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i

July 2008

NASA/TM—2008–215466

Heater Development, Fabrication, and Testing: 

Analysis of Fabricated Heaters

S.M. Bragg-Sitton, R.E. Dickens, J.T. Farmer, J.D. Davis, M.R. Adams, 

J.J. Martin, and K.L. Webster

Marshall Space Flight Center, Marshall Space Flight Center, Alabama

National Aeronautics and

Space Administration

Marshall Space Flight Center • MSFC, Alabama 35812
ii

Available from:

NASA Center for AeroSpace Information

7115 Standard Drive

Hanover, MD 21076 –1320

301–621–0390

This report is also available in electronic form at

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

TRADEMARKS

Trade names and trademarks are used in this report for identification only. This usage does not 

constitute an official endorsement, either expressed or implied, by the National Aeronautics and 

Space Administration.
iii

TABLE OF CONTENTS

1. INTRODUCTION—Historical Thermal Simulator Development ...... 1

1.1 General Requirements ................................................................................................ 1

1.2 Early Refractory Metal Heater Designs ...................................................................... 11

2. CURRENT THERMAL SIMULATOR TESTING AND DEVELOPMENT ................ 13

2.1 Refractory Metal Wire Wrapped Simulator Design ................................................... 13

2.2 Carbon Fiber Braid Simulator Design ........................................................................ 35

2.3 Sealed Thermal Simulator Configurations ................................................................. 40

2.4 Tantalum Wire Wrapped Heater Element .................................................................. 55

2.5 Tungsten Wire Wrapped Heater Element ................................................................... 59

2.6 Material Selection Criteria ......................................................................................... 70

2.7 Initial Conclusions from Thermal Simulator Tests ..................................................... 71

3. DEVELOPMENT OF SUPPORT HARDWARE ............................................................ 72

3.1 Gas Purification System .............................................................................................. 77

3.2 Water-Cooled Calorimeter ......................................................................................... 81

3.3 Data Acquisition System ............................................................................................ 87

3.4 Modified Power Interface ........................................................................................... 91

3.5 Component Cleaning and Chamber Preparation Procedures ..................................... 92

4. ADVANCED THERMAL SIMULATORS: CURRENT STATUS ................................. 96

5. FUTURE TESTING AND DEVELOPMENT PLANS .................................................. 100

5.1 Element Cooling Techniques ...................................................................................... 100

5.2 Advanced Thermal Simulator Development ............................................................... 101

6. CONCLUSIONS ............................................................................................................... 105

APPENDIX A—GRAPHITE ROD HEATER ELEMENT ................................................. 106

APPENDIX B—SPIRAL-WRAPPED SIMULATOR DESIGN, 0.625-IN ASSEMBLY ..... 109

APPENDIX C—SPIRAL-WRAPPED SIMULATOR DESIGN, 0.400-IN ASSEMBLY ..... 116
iv

TABLE OF CONTENTS (Continued)

APPENDIX D—TEMPERATURE-DEPENDENT RESISTIVITY OF SELECTED 

REFRACTORY METALS ........................................................................ 121

APPENDIX E—TANTALUM HEATER ELEMENT MATERIALS ANALYSIS ............. 122

APPENDIX F—RHENIUM HEATER ELEMENT MATERIALS ANALYSIS ................ 136

APPENDIX G—KEY GOVERNING EQUATIONS AND SUPPORTING 

FORMULATIONS .................................................................................... 149

APPENDIX H—THERMAL CROSS-SECTION CALCULATIONS .................................. 152

APPENDIX I—SINDA RESULTS: TEMPERATURE PROFILES .................................... 168

APPENDIX J—THERMAL ANALYSIS: CALORIMETER ............................................... 172

APPENDIX K—THERMOCOUPLE THERMAL ERROR ANALYSIS ............................ 181

APPENDIX L—DESIGN DRAWINGS: MODIFIED POWER INTERFACE ................... 187

APPENDIX M—BASIC MATERIAL CLEANING PROCEDURES ................................. 191

APPENDIX N—PURIFIED INERT GAS CONDITIONING ............................................ 195

APPENDIX O—ADVANCED HEATER ELEMENT DESIGNS: 

CHANGING POWER PROFILE............................................................... 198

APPENDIX P—ADVANCED HEATER ELEMENT DESIGNS: HEXAGONAL

CROSS SECTION ...................................................................................... 202

REFERENCES .................................................................................................................... 204
v

LIST OF FIGURES

 1. Graphite rod heater element: (a) sketch and (b) photograph ...................................... 3

 2. Graphite rod heater elements: (a) Batch of graphite rod heater elements and 

(b) identification of center taper geometry and power input desgin via Mo leads ...... 5

 3. Integration of 48 simulators in SAFE 30 (≈23- by 20-cm (≈9- by 8-in) footprint) ....... 7

 4. Integration of 183 simulators in SAFE 100 (≈26- by 29-cm (≈10.4- by 11.5-in) 

footprint) ................................................................................................................... 7

 5. Integration of 57 simulators in SAFE 100a (≈18- by 16.5-cm (≈7- by 6.5-in) 

footprint) ................................................................................................................... 7

 6. Integration of 37 simulators in DDG test article (≈16- by 18-cm (≈6.25- by 7.1-in) 

footprint). (Note: This core is inserted into a pressure vessel where the wires are 

not accessible after test article integration) ................................................................ 8

 7. Core face seal unit as configured for testing: (a) assembly on the test stand and 

placement of feedthroughs, (b) conflat style front cover plate, and (c) complete 

test configuration (without insulation shown) ............................................................ 9

 8. Integration of power with the LM core design using a core face seal concept: 

(a) core face seal extension showing power feedthroughs and (b) identification 

of ‘fuel tubes’ that house the resistive heater elements ............................................... 10

 9. General configuration of the hairpin heater element design ...................................... 11

10. Generalized assembly configuration (not shown to scale) .......................................... 22

11. Test of W heater element at 3 kW (a) early in test and (b) later in test, revealing 

a spectral shift ........................................................................................................... 23

12. Alumina mandrel after removal of W wire, revealing (a) significant surface 

discoloration and (b) close-up view of contamination ............................................... 25

13. Test of Re heater element at (a) 5 kW and (b) 1.2 kW ................................................. 26

14. Posttest Al2O3 mandrel (a) before and (b) after removal of Re wire, revealing 

slight surface discoloration ........................................................................................ 27
vi

LIST OF FIGURES (Continued)

15. Tungsten wire-braid heater element wrapped on a 1.04-cm (0.410-in) Al2O3

mandrel. Close-up views are provided for (a) the power input end, (b) the 

nonpower end, showing wire wrap return for double helix configuration, 

and (c) the center section showing the five-wire braid ................................................ 29

16. Approximate TC locations on the Mo sheath ............................................................ 29

17. Approximate two-dimensional model applied for W wire element thermal 

analysis (not to scale) ................................................................................................. 31

18. Braid/mandrel/sheath geometry: (a) Annotated layout and (b) finite element 

model of determination of conduction shape factors ................................................. 32

19. Double-ended C fiber braid trail: configuration for rough checkout tests .................. 36

20. Carbonized C fiber braid wrapped in a double helix fashion around 

Al2O3 mandrel ........................................................................................................... 37

21. Basic configuration of the proposed heater/sheath geometry (not to scale) ................ 43

22. Approximate geometry used in the SINDA model .................................................... 47

23. Axial temperature profile for case 2 ........................................................................... 52

24. Axial temperature profile for case 3 ........................................................................... 53

25. Wire transition from the spiral wrap at end of mandrel (complete mandrel design 

with dimensions is included in figure 65) ................................................................... 55

26. Posttest TA heater assembly showing wire failure ..................................................... 56

27. Initial test data for a 0.5-mm (0.020-in) TA three-wire braid heater element 

in a He environment (the element failed at the start of the subsequent test day) ........ 57

28. Tantalum heater element power and resistance on subsequent test days 

for 0.64-mm (0.025-in) three-wire braid: (a) initial test data, July 1, 2005, and 

(b) final test data, July 5, 2005 ................................................................................... 58

29. Tantalum heater element power and resistance on the initial and final test days 

for 0.5-mm (0.020-in) three-wire braid: (a) initial test data, July 6, 2005, and 

(b) final test data, July 19, 2005 .................................................................................. 60
vii

LIST OF FIGURES (Continued)

30. Tungsten wire heater assembly—testing in a vacuum environment: (a) End view 

of element during test and (b) side view of element during test .................................. 61

31. Mandrel distortion due to excessive heating .............................................................. 61

32. Initial test data (day 1) for W three-wire braid: (a) Heater power chamber pressure 

and (b) element resistance and measured chamber and mandrel temperatures ......... 62

33. Final test data for W three-wire braid, 89–100 hr: (a) Heater power and chamber 

pressure and (b) element resistance and measured chamber and mandrel 

temperatures .............................................................................................................. 63

34. Test setup with sheathed heater element fitted inside a water-cooled calorimeter ...... 64

35. Initial and final test data for 0.25-mm (0.010-in) three-wire braided W element: 

(a) Initial test data, July 25, 2005, and (b) final 13 hours of test, August 8, 2005. 

 Plots show total power delivered to the heater element and the calculated 

element resistance ...................................................................................................... 65

36. Test data for 0.25-mm (0.010-in) five-wire W braid (August 25–29, 2005) .................. 66

37. Test data for 0.25-mm (0.010-in) five-wire braid (Test 1, September 13–16, 2005) ....... 67

38. Test data for 0.25-mm (0.010-in) five-wire braid (Test 2, September 16–20, 2005) ...... 68

39. Test data for 0.25-mm (0.010-in) five-wire braid (Test 3, September 16–26, 2005). 

Data were broken into two separate plots due to limitations in Microsoft Excel 

plotting features: (a) Test data, 0–84 hr and (b) test data, 84–140 hr .......................... 69

40. Color optical pyrometer setup ................................................................................... 74

41. Closeup of Mo heater ................................................................................................ 74

42. Sample long-term pyrometer data ............................................................................. 75

43. Sample data with oxidation ....................................................................................... 76

44. Test chamber gas mixture and purification system .................................................... 78

45. Inert gas purification system connected to test chamber ........................................... 80

46. Heater element calorimeter layout ............................................................................. 82
viii

LIST OF FIGURES (Continued)

47. Calorimeter coil fabrication tests .............................................................................. 83

48. Spring-loaded TC unit ............................................................................................... 84

49. Calorimeter test chamber .......................................................................................... 85

50. Calorimeter instrumentation ..................................................................................... 85

51. LabVIEW thermal simulator application .................................................................. 88

52. LabVIEW strip-chart application .............................................................................. 89

53. LabVIEW multigauge server ..................................................................................... 89

54. Flowmeters: (a) LabVIEW gas flowmeter server and (b) installed thermal 

mass flowmeter .......................................................................................................... 90

55. LabVIEW data recorder ............................................................................................ 90

56. Power interface for the DDG-cooled reactor tests ..................................................... 91

57. Modified power interface conceptual design, using 37-pin DDG as a 

baseline concept ........................................................................................................ 92

58. Heater element placement inside DDG core block and TCs would be embedded 

in the Al2O3 spacer region ......................................................................................... 98

59. Thermocouple location on graphite heater element for preliminary testing .............. 99

60. Generalized simulator assembly with core block: cross-sectional view 

(not to scale) .............................................................................................................. 100

61. Heater element with a shifting axial power profile ..................................................... 104

62. Mandrel design for a heater element with a hexagonal cross section ......................... 104

63. Varying diameter graphite rod heater element, 0.375-in OD ...................................... 107

64. Constant diameter graphite rod heater element, 0.305-in OD .................................... 108

65. Mandrel, 0.625-in assembly design ............................................................................ 110
ix

LIST OF FIGURES (Continued)

66. Coupler, 0.625-in assembly design ............................................................................. 111

67. Sheath, 0.625-in assembly design ............................................................................... 112

68. Endcap, 0.625-in assembly design .............................................................................. 113

69. Endcap assembly, 0.625-in assembly design ............................................................... 114

70. Full assembly, 0.625-in assembly design .................................................................... 115

71. Mandrel, 0.400-in assembly design ............................................................................ 117

72. Coupler, 0.400-in assembly design ............................................................................. 118

73. Endcap, 0.400-in assembly design .............................................................................. 119

74. Sheath, 0.400-in assembly design ............................................................................... 120

75. Resistivity of selected refractory metals as a function of temperature 

(data obtained from R. Kapernick, Los Alamos National Laboratory)..................... 121

76. Thermal conductivity of selected gases ..................................................................... 152

77. Axial temperature profile for case 1 ........................................................................... 169

78. Axial temperature profile for case 2 ........................................................................... 169

79. Axial temperature profile for case 3 ........................................................................... 170

80. Axial temperature profile for case 4 ........................................................................... 170

81. Axial temperature profile for case 5 ........................................................................... 171

82. Axial temperature profile for case 6 ........................................................................... 171

83. Cold wall/calorimeter primary tube and fluid circuit ................................................ 174

84. Predicted cold wall temperatures for the case of 6,000 W, He atmosphere, 

19 L/min (5 gal/min) water flow ................................................................................. 175

85. Water temperatures for the case of 6,000 W, He atmosphere, 19 L/min 

(5 gal/min) water flow ................................................................................................ 176
x

LIST OF FIGURES (Continued)

86. Sheath temperature profile for the case of 6,000 W, He atmosphere, 

19 L/min (5 gal/min) water flow ................................................................................. 177

87. Heater temperature profile for the case 6,000 W, He atmosphere, 19 L/min 

(5 gal/min) water flow ................................................................................................ 178

88. Temperature vs. axial position for the sheath, heater, and cold wall for the case 

of 6,000 W, He atmosphere, 19 L/min (5 gal/min) water flow. (Axial position is 

the position along the longitudinal axis of the heater/sheath and zero represents 

the end of the cold wall) ............................................................................................. 179

89. Temperature profile for TC error model; temperature profile corresponds to a 

bead contact conductance of 0.02 W/°C .................................................................... 183

90. Close-up view of exposed TC wires in vicinity of the sheath and primary tube 

cold wall; temperature profile corresponds to a bead contact conductance of 

0.02 W/°C ................................................................................................................... 184

91. Close-up view of conductor between the TC bead and sheath wall; temperature 

profile corresponds to a bead contact conductance of 0.02 W/°C .............................. 185

92. Comparison of TC error versus bead contact conductance ....................................... 186

93. Cross-sectional view showing heater element pin formation for modified 

power interface .......................................................................................................... 187

94. Modified (conceptual) power interface showing positive and negative leads to 

heater element power terminals ................................................................................. 188

95. Modified power terminal connections for a single pin ............................................... 189

96. Modified power terminal connections for a single pin; axial view ............................. 189

97. Modified power interface conceptual design showing series connection between 

neighboring pins ........................................................................................................ 190

98. Oxygen concentration (ppm by volume) in the heat pipe test chambers following 

successive dilutions with (a) UHP He or (b) UHP Ar fill gas ..................................... 196

99. Conceptual design, varying power profile ................................................................. 199

100. Conceptual design, varying power profile ................................................................. 200
xi

LIST OF FIGURES (Continued)

101. Conceptual design, varying power profile ................................................................. 201

102. Conceptual design for hexagonal cross section, spiral-wound mandrel 

heater element ........................................................................................................... 203
xii

LIST OF TABLES

 1. Identification of basic thermal simulator design requirements .................................. 3

 2. Thermal simulator test matrix: current status ........................................................... 14

 3. Maximum calculated element resistance at 140 V compared to measured 

resistance values at prescribed power levels ............................................................... 19

 4. Summary of resistance calculated for select refractory metal wires at 20 °C ............. 20

 5. Measured sheath temperatures (K) for the braided W wire heater assembly. 

(The first temperature entry corresponds to the beginning of the test, the 

and the second to the end of the test) ......................................................................... 30

 6. Calculated mandrel, wire, and sheath temperatures for the braided W wire 

thermal simulator ...................................................................................................... 33

 7. Estimated costs* for additional heater element wire .................................................. 34

 8. Effective gap conductance for different inert gases based on geometry depicted 

in figure 22 (radial gap size: 0.15-cm (0.058-in) properties at 1,273 K) ........................ 45

 9. Relative effect of radiation and gas conduction for baseline simulator geometry ...... 46

10. Thermal performance of combinations of optical property enhancements 

(assuming radiation heat transfer only) ..................................................................... 47

11. Performance of conductive gap filler concepts ........................................................... 48

12. Impact of the sheath on heater temperatures experienced in vacuum testing ............ 49

13. Summary results from SINDA heater simulation (temperatures are in K) ................ 51

14. Suggested materials test matrix. (Work would entail an initial literature review 

to assess potential compatibility issues at certain test temperatures and 

durations, followed by testing of combinations of interest) ........................................ 70

15. SAES MonoTorr phase II 3000 performance guarantee for rare gases ...................... 79

16. Desired advanced thermal simulator characteristics ................................................. 102
xiii

LIST OF TABLES (Continued) 

17. Summary of results from SINDA heater simulation 

(temperatures are given in Kelvin) ............................................................................. 168

18. Comparison of temperatures for various atmospheres and conditions ...................... 180

19. Bounding contact coefficient and conductance between TC bead and 

sheath surface ............................................................................................................ 181
xiv

LIST OF ACRONYMS AND SYMBOLS

Al aluminum

Al2O3 alumina

AlN aluminum nitrate

Al-Ta aluminum-tantalum

Ar argon

BN boron nitride

C carbon

CH4 methane

Cl chlorine

CO carbon monoxide

CO2 carbon dioxide

CTE coefficient of thermal expansion

Cu copper

CVD chemical vapor deposition

DDG direct drive gas

EDS Energy Dispensive Spectroscopy

EFF-TF Early Flight Fission Test Facility

EWI Edison Welding Institute

FBG fiber Bragg grating

FEM finite element model

FESEM Field Emission Scanning Microscope

FLUINT Fluid Integrator
xv

LIST OF ACRONYMS AND SYMBOLS (Continued)

GC gas cooled

GHe gaseous helium

H2 hydrogen gas

HCl hydrochlorine

He helium

Hf hafnium

HF hydrofluoric acid

HF/HNO3 aqua emporis

H2O water

H2O2 hydrogen peroxide

HP heat pipe

ID inner diameter

IR infrared

LM liquid metal

Mo molybdenum

MRI Materials Research Institute

MSFC Marshall Space Flight Center

MWt mega-watt thermal

N nitrogen

N2 nitrogen gas

Na sodium

NaOH sodium hydroxide

Nb niobium

NB-1%Zr niobium-zyrconium
xvi

LIST OF ACRONYMS AND SYMBOLS (Continued)

Ni nickel

NRPCT Naval Reactors Prime Contractor Team

N2 nitrogen gas

O2 oxygen gas

OD outer diameter

ORNL Oak Ridge National Laboratory

PPF power peaking factor

Re rhenium

Re2O7 rhenium oxide

RGA residual gas analyzer

SAFE safe, affordable fission engine

SBIR small business innovative research 

SINDA Systems Improved Numerical Differencing Analyzer

SLIM standard liters per minute

SP space power

Ta tantalum

TC thermocouple

TF trichlorotrifluorethane

TIG tungsten inert gas

TM Technical Memorandum

UDP user datagram protocol

UHP ultra-high purity

VPS vacuum plasma spray

W tungsten

Xe xenon
xvii

NOMENCLATURE

A area

CP specific heat

D diameter

E or ε emissivity function

f factor of exchange

I current

k conductivity

L length of gap

l length of wire

N number of surfaces

P power

Q heat transfer; input power

R resistance

S conductive shape factor

T temperature

V voltage

σ Stefan-Boltzman constant

ρ density

r resistivity
xviii
1

TECHNICAL MEMORANDUM

HEATER DEVELOPMENT, FABRICATION, AND TESTING:

ANALYSIS OF FABRICATED HEATERS

1. INTRODUCTION—Historical Thermal Simulator Development

The thermal simulators (highly designed heater elements) developed at the NASA Marshall 

Space Flight Center (MSFC) Early Flight Fission Test Facility (EFF-TF) are used to simulate the 

heat from nuclear fission in a variety of reactor concepts. When inserted into the reactor geometry, the purpose of the thermal simulators is to deliver thermal power to the test article in the same 

fashion as if nuclear fuel were present. Considerable effort has been expended to mimic heat from 

fission as closely as possible. To accurately represent the fuel, the simulators should be capable of 

matching the overall properties of the nuclear fuel rather than simply matching the fuel temperatures. This includes matching thermal stresses in the pin, pin conductivities, total core power, and 

core power profile (axial and radial). 

1.1 General Requirements

The initial development of the simulators was driven by past space reactor work and 

bounding parameters were purposely chosen to be as challenging as possible. For example, the 

target pin diameter was selected based on the previously developed space power-100 (SP-100) space 

reactor fuel. This would require simulator sizes to be as small as 0.648 cm (0.255 in) in diameter. 

Peak thermal simulator power was based on a very high core thermal power (>2 MW thermal 

(MWt)). Additionally, the simulators must be isolated from the core to prevent electrical shorting and to prevent contamination of the test article by the simulator at high temperatures (due to 

sublimation, outgassing, etc.). Operational requirements for the thermal simulators incorporate 

desired lifetime, thermal cycling, and test environment. The development path for the EFF-TF 

thermal simulators required that the simulators withstand thousands of hours of operation and 

hundreds of thermal (on/off) cycles. The simulators were also required to operate in a vacuum 

environment (≈0.1 to 10–4 Pa; ≈10–3 to 10–6 torr) in the initial simulator checkout testing or in the 

presence of a low-pressure, high-purity gas (helium (He), carbon dioxide (CO2), Argon (Ar), etc.) 

environment. 
2

Significant development was required to determine the best method of providing an electrical hookup to the thermal simulators in all reactor concepts. For the heat pipe- (HP-) cooled 

reactor concept, the simulator leads/power input are not required to penetrate a pressure vessel, 

making hookup much easier to accomplish. However, for both the pumped liquid metal- (LM-) 

cooled and direct gas-cooled (GC-) reactor concepts, power feedthroughs (or the heater/sheath 

assembly in the LM system) must penetrate a pressure vessel to provide power to the simulators. 

In addition, if the simulator leads penetrate into a LM system, then the simulators must be configured to be compatible with the LM both from a materials and an electrical standpoint. This 

becomes increasingly difficult given the electrical conductivity and chemical activity of the LM, 

which can lead to heater shorting or the introduction of impurities into the system. The electrical leads on the simulators must be designed to not be larger (radial footprint) than the simulator 

itself. Due to the large number of simulators that must be integrated in a relatively small footprint 

(as many as 400 in a 38- by 38-cm (15- by 15-in) footprint), the design of the simulator must also 

be incorporated with the core layout and integration process. 

As in most hardware development programs, cost was also a major driver in the development of the thermal simulators used in past electrically heated reactor testing programs at the 

EFF-TF. The development goal was to produce complete heater assemblies at a cost of less than 

$1,000 per simulator; however, this goal may not be achievable for all concepts as the total simulator cost is highly dependent on materials and operating requirements.

Because each reactor concept has different requirements (pin power, axial power profile, 

operating environment, etc.), a single simulator design may not meet the needs of each test concept. Instead, each reactor concept must be approached individually, taking into account the basic 

requirements for the simulator and integration into the core. A decision can then be made for each 

concept, selecting the lowest cost design that is technically acceptable. For example, although the 

team strove to develop simulators with a 0.648-cm (0.255-in) diameter, the safe, affordable fission 

engine (SAFE) 100 core did not require this small simulator size. (The SAFE 100 HP-cooled reactor 

concept has a fuel tube clad with an inside diameter (ID) of 1.4 cm (0.55 in).) Using this approach, 

a graphite heater with alumina (Al2O3) spacers (used to fill the gap between the graphite and the 

internal clad surface, providing electrical isolation and centering the simulator within each fuel tube) 

was selected for the SAFE 100 test series because it was able to meet the technical requirements of 

the test plan at a cost of $200 per heater.

Each new reactor concept that is investigated maintains the basic parameter set identified 

in table 1 when selecting an appropriate simulator design. A successful simulator does not have to 

meet all of these parameters simultaneously, but may require a subset for a specific application.
3

Table 1. Identification of basic thermal simulator design requirements.

Parameter Target Requirement

Temperature 1,400 K (clad) 

Linear heat rate ≈100 W/cm (power per length)

Power density ≈100 W/cm3

Lifetime 10,000+ hours (per simulator, tested individually or in the reactor core concept)

Number of thermal cycles 200 or more (cycles to full temperature and power)

Axial power profile Average pin power peaking of 1.33, cosine distribution

Simulator diameter ≈0.65 cm (0.255 in) and up

The fundamental minimum size achievable for a given design is limited by the area available for power hookup 

and the number of pins that must be connected in a relatively small footprint.

Pin conductivity Match effective fuel pin conductivity—design/age dependent

Pin voltage ≤150 Vdc to avoid voltage breakdown in low-pressure environment

Electrical connections Single ended (power access to only one end of the heater)

Fabrication repeatability Simulator concepts must be sufficiently robust so that minor changes in materials, impurities, or process do not 

significantly affect lifetime or performance of the simulator

Number of pins ≈200 to 500 (per concept/core footprint, operating simultaneously)

Operating environment Vacuum, CO2, GHe, Ar, N2

Compatibility Thermal simulators must not introduce any significant impurities into the test article

To simplify insertion in each reactor test article and to minimize the impact of the thermal simulators on the ability to achieve a prototypic nonnuclear test article, a single-ended heater 

design was selected for use at the EFF-TF. The graphite rod essentially acts as a large resistor, with 

the rod itself axially split down the middle. Power is input at one end of the graphite, and current 

flows along the length of the rod, returning along the opposite half. This configuration is sketched 

in figure 1(a). Alumina pieces are inserted along the center of the two halves of graphite to prevent contact, which would short the element. Additional Al2O3 insulator rings are used at three 

points along the length of the graphite heater element to electrically isolate the heater from the test 

article. A photograph of the complete graphite heater assembly is provided in figure 1(b). 

Al2O3 Insulator Rings (×3)

(Cemented to Graphite) Mo Power Leads

(Cemented into Graphite)

Al2O3 Insulator Shims (×2)

(Cemented to Graphite)

Graphite Heater Element (Split Rod)

(a)

(b)

Figure 1. Graphite rod heater element: (a) Sketch and (b) photograph.
4

In a single-ended design, both power input and output are at the same end of the heater element. To connect the power to the element, molybdenum (Mo) couplers are cemented directly into 

the graphite rod using graphite cement. The coupler reduces the resistance from the interface back 

to the power source to ensure that a majority of the input power goes into the heater itself. The 

coupler serves three purposes:

(1) It acts as the interface to the hot heater wire. 

(2) It reduces the resistance between the power supply and the heater element as a result 

of its material properties.

(3) It lowers the temperature seen by the copper (Cu) lead wires. (The magnitude of the 

reduction is based on the length of the coupler pieces.)

The cold resistance (at room temperature) of the 0.95-cm (0.375-in) graphite rod is ≈1 Ω. 

As the graphite heats up, it initially decreases in resistance, reaching a minimum of ≈0.5 Ω; the 

resistance again increases as temperature further increases, but never fully recovers to the roomtemperature resistance. The characteristic full-load impedance of the power supply used to power 

the heater element is 1.5 Ω, therefore, when a single element is powered by the power supply, it 

operates in a current-limited fashion. To increase the resistance in the test article, the graphite rod 

heaters can be connected in series. Power supply use is most efficient when the resistance of the 

load matches the impedance of the power supply, allowing operation up to the full voltage and 

current capabilities of the supply (150 V, 100 A).

Graphite rod heaters can be designed to provide a flat axial power profile or a prescribed, 

shaped (cosine) axial power distribution. Constant diameter heaters provide uniform heat flux 

along their length. By cutting out a portion of the graphite along the axial center of the heater 

(using either a tapered cut or a squared-off cutout along a portion of the heater length), the power 

profile can be made to simulate that of an operating reactor. The initial set of varying diameter 

heaters had a 0.33-cm (0.130-in) diameter along a 10.2-cm (4.0-in) section at the axial center of the 

heater (engineering schematic provided in app. A). This degree of variation was adequate to show 

that anticipated requirements could be met. Tests have been conducted at the EFF-TF using both 

constant and varying diameter graphite rod heaters. Total operating hours (combined time of all 

individual simulators in all tests) for these simulators is in excess of 20,000 hr and hundreds of 

thermal cycles and a single set of three graphite rod heaters used in cyclic testing of a sodium HP

performed well over 280 thermal cycles and no graphite rod heater has failed in any test to date. 

Figure 2 shows typical graphite elements with a varying axial power profile (center taper). To measure the temperature of the heater element at power, a type C thermocouple (TC) was sandwiched 

between layers of synthetic sapphire (Al2O3 having a single crystal structure versus a multicrystal 

structure) and was inserted at the axial center of the graphite rod, between the split halves of the 

element. The measured element temperature was ≈2,275 K at 5 kW in a vacuum environment.
5

Graphite Heater Assembly

• Single End

• 0.375-in-Diameter Element

• Al2O3 Insulators

Al2O3

Insulator/Spacer

Mo Leads (×2)

(a)

(b)

Center Taper

Providing Axial 

Power Profiling

Figure 2. Graphite rod heater elements: (a) Batch of graphite rod heater elements and 

(b) identification of center taper geometry and power input design via Mo leads.

To date, all tests using the graphite rod heater elements have been 0.95 cm (0.375 in) in diameter (constant or varying), using 1.36-cm (0.535-in) Al2O3 insulator rings to provide electrical isolation from the test article (1.37-cm (0.540-in) fuel tube ID). A more recent graphite rod element design 

reduces the graphite diameter from 0.95 to 0.77 cm (0.375 to 0.305 in), with the outer diameter (OD) 

of the Al2O3 insulating rings of 1 cm (0.410 in). The overall length of the graphite element is 59.7 cm 

(23.5 in). These reduced diameter elements will be employed in the 37 pin LM-cooled reactor design 

that is currently being prepared for test at the EFF-TF. As will be discussed in a later section, the 

power integration into the LM design ensures that alkali metal cannot contact the thermal simulator. Detailed design drawings of the graphite heater elements are included in appendix A.
6

The minimum diameter that can be achieved for graphite elements is limited. The current 

design requires that the Mo couplers (0.32-cm (0.125-in) OD) be screwed directly into the graphite element, and reducing the total element diameter to much less than 0.77 cm (0.305 in) presents 

additional concerns in how to integrate the power. An alternate coupler design might allow the 

coupler cap to slip over the outside of the heater element to integrate the parts, but no design has 

been completed. The strength of the graphite material at reduced diameters also presents a concern, and the manufacturer of the graphite elements (POCO Graphite, Inc.) is reluctant to produce 

elements of significantly reduced diameter. The current design for the shaped graphite heater elements reduces the diameter of the element to just 0.33 cm (0.130 in) for the center and 10.2 cm (4 in) 

along the element’s length. A reduced diameter graphite element may not have sufficient material 

to allow for axial shaping, although this is being investigated further. 

To date, over 350 of these graphite rod simulators have been fabricated and tested with zero 

failures occurring within the element itself. In some cases, simulator refurbishment was required 

to reinstall power leads after multiple test cycles. This led to a redesign of the method of attaching the leads to the simulator elements in order to increase their strength over long periods of 

testing. The constant diameter graphite rod heater elements were used in the HP-cooled SAFE 30 

kW thermal (kWt) test series (0.95-cm- (0.375-in-) OD elements) and will be used in the test circuit 

for the 37 pin liquid metal-cooled reactor design (0.77-cm- (0.305-in-) OD elements). The varying diameter graphite rod heater elements were used in the SAFE 100 (100 kWt) test article, the 

SAFE 100a test article (partial array HP core with gas flow heat exchanger), and the 37 pin direct 

drive gas- (DDG-) cooled test article. See VanDyke et al.,1–5 VanDuyn et al.,6 Godfroy et al.,7 and 

Bragg-Sitton and Forsbacka8 for additional details on some of these systems. 

The electrical integration of the heater elements (thermal simulators) into the reactor core 

test article has evolved significantly from their introduction in the SAFE 30 test article to their use 

in the GC reactor design and their planned integration into the LM-cooled design. Integration 

is a function of the reactor type and operating environment. Simplifying the integration process 

continues to improve by using more flexible wire for power inputs and by developing a capability 

to more easily change out a control zone. The lessons learned from one reactor concept are typically applicable to other concepts. In developing an electrical interface to the core, it is important 

to take into account the total number of heater elements that must be powered in a relatively small 

footprint, where complexity significantly increases as the pin size is reduced and the total number 

of pins increases. Figures 3 though 6 show the integration of the thermal simulators in various 

core configurations.
7

Figure 3. Integration of 48 simulators in SAFE 30 (≈23- by 20-cm (≈9- by 8-in) footprint.

Figure 4. Integration of 183 simulators in SAFE 100 (≈26- by 29-cm (≈10.4- by 11.5-in) footprint).

Figure 5. Integration of 57 simulators in SAFE 100a (≈18- by 16.5-cm (≈7- by 6.5-in) footprint).
8

Figure 6. Integration of 37 simulators in DDG test article (≈16- by 18-cm (≈6.25- by 7.1-in) 

footprint). (Note: This core is inserted into a pressure vessel where the wires are 

not accessible after test article integration.)

The DDG reactor concept uses an electrical interface configuration that can be sealed after 

core integration and ‘flooded’ with gaseous helium (GHe) to fill the gaps between the thermal simulators and the clad, where the ‘clad’ is the inner surface of the test article (also referred to as a ‘fuel 

tube’). A proof-of-concept configuration experiment for a core face seal that would be integrated 

with an HP reactor design is shown in figure 7. Since this unit was not planned for test with an actual 

core, the design includes a series of heaters mounted along the back plane to provide a heat source 

(core simulator). The electrical connections enter from the side allowing front access to complete 

the integration process. A ‘flange’ would be sealed on the front of the apparatus and gas would be 

supplied through the top. Note that the back plate of the fixture simulated a core in this experiment. 

The actual back plate would be removed and the apparatus sealed around a lip of the core.
9

Internal Thermocouple

Feedthroughs (6 TCs)

Power Feedthroughs

• VCR Fitting

• 0.09-in-Diameter Copper Pin

• 12 Feedthroughs Used

Pressurant

Port

Pressure

Transducer

Port

Heater

Elements

(×7)

Front Cover

Plate

Conflat Style

Front Cover

Plate

Vacuum System Line (a)

(c)

(b)

Thermocouples

Figure 7. Core face seal unit as configured for testing: (a) assembly on the test 

stand and placement of feedthroughs, (b) conflat style front cover 

plate, and (c) complete test configuration (without insulation shown).

The core face seal concept is being implemented in the LM-cooled reactor concept that is 

being prepared for testing at the EFF-TF (see fig. 8). Due to the electrical conductivity of the LM, 

the power leads cannot be in contact with the coolant. To prevent the heaters from shorting, the 

fuel clad structure in the stainless steel test article was extended through the plenum. Each tube 

was then welded to the face seal to prevent leakage. With this configuration, the heaters could then 

be installed into each fuel tube without contacting the liquid metal flow.
10

(a) (b)

Fuel Tube

Ports for Power

Feedthrough

LM Core

Power

Feedthroughs

Figure 8. Integration of power with the LM core design using a core face seal concept: 

(a) core face seal extension showing power feedthroughs and (b) identification 

of fuel tubes that house the resistive heater elements.

Not only are improvements constantly being made to simplify integration, but investigation is also ongoing to determine how to better match thermal pin diffusivity (where diffusivity 

is calculated as the conductivity divided by the product of the density and heat capacity), among 

other properties, to that of nuclear fuel without affecting the test article. Steady state and dynamic 

thermal properties of the pin could be better matched to that of a nuclear fuel pin by modification 

of the complete heater element design. A proof-of-concept test using an updated graphite heater 

element design is currently being investigated. In this concept, the heater is sealed within a stainless steel tube using a high-temperature braze, backfilled with GHe, and then electrical leads are 

attached. If successful, this complete thermal simulator assembly will not only be a demonstration 

of the manufacturing process developed to fabricate the heater elements, but it will also demonstrate the ability of sealing an element within a tube identical to the cladding of the core structure 

(stainless steel, in the SAFE 100a case), ensuring there will be no material incompatibility between 

the simulator and the core. The first application makes use of a stainless steel sheath at relatively 

large pin diameters (0.95 cm (0.375 in) for the graphite rod, and 1.4 cm (0.540 in) for the sheath OD) 

to prove the manufacturing concept. In general, if the sealing method works with stainless steel, 

then a similar approach can be applied to refractory metal (Mo) or other metal alloy sheaths/

tubes. After a lengthy series of testing on the stainless steel/graphite assembly, the design can be 

updated to accommodate the desired materials. Thermal analyses of the sheathed heater element 

design and potential refractory metal element designs are discussed in section 2. 
11

1.2 Early Refractory Metal Heater Designs

Early investigation of potential refractory metal heater designs was conducted in 2002. The 

primary goal of the development effort was to build a very small diameter heater element that 

could achieve the power and temperatures necessary for the reactor designs that were being proposed at that time. Very small diameter elements are possible using a hairpin design. In this construction, a double-bore Al2O3 tube is used, that allows a wire to be run down one bore, be bent, 

and returned down the opposite bore, hence the ‘hairpin’ terminology. Figure 9 provides a rough 

sketch of this design. 

Mo-Re Wire

Al2O3

Figure 9. General configuration of the hairpin heater element design.

Several hairpin heater tests were conducted using molybdenum-rhenium (Mo-Re) wire of 

varying diameters. A 50/50 mixture of Mo-Re was selected to increase the ductility of the material 

over that of pure Mo. However, it was noted that, for larger diameter wires, the ductility at room 

temperature was still such that it was difficult to make the necessary bend to return the wire down 

the opposite bore. Heater fabrication at elevated temperature was not investigated. Tested wires 

ranged from 0.64 to 1.3 mm (0.025 to 0.05 in) in diameter. Each heater element required ≈0.6 m (2 ft) 

of wire, as opposed to the 3.7 m (12 ft) required for a spiral-wound heater design. Because of the very 

low resistance of the element (additional wire length is required to increase the resistance), the maximum power that could be deposited in these elements for the available 100 A of current was very 

low. For a 15-kW (150-V, 100-A) power supply, the maximum power that could be achieved with the 

hairpin design was ≈1,600 W. Because the wire was fully surrounded by Al2O3 (there was no active 

heat removal from the element), wire temperatures were very high at these power levels. Additionally, 

a flat axial power profile is inherent to the hairpin design, and no flexibility is available to introduce 

a shaped profile. 

The 1.1-mm (0.045-in) Mo-Re wire element, tested up to a maximum of 1,600 W, was submitted for complete materials analysis by the Metallurgical and Failure Analysis Team at MSFC. 

After removal from the Al2O3 tube, the Mo-Re wire was found to be brittle with a pitted surface. 

Analysis found Mo and some Re deposits on the surface of the Al2O3 insulator, but cross-section 

analysis of the insulator found no reaction between the wire and insulator and no reaction zone was 

found in the Al2O3. The cross section of the Mo-Re wire revealed large grain growth, resulting in 

the embrittlement of the material, and a surface reaction zone that was depleted of Mo-Re. Mo-Re 

grain growth is dictated by time at temperature. Mo-Re is typically fully annealed at temperatures above 1,600 K. Recrystallization of Mo-50%Re can occur at 1,473 to 1,673 K, with variation 

in the recrystallization temperature likely a result of processing history (material microstructure) 
12

and material thickness. The Mo-Re material used in this element test was probably recrystallized 

material because significant grain growth would not be expected below 1,473 K for short times, but 

could occur for very long times. Because large grain growth was observed in the posttest analysis, it 

is likely that the element temperature exceeded 1,473 K for some period during the test. It was noted 

that the preferential oxidation of Mo was likely accelerated by the presence of water vapor in the 

test chamber, indicating the need for multiple purge cycles of the test chamber using dry nitrogen 

(N) or, preferably, high purity He or Ar gas before testing. The addition of a bakeout cycle would 

also help to remove volatile constituents before testing at high temperature.

Because of the various limitations described for the hairpin heater design, it was abandoned in favor of designs that offered additional flexibility in their application and that could 

achieve higher power in each pin.
13

2. CURRENT THERMAL SIMULATOR TESTING AND DEVELOPMENT

Investigation has commenced to assess alternate thermal simulator designs to replace the 

graphite rod heaters discussed in section 1. While the graphite heaters can accommodate most of 

the desired characteristics given in table 1, the limitations in the achievable pin size and the potential for graphite to contaminate a refractory metal test article has driven the simulator research 

effort toward other design concepts, such as refractory wire wrapped heater elements. Preliminary 

results for a carbon (C) fiber braid element will also be discussed. A matrix of the current and 

planned test hardware is provided in table 2.

2.1 Refractory Metal Wire-Wrapped Simulator Design

Refractory metal wire-wrapped heater element designs are investigated to develop a heater 

that will be compatible with refractory metal reactor concepts and to enhance the minimum 

achievable pin diameter. The wrapped wire heater element design considers various refractory 

metal wires wrapped in either a single pass or double pass fashion around a spiral groove etched 

along an Al2O3 mandrel.

Initial tests use a mandrel having a constant pitch groove (for both the single pass and 

double pass helix), resulting in constant power density along the axial length of the heater element. The development effort will also consider cutting the mandrel with a constantly changing 

pitch to generate an axial power profile that mimics that of an operating reactor. Mandrels with a 

constantly changing pitch (corresponding to a prescribed equation determined by reactor designers at Los Alamos National Laboratory) have been produced to demonstrate manufacturability of 

the design, but the current work focuses on using the lower cost, constant pitch mandrels to assess 

materials performance for long-life thermal simulators. Several challenges addressed in concept 

development include the following:

• Identifying a wire manufacturer to make a small diameter refractory metal wire with adequate 

ductility to wrap around the mandrel without ‘unwinding’ once tension is released, but robust 

enough to withstand the long lifetime and high-temperature testing. 

• Designing the electrical connection to the integrated system.

• Achieving power densities >1 kW per element.

• Manufacturing a small diameter mandrel (0.65 cm (0.255 in) with a constantly changing, double 

pass pitch (to provide axial power shaping).

• Overcoming material compatibility issues.
14

Table 2. Thermal simulator test matrix: current status.

Simulator 

Type

Simulator 

Material

Configuration

Adhesives

(Leads, 

Insulators, etc.)

Axial Power 

Profile

(Peak to 

Avg)

Simulator 

Diameters

(Heater Plus 

Sheath, Wire 

Diameter, Full 

Assembly) Length

Chamber 

Environment*

Power 

Level/

Range

(W)

Accumulated 

Test Time

Thermal 

Cycles

Test Status

(Tested 

Configuration, 

Complete/

Planned Tests)

Sheathed 

vs. Bare 

Cover 

Gas

Poco 

Graphite

Graphite rod(Poco Graphite)

Bare rodMo coupler

None

Power leads: Graphite cementAl2O3 standoffs: Al2O3 epoxy

Varying diameter (4-in section cutout at center), ≈1.33

0.375-in-OD, 0.130 in at center 20 in ≈10–5 torr 850 – 5,000 1,000’s hr 100s

SAFE-30SAFE-100SAFE-100aDDG

Varying diameter (4-in section cutout at center), ≈1.33

0.375-in-OD, 0.130 in at center 20 in

Mars environment (CO2)

25 kW total, distributed over 57 heaters (≈440 W per Heater)

≈30 hr 10 SAFE-30

None (constant diameter)

0.375 in 20 in ≈10–5 torr 850 – 1,250 1,000’s hr 100s HP module tests

Brazed couplers/leads

None (constant diameter)

0.305 in 23.5 in Not yet tested Up to 5,000 W Not yet tested

—will be applied in LMR 

test

Graphite rod(Poco Graphite)

Stainless steel sheathed rodMo coupler

None (unsealed)

Power leads: Graphite cementAl2O3 standoffs: Al2O3 epoxy

None (constant diameter)

0.375 in 20 in ≈10–5 torr Not yet built or tested

Sealed 

Tube

Graphite rod(Poco Graphite)

Stainless steel sheathed rod

YesGHe

Seal: High temp. braze, RD material PALARO+3 VWesgo Metals (Morgan Advanced Ceramics)

None

0.375-in graphite,0.540-in stainless steel sheath

20 in Conceptual design

—Not yet built or tested

Wire 

Wrapped

W

Bare element, Al2O3mandrel

None Unsuccessful test due to poor ductility of W wire at larger diameters

—could not be wrapped around mandrel 

(0.040-in wire, with 0.410-in mandrel was attempted)

*Environmental conditions are estimated. No pressure gauges currently on the test chamber

—estimated from pump capabilities.
15

Simulator 

Type

Simulator 

Material

Configuration

Adhesives

(Leads, 

Insulators, etc.)

Axial 

Power 

Profile

(Peak to 

Avg)

Simulator Diameters

(Heater Plus Sheath, 

Wire Diameter, Full 

Assembly) Length

Chamber 

Environment*

Power 

Level/Range

(W)

Accumulated 

Test Time

Thermal 

Cycles

Test Status

(Tested 

Configuration, 

Complete/

Planned 

Tests)

Sheathed 

vs. Bare 

Cover 

Gas

Wire 

Wrapped

W braid

Bare element, Al2O3mandrel

None

Direct mechanical connection to power feedthrough

None

0.010-in-diameter wire, 5 wires, braided together, 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1200 – 3000 125 hr 16

Bare element testing (in progress)

W braid Mo sheathed None, (un- sealed) Direct mechanical connection to power feedthrough None

0.010-in-diameter wire, 5 wires, braided together, 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1200 – 3000 115 hr 12

Sheathed element testing (in progress)

Ta

Bare element, Al2O3mandrel

None

Direct mechanical connection to power feedthrough

None

0.040-in-diameter wire; 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1200 – 4000 150 hr 19

Bare element testing (in progress)

Re

Bare element, Al2O3mandrel

None

Direct mechanical connection to power feedthrough

None

0.040-in-diameter wire; 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1200 – 5000 140 hr 18

Bare element testing (in progress)

Hf

Bare element, Al2O3mandrel

None

Direct mechanical connection to power feedthrough

None

0.040-in-diameter wire; 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1200 – 3000 110 hr 13

Bare element testing (in progress) (did not stabilize

—eliminated)

General Comments: All wrapped wire configurations can be constructed to have the desired power profile by cutting the groove in the Al2O3 mandrel with a constantly varying pitch. In this manner, the axial power profile can be set to any equation desirable, as a function of the selected reactor design. Direct mechanical connection to electrical feedthrough will be replaced by a Mo coupler in future testing. All wire-wrapped configurations met the basic requirement of operating at 1,200 W for 100 hr.

C Fiber 

Braid

C fiber

Bare element, Al2O3mandrel

None Mo coupler None

0.1-in-diameter fiber braid; 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 900 35

Bare element testing (in progress)

C fiber, resin coated

Bare element, Al2O3mandrel, carbonized

None Mo coupler None

0.1-in-diameter fiber braid; 0.410-in mandrel (0.625-in assembly)

20 in 10–5 torr 1,200 10

Bare element testing (in progress)

*Environmental conditions are estimated. No pressure gauges currently on the test chamber; estimated from pump capabilities.

Table 2. Thermal simulator test matrix: current status (Continued).
16

Dozens of refractory metal wire heaters have been built and tested using various wire combinations. While several designs have proven successful, others have failed at high power levels 

due to the inability to efficiently remove heat from the element. Early element tests applied no 

active heat removal from the element. In this configuration, heat loss from the element is by radiation to the cool chamber walls. Because the test chamber was operated at vacuum, no conductive 

or convective cooling was possible. Under these conditions, the element temperatures can become 

very high. The spiral-wrapped refractory metal wire heater elements show promise for small pin 

diameters and excellent axial power profiling capabilities, and tests completed to date indicate that 

the maximum power per pin is only limited by the available power supply voltage. Modification 

of the boundary conditions in the test chamber to better mimic use of the element in the reactor 

simulator test article could allow higher pin power levels to be achieved without exceeding material temperature limitations in the heater elements (addition of heat removal mechanisms similar 

to those in the full test article); possible modifications will be discussed in section 2.3. Research 

and development continues on these element designs in an effort to resolve the issues that have 

been identified, and are reviewed in sections 2.4 through 2.6.

2.1.1 Engineering Design

Heater designs using spiral mandrels were pursued to reduce the overall diameter of the 

heater elements and to investigate the potential of using materials other than graphite in the 

thermal simulator applications to address compatibility concerns for each proposed test article. 

To date, materials tested on the spiral cut Al2O3 mandrels include Re, tantalum (Ta), niobium 

(Nb), tungsten (W), hafnium (Hf), and C fiber, where tests incorporate the use of single wires or 

multiple braided wires. The braided material allows the use of smaller wire diameters, providing 

ductility required for manufacturability. As a result, the braided designs widen the possible range 

of heater element materials and designs, producing flexibility in meeting requirements. The C fiber 

braid design and tests will be discussed in section 2.2.

Initial efforts in the spiral-wrapped heater designs have been directed at developing a better understanding of the performance of potential materials that can be used to construct these 

elements. This provides a knowledge base that can be called on to quickly design a heater that will 

best meet the test article requirements once they have been established. To date, dozens of refractory elements have been built and tested to assess the applicability range of material selection. The 

basic requirement for all spiral-wrapped heater element designs was to produce an element capable of providing stable operation for a minimum of 100 hr at 1,200 W. Tests have been conducted 

over a range of 850 to 5,000 W with the goal power range of 1,200 to 6,000 W per element. Final 

designs for use in refractory metal systems will incorporate a sheath to ensure that oxygen (O2) 

from the Al2O3 mandrel does not contaminate the test article.

2.1.1.1 Heater Assembly: Mandrel Design. The heater assembly designs entail four primary 

components: the heater, electrical connections, sheath, and end cap. For the spiral-wrapped wire 

heater designs, the ‘heater’ consists of a refractory metal wire wrapped around an Al2O3 mandrel. 

To minimize variables from one test to the next, a 1.59-cm (0.625-in) assembly size was used for all 

preliminary material testing, as components required for these assemblies were standard (reducing 

development cost). This corresponds to the engineering drawings provided in appendix B for the 
17

mandrel, coupler, and sheath. For this design, the OD of the Al2O3 mandrel is 1.04 cm (0.410 in), 

with the groove in the mandrel cut with a radius of 1.6 mm (0.063 in) in a double helix fashion. 

The 1.59-cm (0.625-in) dimension corresponds to the OD of the flange on the sheath used with 

this design. Testing completed to date has primarily focused on the unsheathed heater elements to 

establish basic materials performance characteristics. Sheathed testing has been initiated with the 

Ta and W wire braid designs. Engineering specifications on the mandrel (and the sheath) fabrication require that it be held to a straightness of 0.13 mm (0.005 in) over full length. Current designs 

include a 0.064-mm (0.0025-in) radial clearance between the largest diameter on the mandrel (the 

insulator at the end of the wire wrapped region) and the ID of the sheath.

Currently, the smallest full assembly diameter (OD of the sheath flange) design is 1.02 cm 

(0.400 in) and the design includes the coupler and power integration strategy within this footprint. 

The minimum diameter is limited by the amount of power that is required per pin and by the ability to hook up the power within the given diameter. Smaller diameter assemblies have been manufactured, but the power integration issues have not been fully addressed to date. The limitations 

related to and the status of these smaller diameter designs will be discussed briefly in section 2.1.4. 

The 1.02-cm (0.400-in) design is shown in the engineering drawings provided in appendix C for the 

mandrel, coupler, coupler cap, and sheath. For this design, the OD of the Al2O3 mandrel is 0.60 cm 

(0.235 in) (0.65 cm (0.255 in) at the nongrooved end sections), with the mandrel groove cut with a 

radius of 1.2 mm (0.048 in) in a double helix fashion. The 1.02-cm (0.400-in) dimension corresponds 

to the OD of the flange on the sheath used with this design. Although mandrels have been constructed with these dimensions, heater element testing for this configuration has not yet begun. 

Testing has entailed two basic mandrel designs: single-pass and double-pass pitch, and all 

tests to date have been performed on the larger 1.04-cm (0.410-in) mandrel corresponding to the 

1.6-cm (0.625-in) full assembly size. In all cases, the mandrel was constructed from 99.8% pure 

Al2O3. Alumina oxide is commercially available up to a purity of 99.99%. As heater development 

becomes more sophisticated, it may be desirable to obtain higher purity mandrel material to 

lengthen the operational lifetime of the heater elements and to minimize the potential of the heater 

components contaminating the test article.

While several heater elements have demonstrated successful performance, several elements 

‘broke down’ at high power levels, presumably due to over-temperature conditions. Early spiral 

wound heater designs used ‘one-pass’ pitch mandrels, such that the wire was spiraled down the 

length of the mandrel (on the OD) and returned down a hole in the center of the Al2O3 mandrel 

to allow for a single-ended heater element. Failures were noted in the center level at higher power 

levels. Because the center wire was fully surrounded by an Al2O3 insulator, failure was likely due 

to inadequate heat transfer, which resulted in extremely high temperatures in the center wire. To 

date, element tests have applied no active heat removal from the element with heat loss from the 

element being by radiation to the cool chamber walls (≈50 °C). The chamber walls are actively 

cooled by water, but the chamber end caps are not cooled and could be at elevated temperatures 

(≈150 °C). The test chamber was operated at vacuum, preventing the possibility of no conductive 

or convective cooling.
18

The failures noted for the single-pass design led to the development of the double-pass helical 

mandrel. The double-pass mandrel improved the heater performance by bringing the wire completely 

to the outside surface of the Al2O3 (improving the ability to remove the heat from the wire) and by 

increasing the length of the wire used to construct the element (from ≈2.4 to 3.65 m (≈8 to 12 ft)), 

increasing the total resistance of the element by a factor of 1.5. 

Prior to wrapping the Al2O3 mandrel with the desired wire, the mandrel and wire are 

cleaned with methanol and all components are handled with gloves. All tests are conducted in 

vacuum (≈1.3 mPa (≈10–5 torr)). At the start of each run, ≈50 to 100 W is input to the heater element and it is allowed to bake out for a few hours. Testing begins on the following day at 1,200 W 

or higher. (Generally, heater elements reach the target 1,200 W in <30 s.) More extensive cleaning 

procedures will be adopted for higher fidelity element testing and prior to heater integration into 

each test article. Suggested procedures will be discussed in section 3.6.

2.1.1.2 Heater Assembly: Wire Selection. The use of refractory metal wires 0.5 to 1.5 mm 

(0.020 to 0.060 in) in diameter has been assessed with some materials demonstrating fabricability 

issues early on and, as a result, these elements were not tested. Selection of the appropriate wire 

diameter is a balance of resistivity, ductility, and integrity at temperature. As the ductility of the 

wire decreases, it becomes increasingly difficult to wrap the wire down the length of the mandrel. 

Generally, while the larger diameter wires have decreased resistance (such that, in a voltage limited case, the element can pass more current and reach a higher power level for the same voltage 

relative to the small diameter wires), they face limitations in ductility that impact manufacturability. Smaller diameter wires, on the other hand, are more ductile, simplifying heater fabrication, 

but they have increased resistance, so they pass less current in a voltage limited scenario. Additionally, smaller diameter wires are more prone to failure at high temperature due to materials 

limitations (such as temperature-induced grain growth, degradation of material, corrosion, etc.). 

Braiding three or more small diameter wires mitigates the high resistance problem by offering 

parallel paths for the current, providing the capability to reach higher power levels similar to that 

achieved with a single, larger diameter wire while offering the ductility of small diameter wires. 

Use of braided wire configurations widens the range of potential materials that can be used to 

construct the heater elements. The element resistivity could also be adjusted using material doping, but this could generate ‘hot spots’ as a result of nonuniform doping, which could lead to wire 

failure. Maintaining high purity for any refractory wires can increase their operational lifetime. 

Any impurities that might exist in the wire material affect the resistance of the wire and generate 

higher power deposition and higher temperature at the point of the impurity. For a small diameter 

wire, the resulting increase in temperature at impurity points can lead to wire failure more quickly 

than for larger wire diameters. All refractory metal wires have been obtained at the highest purity 

level available, ≈99.9%. (With the exception of Re, all wire has been obtained from ESPI Corp., 

Inc. and Re wire was obtained from Rhenium Alloys, Inc.)

Of the refractory metals, Nb exhibits the most ductility and W the least ductility. Therefore, 

a large 1.5-mm- (0.060-in-) diameter Nb wire was acquired to conduct a feasibility study for using 

this wire size. Because significant difficulty was encountered in wrapping the 1.5-mm (0.060-in) Nb 

wire, other refractory metals, which would have reduced ductility, were not acquired at this large 

diameter. In this case, the Nb was used to ‘bound’ the range of possibilities, eliminating several 
19

potential options from the test scope without additional expenditure of time and funds. The 1.5-mm 

(0.060-in) Nb wire was successfully wrapped around the 1.04-cm (0.410-in) Al2O3 mandrel but has 

not been tested to date. 

The current test setup provides power to the heater element via a 150 V power supply at a 

maximum of 100 A (15 kW maximum). Although multiple power supplies could be used to power 

the heater elements to test to higher voltages, this has not been done to date because operation at 

higher voltages in a vacuum environment can lead to voltage breakdown. The amount of power 

that can be pushed into an element at a given voltage is set by the element’s resistance. From the 

basic Ohm’s law, power (P) is given by the product of the current (I) and the voltage (V), or it can 

be written in terms of the resistance (R) using the relationship that voltage is given by the product 

of the current and the resistance:

 

. (2)

To prevent operation of the power supply near its 150 Vdc limit, it is desirable to determine 

the maximum resistance element that can provide the desired power level at a more reasonable 

level of 140 V. This calculated value is provided in table 3. The right-hand side of the table provides 

experimentally measured resistances of the tested spiral wound elements. For the Hf and W wire 

braid tests at 3 kW, note that the input voltage was pushed to the absolute maximum for the power 

supply, 150 V. To date, most successful refractory wire elements have run for at least 30 hr at a 

power of 800 W (reaching a temperature of ≈1,300 K), while the 0.5-mm (0.020-in) wires appear to 

be limited to ≈850 W. 

Table 3. Maximum calculated element resistance at 140 V compared to measured resistance 

values at prescribed power levels.

Power Level

(W)

Current @ 140 V

(A)

Maximum 

Element 

Resistance 

@ 140 V

(Ω)

Measured Resistance During Test (Ω)

Ta

(1 mm, 0.040 in)

Re

(1 mm, 0.040 in)

Hf

(1 mm, 0.040 in)

W, 

Five-Wire Braid

(0.25 mm, 0.010 in 

No Sheath (× 5))

With 

Sheath

1,200 8.6 16.3 – 3.2

(63 V, 19 A)

3.6

(67 V, 19 A)

6.6

(95 V, 15 A)

5.9

(90 V, 15 A)

2,000 14.3 9.8 3.3

(85 V, 26 A) 

3.6

(85 V, 24 A) – – –

3,000 21.4 6.5

 (7.5 @ 150 V)

3.5

(104 V, 30 A) – 3.9(110 V, 28 A)

7.4*

(150 V, 20 A)

7.3

(150 V, 21 A)

4,000 28.6 4.9 – – 4.0

(130 V, 33 A) – –

6,000 42.9 3.3 – – – – –

*Note that the measured Hf resistance did not stabilize at power—the value provided is estimated from the test runs completed.

At higher power levels, the small diameter wires failed, possibly due to locally elevated temperatures. Elevated temperatures could occur at localized positions due to impurities on the wire 
20

that cause elevated power deposition at that point. Impurities could cause fracture more rapidly 

in small versus large diameter wires. If this is the case, adoption of more rigorous cleaning procedures could remove additional surface impurities and may allow use of smaller diameter wires in 

heater construction.

To reduce the voltage required to provide a set power level, a lower resistance element would 

be required. For instance, to reach a power level of 3 kW, a 5-W element would require 122 V (24 A). 

A 3-W element would require only 95 V (32 A). A lower resistance element is also required to increase pin power for a fixed voltage (P = V2/R). For instance, the relatively high resistance of the 

1-mm (0.040-in) Hf and the five-wire W braid limited the achievable power per pin to ≈3 kW with 

the 150-V power supply. The element resistance can be reduced by selecting a material with a lower 

resistivity (ρR) or by increasing the diameter of the wire used to construct the element. The resistance 

of a wire is given in equation (2):

 

, (2)

where A is the cross-sectional area of the wire and l is the wire length. Table 4 provides the calculated resistance (at 20 oC) of various refractory metals using the resistivity values provided by 

the wire manufacturer, ESPI Corp., Inc. Calculations assume a wire length of 3.6 m (2 ft), the 

approximate length required to wrap the 1.04-cm (0.410-in) mandrel in a double helix fashion. For 

refractory metals, the resistance increases with temperature. (Graphite is unique in the fact that its 

resistance first decreases with temperature before reaching a minimum and then increasing again.) 

Detailed expressions for the resistivity of some of these materials as a function of temperature are 

included in appendix D.

Table 4. Summary of resistance calculated for select refractory metal wires at 20 °C.

Material

Resistivity* @ 20 oC

ρR , (μΩ - cm) Number of Wires

Cold Resistance for 3.6 m Length (Ω)

OD 0.25 mm (0.010 in) OD 0.5 mm (0.020 in) OD 1 mm (0.040 in)

Re 19.3 1 13.94 3.49 0.87

Ta 13.5 1 9.75 2.44 0.61

Nb 12.5 1 9.03 2.26 0.56

W 5.5 5 (braid) 0.79 0.20 0.05

Hf 35.1 1 25.35 6.34 1.58

Mo 5.2 (@ 0 oC) 1 3.76 0.94 0.23

*Resistivity data from manufacturer.9

To balance the tradeoff between resistivity and ductility, 1-mm- (0.040-in-) diameter wire 

was selected as a baseline for initial refractory metal wire heater element tests. Selection of the same 

diameter wire for all material options reduces the number of variables between materials, offering 

the ability to better assess the benefits of each material. Additional wire diameters could be tested 

as the development effort for the heater elements is focused to a defined set of requirements. Results 

from testing Ta and W are discussed in sections 2.4 and 2.5.
21

2.1.1.2 Heater Assembly: Electrical Integration. The electrical connections to the heater 

assembly incorporate the power leads and the junction between the power leads and the heater 

itself, which could be made via direct mechanical connection (used for preliminary testing) 

or by use of a low resistance coupler. The coupler serves three purposes:

(1) It acts as the interface to the hot heater wire.

(2) It reduces the resistance between the power supply and the heater element 

(as a function of its material properties).

(3) It lowers the temperature seen by the Cu lead wires (as a function of the 

length of the coupler pieces).

Heater element design is fundamentally limited by the capability to provide electrical 

hookup to the element. This limitation is established by the physical wire size relative to the element size and the amount of power that can be pushed into the element. To push a large current 

into the element, while maintaining minimum resistance in the power leads, the feed wires become 

too large relative to the element footprint. As the assembly size is reduced, it also becomes increasingly difficult to seal the element (around the diameter and around the small lead wires), a requirement for the sheathed heater designs.

For the graphite rod heater elements, Mo couplers were used to connect the Cu lead wires 

from the power feedthrough to the graphite element. The threads of the coupler are covered with 

graphite cement and then screwed directly into the graphite, securing them into place. Initial 

tests of the refractory wire elements have not used a coupler; instead, the refractory wire elements 

were fed directly from the element to the power feedthrough without using any intermediate, lowresistance wire between the element and the feedthrough. However, it is desirable for all the input 

power to go into the element, requiring lead wires to have very low resistance to minimize the 

power deposition in these components. To accomplish this, a mechanical connection is required 

between the refractory wire and the power lead wires to ensure a good electrical connection. The 

engineering drawings for the redesigned couplers for the 1.6- and 1.02-cm (0.625- and 0.400-in) 

assemblies (corresponding to the 1.04- and 0.60-cm (0.410- and 0.235-in) mandrels, respectively) 

are provided in appendices B and C, respectively. The maximum coupler size is limited by the ID 

of the sheath. The 1.6-cm (0.625-in) design uses a bolted, mechanical hookup, but, the hardware 

used to provide this connection was too large for the 1.02-cm (0.400-in) assembly. As a result, the 

coupler design is modified for the smaller assembly design simply to account for the reduced diameter, and the final design uses a sleeve/compression hookup for the smaller assembly. In both cases 

the coupler is constructed from Mo.

To prevent contamination of the test article by the heater element and to seal an inert gas 

(He) inside the sheath to improve the thermal conductivity across the element, a high integrity 

seal must be developed. The sheath is manufactured as a tube and two end caps are used to provide a closed-end unit. The upper end cap (no power feedthrough) can be welded to the body of 

the sheath before the heater element is inserted. On the opposite end, a metal or ceramic end cap 

that is designed to allow the coupler to feed through is used. This general configuration is shown 
22

schematically in figure 10. The mandrel is inserted into the sheath such that the heater element is 

centered axially, allowing room for the coupler assembly at the power feedthrough end cap. If a 

metal end cap is used, insulating material (electrically nonconductive) must be included around 

the coupler feedthrough surface. The metal cap could then be welded to the sheath around the 

OD, but brazing would still be required at the coupler/cap interface due to the presence of the 

insulating material. Alternately, if a ceramic cap is used, no additional insulating material is 

required around the coupler feedthrough, but it must be sealed by brazing around the OD and at 

the coupler feedthrough. Two concerns emerged with regard to this braze as the design proceeded: 

the desired maximum temperature of the braze joint exceeds that of current braze technology 

and brazing in a He (or other inert gas) environment may not be possible. Possible techniques 

for sealing the heater elements inside a sheath and the associated thermal analysis will be further 

discussed in section 2.3.

Power Leads

End Cap

(Brazed)

He Gap

Mo Coupler (Generic Design Shown) Sheath

Spiral-Wound Heater Element

Over Al2O3 Mandrel

Figure 10. Generalized assembly configuration (not shown to scale).

2.1.2 Hardware Testing

Hardware testing to date has considered Ta, Re, Hf, and W wire heater elements. In all 

cases, the wire was wrapped around the 1.04-cm (0.410-in) Al2O3 mandrel at a constant pitch (producing a flat power profile) and the wire used to form the heater element was mechanically connected directly to the electrical feedthrough. This connection is nonideal and nonprototypic, since 

all the input power from the power supply should go into the spiral-wound element itself rather 

than into the lead wires. However, for material and concept acceptance testing, there is no need 

to build and connect additional couplers. Elements selected for additional tests will incorporate 

the Mo coupler discussed above. After all potential element materials have been tested at 1,200 W 

(minimum test requirement of 1,200 W for 100 hr), the most desirable materials will undergo further testing at higher power levels.

2.1.2.1 Tantalum. Initial tests conducted on Ta wire used a 1-mm (0.040-in) wire wrapped 

around the 1.04-cm (0.410-in) Al2O3 mandrel. Tantalum was tested over a range of 1,200 to 4,000 W 

for a total test time of ≈150 hr. Two Ta heater elements have been tested. The first Ta element was 

tested at 1,200 W for 101 hr, 1,600 W for 17.5 hr, and 2,000 W for 15.5 hr. The second was tested 

at 3,000 W for 50 hr and at 4,000 W for 5 hr. Bakeout of each heater was accomplished by applying ≈50 to 100 W to the element for several hours. Testing began on the following day, so that the 

startup to the desired power level would be from a cold element.
23

For the first Ta heater element run at 1,200 W, resistance measurements made before and 

after a 100-hr run indicated a < 9% increase in the measured cold resistance of the element (≈0.8 

to 0.87 W). After an additional 34 hr at 1,600 to 3,000 W, the measured element resistance (after 

cooling for 12+ hr) had not changed, indicating that Ta was stable before and after operation at 

high power levels. Current and voltage were stable during all runs at power. ‘Stable’ heater operation was defined as an element that maintained approximately constant resistance before and after 

test, indicating that operation at power did not introduce any fundamental changes to the material 

properties. To be defined as stable, the element should also operate at a constant, stable current 

throughout testing at a constant power. The cold Ta element had a measured resistance of ≈0.8 W, 

and at 3 kW, the resistance increased to ≈3 W. From a cold startup condition, the heater was able 

to reach power in <30 s. 

The first Ta heater element was tested both inside and outside of a stainless steel sheath. 

The sheath was unsealed, leaving a vacuum gap between the heater element and the sheath. During the 2-kW run, the heater was removed from the sheath for the remainder of testing. At this 

power level, the element produced a consistent glow from end to end, suggesting uniform performance and temperature along the length of the heater element. The element was run for an 

additional 8 hr at 3 kW. During the 3-kW test, photos of the hot Ta element revealed differential 

heating at localized positions along the helical conductor wire. Figure 11(a) shows the heater element early in testing where there are two points at which the Ta winding does not glow as brightly 

as the adjacent winding. The photo in figure 11(b), taken later in the 3-kW testing period, indicates 

that the differential heating locations have shifted to alternate positions and the overall coloration 

of the image has red shifted. Because some of the element testing was performed inside the stainless steel sheath, there was some concern that the sheath may have contaminated the Ta wire, 

producing the observed differential heating. Posttest materials analysis revealed that a lower temperature was most likely produced due to variations in the emissivity of the element over time.

(a) (b)

Differential Heating

Differential Heating

Figure 11. Test of W heater element at 3 kW (a) early in test and (b) later in test, revealing 

a spectral shift.
24

Test of a second Ta heater, constructed from the same wire lot, was performed to reproduce 

the differential heating observed during testing of the first element. This heater element was run 

bare (no sheath) at 3,000 W for a total of 50 hr. Visual observations indicated that the second Ta 

heater element demonstrated differential heating at high power operation, as was observed during 

testing of the first element. Hence, this behavior was not a result of contamination from the stainless steel sheath in the previous element test. Figure 11 (a) and (b) shows that the location of the 

differential heating on the Ta element moved to new positions at different times during the tests, 

suggesting that it was not a result of fixed surface contamination. Instead, this behavior may be 

due to changes in the surface emissivity as a function of time, which could be caused by oxidation 

or local contamination of the surface, producing an effective local reduction in the element emissivity. With time, the coil could be burning off the contaminant or it could be absorbing emitted 

gases such as O2 (based on the solubility of the material for these gases), resulting in a variable 

surface condition. 

A complete visual inspection, after opening the chamber, revealed that the wire appeared 

to still be ductile. Power was increased to 4,000 W for a duration of 5 hr. Additional testing was 

not possible because the heater was accidentally knocked off the support stand while performing 

activities outside the chamber between tests.

After testing was completed, the Ta heater wire maintained a shiny silver metallic color 

while the white Al2O3 mandrel had a bluish-black discoloration. The first Ta element was submitted for complete materials analysis to assess the overall heater performance. The complete analysis 

report, performed by G. Jerman of MSFC, is included in appendix E. When the wire was removed 

from the mandrel, it began to fracture in a brittle fashion and after approximately one-third of the 

wire had been unwound, the insulator fractured near the base of the mandrel where the Ta enters 

and exits the assembly. After the wire was completely removed from the mandrel, more significant 

discoloration of the Al2O3, which appeared to penetrate into the Al2O3, was noted (fig. 12). The 

complete materials analysis indicated that the discoloration was likely caused by the formation 

of an intermetallic aluminum-tantalum (Al-Ta) coating on the insulator surface. Carbon bearing 

contamination in the heater assembly reduced some of the Al2O3 insulator surfaces under vacuum 

to form metallic Al. The hot Al metal then reacted with the Ta and reduced vaporized tantalum 

oxide to form a conductive intermetallic coating on the Al2O3 insulator. Vapor deposited tantalum oxide on the sight glass window resulted in the overall red shifted image shown in figure 11(b), 

which was taken later in the test.
25

(a) (b) 2 mm

 Figure 12. Alumina mandrel after removal of W wire, revealing (a) significant surface 

discoloration and (b) close-up view of contamination.

Detailed analysis using scanning electron microscopy indicated significant surface contamination on the Ta wire including C, O2, sodium (Na), and chlorine (Cl). The presence of these 

elements suggests that contamination may be due to human handling. This contamination could 

be reduced by thoroughly cleaning the wire to remove any oils or other volatile C bearing contamination before assembly. After assembly, bakeout under vacuum would remove any remaining 

volatile components. Additional cleaning and preparation procedures that may be adopted will be 

discussed in section 3.6. 

The Al2O3 used to construct the mandrel begins to melt at 2,072 °C and the maximum recommended operating temperature of Al2O3 is 1,800 °C. Analysis indicated that localized melting 

may have occurred in regions in which the Ta wire was in intimate contact with the Al2O3 mandrel. Constructing the mandrel from a ceramic with a higher temperature capability such as boron 

nitride (BN) (powder consolidated) or pyrolytic BN (vapor deposited), which does not break down 

until 3,000 °C, may be desirable. However, discussions with manufacturers of solid BN components 

have indicated that the maximum length component that could be fabricated is ≈46 cm (18 in), 

including the length of the ends necessary to hold the component during the fabrication process). 

As a result, elements using solid BN components in their construction may be significantly less flexible in their application than those using Al2O3. Additional higher power tests of braided Ta wire 

in vacuum and He environments, considering both sheathed and unsheathed configurations, are 

discussed in section 2.4.
26

2.1.2.2 Rhenium. Tests conducted on Re wire used a 1-mm (0.040-in) wire wrapped around 

the 1.04-cm (0.410-in) Al2O3 mandrel. The Re heater element demonstrated stable performance at 

1,200 W for the required 100 hr of run time. Operation for >100 hr indicated a reaction between the 

Re wire and the Al2O3 mandrel, but because the heater met the minimum run time requirements 

without failure, the Re heater element is still being considered for further testing. The possibility 

of coating the mandrel with BN, or of constructing the mandrel from BN, is being investigated to 

mitigate the evident materials incompatibility issue between the Re and the Al2O3. Sealing the heater 

in a refractory metal sheath to prevent contamination is also being pursued and will be discussed 

further in section 2.3. Sealing techniques for the sheathed heater design are currently being investigated with industry. 

As noted for the Ta heater, the cold Re element was capable of reaching a power level of 

1,200 W (67 V, 19 A) in <30 s. The measured cold resistance dropped from 0.92 Ω at the start of 

day one to 0.91 Ω at the start of day two. No change in resistance was noted between the second 

and third testing days. The Re heater element was tested for a total of 100 hr at 1,200 W. Current 

and voltage remained stable throughout the element test at a given power level. After meeting the 

basic requirements for an acceptable heater element, power was increased to ensure that the element was not ‘on the edge’ of its performance capabilities. Testing continued for 24 hr at 3,000 W, 

9 hr at 4,000 W, and 8 hr at 5,000 W. Unlike in the previous Ta wire tests, no differential heating 

was observed in the Re element. Photographs taken during the Re element test at 5,000 and 1,200 

W are provided in figure 13. Rhenium has much lower O2 solubility at temperature than does Ta, 

so that the same O2 absorption would not be expected as was postulated from the Ta test results. 

After testing, the Re heater wire maintained a shiny silver metallic color while the white Al2O3

mandrel had a bluish-grey discoloration. 

(a) (b)

Figure 13. Test of Re heater element at (a) 5 kW and (b) 1.2 kW.
27

After test completion, the Re heater element was submitted for complete materials analysis 

to assess the overall heater performance. The complete analysis report, performed by G. Jerman 

of MSFC, is included in appendix F. When the wire was removed from the mandrel it was ductile 

and no wire was broken. After the wire was completely removed from the mandrel, variations in the 

discoloration of the Al2O3 mandrel were noted. Localized surface coloration changes varied from 

bluish-grey to white streaks and dark bluish-grey spots, as shown in figure 14(b). Analysis indicated 

that this discoloration was caused by the deposition of Re oxides. The deposition of these low temperature oxides protected the Al2O3 mandrel from high temperature degradation during the thermal testing. Because the Re oxides preserved the integrity of the mandrel and the conductor wire 

retained a large amount of ductility, the material analysis determined that the Re/Al2O3 combination was superior to the Ta/Al2O3 combination. Impurities contained in the heater element resulted 

in the formation of the Re oxide that actually acted as a protective layer to reduce further damage 

of the wire and mandrel, but because the protective Re oxidation begins to degrade at 1,000 °C, 

operation of an uncleaned Re element above this temperature is not recommended. Although Re 

performed very well in the thermal simulator tests, it was found that each order of Re wire from 

Rhenium Alloys, Inc. varied significantly in its purity levels. As a result, it may be more desirable 

to fabricate heaters with Ta or braided W wire (see sec. 2.1.2.4) due to the reproducibility of the 

specifications for these wires.

(a) (b) 5 mm 5 mm

Figure 14. Posttest Al2O3 mandrel (a) before and (b) after removal of Re wire, revealing 

slight surface discoloration.

2.1.2.3 Hafnium. Tests conducted on Hf wire used a 1-mm (0.040-in) wire wrapped around 

the 1.04- cm (0.410-in) Al2O3 mandrel. Although the Hf met the minimum test requirement of 

1,200 W for 100 hr, the element resistance never stabilized over the run time. Prior to test, the 

element resistance (cold) was ≈1.5 Ω. At the end of each day of testing, the measured element 

resistance had increased, and upon completion of the 100 hr of testing at 1,200 W, the cold resistance had increased to ≈3.5 Ω. Additional testing was performed at 3 kW for 10 hr. Although the 

resistance was relatively stable at 1,200 W, operation at higher power caused the resistance to drop 

slightly throughout the test. 
28

Although no materials analysis was conducted, the continually increasing cold resistance 

could indicate that the Hf element suffered oxidation during test. Alternately, material incompatibilities or impurities in the test components could have altered the microstructure of the wire, 

causing it to become porous and slowly reducing its effective electrical conductivity path. After 

testing, the Hf wire was very brittle.

Because Hf did not demonstrate stable operation, it is not recommended for further testing at 

the current requirement levels. Although stability at power is a necessary requirement for any heater 

element, Hf has not been entirely eliminated from the list of potential heater materials. Depending 

on the final series/parallel configuration of the power hookup, Hf could, in fact, be a desirable choice 

for a reactor simulator that requires lower power density per pin due to its high resistivity. 

2.1.2.4 Tungsten. Construction of W wire-wrapped heater elements has considered use 

of a single 1-mm (0.040-in) wire or braided wire, constructed from five 0.25-mm (0.010-in) wires 

braided together. Initial investigation indicated that the ductility of the 1-mm (0.040-in) W wire 

was such that the wire was very difficult to wrap around the 1.04-cm (0.410-in) mandrel, so no 

heater element was constructed using a single W wire. Use of smaller diameter wires in a braided 

configuration mitigated the manufacturability problem that was encountered with the single wire. 

The number and diameter of wires in the braid can be adjusted to vary the resistivity and currentcarrying capacity of the heater element, per the requirements of a given reactor design. The current work has considered 0.25-, 0.38-, and 0.5-mm (0.010-, 0.015- and 0.020-in) wires because larger 

wire diameters encounter the ductility problem. Heating the W wire while it is being wrapped 

could allow use of larger diameter wires, but this would introduce additional complexity to the 

heater fabrication. 

The braided W wire tests (five 0.25-mm (0.010-in) wires) have accumulated ≈240 hr of test 

time over a total of 30 thermal cycles. Pretest pictures of the braided W wire element are provided 

in figure 15. Tests of the W braid have assessed performance of both the unsheathed (bare) and the 

sheathed heater element. Unsheathed tests included 101 hr of testing at 1,200 W over 13 thermal 

cycles and an additional 24 hr at 3,000 W over 3 thermal cycles. The sheathed tests used a 1.6-cm- 

(0.625-in-) diameter Mo sheath (99.9% pure). The element was slid into the sheath but the sheath 

was not sealed, such that the environment between the heater and sheath and external to the 

sheath were at vacuum (≈1.3 mPa (≈10–5 torr)). At 1,200 W, the measured resistance of the sheathed 

element was ≈6 Ω. 
29

(a)

(c)

(b)

Figure 15. Tungsten wire-braid heater element wrapped on a 1.04-cm (0.410-in) Al2O3

mandrel. Close-up views are provided for (a) the power input end, (b) the nonpower end, showing wire wrap return for double helix configuration, and 

(c) the center section showing the five-wire braid.

For element testing inside the Mo sheath, the sheath was instrumented from end to end 

with type C TCs. A piece of nickel (Ni) foil was first welded to the sheath surface, and the TC 

was then welded to the foil (all welding was performed in air). Measured temperatures indicated 

a range from 950 to 1,400 °C at 1,200 W (variation in temperature along length of heater). The 

exact location of each TC was not measured, but approximate locations are identified in figure 16. 

Because they are bonded to the Mo sheath via Ni foil, the TCs are slightly cooler than the sheath. 

In addition, the TCs act as small fins, locally cooling the Mo and producing a temperature measurement that is somewhat lower than the actual sheath temperature.

Power Leads

Flange

He Gap

1 2 3 4 Mo Sheath

W Wire Braid Wrapped Around

Al2O3 Mandrel

Figure 16. Approximate TC locations on the Mo sheath.
30

During the test, it became apparent that TC 2 may have lost contact with the surface of the 

sheath, as its temperature dropped to ≈630 K (relative to ≈1,200 K in previous test days). Early test 

data (prior to complete loss of TC 2) are provided in table 5. Two measurements are provided for 

each test day: one taken at the beginning of testing, and one taken at the end of the test day prior 

to turning off the power supply. (All temperatures are reported in kelvin.) The significant variability in the TC measurements along the length of the sheath may be related to their location on the 

sheath, but it is more likely (especially in the case of the relatively closely positioned TCs 2 and 3) 

that TCs attached to the outer surface of the sheath may not have provided reliable temperature 

measurement for thermal simulator evaluation. The temperature variation noted in the test data 

from day to day, particularly for TCs 3 and 4, may be due to reduction of the sheath surface emissivity (thereby elevating the temperature necessary to reject the heater power), off-gassing impurities from the sheath exterior surface, or off-gassing impurities from the foil connecting the sheath 

to the TC (thereby limiting cooling of the foil due to radiation). Alternate temperature measurement techniques will be discussed in section 3.1. 

 Table 5. Measured sheath temperatures (K) for the braided W wire heater assembly. 

(The first temperature entry corresponds to the beginning of the test, the 

second to the end of the test).

Power Level 

 (test date)

TC Location 

1 2 3 4

1,350* W

(Feb. 8, 2005)

1,272 1,217 1,598 1,371

1,267 1,217 1,600 1,402

1,350* W

(Feb. 9, 2005)

1,260 1,217 1,593 1,413

1,262 1,208 1,672 1,430

1,350* W

(Feb. 10, 2005)

1,256 1,222 1,708 1,427

1,262 1,207 1,760 1,439

1,350* W

(Feb. 11, 2005)

1,259 1,219 1,785 1,455

1,264 1,202 1,776 1,483

*A total power of 1,350 W corresponds to ≈1,200 W in the heater element 

(assuming losses in the lead wires).

A thermal analysis was initiated after the testing was performed to assess the measured 

temperatures of the W wire braid element and the Mo sheath in the as-tested configuration and to 

determine the impact of varying environmental boundary conditions on the element and sheath 

temperatures. Calculations consider environment conditions both inside and outside the sheath, 

assuming either a vacuum test environment (as in the current experiment setup) or a 9 kPa (70 torr) 

He environment. In a vacuum environment, the only means of heat removal from the element is by 

thermal radiation to the cool chamber walls. If the chamber is backfilled with ultra-high purity gas, 

thermal conduction through the gas reduces the element temperature. 

Analyses were performed using both Mathcad®, a Parametric Technology Corporation product, and Systems Improved Numerical Differencing Analyzer (SINDA)/Fluid Integrator (FLUINT). 

Use of both tools provided a quick check in formulations employed in the Mathcad model. The modeling approaches employed in each tool were similar. The key governing equations and supporting 
31

formulations applied in the thermal analyses are included in appendix G. The simplified cross-section 

geometry shown in figure 17 was the basis of the thermal model, and also, there was an attempt made 

to model the braid/mandrel geometry shown in figure 15(c) (including details of the braid). Additional 

thermal analyses on sealed heater assembly configurations are included in section 2.3.

Mo Sheath

Heater Element

(Spiral-Wrapped Mandrel)

Test Chamber Wall (≈323 K)

Vacuum or GHe

Vacuum or GHe

Figure 17. Approximate two-dimensional model applied for W wire element thermal analysis 

(not to scale).

All calculations assumed that the heater element wire, mandrel, and sheath were each isothermal and that the test chamber was maintained at a constant 323 K. The braid was modeled as 

a cylinder with the same cross-sectional area as the braided wire. The braid was thermally coupled 

to the mandrel and sheath via radiation and gas conduction depending on the case being evaluated. 

The emissivities of braid, mandrel, and sheath were assumed to be temperature independent and 

equal to 0.2, 0.4, and 0.2, respectively. For the radiative heat transfer calculations, the appropriate 

two-dimensional view factors and radiation exchange factors were hand calculated, based on figure 

18(a). For the cases that incorporated gas (He) in the heater element/sheath gap, the gas conductance was assumed constant and equal to 0.427 W/m-K (corresponding to He at 1,273 K). A finite 

element model (FEM) mesh, shown in figure 18(b), was used to determine the two-dimensional conduction shape factors for the W wire-wrapped mandrel. The conduction shape factors and radiative 

exchange factors were then used in the Mathcad and SINDA models to compute the temperatures 

corresponding to the various configurations.
32

(a) (b)

Figure 18. Braid/mandrel/sheath geometry: (a) Annotated layout and (b) finite element model 

of determination of conduction shape factors.

The results of the preliminary thermal analysis at a heater power of 1,200 W are provided in 

table 6. Case 2 (sheathed element, vacuum conditions) corresponds to the test conditions that were 

presented in table 5. A direct comparison of the two test results reveals a mixed picture. Experimental results indicated a sheath temperature ranging from ≈1,256 to 1,785 K along the length of 

the sheath and over several test days. The average measured sheath temperature for each test day 

presented in table 5 ranged from 1,364 to 1,431 K. The predicted sheath temperature, using both the 

SINDA and Mathcad models, was 1,486 K for this test configuration. The lower average measured 

temperature could result from the fact that the TCs are not bonded directly to the Mo sheath, but 

are separated from the sheath by a piece of Ni foil, causing them to be at a slightly lower temperature than the sheath. In addition, the differences in results could be due to differences between the 

emissivity value used in the model and the actual emissivity of the sheath external surface. For 

example, if the emissivity of the sheath were 0.1 (instead of 0.2 used in the model), the sheath temperature estimate would increase to over 1,700 K. 
33

Table 6. Calculated mandrel, wire, and sheath temperatures for the braided W wire 

thermal simulator.

Case Configuration

Calculated Temperature (K)

W Wire Al2O3 Mandrel Mo Sheath

1 Sheath w/He 1,623 1,508 1,368

2 Sheath w/vacuum 2,405 1,863 1,486

3 No sheath w/He 1,271 1,161 N/A

4 No sheath w/vacuum 2,178 1,026 N/A

Thermal analysis results indicate the significant benefit of adding gas to the heater assemblies. The analysis predicts that the addition of thermal conduction across the heater/sheath gap 

will reduce the wire temperature by 782 K (from 2,405 to 1,623 K) in the sheathed heater element 

configuration and by 907 K (from 2,178 to 1,271 K) in the bare heater (no sheath) configuration. 

For both vacuum and He-filled environmental conditions, the sheathed configuration results in 

significantly higher wire and mandrel temperatures than for the equivalent unsheathed geometry. 

This results from the fact that the sheath acts as a radiation shield between the heater element and 

the cool chamber walls, reducing radiation heat transfer from the wire and mandrel. This effect 

is reduced when conduction is present, but the temperature differential between the sheathed and 

unsheathed configurations is still significant. Thermal analyses that consider additional details of 

the heater element and sheath geometry will provide additional insight to the expected temperatures in the heater assemblies. These calculated values could be further verified by incorporating 

enhanced diagnostics on the heater assemblies and on the test chamber itself. Tests of W wirewrapped elements at higher power levels (up to 6 kW) were performed on three-, four-, and fivewire braids using the 0.25-, 0.38-, and 0.5-mm (0.010-, 0.015-, and 0.020-in) wire diameters. Results 

of these tests and the resulting thermal management issues are discussed in section 2.5.

2.1.3 Preliminary Conclusions

Testing of the refractory wire-wrapped heater element designs as of March 2005 indicated 

that the 1.6-cm (0.625-in) assembly with the 1.04-cm (0.410-in) Al2O3 mandrel worked well for 

tests at 1,200 to 5,000 W. This series of testing provided information on desirable materials for the 

heater element, buildup of a complete heater assembly (as it relates to a desirable minimum assembly diameter), and power hookup. Materials recommended for continued development include Re, 

Ta, and W braid. All have successfully run for over 100 hr at 1,200 W. Brief runs of up to 4,000 W 

for Ta, 5,000 W for Re, and 3,000 W for W braid also produced desirable performance. No failures 

have been noted in the baseline tests (1-mm (0.040-in) wire or wire braid) for the single-ended, 

double-helix element designs tested to date. The success of the braided W wire element tests may 

open up the possibility of using other low ductility refractory wires, such as Mo. The 1.02-cm 

(0.40-in) assembly components are currently being procured and will be tested with Re, Ta, and 

W wires. These tests are covered in sections 2.4 through 2.6. If all wires are acceptable for the 

thermal simulator application, cost and ease of fabrication should also be considered in making 

the final material selection before fabricating a large number of heater assemblies for the reactor 

test article. Table 7 provides current estimates for the cost of each of the three refractory wires 

that will undergo further testing.
34

Table 7. Estimated costs* for additional heater element wire.

Wire Diameter W (3N8 Purity) ($) Ta (3N8 Purity) ($) Re (3N Purity) ($)

0.25 mm (0.010 in) 2.00/m 7.40/m 107.00/m

0.38 mm (0.015 in) 3.30/m 7.90/m 70.50/m

0.5 mm (0.020 in) 4.00/m 8.00/m 43.50/m

1.0 mm (0.040 in) — 23.50/m 356.50/m

Comments Smaller W wire 

diameters necessary 

due to low wire ductility. The braided wire 

construction will cost an 

additional $250/braid 

(≈4-m length).

*Quotes obtained from manufacturer.9

Future testing will seek to overcome challenges introduced by further reducing the assembly diameter. Previous testing (in 2003) on smaller element designs indicated significant difficulty 

in further reducing the complete assembly size. Because of the complexity associated with fabricating very small diameter, detailed mandrels and sheaths, it was difficult to find a vendor to 

produce the pieces. Mandrels having an OD of 0.53 cm (0.210 in) were manufactured from Al2O3

with a constantly varying pitch. These mandrels were designed to fit inside a 0.65-cm (0.255-in) 

OD sheath, producing a full heater assembly with a total diameter of 0.65 cm (0.255 in). However, 

tests conducted using 0.5-mm (0.020-in) Re and W wires were not successful. In fabricating the 

element, it was very difficult to hold sufficient tension on the wire so that it did not unwrap from 

the mandrel while the sheath was pulled over the element. Tests of these elements were conducted 

in vacuum, so heat removal from the element was purely by thermal radiation. Tests at 1,200 W 

resulted in wire failure, presumably due to the inability to effectively remove heat from the element. Incorporation of a trapped high-purity gas in the heater element/sheath gap would improve 

the heat transfer by introducing conduction across this gap. Performing the test with the vacuum 

chamber backfilled to low pressure with ultra-high purity (UHP) inert gas (He or Ar) would 

also assist in removing heat from the element. These modifications to the test parameters could 

indicate significantly different results in the potential use of these small diameter heater assemblies. However, development to date has not addressed the issue of power integration with these 

very small diameter heater elements. As discussed previously, the coupler and power integration 

should fit within the diameter of the assembly so that it will not interfere with the neighboring 

heater assembly when integrated into a test article. Very small element diameters introduce a new 

challenge in providing input power to achieve the desired power per pin (to push higher currents 

necessary to reach the desired power levels, the power lead wires become relatively large). This 

challenge in power integration would have to be overcome before a reduction in assembly diameter 

could be considered feasible.
35

The element diameter could be further reduced by removing the requirement for sealing the 

heater element inside a sheath. In this case, the minimum element diameter would be driven by the 

limitations for fabricating the mandrel and the power integration issues discussed above. Although 

the heater would no longer be physically isolated from the test article (electrical isolation would still 

be required to prevent shorting), the achievable pin size could be beneficial. In the current mandrel 

design, at either end of the mandrel section that is grooved for the wire, the mandrel has a slightly 

larger diameter to allow the element to be self-isolating (e.g., unless it relaxes over time, the wire 

physically cannot touch the sheath due to the Al2O3 mandrel ‘standoffs’). For the 1.02-cm (0.40-in) 

assembly, the mandrel diameter at the ends of the grooved section is 0.65 cm (0.255 in). To maintain a 0.13-mm- (0.005-in-) radial clearance, this element could fit into a 0.6-cm (0.270-in) fuel-clad 

surface (allowing a 0.13-mm (0.005-in) tolerance on the mandrel diameter and 0.13 mm (0.005 in) for 

mandrel and fuel-clad straightness). Alternately, the delta between the mandrel ends and the OD of 

the grooved portion could be reduced to bring the sheath closer to the wire element, reducing the 

width of the gap and, consequently, the OD of the sheath. 

The noted interactions between the Al2O3 mandrel and some refractory metal materials 

at elevated temperature (see materials analyses, app. E and F) do not eliminate these materials 

from use in heater construction. With over 100 hr of run time, these elements have performed well 

at the current requirements (1,200 W). Future endeavors will investigate reducing the interaction 

between materials by employing a BN coating over the Al2O3 or replacing the Al2O3 with solid 

BN. Using a BN mandrel, however, introduces additional issues to the heater fabrication. Boron 

nitride has significantly less strength than the Al2O3 currently in use, so at small diameters, it will 

have a tendency to fracture if any significant tension is required to wrap the heater element wire. 

Additionally, as indicated in the above discussions, solid BN is limited to a maximum fabrication 

length of ≈46 cm (18 in). As a result, elements using solid BN components in their construction 

may be significantly less flexible in their application than those using Al2O3 mandrels. Developing a BN coating that could be applied to the Al2O3 and wire using a vapor deposition process for 

electrical insulation and improvement of materials compatibility may be a more promising option 

than employing solid BN.

Investigation of methods for heater encapsulation to prevent contamination to the test 

article will be discussed in section 2.3. 

2.2 Carbon Fiber Braid Simulator Design

Graphite C fiber braid is being investigated for use in thermal simulators in an effort to 

reach small pin sizes (0.65 cm (0.255 in)) and high power levels (1,200 W to 6,000 W per simulator) at high temperatures. One fabrication technique involves taking a C fiber braid, wrapping it 

around a mandrel, and melting the mandrel away while simultaneously carbonizing the fiber. This 

would require the use of a mandrel constructed of a material, such as Al, that could be melted 

near the temperature required for carbonization. Selection of the mandrel material must also take 

into account compatibility with refractory metal sheaths should any traces of the material remain 
36

after the carbonization process. Alternately, the same Al2O3 mandrel design discussed for the 

refractory wire heater elements can be applied for the C fiber braid elements. The challenges 

facing the C fiber braid simulator design include:

• Finding an appropriate method of attaching the C fiber braid which has the approximate 

consistency of yarn, to the Al2O3 mandrel.

• Determining the best carbonizing method (furnace temperature and treatment time).

• Creating a support structure for the mandrel that holds the braid to the mandrel, but does not 

interfere with the mandrel melting process (using an Al mandrel).

• Designing the electrical connection to the integrated system.

• Achieving the desired power levels given the high resistance of the carbon fiber braid. 

Tests run to date using the C fiber braid material have been directed at verifying material 

capabilities at temperature and initial results look promising for the power levels set forth in the 

requirements. Research and development has just begun on this concept, with the initial braid tests 

showing great promise—the primary issues with this concept are related to manufacturability. 

2.2.1 Engineering Design

Preliminary tests were conducted using 3,000 fiber, 0.25-cm- (0.10-in-) diameter C braid 

material acquired from A&P Technology (type RM3793). The tension required to wrap a C fiber 

braid material around a mandrel is much less than that required for a refractory wire, simplifying 

heater fabrication over the design discussed in the previous section. Rough ‘checkout’ tests involve 

wrapping a section of braid material (≈0.6 m (≈2 ft)) around a 2.5-cm (1.0-in) Al2O3 tube and connecting either end to the power supply, creating a double-ended heater design (see fig. 19). Additional tests have been performed using the 1.02-cm- (0.410-in-) OD mandrel (where the complete 

heater assembly would utilize the 1.6-cm (0.625-in) sheath design) that was employed in the refractory metal wire element tests (see fig. 20). 

 Figure 19. Double-ended C fiber braid trail: configuration for rough 

checkout tests.
37

 Figure 20. Carbonized C fiber braid wrapped in a double helix fashion around Al2O3

mandrel.

Power is provided to the C fiber braid assembly via a Mo coupler. The ends of the braid are 

fed through the holes at the end of the Al2O3 mandrel, the coupler is coated with graphite cement, 

and the coupler is then pushed into the holes in the Al2O3 mandrel to bond the coupler to the 

braid material. Copper lead wires are then connected via the Mo coupler. The electrical connection method has not yet been worked out for the carbonized version in which the Al mandrel is 

melted away, leaving only a stiff C coil. 

An acceptable method of carbonizing the braid material, while simultaneously melting 

away the mandrel material (Al), must be developed if this approach is selected. This method will 

require a test fixture to support the mandrel during the carbonization process. The initial attempt 

to produce a mandrel-free carbonized element used a very simple support structure in which the 

element was rested in a V-shaped ‘trough’ such that one end of the element was higher than the 

other. The C fiber braid was dipped in resin (necessary to carbonize the material) and wrapped 

around the Al mandrel. The entire assembly, including the trough support, was then placed in a 

vacuum furnace to melt the Al mandrel. The resin-dipped material carbonizes at 800 °C or higher, 

and the Al melts at 660 °C, and once melted, just a stiff, carbonized braid should remain. In the 

initial test of this carbonization method, the braid was successfully carbonized and the Al melted 

and successfully fell away from the braid material, but a melted pool of Al remained at the lower 

end of the element, leaving ≈90% of the C composite spiral clear with the remaining 10% in the 

pool. Should the C fiber braid material be pursued, alternate, more sophisticated support structure designs will be investigated. 
38

2.2.2 Hardware Testing

To date, versions of the C fiber braid element have been run for over 140 hr. The initial 

checkout tests of the material used a piece of C braid (≈0.6 m (≈2 ft)) that was wrapped around a 

constant diameter (≈2.5-cm (≈1-in)) Al tube and strung between electrical inputs, creating a singlepass spiral, double-ended heater design as shown in figure 19. The mandrel was not grooved to 

establish a specific pitch to the spiral for this initial rough test. This element was run for a total 

of 100 hr at 1,200 W, remained stable throughout the run, and the material maintained a similar 

appearance before and after the 100-hr test. Heating appeared to be even from end to end. At 

140 Vdc from the power supply, the very high resistance of the C fiber element limited the current 

that could be pushed through the element, and the maximum achievable power level at 140 V was 

reduced. Regardless of the heater material and design, in a voltage-limited operational regime, to 

reach 1,200 to 6,000 W at (or less than) 140 V (allowing a 10 V margin before reaching the limitation of the power supply), a resistance of no greater than 16.3 and 3.3 Ω, respectively, is required. 

At the maximum 150 V that could be provided by the power supply, the measured current in the 

element was 6.3 A, yielding a resistance value of 23 Ω. As was noted for the graphite rod heater elements, the resistance of the C fiber braid material decreased with temperature, from 28 Ω at room 

temperature to the measured 23 Ω during test at 950 W. Operation of multiple C fiber braid elements in parallel would reduce the total resistance (two parallel elements would have resistance of 

R = 1/(1/28+1/28) = 14 Ω), but independent operation of elements at high power with a single power 

supply would be limited in the maximum power that could be achieved due to the high resistance.

The second test of a C fiber braid element utilized a braid material coated with resin in an 

attempt to lower the resistance and to allow the braid to be carbonized. This element was constructed using the 1.04-cm (0.410-in) Al2O3 mandrel applied in the refractory metal wire heater 

tests, creating a double-pass, single-ended heater element (using ≈0.3 m (≈12 ft) of braid material 

versus ≈0.6 m (≈2 ft) in the previous test). Photos taken after completion of testing are shown in 

figure 20. The element was coated by dipping the entire assembly (mandrel and braid) in resin. 

This resulted in resin coating the entire surface of the element, including both the spiral-wrapped 

braid and the raised Al2O3 pieces between each rotation of the braid. Because the resin is conductive, resin fully coating the element would create a dead short. To remove the excess resin 

from the mandrel, a buffing wheel was used to clean out the Al2O3 between the wrapped braid. 

In the future, to maintain a consistent mandrel diameter after removal of the excess resin, a better method might be to manufacture the mandrel at a slightly larger than required diameter while 

maintaining the diameter to the lowest surfaces of the groove. After dipping the element in resin, 

it could then be machined down to the desired diameter. With the resin coating, 150 V produced 

a measured current of 7.7 A, yielding a resistance of 19 Ω—a drop of 4 Ω from the previous test. 

However, the resin-dipped C braid did not appear to have even heating (one end of the heater produced a dull glow relative to the remainder of the element), possibly due to an uneven coating of 

resin on the element. This element was tested at 1,200 W for ≈10 hr.

A third C fiber braid element was constructed using the same C braid material and was 

wrapped around the 1.04-cm (0.410-in) Al2O3 mandrel, but no resin was applied to the element 

and the braid was not carbonized. The element demonstrated good performance at 900 W over 35 hr 

of run time. The power supply was set to its maximum 150 V and the current measured through the 

element was ≈6 A, yielding a resistance of 25 Ω. 
39

2.2.3 Preliminary Conclusions 

Tests run to date using the C fiber braid material have been directed toward verifying material capabilities at temperature. Initial results look promising for the power levels set forth in the 

requirements if elements can be operated in parallel in order to reduce the total resistance seen in the 

power supply. Because the Al2O3 mandrels of the other simulator concepts have a possibility of melting at high temperature/high power combinations (the maximum recommended operating temperature of Al2O3 is 1,800 °C), using C fiber material is quite attractive, as the material can be wrapped, 

carbonized, and the mandrel melted away, leaving a true C braid heater. Note, however, that the 

mandrel temperatures will decrease when additional heat removal methods are incorporated in the 

element tests. In addition to potentially higher temperature operation, the tension required to wrap 

C fiber is much less than that required for refractory wire, simplifying heater element fabrication. 

Carbon braid has allowed continuing evaluation of C for the manufacturability of smaller 

diameter heaters, but continued study is necessary. For instance, it was found that when C fiber is 

braided, some of the fibers have a tendency to fracture (a small fraction of the 3,000 fibers used to 

produce the braid), resulting in a ‘hairy’ appearance of the braid material. This is much like taking a 

green stick and bending it without folding it in half, causing some fraying of the stick, but no fracture. 

A similar result is obtained when C fiber is braided and wrapped around a mandrel. Although the 

test elements exhibited this type of fraying, it did not appear to affect the integrity of the material as 

a heating element. However, it is desirable to better understand the behavior of the C braid material 

and to find a potential solution to the fraying problem so that consistent elements can be produced. 

This problem has been discussed with the material manufacturer, A&P Technology, and they have 

suggested that a heavier resin coating on the fiber may reduce fraying. The second C fiber braid heater 

element, which was carbonized, did not have this frayed appearance after the carbonization process.

Research and development has just begun on this concept, but the C fiber braid tests show 

great promise, and the primary issues with this concept are in regard to manufacturability. The 

effort to develop an acceptable C fiber braid heater element was temporarily sidelined to focus on 

the development of refractory metal heater elements. Because some of the proposed reactor concepts 

would incorporate refractory metals in their design, C may not be the material of choice in the heater 

element design due to potential interactions between C and the refractory metal at temperature. 

Refractory metals, however, are also adverse to O2 at elevated temperatures, which may preclude the 

use of an Al2O3 mandrel for any of the proposed heater element designs. Potential incompatibilities 

can be averted by selecting an alternate material for the thermal simulator assembly or by altering its 

geometry or means of isolation from the test article. For instance, a BN coating could be applied to 

the C fiber-wrapped heater assembly using a vapor deposition process to provide electrical insulation 

and to improve materials compatibility. This coated assembly could then be inserted into a sheath of 

appropriate material for compatibility with the test article, or the sheath could also be built up in a 

vapor deposition process, with bonding provided by using a functional gradient between the sheath 

material and the BN coating. If the sealed assembly design is selected, development of an acceptable hermetic seal for the heater/sheath assembly will be critical to its success. This will be discussed 

further in the following section. Additional investigation of the various material options will address 

compatibility issues for the requirements set forth by the selected reactor designs. 
40

2.3 Sealed Thermal Simulator Configurations

Sheathed thermal simulators are considered to prevent contamination of the reactor test 

article from the heater material (testing at high temperature over time can cause the heater to 

degrade) and to better emulate actual nuclear fuel pin designs by matching conductivity across the 

pin. This section addresses the initial design and analysis for a sealed thermal simulator in which 

the heater element is sealed inside a sheath with a low-pressure inert gas fill. 

2.3.1 Engineering Design

Three methods are being investigated to seal the heater inside the sheath with an inert gas 

fill, both to prevent contamination of the reactor simulator test article from the heater element and 

to emulate a better model of a nuclear fuel pin. The processes currently being considered include 

active metal brazing, plating the ceramic coupler cap and then brazing the sheath to it, and plating 

(encapsulation) the ceramic to the metal interface. 

To prevent contamination of the test article and to seal an inert gas (He) inside the sheath, 

a high integrity seal must be developed. The sheath is manufactured as a tube and two endcaps 

are used to provide a closed-end unit. The upper endcap (blank cap, no power feedthrough) can 

be welded to the body of the sheath before the heater element is inserted. On the opposite end, the 

metal or ceramic endcap is designed to allow the coupler to penetrate. This general configuration 

was shown schematically in figure 10. If a metal endcap is used, insulating material (electrically 

nonconductive) must be included around the coupler feedthrough surface. The metal cap could 

then be welded to the sheath around the OD, but it must be brazed to the coupler. The ceramic to 

metal interface could be plated using a vacuum plasma spray (VPS) deposition, which would create a uniform bond without any gaps, ensuring a good seal. Alternately, if a ceramic cap is used, 

no additional insulating material is required around the coupler feedthrough, but it must be sealed 

by brazing around the OD and at the coupler feedthrough. Two concerns emerged with regard 

to this braze as the design proceeded: (1) The desired maximum temperature of the braze joint 

exceeds that of current braze technology and (2) brazing in a He (or other inert gas) environment 

may not be possible.

Finding an appropriate high-temperature braze material to seal the sheath was difficult. 

Current metal to ceramic brazing technology limits the temperature of the braze joint to ≈800 °C, 

significantly lower than the temperature required for the original sheathed heater design (≈1,000 °C). 

In the original design, the end of the mandrel was positioned ≈1.9 cm (≈0.75 in) from the end of the 

sheath, where the braze joint is located. To reduce the temperature at the braze joint, the sheath and 

coupler were lengthened by ≈7.6 cm (≈3 in). Investigation continues to identify a higher temperature 

braze so that the brazed sheath option is fully understood. Under contract with MSFC, Edison 

Welding Institute (EWI) is developing a metalizing seal, where the ceramic is first plated with a 

metal such as Ni and then brazed, providing improved bonding to Al2O3. Additionally, Materials 

Research International (MRI) is working to develop an active metal brazing technique using an 

alloy braze material made by Morgan Advanced Ceramics that can maintain its integrity at high 

temperature.
41

Discussions with potential vendors that would perform the braze operation also led to 

concerns about sealing He inside the sheath. The brazing operation must be performed in a 

vacuum furnace. The current assembly plan would perform this braze with the furnace filled to 

low pressure (≈10 kPa (≈76 torr)) with UHP He gas, allowing the gas to be sealed into the assembly during the sealing process. However, the additional conductivity provided by the He gas in 

the furnace would transfer a significantly greater amount of heat to the walls of the furnace than 

would be transferred by thermal radiation alone. As a result, vendors are presently concerned 

about increased temperatures damaging the vacuum furnace during the brazing procedure. If this 

concern cannot be addressed, an alternate method of filling the assembly with low pressure He 

may be required. 

In an alternate assembly sequence, the nonlead endcap assembly could be used as a gas fill 

location. This could be accomplished by first welding and brazing the power feedthrough end of 

the assembly, and then performing the final weld of the nonlead endcap assembly in an inert gas 

(He) environment. If this is not possible, then an alternate solution would perform all welding and 

brazing of the endcaps in a vacuum environment. The nonlead endcap assembly could then be 

used as a fill location for filling the heater assembly with a partial atmosphere of He gas once the 

ends are sealed and leak checked. A 1/8-in-diameter tube connected to the nonlead endcap would 

be suitable for filling the heater assembly with UHP gas, and after fill, the tube would be flattened 

and electron beam welded to form a leak tight seal.

The primary concern that emerged from this investigation is that the desired maximum 

temperature of the braze joint, ≈1,273 K (≈1,000 °C), exceeds that of current braze technology. 

Although a redesign of the heater sheath and coupler allowed the temperature at the braze joint 

(moving the joint a greater distance from the heated section), alternate solutions were pursued. 

Discussions continue with potential brazing shops in an attempt to develop a hightemperature braze for dissimilar materials that will allow a gas to be sealed inside the assembly. 

Possible vendors include Summation Research, MRI, General Atomics, and Altair Technologies, 

as well as in-house engineering groups. The in-house design group has finalized machine design 

drawings for a brazing jig and potential machine shops to fabricate the part are being identified. 

While industry has successfully brazed the heater assembly components together in a manner that 

relieves the mechanical stress on the electrical connection, a seal that successfully captures He 

inside the sheath has yet to be demonstrated. Lessons learned from industry will be employed in 

the in-house effort, along with a redesign of some of the components. 

Examination of an alternate sealing effort using a mechanical seal option for thermal 

simulator assemblies has been initiated. Although this type of seal would be nonhermetic, it may 

be acceptable if the materials selected for the reactor core and the thermal simulators are compatible over long test times at elevated temperature. If this proves to be a viable option, continued 

pursuit of a high-temperature braze joint may not be required. The focus of the mechanical seal 

development work is on manufacturing components with tight tolerances (to achieve as tight a seal 

as possible) and on investigating commercially available mechanical seals that could be used in the 

heater assembly.
42

2.3.2 Hardware Testing

A sheath design has been developed and will be sealed with a ceramic or metal cap using 

one of the brazing techniques discussed above. Initial tests consider the use of a graphite heater 

element in the sheath, as these heaters have an extensive test history and have been demonstrated 

to be very work-hardy in all test conditions. Future applications could consider other heater element designs, such as the refractory wire wrapped or C fiber braid designs discussed above. The 

sheath material or design may be altered as a function of the application. Electrical hookup varies 

with heater design, and integration in the reactor simulator (test article) must be taken into consideration in the overall design process.

MRI and EWI have not been successful in initial attempts to make the high-temperature 

braze seal. No assemblies have been received to date with He inside. The first attempt by MRI 

revealed a hairline crack in the Al2O3 endcap, and a second attempt appeared to be successful, 

but the resistance measurements made upon receipt at MSFC indicated that all He had leaked out 

of the assembly. Generally, brazing shops do not handle parts that are on the order of the heater 

lengths (≈51 to 64 cm (≈20 to 25 in)) and there is a lack of sufficient knowledge on braze materials that can maintain integrity at the target sustained operating temperature of 1,000 °C. Additionally, vendors are reluctant to perform the braze operation inside a He-filled vacuum furnace 

(required to trap He in the assembly) because of the potential damage to the furnace. (Helium 

increases conduction to the furnace walls, potentially increasing the wall temperature above 

design specifications.)

2.3.3 Preliminary Thermal Analysis

Preliminary calculations and analyses have been conducted to assess the performance of 

high-temperature heater elements as bare elements and inside a sealed sheath. These analyses were 

performed in parallel with the testing program discussed above and, as yet, have not fed back into 

the element design. Thermal analyses will be utilized in refining the simulator designs. The overall 

goal of the test and analysis work has been to identify and evaluate candidate materials and resistive heater configurations to establish a database of design concepts that could be used to meet the 

requirements of the simulator designs once the particular fuel rod design options have been specified for nonnuclear reactor testing.

The initial phase of the simulator thermal analyses has focused on the following tasks:

• Assessment of simulator cross-section design, with emphasis on design aspects such as internal 

geometry, internal conduction, and radiation performance and enhancements. The impact of 

the sheath design on heater temperatures is also considered.

• Assessment of end effects and interface temperatures, particularly on the electrical leads. 
43

As stated, these analyses are preliminary in nature and sophistication, and it is expected 

that, as the development proceeds, the complexity and specificity of these analyses will increase. 

A brief synopsis of the results generated thus far is provided in the following sections. Most of the 

analyses were performed using Mathcad or SINDA/FLUINT. More details of each analysis are 

provided in appendices H and I.

2.3.3.1 Cross-Section Analysis. The primary focus for the simulator cross-section analyses was to identify and evaluate means of enhancing the heat transfer from the resistive heater 

element to the sheath. Figure 21 shows the basic sheathed heater configuration and associated 

dimensions used for the cross-sectional thermal calculations. Improving the heat transfer across 

the gap reduces the heater temperatures that are reached for a given input power due to the 

greater efficiency of transferring the heat into the core. (Input power is selected so that the entire 

heater/sheath assembly mimics the power and temperature for a nuclear fuel pin in a given reactor design.) Reducing the heater temperatures allows consideration of a greater variety of heater 

materials and design options and potentially increases heater element life.

Heater/Sheath Gap

Radius 0.2700 in

Radius 0.2325 in

Radius 0.1750 in Graphite Heater

Element Segments

Refractory Metal Sheath

Figure 21. Basic configuration of the proposed heater/sheath geometry (not to scale).

The heater/sheath length used for the various calculations was 50.8 cm (20 in). While this 

solid graphite heater element configuration was used as the basis for the calculations, most of the 

trends identified in the analysis will still hold for other heater configurations, such as braided or 

solid wire wrapped around a ceramic insulating mandrel. The majority of the thermal analyses 

applied a heater power of 6,000 W to establish a conservative estimate of the maximum temperature differential that might be observed across the heater/sheath gap. All cross-section analyses 

considered a constant sheath temperature of 1,273 K to represent a target heat source (fuel pin) 

temperature that was assumed to be isothermal. In practice, the heater power and desired heat 

source temperature (at the sheath surface) will be specifically selected for each reactor design and 

test condition.
44

Enclosing the heater element in a sheath introduces a series of thermal resistances that 

must be overcome to transfer heat from the heater element into the reactor simulator. While the 

sheath provides some thermal resistance, the relatively high thermal conductivity of the metal 

composing the sheath transfers heat efficiently. The gap between the heater and the sheath and 

possibly the gap between the sheath and reactor core introduce the most significant resistances 

from the heater to the core block. Initial designs consider a heater/sheath gap that is filled with 

either a low conductivity inert gas or is under vacuum conditions. Under vacuum conditions, the 

primary mode of heat transfer is thermal radiation. When the gap is filled with an inert gas, the 

dominance of radiation versus conduction is dependent on temperature, emissivity (of both the 

heater element and sheath), and gap size. As the gap size increases, the benefit of conduction in 

removing heat from the element decreases. While the high operating temperature of the heater 

improves the radiative heat transfer, the thermal resistance is still appreciable. Use of radiative 

heat transfer alone to move heat from the heater to the sheath necessitates relatively high heater 

temperatures to reach the desired sheath temperature and input power. A number of configuration 

changes were identified to enhance the heat transfer across this gap, including:

• Selecting the proper gas to fill the gap.

• Reducing the size of the gap. 

• Increasing the radiative properties (aborptivity and emissivity) of the heater and sheath. 

• Adding conductive gap fillers, such as powders or solid liners.

Each of these suggested modifications are described below with their potential enhancements. The level of enhancement is quantified in terms of two values: the thermal conductance 

and the temperature differential across the gap. The temperature differential is that difference in 

temperature between the heater and the sheath necessary to transfer 6,000 W between the two.

2.3.3.1.1 Improvement of Thermal Conductance Due to Gas Selection. The gas selected 

for use inside the sheath at these extremely high temperatures should be nonreactive to preserve 

the integrity of the heater and sheath materials and to prevent contamination of the test article 

(reactor simulator) if the gas leaks out of the sheath containment. As such, the use of inert gases, 

such as Ar, He, and xenon (Xe), should be considered for inclusion into the sheathed heater design. 

As indicated by the thermal conductance values summarized in table 8, He is by far the most thermally conductive of the gases considered. Note that the values provided are solely based on conduction across the gap and correspond to gas properties at the sheath temperature of 1,273 K. A 

more accurate estimate of the gap conduction would allow the gas conductivity to vary depending 

on the local gas temperature as well as the sheath and heater temperatures. Because gas conductivity typically increases with temperature, the expected thermal conductance would increase if 

the average temperature were applied rather than the temperature at the sheath. Hence, the effective gap conductances presented in table 8 may provide a conservative (low) estimate of the heat 

transfer across the gap.
45

Table 8. Effective gap conductance for different inert gases based on geometry depicted 

in figure 21 (radial gap size: 0.15-cm (0.058-in)) properties at 1,273 K).

Gas in Gap Thermal Conductance (W/K)

He 4.8

Ar 0.56

Xe 0.29

Derived properties: He,10,11 Ar,12 Xe,11,13

2.3.3.1.2 Reduction of Gap Size. The thermal conductance of He, which is an order of 

magnitude higher than that of either Ar or Xe, necessitates a temperature differential of over 1,251 K 

across the gap to drive 6,000 W from the heater to the sheath. To maintain a sheath temperature of 

1,273 K would require the heater temperature to be over 2,400 K. The temperature differential required for Ar or Xe fill gas would be significantly greater. Other gases, such as halogens, can be evaluated, with special attention given to their potential for reacting with the sheath or heater materials. 

2.3.3.1.3 Reduction of Gap Size. The thermal conductance of a gas-filled gap roughly 

scales with the reciprocal of the gap size. As the gap size is reduced, the conductance can increase 

significantly. For example, at the current radial gap size of ≈0.15 cm (≈0.058 in) the gap conductance (considering conduction alone) through the He gas is ≈4.8 W/K, as listed in table 8. However, 

if the radial gap size is reduced to ≈0.25 mm (≈0.010 in), the conductance improves to over 26 W/K 

(gas conduction alone). This reduces the temperature differential from 1,251 K (gap size 0.15 cm 

(0.058 in)) to ≈225 K and results in a heater temperature of ≈1,500 K (versus 2,400 K for a gap size 

of 0.15 cm (0.058 in)) to maintain the sheath at 1,273 K. Appendix H provides a curve of conductance and temperature differential for a number of other gap sizes.

Although small gap sizes improve the heat transfer from the heater element to the sheath, 

small gaps introduce concern over achieving tight manufacturing tolerances, the ability to assemble the thermal simulator, and the effects of differential thermal expansion at elevated temperature, which could result in inadvertent electrical contact between the heater and sheath. Coating 

the heater element with a dielectric material, such as BN, before inserting it into the sheath could 

prevent heater to sheath shorting even if the element were to bow at elevated temperatures. If 

the gap width is reduced by increasing the sheath wall thickness, while maintaining a constant 

sheath OD, then the axial conductance of the sheath will increase due to the additional material. 

This increased axial conductance can diminish any axial temperature gradients (power profile) 

designed into the heater to match the heat distributions inherent in the nuclear fuel rods. Therefore, the complete heater/sheath assembly must be considered when tailoring the heater profile 

to achieve the desired heat flux profiles. 

2.3.3.1.4 Sensitivity to Optical (Thermal Radiative) Properties. At the high temperatures 

achieved in the heater testing for application to nonnuclear testing of a reactor simulator (in concert with the low conductive and convective conditions), radiative and conductive heat transfer 

contribute almost equally to the heat transfer across the heater/sheath gap. Table 9 provides estimates of the temperature differential across this gap for conduction, radiation, and their combined 
46

effects. These calculations assume a gap width of 0.15 cm (0.058 in) per figure 21. For contribution 

due to conduction, the annular gap was filled with He gas at a pressure of 101 kPa (760 torr). For 

the radiation-only calculation, a vacuum gap was assumed. 

Table 9. Relative effect of radiation and gas conduction for baseline simulator geometry.

Configuration Temperature Differential (K)

Gas conduction only 1,251

Vacuum gap (radiation only) 1,242

Gas conduction and radiation 763

The radiative heat transfer calculation in table 9 assumes that the heater and sheath are 

diffuse and gray. Their respective emissivities are 0.5 (representing a mid-range emissivity ceramic 

substrate, such as Al2O3 or BN instead of the graphite heater element) and 0.2 (representing a clean/

slightly polished metallic sheath, such as stainless steel or a refractory metal). Using gas conduction 

in coordination with radiation significantly reduces the temperature differential across the gap, demonstrating the benefit of having a gas filled versus a vacuum gap in the thermal simulator design.

The temperature differential across the gap can be further reduced by increasing the emissivities of the heater and sheath surfaces. Table 10 provides the estimated temperature differentials 

that would result from using various combinations of surface optical properties. The conditions 

used for these calculations are based on the geometry depicted in figure 22. A sheath temperature 

of 1,273 K is imposed, and the sheath is assumed to be isothermal. These calculations suggest a 

few possible design enhancements that could be pursued. First, the emissivity of the sheath inner 

surface may be enhanced. This may be possible by coating the inner surface with a high emissivity material or by chemically reacting (‘oxidizing’ or ‘nitriding’) the surface to create a stable, 

high emissivity surface. Emissivity coatings that may be applicable to the heater design are currently under development via a small business innovative research (SBIR) contract with MSFC. 

Note that chemically modifying the sheath inner surface could adversely affect the sheath if it is 

constructed from a refractory metal. Alternately, the heater element design could be altered or its 

component materials could be selected to enhance the emissive properties. The analyses discussed 

thus far have assumed that heat transfer from the heater to the sheath is relatively insensitive to the 

heater type (a solid graphite resistive element or spiral wound braid or wire, such as C fiber, Re, 

etc.). However, the emissivity of each of these potential heater elements is significantly different, 

so that the radiative heat transfer across the heater/sheath gap is highly dependent on the selected 

materials. The solid graphite used in heater designs tested in multiple test articles at the EFF-TF 

has an inherently high emissivity. For wire/mandrel heater configurations, other high emissivity 

ceramic substrates, such as weapons grade aluminum nitride (AlN), could also be investigated to 

enhance the emissive properties of the heater.
47

Cu Power Leads

Representative Core Surface

He Gap

Nb Sheath

Cool Test Chamber Wall (≈50 oC) Note: Drawing not to scale.

Mo Couplers

Graphite Heater Element

Figure 22. Approximate geometry used in the SINDA model.

The results presented in table 10 indicate that enhancing either the sheath or the heater 

should improve the performance of the heater assembly by improving radiative heat transfer 

across the gap. Improving the sheath inner surface emissivity from 0.2 to 0.7 appears to induce 

a more significant reduction in the temperature differential than changing the emissivity of the 

heater surface. Combining both emissivity enhancements reduces the temperature differential 

still further. As some of these enhancements require coatings or chemical treatments, subsequent 

evaluation and testing are required to ensure chemical compatibility and longevity of the altered 

designs. While the addition of gas to the heater/sheath gap will improve the overall performance of 

the thermal simulator, reducing its temperature and possibly improving the longevity of the heater 

assembly, it could also retard outgassing of the emissivity coatings, improving the integrity of the 

coatings over long test times.

Table 10. Thermal performance of combinations of optical property enhancements (assuming 

radiation heat transfer only).

Heater Emissivity Sheath Emmissivity Temperature Differential (K) Comments

0.5 0.2 1,242 Ceramic heater substrate—slightly polished sheath inner surface

0.5 0.7 840 Ceramic heater substrate—coated or chemically modified sheath 

inner surface

0.9 0.2 1,130 Graphite style heater element or high emissivity ceramic substrate—

slightly polished sheath inner surface

0.9 0.7 637 Graphite style heater element or high emissivity ceramic substrate—

coated or chemically modified sheath inner surface

2.3.3.1.5 Conductive Gap Fillers. An alternative way to further enhance heat transfer across 

the gap is to fill the gap with a more conductive material (nongas). This could entail filling the gap 

with a thermally conductive, electrically isolating powder (BN powder) or inserting a single solid 

liner (or a segmented solid liner) having the same thermal and electrical attributes (thermally conductive, electrically isolating). In a recent microgravity experiment development activity, BN powder was studied as a gap filler in a similar application that required enhanced conduction at high 

temperature.14 This concept has a number of advantages including enhanced thermal performance, 

electrical isolation, geometric conformability, and tolerance of differential thermal expansion. However, previous experience indicated that it was difficult to uniformly distribute and pack the powder 

to prevent axial or circumferential voids, leading to nonuniform heat distribution. The solid liner 

concept is an attempt to address this packing issue. This liner would essentially consist of a ceramic 
48

tube sized to fit in the gap between the heater and the sheath, consuming as much of the gap as 

possible to minimize the resulting circumferential voids between the heater and the sheath. Table 11 

provides preliminary calculations for these two concepts. Both cases use BN ceramic, in a powder or 

solid liner form. The assumed heater/sheath gap (unfilled) is ≈0.15 cm (≈0.058 in) as in the previous 

analyses. Either approach to filling the heater/sheath gap must take into account thermal expansion 

at elevated temperature in producing the final assembly design.

Table 11. Performance of conductive gap filler concepts.

Configuration Temperature Differential (K) Comments

BN powder w/He 220 Properties of He at 1 atm are used—powder assumed uniform

BN liner w/He 228 Liner thickness = 80% of original gap thickness

The suggested conductive gap configurations could significantly reduce the temperature 

differential between the heater and the sheath, allowing the target sheath temperature to be 

maintained at heater temperatures that are quite manageable for a number of materials. While 

both concepts produce similar temperature differentials, it is important to note the sensitivity of 

each to the presence of He. The powder concept will produce slightly degraded, but still useful, 

performance if the He is not present in the interstitial gaps between the grains of powder as was 

observed in previously discussed testing for the microgravity experiment. (This design utilized a 

different powder material.) Although some voids may exist in the absence of He, the acceptable 

performance is likely a result of the particle-to-particle contact that exists whether the He is present or not. On the other hand, if no He is present in the liner concept, then there is only radiation 

across the smaller gaps that are not filled by the liner. Because there are now two serial gaps where 

there used to be only one, the liner acts as a radiation shield, which actually increases the radiative 

resistance and increases the temperature differential. Consequently, the solid liner appears to have 

merit only for a gas-filled gap configuration.

A BN layer could also be added by a method such as plasma vapor deposition, which could 

apply a solid layer to increase the element diameter to correspond to just less than the ID of the 

sheath, allowing it to be inserted with minimal gap. Addition of this layer would not only enhance 

electrical isolation, but could enhance thermal performance by shrinking the gap, as just mentioned, and by increasing the emissivity of the wire or braid being coated. However, consideration 

of differential thermal expansion is critical. Centerless grinding could also be used to grind the 

element to the appropriate size to slide into the sheath if it is built up to a slightly larger diameter. 

This technique can meet ±0.076 mm (±0.003 in) variation on the element OD, allowing the gap 

between the element and the sheath to be as small as possible. Plasma vapor deposition could also 

be used to build up the sheath, setting a fractional gradient to transition from the BN layer to 

the sheath material to establish a strong bond between the materials without any gap. The sheath 

could be built up to a diameter that is slightly greater than that required, so that it could then be 

ground to the proper diameter using the centerless grinding technique. One potential concern 

associated with a full buildup of the simulator using plasma vapor deposition (allowing no gaps 

between component layers) is differential thermal expansion of the various materials used in the 

simulator. Additional analysis would be required to ensure that the individual simulator layers 

would not be subject to cracking during test at elevated temperatures.
49

Each of the suggested modifications to the sheathed thermal simulator design (addition of 

gas, reducing the gap size, improving the emissive properties of the components, or adding conductive gap fillers) was shown to improve the overall heat transfer across the gap and to lower the 

heater temperature necessary to drive the required heat flux. In most cases, potential weaknesses 

of each concept were also briefly identified. While the selection of a specific heater/sheath assembly would likely depend on the specific application, including the desired temperature and power 

levels, the most straightforward design modifications would be to add He gas to the gap (assuming 

the He gas can be contained) and to shrink the gap size either by enlarging the heater diameter or 

by shrinking the sheath ID. The final assembly design will be based on reactor design constraints 

and heater capabilities.

2.3.3.1.6 Impact of the Sheath on Thermal Performance. The impact of encasing the heater 

inside a sealed sheath on the heater temperatures was also assessed. In the typical heater element 

test setup, a bare heater is suspended in the center of a large stainless steel vacuum chamber and 

electricity is applied. The test specimen is allowed to run in this condition for extended periods of 

time, after which it is removed and examined to evaluate its tolerance of high temperature conditions. For the general surface characterization of the test chamber, the inner surface of the chamber is assumed to be uncoated and free of significant oxide buildup or grime, providing a basis 

for estimating surface emissivity. A simple analysis was performed to estimate the test conditions 

for comparison of the expected bare element temperatures with the expected temperatures for the 

element encased inside the sheath and put under the same load. Initial calculations assumed heat 

removal by radiative heat transfer alone. The results of this analysis are provided in table 12. The 

analysis assumed a vacuum environment around the heater and inside the sheath, and as previously 

discussed, gas conduction would help to significantly reduce the sheath configuration temperatures. 

In all cases, calculations assumed a constant diameter graphite rod heater configuration. Similar 

scoping calculations for a braided W wire heater element were presented in section 2.1.2.4. Calculations allowed the sheath temperatures to float while maintaining the vacuum chamber wall temperature at a constant 50 °C.

Table 12. Impact of the sheath on heater temperatures experienced in vacuum testing.

Configuration

Heater Power

(W)

Sheath Temp

(K)

Heater Temp

(K) Comments

No sheath 1,000 – 1,096 –

Sheath 1,000 1,422 1,740 Sheath emissivity = 0.2

No sheath 6,000 – 1,712 –

Sheath 6,000 2,224 2,726 Sheath emissivity = 0.2

Sheath 6,000 1,638 2,057 Sheath emissivity = 0.7

Table 12 summarizes results that indicate that the sheath acts like a radiation shield, 

imposing a significant thermal resistance the value of which depends primarily on its optical 

properties and operating temperature. However, encasing the heater element inside a sheath more 

closely approximates the geometry of a nuclear fuel element for use in the nonnuclear test article 

(while the bare heaters offer convenient testing options, in practice, these heaters will probably 

never be used in a bare configuration). These results also suggest that care should be taken in 
50

selecting the proper sheath and power combinations to avoid over-temperature conditions within 

the sheath or heater materials. Finally, while most of the endurance testing has been completed (in 

most cases without a sheath present), additional high-temperature endurance testing of materials 

may be warranted once the design space (operating temperature, power level, heater design, sheath 

design, etc.) has been narrowed to evaluate these materials at or near use conditions. Note that in 

the test article configuration, the reactor coolant system (LM HPs, pumped gas, or pumped LM) 

will actively remove heat from the surface of the sheath and heat removal will no longer be solely 

by radiation to the vessel wall, significantly reducing the heater and sheath temperatures relative 

to the single element (heater and sheath) testing for the same power level. For instance, the sheath 

temperatures indicated in table 12 would be on the order of 1,400-1,500 K (core block temperature 

≈1,273 K) at full power (6 kW per pin) in a cooled core case, resulting in a lower heater temperature for the same power level.

2.3.3.2 Assessment of End Effects and Heater Lead Wire Thermal Conditions. In addition to examining the gross performance and design of the heater and sheath, the end effects and 

thermal conditions within the electrical lead wires and Mo transition pieces (couplers) are also of 

interest relative to potential material-over-temperature conditions. The present analysis considers 

a graphite rod heater element and Cu wire leads. Molybdenum transition pieces are used between 

the Cu wire and the graphite to reduce the temperature experienced by the Cu leads. 

A simplistic SINDA model was constructed to analyze the thermal conditions within this 

assembly and the basic components of the model are shown in the sketch provided in figure 22. 

The dimensions of the graphite heater rod assembly (heater and sheath) were provided in figure 21. 

A constant diameter graphite rod heater (0.89-cm- (0.350-in-) OD) is assumed for these calculations 

for both uniform and cosine-shaped power distributions. The cosine-shaped axial power distributions apply a peak to average power factor of 1.33. The heater is enclosed in a Nb sheath having the 

dimensions previously cited in figure 21. The heater/sheath gap and the sheath/core block gaps are 

filled with He at 101 kPa (760 torr) and no emissivity enhancements are considered. A brief investigation was performed to determine the effect of gas pressure on the conductance across the heater/

sheath gap. As discussed in section 2.3.1, the intended pressure in the heater/sheath gap is 10 kPa 

(76 torr). For the given gap width (0.15 cm or 0.058 in), the gas is still in the continuum flow regime 

at this pressure. (It begins to approach the mixed flow regime at ≈4 kPa (≈30 torr)) and the difference in gas conductivity at 101 kPa versus 10 kPa is <0.1%.

The graphite heater design is single ended, so the input and output power leads are at the 

same end of the graphite element. Each lead is connected to a 10.2-cm- (4-in-) long Mo transition 

piece (coupler). The details of the connection were not incorporated in the model, but the connection between the Mo and the graphite is assumed to be intimate (or near ideal) with no additional 

resistance modeled between the two materials. Each Mo piece, in turn, connects to a Cu lead 

wire. Both the Mo and the Cu wires are sized to American wire gauge (AWG) 11. The connection between the Mo and the Cu wires is also assumed ideal with no thermal resistance modeled 

between the ends of each piece. The Cu wires were extended to 0.9 m (36 in) in length.
51

As previously stated, the heater is enclosed in the Nb sheath and transfers heat to the 

sheath via thermal radiation and gas conduction with heater and sheath emissivities assumed to 

be 0.9 and 0.3, respectively. In the cross-sectional analyses, the sheath temperature was assumed 

fixed. In this analysis, however, the sheath temperature is allowed to reach a value necessary to 

transfer the heater power to a core boundary condition. The boundary is intended to represent the 

fuel tube (clad) or core block in which the heater/sheath assembly is inserted. In this case, the HP

reactor design, which does not include a fuel clad, is assumed and the boundary condition is set to 

1,325 K, equivalent to the expected boundary condition on a HP-cooled reactor design. The radial 

gap between the sheath and the core block is assumed to be 0.25 mm (0.010 in), and is assumed to 

contain pure He.

In the thermal model, the sheath rejects heat to the representative core boundary by thermal radiation and gas conduction and the core inner surface emissivity is assumed to be 0.3. A 

limited amount of heat may also transfer from the sheath to the boundary through low pressure 

line and point contacts and through end welds or seals, but this is not yet modeled. Thermal 

analysis using radiative and conductive heat transfer (through the He gas) establishes temperature 

trends. However, the current model does not currently include temperature-dependent electrical 

resistivity, so to obtain more realistic correspondence between heater input power and heater 

temperature, subsequent improvements must be incorporated. 

The Mo couplers are assumed to be adiabatic (neither Mo piece is allowed to exchange 

heat with any boundary directly; it only exchanges heat with the heater element and the Cu wire), 

providing a conservative estimate of the temperatures reached in these pieces. The Cu lead wires 

are allowed to radiate to a cool boundary (50 °C). This assumption may be somewhat unrealistic 

given the proximity of the leads to the reactor body, unless active cooling of the electrical interface is employed (as would be the case for a gas-cooled reactor design). Heat generated in the Mo 

couplers and Cu lead wires (by virtue of the ≈60 A current flowing through them) is also included 

in the model. In one case (case 6 in table 13), the ends of the Cu lead wires are assumed to be 

clamped to the cool 50 °C boundary, providing cooling via contact conductance (simulating the 

Cu wire attachment to the power feedthrough at the chamber wall).

Table 13. Summary results from SINDA heater simulation (temperatures are in K).

Case

Cu Lead Mo Transition Heater Element Sheath

Min Max Min Max Min Max Min Max Comments

Case 1 778 1,069 1,107 1,475 1,468 1,471 1,348 1,349 1 kW; PPF*=1

Case 2 778 1,085 1,126 1,547 1,545 1,813 1,362 1,417 3 kW; PPF=1.33

Case 3 778 1,119 1,165 1,707 1,709 1,712 1,394 1,395 3 kW; PPF=1

Case 4 778 1,122 1,168 1,720 1,723 2,153 1,398 1,508 6 kW; PPF=1.33

Case 5 778 1,176 1,232 1,985 1,993 2,001 1,462 1,464 6 kW; PPF=1

Case 6 332 1,121 1,168 1,720 1,723 2,153 1,398 1,508 6 kW; PPF=1.33; cooled lead end

*power peaking factor
52

Based on the design of the graphite heater element, the axial power profile could be uniform (constant heater diameter) or could exhibit a simple cosine distribution along its length 

(smaller heater diameter at the axial center of the heater) to approximate the heat flux profile of a 

fuel rod in a nuclear reactor. The model considers both uniform and shaped power distributions 

but, a constant element diameter (0.89 cm (0.350 in)) is assumed in both cases. The SINDA model 

discretizes the heater into 100 equally sized nodes (0.51 cm (0.2 in) per node). 

At present, the model has not been correlated with test data. Consequently, the characteristics observed in the results are subject to change once the model has been verified. A number of 

cases have been run to characterize the general heater element configuration. Summary results 

are provided in table 13. Plots of temperature versus axial position (for the heater element, sheath, 

couplers, and lead wires) are provided for cases 2 and 3 (3 kW, cosine and flat power distributions, 

respectively) in figures 23 and 24. The corresponding plots for the remaining cases are included in 

appendix I.

2,000

1,800

1,600

1,400

1,200

1,000

800

600

400

200

0

2

1.8

1.6

1.4

1.2

1

0.8

0.6

0.4

0.2

0

Case 2:

Cosine power distribution, 3 kW total heater power

He-filled gap, Nb sheath—no emissivity enhancement,

4-in Mo transition, Cu leads—only radiative cooling

–40 –30 –20 –10 0 10 20

Total power: 3 kW

Peak: 200 W/in

Avg: 150 W/in

Calculated for 0.2-in nodes

Heater node T(K)

Sheath node T(K)

Mo node T(K)

Cu node T(K)

Q into node

Temperature (K)

Normalized Heater Power

Position (in)

Figure 23. Axial temperature profile for case 2.
53

2,000

1,800

1,600

1,400

1,200

1,000

800

600

400

200

0

Case 3:

Uniform power distribution: 3 kW total heater power

He-filled gap, Nb sheath—no emissivity enhancement,

4-in Mo transition, Cu lead—only radiative cooling

–40 –30 –20 –10 0 10 20

Total power: 3 kW

Power density: 150 W/in

Calculated for 0.2-in nodes

Heater node T(K)

Sheath node T(K)

Mo node T(K)

Cu node T(K)

Q into node

2

1.8

1.6

1.4

1.2

1

0.8

0.6

0.4

0.2

0

Normalized Heater Power

Temperature (K)

Position (in)

Figure 24. Axial temperature profile for case 3.

Figure 23 and 24 show the temperature profile for various components of the heater/sheath 

assembly for cases 2 and 3. Both cases assume 3 kW total heater power (Q), applying a cosine 

power distribution with a power peaking factor (peak to average power) of 1.33 (fig. 23), or a uniform power distribution (fig. 24). This input power is calculated at each node-defined heater node 

in the model and is indicated as ‘Q into node’ in the corresponding plots. As discussed previously, 

the model assumes a graphite heater element encased in a Nb sheath (no emissivity enhancements) 

with a He-filled gap and power input is provided via 10.2-cm (4-in) Mo transition pieces (couplers) 

connected to Cu leads wires. A position of zero on the x-axis corresponds to the start of the heater 

and negative position corresponds to the lead wires and transition pieces. 

The current analysis applies a constant cross section for both power distributions. To accomplish a shaped power distribution with a graphite rod heater, the diameter of the element is reduced 

for a portion of its length at the axial center of the element. The estimated temperatures (using a constant cross section) are probably lower than they might be if the cross section had been contoured 

because (1) the constant cross section probably helps reduce temperature gradients by maintaining 

a larger heat flow path, (2) the contoured element has less surface area, driving up the temperatures 

necessary to reject the local heating, and (3) the gap between the graphite and the sheath is increased, 

at the center of the contoured element, reducing the heat transfer by conduction across the gap. 

Because the wire-wrapped heater elements have a constant diameter regardless of power profile 
54

(the profile is accomplished by varying the pitch of the wire wrap rather than by contouring the element diameter), the differential heat transfer for the flat versus shaped power profiles will not present 

the same difficulties as noted for the graphite elements.

In case 2, the sheath temperature varies from 1,362 to 1,417 K in order to reject the heater 

power to the 1,325 K boundary representing the core block. The maximum heater temperature 

reaches 1,813 K. The temperature differential between the sheath and the heater varies with the local 

heat input. A slight elevation is seen in the temperature of the transition piece for this case. The local 

maximum results from the fact that the heat generated in the Mo transition is not allowed to transfer 

directly to any boundary. The Cu wire, on the other hand, exchanges heat radiatively with the cool 

boundary (chamber walls), causing its temperature to drop despite the fact that heat is being generated in it due to electrical resistance. The Cu reaches a fairly uniform temperature once it is well 

away from the coupler. The heater temperatures for case 3, which assumed uniform power distribution, was ≈1,712 K (calculations indicate a range of 1,709 to 1,712 K) and the sheath temperature was 

1,395 K. Because the local heat input was uniform along the length of the element, the temperature 

differential between the sheath and the heater was nearly constant.

In examining the results of the various cases, a number of characteristics are observed. 

First, by allowing the sheath temperature to vary, the sheath and heater temperatures are higher 

than were seen in the cross-sectional analysis as a result of the differences in the established 

boundary conditions. Cross-section analyses fixed the sheath temperature at 1,273 K; the current 

analysis establishes a reactor boundary condition (a core block temperature of 1,325 K) outside 

the sheath and adds a significant resistance between the sheath and the reactor boundary. The 

resulting increased sheath temperature illustrates the significance that the external (to the sheath) 

heat rejection efficiency can have on the operation of the heater assembly. 

Second, the 10.2-cm (4-in) Mo transition pieces appear to sufficiently reduce the temperature extremes seen in the Cu lead wires. The addition of low-pressure, high-purity gas in the test 

chamber would provide additional reduction in the Cu lead temperatures. Given cooling by radiation alone, the Cu temperatures near the Mo interface do not exceed 1,176 K, which is well below 

the melting point of Cu (1,358 K). Further analyses are planned to evaluate the impact of lengthening the Mo couplers to assess the impact of this parameter on the temperature of the braze 

joint. Given that ideal connections were assumed in the current model, the details of the interface 

between the Mo and the Cu leads should be evaluated more closely. Various options may be available to enhance the cooling of the Cu, but this depends strongly on the reactor simulator design 

constraints. In the current model, the Cu emissivity was assumed to be 0.1. More detail is needed 

to examine the thermal impact of the various electrical insulation options under consideration. 

Third, for most of the cases evaluated, the temperature at the end of the Cu opposite the 

heater is allowed to float (boundary condition set to dT/dx = 0 at the end of the Cu lead). This 

is probably not realistic and, in fact, produces a temperature that may adversely impact electrical connections outside the chamber. In case 6, the end was assumed to be clamped in such a 

way that it could be cooled via a contact conductance to the 50 °C boundary (simulating the Cu 

wire attachment to the power feedthrough at the chamber wall). This cooling was based on performance achievable with typical thermal compounds used to cool electronics. In this case, the 
55

clamped end of the Cu cooled down substantially, as expected, but the temperature of the end in 

contact with the Mo transition was not reduced due to the relatively long length of the lead wires. 

Finally, it appears that the heater temperature profile, while impacted slightly by the heat 

transfer through the transition and lead wires, is not very sensitive for the cases evaluated. 

2.4 Tantalum Wire-Wrapped Heater Element

Initial Ta wire-wrapped heater elements were fabricated using a 1-mm (0.040-in) wire 

wrapped around the 1.16-cm (0.460-in) Al2O3 mandrel. As discussed in section 2.1.2.1, these element tests indicted exceptional performance at 1,200 W, with the tested element achieving 101 hr 

of operation in vacuum, accumulated over multiple thermal cycles. The goal of the tests discussed 

below was to reach 6 kW per element and to operate the element for a minimum of 100 hr at this 

elevated power. Both 0.5- (0.020-in-) and 0.64-mm- (0.025-in-) diameter Ta wires, configured in a 

three-wire braid, were tested. 

2.4.1 Element Testing in Vacuum

Four separate attempts were made to test braided Ta wire in vacuum, considering both 

sheathed and unsheathed configurations. In each case, the heater element was constructed from 

three braided Ta wires and both 0.5- (0.020-in-) and 0.64-mm- (0.02-in-) diameter wires were 

attempted. 

2.4.1.1 Test 1: Tantalum Three-Wire Braid, Sheathed, Vacuum Environment. The first and 

second elements tested were encased in a Mo sheath. These tests were terminated when the Ta 

wire shorted out to the sheath during the test. Because of thermal expansion, the wire contacted 

the sheath when it achieved operating temperatures. Therefore, for this wire diameter, the groove 

depth cut into the mandrel (currently 1.6 mm (0.063 in)) should be increased to prevent the wire 

from expanding beyond the edge of the mandrel and coming into contact with the sheath. Additionally, the transition of the wire from the spiral wrap through the hole that prevents unwrapping, shown in figure 25 with design details shown in appendix J, should be improved. Because of 

the limited ductility of the wire, this transition was not smooth in the constructed heater elements, 

which caused the wire to sit slightly above the Al2O3 mandrel and allowing easier contact with the 

sheath. 

Wire Transition to Prevent Unwrapping

Figure 25. Wire transition from the spiral wrap at end of mandrel (complete 

mandrel design with dimensions is included in figure 65).
56

2.4.1.2 Test 2: Tantalum Three-Wire Braid, Unsheathed, Vacuum Environment. The third 

and fourth Ta vacuum tests were performed without the sheath. The third attempt ended when 

the lead wire sagged sufficiently to contact the Ni foil used to support the TC connector. No real 

damage to the heater was observed, but one power lead was slightly shorter after the event. The 

same heater element (0.64-mm (0.025-in) three-wire braid) was reconnected for the fourth attempt. 

In this attempt, 6,000 W was achieved and held for 3.5 hr. However, heat removal by radiation 

alone, resulting from operation in vacuum, was insufficient at this power level. As a result, the 

heater reached excessive temperatures and subsequently melted, as seen in figure 26. Tests of the 

Ta heater element were repeated in a high-purity He gas environment, as discussed in the following section. 

Wire Melt Due to Excessive Temperature

Figure 26. Posttest TA heater assembly showing wire failure. 

2.4.2 Element Testing in Helium Environment

For subsequent Ta element testing, a bottle of UHP He gas was connected to the test chamber to allow for backfill of the chamber to a low pressure. This bottle will be replaced by the gas 

purification system when it is ready for operation (see sec. 3.1). The addition of He to the chamber 

allows heat to be removed from the element by conduction in addition to radiation, significantly 

reducing the operating temperature of the heater element. 

2.4.2.1 Test 1: 0.5-mm (0.020-in) Tantalum Three-Wire Braid, Sheathed, Helium Environment. A 0.5-mm (0.020-in) Ta three-wire braid element was assembled using the 1.17-cm- (0.460-in-) 

diameter Al2O3 mandrel. This element was installed inside a 1.37-cm (0.540-in) Mo sheath for 

testing at 5.5 kW. After a ramp up to the desired 5.5 kW over ≈45 min, the element operated for 

≈7 hr with no observed problems (fig. 27). Shutdown was required at the end of the work day. On 

the subsequent test day, the element failed when it was brought up to 6 kW of power from a cold 

condition at the start of the test. At low temperature, the resistance of a refractory metal element 

is low and then increases as the element temperature increases. Because the full power was applied 

at the start of the test, it is likely that the element experienced significant thermal stress as the 

temperature rapidly increased, resulting in element failure. These results suggest that, in order to 

achieve high power levels, the wire material requires a ramp time to reach the appropriate power 

level. This allows the element temperature and, correspondingly, its resistance, to slowly increase, 

minimizing the thermal stress places on the element. Future testing will determine the required 
57

ramp time. Testing at 1,200 W did not require power ramping at the start of test so the full desired 

power could be applied to a cold heater element due to the smaller change in temperature over this 

power range. In contrast to the refractory metal elements, C is unique in that it drops in resistance 

as it heats up. Previous testing experience with C fiber braid and graphite rods demonstrated that 

C-based elements are very robust under all thermal conditions. If C braid can be acquired in the 

resistance range needed for thermal simulators, power ramping may not be required to achieve the 

6 kW power level. This would potentially allow the application of full power at the start of the test 

with no ramp time.

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0:00:00 1:12:00 2:24:00 3:36:00 4:48:00 6:00:00 7:12:00 8:24:00 9:36:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

Elapsed Time (hr)

 Figure 27. Initial test data for a 0.5-mm (0.020-in) TA three-wire braid heater element in 

a He environment (the element failed at the start of the subsequent test day).

Two additional Ta elements were tested after the initial power ramping failure discussed 

previously. In these tests, which considered three-wire braided elements using wire diameters of 

0.64 and 0.5 mm (0.025 and 0.020 in), power was increased gradually to allow the heater temperature to increase at a more reasonable rate. Both elements were tested inside a Mo sheath with the 

chamber backfilled to ≈6 kPa (≈45 torr) He to provide additional heat removal from the element. 

However, the same size mandrel was used in these assemblies (1.17-cm (0.460-in) OD, 1.65-mm 

(0.065-in) groove depth) as in the previous element tests. Therefore, although testing in an inert gas 

environment was expected to significantly reduce the maximum temperature of the heater element 

wire, it was uncertain whether or not there would be sufficient clearance for the wire at 6 kW.
58

2.4.2.2 Test 2: 0.64-mm (0.025-in) Tantalum Three-Wire Braid With Power Ramping, 

Sheathed, Helium Environment. A three-wire braid element using 0.64-mm- (0.025-in-) diameter 

Ta wire was tested on July 1, 2005, and July 5, 2005. On the first day of test, the desired power level 

of 6 kW was approached very slowly, reaching the full power after ≈3 hr. The element continued 

to operate an additional hour of run time at the full power level. On the second test day, the element failed ≈45 min into the test, during the power ramp procedure. Heater power and calculated 

element resistance for each test day are shown in figure 28. This failure occurred when the heater 

element wire came in contact with the sheath due to excessive thermal expansion. This wire diameter will not be tested further until redesigned mandrels having deeper grooves are procured.

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0:00:00 1:12:00 2:24:00 3:36:00 4:48:00 6:00:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(a) Elapsed Time (hr)

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0:00:00 0:07:12 0:14:24 0:21:36 0:28:48 0:36:00 0:43:12 0:50:24 0:57:36

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(b) Elapsed Time (hr)

 Figure 28. Tantalum heater element power and resistance on 

subsequent test days for 0.64-mm (0.025-in) threewire braid: (a) Initial test data, July 1, 2005, and 

(b) final test data, July 5, 2005.
59

2.4.2.3 Test 3: 0.5-mm (0.020-in) Tantalum Three-Wire Braid With Power Ramping, 

Sheathed, Helium Environment. A final Ta element was assembled using a three-wire braid 

constructed from 0.5-mm- (0.020-in-) diameter wire. Due to the previous failures of Ta at high 

power levels, the stated goal of this test was to first achieve 100 hr of operation at 3.6 kW before 

moving to the higher power levels. This testing commenced on July 6, 2005, and ran until the 

element failed after 71 hr of operation at 3.6 kW on July 19, 2005 (fig. 29). Because the automated control system was not yet operational, the element was started and shut down on each 

test day. (Later tests have incorporated continuous runs with the addition of automated control 

and safety systems.) The calculated element resistance was relatively stable at ≈3.3 Ω, indicating that no fundamental changes occurred in the element from one thermal cycle to the next. 

Element failure occurred at 71 hr when the data acquisition system reported an open circuit. 

Posttest analysis of the assembly indicated that the heater element wire had welded itself to the 

interior surface of the sheath. The automated data acquisition system was not yet available during this test. Observations, however, indicated that a brief disruption may have occurred in the 

water flow to the test chamber, which may have resulted in an elevated heater temperature. If 

this is the case, the increased temperature may have resulted in slightly greater thermal expansion of the wire, causing it to contact the sheath. 

Tantalum wire followup testing may consider a braid composed of 0.69-mm- (0.027-in-) 

diameter wire; however, this will require a mandrel with deeper grooves to accommodate the 

increased wire thickness. A short trial test was attempted with the 0.69-mm- (0.027-in-) diameter 

wire (and an existing mandrel with a 1.6-mm (0.063-in) groove depth), resulting in a short circuit 

condition to the Mo heater sheath. A new mandrel design to evaluate the larger braids has been 

produced and submitted to the procurement process and this configuration will allow the braid 

to lie flatter in the mandrel groove, minimizing the potential for shorting to the sheath. In addition, this new deeper groove design introduces the possibility of developing a mechanical seal 

configuration (designed to seal the heater inside the sheath) for a heater assembly with an OD as 

small as 1.4 cm (0.540 in) at the sheath OD (1.6-cm (0.625-in) flange). Element sealing options will 

be discussed in section 3.4.

2.5 Tungsten Wire-Wrapped Heater Element

Initial investigation of W wire-wrapped heater elements considered 0.25-, 0.38-, and 

0.5-mm (0.010-, 0.015-, and 0.020-in) W wire. Use of larger wire diameters was not feasible due 

to the reduced ductility of W as the wire diameter is increased. A five-wire braid, using 0.25-mm 

(0.010-in) wire, wrapped around the 1.17-cm (0.460-in) Al2O3 mandrel accumulated 240 hr of 

test time over 30 thermal cycles, meeting the minimum requirement of 100 hr at 1,200 W in an 

unsheathed, vacuum environment. Additional testing of the five-wire braid in the initial test series 

was performed at higher power levels (3,000 W, 24 hr) in a sheathed configuration. Details of these 

early W element test results can be found in section 2.1.2.4. Based on initial success with 

W wire, it was recommended for further testing at higher power levels.
60

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Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(a) Elapsed Time (hr)

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Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(b) Elapsed Time (hr)

Figure 29. Tantalum heater element power and resistance on the initial and final test days for 

0.5-mm (0.020-in) three-wire braid: (a) Initial test data, July 6, 2005, and (b) final 

test data, July 19, 2005.

Tests of W wire-wrapped elements at higher power levels (up to 6 kW) were performed on 

three-, four-, and five-wire braids using the 0.25-, 0.38-, and 0.5-mm (0.010-, 0.015-, and 0.020-in) 

wire diameters. Results of these tests indicate that W is capable of reaching these high power levels over long periods of time, but that thermal management becomes a significant concern as the 

power level is increased.
61

2.5.1 Bare Element Testing in Vacuum

Vacuum testing was performed on a three-wire braid, using 0.5-mm- (0.020-in-) diameter W 

wire, wrapped around a 1.17-cm (0.460-in) Al2O3 mandrel (see fig. 30). This element operated for 

100 hr at 6,000 W in an unsheathed configuration under a 1.3-mPa (10–5-torr) vacuum. A check 

of the cold resistance of the heater prior to running on each test day indicated that the resistance 

was stable, measuring 0.40 to 0.43 Ω. At this elevated power level, thermal management becomes 

a concern for the mandrel if testing is performed in a vacuum environment. Because heat was 

only being removed by radiation, the mandrel became extremely hot, and sagged in unsupported 

regions, as shown in figure 31. However, despite the drastic change in shape from the start of test, 

the heater did not fail. If tested in the sheathed configuration, the additional increase in temperature could cause the mandrel to melt if additional heat transfer mechanisms are not added to the 

test configuration. Further testing could be conducted to better establish the limits of the heater 

element design and material selections. The addition of inert gas to the test chamber can drastically reduce temperatures, allowing operation at much higher power levels. 

(a) (b)

 Figure 30. Tungsten wire heater assembly—testing in a vacuum environment: 

(a) End view of element during test and (b) side view of element 

during test.

Figure 31. Mandrel distortion due to excessive heating.
62

Data for the initial and final days of testing are shown in figures 32 and 33. During the first 

day of testing, a slow heating rate was applied to bakeout the heater element. Two TCs were operational during test. One TC was lying on the bottom of the test chamber (TC3) and an additional 

TC was placed on the mandrel (TC5). The melting point of Al2O3 is 2,313 K (2,040 °C). The plastic 

deformation of the mandrel indicates that the measured mandrel temperature of ≈2,300 °C on the 

first day of testing was likely correct. However, later test data suggest that neither TC was reading 

accurately. Although TC3 maintained a higher temperature measurement of ≈2,300 °C throughout testing, the mandrel temperature reported by TC5 dropped to less than 1,000 °C, as shown in 

figure 33(b), despite the observed plastic deformation, suggesting possible TC failure. Data shown 

in figures 32(b) and 33(b) indicate that the element resistance was stable throughout the test at 

6,000 W. The total of 100 hr of testing was accumulated over eight thermal cycles. 

1.0E+00

1.0E–01

1.0E–02

1.0E–03

1.0E–04

1.0E–05

1.0E–06

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Chamber Pressure

Heater Power

Pressure (torr)

Power (W)

(a) Elapsed Time (hr)

4

3.5

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0

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1,500

1,000

500

0

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Element Resistance

Temperature, Bottom of Chamber (TC3)

Temperature, Mandrel (TC5)

Resistance (ohms)

Temperature (C)

(b) Elapsed Time (hr)

 Figure 32. Initial test data (day 1) for W three-wire braid: (a) Heater 

power chamber pressure and (b) element resistance and 

measured chamber and mandrel temperatures.
63

1.0E+00

1.0E–01

1.0E–02

1.0E–03

1.0E–04

1.0E–05

1.0E–06

1.0E–07

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0

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Chamber Pressure

Heater Power

Pressure (Torr)

Power (W)

(a) Elapsed Time (hr)

4

3.5

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0

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Element Resistance

Temperature, Bottom of Chamber (TC3)

Temperature, Mandrel (TC5)

Resistance (ohms)

Temperature (C)

(b) Elapsed Time (hr)

 Figure 33. Final test data for W three-wire braid, 89–100 hr: (a) Heater 

power and chamber pressure and (b) element resistance and 

measured chamber and mandrel temperatures.

Results of the 0.5-mm (0.020-in) three-wire braid W element vacuum test indicate that a 

refractory wire element can be designed and built to achieve the highest stated power requirement 

under extreme conditions (well above expected operating temperatures in a core array and given 

element deformation). In this case, heat removal by radiation alone allowed the heater element 

temperature to become excessive, resulting in the mandrel deformation that was seen in figure 31. 

This result indicates that improved thermal management (operation of the heater in a low pressure 

inert gas environment) should be implemented at these high power levels. Additionally, differential 

heating was noted in the coils on the power inlet side of the center heater element support. Every 

other coil from the support back to the power input appears dimmer, as seen in figure 30(b). This 

suggests that the coil could be losing some power due to contact with the support or, given that the 

differential heating appeared to wander during the test, it is very likely that the observed differential heating results from emissivity changes in the element at elevated temperatures.
64

2.5.2 Sheathed Element Testing in Helium Environment

All subsequent tests of W wire-wrapped elements were performed with the heater element 

inside a Mo sheath and with the vacuum chamber backfilled with high purity He gas to ≈5.3–6 kPa 

(≈40–45 torr). Tests considered a 0.25-mm (0.010-in) three-wire braid, 0.25-mm five-wire braid, and 

a 0.38-mm (0.015-in) four-wire braid. All wires were wrapped around a 1.17-cm (0.460-in) Al2O3

mandrel and were tested inside a 1.37-cm- (0.540-in-) diameter Mo sheath. All sheathed element 

tests were performed with the assembly housed inside a water-cooled calorimeter. This calorimeter, 

shown in figure 34, was a basic shell-in-tube design having a single flow path. It was not designed to 

achieve a specified sheath temperature. Development of a highly designed calorimeter for application to thermal simulator testing is discussed in section 3.2, but testing of this calorimeter will not 

be covered in this Technical Memorandum (TM). 

Heater Power Leads

Sheathed 

Heater Element

Water Flow

Calorimeter

Water Inlet

Water Outlet

 Figure 34. Test setup with sheathed heater element fitted 

inside a water-cooled calorimeter.

2.5.2.1 Test 1: 0.25-mm (0.010-in) Tungsten Three-Wire Braid With Power Ramping, 

Sheathed, Helium Environment. The initial objective of the 0.25-mm (0.010-in) three-wire braid test 

was to achieve 100 hr of testing at 2.4 kW. Because the test configuration differed from previous 

tests, it was desirable to first study W element operation at lower power levels before stepping up 
65

to the ultimate goal of 6 kW. In addition, the total power that can be delivered to the element using 

a single power supply is limited by the element resistance and the maximum voltage that could be 

delivered (150-V, 100-A supply). A total of ≈108 hr were acquired at 2.4 kW over 10 thermal cycles, 

with the ramp to 2.4 kW taking ≈1 hr on each test day. (Test data are provided in fig. 35). Testing 

was performed in an ≈6 kPa (≈45 torr) He environment. To ensure high-purity He in the test chamber, the chamber was pumped to vacuum (≈1.3 mPa (≈10–5 torr)) and backfilled with He three times 

prior to the start of test. This procedure significantly reduces the potential contamination in the 

test chamber. The measured cold resistance of the element, taken prior to the start of each test run, 

was consistent at ≈1.21 Ω and measured resistance at 2.4 kW was also consistent at ≈9.2 Ω throughout testing.

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Heater Power

Resistance (ohms)

Power (W)

(a) Elapsed Time (hr)

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Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(b) Elapsed Time (hr)

 Figure 35. Initial and final test data for 0.25-mm (0.010-in) three-wire 

braided W element: (a) Initial test data, July 25, 2005, and 

(b) final 13 hr of test, August 8, 2005. Plots show total 

power delivered to the heater element and the calculated 

element resistance.
66

2.5.2.2 Test 2: 0.25-mm (0.010-in) Tungsten Five-Wire Braid With Power Ramping, 

Sheathed, Helium Environment. Upon successful operation of the three-wire braid for more than 

100 hr at 2.4 kW, a 0.25-mm (0.010-in) five-wire W braid was assembled for test at 3.6 kW in the 

same test environment (Mo sheath, 6 kPa (45 torr) He, water-cooled calorimeter). The additional 

two wires in the braid configuration dropped the element resistance at temperature from 9.2 to 

6 Ω, increasing the maximum power that could be delivered to the element by the 150-V supply; 

at 6 Ω, the 15 kW supply could provide up to 3.75 kW to the element. Testing from August 12 to 

September 2, 2005, recorded more that 100 hr of test time at 3.6 kW. After completing the initial 

requirement of operation for a minimum of 100 hr, this element was used to assist in debugging 

the test automation software, allowing additional test time to be accumulated. In total, this element was allowed to reach over 300 hr of total test time, accumulating over 40 thermal cycles, 

without experiencing element failure. This is a significant improvement over previous tests, indicating the effect of active heat removal (inert gas environment and operation in a water-cooled 

calorimeter) to achieve operating temperatures that are more representative of a heater element 

operated in a reactor core array. The longest single run, using the automated test operation and 

data acquisition system, was initiated on August 25, 2005, and ran for 95 hr, accumulating a total 

of 10 thermal cycles on the heater element. Data for this test run are shown in figure 36. The final 

element shutdown is not shown in this plot due to a limitation in the total number of data points 

that can be displayed by Microsoft Excel in a single plot. The plot indicates that the element resistance remained relatively stable at ≈6 Ω throughout the duration of the cyclic test that was consistent with the calculated resistance at power on other test days as well. 

7

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0:00:00 2:24:00 4:48:00 7:12:00 9:36:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

Elapsed Time (hr)

Figure 36. Test data for 0.25-mm (0.010-in) five-wire W braid (August 25–29, 2005).
67

2.5.2.3 Test 3: 0.38-mm (0.015-in) Tungsten Four-Wire Braid With Power Ramping, Sheathed, 

Helium Environment. An additional W heater element was assembled for testing at 5.0 kW using 

0.38-mm- (0.015-in-) diameter, four-wire braid. Braiding the 0.38-mm-diameter W wire was difficult. Although no problems were encountered with the 0.25-mm or 0.5-mm wire, the 0.38-mm 

wire experienced breakage at random lengths during the braiding process. There was no evidence 

of problems with the braiding machine, suggesting potential variability in the wire manufacturing 

process. The manufacturer was questioned as to the potential fabrication differences that might 

exist between the different wire diameters, but no response was received. A total of 3.7 m (12 ft) 

of braided wire is required to wrap the 1.17-cm (0.460-in) mandrel. This length was achieved after 

several failed attempts and the assembled heater element was installed for testing on September 9, 

2005. All testing was conducted with the element installed inside a Mo sheath and with the entire 

assembly fitted inside the water-cooled calorimeter shown in figure 34. The chamber was again 

backfilled to 6 kPa (45 torr) with high-purity He gas. The heater element burned out after only 4 hr 

of testing. The control system detected an over-current condition, and the power to the heater was 

automatically shut down before the circuit opened, per specifications set in the automated control 

system. 

A second 0.38-mm (0.015-in) four-wire braid element was assembled and installed for testing 

on September 13, 2005. Initially operated at 4 kW for ≈16 hr, the heater power was then increased to 

5 kW in an attempt to reach over 100 hr at the elevated power level. This element failed during its 

fifth thermal cycle after operating for 36 hr at 5 kW. Corresponding data are shown in figure 37.

10

9

8

7

6

5

4

3

2

1

0

6,000

5,000

4,000

3,000

2,000

1,000

0

0:00:00 12:00:00 24:00:00 36:00:00 48:00:00 60:00:00 72:00:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

Elapsed Time (hr)

Figure 37. Test data for 0.25-mm (0.010-in) five-wire braid (test 1, September 13–16, 2005).
68

A third 0.38-mm (0.015-in) four-wire braided W element was tested during September 16–26, 

2005) (figs. 38 and 39). This element was operated continuously throughout this time period, utilizing the automated control and data acquisition system. The element was first allowed to operate 

for more than 130 hr at 5 kW (14 thermal cycles) and for an additional 100 hr at 5.4 kW (11 thermal 

cycles). Hence, despite the initial difficulty noted with regard to wire breakage and failure of the 

first two 0.38-mm four-wire braid W elements, this configuration was operated well in excess of the 

minimum required test time of 100 hr at power levels at or above 5 kW. It is possible that the previous two 0.38-mm wire elements had microbreaks in the braided wire that resulted in early element 

failure. Additionally, the refractory wire elements were hand wrapped on the mandrel. Varying 

amounts of tension applied during the wrapping process can significantly affect the stress placed on 

the wire and can affect how tightly the wire lies on the spiral-cut mandrel. The variability incurred 

in a manual process could affect the success or failure of the resulting element during testing. The 

reproducibility of the refractory material and the assembled heater elements must be assessed if 

these heaters are to be used in a full core array. Element reproducibility was very good for earlier 

tests of graphite rod elements manufactured to specifications set by MSFC (see sec. 2.2), suggesting 

that the use of graphite elements should be reassessed if materials compatibility issues do not exist 

with the final core design. 

10

9

8

7

6

5

4

3

2

1

0

6,000

5,000

4,000

3,000

2,000

1,000

0

0:00:00 24:00:00 48:00:00 72:00:00 96:00:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

Elapsed Time (hr)

Figure 38. Test data for 0.25-mm (0.010-in) five-wire braid (test 2, September 16–20, 2005).
69

5

4.5

4

3.5

3

2.5

2

1.5

1

0.5

0

6,000

5,000

4,000

3,000

2,000

1,000

0

0:00:00 12:00:00 24:00:00 36:00:00 48:00:00 60:00:00 72:00:00 84:00:00 96:00:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(a) Elapsed Time (hr)

5

4.5

4

3.5

3

2.5

2

1.5

1

0.5

0

6,000

5,000

4,000

3,000

2,000

1,000

0

72:00:00 84:00:00 96:00:00 108:00:00 120:00:00 132:00:00 144:00:00 156:00:00

Element Resistance

Heater Power

Resistance (ohms)

Power (W)

(b) Elapsed Time (hr)

 Figure 39. Test data for 0.25-mm (0.010-in) five-wire braid (test 3, 

September 16–26, 2005). Data were broken into two 

separate plots due to limitations in Microsoft Excel 

plotting features: (a) Test data, 0–84 hr and (b) test 

data, 84–140 hr.
70

2.6 Material Selection Criteria

Heater material compatibility issues are being investigated for a number of potential core, 

sheath, and heater element materials. Specific issues are contamination of the core materials from 

heater assemblies, coefficient of thermal expansion (CTE), and interfacial diffusion during long 

time exposures at high temperatures. Quick, cost-effective testing of subscale samples by inductive 

and resistive heating is planned for existing materials test chambers at the Propulsion Research 

Laboratory at MSFC. A matrix of potential core, heater (wire, mandrel, central rod, power connections, etc.), and sheath materials has been established as a baseline to prepare a comprehensive 

assessment of materials compatibility issues over time (table 14). If compatibility issues are not 

expected to arise over the planned lifetime of a heater element (hundreds of hours), these material 

combinations may preclude the need for developing a hermetic seal for the heater element (sealing the central element inside the sheath). In all cases, it is expected that a mechanical seal will 

be necessary to develop elements that can properly mimic the static and dynamic characteristics 

of a nuclear fuel element, allowing layering of materials in the overall buildup of the element, but 

removing the need for a hermetic seal may significantly reduce the cost of developing an advanced 

thermal simulator.

Development of a thermal simulator that more closely mimics the characteristics of a 

nuclear fuel element, taking into account various material compatibility issues, will be discussed 

further in section 3. 

Table 14. Suggested materials test matrix. (Work would entail an initial literature review 

to assess potential compatibility issues at certain test temperatures and durations, 

followed by testing of combinations of interest.)

 

Poco Graphite

C Fiber 

Braid Al2O3 BN W Ta Re Mo

Mo-47.5%Re

Astar

Ta-10%W

SiC

FS85

Potential Heater 

Element and Sheath 

Materials

Potential Core 

Materials
71

2.7 Initial Conclusions From Thermal Simulator Tests

After completion of the initial test series at 1,200 W, refractory materials recommended for 

continued development and testing included Re, Ta, and W. The Re material has not been received 

to date and was not included in the current round of testing at higher power levels. The supplier 

has missed delivery dates in the past and has not proven reliable in providing the necessary material. Therefore, due to the limited supply, expense, and available quality of the material, Re is not 

recommended for future use in thermal simulator development.

The results presented in this TM suggest that W is the most desirable material for thermal 

simulator assembly should the reactor design dictate refractory metal elements. The final W simulator assembly achieved well over the stated test requirements, operating for more than 100 hr 

at 5 kW and for an additional 100 hr at 5.4 kW without failure. Tantalum element tests failed to 

achieve the minimum requirements at high power levels. Although none of the tested elements 

could go from zero to full power (5–6 kW) in a single step, given either a vacuum environment 

(sec. 2.4.1) or an inert gas environment (sec. 2.4.2.1), power ramping over ≈1 hr was successful. 

Additional testing must be performed to determine the maximum power ramp rate from a cold 

condition that can be applied without element failure. Once at operating temperature, the power 

to the element can be turned to zero and reset to the full power level without failure, allowing 

thermal cycling during the extended test period. During the automated power cycling, the element 

temperature was not allowed to decrease between power cycles. 

It may also be appropriate to investigate hardening effects during wire braiding. Three 

different W wire diameters were investigated: 0.25, 0.38, and 0.5 mm (0.010, 0.015, and 0.020 in). 

Although each of these wire diameters was braided (using three, four, or five strands per braid), 

the 0.38-mm wire had problems with breakage during the braiding process. Braiding was attempted 

with two separate lots of each tested diameter, with the same results in each case. This variability 

between potential wire diameters should be further investigated if W is selected for use in core 

thermal simulators.

Aside from reproducibility of the material and fabricated elements, one potential problem 

noted with the use of W wire was the limited material ductility. It may be desirable to investigate 

the possibility of increasing the W ductility to improve fabricability of the heater elements. If a 

blend of materials (doping) is necessary to improve ductility, however, localized hot spots caused 

by uneven distribution of materials could become a concern. 

Previous testing indicated that graphite rod heater elements are significantly more robust 

than any of the tested refractory wire elements, allowing instantaneous power changes without 

affecting the integrity of the heater. Element variability was also avoided in earlier testing of 

graphite rod elements that were manufactured to specifications set by MSFC (see sec. 2.2). The 

use of graphite heater elements should be reassessed if materials compatibility issues do not exist 

with the final core design.
72

3. DEVELOPMENT OF SUPPORT HARDWARE

In early 2005, heater element testing was performed with a minimum of diagnostics on the 

test chamber or on the elements themselves. The current test chamber is equipped with a roughing 

pump and a turbopump to provide pump down to vacuum. Vacuum gauges were present on the 

chamber to make regular pressure measurements during earlier testing and development. Based 

on these earlier measurements, it is estimated that the current tests were conducted at a pressure on the order of 1.3 mPa (10–5 torr). Future testing will incorporate vacuum gauges and other 

pressure gauges to allow for specification of chamber conditions for testing either at vacuum or 

in a high-purity inert gas environment. To properly conduct thermal simulator testing in relevant 

environments, a significant amount of support hardware must be developed. Discussions with the 

Naval Reactors Prime Contractor Team (NRPCT) indicate that the desirable sheath temperatures 

could range from 1,100 to 1,700 K at power. Achieving specified sheath temperatures requires that 

the test environment be manipulated through the addition of gas to the test chamber (providing 

conductive heat transfer from the element) and by the use of a highly designed noncontact calorimeter to actively remove heat from the element under test.

Data acquisition and control systems were installed in late 2005 and activated on the thermal simulator test chamber. The system is capable of recording temperatures and pressures, and 

provides automated control for the heater power supply. All future tests will be conducted with 

the data acquisition system operational. Chamber pressure is determined using three pressure 

gauges to cover the range of pressures expected in the chamber: a capacitance manometer is used 

for pressures from 0.1 to 100 kPa (1 to 1,000 torr), a convector from 0.1 to 100 Pa (10–3 to 1 torr), 

and a lower range convector from 10–5 to 0.1 Pa (10–7 to 10–3 torr). The data acquisition system is 

currently set up to record eight temperatures, with wiring for type C TCs run up to the chamber 

feedthrough. These TCs can be connected to the heater assembly or to other structures in the test 

chamber, as desired for a given experiment. The control system is configured to remotely operate 

the heater power supply, with read back to record the actual power delivered to the assembly. The 

control system also has the capability to establish set points for voltage and current, and if either 

should exceed the predetermined set point, the power can be automatically shut down. To operate the test chamber in a round-the-clock fashion, additional cutoffs must be programmed (a flow 

switch would be required to monitor water flow through the test chamber walls because loss of 

water would require a power down of the experiment), but the currently installed data acquisition 

and control systems will provide for higher fidelity in the future thermal simulator tests.

To better assess heater performance, diagnostics should be incorporated on the heater 

assemblies to provide data that can be compared to thermal analysis results. To monitor overall 

operation of each heater assembly during initial testing (outside of the full test article), temperature 

measurements are required. One test of the graphite heater element (bare) included a type C TC 

placed at the center of the element between two pieces of synthetic sapphire. Some tests attempted 

to obtain a sheath temperature using a TC tack welded to the outer surface of the sheath. This connection, however, failed during the test, and temperature measurements were not accurate. It may 
73

be desirable to incorporate a noncontact method of temperature measurement. Previous tests used 

an infrared (IR) camera, which viewed the heater through one of the zinc selenide viewing windows, to estimate element temperature. Future testing may incorporate an IR camera, and/or an 

optical pyrometer could be positioned to view the test article through one of the viewing windows 

to obtain a temperature measurement at a specific location on the element.

TCs provide inexpensive temperature measurement, but they have a number of negative 

issues related to their use. First, they are a contact measurement technique requiring direct attachment to the surface of the heater assembly. This can be troublesome since they can 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 Ni foil, is typically 

required. The foil is bonded to the refractory metal surface and the TC is then bonded to the foil. 

This intermediate material could diffuse into the sheath material over long periods of testing, 

potentially resulting in undesirable and/or unknown effects. Type C TCs, which are available with 

a 26% Re content (the balance being W), may have sufficient Re to allow them to be spot welded 

directly to the sheath surface, but testing would be required. However, the use of spot welding in 

general is not an recommended, as it can result in pitting of the heater sheath surface and potentially embed or trap impurities. In addition, the TC leads act as fins, locally cooling the location 

where the measurement is taken, and the long-term degradation of the TC wires over time at high 

temperature also introduces unknowns. If TCs are employed, they will be used in pairs to provide 

backup at each measurement location. 

The two-band optical pyrometer is a remote unit-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, IR sensing device uses a ratio technique between the 

emissivity at two overlapping wavelength ranges to determine temperature rather than using the 

emissivity at just a single wavelength. The result is a temperature measurement that is somewhat 

independent of emissivity, target size/shape, and dust/contamination on the windows in the optical path. The noncontact technique simplifies setup (no internal TC wires) and eases operations 

involving loading and unloading of the heater elements. 

The range of wavelengths used by the two-band pyrometers typically falls between 0.70 and 

1.15 μm. (For example, the Ircon model M5-R uses the ranges of 0.75 to 1.05 and 1.0 to 1.1 μm). As 

such, the type of sight glass selected is an important factor in the unit’s performance. The suggested 

view port material type is quartz fused-silica (available from vacuum product vendors such as Kurt 

J. Lesker). Zinc selenide windows have a wavelength of 8 to 14 μm, appropriate for infrared thermometry used previously but not applicable in combination with the suggested 0.70–1.15 μm pyrometer, requiring that the existing viewing windows in the test chamber be changed if optical pyrometers 

are adopted. Figure 40 illustrates a general laboratory pyrometer setup used to assess the operation 

of a two-band unit to support another program. The heated element is a Mo-sheathed graphite 

heater operated in vacuum and equipped with several type K TCs. Figure 41 shows a closeup of the 

heater area being monitored where the TC is visible at the top, and spot welded to the Mo sheath 

using an intermediate Ni foil.
74

Two-color

Pyrometer

Sight Glass

Heater

Element

Figure 40. Color optical pyrometer setup.

Ni Foil

Used to

Attach the TC 

Type-K TC 

Al2 O3 Tube 

Over TC Leads

Mo Heater Tube

Figure 41. Closeup of Mo heater.

Heating tests were performed to examine trends between the pyrometer and TC rather than 

absolute values of temperature (due to time and equipment limitations). These ‘heat and hold’ tests 

were conducted in a vacuum environment (≈1.3 mPa (≈10–5 torr) or lower), and an emissivity slope 

of 1.06 was used for the pyrometer (typical value for Mo suggested by manufacturers). Figure 42 

shows temperature traces over ≈5 hr for both the pyrometer and TC (type K). There is a difference 

of ≈42 °C between the readings, but both readings track very well over the interval. The TC measurements are unadjusted for losses, which typically produce lower readings (fin effect, etc.). The 

stated accuracy of the pyrometer and TC are ±0.5 and ±2.0%, respectively. 
75

1,500

14,00

1,400

1,350

1,300

1,250

1,200

2,800 4,800 6,800 8,800 10,800 12,800 14,800 16,800 18,800 20,800

TC

Pyrometer Notes:

1. Vacuum low 10–5 torr range

2. Power constant at 880 W

3. Emissivity ratio maintained constant at 1.06

4. TC readings are direct from junction (no adjustment)

5. Clean Mo surface condition

6. Testing was focused on examining trends only (not absolutes)

∆T = 42 oC

Temperature (K)

Time (s)

Figure 42. Sample long-term pyrometer data.

An additional test examined the effect of oxidation on the temperature readings by injecting air into the vacuum chamber and monitoring the resulting trends. Figure 43 illustrates one 

of these cycles in which air was briefly introduced while the turbo pumps continued to operate 

(causing the vacuum level to return to ≈1.3 mPa (≈10–5 torr) fairly quickly after injection). The 

Mo tube was already slightly oxidized at the beginning of this test due to previous air exposure 

experiments. Very small amounts of air injection resulted in higher sheath temperatures, primarily due to better thermal conduction from the hot graphite heater to the sheath. The temperature 

peaked and then dropped off again as the gas was evacuated (typically to a lower temperature due 

to some oxidation). In the case of a large air injection, the sheath temperature dropped rapidly, 

stabilizing at a much lower temperature due to the increased oxidation. (The vacuum level rapidly 

returned to ≈1.3 mPa (≈10–5 torr) after injection). After the large air injection, the Mo tube was 

dark gray (when inspected posttest at room temperature). During these transients, the pyrometer 

and TC readings trended in the same directions at similar response rates, and as expected, there 

was a variation in their difference. However, these differences were within the accuracy of the 

diagnostics and were reasonable considering no calibration or corrections were made for thermal 

losses, etc.
76

1,500

1,400

1,300

1,200

1,100

1,000

900

800

10,000 10,500 11,000 11,500 12,000 12,500 13,000

TC

Pyrometer Notes:

1. Vacuum low 10–5 torr range (except when air was briefly introduced)

2. Power constant at 875 W

3. Emissivity ratio maintained constant at 1.06

4. TC readings are direct from junction (no adjustment)

5. Mo surface was degraded significantly by oxidation when air was injected

 (initially the surface was slightly oxidized at the end surface very dark gray)

6. Testing was focused on examining trends only (not absolutes)

∆T = 44 oC

∆T = 15 oC

Power Increased 

to 1,380 W

Large Amount of Air

Injected Into Vacuum

Chamber

Very Small Amount 

of Air Injected Into 

Vacuum Chamber

∆T = 2 oC

Temperature (K)

Time (s)

Figure 43. Sample data with oxidation.

To improve the absolute temperature measurement for a two-band optical pyrometer the 

emissivity slope for the material of interest must be determined. These ratios are available in 

literature for common materials, but they must be determined experimentally for special cases or 

applications. One possible method for determining this parameter is to fabricate a small blackbody cavity of the material to be tested. The cavity could be a piece of 2.5-cm- (1-in-) diameter 

tubing (of the correct material) with sealed ends and a small, 0.64-cm- (0.25-in-) diameter opening 

at one end. The internal surface finish should have a roughness of >0.25 μm to produce a diffuse 

emission. The tube should be equipped with a TC (with a lead wire length of at least 15 diameters 

inside the heated cavity to minimize loss effects) and an external heater. This assembly should be 

fully insulated (with the 0.64-cm hole exposed) and placed in a small vacuum chamber with a sight 

glass. With the pyrometer looking into the blackbody and focused in the 0.64-cm hole, the emissivity ratio is unity and the pyrometer and TC temperature readings should read essentially the 

same (within their accuracy bands). Shifting the pyrometer focal point so that it is on the outer 

surface of the tube next to the hole (a graybody) results in a temperature shift. The pyrometer’s 

emissivity ratio is adjusted in this configuration until its reported temperature agrees with that 

found during the blackbody measurement. This approach assumes that the tube is isothermal 

(with no drastic temperature drop across the thickness of the tube, a good assumption, if care is 

taken with the layout of the heaters and overall geometry).

A typical two-band system (such as the Ircon series 5R model 1410 with a temperature 

range of 600 to 1,400 °C) retails for $3,100 (estimate received in 2005). This device is equipped with 

lens options that allow for measurement spot sizes as small as 0.3-cm- (0.12-in-) diameter at a distance of 30 cm (12 in) and sighting of the measurement location is done through the lens for easy 

alignment.
77

The current recommendation is to incorporate both a limited number of TCs and a 

two-band optical pyrometer to measure the sheath surface temperature during testing. This will 

reduce the overall effort involved in setting up, attaching, and maintaining TCs on the assembly 

sheath and will offer duplication of temperature measurements as the pyrometer is being tested. 

In addition, the two-band system will be completely external to the chamber, allowing for rapid 

replacement or troubleshooting should it be required. TCs will also be incorporated to measure 

environmental conditions on the support stand and the test chamber.

In summary, a significant amount of support hardware must be developed to properly conduct thermal simulator testing in relevant environments and at characteristic temperatures prior 

to full core or partial core array testing. 

3.1 Gas Purification System

The gas purification system adopted for use with the thermal simulator test chamber was 

originally designed for use with lifetime testing of refractory metal HPs that utilize a LM working 

fluid. Therefore, it was designed to achieve very high gas purity (<1 ppb O2 contamination) over 

extended test times. The system incorporates two independent gas purifiers to accomplish this 

goal: a once-through SAES Pure Gas, Inc. MicroTorr® purifier that is used to initially cleanse the 

gas from the supply bottles and a recirculating SAES MonoTorr® purifier that is used to continually filter contaminants from the chamber fill gas as it is recycled through the system during testing. The gas purification system is designed to operate with any desired mixture of He and Ar 

gas (set by the partial pressure of each gas), which can be premixed from independent K-bottles. 

The premixed gas is contained in a mixed gas bottle and once the recirculating gas system is 

brought up to the desired gas pressure, the supply bottle is isolated and the mixed gas is continuously cycled through the test chamber and gas purification system. The design of the gas mixture 

and purification system, shown in figure 44, is such that the recirculating gas system can be fully 

isolated from the gas mixing/fill system using the hand valve G-HOV9, allowing the gas mixture 

bottle to be charged with an appropriate He/Ar mixture prior to test chamber fill. 
78

Turbo Vacuum

Pump and ISO Valve

Test Chamber

With Heat Pipes

1/2-in Tubing

Recirculating

Flow Path

G-RV4

Set 20 psia

0-1,000

Torr G-HOV10 Gas G-HOV9

Pump

G

 P

G-TC1

G-HOV7

G-HOV8

G-HOV6

G-HOV15

G-HOV14

G-HOV16

G-HOV13

G-HOV12 RGA

Identifier Key:

HOV: Hand Operated Valve

HOR: Hand Operated Regulator

RV: Pressure Relief Valve

RGA: Residual Gas Analyzer

TC: Thermocouple

FM: Flowmeter

G-HOV11

G-RM1 G-TC2

TC FM

G-HOV4

G-HOV5

G-HOV17

G-HOV2

G-HOV3

0-150 PSIG

1/4-in Tubing

Bleed Valves

Bypass 

Valve

MicroTorr

Purifier

Ar 

Supply

K-bottle

He 

Supply

K-bottle

Gas Mixture

Bottle

MonoTorr

Purifier

Evacuation

Port

Duplicate Fill

System for Small

Test Chamber

G-RV1

Set 120 psig

Set 120 psig

G-RV2

Set 120 psig

Set 120 psia G-RV3

G-RV5

0–150 psi

G-HOV1 G-HOR1

G-HOR2

0–4,000 psi

0–150 psi 0–4,000 psi

Figure 44. Test chamber gas mixture and purification system.

The MicroTorr Purifier, which has an advertised performance for purifying both He and 

Ar to a final O2 concentration of 1 ppb, is employed on the inlet flow line. With a vacuum chamber 

initially filled with air, the O2 concentration in the test chamber can be significantly reduced over 

multiple gas dilution cycles with high purity He or Ar. After only two dilution cycles starting from 

atmospheric pressure, the final O2 impurity concentration is just 0.1 ppb at a 10 kPa (76 torr) test 

chamber pressure. This condition meets the acceptable O2 concentration (≈0.28 ppb) for testing 

Mo-Re alloys (per the design specification set for lifetime HP testing) and is well within an acceptable range for testing prospective thermal simulators. 

Within the MonoTorr purifier, getter materials are used to irreversibly trap gaseous impurity 

molecules that are captured on the surface of the materials and, upon heating, diffuse into the bulk 

of the getter. Note that if the getters inside the MonoTorr are inadvertently exposed to air, they will 

be immediately filled, and the unit will require extensive overhaul before it can be applied again in 

the purification loop. The internal heater within the MonoTorr device has an operating temperature of 400 °C and, as a result, gas exiting the MonoTorr will have an elevated temperature up to as 

much as 400 °C, which may require that the lines be cooled by water circulation downstream of the 

MonoTorr. For this reason, the gas pump is located upstream of the MonoTorr purifier to prevent 

overheating of the pump internals. When used with Ar or He, the MonoTorr purifier can remove 

molecules of water (H2O), O2, hydrogen gas (H2), carbon monoxide (CO), CO2, nitrogen gas (N2), 

and methane (CH4). The performance of the purifier is dependent on the pumping speed at which 

it is operated. Table 15 provides a summary of the performance guarantee for the SAES MonoTorr 

Phase II 3000 for rare gases. 
79

Table 15. SAES MonoTorr phase II 3000 performance guarantee for rare gases.

Impurity Pumping Rate

0–20 slpm (ppb) 20–50 slpm (ppb)

O2 < 1 < 1

H2O < 1 < 1

CO < 1 < 1

CO2 < 1 < 1

N2 < 1 < 10

H2 < 1 < 10

CH4 < 1 < 10

When the gas purification system is incorporated in thermal simulator testing, the test 

procedure will include an initial bakeout of the system (under vacuum), which will assist in driving out volatile impurities (primarily water, if the system is clean and degreased) from the test 

components and gas lines to reduce additional impurities in the system. An endothermic process, 

desorption (or outgassing) is accelerated by increased temperature. The rate at which gas appears 

to emanate from a surface is referred to as the outgassing constant and this value is usually given 

in torr-L/s-cm2. It is advisable to begin a degassing program by first pumping down the system 

at room temperature to remove physically adsorbed water before commencing the baking cycle. 

If heat is applied at atmospheric pressure, it could result in activated chemisorption of physically 

adsorbed gas, which would require a prolonged heating cycle to remove. 

Residual gases in a vacuum chamber derive from the original gas content of the chamber, 

gas emission from the chamber walls, or hardware in the chamber, or leaks from the outside of the 

chamber. To determine the actual gas content in the test chamber and to assess the effectiveness 

of the bakeout and gas purification processes, a residual gas analyzer (RGA) should be employed 

to verify the partial pressure of various gaseous components in the test chamber throughout test. 

A DYCOR® DymaxionTM Mass Spectrometer (an Ametek Process Instruments product), model 

DM100M) will be employed in the test system. The DYCOR mass spectrometer utilizes a quadrupole mass analyzer. A hot filament in the RGA is used to create electrons of a suitable energy 

that then generate a stream of ions, where the rate depends on the pressure, temperature, and 

species of the individual molecules. The stream of ions is then electrostatically focused toward 

the mass filter. A quadrupole mass filter consists of four metal rods having a time-varying electrical voltage applied, selected to only allow ions of a particular mass to enter along the axis and to 

pass through to the opposite end. After passing through the mass filter, ions are focused toward a 

Faraday cup and the current is measured using a highly sensitive ammeter. The resulting signal is 

proportional to the partial pressure of the ion species that was passed by the mass filter. Because 

the current produced by residual gases is very small, a sensitive signal amplifier is also required 

in the system to detect very small partial pressures of the various gases that may be present. The 

DYCOR DM100M can be used to detect masses up to 100 AMU at a minimum detectable partial 

pressure of 6.7× 10–10 Pa (5× 10–12 torr).
80

The gas purification system has been installed in the MSFC thermal simulator test laboratory and is connected to the thermal simulator test chamber, as shown in figure 45. This system 

is intended to purify the gas (reduce the potential for contaminates) contained within the test 

chamber. The test chamber gas thermally couples the calorimeter and heater element by means 

of a static gas gap to achieve the required power/temperature balance. A complete leak check was 

conducted and all leaks on the low pressure side of the system (recirculating section) were located 

and repaired. 

Gas Purification System Test Chamber

Gas Feed 

Lines

Gas Bottles

Control

Valves

MonoTorr

Element

Figure 45. Inert gas purification system connected to test chamber.

Flow, temperature, and pressure measurement instrumentation on the gas purification rack 

have been tied into a data acquisition system and initial checkout tests have been performed to 

evaluate gas loading and flow performance of the system. In an effort to conserve its capacity, the 

MonoTorr purifier element has not been activated in tests performed to date. It will be employed at 

a later date after the RGA has been integrated with the system and all gas purification system lines 

have been bakedout to minimize potential contaminants in the system. Initial checkouts were performed with air and pure He to determine flow performance in the gas purification system. Results 

of air flow tests at a pressure of 1 atm indicate a volumetric flow rate of ≈20 L/min, in agreement with 

manufacturer reported capability. As operating pressure is reduced, the volumetric flow (as reported 

by the thermal mass flowmeter) remains approximately constant, as expected, since the pump is a 

diaphragm design with a constant volume displacement. However, below 20 kPa (150 torr), the flowmeter indicates substantial drops in the volumetric flow, reporting ≈9 L/min at 6.7 kPa (50 torr). It is 

currently believed that the actual volumetric flow rate is approximately constant even at the reduced 

operating pressure and the reduction in the reported flow rate could be attributable to error in the 

thermal flowmeter since it is being operated well below its designed operating and calibrated pressure range. Note that accurate flow measurements are not important for this application since the 

flowmeter is only intended to provide an indication of flow in the purification loop, much like a 

flow switch. 
81

For the baseline He case at a test chamber pressure of 13 kPa (100 torr), the gas inventory 

was ≈0.012 kg. This gas was cycled throughout the purifier at a rate of 0.023 kg/hr, resulting in 

about two gas exchanges per hour. To monitor the gas conditioning process, the DYCOR RGA 

was set up to sample the gas purification loop and thermal simulator chamber. The RGA was 

successfully operated after resolving some initial communication problems between the analyzer 

and the control computer. The RGA was used to sample the recirculation line with He present in 

the system and the unit was used to display the partial pressure of each species in the system. At 

this time, however, the circulation system tubing has not been vacuum baked, resulting in a high 

background floor (as read on the RGA). Heater tape will be applied to all tubing and components 

on the gas purification loop and the system bakedout for a minimum of 24 hr prior to use with the 

thermal simulator test chamber and prior to incorporation of the MonoTorr purifier (testing to 

date has used the MonoTorr bypass flow path). The SAES MonoTorr purifier, which will be used 

to scrub the recirculation flow, has been operated briefly to ensure proper operation of the complete purification loop. Initial results indicate that it has a higher than anticipated pressure drop, 

resulting in a lower flow rate which corresponds to an increase in the time required to fully circulate the gas inventory in the system, but this is not expected to adversely affect the proper operation of the test chamber. 

Initial operational procedures have been laid out and checkout operations of the gas purification system have been initiated, as previously discussed. Actual use of the purification system 

to support day-to-day operations will be held off until additional operating experience has been 

gained so that its implementation will not impact currently scheduled test activities. 

3.2 Water-Cooled Calorimeter

As discussed in the test results in section 2.5.2, a noncontact water-cooled calorimeter has 

been specifically designed to allow testing of individual thermal simulators in a relevant thermal 

environment. This calorimeter is designed to achieve a specific sheath temperature at a given 

power level, simulating the conditions that the thermal simulator assembly would see in the full 

core assembly. A series of thermal calculations was performed to initially size a calorimeter for 

anticipated simulator testing. Coupling of the calorimeter to the heater element is controlled by 

a static gas gap containing He, Ar, or a mixture of these gases. Water flow rates through the calorimeter are assumed to range from 0.4 to 7.6 L/min (0.1 to 2 gpm) for initial calculations. In shifting from pure He gas to a mixture of He and Ar, it is noted that the reduced thermal conductivity 

arising from the addition of Ar to the gas mix increases the sheath temperature by a few hundred 

degrees kelvin depending on the power levels and the mixture ratio. Desired boundary conditions 

for heater testing were obtained from NRPCT. The current state of the reactor designs indicate 

that the fuel pin OD temperature could range from 1,100 to 1,700 K, with power per pin of up 

to 6,000 W. However, it is anticipated that the final design will set the fuel OD temperature closer 

to 1,400 to 1,500 K. The calorimeter design attempts to meet this range of conditions, but it is 

likely that multiple calorimeters will be necessary to cover the full range of test parameters. 

The current calorimeter concept uses a series of Cu coils attached to a central Cu tube 

that completely surrounds the sheathed heater element. Operational conditions include holding 

a given power throughput in the heater (up to 6 kW) while maintaining a prototypic fuel pin surface 
82

temperature on the heater element (up to 1,700 K, as specified by the reactor design). The individual 

Cu coils that form the calorimeter assembly are positioned with a slight separation to provide access 

to the surface of the heater sheath for TC and optical pyrometer measurements.

A thermal and flow analysis was performed for the initial noncontact water calorimeter 

design. Key operating parameters included: sizing of the conduction gas gap width, cooling water 

flow rate, water temperature rise, water pressure drop, and bulk Cu operating temperature. The 

design heater boundary conditions used in this analysis include a power level of 6,000 W at temperatures of 1,700, 1,400, and 1,100 K. The basic engineering layout for a wrapped coil type configuration 

uses off-the-shelf Cu tubing and hardware components. Two methods were identified to meet the 

heater operating power and temperature conditions. First, each test boundary condition could be 

accomplished by fabricating three different calorimeter units with gas gap widths set to meet each 

of the operating temperature conditions. Alternately, a fixed gas gap geometry could be used and 

the desired sheath temperature could then be achieved by varying the composition of the conduction 

gas to adjust the gas conductivity. Figure 46 illustrates the basic calorimeter layout that includes six 

individual cooling coils separated by ≈0.64 cm (≈0.25 in) to allow for TC/optical pyrometer access. 

Four coils have 13.5 revolutions of Cu tubing while the remaining two have 6.5 revolutions. To test 

fabrication techniques, 4-in-long coil segment samples were wrapped on the central support tube 

and brazed in place (fig. 47). Initial trials using a furnace braze method resulted in good bonding of 

the 0.64-cm- (0.25-in-) diameter flow coils to the central tube and each other; however, the tri-cusp 

area between coils was not filled due to the limited amount of braze material which could be added. 

A hand-brazing technique was then employed (allowing more filler braze material to be added) and 

this technique resulted in good bonding between adjacent coil tubes and nearly filled all the tricusp 

cavities. Due to the high thermal conductivity of Cu and the close packing of the cooling coils, complete filling of the tricusp gaps is not necessary to meet the desired operating conditions.

Heater Element Calorimeter Design

Gas Gap Coupling (Heater to Calorimeter)

6,000–W Nominal Power Extraction (With Heater at 1,700 K)

Water Cooling Flow (Nominal 0.2 gpm)

Cu Construction

Coiling Coils (×6)

Coil Gap

(Instrumentation)

Water Outlet

Manifold

Water Inlet

Manifold

Central Tube

Mo Sheathed

Heater Element

Figure 46. Heater element calorimeter layout.
83

4-in Cooling Coil Segment

(Before Brazing)

Partial Coil Cross Section

After Brazing

Filled Tricusp Region

Between Tubes

Partial Void 

Tricusp

Calorimeter Coil Wrapping and Brazing Tests

Figure 47. Calorimeter coil fabrication tests.

After initial success with the hand-brazing technique, it was noted that the high temperatures required for this braze resulted in a reduction of the central tube strength, introducing 

difficulty in maintaining concentricity between the heater element and the calorimeter. Therefore, 

low-temperature hand-brazing and soldering tests were conducted to further perfect the technique 

of attaching the Cu coils to the central calorimeter tube. Due to difficulties encountered during 

repeated brazing attempts, the initial as-built configuration did not include braze or solder connecting the coils to the central tube. There may be sufficient contact between the coils and the tube 

by virtue of the wrapping procedure and this contact conductance may be complimented by the 

existence of a very thin, sufficiently conductive gas gap. Testing is necessary to fully assess these 

assumptions. Differential thermal expansion between the central tube and the water-filled coils 

might enhance contact pressure and the conductance between the coils and the primary tube. If 

this concept does not provide sufficient cooling during initial testing, additional braze applications 

will be attempted or other coupling techniques (such as thermal epoxy or thermal grease) will be 

evaluated. To complete the connection to the Cu coil assemblies, silver solder is used to connect 

the water supply/return headers.

 

Thermal analysis, performed in parallel to the assembly of the initial calorimeter concept, 

assumes some preliminary values for coil-to-central-tube contact and gas gap conductance. This 

analysis indicates that for a pure He gas gap between the coils and the primary tube, sufficient 

cooling is available to extract 6,000 W while maintaining the maximum sheath temperature at or 

near the 1,700 K target. The analysis further shows that if the gap contains pure Ar or vacuum 

conditions, excessive sheath and heater temperatures will ensue. Unless initial testing indicates 

that the preliminary assumptions in the model are too conservative, some means of directly coupling the coils to the central tube should be implemented to allow lower sheath temperatures to be 

achieved, or to provide any ‘tunability’ to reach alternative conditions. Analysis results also suggest that a thicker wall central tube with a smaller sheath/central tube gap would be helpful 
84

in meeting desired target ranges. The initial calorimeter thermal testing should help to quantify 

the thermal conductances and provide a basis for correlation of the calorimeter thermal model to 

the actual test hardware. More detailed discussion of the thermal model and results are provided 

in appendix J.

Spring-loaded TCs were built to allow monitoring of the heater element sheath temperature at 

five locations. Figure 48 illustrates the TC probe layout in which exposed bead TC probes will be used 

in the final assembly. The spring loaded TC allows the assembly to be fitted to the calorimeter and 

ensures that direct contact between the TC and sheath will be maintained. However, initial thermal 

analysis suggests that the TC thermal conduction error for this configuration can be very large, ranging from tens to hundreds of degrees kelvin. A detailed analysis of the potential error associated with 

spring-loaded TCs is provided in appendix K. Additional temperature measurements will be obtained 

at the same axial positions (rotated by 180o) using two-band optical pyrometers to better assess the 

element temperature. The calorimeter test chamber and the instrumented test configuration are 

shown in figures 49 and 50.

Assembled Type-K TC Spring-Loaded Probe

Mounted on Calorimeter TC Bracket Assembly

TC Probe Tip to be an Exposed Bead in Final Version

Figure 48. Spring-loaded TC unit.

The heater element will be centered in the calorimeter using thin Mo spacer rings. The 

selected approach uses two spacer rings, one placed at each end of the sheath in the unheated 

regions. The Mo heater sheath is not expected to sag significantly at the planned operating temperatures, so the need for a central spacer ring is not anticipated at this time, but could easily be 

added if necessary. The Mo spacer rings were fabricated from 0.25-mm- (0.010-in-) thick plate 

stock and machined to a knife edge along the contact perimeter in order to minimize heat transfer 

through the ring itself. An approximate 0.13-mm (0.005-in) clearance is provided on the inner and 

outer ring diameters to allow for easy assembly and to account for thermal expansion.
85

Vacuum System

Installed Heater

Calorimeter

Pyrometers (×5)

Heater Power Feeds

TCs

Figure 49. Calorimeter test chamber.

Heater 

Element

Water 

Coolant

Calorimeter 

Shell TCs

TC Probes (×5)

Coolant 

Flow TCs

Support TCs

Figure 50. Calorimeter instrumentation.
86

The thermal simulator test chamber with integrated calorimeter assembly has been set 

up and auxiliary support systems such as vacuum hardware, inert gas supply, and water-cooling 

components have been connected. To measure coolant water flow rate, a paddle wheel style flowmeter was installed. The test chamber shown in figure 49 is separate from the one currently used 

for refractory wire-wrapped heater testing discussed in section 2. Once calorimeter checkout tests 

are complete, the two independent thermal simulator test chambers will be used in heater checkout and lifetime testing. Figure 49 shows the current setup with the calorimeter and heater element 

installed. TC and optical pyrometer instrumentation have been attached to select calorimeter 

locations to monitor its operation, allowing comparison with thermal models. Figure 50 provides 

a closeup of the test configuration with a number of instrumentation locations identified. Before 

use in the current thermal simulator testing, the calorimeter will be fully checked out using proven 

heater elements to minimize any potential impact on the thermal simulator test program. For initial testing, a 46-cm- (18-in-) long graphite heater element, powered by a 15-kW dc power supply, 

has been installed. At the expected operating temperature, heater element resistance will be ≈0.6 Ω 

dissipating 6 kW, limited by the power supply maximum current output of 100 A (operating voltage would be 60 Vdc in this condition).

A successful low power checkout test of the power supply and data system was performed, 

bringing the heater to a temperature of 50 °C. (The test chamber was open to the air during these 

tests.) Following the checkout, the calorimeter assembly and test chamber were He leak checked. 

Results indicate that all vacuum chamber penetrations and connections are leak tight (as measured 

to 10–10 Std cm3/s He); however, leaks were noted on the calorimeter assembly, which includes several mechanical fittings and low-temperature braze joints. A water bubble test (with the calorimeter 

assembly pressurized) was used to isolate the visible leaks and they were repaired. The calorimeter 

was reinstalled in the test chamber and He leak checked to verify its integrity. The low-temperature 

braze joints on the Cu calorimeter fittings were suspect due to their close proximity to each other 

since the high thermal conductivity of Cu tends to significantly heat all the joints around the repair 

location. To address this issue should future repairs be required, an alternative to brazing is to perform a tungsten inert gas (TIG) weld at the joint locations. A sample joint (which was brazed) was 

TIG welded to test the welding process and to examine the interaction of the braze material with 

the weld. The operation was very successful, producing a final weld that looked good and passed 

a He leak test on the first try. Therefore, welding will be used to repair or manufacture new calorimeter components.

During a second checkout test, the test chamber was evacuated and the calorimeter 

assembly water cooling system with integrated flowmeter was temporarily connected to a facility potable water connection. Checkout testing indicated that the flow could be adjusted (using 

a throttling hand valve on the calorimeter exit) from 0.95 to 13.2 L/min (0.25 to 3.5 gal/min) at a 

service pressure of ≈0.4 MPa (≈60 psi). This range of flow rate is expected to be sufficient for the 

planned tests. However, the thermal model calculations, provided in full in appendix K, assume 

a maximum water flow rate of 19 L/min (5 gal/min) and a maximum service pressure of 0.55 MPa 

(80 psi). During the checkout testing, the input heater power was held at 500 W, resulting in a 

heater sheath temperature of ≈1,073 K (≈900 °C) and the Cu calorimeter temperature remained 

essentially uniform, ranging from 302 K (29 °C) at the water inlet to ≈308 K (≈35 °C) on the 
87

uncooled portion of the shell (midway between cooling coils). During this operation, the chamber 

vacuum level initially spiked into the 1.3 mPa (10–5 torr) range, followed by a slow fall into the 

≈0.5 mPa (≈mid 10–6 torr) range. This response is typical when baking out vacuum systems. These 

initial test results, performed in a vacuum environment, have not yet been used to correlate thermal model results in order to determine the actual contact and gap conductances. As shown in the 

results presented in appendix J, the thermal model predicts a maximum sheath temperature that 

is appreciably higher than that measured. Subsequent examination is warranted to understand the 

differences between the experimental and predicted results, but the lower measured temperature is 

promising with regard to achieving sheath temperatures in the range of interest. 

3.3 Data Acquisition System

A computer controlled user interface (data acquisition and control) has been developed for 

automated thermal simulator testing, incorporating multiple safety cuts to allow round-the-clock 

testing in the absence of support personnel. The control interface is comprised of five LabVIEW® 

(a National Instruments product) applications that include: main thermal simulator application, 

strip-chart application, multigauge server, gas flowmeter server, and data recorder.

The thermal simulator application, shown in figure 51, provides the following functionality:

• Manual control of the heaters, where the user specifies a current and voltage.

• Automatic power control, where the application automatically adjusts the voltage setpoint to 

achieve the user-specified power setting.

• Automatic control of heater power over time to a user-specified profile for a user-specified number of cycles, allowing the heaters to be cycled repeatedly over a long period of time without 

user intervention.

• Checks for out-of-limit conditions. Each limit violation is indicated on the display. If a limit 

violation occurs when cutoffs are enabled, the application automatically interrupts the power 

to the heater.

Out-of-limit conditions are established such that the power to the heater simulator is automatically shut down if any one of the limits is exceeded. Currently, the application can monitor 

the following:

• Up to 20 TCs for individually specified limits.

• Chamber pressure for a specified upper and lower limit (shutdown will also occur in the event 

of a loss of signal from the vacuum gauge server).

• Heater element power for a specified percent tolerance over the setpoint.

• Calculated element resistance for a specified percent tolerance relative to a specified resistance.
88

Figure 51. LabVIEW thermal simulator application.

• Mass flow rate of gas through the recirculating gas purification system below a specified limit 

(shutdown will also occur in the event of a loss of signal from the gas flowmeter server).

• Mass flow of cooling water through the walls of the thermal simulator test chamber and the 

heater calorimeter (shutdown will occur given a low or no flow condition).

The application monitors measurements for limit violations and updates display indicators 

appropriately. If cutoffs are enabled and an out-of-limit condition occurs, the power to the heater 

is immediately shut down and the limit indicators are latched. The indicators are unlatched when 

cutoffs are disabled by user intervention. Before enabling cutoffs, the user must verify that all 

limit indicators are off to prevent the cutoff from immediately tripping. Additionally, a signal light 

was installed to indicate the system status: when the system cutoffs are enabled and no cutoff has 

occurred, the light is green; during a cutoff condition, the light flashes red, and when cutoffs are 

disabled, the light is off.

The strip-chart application, shown in figure 52, plots up to 24 hours of pressure, TC, voltage, 

current, resistance, and power data. All data are recorded to a data file for post processing.
89

Figure 52. LabVIEW strip-chart application.

The multigauge server, shown in figure 53, acquires data from the vacuum gauges and uses 

the user datagram protocol (UDP) to distribute the data to other applications. The multigauge 

server handles the serial communication interface with the multigauge vacuum controller.

Figure 53. LabVIEW multigauge server.

The gas flowmeter server, shown in figure 54, acquires data from a thermal mass flowmeter 

and converts the reading from standard liters per minute (SLM) of N2 to the actual mass flow of 

a mixture of two gases and then uses the UDP to distribute the results to other applications. The 

conversion uses the mole fraction and the specific heat of each gas to compute the specific heat of 

the mixture. The ratio of the specific heat of N2 to that of the gas is then used to convert from the 

standard N2 reading to the actual reading for the mixture. The specific heat is necessary for the 

conversion because the meter is a thermal mass flowmeter and the sensor reading is proportional 

to the specific heat of the gas as well as the mass flow.
90

(a) (b)

Figure 54. Flowmeters: (a) LabVIEW gas flowmeter server and (b) installed thermal mass flowmeter.

The data recorder, shown in figure 55, acquires data from the data acquisition hardware, 

the multigauge server, the gas flowmeter server, and the thermal simulator applications. These 

data are then recorded to a tab delimited file, which can be loaded into a spreadsheet. The 

recorder supports various user-specified record rates and allows the user to specify comments 

to be inserted into the data.

Figure 55. LabVIEW data recorder.
91

3.4 Modified Power Interface

In response to the 2005 NRPCT downselect to a direct gas-cooled reactor design, the 

power interface to the electrically heated reactor module was reinvestigated. The original power 

interface used in testing the SAFE HP-cooled reactor and the DDG cooled reactor allowed heater 

power leads to extend from the base of the core in order for Cu wire to be connected for power 

input to the heaters. This configuration is shown in figure 56 for the DDG test series. To minimize 

impact on the gas flow through the inlet plenum in gas-cooled reactor designs, a conceptual design 

for a modified power interface was developed for implementation in the next phase of thermal simulator designs. This power interface will minimize the interference of the electrical system with the 

flow plena, better mimicking an actual reactor design. The current design assumes single-ended 

heater elements (power in/out at the same end), removing the need to account for axial thermal 

expansion from the power interface design. Additionally, the proposed concept assumes that the 

power interface is on the inlet flow plena, as in the original DDG design. This assumption allows 

the use of existing power feedthroughs in the DDG pressure vessel, but power feedthroughs could 

be moved to the exit flow plenum if it is desired in subsequent reactor prototypes. 

Connectors

Core

Thermal 

Simulators

Figure 56. Power interface for the DDG-cooled reactor tests.

Initial conceptual designs have been prepared for a 2.5-cm- (1.0-in-) diameter heater (outer 

sheath diameter, corresponding to the outer clad diameter), within the range of potential fuel 

pin sizes in the reactor designs under evaluation by NRPCT. The power interface incorporates a 

‘motherboard’ type design to bring power into the core with minimal disruption to the gas flow. 

The conceptual design appears to be feasible, but could become significantly more complicated 

with a larger number of fuel pins. The 37-pin DDG layout was used as a baseline for the initial 

design study. Figure 57 shows the conceptual design; detailed design drawings are included in 

appendix L.
92

Positive: 0.070-in Thickness

Negative: 0.070-in Thickness Isolator: 0.125-in Thickness

Figure 57. Modified power interface conceptual design, using 37-pin DDG as a baseline concept.

3.5 Component Cleaning and Chamber Preparation Procedures

The current test preparation procedures do not incorporate rigorous cleaning of test components, but some rough cleaning is performed. Prior to wrapping the Al2O3 mandrel with the 

desired wire, both the mandrel and the wire are cleaned with methanol and all components are 

handled with gloves. The chamber startup and operation procedures are as follows: 

(1) Initialize roughing pump on the vacuum chamber. Allow to run for ≈1 hr.

(2) Initialize turbopump—it will take ≈10 min to get to operational conditions. It will take 

several hours for the pressure to reduce to an acceptable level if pumping down from atmospheric 

conditions. Between testing days, the turbopump is turned off but the chamber is not opened (so 

that an acceptable pressure is reached in a relatively short time on subsequent test days). Note: 

The current chamber configuration does not incorporate any vacuum pressure gauges. Pressure is 

estimated to be ≈10–5 torr based on past experience with this pump and chamber.

(3) Turn on cooling water to cool the chamber walls.

(4) Turn on power supply. The heater elements operate in a voltage-limited mode, so the current on the supply is turned to its maximum 100 A and power is controlled by adjusting voltage.
93

(5) For initial testing of an element, apply 50–100 W and allow the element to bake out for 

a few hours. For subsequent test days, set the power supply to the desired power level (1,200 W).

(6) At the end of test day, turn off power supply and pumps. Allow the element temperature to drop below ≈200 °C before turning off chamber cooling. (Note: It will take ≈30 min for 

a bare element and ≈1 hr for a sheathed element to cool to the desired temperature.)

More extensive cleaning procedures will be adopted for higher fidelity element testing and 

prior to heater integration into each test article. 

Posttest materials analysis of the W and Re heater elements indicates that the lack of thorough cleaning procedures may have impacted test results due to the presence of various residues 

on the heater assembly. To minimize the possibility of contaminants affecting the test results, 

more extensive cleaning procedures should be adopted for all components that will be in the test 

chamber: mandrel, heater wire, sheath, chamber, and support structure. Extensive cleaning will 

also be required before heater assemblies are integrated in the reactor test article to prevent any 

contamination of the test article. Proposed cleaning procedures for various materials are included 

in appendix M. Each procedure includes a series of steps that use FreonTM (a DuPont product), 

chemical cleaning, and ethanol to remove any grease, oil, and other contaminants from the material. Vacuum firing is also employed to remove moisture (water vapor) from the material.

In addition to thorough cleaning, all components should be fully bakedout to remove all 

water vapor and other residual contaminants from the chamber and other structures. Tests will be 

conducted in a vacuum environment or in a low-pressure, high-purity inert gas environment per 

the test requirements and the boundary conditions that will exist when the heater assemblies are 

inserted into the full test article. The proposed procedure is derived from chamber preparation 

procedures outlined for HP life testing at the EFF-TF, with modification made for this specific 

application. Step 14 can be omitted if testing is to be performed in a vacuum environment. This 

procedure assumes that additional instrumentation has been incorporated in the test chamber. 

The procedure is as follows:

(1) Clean heater unit using the approved procedure (see app. M) and handle only with 

powder-free, alcohol-cleaned surgical gloves. 

(2) Wipe down the test chamber with Freon and alcohol to remove any potential grease, 

solvents, fingerprints, etc. Also perform a complete visual inspection to verify that no foreign 

materials are present. Allow the test chamber to dry for a minimum of 1 hr.

(3) Install cleaned heater assembly into the test chamber support bracket and secure.

(4) Close the test chamber and secure all flanges using appropriate tightening sequence.
94

(5) Evacuate test chamber and inert gas feed system lines to 0.1 mPa (10–6 torr) or lower 

and perform a He leak check to a sensitivity of 10–10 std cm3/s He on all flanges, feedthroughs, and 

view ports. Tighten any leaky fittings using the appropriate tightening sequence, exchanging any 

leaky view ports. Note: He leak checking is a fairly time-consuming process and it is not necessary 

to perform this check prior to every test. Leak checks will be performed on initial chamber checkout and will be repeated periodically (approximately monthly) to ensure that the chamber is still 

leak tight.

(6) Turn on the test chamber wall and inert gas feed system heater tapes and bring to 

a temperature of ≈200 °C (≈575 K) to outgas all trapped water vapor and other volatiles. 

(7) Measure the cold resistance of the heater assembly prior to heat-up. (This will be compared to the assembly resistance after outgassing, during test, and after test to assess stability of 

the heater element.) Turn on the heater element power supply and heat the element to ≈200 to 400 

°C (≈575 to 775 K) by applying a low-power level to the element (≈50–100 W). Hold at this heated 

condition for ≈24 hr or until pressure reaches the ≈0.01 mPa (≈mid 10–7 torr) range. Periodically 

monitor test chamber vacuum pressure. Note: Test chamber pressure should increase during heating and then begin to fall after a steady temperature has been maintained for 1 to 2 hr. If pressure 

continues to rise or does not drop, there may be a leak due to heating of the flanges. Perform a leak 

test while at temperature to locate the leak and tighten the offending flange.

(8) Verify that all data systems are operational and the temperature measurement system 

is properly set up (two-band optical pyrometer is focused on the desired heater assembly location 

and/or TCs are registering reasonable values). 

(9) Increase the power to the heater element to continue the outgassing process. Power 

can be increased using a linear ramp over time, up to approximately two-thirds of the desired pin 

power level (800 W for a 1,200 W assembly test, reaching ≈800 to 900 K). Note: The test chamber 

pressure will increase as additional outgassing occurs. As the temperature is increased, the chamber pressure will drop again when the element is no longer outgassing.

(10) Maintain the heater at a constant temperature for ≈2 hr and monitor the test chamber 

pressure. The pressure should begin to fall off after reaching a high value.

(11) Turn off power system and all tape heaters and monitor the test chamber vacuum level 

as the system cools. The pressure should drop approximately one decade after the entire system is 

cooled to room temperature.

(12) Isolate the inert gas system from the vacuum chamber and charge the system with the 

required inert gas mixture ratio. (Typical gasses will be He and/or Ar, with the desired percent Ar 

specified by weight.) This is in preparation for multiple dilution cycles of the test chamber to sweep 

the environment. A discussion of the effect of multiple dilution cycles using UHP gas on the level 

of contaminants in the test chamber is included in appendix O.
95

(13) Isolate the test chamber from the vacuum system and then introduce the inert gas 

mixture purge into the test chamber to a pressure of ≈9 kPa (≈70 torr). Isolate the test chamber 

and let stand for 2 min.

(14) Isolate the inert gas system and open the vacuum system to remove the inert gas 

mixture from the test chamber. Allow the vacuum to reach the 0.1 mPa (10–6 torr) range.

(15) Perform two more inert gas dilution cycles, and then bring the test chamber pressure 

to ≈9 kPa (≈70 torr). (Note: This step is omitted for vacuum testing.)
96

4. ADVANCED THERMAL SIMULATORS: CURRENT STATUS

Advanced simulator development includes efforts to embed instrumentation (TCs) within 

the thermal simulator structure and to develop simulators that better mimic the behavior of 

nuclear fuel elements. Discussions and initial investigations on advanced thermal simulator design 

and development have commenced. The current plan is to develop simple, advanced simulator 

models in-house to test techniques and methods for embedding instrumentation within the thermal simulator assembly and to assess fabrication ability with regard to matching key properties 

of nuclear fuel elements. Test articles will begin with graphite heater element cores. Rather than 

inserting the graphite element into a metal sheath with a gas fill in the heater/sheath gap, attempts 

will be made to fill it with an insulating powder to improve the conductivity across the thermal 

simulator. This technique was previously attempted by engineers that developed thermal simulators for testing at the Oak Ridge National Laboratory (ORNL) in the 1970s and 1980s.15,16 These 

experiments led to limited success due to the method in which the filler material was inserted 

between the element and sheath. In the ORNL heater elements, the insulating powder was poured 

into the gap, and tamped down periodically to remove voids, but no verification was performed 

to ensure that no voids existed. In addition to using the insulating powder layer, the future TCs 

will be embedded inside the sheath structure. External TCs and pyrometers will also be used to 

measure sheath temperature and to then assess the relationship between the internal and external 

measured temperatures.

In developing advanced simulators, the most important parameter to match is the dynamic 

response of the nuclear fuel element rather than the static characteristics, such as conductivity. 

It may be desirable to maximize the conductivity across the pin (radially), allowing significant 

capability to get heat out of the simulator and minimizing the temperature of the central element. 

The simulator design will then attempt to match dynamic characteristics at the outer radius of 

the simulator (at the sheath surface) to those at the outer radius of the nuclear fuel element (at the 

outer surface of the fuel clad structure). 

Initial attempts at developing advanced simulators will take advantage of materials available in-house. This will include use of graphite heater elements with a buildup of ceramic material 

around the element with TCs embedded during the buildup process. Characteristics of various 

powdered insulating materials will be assessed to determine the best option for simulator construction. Methods of applying the insulating material will also be investigated. One option is to 

simply fill the gap with powder, tamping down the powder periodically to remove any voids that 

might exist. However, this method limits the ability to predict the packing density in the heater/

sheath gap, which could lead to localized heating due to poor conductivity. Other possible techniques include VPS or chemical vapor deposition (CVD). If a plasma spray process is adopted, 

it may be possible to build up the heater element with the TCs already positioned in the heater/

sheath gap, but this remains undetermined. The potential of using a metal ‘shield’ between the 
97

heater and sheath to improve radial pin conductivity will also be considered as an option in heater 

element buildup. An analysis of the potential materials selection and layering will be completed, 

along with a materials compatibility study.

Thermal analysis associated with advanced thermal simulator development will entail 

an assessment of the axial heat flux at various layers of the proposed element to determine the 

ultimate external heat flux distribution at the outer sheath surface, both during steady state and 

dynamic conditions. This analysis will be run in parallel to initial element design and buildup and 

will feed into the future development of advanced simulators.

Initial advanced thermal simulator configurations can be evaluated using graphite heater 

elements. The samples would be fabricated using powder materials such as AlN, W, or Mo to fill 

the gap between the graphite heater and Mo sheath. However, caution should be taken when using 

W or Mo filler materials because, while both are thermally conductive, they are also electrically 

conductive and will require some type of insulator layer to prevent electrical contact between the 

heater and sheath. Alumina nitrate has a much higher thermal conductivity (140–200 W/m-K) 

than ceramics currently being used in simulator construction, such as Al2O3 (18 W/m-K) or BN 

(40 W/m-K). The thermal conductivity values decrease at temperatures above 1,273 K (1,000 °C) 

but still provide some improvement over Al2O3 or BN. 

Alumina nitrate is also a good electrical insulator. Similar materials have been used on 

VPS Mo and Mo-Re HPs that had an Al2O3/AlN composite material on the OD for electrical 

isolation. If necessary, a thin, dense AlN coating (applied by a technique such as sputtering, CVD, 

etc.) can first be deposited on the graphite to electrically isolate the heater. This approach might 

allow the use of a metallic powder or solid gap filler with higher thermal conductivity. Tungsten 

and Mo, for instance, both have decent thermal conductivity as compared to ceramics. A spray 

deposition approach may also allow for a milling process that could be used to minimize the gap 

between the heater and the sheath to very small clearances. By overapplying the outer isolating 

layer, then machining the OD to match the ID of the sheath (keeping in mind the tolerances), very 

small gaps might be achieved depending on the machining tolerances of the sheath and the heater. 

If this small gap can be filled with He, then heater to sheath thermal conductances could be significantly improved over existing larger gap configurations. On the other hand, the sheath material 

could be sprayed directly onto the outer layer of the isolator if differential thermal expansion of 

the materials does not present potential problems.

Advanced fabrication techniques are also being investigated to fabricate assemblies with 

embedded TCs. These could be embedded within the powder material between the graphite heater 

element and the sheath. This can be done using processes such as VPS, CVD, or electroforming to 

close out machined grooves in refractory metal sheath materials. Net shaped multimaterial components are also possible. 
98

Temperature measurements inside a heater element assembly could be accomplished using 

very fine TCs embedded in the insulation around the heater element, or fiber optic probes might 

be inserted along the length to obtain a complete description of the temperature in this region at 

multiple axial positions. Figure 58 shows the heater element placement inside the DDG core block, 

identifying the Al2O3 spacer region that could be used to embed TCs.

Power Leads Core Block

Annular Flow Gap

Alumina Spacer

Sheath/Clad

Graphite Heater

Figure 58. Heater element placement inside DDG core block and TCs would be embedded 

in the Al2O3 spacer region.

Fiber optic temperature probes developed by Luna Innovations, Inc. are constructed with 

Bragg gratings spaced evenly (standard fiber optic design allows measurement of temperature 

at 1-cm increments) along the sensing length to obtain detailed temperature profiles. These fiber 

optics currently have a temperature limitation of 1,000 °C for testing over extended times and 

may not survive over long test times at elevated temperatures in the DDG core. Luna fiber optic 

probes tested on the SAFE-100 demonstrated superior ability in generating three-dimensional 

temperature profiles of the core.17 Luna is currently working to extend the maximum operating 

temperature of the fiber optic sensors, which may qualify them for use as embedded temperature 

instrumentation in future advanced heater elements. Previous tests with these fiber optics were 

conducted by Luna, who brought the support hardware necessary to obtain and record data from 

the fiber optic probes to MSFC for temporary use. The EFF-TF is currently working to obtain an 

in-house capability to use fiber optic sensing techniques in future testing applications.

An initial series of tests was conducted to examine the possibility of embedding very fine 

(52 AWG) TCs between the heater and sheath. Because TCs are relatively inexpensive and are available in a wide range of sizes and configurations, it was feasible to test a proposed configuration in a 

relatively short period of time. A graphite heater element, fitted with three Al2O3 spacer rings, was 

used for the test. Six 0.254-mm- (0.010-in-) diameter, Inconel®- (a Special Metals Corporation product) sheathed type K TCs with an ungrounded junction were procured to assess their applicability 

in planned DDG testing. Graphite heater elements used in the DDG testing are 0.985 cm (0.375 in) 

in diameter and are single ended (designed so that both inlet and outlet power leads are on a single 

end). These heater elements fit inside a 1.4-cm- (0.55-in-) diameter sheath, the OD of which forms 

the annular flow path. The heater is separated from the sheath using three Al2O3 spacer rings to 

prevent contact between the graphite element and the metal sheath. The junction end of a TC was 

bonded to the graphite element at each Al2O3 ring using Al2O3 cement to obtain three temperature
99

Graphite Heater

Established TCs

TC Lead Wires

Heater Power Leads

Figure 59. Thermocouple location on graphite heater element for preliminary testing.

measurements along each element, as shown in figure 59. During testing, several problems were 

noted with this configuration. Because the TC sheath was made of an electrically conductive material (Inconel), bonding the sheath to the heater in more than one location set up a parallel current 

path in the Inconel sheath. The current required to operate the heater element caused the sheath 

to vaporize and the 52-gauge TC wires did not survive. Future investigation will consider alternate 

methods of embedding TCs within the heater elements to obtain approximate ‘fuel’ temperature 

measurements (heater temperature measurement that can be directly correlated to the expected fuel 

temperature at comparable conditions). 
100

5. FUTURE TESTING AND DEVELOPMENT PLANS

Future thermal simulator work will extend testing on the current materials and assembly 

designs, incorporating improved procedures to ensure that boundary conditions on the individual 

heater assembly tests will match those experienced when the heater assembly is incorporated in a 

reactor test article. These improvements include: enhancement of engineering in the incorporation 

of diagnostics (temperature and pressure) on the heater assembly and test chamber, improvement 

of the automated control system, refinement of cleaning and chamber preparation procedures, 

conduction of tests in both vacuum and high purity inert gas environments, incorporation of a 

heat removal method to better mimic the assembly boundary conditions in the full test article 

(using a calorimeter assembly around the heater assembly), and development of advanced thermal 

simulator designs that will seek to better match key properties of nuclear fuel elements and that 

will incorporate advanced instrumentation inside the heater assembly.

5.1 Element Cooling Techniques

Early testing was conducted at vacuum, to ensure that the only method of heat removal 

from the heater element was by thermal radiation to the cool chamber walls. This condition, 

however, is not typical of the boundary conditions that would be present for a heater assembly 

inserted into a reactor test article. Although each potential reactor design will have slightly different boundary conditions, attempts will be made to adopt a test configuration that can be made to 

simulate the conditions that would exist if the heater assembly were located in the core block test 

article. Testing at vacuum represents the most extreme condition that might be experienced by the 

heater. In any of the reactor designs, the heater elements would be actively cooled by the reactor 

coolant system (LM HP, pumped gas, or pumped LM), more efficiently removing heat from the 

element than by radiative cooling alone. A sketch of a single heater assembly (heater element plus 

sheath) as it fits into the full test article (represented by the core block) is shown in figure 60.

GHe

Clad/Liner

(Not Included in all Designs)

Core Block

Sheath

Heater Element

(Graphite Rod or

Spiral-Wrapped Mandrel)

GHe, 

or BN Powder

Gas or LM

Figure 60. Generalized simulator assembly with core block: cross-sectional view (not to scale).
101

In the electrically heated configuration, the heater element and the sheath—the full thermal 

simulator assembly—are representative of the nuclear fuel element. This sheathed assembly would 

slide into (a) the clad/liner (bonded) for the gas-cooled reactor or LM-cooled reactor or (b) the 

core block itself for the HP-cooled reactor design. In each case, the boundary condition that will be 

held in the separate heater assembly tests is the temperature of the inner surface of the core block—

whether there is a clad structure between this and the heater assembly (sheath) or not. The currently 

assumed design point temperature for the HP core block will be ≈1,325 K and for the gas-cooled 

reactor it will be ≈l,225 K. These boundary conditions can be held in the single heater assembly 

tests by using an adjustable heat removal device, such as a calorimeter. Calculations required to 

design the calorimeter will initially assume that the sheath/clad or the sheath/block gap will be 

≈0.25 mm (≈0.010 in) (radial gap). 

For the individual heater assembly tests to examine heater checkout and characterization, 

the test configuration may or may not include the clad structure (using materials equivalent to 

what will be in the complete test article) in order to get the correct boundary conditions for the 

test. If no clad structure is incorporated, the calorimeter design could allow for operation at the 

temperature of the inner surface of the clad, calculated from the core block temperature and conduction through the clad/block gap and the clad/liner itself (in designs that include a clad in the 

full reactor core). This would be simpler for the single element tests, and should allow for testing 

at prototypic boundary conditions. Additional testing has been conducted in a He environment 

of 6 kPa (47 torr) and also with incorporation of a calorimeter, as noted in section 3.2. 

5.2 Advanced Thermal Simulator Development

Development of more advanced thermal simulators will focus on designing a heater element that more closely mimics the characteristics of a nuclear fuel element, both at steady state and 

during transient conditions, to allow improved capability in testing and analysis of a fission reactor using electrically heated elements to mimic nuclear fuel rods. This will take into account both 

materials selection and heater element design in an attempt to match radial pin conductivity, element 

surface temperatures, and the time-dependent temperature response of nuclear fuel. Although it is 

unlikely that a single design will meet each of these characteristics simultaneously, each stage in the 

development of advanced heater elements keeps these characteristics as a focal point. 

Current thermal simulator development work is directed toward meeting requirements 

established for power per pin, axial flux profile, maximum assembly diameter, materials compatibility, and operational lifetime. Proposed advanced thermal simulator development will look to 

achieve these basic requirements while attempting to match significant thermal properties to that 

of an actual fuel element. Properties that will be considered include thermal conductivity, heat-up 

rate (which can be expressed as the product of density and specific heat (ρCp)), thermal diffusivity, and thermal expansion characteristics. The design will also include advanced instrumentation 

housed within the thermal simulator for analysis of the axial temperature distribution during test. 

Criteria that should be established prior to significant investment in thermal simulator and internal temperature sensor development are summarized in table 16.
102

Table 16. Desired advanced thermal simulator characteristics.

Total power per pin ≈1 – 6 kW

Linear heat rate ≈100 W/cm (dependent on selected reactor design)

Pin power density ≈100 W/cm3

(assumes 15-yr operation, 6% burnup, 1 MW-d/g energy from fission)

Axial power profile Peaking factor per reactor specifications, ≈1.33 used in graphite rod designs 

Cosine distribution, per reactor specifications

Pin dimensions 

 Diameter 

 Length

Target range: 

 ≈0.65-cm (≈0.255-in) minimum to 2.5-cm (1-in) maximum (selected based on 

 reactor design) 

 ≈50 cm nominal

Operating temperature 

 Centerline 

 Outer surface (sheath OD)

≈1,700 K nominal, reactor fuel dependent 

≈1,400 K nominal, reactor design dependent

Operational environment 

 Vacuum/pressure 

 Fill gas

Testing to include: 

 Operation in 10–5 Pa (10–7 torr) vacuum 

 Operation in ≈10 kPa (≈70 torr) He or Ar environment (UHP, O2 < 1 ppm)

Operational lifetime (for cyclic testing) ≈ 10,000 hr

Basic geometry 

 Heater element/sheath 

 Heater element/sheath/clad

Single-ended design

Clad structure dependent on reactor design

Effective thermal conductivity (centerline to inner clad diameter) Thermal simulator will attempt to match—value dependent on selected reactor fuel

Specific heat (as a function of temperature)

Density (as a function of temperature)

Effective ρCp (heater element + sheath)

Thermal expansion characteristics (effective)

The proposed advanced thermal simulator work will incorporate state-of-the-art techniques to fabricate the thermal simulator, incorporating advanced temperature sensors in the 

buildup process that will allow simultaneous determination of temperature at multiple axial positions. If desired, temperature measurements will be considered at both the centerline and the outer 

radius (inner clad wall) of the simulator. Inclusion of internal diagnostics becomes increasingly 

difficult as the size of the heater assembly is further reduced, and internal diagnostics may not be 

possible for all heater assembly configurations.

5.2.1 Plasma Spray Process

This effort will examine the use of the VPS process to build up a thermal simulator. This 

process begins with a mandrel and builds the simulator using multiple layers of material with layer 

thickness and porosity being defined in the process. In the layup of each material, materials can be 

bonded using a functional gradient, where the two adjacent materials are mixed over a few layers 

to guarantee that a bond is achieved. Depending on the layer thickness and overall element size, 

grooves can be left in preselected positions during the VPS process to allow placement of instrumentation (along the centerline and radius of the heater elements). Inconel tubing can be inserted 

into these channels, which can then be used to position temperature sensors (type C TCs or fiber 

optic Bragg grating sensors). Options for directly building up tubing during the VPS manufacturing process will also be investigated. The VPS layup of the heater element would be performed 
103

in a low pressure (≈10 kPa (≈76 torr)), UHP He environment, trapping the He inside the element 

and improving the conductivity across the element. Heater leads could be embedded in the heater 

element during the VPS process. (The current proposal would include three-quarters of an inch of 

the heater lead embedded in the element.)

5.2.2 Advanced Temperature Sensors

The fiber Bragg grating (FBG) temperature sensors developed by Luna Innovations and 

previously tested in the SAFE-100 and 100a cores allow detailed determination of axial temperature distribution with a sensor that is only 0.2 mm (0.008 in) in diameter, minimizing both the 

number and size of the penetrations necessary to fully characterize the temperature distribution 

in the simulated reactor core or in a thermal simulator.18,19 To protect the sensor itself, these fibers 

are encased in a stainless steel sheath that can be as small as 0.5 mm (0.02 in) in diameter. The 

FBG sensors are currently limited to a maximum temperature of ≈1,000 °C, but can be pushed to 

as high as ≈1,100 °C. Due to the temperature limitation, use of the FBG sensors to characterize 

the internal temperature of the thermal simulators during operation may not be a current option, 

but the sensors could be included in the representative core block in the heater element test chamber to assess temperature distributions along its length. Luna Innovations is seeking to extend 

the temperature range over which the FBG sensors are operational and successful advanced FBG 

sensor designs may be incorporated into the thermal simulators at a later date. As an alternative, 

more traditional means of embedded temperature measurement devices, such as high-temperature 

type C TCs, will be pursued in the initial phase of this proposed work.

5.2.3 Power Profiling

In an operating nuclear reactor, the axial power profile approximates a cosine distribution. For a selected reactor design, heater elements can be developed to match the corresponding 

axial power profile and the radial power profile of an operating reactor can be matched in a nonnuclear test article by operating the heater elements in independent zones using multiple power 

supplies. The advanced heater development work will also entail investigation into methods of 

better matching the power profile of the heater element to the reactor fuel elements, both during 

nominal and off-nominal operation. A conceptual design of a heater having two power profiles 

has been generated and is shown in figure 61. The complete design drawing is included in appendix O. The proposed design is essentially constructed from two heater elements, with the smaller 

diameter element fitting completely inside the larger diameter element. This design would allow 

the power density to be shifted quickly during test to simulate a failure condition. For instance, in 

a reactor that is controlled by sliding reflectors, a stuck slider could result in an off-nominal axial 

power profile. If the shifting power profile design cannot be manufactured, or if it is found that 

the center element would reach temperatures that exceed materials limitations, then several heater 

elements could be constructed with different axial power profiles. These elements would then be 

interchangeable in the reactor simulator, with changeout being achieved over 2 to 3 days, to test 

off-nominal conditions. 
104

Figure 61. Heater element with a shifting axial power profile.

To allow for rapid transition from one test configuration to another, developing methods 

that will allow the heater elements to be removed and replaced in the reactor simulator in a simpler, more efficient fashion without damaging the elements or the input power leads will be investigated. Extended testing using the graphite rod heater elements (including hundreds of thermal 

cycles, testing up to high temperatures and power densities, and significant handling by the experimenters) has also made it evident that it will be necessary to develop a stronger lead connection 

mechanism that will endure extended testing and handling of the heater elements.

Alternate heater assembly cross sections are also being investigated. To date, all heater elements and sheathed assemblies have had a circular cross section. Depending on the reactor design, 

it may be desirable to develop a heater assembly with a noncircular cross section to better simulate 

the heat that would be produced in the corresponding nuclear fuel rod. A conceptual design for a 

hexagonal fuel rod has been developed and is shown in figure 62. A complete design drawing for 

the hexagonal heater element is included in appendix P.

Figure 62. Mandrel design for a heater element with a hexagonal cross section.
105

6. CONCLUSIONS

The materials testing performed to date has provided a basic platform for future heater 

assembly development for simulated reactor testing. This series of testing has focused on establishing a minimum achievable assembly size, developing a sealing method to improve conductivity 

across the assembly, and preventing contamination of the reactor test article. Future testing will 

test the current heater elements to higher power levels (6,000 W/element), and will seek a broader 

materials list that could potentially be applied in heater elements in a variety of reactor test configurations. Investigation will continue to identify a method that reduces interaction between the 

heater materials and the test article in addition to developing an acceptable sealing procedure. 

The current thread of thermal simulator development was concluded prior to completion of 

the stated objectives due to a change in program direction and a stop work order from NRPCT. To 

date, significant work has been completed on the engineering of refractory metal heater elements 

that may be suitable for use in a refractory metal core block to avoid any potential contamination 

of the core by the thermal simulators. However, should compatibility issues not be a concern based 

on the materials selection in the final core design, graphite heater elements have been demonstrated 

to be more robust and more reproducible in their fabrication than the corresponding refractory 

wire-wrapped heater elements. Planned future work entails the development of highly instrumented 

thermal simulators that better mimic the operation of nuclear fuel pins. Sample advanced simulator 

elements can be tested individually in relevant environments to assess their characteristics relative 

to thermal analysis predictions, and can then be applied in partial or full-array reactor core testing. Engineering of thermal simulators that better mimic nuclear fuel pins, both in their static and 

dynamic characteristics, can offer significant enhancement to realistic nonnuclear test methodologies, as discussed in Bragg-Sitton and Webster.19 Initial design work has also been completed on 

a power interface assembly that minimizes the impact on the gas flow plena in a direct gas-cooled 

reactor design, per the down-select made by NRPCT in mid-2005. Each additional step toward 

developing an electrically heated test article (and integrated system) that is more prototypic of the 

planned nuclear system significantly enhances the value of the data acquired from the test, offering 

a vast knowledge base to be developed at a relatively small fiscal investment and over a short period 

of time as compared to a full nuclear test.
106

APPENDIX A—GRAPHITE ROD HEATER ELEMENT

Figures 63 and 64 show the varying diameter (0.375-in-OD) and constant diameter 

(0.305-in-OD) graphite rod heater elements, respectively.
107

Figure 63. Varying diameter graphite rod heater element, 0.375-in OD.
108

Figure 64. Constant diameter graphite rod heater element, 0.305-in OD.
109

APPENDIX B—SPIRAL WRAPPED SIMULATOR DESIGN, 0.625-IN ASSEMBLY

Figures 65–70 show the mandrel, coupler, sheath, endcap assembly, and full assembly, 

respectively, (0.625-in) for the spiral-wrapped simulator design.
110

Figure 65. Mandrel, 0.625-in assembly design.
111

Figure 66. Coupler, 0.625-in assembly design.
112

Figure 67. Sheath, 0.625-in assembly design.
113

 

Figure 68. Endcap, 0.625-in assembly design.
114

Figure 69. Endcap assembly, 0.625-in assembly design.
115

Figure 70. Full assembly, 0.625-in assembly design.
116

APPENDIX C—SPIRAL-WRAPPED SIMULATOR DESIGN, 0.400-IN ASSEMBLY

The mandrel coupler, endcap, and sheath are shown in figures 71–74, respectively, for the 

0.400-in assembly for the spiral-wrapped simulator design.
117

Figure 71. Mandrel, 0.400-in assembly design.
118

Figure 72. Coupler, 0.400-in assembly design.
119

Figure 73. Endcap, 0.400-in assembly design.
120

Figure 74. Sheath, 0.400-in assembly design.
121

APPENDIX D—TEMPERATURE DEPENDENT RESISTIVITY OF SELECTED 

REFRACTORY METALS

The temperature-dependent resistivity of refractory metals of interest in the current work 

are summarized in equations D-1 to D-3, where resistivity r is given in 10–6 Ω-cm and temperature 

T is in K:

Tantalum: (3)

Tungsten: (4)

Rhenium: . (5)

Figure 75. Resistivity of selected refractory metals as a function of temperature. (Data 

obtained from R. Kapernick, Los Alamos National Laboratory.)
122

APPENDIX E—TANTALUM HEATER ELEMENT MATERIALS ANALYSIS

Appendix E is a NASA memo dated December 20, 2004, regarding materials analysis of 

the Ta heater element.
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1

December 20, 2004

(EM30 2004-092)

TO: ALLIED/R. Dickens

FROM: EM30/G. Jerman

SUBJECT: Failure Analysis of a Discolored Helical Alumina Insulator, Wire Wrapped by a 

 Tantalum Heater Element (EM30 2004-726)

Introduction

A 0.040 inch diameter tantalum (Ta) wire heater element was tested at high temperature in 

a vacuum chamber to simulate nuclear fission heating. During testing, images of the hot Ta 

element showed differential heating at localized positions along the helical conductor wire that 

twisted around a grooved alumina (Al2

O3) insulator. Figure 1 shows the heater element early in 

its testing. There are two places where the Ta winding does not glow as bright as the adjacent 

winding. Figure 2 shows the heater element later in the test. The locations of differential 

heating have shifted to alternate positions and the overall coloration of the image has red shifted. 

No voltage or current anomalies were noted during heater operation. Upon removal from the 

test chamber, the white alumina insulator had bluish-black discoloration. The Ta heater wire 

remained a shiny silver metallic color. The heater element was then submitted to the EM30 

Materials Analysis Branch to determine the cause of the differential wire heating and the 

insulator discoloration.

Optical Microscopy

In order to analyze the Ta heater wire and the alumina insulator separately, the Ta wire was 

unwound from the helical insulator. After unwinding the Ta wire a few turns, the wire began 

to fracture in a brittle fashion. Large variations in discoloration were noted on the insulator 

surfaces as the Ta wire was removed. Figures 3, 4, and 5 show localized insulator surface 

coloration changes including metallic spheroid deposits. After unwinding about a third of the 

way up the insulator, the insulator fractured. The fracture occurred near the base of the insulator 

where the Ta wire enters and exits the assembly. Discoloration also appears to have penetrated 

into the alumina as shown in figure 6. A second fracture was then induced in the alumina 

insulator to liberate a small segment for further analysis. Pieces of new and thermally tested Ta 

wire were also cut for further analysis. Figures 7 and 8 show the surface condition of the new 

and heated wire while figures 9 and 10 show the cross sections. Although the surface texture of 

the heated wire did not change, large grain growth can be seen in the wire cross section. Such 

a large average grain size with respect to the wire diameter probably caused the observed wire 

embrittlement during disassembly.

National Aeronautics and

Space Administration

George C. Marshall Space Flight Center

Marshall Space Flight Center, AL 35812

Reply to Attn of:
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2

Scanning Electron Microscopy 

Surface morphology and chemistry information were collected from the insulator and conductor 

wires in a Field Emission Scanning Electron Microscope (FESEM). Figures 11 and 12 show 

secondary electron images of the helical insulator at a location that contained both discoloration 

and metallic spheroid deposits. Normally alumina is an insulator which requires conductive 

coating for FESEM imaging. Nonconductive samples collect an electrical charge from the 

electron beam and appear bright white under normal circumstances. The majority of the 

surface area in figures 11 and 12 have significant conductivity. Only the tops of the helical 

ridges and a limited portion of the valley were nonconductive like alumina should be. Because 

of the unexpected electrical conductivity, Energy Dispersive x-ray Spectroscopy (EDS) was 

performed in order to document the surface chemistry that was allowing conductivity. Figure 13 

shows the chemistry of the general discoloration that covers most of the insulator. The general 

discoloration contains a combination of metallic aluminum (Al) and Ta. Metallic Al and Ta can 

form a variety of intermetallic compounds. Figure 14 shows a region of discontinuity in the 

general discoloration. This region contains more Al than the general discoloration, but it also 

contains some oxide. Figure 15 shows the metallic spheroids that were seen optically in figure 

5. They are composed of pure metallic Al. In figure 16, a region that charged electrically was 

analyzed. It was composed mainly of alumina, but it also contained Ta which was also possibly 

in oxide form. Finally, the new and used Ta conductor wires were analyzed by FESEM/EDS. 

The surface of the new wire is shown in figure 17. It is composed of pure Ta, but some traces 

of carbon and oxygen are present. Figure 18 shows the same Ta wire after thermal testing. The 

surface morphology has changed little and small amounts of carbon and oxygen are still present. 

Figure 19 shows the composition of surface contaminants on the new Ta wire. The contaminant 

contains significant amounts of carbon, oxygen, sodium, and chlorine which indicates it may be 

related to human handling. 

Discussion of Differential Heating and Image Red Shift

There are three reasons why local sections of the Ta heater wire glowed less than neighboring 

wire locations during operation: current loss, thermal dissipation, or emissivity changes. No 

significant roughness or chemical changes were noted on the Ta wire surfaces, therefore the 

emissivity of the wire did not change during testing. No emissivity change means no red shift. 

A large temporary change in thermal dissipation would require a large localized improvement 

then loss of thermal conductivity in the insulator, which is highly unlikely. The only likely 

mechanism for the observed localized change in wire heating is loss of electrical current. 

Alumina is an insulator, but FESEM/EDS analysis showed a conductive metallic Al-Ta coating 

covers the insulator surface. If the coating was thick enough, a parallel electrical circuit along 

the insulator surface could draw current away from the Ta wire temporarily. Current flow to 

ground would not change, so from the heater instrumentation standpoint, no current or large 

voltage change would be noticeable. Since the conductor was producing a small amount of Ta 

oxide vapor, it would coat the entire vacuum chamber including the sight glass window. A vapor 

deposited oxide coating would limit light transmittance through the sight glass resulting in the 

red shifted image.

Discoloration Discussion 

The presence of pure metallic Al on the insulator indicates the alumina was being reduced to 
125

3

metallic Al and it was reacting with Ta to form an intermetallic layer on the insulator surface. 

Alumina is known to break down in a vacuum environment in the present of carbon. Since the 

new and tested Ta wires contained measurable amounts of surface carbon contamination, this 

was a possible source of degradation. Metallic Ta melts at 2996 oC but its oxide melts at 1872 o

C. The calculated operational temperature of the tantalum wire could have varied between 1933 

and 2500 o

C. In this temperature range, the wire would not have melted, but any oxide formation 

would have vaporized then condensed on lower temperature surfaces. Since the insulator kept 

its shape and no wide spread slumping or distortion was seen in the insulator rod, the general 

insulator temperature was probably close to the targeted 1000 oC output temperature of the heater 

assembly. Alumina melts at 2072 oC, so it is possible that some localized insulator melting 

occurred where the Ta wire was in intimate contact with the insulator surface. The maximum 

recommended operating temperature for alumina is 1800 o

C, so the use of alumina with such 

high contact temperatures is not recommended.

Conclusions

1. Discoloration of the alumina insulator was caused by the formation of an intermetallic 

 Al-Ta coating on the insulator surface.

2. Carbon bearing contamination in the heater assembly reduced some of the alumina 

 insulator surfaces under vacuum to form metallic aluminum.

3. The hot aluminum metal then reacted with and reduced vaporized tantalum oxide to form 

 a conductive intermetallic coating on the alumina insulator.

4. Vapor deposited tantalum oxide on the sight glass window resulted in the later red shifted 

 test image.

Recommendations

1. The tantalum wire and insulator should be thoroughly cleaned to remove any oils or other

 volatile carbon bearing contamination before assembly.

2. After assembly, the heater should be baked out under vacuum at approximately 150 o

C in 

 order to vaporize any remaining volatile components without causing decomposition of 

 organic molecules that could lead to carbon deposits.

3. The helical insulator should be manufactured from a ceramic with higher temperature 

 capability such as boron nitride (powder consolidated) or pyrolytic boron nitride (vapor

 deposited) which doesnʼt break down until 3000 oC.

_______________________________

Gregory A. Jerman

EM30 Materials Analysis
126

4

Figure 1. Initial thermal test image showing differential heating.

Figure 2. Later thermal test image showing a red color shift and differential heating.

differential heating

differential heating
127

5

Figure 3. Surface condition of the alumina insulator, after testing, with the Ta wire 

 removed. 

Figure 4. Surface reaction zone on the alumina insulator, after testing, with the Ta wire 

 removed. (20x)
128

6

Figure 5. Metallic droplets in the alumina insulator surface reaction zone. (100x)

Figure 6. Penetration of surface discoloration into the insulator cross sectional thickness. 

 (10x)
129

7

Figure 7. Optical view of the new Ta wire surface condition. (50x)

Figure 8. Optical view of the tested Ta wire surface condition. (50x)
130

8

Figure 9. Optical cross sectional view of the new Ta wire grain structure. (50x)

Figure 10. Optical cross sectional view of the tested Ta wire grain structure. (50x)
131

9

Figure 11. Low magnification FESEM view of the helical groove in the alumina support 

 insulator. (20x)

Figure 12. Low magnification FESEM view of the deposits at the bottom of the helical 

 groove in the alumina support insulator. (50x)
132

10

Figure 13. FESEM/EDS analysis of the uniform coating that covers the most of the helical 

 alumina insulator groove.

Figure 14. FESEM/EDS analysis of the discontinuous coating that covers a small portion of 

 the helical alumina insulator groove.
133

11

Figure 15. FESEM/EDS analysis of the spheroidal formations at the bottom of the helical 

 alumina insulator groove.

Figure 16. FESEM/EDS analysis of the region that electrically charges at the bottom of the 

 helical alumina insulator groove.
134

12

Figure 17. FESEM/EDS analysis of new Ta wire from the same lot installed on the helical 

 alumina insulator.

Figure 18. FESEM/EDS analysis of the thermally tested Ta wire that was installed on the 

 helical alumina insulator.
135

13

Figure 19. FESEM/EDS analysis of surface contamination on the new Ta wire from the same 

 lot installed on the helical alumina insulator.
136

APPENDIX F—RHENIUM HEATER ELEMENT MATERIALS ANALYSIS

Appendix F is a NASA memo dated February 4, 2005, concerning Re heater element 

materials analysis.
137

1

February 04, 2005

(EM30 2005-006)

TO: MORGAN/P. Salvail

FROM: EM30/G. Jerman

SUBJECT: Analysis of a Discolored Helical Alumina Insulator, Wire Wrapped by a 

 Rhenium Heater Wire (EM30 2005-0004)

Introduction

A 0.040 inch diameter rhenium (Re) wire, twisted around a helically grooved alumina (Al2

O3) 

insulator, was resistively heated in a vacuum chamber to simulate nuclear fission heating. 

During testing, images of the hot Re element were taken at different power levels. Figure 1 

shows the heater element early in its testing at a 1200 W power level. Figure 2 shows the heater 

element later at a 5000 W power level. Unlike the previous testing of a tantalum (Ta) heater 

element, documented in EM30 memorandum 2004-092, no differential heating was seen with the 

Re element. Upon removal from the test chamber, the white alumina insulator had a bluish-grey 

discoloration. The Re heater wire remained a shiny silver metallic color. The heater element 

was submitted to the EM30 Material Diagnostics Team to determine the cause of the insulator 

discoloration and document the condition of the alumina insulator and the Re heater wire.

Optical Microscopy

The Re wire and the helical alumina insulator were separated for analysis. Figure 3 shows the 

heater assembly before separation. During unwinding, the Re wire was found to be ductile. No 

wire breakage occurred during unwinding. A small segment was cut from the middle of the 

heated Re wire for later analysis. A segment of new Re wire was also available for comparison. 

After removal of the conductor wire, variations in surface discoloration were noted on the 

alumina insulator. Figures 4, 5, and 6 show localized insulator surface coloration changes from 

general bluish-grey to white streaks and dark bluish-grey spots. In order to determine if the 

surface discoloration affected the alumina through its thickness, a small segment of the insulator 

was fractured. Figure 7 shows the insulator cross section. The general surface discoloration 

did not affect the alumina insulator thickness, but some shallow interactions did occur at the 

localized dark bluish-gray spots. Figures 8 and 9 show the surface condition of the new and 

heated wire and figures 10 and 11 show the wire cross sections. The surface texture of the Re 

wire did not change significantly after heating, but some grain growth did occur.

Scanning Electron Microscopy 

Surface morphology and chemistry information were collected from the insulator and conductor 

National Aeronautics and

Space Administration

George C. Marshall Space Flight Center

Marshall Space Flight Center, AL 35812

Reply to Attn of:
138

2

wires in a Field Emission Scanning Electron Microscope (FESEM). Energy Dispersive x-ray 

Spectroscopy (EDS) was used to identify the surface chemistry in different regions of interest. In 

figure 12, FESEM/EDS identified the main constituents of the general bluish-grey discoloration 

as rhenium, aluminum, and oxygen. Figure 13 shows the surface morphology and chemistry of 

a dark bluish-grey spot. Although similar in color to the general surface discoloration, only a 

small amount of rhenium was detected on the alumina insulator. Boron and tungsten were also 

identified in relation to the dark bluish-grey discoloration. Figure 14 shows a region of white 

discontinuity in the general discoloration. This region was identified as mainly alumina with 

a little rhenium present. Figure 15 shows FESEM/EDS results for the new Re conductor wire. 

General EDS of the new Re wire surface revealed some surface contamination. Figure 16 shows 

the EDS results from a single location of surface contamination on the new wire. Significant 

amounts of sulfur, tungsten, oxygen, carbon, potassium, calcium, and chlorine were present in 

the contamination. Figure 17 shows FESEM/EDS results for the heated Re conductor wire. 

General EDS of the surface revealed some contamination was still present after heating the Re 

wire. Figure 18 shows the EDS results from a single location of surface contamination on the 

tested wire. Significant amounts of carbon, sulfur, chlorine, and potassium, are still present after 

heating the Re wire.

Discussion of Insulator Discoloration

The general bluish-grey discoloration on the alumina insulator was caused by vapor deposition of 

rhenium oxides. White streaks in the discoloration probably formed due to contact with the hot 

conductor which prevented condensation of the vaporized rhenium oxide on the alumina surface. 

The lack of rhenium in the dark bluish-grey spots and the presence of tungsten and boron 

indicate these spots probably formed from vaporizing surface contamination on the Re wire. 

Since there was some depth to the interactions at the dark bluish-grey spots, there may have been 

some limited electrical arching or high temperature interactions at conductor contact points on 

the alumina. None of the discoloration significantly affected the thickness of alumina, so the 

insulatorʼs properties were not affected by rhenium oxide deposition. The discolored surfaces 

were somewhat conductive in the FESEM, so the bluish-grey deposits probably contained 

some metallic Re, but the significant presence of oxygen suggests no fully metallic layers were 

deposited on the insulator during heating.

Discussion of the Re Conductor Wire

The rhenium wire surfaces contained significant volatile contamination which could negatively 

impact the alumina insulator. The presence of tungsten in the Re wire surface contamination 

indicates the Re wire was probably contaminated by debris left in wire drawing equipment 

utilized to draw both tungsten and rhenium wire. This contamination, however, did not play a 

significant role in insulator discoloration, and it did not result in degradation of the conductor 

wire. The observed ductility of the Re wire after testing precludes embrittlement due to diffusing 

surface contamination. Optical cross sections of the new Re wire found significant cracking. 

This cracking was probably due to over working the material during wire drawing. Subsequent 

heating of the wire healed many of the internal cracks, but small surface cracks are still present 

which could lead to later wire breakage

.
139

3

Comparison of Tantalum vs. Rhenium Heaters

Previously tested Ta conductor wires and Re conductor wires both had surface contamination 

before use. Only the Ta wire degraded the alumina insulator. Why? The initial Re oxide 

vaporizes at 145 oC while the first Ta oxide vaporizes at 1870 oC. Thus rhenium oxide is present 

on the surface of the alumina insulator up to 1000 oC before it starts to decompose. The presence 

of a Re oxide layer would provide a buffer to alumina decomposition in the presence of carbon 

contamination under vacuum. In the Ta heater, alumina was reduced to metallic aluminum in the 

presence of carbon contamination before Ta oxide was vaporized and deposited on the insulator.

Conclusions

1. The general bluish-grey discoloration on the alumina insulator was caused by the 

 deposition of rhenium oxides.

2. The deposition of low temperature rhenium oxides protected the alumina insulator from 

 high temperature degradation during thermal testing.

3. Because Re oxides preserved the alumina insulator integrity and the Re conductor wire 

 retained a large amount of ductility, the Re/alumina combination was superior to the 

 Ta/alumina combination.

Recommendations

1. Because a small amount of vapor deposited rhenium oxide was beneficial to preserving 

 the alumina insulator, and surface contamination on the conductor wire did not 

 significantly affect the insulator, no operational changes are recommended.

2. Because the protective rhenium oxidation begins to degrade at 1000 o

C, operation of the 

 Re/alumina heater above 1000 oC is not recommended.

_______________________________

Gregory A. Jerman

EM30 Material Diagnostics Team
140

4

Figure 1. Initial Re wire thermal test image at 1200 W power setting.

Figure 2. Later Re wire thermal test image at 5000 W power setting.
141

5

Figure 3. Re wire heater element as received. (7.1x) 

Figure 4. Surface discoloration after Re wire removal. (7.1x)
142

6

Figure 5. Surface discoloration zone 1 on the alumina insulator. (40x)

Figure 6. Surface discoloration zone 2 on the alumina insulator. (40x)
143

7

Figure 7. Fractured alumina cross section showing no through thickness interaction with

 the surface discoloration. (10x)

Figure 8. Optical view of the new Re wire surface condition. (50x)
144

8

Figure 9. Optical view of the tested Re wire surface condition. (50x)

Figure 10. Optical cross sectional view of the new Re wire grain structure. (50x)
145

9

Figure 11. Optical cross sectional view of the tested Re wire grain structure. (50x)

Figure 12. FESEM/EDS analysis of the uniform bluish grey coating that covers the most of 

 the helical alumina insulator groove.
146

10

Figure 13. FESEM/EDS analysis of the dark bluish grey spot in the valley of the helical 

 alumina insulator groove.

Figure 14. FESEM/EDS analysis of the white streak in the valley of the helical alumina 

 insulator groove.
147

11

Figure 15. FESEM/EDS analysis of new Re wire from the same lot installed on the helical 

 alumina insulator.

Figure 16. FESEM/EDS analysis of the surface contamination on the new Re wire.
148

12

Figure 17. FESEM/EDS analysis of the thermally tested Re wire that was installed on the 

 helical alumina insulator.

Figure 18. FESEM/EDS analysis of surface contamination on the thermal tested Re wire.
149

APPENDIX G—KEY GOVERNING EQUATIONS AND SUPPORTING FORMULATIONS

Concentric cylindrical gap conductive conductance (typically in W/K)

,

(6)

where L is the length of the gap, Dinner and Douter are the inner and outer diameters of the gap, 

and k is the conductivity of the media in the gap. 

Concentric cylindrical gap radiative conductance (typically in W/K)

, (7)

where A is the surface area of the inner surface that faces the gap, σ is the Stefan-Boltzman constant, and E is the emissivity function that takes into the account the two-dimensional geometry 

of the concentric cylinders. This assumes the axial length of the two opposing sides of the gap are 

very long compared the gap thickness. The emissivity function is: 

, (8)

where εinner and εouter are the emissivities of the inner and outer surfaces, respectively, encompassing the gap. This assumes that the surfaces are gray and diffuse. When there are more than two 

concentric cylinders making a series of gaps, these (radiative) gap conductances can be added in a 

serial fashion to obtain an overall conductance, which might be used, for example, to estimate the 

heat transfer between a heater wire and chamber boundary through a sheath. 

Heat balance equation

The conductances are then used to determine temperatures by solving equations such as 

the following for either temperature or heat transferred:

, (9)

where Tinner and Touter are the temperatures of the inner and outer surface and the Q is the heat 

transferred between the surfaces. 
150

Conductive shape factors:

When the geometry is more complex than a concentric cylinder or slab, determining the 

conductive conductances can be performed by using what is referred to a conduction shape factor 

S as in the following equation:

, (10)

where T1 and T2 are the temperatures of two surfaces or thermal nodes, k is the thermal conductivity of the material between the two surfaces, Qcon,12 is the net heat transferred by conduction 

between surfaces 1 and 2, and S12 is the conductive shape factor that takes into account the area 

of the surfaces and geometry of the thermal path between 1 and 2. In the current analysis, the 

shape factors were obtained using a partial finite element model of the gap between the heater 

wire, mandrel, and enclosing sheath. Shape factors were used to perform a preliminary level 

analysis or as part of a closed-form parametric solution to support trade studies. Because of the 

simplifications in the shape factors’ development and computation, their use is subject to a number of restrictions and caution should be used when applying them. For more accurate estimates, 

complete finite difference and finite element solutions can be used. 

Radiative exchange factors:

When the geometry is more complex than flat plates or concentric cylinders and when radiation leaving or striking a surface is due to emission and reflection, radiative exchange factors can 

be computed that accurately (again, under certain restrictions) model the entire radiative transfer 

between an assembly of discrete surfaces. The following equation depicts, in a general sense, what 

is meant by radiative exchange factors:

, (11)

where Qrad,i is the net transfer into and out of a surface i, by radiative exchange with other surfaces 

in the enclosure, Ai is the surface area of the specific surface i, fij is the radiative exchange factor 

from surface i to j and encompasses the surface emissivities and geometries of the various paths by 

which energy can be radiated from i and reach surface j, Ti and Tj are the surface temperatures of 

surfaces i and j, respectively, N is the number of surfaces that contribute to the radiative enclosures, 

and σ is again the Stefan-Boltzman constant. The f

ijs are derived from calculated view factors and 

surface emissivities. These view factors and exchange factors can be calculated by hand per methods described in texts such as Thermal Radiation and Heat Transfer by Siegel and Howell,21 as was 

performed in this case, or using a computer program such as TRASYS or Thermal Desktop or by 

a finite element modeler/solver such as FIDAP, as was used in this case as a backup solution.
151

Heat balance equation part 2:

The radiative exchange factors and conductive shape factors can be combined along with 

an estimate of the heat generated in a component such as the heat in the wire arising from an 

applied voltage and the electrical resistance to compute the temperatures of the various components. The governing equation follows:

, (12)

where Qgen,i is the heat generated within a surface or thermal node i, and Sijs are the conduction 

shape factors between nodes i and j. This equation can be written for each of the N surfaces in the 

enclosure then expanded and tailored, resulting in a set of equations that can be solved iteratively 

to obtain the unknown temperatures. This was performed in the Mathcad solution approach. For 

the SINDA solution, the radiative exchange factors and the conductive shape factors are submitted as part of the model; the conductors (both radiative and conductive) were calculated and the 

SINDA solution routines were applied to solve for temperatures. 
152

APPENDIX H—THERMAL CROSS-SECTION CALCULATIONS

The thermal conductivities of selected gases, at temperature, are provided in figure 76. Gas 

properties were obtained from: Incropera and DeWitt, Arp et al., Manglik, Bailey, and Jain.10-14

0

0.05

1

0.15

2

0.25

3

0.35

4

0.45

5

0 500 1,000 1,500 2,000 2500

Temperature (K)

Air

Helium

Argon

Xenon

Conductivity (W/m-K)

Figure 76. Thermal conductivity of selected gases.
153

Summary of MathCad Calculations for Analysis of Sheathed Graphite Heater Elements

One Dimension, Cross-Sectional Heat Transfer Calculations

Heater/Sheath Geometry

IDcore := 0.545 in

Lsheath := 20in

ODsheath := 0.540 in

IDsheath := .465in IDsheath = 11.8 mm

 

ODheater := .350in ODheater = 8.9 mm

thsheath

ODsheath − IDsheath

2 := thsheath = 1 mm

Geometric Calculations 

radialgap

IDsheath − ODheater

2 := radialgap = 0.1 in

Asrfsheath := π ⋅ODsheath ⋅Lsheath radialgap = 1.5 mm

Asrfheater := π ⋅ODheater ⋅Lsheath

Design Conditions: Heater Power and Sheath Temperature

Tsheath := (1000 + 273.15 )K

Qheater := 6000W
154

Gap Heat Transfer Calculations: Gas Gap, conduction only

Gas in gap is helium

Helium thermal conductivity 

Read in data from general gas properties spreadsheet 

	 (summarized in fig. 75)

TKheND:= KheND

:=

Adjust units of imported data

TKhe := TKheND ⋅K Khe KheND

W

m⋅K := ⋅

Radial Conductor, G

Tgap := Tsheath

Kgas := linterp(TKhe , Khe ,Tgap) Kgas 0.4 kgm

s

3

K

=

Gheonly

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= Gheonly 4.798

W

K =

Theater Qheater

Gheonly := + Tsheath Theater = 2524 K

ΔThs := Theater − Tsheath ΔThs = 1251 K

eND:=
155

Gap Heat Transfer Calculations: Gas Gap, conduction only

Gas in gap is Argon

 Argon thermal conductivity 

 Read in data from general gas properties spreadsheet

TKarND:= KarND

:=

Adjust for units of imported data

TKar := TKarND ⋅K Kar KarND

W

m⋅K := ⋅

Radial Conductor, G

Tgap := Tsheath

Kgas := linterp(TKar , Kar ,Tgap) Kgas 0.1

W

m⋅K =

Garonly

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= Garonly 0.56

W

K =

Theater Qheater

Garonly := + Tsheath Theater = 11938 K

ΔThs := Theater − Tsheath ΔThs = 10665 K
156

Gap Heat Transfer Calculations: Gas Gap, conduction only

Gas in gap is Xenon

Argon thermal conductivity 

Read in data from general gas properties spreadsheet

TKxeND:= KxeND:=

Adjust for units of imported data

TKxe := (TKxeND + 273.15 )⋅K Kxe KxeND

W

m⋅K := ⋅

Radial Conductor, G

Tgap := Tsheath

Kgas := linterp(TKxe , Kxe ,Tgap) Kgas 0.018

W

m⋅K =

Gxeonly

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= Gxeonly 0.21

W

K =

Theater Qheater

Gxeonly := + Tsheath Theater = 30393 K

ΔThs := Theater − Tsheath ΔThs = 29120 K
157

Effect of gap size

	 Since the OD of the sheath is fixed; the gap size can be reduced by thickening the sheath –

	 decreasing the sheath ID or by increasing the heater diameter, which requires redesign of

	 of the resistive element.

 IDshth := 0.465 in, 0.460 in.. 0.355 in Kgas := linterp(TKhe , Khe ,Tsheath)

Ggap(IDsheath)

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= graph units: x -> m; y ->W/K

0.009 0.0095 0.01 0.0105 0.011 0.0115 0.012 0

50

100

150

Ggap(IDshth)

IDshth

ΔT(IDsheath, Qheater) Qheater

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:=

 

	 	 graph units: x -> m; y ->K or C

0.370 in = 0.0094 m Ggap(.370in) 26.7

W

K = ΔT(.370in, 6000W) = 225.1 K
158

Gap Heat Transfer Calculations: Vacuum Gap, radiation only

Assumed heater and wall emissivities and important constants

σ 5.67 10− 8 ⋅ W

m2K

4 ⋅

:= Ehs(εh, εs) 1

εh

1 − εs

εs

ODheater

IDsheath

⎛

⎜

⎝

⎞

⎟

⎠ + ⋅ ⎡

⎢

⎣

⎤

⎥

⎦

− 1

:=

εh := .9 Heater surface emissivity;

	 	 typical graphite emissivity

 εs := .7 Sheath ID emissivity

	 	 typical oxidized stainless steel; or potential emissivity

	 	 enhanced refractory surface

 Theater Qheater

Asrfheater⋅Ehs(εh, εs)⋅σ Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 1910 K

	 Effective Radial conductor G normalized for the heater and sheath conditions;

	 note that this radial conductor is highly temperature dependent.

Gradonly Asrfheater⋅Ehs(εh, εs)⋅σ ⋅(Theater + Tsheath) Theater2 Tsheath2 := ( + )

Gradonly 9.4

W

K =

ΔThs := Theater − Tsheath ΔThs = 637 K
159

Gap Heat Transfer Calculations: Vacuum Gap, radiation only, different radiative properties

	 Alternative emissivity combination #1

εh := .9 Heater surface element emissivity

	 	 Typical graphite emissivity

εs := .2 Sheath ID emissivity

	 	 Represents polished moly, tantalum sheath;

	 	 uncertain of long term of stability of oxidized

surfaces (TaO2, MoO2, etc)

Theater Qheater

Asrfheater⋅Ehs(εh, εs)⋅σ Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 2403 K

	 Effective Radial conductor G normalized for the heater and sheath conditions;

	 note that this radial conductor is highly temperature dependent.

Gradonly Asrfheater⋅Ehs(εh, εs)⋅σ ⋅(Theater + Tsheath) Theater2 Tsheath2 := ( + )

 Gradonly 5.3

W

K =

ΔThs := Theater − Tsheath ΔThs = 1130 K
160

Gap Heat Transfer Calculations: Vacuum Gap, radiation only, different radiative properties

	 Alternative emissivity combination #2

εh := .5 Heater surface element emissivity

	 	 Carbon filament, could also represent ceramics

	 	 such as boron nitride, alumina.

 εs := .2 Sheath ID emissivity

	 	 Represents polished moly, tantalum sheath;

	 	 uncertain of long term of stability of

	 	 oxidized surfaces (TaO2, MoO2, etc)

Theater Qheater

Asrfheater⋅Ehs(εh, εs)⋅σ Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 2515 K

	 Effective Radial conductor G normalized for the heater and sheath conditions;

	 note that this radial conductor is highly temperature dependent. It is also dependent

	 on geometry somewhat, although not as much.

Gradonly Asrfheater⋅Ehs(εh, εs)⋅σ ⋅(Theater + Tsheath) Theater2 Tsheath2 := ( + )

Gradonly 4.8

W

K =

ΔThs := Theater − Tsheath ΔThs = 1242 K
161

Gap Heat Transfer Calculations: Vacuum Gap, radiation only, different radiative properties

	 Alternative emissivity combination #3

Heater surface element emissivity

 εh := .5 Carbon filament, could also represent ceramics

such as boron nitride.

Sheath ID emissivity

 εs := .7 typical oxidized stainless steel; or potential emissivity

	 	 enhanced refractory surface

Theater Qheater

Asrfheater⋅Ehs(εh, εs)⋅σ Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 2113 K

	 Effective Radial conductor G normalized for the heater and sheath conditions;

	 note that this radial conductor is highly temperature dependent. It is also dependent

	 on geometry somewhat, although not as much.

Gradonly Asrfheater⋅Ehs(εh, εs)⋅σ ⋅(Theater + Tsheath) Theater2 Tsheath2 := ( + )

Gradonly 7.1

W

K =

ΔThs := Theater − Tsheath ΔThs = 840 K
162

Gap Heat Transfer Calculations: Gap radiation and conduction combined

Th := Theater εh := .5 εs := .2

Th root Gheonly⋅(Th − Tsheath) Asrfheater⋅σ ⋅Ehs(εh, εs) Th4 Tsheath4 + ⋅( − )− Qheater ,Th ⎡

⎣ ⎤ := ⎦

Th = 2036 K

Gcomb Qheater

Th − Tsheath := Gcomb 7.9

W

K =

ΔThs := Th − Tsheath ΔThs = 763 K
163

Gap Heat Transfer Calculations: Conductive gap filler; ceramic powder w/ helium

Boron Nitride conductivity dat

Computation of packed bed conductance; based on scott gilley’s

	 method/research in development of microgravity processing SACA’s with power filled gaps

BSF := .5 BSF => Bulk Solid Fraction

XN := 11.6 ⋅BSF

Kgas := linterp(TKhe , Khe ,Tgap) Kgas 0.4

W

m⋅K =

Kpwdr := linterp(TKbnsg , Kbnsg ,Tgap) Kpwdr 46.4

W

m⋅K =

Kbulk Kgas 0.5 ⋅XN ⋅BSF Kpwdr

Kpwdr − Kgas ⋅

Kpwdr ln Kpwdr

Kgas

⎛

⎜

⎝

⎞

⎟

⎠ ⋅

Kpwdr − Kgas − 1

⎛

⎜

⎜

⎝

⎞

⎟

⎟

⎠ ⋅ + 0.2146

⎡

⎢

⎢

⎣

⎤

⎥

⎥

⎦ := ⋅

Kbulk 2.4

W

m⋅K =

Equation for Kbulk is based on Hengst-Kaganer formulation for Keff/Kgas

GBNfiller

2⋅π ⋅Lsheath⋅Kbulk

ln

IDsheath

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= GBNfiller 27.2

W

K =

Theater Qheater

GBNfiller := + Tsheath Theater = 1494 K

ΔThs := Theater − Tsheath ΔThs = 220 K
164

Gap Heat Transfer Calculations: Conductive gap filler; ceramic liner w/ helium

Calculate thickness of liner

th := 0.8 ⋅radialgap th = 0.046 in

Calculate thickness of gaps between liner and heater and liner and sheath

gapsmall radialgap − th

2 := gapsmall = 0.006 in

IDliner := ODheater + 2⋅gapsmall IDliner = 0.361 in

ODliner := ODheater + 2⋅gapsmall+ 2⋅th ODliner = 0.454 in

Ghesmallinner

2⋅π ⋅Lsheath⋅Kgas

ln

IDliner

ODheater

⎛

⎜

⎝

⎞

⎟

⎠

:= Ghesmallinner 42.161

W

K =

Ghesmallouter

2⋅π ⋅Lsheath⋅Kgas

ln

IDsheath

ODliner

⎛

⎜

⎝

⎞

⎟

⎠

:= Ghesmallouter 54.43

W

K =

Kliner := Kpwdr

Glineronly

2⋅π ⋅Lsheath⋅Kliner

ln

ODliner

IDliner

⎛

⎜

⎝

⎞

⎟

⎠

:= Glineronly 653.02

W

K =

GBNliner

1

Ghesmallinner+ Gradonly

1

Ghesmallouter+ Gradonly + 1

Glineronly + ⎛

⎜

⎝

⎞

⎟

⎠

− 1

:=

 

GBNliner 26.277

W

K =

Theater Qheater

GBNliner := + Tsheath Theater = 1501 K

ΔThs := Theater − Tsheath ΔThs = 228 K

GBNlinernogas

1

Gradonly

1

Gradonly + 1

Glineronly + ⎛

⎜

⎝

⎞

⎟

⎠

− 1

:= GBNlinernogas 3.551

W

K =

Theater Qheater

GBNlinernogas := + Tsheath Theater = 2963 K

ΔThs := Theater − Tsheath ΔThs = 1690 K
165

Task 1a: effect of sheath on heater temperatures–low power

Qheater := 1000W

Tc := (50 + 273.15 )⋅K Temperature of Vacuum Chamber Wall

εh := .9 heater emissivity; represents graphite heater type

	 	 	 radiative performance

εc := .2 chamber wall emissivity; represents clean ss wall

IDchamber := 3ft Approximate interior diameter of vacuum chamber

	 	 	 housing heater and sheath

εs := .2 sheath interior and exterior emissivity representing

	 	 	 bare (clean) stainless, moly or tan

No sheath configuration

Ehc(εh, εc) 1

εh

1 − εc

εc

ODheater

IDchamber

⎛

⎜

⎝

⎞

⎟

⎠ + ⋅ ⎡

⎢

⎣

⎤

⎥

⎦

− 1

:=

Theater Qheater

Asrfheater⋅σ ⋅Ehc(εh, εc) Tc

4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:=

 Theater = 1096 K

With Polished stainless steel sheath surrounding heater

Esc(εs, εc) 1

εs

1 − εc

εc

ODsheath

IDchamber

⎛

⎜

⎝

⎞

⎟

⎠ + ⋅ ⎡

⎢

⎣

⎤

⎥

⎦

− 1

:=

Tsheath Qheater

Asrfsheath⋅σ ⋅Esc(εs, εc) Tc

4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:=

 Tsheath = 1422 K

Theater Qheater

Asrfheater⋅σ ⋅Ehs(εh, εs) Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:=

 Theater = 1742 K
166

Task 1a: effect of sheath on heater temperatures - high power

Qheater := 6000W

	 Same operical properties and chamber temperature conditions

With no sheath

Theater Qheater

Asrfheater⋅σ ⋅Ehc(εh, εc) Tc

4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 1712 K

With Polished stainless steel sheath surrounding heater 

Tsheath Qheater

Asrfsheath⋅σ ⋅Esc(εs, εc) Tc

4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Tsheath = 2224 K

Theater Qheater

Asrfheater⋅σ ⋅Ehs(εh, εs) Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 2726 K

With sheath having modified optical properties

εs := .7

Tsheath Qheater

Asrfsheath⋅σ ⋅Esc(εs, εc) Tc

4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Tsheath = 1638 K

Theater Qheater

Asrfheater⋅σ ⋅Ehs(εh, εs) Tsheath4 + ⎛

⎜

⎝

⎞

⎟

⎠

1

4

:= Theater = 2057 K
167

Temperature differential across sheath

Kss 28

W

m⋅K :=

Knb 69

W

m⋅K := ⋅ IDshth := ODheater + 0.005 in

Kmo 96

W

m⋅K := Gsheath(Ksheath ,IDshth)

2⋅π ⋅Lsheath⋅Ksheath

ln

ODsheath

IDshth

⎛

⎜

⎝

⎞

⎟

⎠

:=

Thick sheath

ΔTsheathss

6000W

Gsheath(Kss ,IDshth) :=

 ΔTsheathss = 28.2 K

ΔTsheathnb

6000W

Gsheath(Knb ,IDshth) :=

 ΔTsheathnb = 11.4 K

ΔTsheathmo

6000W

Gsheath(Kmo ,IDshth) :=

 ΔTsheathmo = 8.2 K

Thin Sheath

ΔTsheathss

6000W

Gsheath(Kss ,IDsheath) :=

 ΔTsheathss = 10 K

ΔTsheathnb

6000W

Gsheath(Knb ,IDsheath) :=

 ΔTsheathnb = 4.1 K

ΔTsheathmo

6000W

Gsheath(Kmo ,IDsheath) :=

 ΔTsheathmo = 2.9 K
168

APPENDIX I—SINDA RESULTS: TEMPERATURE PROFILES

SINDA heater simulation results are shown in table 17.

Table 17. Summary of results from SINDA heater simulation (temperatures are given in Kelvin).

Case

Cu Lead Mo Transition Heater Element Sheath

Min Max Min Max Min Max Min Max Comments

Case 1 778 1,069 1,107 1,475 1,468 1,471 1,348 1,349 1 kW; *PPF=1

Case 2 778 1,085 1,126 1,547 1,545 1,813 1,362 1,417 3 kW; PPF=1.33

Case 3 778 1,119 1,165 1,707 1,709 1,712 1,394 1,395 3 kW; PPF=1

Case 4 778 1,122 1,168 1,720 1,723 2,153 1,398 1,508 6 kW; PPF=1.33

Case 5 778 1,176 1,232 1,985 1,993 2,001 1,462 1,464 6 kW; PPF=1

Case 6 332 1,121 1,168 1,720 1,723 2,153 1,398 1,508

6 kW;

PPF=1.33;

cooled lead end

* PPF corresponds to the peak to average power peaking factor.
169

Figure 77. Axial temperature profile for case 1.

Figure 78. Axial temperature profile for case 2.
170

Figure 79. Axial temperature profile for case 3.

Figure 80. Axial temperature profile for case 4.
171

Figure 81. Axial temperature profile for case 5.

Figure 82. Axial temperature profile for case 6.
172

APPENDIX J—THERMAL ANALYSIS: CALORIMETER

Thermal Analysis of the Cold Wall Calorimeter and Proposed Sheath Instrumentation for Fuel Pin 

Thermal Simulator Testing

Abstract

In order to perform high power testing of the fuel pin thermal simulator concepts, a cold 

wall calorimeter was designed and constructed. This cold wall calorimeter was intended to remove 

a large amount of heat (~6000 W) from a single simulator while maintaining the simulator temperature at levels more representative of actual-use conditions in an electrically heated reactor 

core array. The calorimeter is also designed to allow for improved instrumentation in single element testing, necessary to determine element temperature at various positions and to measure the 

total heat removal by the calorimeter. While preliminary thermal analyses were considered in this 

design, a more detailed thermal analysis of the as-built design was required to verify the operation 

of the calorimeter and to determine operating ranges and thermal behavior in different ambient 

conditions. The analysis data will be compared to test data to correlate the model for use in other 

simulator analyses. An additional analysis was performed in an attempt to quantify the conduction error inherent in the compression-based thermocouples intended to measure the temperature 

of the simulator external sheath surface.

Calorimeter and Thermal Model Description

The calorimeter is composed of a primary large diameter copper tube (ID = 0.68", OD = 

0.75", Length = 24"), surrounded by six coiled copper tube water loops, and associated instrumentation. Each of these water loops is connected in parallel to a central water supply loop. The 

loops are helically wrapped around the primary tube with spacing between the loops to allow both 

thermocouple and optical pyrometer access to the fuel pin simulator sheath through small holes 

in the primary tube. The coils are not physically attached to the primary tube but are wrapped to 

minimize the gap between each coil and the primary tube outer diameter. There is limited contact 

between the coiled loops and the primary tube, but the amount of contact (contact area and pressure) is not controlled nor has it been quantified.

The analysis model consists of both hydraulic and thermal elements. The hydraulic element serves to estimate the distribution of mass flow and the ensuing temperature increase in 

water that flows in each coil as it removes heat from the portion of the primary tube (or cold wall) 

to which that coil is thermally coupled. The hydraulic element consists of sets of fluid nodes and 

paths representing each coil in addition to supply and return lines and a mass flow source. There 

are also thermal ties associated with each fluid node and the adjacent thermal node to determine 

the amount of heat that is convected away from the surface by each node in the water circuit. The 
173

thermal element consists of sets of thermal nodes and connecting conductive and radiation conductors that represent the following components: copper tubing in the water lines and primary 

tube, molybdenum fuel pin simulator sheath, graphite rod heater, and associated alumina spacers 

that electrically isolate the graphite from the sheath.

The thermal model considers the convective heat transfer from the cold wall to the water 

loops, conductive heat transfer within the various solid components, conductive heat transfer 

across any narrow gas gaps and radiative heat transfer between components and to the environment. The ambient conditions considered pure helium gas, pure argon gas and vacuum. Appropriate constant and temperature dependent material properties (e.g. thermal conductivity and 

total emissivity) were included in the model. While density and specific heat were also included in 

the model, only steady state analyses were performed. 

The thermal and hydraulic model, shown in figure 83, was created using Thermal Desktop 

(by Culimore and Ring). Only the primary tube and the fluid circuit are shown; all other components are located concentrically within the cold wall. The blue arrows in the figure represent the 

fluid path, and each arrow adjacent to the cold wall represents one loop around the cold wall. The 

radiative exchange calculations were performed using RADCAD, a module of Thermal Desktop, 

and were incorporated into the thermal/hydraulic solution computed by SINDA/FLUINT.

Figures 84–87 provide typical contour results for the cold wall, sheath and heater for the 

case of a helium environment, a heater power of 6000 W, and a total cooling water mass flow rate 

of 5 gal/min. A more quantitative depiction of the results for this case is provided in figure 88.

The model assumes an inlet water temperature of 293 K (20 °C). The outlet water temperature from each branch is slightly different, as indicated by the contours applied to the fluid 

nodes in figure 85. After all branches merge, however, the temperature of the water flow return is 

approximately 317.7 K (24.7 °C). The cold wall contour is provided here for reference; its scale is 

not shown in figure 85 but is identical to that shown in figure 84.

Banding is seen in the cold wall temperature profiles in both the contour plots (fig. 84–87) 

or the quantitative x-y plot (fig. 88). This characteristic results from the spacing between coils 

where no direct cooling is applied to the cold wall. This is the area where small holes are placed 

in the cold wall for the sheath thermocouples and optical pyrometers. The heat received from the 

sheath in these ‘uncooled’ regions conducts axially to the nearest cooled portions of the primary 

tube where it is removed by the water flow. The effect of this banding on the sheath and heater 

temperature profiles appears to be negligible.
174

Figure 83. Cold wall/calorimeter primary tube and fluid circuit.
175

Note: The colors applied to the fluid paths represent flow rate (scale not provided).

Figure 84. Predicted cold wall temperatures for the case of 6,000 W, He atmosphere, 19 L/min 

(5 gal/min) water flow.
176

Figure 85. Water temperatures for the case of 6,000 W, He atmosphere, 19 L/min (5 gal/min)

water flow.
177

Figure 86. Sheath temperature profile for the case of 6,000 W, He atmosphere, 19 L/min 

(5 gal/min) water flow.
178

Note: The colors applied to the fluid paths represent flow rate (scale not provided).

Figure 87. Heater temperature profile for the case 6,000 W, He atmosphere, 19L/min (5 gal/min) 

water flow.
179

Figure 88. Temperature vs. axial position for the sheath, heater, and cold wall for the case 

of 6,000 W, He atmosphere, 19L/min (5 gal/min) water flow. (Axial position is the 

position along the longitudinal axis of the heater/sheath and zero represents the 

end of the cold wall.)

Table 18 contains a comparison of predicted temperatures for the three different ambient 

conditions (He, Ar, or vacuum), three heater power levels, and three total cooling water flow rates. 

Results for the argon and vacuum conditions at 6,000 W heater power indicate excessive heater 

and sheath temperatures, as well as elevated cold wall temperatures. These temperature estimates 

may be conservative in that the cold wall (and, in turn, the sheath and the heater) is coupled to the 

cooling coils primarily by gas gap conduction and radiation; i.e. negligible metal-to-metal contact conductance is modeled between the cooling coils and the primary cold wall. Consequently, 

when the gap is filled with a low conductivity gas such as argon or a vacuum exists, the conductance between the primary tube and cooling tubes is small or nonexistent. The assumption of 

minimal metal-to-metal contact may be overly conservative as there is probably some contact 

between the coils and the primary tube. This contact can be estimated from early testing and 

results and implemented into the thermal model which can then be used to revise the predicted 

temperatures. If the assumption of poor or no contact is found to be valid through initial testing, it may be advisable to improve the coupling between the cooling coils and primary tube. In 

the original calorimeter design, the cooling coils were brazed to the primary tube to enhance this 

conductance. Early attempts to apply this braze were inconsistent and unrepeatable. Additional 
180

development may still be necessary. Alternatively, a thermally conductive glue (e.g. Ecco bond) 

or vacuum compatible thermal grease (e.g. arctic alumina or other) could be applied in the gap 

between the coils and primary cold wall to improve this coupling and lower cold wall, sheath and 

heater temperatures. 

Table 18. Comparison of temperatures for various atmospheres and conditions.

Configuration/Conditions

Heater 

Temperature

Sheath 

Temperature

Cold Wall 

Temperature

Water 

Temp Branch Flow Rate

Ambient 

Condition

Heater 

Power 

(W)

Total Flow 

Rate 

(gal/min)

Tmax 

(oC )

Tmin 

(oC )

Tmax 

(oC )

Tmin 

(oC )

Thot 

(oC )

Tcold 

(oC )

Tmax 

(oC )

Max 

(gal/min)

Min 

(gal/min)

Helium 6,000 5 2,108 1,697 1,441 450 135 36 27 1.134 0.676

Argon 6,000 5 3,810 3,095 2,991 1,855 539 240 27 1.134 0.676

Vacuum 6,000 5 4,141 3,441 3,395 2,350 1,511 967 26 1.135 0.676

Helium 1,000 5 632 534 408 161 41 24 21 1.136 0.677

Argon 1,000 5 1,948 1,674 1,431 968 141 81 21 1.136 0.677

Vacuum 1,000 5 2,534 2,162 2,062 1,630 849 639 21 1.136 0.677

Helium 6,000 1 2,113 1,699 1,448 454 150 39 56 0.223 0.136

Helium 1,000 1 634 536 410 161 43 24 26 0.225 0.137

Vacuum 500 3.5 2,078 1,806 1,680 1,389 657 524 20.4 .827 .476

Model results indicate that at a heater power of 6,000 W, the calculated sheath temperatures still slightly exceed the maximum target temperature of 1,700 K (1,427 °C) even when high 

conductivity helium is used to fill the gaps between the primary tube and the sheath. This precludes the possibility of ‘tunability’ to achieve a specific sheath temperature below 1,700 K (1,427 

°C). If lower sheath temperatures are required at the higher power settings, then additional cooling techniques or higher conductivity gap fillers will be needed. Alternatively, shrinking the gas 

gap between the primary tube and sheath would increase the conductance between the sheath and 

the primary tube, resulting in the possibility for lower sheath temperatures. Further, an increased 

axial conductance down the cold wall would decrease temperature banding in the results. To 

achieve both of these enhancements, a thicker walled tube with same outer diameter could be 

used. The tube used to assemble the current calorimeter was the thickest available off-the-shelf 

when the system was assembled; a thicker primary tube may require a customized order or may 

require the adoption of a variety of approaches to fabricate a cold wall for this application. 
181

APPENDIX K—THERMOCOUPLE THERMAL ERROR ANALYSIS

L.1 Thermal Model

The primary purpose of the development of the thermocouple thermal model was to evaluate

the potential error associated with spring loaded thermocouples, similar to those depicted in figure14,

employed in the calorimeter to measure sheath surface temperature. While thermocouple error can result

from a number of causes, this modeling and analysis effort focuses on those factors that affect how close

the thermocouple bead temperature is to the temperature of interest: that of the sheath surface.

In order for the bead temperature to closely resemble the sheath surface temperature, the bead

must be closely thermally coupled to the surface of interest and uncoupled from the surroundings as

much as possible. Therefore the thermal model considers these two key aspects: the contact conductance

between the thermocouple bead and the sheath surface and the heat losses from the bead to other parts of

the thermocouple and the surroundings. 

The contact and the associated thermal contact conductance between the bead and the sheath

surface depend on a number of characteristics. Key characteristics include the contact force and area (i.e.,

contact pressure), the relative hardness and roughness of the contacting surfaces, thermal conductivity

of both the bead and sheath surface materials and the presence and properties of any intervening media

(gas, liquid, etc.) between the two contacting surfaces. Typically, the contact conductance increases as

the contact pressure increases, the materials become softer and smoother, or the materials on either side

of the interface become more conductive. The presence of a gas such as air or helium can also significantly increase the contact conductance. The compilation of these parameters makes it very difficult to

predict the contact conductance. Therefore, the analysis presented here considers a number of cases in

an attempt to bound the estimated error. The conductance values are provided in table 19. The conductances for cases B and D were based on a 5 MPa contact pressure (an estimated 1.0 N contact force) and

bead contact area of ≈0.2 mm2 in addition to the conditions stated in the table. The other conditions were

included as bounding cases. All values of the contact coefficient were obtained from empirical data found

in Manglik (1984); the associated contact conductances were derived by dividing the contact area by the

contact coefficient.

Table 19. Bounding contact coefficient and conductance between TC bead and sheath surface.

Case Contact Configuration

Contact Coefficient 

(°C*M2/KW) Conductance (W/°C)

A High resistance 1 0.0002

B Nickel on stainless steel (vacuum) 0.6 0.0003378

C Medium resistance 0.1 0.002

D Aluminum on Aluminum (air) 0.05 0.004054

E Low Resistance 0.01 0.02
182

In the case of the spring loaded K type thermocouples, the thermocouple bead consists predominately of nickel and the sheath has, thus far, typically consisted of molybdenum. Because molybdenum is

more conductive and is similar to or softer than stainless steel, the contact conductance should be higher

than that in case B. However, both nickel and molybdenum are harder and less conductive than aluminum, so the conductance will likely be less than that in case D (a helium ambient environment might

raise the conductance a bit compared to that of air). As there is an appreciable amount of scatter in the

empirical data in Manglik (1984) and since the exact conditions of bead and sheath surface contact are

unknown, it is important to keep in that these are only estimates. In addition, since the potential exists to

operate the calorimeter in different ambient gases, the contact conductance for this spring loaded thermocouple may vary. Finally, while it may be tempting to apply more contact force to increase the contact

pressure and the associated contact conductance, excessive force can be applied that can buckle or break

the exposed lead wire or bead connections. None the less, the 1N force used to obtain these numbers was

only an estimate of that force actually applied by the spring, and improvements in contact conductance

could be realized if the actual force was or could safely be made higher than this value.

The thermal losses to the surroundings and along the thermocouple leads were modeled using

a geometric thermal model that considers solid conduction and surface radiation. The full model, as is

apparent from figures 89–91, consists of a sheath wall held constant at 1,441 °C, the cold wall held constant at 135 °C, and a representative thermocouple that contacts the sheath and passes through the cold

wall. The thermocouple is in turn composed of a bead, an exposed wire lead section, a long metallic

sheathed lead section, a spring section and a plastic connector. Heat is exchanged between the thermocouple sheath via a contact conductance, as listed in table 19, as well as by radiation; heat is also exchanged

between the thermocouple and the cold wall and the surrounding ambient environment via radiation.

L.2 Predicted Thermocouple Error 

Figures 89–91 depict contour plots showing temperatures for the highest conductance case considered (case E). Figure 89 shows the complete mode and figures 90 and 91 show successively zoomed in

shots to provide a closer view at the exposed lead and contact regions. A temperature scale is provided

for reference.
183

Figure 89. Temperature profile for TC error model; temperature profile 

corresponds to a bead contact conductance of 0.02 W/°C.
184

 Figure 90. Close-up view of exposed TC wires in vicinity of the sheath and primary 

tube cold wall; temperature profile corresponds to a bead contact 

conductance of 0.02 W/°C.
185

	 Figure 91. Close-up view of conductor between the TC bead and sheath wall; temperature 

profile corresponds to a bead contact conductance of 0.02 W/°C.

Figure 92 summarizes the percent error in temperature for the various conductance cases. The

percent error calculation technique is shown on the figure. Tbead is the temperature calculated for the

thermal node on the wire closest to the sheath surface. As can be seen from the summary data, the

temperature error can vary quite a bit for the range of conductance values evaluated. For example, the

predicted error varies from near 60% for case B to less than 20% for case D.

Based on this information, a means for improving the overall conductance and the certainty of

that conductance needs to be considered to avoid the ambiguity that might result from using this configuration of thermocouple contact and to improve the overall usefulness and reliability in sheath temperature

instrumentation. 
186

Figure 92. Comparison of TC error versus bead contact conductance.

References

Manglik, R. (technical editor), “Imperfect Metal-to-Metal Surface Contact (May 1991 Update),” Heat 

Transfer Data Book, Section 502.5, Genium Publishing Corporation, 1984.
187

APPENDIX L—DESIGN DRAWINGS: MODIFIED POWER INTERFACE

Figures 93–97 show design drawings for the modified power interface.

Figure 93. Cross-sectional view showing heater element pin formation for modified 

power interface. 
188

Positive (red)

Negative (orange)

Positive

Isolator Negative

Figure 94. Modified (conceptual) power interface showing positive and negative 

leads to heater element power terminals.
189

Positive Lead

Negative Lead

Figure 95. Modified power terminal connections for a single pin.

Figure 96. Modified power terminal connections for a single pin; axial view.
190

	 Figure 97. Modified power interface conceptual design showing series connection

between neighboring pins.
191

APPENDIX M—BASIC MATERIAL CLEANING PROCEDURES

All components that will be used in the heater assembly and test chamber should be cleaned prior

to initial checkout testing and prior to integration with any reactor test article. The following steps outline

the basic approach to cleaning various materials that may be used in these components. In all procedures,

the terms ‘wash’ and ‘rinse’ constitute full immersion in fluid. The procedures differentiate between

simple cleaning (to remove oil, dirt and grease) and pickling. During the pickling process, contaminant

metals, especially low melting point metals, are removed. In this case, the cleaning process will include

an additional immersion in aqua regia. The more aggressive aqua emporis (Hf/HNO3) immersion should

not be necessary and is not as desirable, as it actually removes some of the prime metal. The proposed

procedures are employed by Applied Methods and Materials in handling by stainless steel and refractory

metal heat pipe components and should be applicable to the heater assembly components.

Basic Cleaning Procedure – All Metals 

This Basic Cleaning procedure is to be used for parts requiring minimal cleaning (e.g. freshly

machined parts with light contamination). Note: This should not be taken to preclude the requirement for 

acid pickling if specified.

1. Wash in Freon, acetone or approved substitute until all signs of oil /grease contaminants have 

been removed.

2. Wash in hot Tri-sodium phosphate solution.

3. Soak in solution of caustic cleanser (11 parts de-ionized water, 1 part NaOH, 1 part hydrogen 

peroxide).

4. Flush in hot distilled water.

5. Repeat step 2 three times.

6. Rinse piece for 5 minutes in hot distilled de-ionized water inside an ultrasonic cleaner.

7. Rinse piece for 5 minutes in distilled water using ultrasonic vibration.

8. Rinse piece for 5 minutes in ethanol using ultrasonic agitation.
192

Stainless Steel Cleaning Procedure

This procedure applies to austenitic stainless steel in the as-milled condition. It may be used for

screen and wire cloth for heat pipes, tubing, plate, and other forms of material that although machined in

part, contain surfaces in the as-milled condition.

1. Wash part in Freon TF in an ultrasonic cleaner for at least 5 minutes to remove all signs of 

grease.

2. Wash part in ultrasonic cleaner containing a caustic solution consisting of 11 parts (by volume) 

de-ionized water, 1 part sodium hydroxide, and 1 part hydrogen peroxide for up to 5 minutes.

3. Wash part in hot de-ionized water for at least 5 minutes.

4. Repeat Steps 2 and 3 three times.

5. Wash part in hot de-ionized water in ultrasonic cleaner for at least 5 minutes.

5. Wash part in ethanol in ultrasonic cleaner for at least 5 minutes.

6. Proceed to vacuum baeout; establish a pressure of 10-6 torr and a temperature 50 °C over the 

maximum operating temperature, baeout for a minimum of 3 hours.

Nb-1%Zr Cleaning Procedure (may be adopted with modifications for Ta and W)

This procedure may be used for Nb-1%Zr material, and may be adopted for Ta, assuming that a

small amount of metal loss is tolerable. For tungsten alloys aqua regia should be used as a substitute for

the Hf/HNO3, as the rate of attack on the base metal is much slower. This procedure can be applied to

screen and wire cloth (as would be employed in heat pipe wicks), tubing, plate, and other forms of material that, although machined in part, contain surfaces in the as-milled condition.

1. Wash part in Freon TF in an ultrasonic cleaner for at least 5 minutes to remove all signs of 

grease.

2. Rapidly dunk part (initially for 5 seconds) in cleaning solution consisting of two parts (by 

volume) HNO3, two parts deionized water, and one part hydrofluoric acid (HF). Part should 

be immersed long enough to thoroughly remove scale, but not long enough to remove 

measurable amounts of material from the part.

3. Flush part in hot de-ionized water.

4. Rinse part in hot de-ionized water in ultrasonic cleaner for at least 5 minutes.

5. Rinse part in ethanol in ultrasonic cleaner for at least 5 minutes.

6. Proceed to vacuum baeout; establish a pressure of 10-8 torr and a temperature 50 °C over the 

maximum operating temperature, baeout for a minimum of 3 hours.
193

Tantalum Cleaning Procedure (the Triple Threat)

Alternately, the following procedure may be used for tantalum, assuming that a small amount of

metal loss is tolerable. This procedure can be applied to screen and wire cloth (as would be employed in

heat pipe wicks), tubing, plate, and other forms of material that, although machined in part, contain surfaces in the as-milled condition.

1. Wash part in Freon TF in an ultrasonic cleaner for at least 5 minutes to remove all signs of 

grease.

2. Rapidly dunk part (initially for 5 seconds) in cleaning solution consisting of two parts (by 

volume) sulfuric acid, one part nitric acid, and one part hydrofluoric acid (HF). Part should 

be immersed long enough to thoroughly clean, but not long enough to remove measurable 

amounts of material from the part.

3. Flush part in hot de-ionized water.

4. Rinse part in hot de-ionized water in ultrasonic cleaner for at least 5 minutes.

5. Rinse part in ethanol in ultrasonic cleaner for at least 5 minutes.

6. Proceed to vacuum baeout; establish a pressure of 10-8 torr and a temperature 50 °C over the 

maximum operating temperature, baeout for a minimum of 3 hours.

Mo-Re Alloy Cleaning Procedure

This procedure may be used for molybdenum and rhenium alloys.	 It can be applied to screen

and wire cloth (as would be employed in heat pipe wicks), tubing, plate, and other forms of material that,

although machined in part, contain surfaces in the as-milled condition.

1. Wash piece in PF solvent until all signs of grease have been removed.

2. Soak piece for 1 to 2 minutes in 1 part by volume HCl and 1 part by volume de-ionized water to 

remove residual iron surface impurities.

3. Soak piece for 5 minutes in caustic cleaning solution consisting of 11 parts by volume de-ionized water, 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 de-ionized water for at least 5 minutes.

5. Repeat step 3 and 4 three times.

6. Rinse piece for 5 minutes in ethanol inside an ultrasonic cleaner.

7. Vacuum-fire at 200 to 400 oC for 12 to 24 hours at a vacuum level of 10-6 torr or lower.
194

An additional pickling procedure can be used to chemically remove surface oxides (scales) 

and other surface contaminants by immersion in an aqueous acid solution.

Mo-Re Alloy Pickling Procedure

1. Submerge in mixture of acetic, nitric and HF in a ratio of 10:4:1.

2. Immerse in pure HCl to remove staining.

3. Rinse with de-ionized water.

4. Rinse with alcohol.

5. Following pickling, handle components with clean gloves.

Alumina

A detailed cleaning procedure for alumina is currently being investigated; this procedure is likely

to include the following steps:

1. Abrasive blasting using glass bead or alumina grit (alumina grit would produce higher erosion).

2. Cleaning in acetone, possibly in an ultrasonic cleaner.

3. Dry in a furnace (in air) up to 1000 oC.
195

APPENDIX N—PURIFIED INERT GAS CONDITIONING

The following discussion was adapted from the purified inert gas conditioning system that will 

be adopted for heat pipe life testing, as discussed in Report ER11-05-WI2-003, Section 5 (Martin, et 

al., 2005). Because similar materials and fill gases may be applied in the initial heater assembly check 

out tests, the analysis is applicable to the present discussion.

The ambient oxygen concentration is critical to the lifetime of refractory metals operated at 

increased temperature for a relatively long time. For the proposed testing, refractory metal components will be operated in either a vacuum environment or in a low-pressure, ultra high purity 

(UHP) noble gas environment (argon and helium mixed to a predetermined partial pressure of 

each gas). To achieve the desired gas purity, successive dilutions and pump down of the system are 

required. The pump down 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 generally accepted vacuum level for testing Mo-Re alloys is in the 10-6 torr range; given 

a direct pump down from air to 10-6 torr, the oxygen concentration is 0.28 ppb in the vacuum test 

environment. Hence, the target maximum oxygen concentration in the test chambers will be 0.28 

ppb at the desired operating pressure of approximately 76 torr (the set pressure used in calculating 

gas conditioning requirements).

Sizing and Performance

As discussed in Report ER11-04-WI2-1.1 (Martin, 2004), a series of 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 the air filling the 

test chambers (20.95% oxygen by volume). Calculations were performed assuming three cylindrical test chambers having approximate dimensions of 24" diameter (23.625" ID) by 36" length (total 

volume 48000 in3 (0.78 m3) for all three chambers), which will be constructed for use in heat pipe 

life tests. All heat pipe life testing will be performed with a low pressure fill of mixed helium and 

argon in the test chambers. The UHP He and Ar, purchased from Sexton Supply (Huntsville, AL), 

have a guaranteed minimum purity of 99.999%. The UHP He has a maximum oxygen content of 

3 ppm (by volume); UHP Ar has a maximum oxygen content of 1 ppm (by volume).

To obtain the oxygen concentrations shown in figure 93, the test chambers are first pumped 

from atmospheric pressure to 10-3 torr (e.g. operating only the roughing pump connected to the 

test chambers). Note that the plots correspond to a 76-torr chamber pressure, the approximate 

pressure at which the heat pipe life tests will be performed. Initially pumping the chamber to a 

pressure below the desired operating pressure results in a drastic reduction in the oxygen concentration from 209,500 ppm to 0.3 ppm at 10-3 torr; direct pump down to only 76 torr reduced the 

oxygen concentration to ~2x104 ppm, as shown in figure 89 prior to fill with ultra high purity gas 

at 76 torr. After this initial pump down, the chambers were back-filled to 76 torr of UHP He or Ar 
196

and again pumped down to 10-3 torr to further reduce the oxygen concentration in the test chambers. After just two dilutions with UHP gas (He or Ar), the test chambers reach the minimum 

purity level achievable given the purity of the supply gas. 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. The test procedure will also include an initial bakeout of 

the system (under vacuum), which will assist in driving out volatile impurities (primarily water if 

the system is thoroughly cleaned and degreased) from the test components and fill lines to reduce 

additional impurities in the system. 

1.00E-02

1.00E-01

1.00E+00

1.00E+01

1.00E+02

1.00E+03

1.00E+04

1.00E+05

0 1 2 3 4

# of Dilutions

O

2

 (ppm by volume) at 76 torr

 

1.00E-02

1.00E-01

1.00E+00

1.00E+01

1.00E+02

1.00E+03

1.00E+04

1.00E+05

0 1 2 3 4

# of Dilutions

O

2

 (ppm by volume) at 76 torr

(a) (b)

Figure 98. Oxygen concentration (ppm by volume) in the heat pipe test chambers following 

successive dilutions with (a) UHP He or (b) UHP Ar fill gas.

A smaller test chamber will be utilized for heater assembly testing (approximate diameter 

of 24", length of 72", yielding a total volume of 5400 in3 (0.09 m3). Because the same supply gases 

will be used in this chamber, the ultimate minimum purity level of the fill gas will be limited to 

the same levels shown in figure 93 (0.3 ppm O2 for He, 0.1 ppm O2 for Ar) without additional gas 

purification.

As discussed in Report ER11-04-WI2-001.1 (Design of a Refractory Metal Life Test Heat 

Pipe and Calorimeter, Section 4.2.1), the gas mixture used in the heat pipe tests will be selected 

to achieve the desired gas conductivity across the gap between the heat pipe condenser and the 

calorimeter. A similar approach will be adopted in the heater assembly tests. In this case the gas 

mixture will be selected to achieve the desired conductivity between the heater assembly and the 

representative core surface (fuel clad or core block), and between this surface and the calorimeter 

that will be designed to remove heat from the heater assembly in the absence of the full reactor 

test article (which would include active cooling). The gas mixture selected for the lower power 

(1 – 4 kW) heat pipe tests was He-32%Ar (molar fraction). The increased heat flux for the 5 kW 

heat pipe test and the minimum gap width established for fabrication purposes (0.020") requires 
197

that the Ar content in the gas mixture be reduced to 6% (molar fraction) for this test. The partial 

pressure of each gas in the mixture is given by the molar fraction multiplied by the total gas pressure. For instance, for a total pressure of 100 psig and a 32% Ar mix, the partial pressure of argon 

will be 32 psig and the partial pressure of helium will be 68 psig. Partial pressures will be used to 

establish the proper gas ratios in the helium/argon mixing procedure.

Gas Purification

Additional purification of the UHP gases will be required to meet the required oxygen concentration for long term testing of refractory metals, established at less than 0.28 ppb (comparable 

to 10-6 torr) for testing Mo-Re alloys. To accomplish this purification a SAES ambient temperature MicroTorr gas purifier, which has an advertised performance for purifying both helium and 

argon to a final oxygen concentration of 1 ppb, can be used to clean the incoming gases prior to 

entry into the test chamber. A recirculating gas system, which would include additional purification using a SAES MonoTorr point-of-use purifier, will not be considered at this time. Additional 

information on gas purification systems can be found in Report ER11-04-WI2-003.

The design of the gas mixture and purification system is such that the recirculating gas 

system can be fully isolated from the gas mixing/fill system using the hand valve G-HOV9, allowing the gas mixture bottle to be charged with an appropriate helium/argon mixture prior to test 

chamber fill. The small test chamber fill system will be connected to the main fill system just after 

the MicroTorr Purifier (G-HOV5 opens the fill line to this chamber), such that the same He and Ar 

supply bottles and initial purification system may be used for both systems. The pre-mixed fill gas 

can be mixed to the desired partial pressures of helium and argon to meet both test requirements 

using separate gas mixture bottles and recirculating gas systems.

References for Appendix O:

Martin, J.J., Reid, R.S., and Bragg-Sitton, S.M. (2004), Design of Refractory Metal Life Test Heat 

Pipe and Calorimeter, ER11-04-WI2-001.1, December 16, 2004.

Martin, J. J., Bragg-Sitton, S.M., Reid, R.S., Stewart, E. and Davis, J. (2005), Design of Refractory 

Metal Heat Pipe Life Test Environment Chamber, Cooling System, and Radio Frequency Heating 

System, NASA MSFC Report submitted to NRPCT, Report Number ER11-05-WI2-003, 

February 22, 2005.
198

APPENDIX O—ADVANCED HEATER ELEMENT DESIGNS: 

CHANGING POWER PROFILE

Figures 99–101 are conceptual designs with varying power profiles.
199

Figure 99. Conceptual design, varying power profile.
200

Figure 100. Conceptual design, varying power profile.
201

Figure 101. Conceptual design, varying power profile.
202

APPENDIX P—ADVANCED HEATER ELEMENT DESIGNS:

HEXAGONAL CROSS SECTION

Figure 102 is a conceptual design for the hexagonal cross section.
203

Figure 102. Conceptual design for hexagonal cross section, spiral-wound mandrel heater element.
204

REFERENCES

 1. VanDyke, M.K.; Houts, M.G.; Pederson, K.; et al.: “Phase I Space Fission Propulsion System 

Testing and Development Process,” Space Technology and Applications International Forum, 

American Institute of Physics, Albuquerque, NM, Vol. 552, pp. 837–842, 2001. 

 2. VanDyke, M.K.; Houts, M.G.; Godfroy, T.J.; et al.: “Phase I Space Fission Propulsion System 

Testing and Development Progress,” Space Technologies and Applications International Forum, 

American Institute of Physics, Albuquerque, NM, Vol. 608, pp. 692–697, 2002.

 3. VanDyke, M.K.; Houts, M.G.; Godfroy, T.J.; et al.: “Test Facilities in Support of High Power 

Electric Propulsion Systems,” Space Technology and Applications International Forum, American Institute of Physics, Albuquerque, NM, Vol. 654, pp. 451–456, 2003.

 4. VanDyke, M.K.; Martin, J.J.; and Houts, M.G.: “Overview of Nonnuclear Testing of the Safe, 

Affordable 30-kW Fission Engine Including End-to-End Demonstrator Testing,” NASA/TM—

2003–212930, Marshall Space Flight Center, AL, November 2003.

 5. VanDyke, M.K.: “Early Flight Fission Test Facilities (EFF-TF) To Support Near-Term Space 

Fission Systems,” Space Technology and Applications International Forum, American Institute 

of Physics, Albuquerque, NM, Vol. 699, pp. 713–719, 2004. 

 6. VanDuyn, L.B.; Poston, D. I.; and Reid, R.S.: “Analysis of the SAFE-30 Resistance-Heated 

Test Data,” Space Technologies and Applications International Forum, American Institute of 

Physics, Albuquerque, NM, Vol. 608, pp. 698–710, 2002.

 7. Godfroy, T.J.; Bragg-Sitton, S.M.; and VanDyke, M.K.: “Thermally Simulated Testing of a 

Direct-Drive Gas-Cooled Nuclear Reactor,” International Congress on Advances in Nuclear 

Power Plants (ICAPP-2003), Cordoba, Spain, Paper 3343, 2003.

 8. Bragg-Sitton, S.M.; and Forsbacka, M.: “Application of a Virtual Reactivity Feedback Control Loop in Non-Nuclear Testing of a Fast Spectrum Reactor,” International Congress on 

Advanced Nuclear Power Plants (ICAPP-2004), American Nuclear Society, Pittsburgh, PA, 

pp. 2259–2268, 2004.

9.	 ESPI Corporation, <www.espimetals.com>, March, 2005.

10.	 Incropera, F.P.; and DeWitt, D.P.: Fundamentals of Heat and Mass Transfer, 3rd Ed., 981 pp.,

John Wiley & Sons, New York, NY, 1990.

11. Arp, V.; McCarty, B.; and Fox, J.: “GASPAK®,” Ver.3.401, Cryodata, Inc., October 16, 2002.
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12. Manglik, R. (ed.): Fluid Flow Data Book, Genium Publishing Corp., Schenectady, NY, 

Section 410.2, pp.1–6, 1982.

13. Jain, P.C.: “The Prediction of Thermal Conductivity of Xenon,” Journal of Physics D: 

Applied Physics, Vol. 11(17), printed in Great Britain, The Institute of Physics, 1978.

14. “Thermal Design Data Book for Coupled Growth in Hypermonotectics and Particle 

Engulfment and Pushing by Solidifying Interfaces Sample Ampoule Cartridge Assembly,” 

Teledyne Brown Engineering Inc., Report #480MMSC0071, Huntsville, AL, June 2003.

15. Ott, L.J.; and McCulloch, R.: “Overview of Fuel Rod Simulator Usage at ORNL,” in 

Proceedings of Space Technology and Applications International Forum (STAIF-2004), 

American Institute of Physics, Albuquerque, NM, Vol. 699, pp. 703–712, 2004.

16. Ott, L.J.; and McCulloch, R.: “Electric Fuel Rod Simulator Fabrication at ORNL,” in 

Proceedings of Space Technology and Applications International Forum (STAIF-2004), 

American Institute of Physics, Albuquerque, NM, Vol. 699, pp. 720–727, 2004.

17. Stinson-Bagby, K.L.; and Fielder, R.S.: “Fiber Bragg Gratings for High-Temperature 

Thermal Characterization,” International Congress on Advances in Nuclear Power Plants 

(ICAPP-2004), Pittsburgh, PA, Paper 4299, pp. 2288–2294, 2004.

18. Fielder, R.S.; Klemer, D.; and Stinson-Bagby, K.L.: “High-Temperature Fiber Optic Sensors, 

an Enabling Technology for Nuclear Reactor Applications,” International Congress on Advances 

in Nuclear Power Plants (ICAPP-2004), Pittsburgh, PA, Paper 4298, pp. 2295–2305, 2004.

19. Bragg-Sitton, S.M.; and Webster, K.: “Application of Simulated Reactivity Feedback in 

Nonnuclear Testing of a Direct-Drive Gas-Cooled Reactor,” NASA/TM-2007-21495, 

Marshall Space Flight Center, AL, 2007. 

20. Siegel, J.R.; and Howell, R.: Thermal Radiation Heat Transfer, 4th Ed., 1072 pp., Taylor and 

Francis, New York, NY, 2002.
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NSN 7540-01-280-5500 Standard Form 298 (Rev. 2-89)

Prescribed by ANSI Std. 239-18

298-102

Unclassified Unclassified Unclassified Unlimited

Heater Development, Fabrication, and Testing: Analysis 

of Fabricated Heaters

S.M. Bragg-Sitton, R.E. Dickens, J.T. Farmer, J.D. Davis, 

M.R. Adams, J.J. Martin, and K.L. Webster

George C. Marshall Space Flight Center

Marshall Space Flight Center, AL 35812

National Aeronautics and Space Administration

Washington, DC 20546–0001

Prepared by the Propulsion Systems Department, Engineering Directorate

Unclassified-Unlimited

Subject Category 18

Availability: NASA CASI 301–621–0390

Thermal simulators (highly designed heater elements) developed at the Early Flight Fission Test Facility (EFF-TF) are used to simulate the heat from nuclear fission in a variety of reactor concepts. 

When inserted into the reactor geometry, the purpose of the thermal simulators is to deliver thermal 

power to the test article in the same fashion as if nuclear fuel were present. Considerable effort has 

been expended to mimic heat from fission as closely as possible. To accurately represent the fuel, the 

simulators should be capable of matching the overall properties of the nuclear fuel rather than simply 

matching the fuel temperatures. This includes matching thermal stresses in the pin, pin conductivities, 

total core power, and core power profile (axial and radial). This Technical Memorandum discusses 

the historical development of the thermal simulators used in nonnuclear testing at the EFF-TF and 

provides a basis for the development of the current series of thermal simulators. The status of current 

heater fabrication and testing is assessed, providing data and analyses for both successes and failures 

experienced in the heater development and testing program. 

224

M–1235

July 2008 Technical Memorandum

NASA/TM—2008–215466

fission power systems, nonnuclear testing, thermal simulators, space power, early flight fission test facility, heater elements for nuclear simulation
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