

Mass Fluence (kg cm–2)

Temperature at Evaporator Exit (K)

F(–4)

F(4)

F(–2)

F(2)

F(–3)

F(1)

F(–1)

F(3)

F(0) G(7)

G(6)

G(5)

G(4)

G(3)

G(2)

G(1)

1,100 1,200 1,300 1,400

Fisher

G68–80

Design

Figure 33. Option A life test design compared Figure 34. Option A life test matrix assuming

to boiling limits at two-nucleation a maximum 1-μm nucleation

site radii. site radius.
54

APPENDIX B—HEAT PIPE WICK PROCUREMENT SPECIFICATION

B.1 MSFC IFMP Procurement Specification for Heat Pipe Wick Assembly Fabrication

Description—Manufacture complete heat pipe wick assemblies from refractory metal material. 

Quantity—18 units.

Material—Government-supplied Mo-5%Re mesh and 0.5-in-diameter Mo-44.5%Re rod stock.

Delivery—ARO 18–22 wk. 

Potential Suppliers:

• Advanced Methods and Materials

1190 Mountain View-Alviso Rd.

Suite P

Sunnyvale, CA 94089

POC: Peter Ring (408–745–7772)

• Refrac Systems

7201 W. Oakland St.

Chandler, AZ 85226–2434

POC: Norm Hubele (480–940–0068)

Source—Competitive (not planning on using a sole source).

B.2 Wick Assembly Fabrication Specifications

After final acceptance of completed wick assemblies, they will be shipped to an MSFC-specified 

contractor facility. Previous experience fabricating similar hardware components using refractory metal 

(Mo-Re alloy) mesh material is a significant technical benefit for vendors. The vendor will state relevant 

experience, procedures, etc. in their response with references of past work and the overall success of these 

past projects. Heat pipe wick assembly will be fabricated with specifications listed in section B.2.1.

B.2.1 General Specifications 

(1) The wick will be fabricated from Mo-5%Re wire cloth with a 400 × 400 mesh and nominal wire 

diameter of 0.001 in provided by NASA or its representative.
55

(2) The wire cloth will be chemically cleaned before processing to remove trace impurities such 

as iron and carbon. The chemical cleaning steps will be discussed with and approved by MSFC. A sample 

procedure is provided in section B.3. 

(3) During cleaning, the wire cloth material will also be vacuum- or hydrogen-fired (hydrogen 

dewpoint less then –77 °C), the nature of which will be discussed with and approved by MSFC. During 

heat treatments, the material will be kept within the cumulative time-temperature envelope, shown in 

section B.4, to prevent recrystallization of the Mo-5%Re material. (This includes multiple heat-treat steps 

at different temperatures.) The total number of heat treatments and temperature/time of exposures must be 

discussed with and approved by MSFC. A sample fabrication process flow is provided in section B.5.

(4) The finished wicks will be composed of seven layers of mesh material that has been drawn or 

swaged with appropriate procedures discussed with and approved by MSFC to the final dimensions as follows: 

0.481-in outside diameter and 0.453-in inside diameter (dimensional tolerance specified in fig. 35, vendor 

will discuss capabilities with MSFC). The fabricated wick length will be sufficiently long enough to allow 

material samples to be taken at various fabrication steps (as necessary) for purity measurements and cut 

to provide a number of final wick products. Estimated wick length is 12.5 in; final exact length will be 

discussed and approved by MSFC.

(5) Wick straightness will be 0.01 in/ft maximum with wick outside diameter variation not to 

exceed +0.000/–0.005 in. Ovality (diameters at any cross section) will not violate the tolerance set by the 

wick outside diameter.

(6) The wick maximum pore radius will be <35 μm once final drawing/swaging has been 

completed. Bubble-point tests with pure ethyl alcohol will be performed on each wick tube to verify pore 

radius after fabrication of the wick. Error limits for radius measurement will be within 10% of measured 

value. Any wick that does not meet the pore size requirement will be marked as defective and set aside. 

(7) The condenser end of the wick will remain open.

(8) Fabricate evaporator end plugs from Mo-44.5%Re alloy material supplied by MSFC and 

custom fit to each wick tube that passes bubble-point test criteria. The vendor will provide design options 

for the wick evaporator plug that are suitable for crimp attachment of the wick (engineering drawing in 

fig. 35 serves as a baseline). The proposed design/attachment technique will be supplied to MSFC for 

discussion and approval.

(9) The maximum pore radius of the final wick assembly (wick with attached evaporator plug) 

will be <35 μm. Bubble-point tests with pure ethyl alcohol will be performed on each final wick assembly 

to verify pore size. Any wick assembly that does not meet the pore size requirement will be marked as 

defective and set aside. The wick tube will be cleaned by hydrogen firing with a hydrogen dewpoint of less 

than –77 °C) as discussed and approved by MSFC (any heating process will conform to predetermined 

time-temperature requirements so as not to result in recrystallization of the Mo-5%Re mesh; see item (3).
56

Figure 35. Wick assembly drawing.

(10) Material chemical composition tests (measured in weight parts per million) to determine purity 

will be performed on the wick material at major manufacturing steps to assess introduction of impurities and 

quality of the final product (using a LECO elemental analyzer, GDMS, or other suitable method). Initially, 

the as-received material will be randomly sampled. Additional sampling intervals during wick development 

(first several units) include required steps (b) and (e) and the recommended but not required steps (a), (c), 

and (d) as follows:

(a) Random sampling of mesh material after chemical cleaning.

(b) Random sampling of mesh material after vacuum/hydrogen cleaning.

(c) Sample wick material after mandrel dissolution.

(d) Sample wick assembly after plug attachment.

(e) Sample final wick assembly after bubble-point testing and hydrogen-fire cleaning using 

any excess wick material on the wick assembly or any excess wick tube material from the lot that has 

accompanied the process as a witness sample. Hydrogen gas used in firing will have a dewpoint less than 

–77 °C.
57

Required sample intervals for wick production are as follows: 

(a) Sample final complete wick assembly after bubble-point testing and hydrogen-fire cleaning 

using any excess wick material on the wick assembly or any excess wick tube material from the lot that 

has accompanied the process as a witness sample. Hydrogen gas used in firing will have a dewpoint less 

than –77 °C.

The maximum acceptable and target concentrations of each impurity are listed in table 16. Error 

limits for purity measurements will be within 10% of measured value. The chemical composition will be 

documented for all testing and provided to MSFC.

Table 16. Impurity chemical composition (acceptable) final wick assembly.

Element

Maximum

Acceptable

(wppm)

Target

(wppm)

O

C

N

H

Ca

Fe

Ni

W

Si

Mg

Mn

Co

Sn

B

220

85

100

50

7

130

160

300

30

10

4

30

20

20

<25

<20

<5

<3

<7

<50

<50

<100

<10

<10

<4

<30

<20

<10

(11) Figure 35 provides an engineering drawing for the wick/wick assembly. 

(12) Wick assemblies will be cut to the specified final length prior to packaging and shipment. 

The residual wick material (2 in or more drawn from each processed assembly) will be packaged with 

appropriate identification so that it can be matched to the final wicks produced from that drawing 

process. The residual wick material will be used to document the purity of each wick assembly at an 

MSFC-designated contractor facility. 

(13) Final wick assemblies will be delivered to an MSFC-designated contractor facility. Wick 

assemblies (with cutoff residual material) will be transported inside inert gas carriers in the uncrystallized 

condition. Witness tests on the residual material will be conducted at the MSFC-designated contractor 

facility to verify wick assembly purity (LECO elemental analyzer or other suitable method) and on the 

wick assemblies to verify maximum pore size (bubble point). For final review/acceptance, the vendor will 

generate and supply to MSFC “how they did it” documentation for the assembly operation (no proprietary 

processing) with the results of purity testing, dimensional tolerance checks, bubble-point testing, cleaning, 

and final assembled weight for each completed heat pipe.
58

B.3 General Cleaning Procedure for Mo and Mo-Re Parts

(1) Wash the piece in PF solvent, an R113 substitute, until all signs of grease have been removed.

(2) Soak piece for 1–2 min in one part by volume HCl and one part by volume deionized water 

(target of 10 MΩ-cm or better) to remove residual iron surface impurities.

(3) Soak piece for 5 min in caustic cleaning solution consisting of 11 parts by volume deionized 

water (target of 10 MΩ-cm or better), 1 part by volume NaOH, and 1 part by volume H2O2. Remove piece 

from caustic bath. Replenish or replace solution as required.

(4) Wash part in hot deionized water (target of 10 MΩ-cm or better) for at least 5 min.

(5) Repeat steps (3) and (4) three times.

(6) Rinse the piece in ethanol inside an ultrasonic cleaner for 5 min.

(7) Vacuum fire or hydrogen fire part as appropriate using not-to-exceed time-temperature relation 

given in section B.4. Hydrogen gas used in firing will have a dewpoint less than –77 °C.

B.4 Time-Temperature Relation for Mo-5%Re Material

To prevent recrystallization of the Mo-5%Re wick materials, the cumulative time at temperature 

must be assessed for all expected heat treatments, as shown in table 17. For multiple heat-treat steps at 

different temperatures, the following relation will apply:

τ i / τ m Ti ( ) ⎡

⎣ ⎤ ∑ ⎦ < 1 , (45)

where τi is the time at temperature, Ti for a given step, and τm is the maximum time at temperature for the 

given step.
59

Table 17. Mo-5%Re cumulative time at temperature.

Temperature 

(K)

Maximum Time

at Temperature

(min)

1,000

1,010

1,020

1,030

1,040

1,050

1,060

1,070

1,080

1,090

1,100

1,110

1,120

1,130

1,140

1,150

1,160

1,170

1,180

1,190

1,200

1,210

1,220

1,230

1,240

1,250

1,260

1,270

1,280

1,290

1,300

2,049.0

1,448.9

1,031.6

739.3

533.2

387.0

282.6

207.6

153.3

113.9

85.1

63.9

48.2

36.5

27.9

21.3

16.4

12.7

9.8

7.7

6.0

4.7

3.7

2.9

2.3

1.9

1.5

1.2

1.0

0.8

0.6

B.5 Sample Process Flow for Heat Pipe Wick Fabrication Simplified Process Flow

(1) Inspect and lay out supplied mesh (received condition).

• Required random material chemical/purity sampling.

(2) Chemically clean supplied mesh.

(3) Vacuum-/hydrogen-fire supplied mesh.

• Required random material chemical/purity sampling of supplied mesh material 

(during the development process of first several assemblies).

(4) Fabricate mandrels and sheaths.
60

 (5) Clean mandrels and sheaths.

 (6) Mount mesh material on mandrel.

 (7) Draw.

 (8) Clean any lubricants used in process.

 (9) Remove mandrel and sheath.

(10) Bubble-point formed wick tube.

(11) Attach evaporator plug to wick tube.

(12) Bubble-point wick tube.

(13) Hydrogen-fire wick assembly.

• Required material chemical/purity sampling of extra wick material carried through 

the process.

(14) Pack unit for shipment to MSFC.
61

APPENDIX C—HEAT PIPE MATERIAL PROCUREMENT SPECIFICATION

C.1 IFMP Procurement Specification for Mo-Based Refractory Material

Description—Procure refractory material—Tube and rod stock.

Material—Refractory Mo-44.5%Re or Mo-47.5%Re material.

Quantity—Listed in specification section.

Delivery—ARO 16–20 wk. 

Potential Suppliers:

• Rhenium Alloys, Inc.

1190 Mountain View-Alviso Rd.

1329 Taylor Street

Elyria, OH 44036–0245

POC: Todd Leonhardt (440–365–7388 ext. 243)

• Rembar Compay, Inc.

P.O. Box 67

67 Main Street

Dobbs Ferry, NY 10522

POC: Walter Pastor (914–693–2620)

• Schwarzkopf Technologies Corp.

115 Constitution Blvd.

Franklin, MA 02038

POC: Sales (800–782–6659)

• H.C. Starck

460 Jay St.

Coldwater, MI 49036

POC: Sales (201–438–9000)

Source—Competitive (not planning on using a sole source).
62

C.2 Material Specification

Mo-Re alloy tube and rod stock will be produced with a baseline Re composition of either 44.5% 

or 47.5% (weight percentage—the final composition to be specified by MSFC). The material for the rod 

and tube products can be formed by powder metallurgy or arc cast, provided that the final drawn/swaged 

and machined tube and rod stock is fully dense. Previous experience manufacturing similar material shapes 

from various Mo-Re alloys is a significant technical benefit for the vendor. The vendor will state relevant 

experience, capability, and manufacturing specifications/guidelines in their response.

Material processing specifications will include the following:

• Straightness and ovality—ASTM A269, A213, or equivalent.

• Mo alloy (rod, bar, wire)—ASTM B387.

• Estimate the average grain size of metals—ASTM E112.

• Density measurement/microstructure technique/specification—determine ductility of final product, 

procedure/specification to be provided by the vendor.

• Porosity measurement technique/specification—determine ductility of final product, procedure/ 

specification to be provided by the vendor.

• Bending test technique/specification—determine ductility of final product, procedure/specification to be 

provided by the vendor.

C.2.1 Dimensional 

Material stock will conform to the following requirements (vendor will state what they are capable 

of providing):

• Permissible variations in outside diameter ±0.005 in.

• Permissible variation in wall thickness for tubing is ±10%.

• Permissible ovality variations in tubing diameter at any cross section cannot exceed the variation in 

outside diameter given above.

• Permissible straightness is 0.010-in/ft length for 5/8-in-diameter tubes and 0.05 in/ft for 1/4-in-diameter 

tubes.

• Surface roughness on all tube and rod stock will be 64–125 μin.

• The use of centerless grinding results in a permissible diameter reduction of the outside diameter to 

±0.003 in; however, permissible variation in wall thickness must be maintained in the case of tubing.
63

C.2.2 Metallurgical

The Mo-Re alloy material for both tubing and rod stock will be processed to conform to the 

following requirements:

• All material stock will have no interconnecting porosity that would serve as a leak path to He.

• All material stock will be uniform and free of defects/imperfections such as draw marks, cracks, splits, 

voids, dimples, etc. on both the inner and outer surfaces impacting the use and performance of the parts.

• All material stock will be free of surface and embedded contamination such as residues and films that 

might be introduced during the fabrication processes, handling, furnace treatments, and chemical 

cleaning/rinsing.

• All material stock will be furnished in the annealed temper condition with a hardness range to be 

provided by the vendor to MSFC for approval.

• All material stock can show no more than 1% microstructure as deformation twins of platelet morphology. 

For the Mo-47.5%Re alloy, it can contain up to a 5% nodular sigma phase (hard and brittle intermetallic 

compound). For the special alloy Mo-44.5%Re, no sigma phase is present.

• The vendor will provide details regarding the smallest average grain size they are capable of providing 

in all products (a maximum average grain size of 0.0015 in or smaller is desirable). The vendor will 

identify a method to be used to quantify the grain size and discuss with MSFC. 

• Ductility tests will be performed to verify ductility of product (from material batch). For tubing, this 

could possibly consist of cutting out a piece of the wall and then bending to a specification. The vendor 

will determine method and then discuss with MSFC.

• For all Mo-Re alloys, the percentage of Re will be ±0.3% by weight or better.

C.2.3 Material Density Sampling 

Random sampling will be performed for each material batch from which the tubing and rod stock 

is produced. In addition, density sampling will be performed on all final drawn/swaged tube and rod stock 

to determine effect of processing—documenting the final product density with a target of fully dense. The 

vendor will provide technique/specification and discuss with/get approval from MSFC.

C.2.4 Chemical Composition 

Random material purity sampling will be performed for initial material batch from which the tube 

and rod stock is produced. In addition, purity tests on all final drawn/swaged and cleaned tube and rod 

stock (of drawn length) will be performed, documenting the final purity of the product. The test method 

will conform to LECO elemental analyzer and GDMS or other acceptable methods (will be discussed with 

and approved by MSFC) to examine both bulk and surface compositions. The maximum acceptable and 
64

target concentrations of each impurity for powder metallurgy material is listed in table 18. Error limits for 

purity measurements will be within 10% of measured value. The balance of material will be Mo and Re 

with the weight percentage of Re at ±0.3% of the prescribed composition (44.5% or 47.5%).

Table 18. Impurity chemical composition (acceptable) final heat 

Mo-Re alloy material (PM).

Element

Maximum

(wppm) 

Target 

(wppm)

O

C

N

H

Ca

Fe

Ni

W

Si

Mg

Mn

Co

Sn

B

20

20

30

25

7

75

50

300

30

10

4

30

20

20

<10

<10

<3

<1

<5

<20

<20

<100

<10

<10

<4

<30

<20

<10

For arc cast Mo-Re alloy material, the concentration of oxygen will be maintained at 20 ppm while that of 

carbon can range from 40 to 80 ppm.

C.2.5 Cleaning 

To ensure the purest material stock possible, cleaning will be performed as necessary during 

processing. The material can be chemically, vacuum/hydrogen fired (hydrogen dewpoint less than –77 °C), 

or other to accomplish this task. Specific selection and documented procedures will be provided by the 

vendor and discussed with MSFC for approval (a general sample chemical cleaning specification is 

provided in section C.4). 

C.2.6 Leak Checking 

Leak checking will be performed to verify leak tightness of all heat pipe components. Tubing and 

welded assemblies will be checked with an He mass spectrometer leak detector with a sensitivity capable of 

detecting leaks rates of 1×10–10 std-cc/s of He or lower. All of the tubing (100%) will be leak tested and any 

defective units identified. Corrective action to repair units with leakage will be discussed and approved 

by MSFC. 

C.2.7 Final Acceptance and Packaging 

All tube and rod stock will be verified to conform to required size, quantity, purity, density, leak 

tightness, and microstructure. Process controls will be in place that document these items and enable 

tracing of material lots. This documentation will be provided to MSFC for acceptance. All materials 
65

will be packaged in containers that are accepted by common carriers for safe transport and provide the 

following: (1) Sufficient protection to protect against damage during shipment, (2) cleanliness of the 

product and prevention of oxidation, and (3) safe transport at the lowest cost rate. All items will be shipped 

to either MSFC or a designated contractor facility. 

C.3 Material Size and Quantity Specifications

The quantity for all items has been increased above the required estimate by 5% to account for 

production waste and machining as follows with the total price estimated on best available data: 

• Item 1

– Quantity = 246 in—Final product in 13.25-in lengths or multiples.

– OD = 0.625-in tube.

– Wall = 0.05 in.

• Item 2

– Quantity = 38 in—Final product lengths in multiples of 2 in. 

– OD = 0.625-in tube.

– Wall = 0.03 in.

• Item 3

– Quantity = 86 in—Final product lengths in multiples of 1.5 in. 

– OD = 0.5-in tube.

– Wall = 0.03 in.

• Item 4

– Quantity = 114 in—Final product lengths in multiples of 6 in. 

– OD = 0.25-in tube.

– Wall = 0.03 in.

• Item 5

– Quantity = 48 in—Final product lengths in multiples of ≈12 in.

– OD = 0.625 in rod.

• Item 6

– Quantity = 20 in—Final product lengths in multiples of ≈12 in.

– OD = 0.5-in rod.

• Item 7

– Quantity = 20 in—Final product length in multiples of ≈12 in.

– OD = 0.05-in rod.

• Item 8

– Quantity = 20 in2—Minimum size to cut out 1-in2 items.

– THK = 0.03-in sheet.
66

C.4 General Cleaning Procedure for Mo and Mo-Re Parts

(1) Wash the piece in PF solvent, an R113 substitute, until all signs of grease have been removed.

(2) Soak piece for 1–2 min in one part by volume HCl and one part by volume deionized water 

(target of 10 MΩ-cm or better) to remove residual iron surface impurities.

(3) Soak piece for 5 min in caustic cleaning solution consisting of 11 parts by volume deionized 

water (target of 10 MΩ-cm or better), 1 part by volume NaOH, and 1 part by volume H2O2. Remove piece 

from caustic bath. Replenish or replace solution as required.

(4) Wash part in hot, deionized water (target of 10 MΩ-cm or better) for at least 5 min.

(5) Repeat steps (3) and (4) three times.

(6) Rinse piece for 5 min in ethanol inside an ultrasonic cleaner.
67

APPENDIX D—HEAT PIPE FABRICATION/ASSEMBLY PROCUREMENT SPECIFICATION

D.1 IFMP Procurement Specification for Fabrication of Mo-Alloy Heat Pipe Units

Description—Procure assembled refractory metal Mo alloy-based heat pipe units.

Material—Refractory material and wick assemblies supplied by MSFC (unless otherwise specified in the 

specifications). 

Quantity—18 units.

Delivery—ARO 16–20 wk. 

Potential Suppliers:

• Advanced Methods and Materials

1190 Mountain View-Alviso Rd.

Suite P

Sunnyvale, CA 94089

POC: Peter Ring (408–745–7772)

• Refrac Systems

7201 W. Oakland St.

Chandler, AZ 85226–2434

POC: Norm Hubele (480–940–0068)

• Lockheed-Martin

Denver, CO

POC: Suraj Rawal (303–971–9378)

Source—Competitive (not planning on using a sole source).

D.2 Heat Pipe Layout Specification

Heat pipe units will be fabricated from Mo-Re alloy with a baseline Re composition of either 

44.5% or 47.5%. The engineering drawings, provided in figures 36 and 37, detail the hardware layout. The 

Mo-Re heat pipe material and completed wick assemblies will be supplied by MSFC or its representative. 

The heat pipes will also be equipped with a getter pack located in the evaporator plug. The contractor will 

purchase the Zr materials required for getter construction as part of this contract. The final output is a fully 

assembled and cleaned heat pipe unit that is ready for loading with an alkali metal working fluid. 
68

It is a significant technical benefit if the vendor has previous experience manufacturing similar 

material shapes from various Mo-Re alloys. The vendor will state relevant experience, capability, and 

manufacturing specifications/guidelines in their response.

D.2.1 Material and assembly specifications include the following:

• Straightness and ovality—ASTM A269 (or equivalent).

• Mo alloy (rod, bar, wire)—ASTM B387 (or equivalent).

• Other applicable machining techniques/specifications:

– Determine ductility of final product, procedure/specification to be provided by the vendor.

• Sample EB welds (or other applicable process).

– Performed to verify room temperature ductility of final closeout welds; vendor will supply methods to 

test and verify. 

Figure 36. Heat pipe unit layout.
69

Figure 37. Heat pipe getter pack.

D.2.2 Dimensional

The material stock should conform to the following target requirements (as provided by MSFC 

material vendor):

• Permissible variations in outside diameter ±0.005 in.

• Permissible variation in wall thickness for tubing is ±10%.

• Permissible ovality variations in tubing diameter at any cross section cannot exceed the variation in 

outside diameter given above.

• Permissible straightness is 0.01 in/ft length for 5/8-in-diameter tubes and 0.05 in/ft for 1/4-in-diameter 

tubes. The use of centerless grinding provides for a permissible variation in the outside diameter to 

±0.003 in; however, the permissible variation in wall thickness must be maintained in the case of tubing. 

Deviations will be documented and discussed with MSFC for determination of course of action by the 

heat pipe fabrication vendor (machining, straightening, material rejection, etc.).
70

D.2.3 Surface Roughness

All tube and rod stock provided by MSFC will have a 64–125 μin surface finish in the as-received 

condition. The contractor during fabrication will machine the outer surface of all heat pipe components 

(tubeand rod) to a surface finish of 64 μin or better. The inner surface of all heat pipe components (tube and 

rod) will be machined to a surface finish of 32 μin or better. The exception is the inner surface of the heat pipe 

outer tube (dwg No. 100–HEAT PIPE, item 104) that will be honed to a surface finish of 1–16 μin.

D.2.4 Wick Assembly Verification

The wick assemblies will be received in ready-to-install condition from either MSFC or a specified 

contractor. Each assembly will include a residual wick piece 2 in or longer to be used in purity tests. These 

assemblies will be provided by MSFC in a clean condition and contained in inert gas carriers with the Mo5%Re wick material in an uncrystallized condition. Upon receipt, the contractor will visually inspect the 

wick assemblies to identify potential shipping damage. During inspection, the wicks will be handled in a 

clean, inert environment to minimize potential contamination. After inspection, the wick will be stored in 

the inert gas carriers in a safe location. Two specific tests on the as-received, completed wick assemblies 

will be conducted by the contractor as follows:

(1) Bubble-point testing will be performed using pure ethyl alcohol with a maximum acceptable 

pore radius of 35 μm. Any wick assembly that does not meet the pore size requirement will be identified 

and set aside. This process will be documented and results provided to MSFC.

(2) Material chemical analysis will be performed on a section of residual wick material and provided 

with each lot of wick assemblies. Residual material that serves as a witness sample should be marked using 

a LECO elemental analyzer, GDMS, or other suitable method to identify which assemblies it belongs 

with, the composition documented, and the results provided to MSFC. The maximum acceptable and 

target concentrations of each impurity for the MSFC-supplied wick assemblies are listed in table 19. The 

as-received wick assembly chemical analysis will be available to the vendor for comparison.

Table 19. Impurity chemical composition (acceptable) 

wick assembly.

Element

Maximum

Acceptable

(wppm)

Target

(wppm)

O

C

N

H

Ca

Fe

Ni

W

Si

Mg

Mn

Co

Sn

B

220

85

100

50

7

130

160

300

30

10

4

30

20

20

<25

<20

<5

<3

<7

<50

<50

<100

<10

<10

<4

<30

<20

<10
71

D.2.5 Chemical Composition 

Material purity sampling will be performed (all but getter pack/wick) during the assembly 

development phase for the first several heat pipe units processed. Material purity tests (measured in 

weight parts per million) will be performed on components at major assembly steps to determine the 

introduction of impurities and the quality of the final product. The test method will be LECO elemental 

analyzer, GDMS, or other acceptable method to examine bulk and surface impurities (will be discussed 

with and approved by MSFC). The maximum acceptable and target concentrations of each impurity for the 

MSFC-supplied materials is listed in table 20. The as-received material chemical analysis will be available 

to the heat pipe assembly contractor and serve as a baseline to monitor the fabrication process (objective 

is to have no to minimal pickup of impurities during processing). Sampling intervals during development 

of the process (first several heat pipe units) includes the following potential steps:

• Random sampling of supplied stock (based on lots) recommended.

• Required—Surface sampling of assembled heat pipe components prior to final assembly (just prior 

to welding).

• It is recommended that a sample material item subjected to similar machining, cleaning, and other 

processing serve as a witness piece and be carried through the initial production sequence and monitored 

to get an idea of any contamination pickup.

• Sampling interval for heat pipe production:

– Surface sampling of complete assembled heat pipes (random locations) after final cleaning and just 

prior to packaging for shipment.

The fabrication process will be devised so as to introduce no additional impurities above the 

as-received condition. Error limits for purity measurements will be within 10% of the measured value.

Table 20. Impurity chemical composition (acceptable)

final heat Mo-Re alloy material (PM).

Element

Maximum 

(wppm)

Target

(wppm) 

O

C

N

H

Ca

Fe

Ni

W

Si

Mg

Mn

Co

Sn

B

20

20

30

25

7

75

50

300

30

10

4

30

20

20

<10

<10

<3

<1

<5

<20

<20

<100

<10

<10

<4

<30

<20

<10
72

D.2.6 Cleaning 

Each of the heat pipe components and completed assemblies will be chemically cleaned and/

or vacuum-hydrogen fired as specified in procedures (to be discussed with and approved by MSFC). A 

sample general Mo-Re chemical cleaning specification is provided in section D.3. Each individual part 

will be cleaned by vacuum-hydrogen firing (does not include the wick/getter pack) once final machining is 

completed and prior to EB welding. The final heat pipe configuration with installed wick assembly will be 

cleaned by hydrogen firing (to be discussed with and approved by MSFC). During any heat treatments that 

involve the wick assembly, the unit will be kept within the cumulative time-temperature envelope, shown in 

section D.4, to prevent recrystallization of the Mo-5%Re material. (This includes multiple heat-treat steps 

at different temperatures and past history.) The total number of heat treatments and temperature/time of 

exposures must be discussed with and approved by MSFC. After final cleaning of the complete heat pipe 

units, they will be charged (internally) with 1-atm, dry, oil-free Ar (99.999% pure) and the fill stem sealed 

using a Swagelok compression-type fitting. Specific chemical cleaning procedures for the Zr getter pack 

with to-be-approved process that will be discussed with and approved by MSFC. Hydrogen gas used in 

firing to have a dewpoint less than –77 °C.

D.2.7 Leak Checking 

Leak checking will be performed to verify leak tightness of all heat pipe components. Tubing and 

welded assemblies will be checked with an He mass spectrometer leak detector with a sensitivity capable 

of detecting leaks rates of 1×10–10 std-cc/s of He or lower. All components (100% of the heat pipe units) 

will be leak tested and any defective units identified. Corrective action to repair units identified to leak 

will be discussed with and approved by MSFC. A minimum of three leak-check intervals (to be used on 

all heat pipes) are identified:

(1) After EB welding the heat pipe outer tube (dwg. No. 100–HEAT PIPE, item 104) to the 

evaporator plug (dwg. No. 100–HEAT PIPE, item 101), the wick assembly has not been installed.

(2) After EB welding the fill stem tube (dwg. No. 100–HEAT PIPE, item 103) to the condenser 

plug (dwg. No. 100–HEAT PIPE, item 102).

(3) After closeout the complete heat pipe assembly—The final EB weld of the heat pipe outer tube 

(dwg. No. 100–HEAT PIPE, item 104) to the condenser plug (dwg. No. 100–HEAT PIPE, item 102).

D.2.8 Final Acceptance and Packaging 

Documentation will be generated for all specific process steps/operations including how they did it 

(no proprietary processing), results of purity, dimensional tolerances checks, leak testing, cleaning, and final 

assembled weight; documentation for each completed heat pipe will be provided to MSFC for final review/

acceptance. The complete heat pipe units, after acceptance, will be packaged inside inert gas protective 

transport carriers (type of transport will be discussed with MSFC); the Mo-5%Re wick assemblies will 

remain in the uncrystallized condition (cumulative time/temperature heat treatments that have involved the 

wick structure have not exceeded limits to be discussed and approved by MSFC). Additionally, section D.5 

lists a sample process flow for the fabrication and assembly process.
73

D.3 General Cleaning Procedure for Mo and Mo-Re Parts

(1) Wash piece in PF solvent, an R113 substitute, until all signs of grease have been removed.

(2) Soak piece for 1–2 min in one part by volume HCl and one part by volume deionized water 

(target of 10 MΩ-cm or better) to remove residual iron surface impurities.

(3) Soak piece for 5 min in caustic cleaning solution consisting of 11 parts by volume deionized 

water (target of 10 MΩ-cm or better), 1 part by volume NaOH, and 1 part by volume H2O2. Remove piece 

from caustic bath. Replenish or replace solution as required.

(4) Wash part in hot, deionized water (target of 10 MΩ-cm or better) for at least 5 min.

(5) Repeat steps (3) and (4) three times.

(6) Rinse piece for 5 min in ethanol inside an ultrasonic cleaner.

(7) Vacuum fire or hydrogen fire part as appropriate using not-to-exceed time-temperature relation 

given in section D.4). Hydrogen dewpoint will be less than –77 °C.

D.4 Time-Temperature Relation for Mo-5%Re Material

To prevent recrystallization of the Mo-5%Re wick materials, the cumulative time at temperature 

must be assessed for all expected heat treatments, as shown in table 21. For multiple heat-treat steps at 

different temperatures, the following relation will apply: 

τ i / τ m Ti ( ) ⎡

⎣ ⎤ ∑ ⎦ < 1 , (46)

where τi is the time at temperature, Ti for a given step, and τm is the maximum time at temperature for the 

given step.
74

Table 21. Mo-5%Re cumulative time at temperature.

Temperature 

(K)

Maximum Time 

at Temperature 

(min)

1,000

1,010

1,020

1,030

1,040

1,050

1,060

1,070

1,080

1,090

1,100

1,110

1,120

1,130

1,140

1,150

1,160

1,170

1,180

1,190

1,200

1,210

1,220

1,230

1,240

1,250

1,260

1,270

1,280

1,290

1,300

2,049.0

1,448.9

1,031.6

739.3

533.2

387.0

282.6

207.6

153.3

113.9

85.1

63.9

48.2

36.5

27.9

21.3

16.4

12.7

9.8

7.7

6.0

4.7

3.7

2.9

2.3

1.9

1.5

1.2

1.0

0.8

0.6

D.5 Sample Process Flow for Heat Pipe Fabrication/Assembly

D.5.1 Simplified Process Flow

(1) Inspect and lay out supplied materials (received condition).

(2) Machine parts.

(3) Chemically clean and or vacuum-hydrogen fire materials.

• Required—Random material chemical/purity sampling of supplied material or sample 

witness pieces carried through the operations (during development process of first several 

assemblies).

(4) Fabricate getter pack and assemble in evaporator plug.
75

(5) Weld heat pipe assembly, except for condenser/fill stem plug to heat pipe body.

(6) Insert wick structure.

(7) Final weld of condenser/fill stem plug to heat pipe body.

(8) Clean unit with hydrogen firing:

• Required—Material chemical/purity sampling of surface or sample material pieces carried

through the assembly operation.

(9) Pack unit for shipment to MSFC.

D.5.2 Cost Estimation

• Item 1: Evaporator plug/condenser plug/fill stem—Machine, clean, and EB-weld fill stem to condenser plug

– Quantity = 18 of each.

• Item 2: Getter assemblies—With spot welds into the evaporator plugs

– Quantity = 18 of each.

• Item 3: Heat pipe tube–Honing, machining, clean, EB weld to evaporator plug and leak check

– Quantity = 18 of each.

• Item 4: Heat pipe tube final assembly—Wick, cleaning, leak check, and final cleaning/sampling 

– Quantity = 18 of each.

• Item 5: Fill stem closeout tube—Machine, assemble, leak check, and clean 

– Quantity = 18 of each.

• Item 6: Final checkout/packaging and shipping

– Quantity = 18 of each.
76

APPENDIX E—MATERIAL PROPERTIES USED IN ANALYSIS

E.1 Helium/Argon Mixture—Thermal Conductivity Property

The ideal gas approximation is used for the equation of state and specific heat of the He-Ar gas 

mixture in the gap between the heat pipe and the calorimeter. Thermal conductivities for He-Ar mixtures 

in the gap are found in data from Touloukian as follows:25

kHe = 0.0025672T 0.716 , (47)

kAr = 1.39 ×10−4T 0.852 −1.5 ×10−8 (T − 300)(T − 300) ⎡

⎣ ⎤⎦ 1+ 2 ×10−8 ⎡ P ⎣ ⎤

⎦ , (48)

and

kHe−Ar = f

ArkAr + fHekHe − kAr + kHe ( ) fAr fHe

1.5 . (49)

Properties for pure He are given in table 22 and figures 38–40, 316 stainless steel in table 23 and 

figures 41–45, Mo-44%Re in table 24 and figures 46–48, and water in table 25 and figures 49–53.26–28

Table 22. Pure He properties at 76 torr.

Temperature

(K)

Density

(kg/m3)

Kinematic

Viscosity (ν)

(m2/s)

Thermal

Conductivity (k)

(W/m-K)

Specific

Heat (Cp)

(J/kg-K)

300

400

500

700

1,000

1,300

1,500

0.016260

0.012190

0.009756

0.006968

0.004878

0.003752

0.003252

1.224×10–3

1.993×10–3

2.911×10–3

5.152×10–3

9.471×10–3

1.482×10–2

1.891×10–2

0.155

0.189

0.221

0.280

0.360

0.434

0.480

5,193

5,193

5,193

5,193

5,193

5,193

5,193
77

Density (kg/m3)

Temperature (K)

 =4.8751×T –0.999

0.018

0.016

0.014

0.012

0.01

0.008

0.006

0.004

0.002

0

0 500 1,000 1,500 2,000 2,500

ρ

Figure 38. Helium density at 76 torr (1.5 psia).

Kinematic Viscosity (m2/s)

Temperature (K)

 =6.004×10–9×T 2 + 3.991×10–6×T – 5.531×10–4

3.5×10–2

3×10–2

2.5×10–2

2×10–2

1.5×10–2

1×10–2

5×10–3

0

0 500 1,000 1,500 2,000 2,500

ν

Figure 39. Helium kinematic viscosity at 76 torr (1.5 psia).
78

Thermal Conductivity (W/m-K)

Temperature (K)

k=–4.72×10–8×T 2 + 3.54×10–4×T – 5.45×10–2

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0

0 500 1,000 1,500 2,000 2,500

Figure 40. Helium thermal conductivity at 76 torr (1.5 psia).

Table 23. 316 stainless steel properties.

Temperature

(K)

Density

(kg/m3) Emissivity*

Thermal

Conductivity

(W/m-K)

Specific

Heat

(J/kg-K)

Coefficient

Thermal

Expansion

 (1/K)

273

300

400

500

600

700

800

900

1,000

1,100

1,200

1,300

7972.70

7960.74

7916.46

7872.18

7827.89

7783.61

7739.32

7695.04

7650.75

7606.47

7562.19

7517.90

0.33

0.34

0.37

0.41

0.44

0.47

0.50

0.54

0.57

0.60

0.63

0.67

12.99

13.46

15.14

16.77

18.35

19.88

21.34

22.76

24.11

25.40

26.64

27.81

456.39

467.69

502.87

529.17

548.71

563.59

575.91

587.78

601.31

618.61

641.77

672.91

1.49×10–5

1.52×10–5

1.60×10–5

1.67×10–5

1.74×10–5

1.79×10–5

1.83×10–5

1.87×10–5

1.90×10–5

1.93×10–5

1.95×10–5

1.97×10–5

* Number 8 finish and degraded 33%.
79

Conductivity (W/m-K)

Temperature (K)

k=–3×10–6×T 2 + 0.0189×T + 8.0402

30

28

26

24

22

20

18

16

14

12

10

0 200 400 600 800 1,000 1,200 1,400

Figure 41. 316 stainless steel thermal conductivity.

Specific Heat (J/kg-K)

Temperature (K)

Cp=3.51×10–7×T 3 – 8.65×10–4×T 2 + 8.27×10–1×T + 2.88×102

700

650

600

550

500

450

400

0 200 400 600 800 1,000 1,200 1,400

Figure 42. 316 stainless steel specific heat.
80

Density (kg/m3)

Temperature (K)

 =–0.4428×T+8093.6

8,000

7,900

7,800

7,700

7,600

7,500

7,400

0 200 400 600 800 1,000 1,200 1,400

ρ

Figure 43. 316 stainless steel density.

Coefficient of Thermal Expansion (1/K)

Temperature (K)

CTE = –3.783×10–12×T 2 + 1.048×10–8×T + 1.238×10–5

2×10–5

1.9×10–5

1.8×10–5

1.7×10–5

1.6×10–5

1.5×10–5

1.4×10–5

0 200 400 600 800 1,000 1,200 1,400

Figure 44. 316 stainless steel CTE.
81

Emissivity

Temperature (K)

ε=4.11×10–7×T 2 – 2.88×10–4×T + 3.42×10–1

0.7

0.65

0.6

0.55

0.5

0.45

0.4

0.35

0.3

0.25

0.2

0 200 400 600 800 1,000 1,200 1,400

Figure 45. 316 stainless steel emissivity.

Table 24. Mo-44%Re properties.

Density (273 K) = 13.5 g/cm3

Specific Heat (283 K) = 200 J/kg-K

Temperature 

(K) Emissivity*

Thermal

Conductivity

 (W/m-K)

Coefficient

Thermal

Expansion

(1/K)

273

300

400

500

600

700

800

900

1,000

1,100

1,200

1,300

1,400

1,500

1,600

1,700

1,800

1,900

2,000

2,100

2,200

0.051

0.055

0.067

0.079

0.091

0.102

0.113

0.124

0.134

0.144

0.153

0.162

0.171

0.180

0.188

0.196

0.203

0.210

0.217

0.223

0.229

41.04

41.96

45.21

48.20

50.94

53.42

55.65

57.62

59.33

60.79

61.99

62.94

63.63

64.07

64.25

64.18

63.85

63.27

62.43

61.34

59.99

5.84×10–6

5.85×10–6

5.9×10–6

6×10–6

6.13×10–6

6.3×10–6

6.51×10–6

6.76×10–6

7.04×10–6

7.36×10–6

7.73×10–6

8.12×10–6

8.56×10–6

9.03×10–6

9.55×10–6

1.01×10–5

1.07×10–5

1.13×10–5

1.2×10–5

1.27×10–5

1.34×10–5

* Grit-blasted surface.
82

Thermal Conductivity (W/m-K)

Temperature (K)

k=–1.277×10–5×T 2 + 4.141×10–2×T + 3.069×101

70

65

60

55

50

45

40

35

30

0 500 1,000 1,500 2,000 2,500

Figure 46. Mo-44%Re thermal conductivity.

Emissivity

Temperature (K)

ε=–1.764×10–8×T 2 + 1.358×10–4×T + 1.565×10–2

0.25

0.2

0.15

0.1

0.05

0

0 500 1,000 1,500 2,000 2,500

Figure 47. Mo-44%Re emissivity.
83

Coefficient of Thermal Expansion (1/K)

Temperature (K)

CTE = 1.894×10–12×T 2 – 7.528×10×T + 5.901×10–6

1.4×10–5

1.3×10–5

1.2×10–5

1.1×10–5

1×10–5

9×10–6

8×10–6

7×10–6

6×10–6

5×10–6

4×10–6

0 500 1,000 1,500 2,000 2,500

Figure 48. Mo-44%Re CTE.

Table 25. Water properties at 0.448 MPa (65 psia).28

Temperature

(K)

Density

(kg/m3)

Absolute

Viscosity 

(kg/m-s)

Thermal

Conductivity

(W/m-K)

Specific

Heat

(J/kg-K)

280

290

300

310

320

330

340

350

360

370

380

390

400

410

420

1,000.10

998.96

996.71

993.54

989.58

984.94

979.69

973.88

967.56

960.75

953.48

945.76

937.59

928.98

919.93

1.433×10–3

1.084×10–3

8.538×10–4

6.936×10–4

5.771×10–4

4.896×10–4

4.221×10–4

3.689×10–4

3.262×10–4

2.915×10–4

2.628×10–4

2.388×10–4

2.187×10–4

2.015×10–4

1.867×10–4

0.574

0.593

0.610

0.626

0.640

0.651

0.661

0.668

0.674

0.678

0.681

0.683

0.684

0.684

0.683

4,199.5

4,185.5

4,179.7

4,178.4

4,179.7

4,182.9

4,187.5

4,193.7

4,201.6

4,211.4

4,223.3

4,237.8

4,255.0

4,275.2

4,298.9
84

Density (kg/m3)

Temperature (K)

 = –0.0028×T2 + 1.3882T + 833.38

1×103

9.9×102

9.8×102

9.7×102

9.6×102

9.5×102

9.4×102

9.3×102

9.2×102

9.1×102

275 300 325 350 375 400 425 450

ρ

Figure 49. Water density at 0.448 MPa (65 psia).

Thermal Conductivity (W/m-K)

Temperature (K)

k = –8.167×10–6×T2 + 6.462×10–3×T – 5.933×10–1

7×10–1

6.8×10–1

6.6×10–1

6.4×10–1

6.2×10–1

6×10–1

5.8×10–1

5.6×10–1

275 300 325 350 375 400 425 450

Figure 50. Water thermal conductivity at 0.448 MPa (65 psia).
85

Viscosity (kg/m-s)

Temperature (K)

 = 9.564×10–12×T4 –1.433×10–8×T3 + 8.058×10–6×T2 – 2.019×10–3×T +1.908×10–1 

1.6×10–3

1.4×10–3

1.2×10–3

1×10–3

8×10–4

6×10–4

4×10–4

2×10–4

0

275 300 325 350 375 400 425 450

μ

Figure 51. Water absolute viscosity at 0.448 MPa (65 psia).

Specific Heat (J/kg-K)

Temperature (K)

Cp = 0.0111×T 2 – 7.0364×T + 5295.2

4,320

4,300

4,280

4,260

4,240

4,220

4,200

4,180

4,160

275 300 325 350 375 400 425 450

Figure 52. Water specific heat at 0.448 MPa (65 psia).
86

Pressure (Pa)

Temperature (K)

P=9.946×101×T 2 –7.167×104×T + 1.3×107

7×105

6×105

5×105

4×105

3×105

2×105

1×105

0

370 380 390 400 410 420 430 440

Figure 53. Water boiling point.
87

APPENDIX F—POWER EXTRACTION PERFORMANCE WORKSHEET

Tables 26–28 provide a quick geometry and performance assessment for water calorimeter heat 

pipe life time testing with a cooling film annulus of 0.010, 0.015, and 0.020 in. 
88

Table 26. Calorimeter geometry and performance for a cooling film annulus = 0.010 in. 
89

Table 27. Calorimeter geometry and performance for a cooling film annulus = 0.015 in. 
90

Table 28. Calorimeter geometry and performance for a cooling film annulus = 0.020 in. 
91

APPENDIX G—HEAT PIPE/CALORIMETER ENVIRONMENTAL LOSSES

Tables 29 and 30 provide a quick assessment of environmental losses from the heat pipe and 

calorimeter unit using environments of pure He and He-32%Ar.

Table 29. Heat pipe/calorimeter thermal loss spreadsheet for pure He environment.
92

Table 30. Heat pipe/calorimeter thermal loss spreadsheet for He-32%Ar environment.
93

APPENDIX H—THE GENERALIZED FLUID SYSTEM SIMULATION PROGRAM

GFSSP has gone through extensive V&V since the first version was released in 1996. The following 

three methods have been used for V&V: (1) Comparison with classical analytical and numerical solution, 

(2) comparison with other code, and (3) comparison with test data.29–34 GFSSP (version 3) was a joint 

winner of the 2001 NASA Software of the Year Award and has been patented. Concepts NREC Inc. has 

been licensed for GFSSP commercialization. Further information is available online, <http://mi.msfc.

nasa.gov/GFSSP/index.shtml>.

GFSSP was developed and used during the design of the Simplex turbopump at MSFC. Predictions 

matched well to test data. However, the code was more extensively used during the design of the Fastrac 

turbopump at MSFC in which both steady state and transient calculations were performed. The Fastrac 

engine is a 60,000-lbf thrust liquid oxygen and rocket propellant 1 engine developed by MSFC. When the 

Fastrac turbopump was being designed, GFSSP was used to predict the steady-state internal flows and net 

axial thrust. This turbopump has since been tested at MSFC. Data from this testing program have been 

used to validate the transient capabilities of GFSSP.

Modeling a component test of the Fastrac turbopump provided the opportunity to compare the 

GFSSP-predicted flow field with measured pressures and temperatures. The boundary conditions for this 

transient model were derived from measured pressure and temperature data taken from a Fastrac turbopump 

component test conducted at MSFC. Results from the steady-state GFSSP model are shown in figure 54. 

Additionally, a representative plot comparing results from the GFSSP transient model to test data is shown in 

figure 55. In both the steady state and transient cases, good agreement (within 10%) was observed between 

the model predictions and test data at all model nodes shown. 
94

58

N/A

48

N/A

265

262

–1%

259

259

–0%

856

816

–5%

161

57

292

57

917

57

14.7

70

14.7

70

848

–287

247

–289 68

–290

##

##

O2/He 

Exit

RP/He 

Exit

175

67

729

690

–5%

##

##

#%

449

487

+8%

0.008

0.008

–0%

0.008

0.008

–0%

70

69

–1%

675

678

+0%

RP

RP and He

He

O2 and He

O2

Boundary Pressure (psia)

Boundary Temperature (°F)

Measured Pressure (psia)

Calculated Pressure (psia)

% Difference

Measured Flowrate (lbm/s)

Calculated Flowrate (lbm/s)

% Difference

+ Thrust 60k100.out 3/99

Net Predicted Axial Thrust = – 810 lbf

*Measured and node location actually downstream of turbopump

Figure 54. Steady-state results from the Fastrac turbopump GFSSP model.

900

800

700

600

500

400

300

200

100

Pressure (psia)

Time (s)

211

212

111

Actual

Predicted

42 43 44 45 46 47 48 49 50

Figure 55. Representative transient pressure results from 

the Fastrac turbopump GFSSP model.
95

REFERENCES

1. Feber, R.C.; and Merrigan, M.A.: “Thermochemical Modeling of Mass Transport in High-Temperature 

Heat Pipes,” Proc. VI Int. Heat Pipe Conf., Grenoble, France, 1987.

2. Sena, J.T.; and Merrigan, M.A.: “Niobium 1% Zirconium Potassium and Titanium Potassium 

Life-Test Heat Pipe Design and Testing,” M.S. El-Genk and M.D. Hoover (eds.), Proc. Seventh

Symposium on Space Nuclear Power Systems, CONF–900109, Albuquerque, NM, January 7–11, 1990.

3. Deming, S.N.; and Morgan, S.L.: Experimental Design: A Chemometric Approach, Elsevier, 

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Design of Refractory Metal Life Test Heat Pipe and Calorimeter

J.J. Martin, R.S. Reid, and S.M. Bragg-Sitton

George C. Marshall Space Flight Center

Marshall Space Flight Center, AL 35812

National Aeronautics and Space Administration

Washington, DC 20546–0001

Unclassified-Unlimited

Subject Category 20

Availability: NASA CASI (443–757–5802)

Prepared by the Nuclear and Advanced Propulsion Development Branch, Engineering Directorate 

M–1284

Technical Publication

NASA/TP—2010–216435

alkali metal, sodium, heat pipes, refractory, metal, molybdenum, rhenium, radio frequency heating, fission 

power, space nuclear, life testing

01–08–2010

UU 116

STI Help Desk at email: help@sti.nasa.gov

NASA

U

Heat pipe life tests have seldom been conducted on a systematic basis. Typically, one or more heat pipes are built and tested for an extended period at a 

single temperature with simple condenser loading. Results are often reported describing the wall material, working fluid, test temperature, test duration, 

and occasionally the nature of any failure. Important information such as design details, processing procedures, material assay, power throughput, and 

radial power density are usually not mentioned. We propose to develop methods to generate carefully controlled data that conclusively establish heat 

pipe operating life with material-fluid combinations capable of extended operation. The test approach detailed in this Technical Publication will use 

16 Mo-44.5%Re alloy/sodium heat pipe units that have an approximate12-in length and 5/8-in diameter. Two specific test series have been identified: 

(1) Long-term corrosion rates based on ASTM–G–68–80 (G-series) and (2) corrosion trends in a cross-correlation sequence at various temperatures 

and mass fluences based on a Fisher multifactor design (F-series). Evaluation of the heat pipe hardware will be performed in test chambers purged 

with an inert purified gas (helium or helium/argon mixture) at low pressure (10–100 torr) to provide thermal coupling between the heat pipe condenser 

and calorimeter. The final pressure will be selected to minimize the potential for voltage breakdown between the heat pipe and radio frequency (RF) 

induction coil (RF heating is currently the planned method of powering the heat pipes). The proposed calorimeter is constructed from a copper alloy 

and relies on a laminar flow water-coolant channel design to absorb and transport energy.

U U STI Help Desk at: 443–757–5802

National Aeronautics and

Space Administration

IS20

George C. Marshall Space Flight Center

Marshall Space Flight Center, Alabama

35812