Salvatore M. Oriti

Glenn Research Center, Cleveland, Ohio

Extended Operation of Stirling Convertors

in a Thermal Vacuum Environment

NASA/TM—2006-213984

February 2006

AIAA–2005–5747
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Salvatore M. Oriti

Glenn Research Center, Cleveland, Ohio

Extended Operation of Stirling Convertors

in a Thermal Vacuum Environment

NASA/TM—2006-213984

February 2006

National Aeronautics and

Space Administration

Glenn Research Center

Prepared for the

Third International Energy Conversion Engineering Conference

sponsored by the American Institute of Aeronautics and Astronautics

San Francisco, California, August 15–18, 2005

AIAA–2005–5747
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NASA/TM—2006-213984 1 

Extended Operation of Stirling Convertors in a 

Thermal Vacuum Environment 

Salvatore M. Oriti 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

A 110 watt Stirling Radioisotope Generator (SRG110) is being developed for potential use on future NASA 

exploration missions. The development effort is being performed by Lockheed Martin under contract to the 

Department of Energy (DOE). Infinia, Corp. supplies the free-piston Stirling power convertors, and NASA Glenn 

Research Center (GRC) provides support to the effort in a range of technologies. This generator features higher 

efficiency and specific power compared to alternatives. One potential application for the generator would entail 

significant cruise time in the vacuum of deep space. A test has been initiated at GRC to demonstrate functionality of 

the Stirling convertors in a thermal vacuum environment. The test article resembles the configuration of the 

SRG110, however the requirement for low mass was not considered. This test demonstrates the operation of the 

Stirling convertors in the thermal vacuum environment, simulating deep space, over an extended period of 

operation. The status of the test as well as the data gathered will be presented in this paper. 

Nomenclature 

LMA Lockheed Martin Astronautics 

DOE Department of Energy 

We Watt electric 

GPHS General Purpose Heat Source 

SRG110 Stirling Radioisotope Generator 110 W 

RTG Radioisotope Thermoelectric Generator 

TDC Technology Demonstration Convertor 

GRC Glenn Research Center 

LN2 Liquid Nitrogen 

TRL Technology Readiness Level 

I. Introduction 

 Lockheed Martin Astronautics (LMA) has been contracted by the Department of Energy (DOE) to develop a 

radioisotope powered generator for potential use in future exploration missions1

. The generator would nominally 

produce 110 We from two Plutonium-238 General Purpose Heat Source (GPHS) modules. The generator would use 

two Stirling cycle machines to convert thermal energy from the GPHS modules into electrical power, and has thus 

been named 110 watt Stirling Radioisotope Generator (SRG110). The SRG110 has potential multi-mission 

applications for Mars surface (i.e., in atmospheres) as well as deep space missions. The SRG110 offers the potential 

for higher efficiency and specific power than Radioisotope Thermoelectric Generators (RTG’s). LMA’s design for 

the SRG110 is shown in Fig. 1. The design consists of two 55 We Stirling convertors each being supplied heat by a 

GPHS module. 

The power convertors used in the SRG110 were designed and manufactured by Infinia, Corp. of Kennewick 

WA. The convertor was designated the Technology Demonstration Convertor (TDC) while being developed under 

contract to DOE2

. A schematic of the 55 We TDC is shown in Fig. 2. Infinia, Corp. is presently a subcontractor to 

LMA as the development of the SRG110 continues. 

 NASA Glenn Research Center (GRC) has been supporting LMA’s development of the SRG1103

. Exploration 

missions that may utilize the SRG110 could involve continuous operation of the generator for upwards of 14 years. 

One of the ongoing experiments at GRC has two TDC’s being operated in a thermal vacuum environment. The 

purpose of this experiment is to provide data for TDC’s operating in a configuration and environment similar to the 

SRG110. 

The experiment makes use of a vacuum chamber (Fig. 3) to simulate the vacuum of space, and a liquid nitrogen 

(LN2) cold shroud to simulate space-like temperatures. Per request of NASA Headquarters, the TDC’s are planned 

to be operated in the thermal vacuum configuration for three years. 
NASA/TM—2006-213984 2 

 

HEAT SOURCE

STIRLING 

CONVERTER

ASSEMBLY

CONTROLLER

COLD SIDE

ATTACHMENT

S/C INTERFACE (4)

HOUSING

GAS MANAGEMENT

VALVE

PRESSURE RELIEF

DEVICE

INSULATION

SYSTEM

ALTERNATOR

COOLING STRAPS

HEAT SOURCE

SUPPORT

ASSEMBLY

Figure 1.—Schematic of the SRG110 design by LMA.

Figure 3.—Vacuum facility 67 at NASA GRC.

Linear Alternator – converts 

linear motion to electric power

Piston Flexures - support and 

spring the Piston/Alternator 

assembly Non-Contacting Seals -

isolate gas working spaces

Displacer Flexures - support 

and spring the displacer 

Displacer – shuttles the working 

fluid between expansion and 

compression spaces 

Cold End Hx -

rejects heat from 

the working fluid

Moving 

Components in 

blue

Linear Alternator Mover, 

reciprocates with the piston

Regenerator - improves 

engine cycle efficiency

Hot End Hx - heats 

the working gas

Power Piston -

reciprocates and 

drives the alternator Displacer Rod - drives the 

displacer

Linear Alternator – converts 

linear motion to electric power

Piston Flexures - support and 

spring the Piston/Alternator 

assembly Non-Contacting Seals -

isolate gas working spaces

Displacer Flexures - support 

and spring the displacer 

Displacer – shuttles the working 

fluid between expansion and 

compression spaces 

Cold End Hx -

rejects heat from 

the working fluid

Moving 

Components in 

blue

Linear Alternator Mover, 

reciprocates with the piston

Regenerator - improves 

engine cycle efficiency

Hot End Hx - heats 

the working gas

Power Piston -

reciprocates and 

drives the alternator Displacer Rod - drives the 

displacer

Figure 2.—Infinia, Corp. 55 We Technology Demonstration Convertor (TDC).
NASA/TM—2006-213984 3 

II. Test Setup 

The test article consists of two Stirling TDC’s (#5 and #6) mounted in the dual-opposed configuration. This 

means that the two convertors are rigidly attached through their mounting flanges and their alternator sections are 

oriented toward the center as shown in Fig. 4. This configuration allows for dynamically balanced operation 

because the pistons move synchronously, but in opposite directions. Heat is provided to the Stirling cycle via an 

electric heater coupled to a heat collector (Fig. 5a). The heat collector is a cup-shaped component machined from 

nickel 201 that directs the thermal energy from the rectangular heater to the circumferential heat input zone of the 

TDC heater head. Heat is rejected from the Stirling cycle via a flange with attached radiation panels. The radiation 

panels then dissipate the heat to the liquid nitrogen shroud surrounding the entire test article. 

The heat rejection flange is an assembly of nickel and copper sections (Fig. 5b). The nickel section is brazed to 

the heat rejection zone of the TDC heater head. Four triangular copper sections are then assembled onto each nickel 

section to increase the effective thermal conductivity of the flange. The copper pieces also form the four flat sides to 

which the radiation panels are attached. T-gon® 800 thermal interface material is installed between the contacting 

TDC’s

Radiation 

Panels Min-K 

Insulation

Heat Rejection 

Flange

Figure 4.—TDC’s in thermal vacuum configuration.

Heat Collector

Boralectric Heater

Backing Plate

TDC Heater Head

 

TDC Heater Head

Copper

Nickel Boralectric 

Heater

 (a) (b) 

Figure 5.— (a) Heat input hardware. (b) Heat rejection hardware.
NASA/TM—2006-213984 4 

surfaces. This material consists of a 0.005 inch thick graphite sheet that helps reduce thermal contact resistance. 

The radiation panels are made of aluminum and coated with MIL-1-8625 to increase their surface emissivity. The 

hot-end and regenerator sections are insulated by Min-K® 1302, a machinable ceramic insulation manufactured by 

Thermal Ceramics, Inc. 

III. Measured Versus Predicted Data 

 A system-level thermal analysis was performed to obtain temperature estimates of various components of the test 

article during steady-state operation. The ANSYS® three-dimensional model consisted of one convertor with its 

supporting hardware (radiation panels, heat rejection flange, heat collector, and insulation.). Figure 6a illustrates the 

applied boundary conditions used in the analysis. A heat load of +235 W was applied to the face of the heat 

collector, evenly distributed over the area occupied by the heating element. A temperature of 650°C was applied at 

the heat collector – heater head interface. A heat load of -220 W was applied to the inner diameter of the heater 

head, along the region occupied by the hot heat exchanger. A heat load of +170 W was applied to the inner diameter 

of the heater head along the region occupied by the cold heat exchanger. The heat generated in the alternator was 

modeled by applying a heat load of 10 W to the inner diameter of the alternator housing (alternator not shown). 

Each heat load boundary condition was applied in the form of heat flux applied over each respective area. The heat 

flux values were calculated by dividing each heat load by the appropriate area. A space node at a temperature of 

-109°C was used to model the LN2 shroud. Theses boundary conditions were determined from full power operating 

parameters measured during previous TDC testing at GRC. 

 Figure 6b compares data from the thermal analysis to data gathered during a comparable operating point in 

thermal vacuum. The axial temperature gradient along the heat collector agrees closely with the prediction. The 

measured rejection temperatures were 4% higher than those predicted while the radiator panel temperatures were 

0.7% lower*

. This suggests the thermal contact resistances between mating components of the heat rejection 

hardware is not negligible, as was assumed in the thermal analysis. The pressure vessel temperatures differ greatly 

from the predictions. The analytical results suggested the aft end of the pressure vessel (PV top in Fig. 6b), would 

be 63°C cooler than the fore end (PV bottom in Fig. 6b). The measured data shows a negligible difference between 

these two temperatures. Also, the maximum measured pressure vessel temperature is 6% higher than the prediction. 

One possible explanation for the inaccuracy is that the motion of the helium working fluid in the pressure vessel was 

not modeled. The helium in the bounce space experiences an oscillating motion which enhances heat transfer 

between the aft and fore ends of the pressure vessel. This may account for the reduced temperature gradient along 

the axial length. The minimum attainable LN2 shroud temperature was 38% lower than that used for the analysis, 

suggesting that the shroud has a larger heat dissipation capacity than predicted. 

 

*

 Temperature comparisons in this section are made using the absolute temperature scale. 

-109°C

-109°C

10W 170W 220W 235W

 

Predicted

Measured

Cold Shroud

-109°C Avg.

-171°C Avg.

PV Top

-13.5°C

69.4°C

PV 

Bottom

49.3°C

68.3°C Rejection 

Temp 2

65.5°C

78.7°C

Rejection 

Temp 1

65.5°C

79.1°C

Rad. Panel 

Temps

55.2°C Avg.

52.8°C Avg.

Gradient

26°C

25.7°C

 (a) (b)

Figure 6.—(a) Thermal FEA boundary conditions. (b) Predicted vs. measured data. 
NASA/TM—2006-213984 5 

Electrical 

Connections

Backing Plate Fasteners 

(hold heater against 

heat collector face)

Figure 7.—Hot-end fastener redesign.

IV. Hot-End Fasteners 

 The initial test article design utilized molybdenum fasteners for the heater electrical connection and for the 

backing plate preload. The fasteners functioned during the initial checkout but lost preload upon return to ambient 

temperature. Analysis indicated the failure was caused by thermally induced stresses from differing thermal 

expansion coefficients. The thermal expansion coefficient of the heater is approximately 7.4 times that of 

molybdenum. The heater and fasteners both experienced a temperature change greater than 600°C. This induced a 

stress higher than the yield strength of molybdenum at that temperature, resulting in a permanent deformation. The 

connections were redesigned to improve their tolerance of the high temperatures. The goal of the new design was to 

minimize the stiffness of the fastening system and thus minimize the thermally induced stress. This was 

accomplished by maximizing the fastener lengths and minimizing the thickness of the spacers that filled each gap 

created by the longer fasteners (Fig. 7). The maximum length of the fasteners was restricted to 0.75 inches by the 

physical space available under the insulation cap. The spacers were fabricated from solid 1⁄4 inch OD rod, with the 

minimum wall thickness limited to 0.015 inches by machining capability. The fasteners and spacers were made of 

A286 stainless steel; a precipitation hardened alloy exhibiting good yield strength and rupture life at the temperature 

range of interest. 

V. Extended Operation Performance Data 

The initial operating point (Fig. 8, label A) was limited to a 513°C hot-end temperature by several factors. It 

became evident early in thermal vacuum testing that the pressure vessel cooling was insufficient. At point A, the 

pressure vessel temperature exceeded 70°C and came close to the imposed limit of 75°C. The displacer was 

operating at its design amplitude while the piston was operating at only 5.75 mm; slightly less than its design 

amplitude of 6.00 mm. The rejection temperature also reached its design point of 80°C prematurely, indicating 

insufficient heat rejection capacity. All of these parameters prevented operation at the hot-end design temperature of 

650°C. This scenario was alleviated by reducing the piston amplitude to 5.45 mm via the variable zener-diode 

controller while simultaneously increasing the hot-end temperature. This combination allowed an increase in hotend temperature to 604°C while maintaining the rejection and pressure vessel temperatures (Fig. 8, label B). The 

power output increased slightly due to the greater Carnot efficiency from the higher temperature ratio. The hot-end 

fasteners survived for 157 hours at the 604°C temperature, at which time one of the spacers on the TDC 5 heater 

electrical connection melted, precipitating an automatic shutdown. This is depicted in the performance graphs by 

the sharp decline of all parameters shortly after point B. The hot-end fasteners were replaced and operation 

resumed. The data between points B and C was erratic due to multiple shutdowns related to facility issues such as 

power outages, vacuum leaks, and LN2 system failures. At point C, steady operation was achieved, but at a reduced 

hot-end temperature of approximately 430°C to prevent another fastener failure. The replacement fasteners have 

survived for over 1860 hours at this temperature, but there is evidence of fastener failure on the TDC 6 heater. The 

real-time average hot-end temperature plot of TDC 6 fluctuates by ±0.5°C in a sinusoidal fashion at a frequency of 
NASA/TM—2006-213984 6 

1/80th Hz. This may be the result of loss of preload of the backing plate fasteners. This scenario is confirmed by 

the fluctuations in heater electrical resistance. The resistance of the pyrolytic graphite element decreases linearly 

with temperature. The electrical resistance also fluctuated, indicating a temperature change of the heater in accord 

with loss of thermal connectivity to the heat collector face. It is unknown why the cycle repeats rather than reaching 

a static deflection. Steady-state operation was also disrupted by helium leakage. The charge pressure of TDC 5 

dropped about 1.3 psig per day, and required daily adjustment. The charge pressure of TDC 6 dropped at a slower 

rate, and required adjustment only once per week. The leak on TDC 5 has been attributed to a faulty weld on the 

feedthrough ring used to pass the piston position signal through the convertor housing (Fig. 9). Plans have been 

implemented to repair the deficiencies in the test article. The radiation panels will be replaced with larger units to 

allow more heat dissipation at the same 80°C rejection temperature. The heater electrical connection will be 

replaced with an entirely new mechanism using a spring loaded lead rather than a threaded fastener. This will 

enhance the reliability of the electrical connection during temperature cycles by better accommodating the thermal 

expansion of the heater. The material of the backing plate and its fasteners will be changed to Inconel 718® to better 

maintain clamping force at the high temperature. The pressure vessel will be modified to increase its emissivity by 

adding a layer of Kapton® tape. This will allow the same amount of heat to be dissipated at a lower surface 

temperature. The faulty feedthrough ring will be replaced with a spare unit. 

Avg. Hot-end Temp

100

200

300

400

500

600

700

0 500 1000 1500 2000 2500 3000 3500

Hours of Operation

Temperature, °C

TDC 5

TDC 6

A

B

C

Power OUT

0

10

20

30

40

50

60

0 500 1000 1500 2000 2500 3000 3500

Hours of Operation

Power, W

TDC 5

TDC 6 A

B

C

Avg. Rejection Temp

0

10

20

30

40

50

60

70

80

90

0 500 1000 1500 2000 2500 3000 3500

Hours of Operation

Temperature, °C

TDC 5 A TDC 6

B

C

Avg. Pressure Vessel Temp

0

10

20

30

40

50

60

70

80

0 500 1000 1500 2000 2500 3000 3500

Hours of Operation

Temperature, °C

TDC 5

A TDC 6

B

C

Figure 8.—Extended operation performance data.
NASA/TM—2006-213984 7 

Alternator housing

Piston housing

Welded 

feedthrough

Figure 9.—Feedthrough ring.

VI. Conclusion 

An experiment has been initiated at GRC to investigate operation of two TDC’s in a thermal vacuum 

environment. Performance data gathered thus far has been presented, as well as the deficiencies encountered with 

the supporting hardware. As of July 1st 2005, TDC’s #5 and #6 have accumulated 3200 hours of extended operation 

in the thermal vacuum configuration. After modifications to the supporting hardware have been made, the operating 

point will be adjusted as close as possible to the design point (650°C hot-end, 80°C cold-end, 6.00 mm piston 

amplitude). A milestone has been established to accumulate 5000 hours of operation in thermal vacuum by the end 

of fiscal year 2005. The overall goal of the experiment is to accumulate 25,000 hours of operation in thermal 

vacuum. 

References 1

"Stirling Radioisotope Generator for Mars Surface and Deep Space Missions", Robert D. Cockfield and Tak S. 

Chan, 2002 Intersociety Energy Conversion Engineering Conference, Washington, DC, Paper No. 20188. 

2

"Stirling Convertor Performance Mapping and Test Results for Future Radioisotope Power Systems", Songgang 

Qiu and Allen A. Peterson, 2003 International Energy Conversion Engineering Conference, Porthmouth, VA. 

3

"Extended Operation of Stirling Convertors", Mary Ellen Roth, Jeffrey G. Schreiber, and Stephen V. Pepper, 2004 

Second International Energy Conversion Engineering Conference, Providence, RI. Paper No. AIAA 2004-5508
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13

Extended Operation of Stirling Convertors in a Thermal Vacuum Environment

Salvatore M. Oriti

Space power experimentation; Stirling; SRG; Thermal vacuum

Unclassified -Unlimited

Subject Category: 20

Prepared for the Third International Energy Conversion Engineering Conference sponsored by the American Institute

of Aeronautics and Astronautics, San Francisco, California, August 15–18, 2005. Responsible person, Salvatore M.

Oriti, organization code RPT, 216–433–2066.

A 110 watt Stirling Radioisotope Generator (SRG110) is being developed for potential use on future NASA exploration

missions. The development effort is being performed by Lockheed Martin under contract to the Department of Energy

(DOE). Infinia, Corp. supplies the free-piston Stirling power convertors, and NASA Glenn Research Center (GRC)

provides support to the effort in a range of technologies. This generator features higher efficiency and specific power

compared to alternatives. One potential application for the generator would entail significant cruise time in the vacuum of

deep space. A test has been initiated at GRC to demonstrate functionality of the Stirling convertors in a thermal vacuum

environment. The test article resembles the configuration of the SRG110, however the requirement for low mass was not

considered. This test demonstrates the operation of the Stirling convertors in the thermal vacuum environment, simulating

deep space, over an extended period of operation. The status of the test as well as the data gathered will be presented in this

paper.