Salvatore M. Oriti

Glenn Research Center, Cleveland, Ohio

Update on Extended Operation of Stirling

Convertors in Thermal Vacuum at NASA

Glenn Research Center

NASA/TM—2006-214424

November 2006

AIAA–2006–4062
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Salvatore M. Oriti

Glenn Research Center, Cleveland, Ohio

Update on Extended Operation of Stirling

Convertors in Thermal Vacuum at NASA

Glenn Research Center

NASA/TM—2006-214424

November 2006

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Prepared for the

Fourth International Energy Conversion Engineering Conference and Exhibit (IECEC)

sponsored by the American Institute of Aeronautics and Astronautics

San Diego, California, June 26–29, 2006

AIAA–2006–4062
Acknowledgments

The work described in this paper was performed for the Science Mission Directorate (SMD) and the Radioisotope Power

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NASA/TM—2006-214424 1

Update on Extended Operation of Stirling Convertors in 

Thermal Vacuum at NASA Glenn Research Center 

Salvatore M. Oriti 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Abstract 

The U.S. Department of Energy (DOE), Lockheed Martin Space Systems (LMSS), Infinia Corporation, and 

NASA Glenn Research Center (GRC) have been developing a Stirling Radioisotope Generator (SRG) for use as a 

power system on space science missions. This generator would make use of Stirling cycle energy conversion to 

achieve higher efficiency than currently used alternatives. A test has been initiated at GRC to demonstrate 

functionality of Stirling conversion in a thermal vacuum environment over an extended period of time. The test 

article resembles the configuration of the SRG, but was designed without the requirement of low mass. Throughout 

the 8700 cumulative hours of operation, modifications to the supporting hardware were required to attain the desired 

operating conditions. These modifications, the status of testing, and the data recorded will be discussed in this paper. 

Nomenclature 

DOE Department of Energy 

GPHS General Purpose Heat Source 

GRC Glenn Research Center 

LMSS Lockheed Martin Space Systems 

LN2 liquid nitrogen 

PID proportional integral derivative 

psig pounds per square inch, gauge 

RTG Radioisotope Thermoelectric Generator 

SRG Stirling Radioisotope Generator 

TDC Technology Demonstration Convertor 

We Watt electric 

Wth Watt thermal 

I. Introduction 

Lockheed Martin Space Systems (LMSS) has been contracted by the Department of Energy (DOE) to develop a 

radioisotope-powered generator for potential use on future space science missions (ref. 1). The generator would use 

two Stirling cycle machines to convert thermal energy from General Purpose Heat Source (GPHS) modules into 

electrical power, and was thus named the Stirling Radioisotope Generator (SRG). The SRG has potential for multimission applications including Mars, lunar, and deep space. The SRG offers the potential for higher efficiency than 

Radioisotope Thermoelectric Generators (RTGs) reducing the required amount of radioisotope by a factor of 4. 

LMSS’s design of a 110-We SRG is shown in figure 1. The design consists of two Stirling convertors each being 

supplied heat by one GPHS module. The power convertors used in the SRG in figure 1 were designed and 

manufactured by Infinia, Corporation of Kennewick, Washington. The convertor was designated the Technology 

Demonstration Convertor (TDC) while being developed under contract to DOE (ref. 2) and has been further 

developed by Infinia, Corporation under subcontract to LMSS. 

NASA Glenn Research Center (GRC) has been supporting LMSS’s development of the SRG (ref. 3). 

Exploration missions that may utilize the SRG could involve continuous operation of the generator for upwards of 

14 years. One of the ongoing experiments at GRC consists of two TDCs operating continuously in a thermal vacuum 

environment. This experiment has several purposes. It provides data for extended operation of the convertors over 

thousands of hours to examine potential aging effects that would degrade the electrical power generation ability. 

Some potential sources of aging include high-temperature material creep, alternator magnet degradation, working 

fluid contamination from internal component outgassing or external impurity permeation, and mechanical wear of 

moving components. The experiment provides data to validate the various thermal models that were developed 

during the design of the test article’s heat input and rejection systems, which are described later. It also provides data 
NASA/TM—2006-214424 2

for the convertors operating in a thermal environment similar to the SRG. The thermal vacuum configuration more 

closely emulated the thermal interfaces of a flight system than typical in-air setups previously implemented at GRC 

(ref. 3). In-air setups have used cartridge heaters bonded to the hot-end of the convertor for heat input and a water 

jacket on the rejection end for cooling. In a flight system, the heat input interface would be made to a GPHS module, 

while the heat rejection interface would be coupled to a radiator. The test article was designed with these interfaces 

in mind as described in the following section. 

II. Test Article Setup 

The experiment makes use of a vacuum chamber to simulate the vacuum of space, and a liquid nitrogen (LN2) 

cold shroud to simulate spacelike sink temperatures. The test article consists of two Stirling TDCs (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 figure 2. Heat is supplied 

via an electric heater coupled to a heat collector (fig. 3(a)). The heat source is a BoralectricTM heater, composed of a 

pyrolytic graphite trace encapsulated by pyrolytic boron nitride acting as an electrical insulator. 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. The Boralectric (GE Advanced Ceramics) heater is held against the heat 

collector face by a plate. Heat is rejected from the Stirling cycle via a flange attached to radiation panels. The 

radiation panels dissipate the waste heat to the LN2 shroud surrounding the entire test article. The heat rejection 

flange is an assembly of nickel and copper sections (fig. 3(b)). The nickel section was brazed to the heat rejection 

zone of the TDC heater head. Four triangular copper sections were 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 attach. T-gon 800 thermal interface material was used between the contacting surfaces. This 

material is a 0.005-in.-thick graphite sheet that helps reduce thermal contact resistance. The radiation panels were 

fabricated out of an aluminum sheet and coated with ECP–2200 to increase surface emissivity. The hot-end and 

regenerator sections were insulated by Min-K 1302, a machinable ceramic insulation. ANSYS thermal models of the 

support hardware were constructed for design purposes. The heat collector geometry was optimized using a twodimensional, axisymmetric, parametric thermal model. The heat rejection flange was optimized using a parametric 

thermal model to determine the nickel and copper section thicknesses. The radiator panel geometry was determined 
NASA/TM—2006-214424 3

using a three-dimensional thermal model. The intent of these designs was to permit operation of the convertors at the 

nominal conditions of 650 °C maximum hot-end temperature, 80 °C rejection temperature, 220 Wth net heat input, 

65 We power output, and 6.0-mm piston amplitude. 

III. Hardware Improvements 

TDCs 5 and 6 have been operating in a thermal vacuum environment since November 2004. Several deficiencies 

in the initial test article build became apparent that prevented operation at the nominal conditions. Steps were taken 

to remedy these deficiencies, as described in the following sections. 

A. Pressure vessel cooling 

The initial test article build included no specific provision for cooling the pressure vessel other than inherent 

radiation from the as-machined stainless steel surface of the pressure vessel to the LN2 shroud. This was found to be 

insufficient and prevented operation of the convertors at design power and temperature. The initial operating point 

was limited to a hot-end temperature of 513 °C and 53 We output. At this condition, the pressure vessel temperature 

was operating at 70 °C, close to the imposed limit of 75 °C. The pressure vessel heat load originates from both the 

inefficiency of the alternator (11 Wth at full output power) and the thermal coupling to the convertor rejection zone. 

The only path of heat rejection from the pressure vessel was radiation from the surface to the LN2 shroud. To reduce 

the pressure vessel surface temperature and enable higher power operation, the surface emissivity was increased by 
NASA/TM—2006-214424 4

adding a layer of Kapton tape. The emissivity of bare stainless steel was measured at 0.15 while stainless steel with 

a layer of 0.0025-in.-thick Kapton tape was measured at 0.83; approximately 5.4 times higher. Using these values, 

analysis was performed to estimate the effect of adding Kapton tape to the pressure vessel surface. The analysis 

suggested that increasing just 10 percent of the surface’s emissivity by a factor of 5.4 would reduce the temperature 

from 70 to 50 °C; a reduction of 20 °C. However, when this idea was implemented, the pressure vessel temperature 

was only reduced by 5 °C. In light of this result, the remaining accessible area of the pressure vessel was covered 

with the Kapton tape layer as shown in figure 4, which resulted in a 13 °C drop. The addition of Kapton to the 

majority of the pressure vessel surface permitted operation at the desired conditions. At an average hot-end 

temperature of 630 °C and a power output of over 60 We per convertor, the pressure vessel temperature was 

maintained slightly below 60 °C. 

B. Rejection zone cooling 

During initial extended operation, the rejection temperature reached its design point of 80 °C prematurely, before 

full design output power was attained. Data recorded during initial operation showed the radiation panels maintained 

the rejection temperature at 80 °C at a power output level of 53 We per convertor. The original radiator panels were 

8.5 in. long (fig. 5(a)). These were replaced by panels of the same shape, but 11 in. long (fig. 5(b)). The larger 

panels provided a 29 percent increase in radiator area. Analysis was performed to estimate the effect of the added 
NASA/TM—2006-214424 5

area before installation. A thermal circuit model of the rejection hardware was constructed based on existing 

performance data with the shorter set of panels. This model was used to predict rejection temperatures over a range 

of output power levels with the added panel area. The analysis indicated the larger panels would be capable of 

maintaining the rejection temperature at 82 °C at a power output level of 70 We per convertor. After the larger 

panels were installed, the operating conditions were adjusted to achieve maximum output power by increasing the 

hot-end temperature. Data recorded at this point showed that the new panels maintained a 70 °C rejection 

temperature at 64 We output. This was in accord with the predictions generated by the analysis. 

C. Heater attachment hardware 

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

backing plate preload. The fasteners functioned properly during the initial checkout up to 575 °C, 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 Boralectric 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 (rev. 1) to improve their tolerance of the 

high temperatures. The goal of the new design was to minimize the stiffness of the fastening system thus minimizing 

the thermally induced stress. This was accomplished by increasing the fastener lengths (fig. 6). Spacers were 

required to fill in the gaps created by the longer fasteners. The thickness of these spacers was minimized to reduce 

their stiffness as well. The fasteners and spacers were made of A286 stainless steel. However, design revision 1 did 

not function as desired. After just 157 hr of operation at an average hot-end temperature of 604 °C, one of the 

spacers on the TDC 5 heater electrical connection melted. The heater attachment hardware was again redesigned 

(rev. 2). This revision utilized a spring-loaded electrical lead rather than a threaded fastener to supply electricity to 

the heater (fig. 7). This enhanced the reliability of the electrical connection during temperature cycles by better 

accommodating the thermal expansion of the heater. The structure to apply the spring load to the lead was situated 

outside the insulation to avoid the high-temperature environment. The only component exposed to the high 

temperature was the lead itself. The spring load of the electrical leads also supplemented the backing plate to hold 

the heater against the heat collector face. The material of the backing plate was changed from nickel 201 to Inconel 

718 for added strength. The material of the backing plate fasteners and spacers was changed from A286 stainless 

steel to Inconel 718. The spring-loaded electrical connection permitted operation at a maximum hot-end temperature 

of 650 °C as desired. At full input power and a 650 °C maximum hot-end temperature, the Boralectric heater 

operated at 715 °C. The spring-loaded connection has survived at this condition for over 2800 hr. 
NASA/TM—2006-214424 6

D. Helium seals 

Initial extended operation was also disrupted by helium leakage. The charge pressure of TDC 5 dropped about 

1.5 psi per day, and was adjusted almost daily. The charge pressure of TDC 6 dropped to about 0.2 psi per day, and 

was adjusted once per week. The TDCs were not hermetically sealed, and utilized o-rings to seal the helium working 

fluid at the flange joints. Thus there was some finite leakage that occurred because of permeation or imperfect o-ring 

compression. An additional leak on TDC 5 was discovered at a faulty weld on the feedthrough ring used to pass the 

piston position signal through the convertor housing (fig. 8). The faulty feedthrough ring was replaced with a spare 

and the leak rate was reduced by a factor of 2.4 as shown in table 1. 
NASA/TM—2006-214424 7

TABLE 1.—TEST ARTICLE HELIUM LEAK RATES 

Charge pressure leak rate TDC 5 TDC 6 

Before hardware mods (psig/day) –1.51 –0.21 

After hardware mods (psig/day) –0.62 –0.34 

IV. Extended Operation Performance Data 

Extended operation was initiated in November 2004. The initial operating point was limited to a hot-end 

temperature of 513 °C because of the inadequate cooling of the rejection zone and pressure vessel. An attempt was 

made to operate at a higher hot-end temperature of 604 °C while simultaneously reducing the piston amplitude to 

maintain the rejection and pressure vessel temperatures. However, the revision 1 heater electrical connection did not 

survive more than 157 hr at the higher temperature. This failure is depicted at 1083 hr (label 1) of operation on the 

performance graphs of figure 9(a) and (b). The heater electrical connection hardware was replaced with spare units, 
NASA/TM—2006-214424 8

and operation resumed at a lower hot-end temperature of 420 °C. This condition continued from 1083 to 4320 hr 

(label 3) while the revision 2 heater attachment hardware was designed and fabricated. The data between labels 2 

and 3 were erratic due to multiple shutdowns related to facility issues such as power outages, vacuum leaks, and LN2

system failures. Data between 4320 and 4637 hr (label 4) depict low-temperature operation after installation of the 

revision 2 heater hardware, the larger radiation panels, the replacement feedthrough ring, and the pressure vessel 

Kapton tape. The low-temperature checkout procedure restricted the maximum hot-end temperature to 500 °C to 

prevent oxidation of the stainless steel regenerator. This was necessary because TDC 5 was exposed to air during the 

replacement of the feedthrough ring, and a vacuum bakeout had not yet been performed. Data starting at 4637 hr 

(label 5) depict post-vacuum bakeout operation at which point it was permissible to operate at a hot-end temperature 

greater than 500 °C. The hot-end temperature was increased gradually until any hot-end thermocouple read 650 °C. 

This was achieved at 5842 hr (label 6). While increasing the hot-end temperature, it was necessary to reduce the 

piston amplitude to maintain the displacer amplitude below its maximum value. Table 2 summarizes the operating 

conditions attained at 5842 hr. Because of the use of a heat collector, an axial temperature gradient existed along the 

hot-end zone. This is the reason for the difference between average and maximum hot-end temperature. Operation at 

full output power and a maximum hot-end temperature of 650 °C yielded an average hot-end temperature of 

approximately 630 °C. True steady-state operation at the desired operating point was not achieved until after 5842 hr 

(label 6) of operation. After this point, the operating conditions were maintained at those shown in table 2. As of 

June 1, 2006, the TDCs have accumulated over 2800 hr of operation at these conditions. The fluctuation in power 

output can be attributed to the fluctuation in charge pressure. The charge pressure was periodically adjusted to 

maintain its nominal value of 365 psig to compensate for helium leakage. The fluctuating charge pressure also 

caused fluctuations in the rejection and pressure vessel temperatures since these components rely on passive cooling. 

The hot-end temperature does not exhibit these fluctuations because it was controlled by a closed proportionalintegral-derivative (PID) loop. The hot-end insulation losses were not measured. On in-air setups, the insulation 

losses have been characterized by replicating the operating temperatures on a nonoperating convertor and measuring 

the heat input. On the thermal vacuum setup, the thermal conditions could not be replicated without an operating 

convertor because of the use of passive temperature control on the rejection and pressure vessel sections. The 
NASA/TM—2006-214424 9

insulation losses in thermal vacuum were estimated using finite element analysis. The analysis indicated that at 

maximum hot-end temperature the losses total 30 Wth. This loss was applied as a constant value to all operating 

conditions. Thus the net heat input equals the measured gross heat input minus 30 Wth for all data points. The 

2800 hr of operation following label 6 are most useful for examining long-term aging effects. The data in this region 

showed no decay in power or efficiency. A linear curve fit of TDC 5 power output in this region actually has a 

positive slope of 8e-5 We per day. However, more hours are necessary to accurately evaluate the long-term stability 

of power generation. 

TABLE 2.—MAXIMUM OPERATING CONDITIONS ATTAINED 

1/17/06 8:42 AM TDC 5 TDC 6 

Hot end temperature, °C 629.8 629.6 

Rejection temperature, °C 69.9 69.0 

Pressure vessel temperature ,°C 58.9 55.8 

Charge pressure, psig 363.5 364.1 

Piston amplitude, mm 5.44 5.40 

Power output, We 63.6 62.2 

Cold wall temperature, °C –168.1 

V. Summary 

An experiment has been initiated at GRC to investigate operation of two Stirling power convertors in a thermal 

vacuum environment. Performance data gathered thus far has been presented, as well as the difficulties encountered 

with the supporting hardware. As of June 1, 2006, TDCs 5 and 6 have accumulated over 8700 hr of extended 

operation in the thermal vacuum configuration. After modifications to the initial support hardware were made, the 

operating point was adjusted as close as possible to the design point. The actual operating conditions achieved were 

630 °C average hot-end, 70 °C rejection, 5.45-mm piston amplitude, and 64 We power output. 

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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WBS 138494.04.01.01

15

Update on Extended Operation of Stirling Convertors in Thermal Vacuum at

NASA Glenn Research Center

Salvatore M. Oriti

Stirling; SRG; Radioisotope; Thermal vacuum

Unclassified -Unlimited

Subject Category: 20 or

Prepared for the Fourth International Energy Conversion Engineering Conference and Exhibit (IECEC) sponsored by the

American Institute of Aeronautics and Astronautics, San Diego, California, June 26–29, 2006. Responsible person,

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

The U.S. Department of Energy (DOE), Lockheed Martin Space Systems (LMSS), Infinia Corporation, and NASA Glenn

Research Center (GRC) have been developing a Stirling Radioisotope Generator (SRG) for use as a power system on space

science missions. This generator would make use of Stirling cycle energy conversion to achieve higher efficiency than

currently used alternatives. A test has been initiated at GRC to demonstrate functionality of Stirling conversion in a thermal

vacuum environment over an extended period of time. The test article resembles the configuration of the SRG, but was

designed without the requirement of low mass. Throughout the 8700 cumulative hours of operation, modifications to the

supporting hardware were required to attain the desired operating conditions. These modifications, the status of testing, and

the data recorded will be discussed in this paper.