Paul C. Schmitz

Power Computing Solutions, Inc., Avon, Ohio

L. Barry Penswick

Sest, Inc., Middleburg Heights, Ohio

Richard K. Shaltens

Glenn Research Center, Cleveland, Ohio

Lunar Surface Stirling Power Systems

Using Isotope Heat Sources

NASA/TM—2010-216352

July 2010

AIAA–2009–4578
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Paul C. Schmitz

Power Computing Solutions, Inc., Avon, Ohio

L. Barry Penswick

Sest, Inc., Middleburg Heights, Ohio

Richard K. Shaltens

Glenn Research Center, Cleveland, Ohio

Lunar Surface Stirling Power Systems

Using Isotope Heat Sources

NASA/TM—2010-216352

July 2010

AIAA–2009–4578

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Prepared for the

Seventh International Energy Conversion Engineering Conference (IECEC)

sponsored by the American Institute of Aeronautics and Astronautics

Denver, Colorado, August 2–5, 2009
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The work described in this paper was performed for the Exploration Systems Mission Directorate (ESMD),

which provided funding for these projects.

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Level of Review: This material has been technically reviewed by technical management. 
NASA/TM—2010-216352 1 

Lunar Surface Stirling Power Systems 

Using Isotope Heat Sources 

Paul C. Schmitz 

Power Computing Solutions, Inc. 

Avon, Ohio 44011 

L. Barry Penswick 

Sest, Inc. 

Middleburg Heights, Ohio 44130 

Richard K. Shaltens 

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Abstract 

For many years, NASA has used the α decay of plutonium-238 (Pu-238) (in the form of the General 

Purpose Heat Source (GPHS)) as a heat source for Radioisotope Thermoelectric Generators (RTGs), which 

have provided electrical power for many NASA missions. While RTGs have an impressive reliability record 

for the missions in which they have been used, their relatively low thermal to electric conversion efficiency 

and the scarcity of plutonium-238 (Pu-238) has led NASA to consider other power conversion technologies. 

NASA is considering returning both robotic and human missions to the lunar surface and, because of the 

long lunar nights (14.75 Earth days), isotope power systems are an attractive candidate to generate electrical 

power. NASA is currently developing the Advanced Stirling Radioisotope Generator (ASRG) as a candidate 

higher efficiency power system that produces greater than 160 W with two GPHS modules at the beginning 

of life (BOL) (~32% efficiency). The ASRG uses the same Pu-238 GPHS modules, which are used in RTG, 

but by coupling them to a Stirling convertor provides a four-fold reduction in the number of GPHS modules. 

This study considers the use of americium-241 (Am-241) as a substitute for the Pu-238 in Stirlingconvertor-based Radioisotope Power Systems (RPS) for power levels from tens of W to 5 kWe. The 

Am-241 is used as a substitute for the Pu-238 in GPHS modules. Depending on power level, different 

Stirling heat input and removal systems are modeled. It was found that substituting Am-241 GPHS modules 

into the ASRG reduces power output by about one-fifth while maintaining approximately the same system 

mass. In order to obtain the nominal 160 W of electrical output of the Pu-238 ASRG requires 10 Am-241 

GPHS modules. Higher power systems require changing from conductive coupling heat input and removal 

from the Stirling convertor to either pumped loops or heat pipes. Liquid metal pumped loops are considered 

as the primary heat transportation on the hot end and water pumped loop/heat pipe radiator is considered for 

the heat rejection side for power levels above 1 kWe. 

Nomenclature 

GPHS General Purpose Heat Source 

MLI multi-layer insulation 

RPS Radioisotope Power Systems 

RTG Radioisotope Thermoelectric Generator 

FeNdB iron neodymium boron

SmCo samarium cobalt 

ASRG advanced Stirling radioisotope generator 

BOL Beginning-of-Life 
NASA/TM—2010-216352 2 

EOL End-of-Life 

DOE Department of Energy 

Mar-M247 super alloy 

SOA State-of-the-Art 

LM Lockheed Martin Corp. 

Introduction 

NASA is currently studying both manned and unmanned missions to the lunar surface. The lunar 

surface is a challenging environment for power systems due to its wide fluctuations in temperature and its 

long nighttime period (29.5 earth day/night cycle). Because of these very long night periods, any solararray-based system must have an energy storage system that provides power during this 14.75-day period. 

Isotope power systems offer an attractive alternative because of their relatively constant power output 

over the lunar day/night cycle. Although plutonium-238 (Pu-238) has many attractive features, it is 

possible to use other isotopic heat sources with Stirling convertors to generate electrical power. This study 

considered replacing the Pu-238 in the General Purpose Heat Source (GPHS) with americium-241 

(Am-241) and then conceptually integrating these new heat sources with Stirling convertors. Strontium90 (Sr-90) is also considered for the Advanced Stirling Radioisotope Generator (ASRG) to show the 

effect on power output due to the relatively short half-life of the Sr-90. 

Isotope Candidates 

The GPHS has been used for many deep space missions when there is a lack of adequate solar 

illumination to power solar cells (Ref. 1). It is a Department of Energy (DOE) standardized thermal 

source that produces approximately 250 W of thermal power at the beginning of life (BOL). Dimensions 

of a GPHS module are shown in Table 1 (Ref. 2). 

TABLE 1.—GPHS DIMENSIONS 

Height 5.92 cm 

Width 9.32 cm 

Length 9.95 cm 

PuO2 is the molecule of Pu-238 that is used as the fuel for the GPHS. PuO2 is placed in four iridium 

capsules and surrounded by a graphite shell to form each GPHS module (Refs. 3 to 6). Figure 1 shows a 

drawing of a GPHS. Pu-238 is attractive because most of its radioactive decay energy comes from an 

alpha emission and it has a long half-life (87 yr). Relatively low amounts of neutron emission come from 

both spontaneous fission and (α,n) reactions, which result from the interactions of the high energy alpha 

particles with low atomic mass materials. The iridium capsule prevents the alpha particles from leaving 

the fuel pellet (and interact with the surrounding graphite) but interactions with both O17 and O18 in the 

PuO2 mixture does produce some neutron flux. Production of Pu-238 is commonly done by neutron 

irradiation of neptunium-237 (Np-237) in a high-flux reactor. The product of this irradiation is Np-238 

that decays (2.117 day half-life) via beta emission into Pu-238. Currently we buy our Pu-238 from the 

Russians when used with the GPHS–RTG (Radioisotope Thermoelectric Generator) systems. 

Because of the scarcity of Pu-238 (Ref. 7) other candidates have been considered (Ref. 8). 

Reference 6 outlines the expected power output, mass, and other physical characteristics when different 

isotopes replace the PuO2 in GPHS modules. Although there are many possible isotope replacements, 

Am-241 and Sr-90 are two candidates each with greater availability than Pu-238. Notable differences 

between the three isotopes are half-life, power density, and radiation emission. Am-241 has a specific 

power that is about one-fifth that of Pu-238 while its half-life is about 5 times as great (87.5 vs. 

433 yr). Both Pu-238 and Am-241 produce relatively flat power profiles over the assumed 10-yr 
NASA/TM—2010-216352 3 

Figure 1.—GPHS Module. 

mission. Am-241 does have a gamma emission with photon energy of 59.5 KeV that can be easily 

shielded. Am-241 is produced from transuranic waste and is readily available. 

Sr-90 is not practical as a direct substitute for Pu-238 without additional shielding due to the emission 

of secondary gamma rays (2 MeV from the yttrium-90 (Y-90)) and the short-lived daughter nuclei from 

the beta decay of Sr-90. Additionally, Sr-90 with its relatively short half-life when compared with Pu-238 

or Am-241 loses approximately one-third of its thermal output over the 10-yr mission life used in this 

study. 

Many different molecules can be made from the isotopes considered above. In this study GPHS 

performance is estimated with Am-241 in the form of AmO2 and Sr-90 in the form of SrO. Other possible 

candidates can be found in Reference 9. Table 2 provides a summary of the isotope candidates. 

Figure 2 shows the heat output as a function of time for a GPHS module filled with the isotopic 

compounds PuO2, AmO2, and SrO. Sr-90 has a much higher specific power (W/g) but its lower density 

combined with the volume available in the GPHS module limits the thermal output. SrO and AmO2 when 

placed in GPHS modules have very similar power outputs of 42.8 and 56.9 W, respectively. Because of 

their similar thermal power outputs and AmO2 lower gamma radiation, Am-241 is selected as the isotope 

for further study. A brief analysis using Sr-90 is provided to show some of the difficulties in integrating a 

power source where its half-life becomes a significant fraction of the operational life of the system. For 

this analysis heater head temperatures were limited to 1123 K, which is consistent with the current 

State-of-the-Art MarM247 super alloy currently being used in the ASRG. All of the isotope compounds 

considered here have melting points well above this limit and should have no impact on the selection 

criteria. 

TABLE 2.—ALTERNATIVE ISOTOPES SUMMARY 

Isotope Compound 

form 

Half-life, 

yr 

Melting 

temperature, 

K 

W per cc, 

W/cm3

W per, 

g 

Class of 

emitter 

Heat generated 

BOL, 

W 

Pu-238 PuO2 87.7 2673 3.9 0.39 α 250 

Am-241 AmO2 432.7 2273 0.088 0.097 α a56.9 

Sr-90 SrO 28.8 1661 0.66 0.128 β a42.8 a

Based on keeping iridium capsule volume constant and filling with isotopic compound. 
NASA/TM—2010-216352 4 

Figure 2.—Thermal power output of a GPHS with various isotopes. 

Figure 3.—ASRG.

System Layout 

In the ASRG, the GPHS modules are placed at the two ends of the ASRG and the heat is conductively 

coupled into the heater head of the Stirling convertor (see Fig. 3). Heat is removed via a solid conductor 

from the cold end of the Stirling out to the housing/radiator surface. As electrical power requirements 

grow, additional heat and GPHS modules are required. For heat input, one-way to arrange the additional 

GPHS modules is to circle the Stirling convertor heater head (see Fig. 4 and more detailed discussion in 

the Stirling Convertors section of this paper). Conductive coupling is practical until the resistive losses 

and mass of the conductive interface become prohibitive. For this study, the maximum number of GPHS 

modules for conductive coupling was limited to six per Stirling convertor. 

For power levels when conductive coupling is no longer practical, a pumped loop is used to transport 

heat to and from the Stirling convertor. On the hot side, NaK is pumped over a GPHS to a NaK heat 

exchanger and then to the Stirling convertor using an electromagnetic pump. Heat is removed via a 

flowing water loop. Notable differences between the Pu-238 and the Sr-90 and Am-241 are the number of 

GPHS modules required to produce any specific power level and its decay profile. Table 3 shows the 

method of heat transport as a function of the number of GPHS modules and isotope. 
NASA/TM—2010-216352 5 

 

TABLE 3.—HEAT SOURCE AND HEAT REJECTION INTERFACE METHOD 

Number of GPHS 

Modules per convertor 1 2 to 6 >6 

Power level, W 

Pu-238 Am-241/ 

Sr-90 Pu-238 Am-241 Pu-238 Am-241 

160 ~20 ~300 to 

1000 60 to 240 >1120 >210 

GPHS to Stirling 

interface Cond Cond Cond Cond PL PL 

Stirling to radiator 

interface Cond Cond HP or PL Cond PL PL 

Cond = Conductive Coupling 

HP = Heat Pipe Heat Transport 

PL = Pumped Loop 

GPHS to Stirling Coupling 

Conductive Coupling 

The ASRG GPHS module is located forward of the heater head and coupled to it via a “hot shoe” that 

acts as the thermal interface between the two components (Ref. 10). The heat transfer mechanism is used 

solely via conduction through the hot shoe material that is generally a nickel-200 (Ni-200) series alloy 

because of their relatively high thermal conductivity. 

Conductive coupling is practical provided the heat flux and distance between the GPHS and Stirling 

is not too great. Figure 5 shows that as the number of GPHS modules grows, the distance from the surface 

of the GPHS module to the heater head must also grow. This additional distance and its associated 

temperature drops and the mass of the conductive plate and insulation eventually make a conduction 

coupling prohibitive. 

Pumped Loop Integration of Heat Source 

For Pu-238 with pumped loop heat transport systems, three power levels (1, 2, and 5 kWe) were 

modeled using both Pu-238 and Am-241. Figure 6 shows the fundamental component layout for this 

system (Ref. 11). 

The proposed conceptual LM heat exchanger, shown in Figure 7, is configured to take advantage of 

the relatively low heat flux levels produced by the alternative isotope GPHS modules. The modules are 

clamped against a flat plate Heat Exchanger through which the LM flows. The internal LM flow passage 

geometry is optimized via the passages formed in a machined plate that is brazed between to the exterior 

face sheets that are in contact with the GPHS modules. It is important to note that the GPHS modules are 

not exposed to the LM flow and as such the potential venting/out gassing of the GPHS module cannot 

Figure 4.—A three GPHS module/ 

convertor configuration. 

Figure 5.—Five and six GPHS 

geometries for a SRG. 
NASA/TM—2010-216352 6 

Figure 6.—Pumped loop heat transport system configuration. 

Figure 7.—GPHS module/LM HX configuration. 
NASA/TM—2010-216352 7 

contaminate the LM flow circuit. Since the structural loads are relatively low, the materials in contact 

with the LM working fluid can be fabricated from fully compatible materials such as stainless steel-316 

(SS-316). In higher stressed areas this compatible material can be clad to the desired higher strength 

materials such as Ni-based superalloys. 

At the interface with the convertor hot-end HX, the LM working fluid would flow through passages 

configured to ensure the proper heat transfer between the fluid and the Stirling convertor HX wall. A 

number of configurations for the LM HX are available ranging from a simple annular flow passage to 

more complex flow geometries. An important aspect in the implementation of either of these 

configurations is the design of the LM manifold system, which must ensure that the LM flow is uniform 

within the core of the HX while also minimizing the pump power required due to the HX pressure drop. 

Stirling Convertors 

A “family” of Stirling convertors was used in the evaluation of the Lunar Power System (LPS). These 

were derived directly from either existing hardware (in the case of the ASRG and 1 kW), operating 

existing hardware well off their nominal operating point (the low power ASRG), or defined via well 

proven scaling rules for Stirling convertors using the latter hardware as a baseline. For example, scaling to 

a very low power convertor and to the 2- and 5-kW high-power variants. Utilizing this approach it is a 

straightforward matter to scale this technology to different power levels and waste heat rejection 

temperatures, which will allow the convertors to be evaluated within the overall LPS context. The 

following material discusses a number of technical aspects of these convertors, various operating 

constraints, and potential changes required for a LPS application. 

The convertor technology utilized in the LPS evaluation is derived from that employed in the 

Advanced Stirling Convertor (ASC) currently under joint development by the industry team of Sunpower, 

Inc., LM, and NASA GRC. This convertor and the associated technology emphasizes the use of high 

heater head temperatures (1123 K (850 C) versus the current widely employed 923 K (650 C)) 

(Ref. 12), high specific power levels (75 to 100 We/kg), and electrical output to thermal input efficiencies 

of approximately 60 percent of Carnot producing on the order of 160 W of electrical power when 

incorporating two standard Pu-238-fueled GPHS modules. 

The existing ASRG hardware was considered for application directly into two systems that employed 

a varying number of GPHS modules with the alternative isotope. The first technique was to approach or 

meet the current power output of the ASRG via the use of a large number of Am-241 GPHS modules 

coupled to a modified ASRG heater head, as shown schematically in Figure 5. The other option was to 

employ a single modified Am-241 GPHS to a single ASRG convertor heat collector. In this case the 

output power would be about 1/5 of an ASRG convertor using Pu-238. While this latter approach 

dramatically reduced program risk, the performance of the convertor system is significantly lower than 

the nominal values of the current ASRG. This is due to the fact that in the small, low-power convertors, 

the parasitic thermal losses through the mechanical hardware itself can be significant. This occurs 

independent of the operating power level of the convertor. 

The alternative is to simply “scale” the current ASRG design to a lower power level where the 

parasitic losses are resolved. However, in some cases it may not be possible to totally eliminate these 

size-related effects for reasons such as manufacturing limits that define minimum allowable wall 

thickness, which are above those desired. These various technical issues were incorporated into the 

scaling techniques and provide a relationship considering desired power level and convertor operating 

temperatures versus convertor efficiency, as shown in Figure 8, where the reference value (1.0) is on the 

order of 90 W. These relationships were used in the LPS evaluation process. 

An additional issue that must be considered for the conductively coupled configurations discussed 

above is heat flux capabilities of the alternative isotopes. The temperature limits for the GPHS modules 

are set by iridium alloy capsule cladding placed around the Pu-238 fuel. Using the largest face (9.32 by 

9.72 cm), the maximum heat flux out of a single GPHS module at BOL with insulation on the other sides 
NASA/TM—2010-216352 8 

Figure 8.—Impact of convertor power level on conversion efficiency. 

is 2.69 W/cm2 using Pu-238 fuel (Ref. 13). This drops to 0.61 W/cm2

 for Am-241 and 0.46 W/cm2 for 

Sr-90. In contrast to the low heat flux from the GPHS modules is that of the Stirling convertor heater 

head, which requires an input heat flux of about 12 to 15 W/cm2

. It was shown with >6 GPHS modules 

conductively coupling GPHS to Stirling convertors became mass prohibitive (Ref. 14). Because of the 

lower heat flux of both the Am-241 and Sr-90 systems, this has significant impact on the potential power 

levels for the conductively coupled configurations. 

With the introduction of the pumped loop heat transport mechanism using a LM working fluid, the 

issues of the thermal “coupling” between the convertor’s hot-end heat exchanger and the isotope heat 

sources change dramatically for the higher power convertors evaluated within LPS. With this HTS 

configuration the LM side HX characteristics can be tailored to those of the convertor’s helium heat 

exchange process resulting in a highly effective overall heat transfer process. 

For higher power levels, two evaluation paths were considered. One option for the evaluation of the 

1-KWe class of power convertors employing a pumped liquid metal loop was utilizing a well-defined 

1-KWe Sun Power, Inc., (shown in Fig. 9) convertor as a reference point. This specific convertor is of a 

very robust design developed initially for residential micro co-generation applications and operates at hotend temperatures in the range of 775 to 825 K. All of the fundamental convertor technology employed 

such as the linear alternator design, piston/displacer assembly, gas bearing system, etc., are exactly the 

same as those that would be employed in a high-power space-based design and are currently used in the 

ASRC convertor. This convertor has been modified to incorporate a pumped LM hot-end heat exchanger, 

as shown in the following photograph, and currently is installed in a LM pumped loop test facility located 

at Marshall Space Flight Center (MSFC). Testing is expected to commence in the June 2009 time period. 

The primary goal of this testing to fully define the operating and performance characteristics of the LM 

heated version to those of the well-defined 1-kW engine. The latter has been tested extensively at NASA 
NASA/TM—2010-216352 9 

GRC under well-controlled conditions. As can be seen in the photograph, the only change to the basic 

convertor is the 

Figure 9.—1-kWe convertor modified for use with LM pumps. 

addition of the sheet metal (SS-316) inlet/outlet manifold that directs the LM over the existing convertor 

heater head. This same basic configuration would be employed in the alternative isotope-based system. 

Because of the design goals for the basic convertor, the unit is heavier and has lower performance than 

desired for the LPS. 

While the above configuration represents one approach, the primary approach employed in the LPS 

evaluation was to scale up in power utilizing the higher hot-end material capabilities demonstrated as in 

the ASRG convertor. This scale up process is actually easier to carry out since the majority of the issues 

discussed above that occur in scaling to lower powers are eliminated. For the specific convertors 

considered, it was assumed that the Mar-M-247 material technology would be applied. Because of 

potential interactions between this material and the LM working fluid, it was assumed that a compatible 

material, such as SS-316, was clad to the LM exposed surfaces. This cladding does not provide any 

structural support. Because of the high operating temperature capability of this material, it is possible to 

retain conventional cycle temperature ratios but also provide higher rejection temperatures that greatly 

assist in the radiator sizing. Based on these scaling rules, a set of performance relationships were 

developed, which coupled the output power level and hot-end operating temperatures to the convertor 

efficiency. The specific power of the Stirling convertor itself improves somewhat with power level; 

however, this change is relatively small in comparison to the significant mass changes that occur in the 

convertor-related subsystems involving the multiple GPHS to convertor integration and waste heat 

radiator (Ref. 15). Therefore, the focus of the convertor evaluation was in the areas of integration with the 

GPHS heat source, vibration isolation options, and convertor impact on LPS configurations. 

One issue that must be considered is that of all the “reference” convertor configurations discussed 

above, the linear alternators employ high-performance FeNdB magnet technology. Due to this magnet’s 

operating temperature limits, the convertor’s rejection temperature is constrained to less than 

approximately 390 to 400 K (115 to 125 C). In many of the cases considered the worst-case lunar sink 

temperature is 340 K (67 C) allowing only a 50 K temperature difference between the cold end and sink 

temperature to reject heat. In order to allow adequate margin and to maintain a reasonable size on radiator 

area, samarium cobalt (SmCo) magnets were used for all convertors modeled. SmCo magnet/alternator 

technology limits rejection temperatures to somewhat less than 550 K (250 C) and in all cases the 

maximum Stirling cold-end temperature was set at 530 K. Current assessment is that there is little impact 

on alternator mass and none on efficiency if SmCo magnets are used at temperatures between 400 K 

(127 C) and 550 K (277 C) when compared with the FeNdB magnets operating at their lower 

temperature limits. 
NASA/TM—2010-216352 10 

Heat Rejection 

The lunar surface provides a challenging environment for power systems operation due to the long 

duration of the lunar night and the large diurnal changes in environment temperature. At latitudes less 

than 85°, the lunar day/night cycle is relatively constant in duration and averages about 14.74 days 

(354 hr); at latitudes greater than 85°, the local terrain plays an important role in illumination. 

Figure 10 shows the length of the lunar night as a function of latitude, from 0° to 85°. During the day 

the lunar soil is a good solar absorber and emits its radiation back to the environment at temperatures near 

the operating temperature of the heat rejection radiators. Peak soil temperatures at the equator can reach 

over 370 K at local noon. 

Sun position in the sky must also be known to calculate the direct insolation incident on the spacecraft 

components. To estimate the effects of these local environmental conditions on the RPS heat rejection 

system is assigned an orientation on the lunar surface. For the ASRG and other conductively coupled 

systems a cylinder is used to approximate the housing/radiator. These conductively coupled systems are 

oriented such that the piston motion is in line with the gravity vector. For the higher power systems, when 

the orientation of the radiator and the Stirling system is easily decoupled both vertical and horizontal 

radiators are considered. In the past, studies have analyzed the advantages and disadvantages of radiator 

orientation on the lunar surface (Refs. 16 and 17). From these studies, it was found that for vertically 

oriented radiators the maximum and minimum sink temperatures during the lunar day/night cycle was 

about 100 to 314 K and between 100 and 270 K for horizontal radiators that do not see the lunar surface. 

The trade between a two-sided radiator operating at a higher sink temperature and a single-sided radiator 

rejecting heat at a lower temperature may provide different answers depending on isotope availability or 

if the system is mass or area constrained. In addition, the large change in sink temperatures will result in 

changes in both the power output of the system and the temperature range requirements of the heat 

rejection system. For this study a vertical radiator is used, that is, a rectangular plane with one edge 

touching the lunar surface and the other edge skyward. Estimates of the surface and sky view factors on 

the radiator combined with their optical properties are used to generate an energy balance, which provide 

estimates of the radiator temperatures as a function of Sun location. 

Additionally it is necessary to estimate the impacts of the lunar dust on the power system. Typically, 

solar emissivity/absorptivity for an orbiting platform treated with a high emissivity coating after a 10-yr 

life span in low Earth orbit (LEO) will be close to 0.9 and a thermal absorptivity of about 0.15. Based on 

earlier work between Jet Propulsion Laboratory (JPL) and NASA GRC the thermal absorptivity of the 

radiator due to lunar soil contamination was estimated be at worst up to 0.5 with little change in solar 

emissivity (Ref. 18). Using these values for absorptivity and emissivity, effective sink temperatures can 

be found as a function of time of day. Figure 11 shows a plot of sink temperature as a function of time of 

day for a vertically oriented cylindrical radiator located at the equator. Peak sink temperature is about 

340 K and occurs twice just before and after local noon and nighttime sink temperature of 100 K. If the 

radiator dust contamination is not an issue and the thermal absorptivity is what we find in space then the 

peak temperature drops to 320 K. 

The most direct method of removing the waste heat from the Stirling convertor is to have a 

conductive coupling between the cooler section of the Stirling and a heat rejection surface. This method is 

used in the ASRG and consists of a cylindrical ring with the inner portion contacting the Stirling cooler 

and the outer surface of the ring in contact with the ASRG housing/radiator. The advantage of this 

configuration is obviously its simplicity. As the amount of heat rejected increases the trade between 

material thicknesses (and thus mass) an allowable temperature drop eventually favors other heat transport 

augmentation methods. Two other methods of heat removal are considered: heat pipes to move the waste 

heat from the Stirling cold end to the radiator surface and to flow a liquid over the cold end of the Stirling 

and pumping it to a radiator. 

In this analysis, heat pipes are used either as the primary heat transport device from the Stirling cold 

end to the surface of a radiator or as a heat spreader connected the liquid pumped loop on the evaporator 

side and to the radiator fins on the other. 
NASA/TM—2010-216352 11 

Figure 10.—Mean lunar night duration as a function of latitude. 

Figure 11.—Sink temperature as a function of time of lunar day. 
NASA/TM—2010-216352 12 

Figure 12.—Radiator panel layout. 

For a pumped loop heat transport system, water is passed over the Stirling cold end, transported out to 

the radiator panels, goes through the pumps, and returns to the Stirling convertor. Stirling convertors 

operate best when the inlet to exit coolant temperature difference is kept to a minimum. In general the 

temperature rise of a fluid used to remove the waste heat should be about 25 K or less to ensure no 

Stirling cycle performance penalty. The radiator panel design, see Figure 12, consists of water heat pipes 

sandwiched between two outer face sheets. Panel mass was approximately 3.5 kg/m2

 for the cases shown 

for the two-sided radiator. This areal mass does not include the fluid ducts, fluid, or pumps, which are 

accounted for separately. The pump design selected is scaled from other space pumps and is scaled both 

in efficiency and mass to meet the pressure drop and flow rate requirements of the system. 

Analysis 

A modeling tool has been developed to size Stirling RPS systems that operate under a variety of 

environmental conditions and a wide range of power outputs. At each required power level cold-end 

temperature is varied to explore the trade between efficiency and mass. In order to minimize the amount 

of isotope required the thermodynamic cycle requires as low a waste heat rejection temperature as 

possible. The penalty for this low rejection temperature is that the area required to reject the waste heat 

increases with decreasing rejection temperature. This trade between the isotope source and radiator size 

leads to mass minimums for the overall system. After the design point is set, these component sizes are 

used to find off-design performance in both different environments and as the isotope decays in time. 

ASRG on the Lunar Surface 

The ASRG using Pu-238 is currently estimated to produce ~160 W direct current (dc) when using the 

high-strength alloy MarM-247 as its heater head material and placed in a deep space environment 

(Ref. 19). Hot and cold side temperatures for the ASRG are 850 (1123)/90(363) C (K), respectively. 

Each ASRG comprises two ASCs with each producing 88 W (2·88 = 176 W total) of alternating current 

(ac) power, which is sent through power conditioning and control electrics that monitor the ASRG and 

converts the 102-Hz power to 28-V dc (nominal). Estimates for the Power Conditioning and Control 

(PC&C) at 164 W of throughput are estimated to be approximately 12 W. Total dc power output for the 

ASRG under the above referenced conditions and with the 12 W for the PC&C is 164 W. Calculations are 

made as to how the ASRG would perform during both day/night cycles and over its required 10-yr life. 

Control of the ASRG is very important as to how the system will behave at off-nominal conditions. For 

these cases it was assumed the controller maintained a constant heater head temperature during the entire 

life of the mission. Other control schemes will be discussed later. Table 4 shows a summary of both the 

ASRG design point performance as quoted from NASA and estimates of how it will perform under the 

lunar environmental conditions. Power output during daylight hours falls from the nominal 164 (estimate) 

to 143 W due to the higher temperatures of the surrounding environment. Figure 13 shows both power 

output and Stirling efficiency over a lunar day/night cycle. 
NASA/TM—2010-216352 13 

TABLE 4.—PERFORMANCE SUMMARY USING VARIOUS ISOTOPES IN GPHS MODULE HOUSING 

Description ASRG 

Pu-238 

ASRG 

Am-241 

ASRG 

Sr-90 

Mission duration, yr 10 10 10 

Stirling hot end temp, °C 850 850 850 

Stirling cold end temp, °C 90 47 44 

BOM heat, W 500 112 85.6 

AC output, W 176 38 29 

PC&C power, W 12 9 9 

BOM DC power-nominal, W 164 29 19 

EOM DC power-nominal, W 148 28 13 

BOM daytime power (est.), W 143 25 17 

BOM night power (est.), W 172 34 25 

No. GPHS modules 2 2 2 

Mass, kg 20.2 20.2 19.4 

Specific power, W/kg 8.1 1.4 1.0 

BOM efficiency, percent ~32% ~26% ~23% 

Diameter, cm 46 46 46 

Length, cm 72 72 72 

Figure 13.—Power output and efficiency as a function of time of day. 

Power output increases during the night from 164 to 172 W as the environment temperature drops. As the 

isotope decays, thermal power output from the GPHS modules decrease and thereby reduces the power 

output of the ASRG. Nominal power for a 10-yr mission drops from 164 to 148 W while Stirling cold-end 

temperature drops from 360 to 350 K. Figure 14 shows power output and Stirling cold-end temperature as 

a function of mission time for the Pu-238 ASRG.
NASA/TM—2010-216352 14 

Figure 14.—Nominal power output as a function of mission. 

ASRG With Alternative Isotopes 

Direct replacement of the two Pu-238 GPHS modules with either Am-241 or Sr-90 GPHS filled 

modules was outlined earlier. Using Sr-90, system mass was reduced by 0.8 kg because of the decrease in 

GPHS module mass due to the lower density of the SrO fuel. Radiator, hot and cold side flanges, Stirling 

convertor, and all of the other components are assumed to be the same as in the ASRG. BOL power 

output is reduced significantly with the Am-241 system producing 29 W nominally and the Sr-90-based 

ASRG producing 20 W. Table 4 shows a summary and comparison of the Pu-238-, Am-241-, and Sr-90-

based ASRGs. Controller power consumption is reduced but is a greater fraction of total power because 

although ac/dc conversion power throughput is reduced, the monitoring and system housekeeping power 

consumption remains constant. Because of the very long half-life of the Am-241 both power output and 

cold-end temperature remain nearly constant. Figures 15 and 16 show power output of the Am-241 

ASRG as both a function of time and diurnal lunar cycle. Power output varies for the lunar day/night 

cycles in a similar fashion to that of the Pu-238-based system. Maximum power occurs during lunar night 

and is 34 W while minimum power occurs in midmorning and mid-afternoon with the ASRG providing 

25 W of dc power. Sr-90-based ASRG, because of its short 28-yr half-life has a more rapid drop in power 

output than either of the two other systems. Power output under nominal conditions falls from 

19 W BOL to 13 W at 10 yr. BOL power output varies from 25 W at night to 17 W during the day. 

Figures 17 and 18 show the power output changes. 
NASA/TM—2010-216352 15 

Figure 15.—Power output and Stirling cold-end temperature for an Am-241 GPHS ASRG. 

Figure 16.—Am-241 ASRG power and convertor efficiency as a function of lunar day/night cycle. 
NASA/TM—2010-216352 16 

Figure 17.—Power output and Stirling cold-end temperature for a Sr-90 GPHS. 

Figure 18.—Sr-90 ASRG power and convertor efficiency as a function of lunar day/night cycle. 
NASA/TM—2010-216352 17 

TABLE 5.—COMPARISON OF ALTERNATIVE ISOTOPE SYSTEMS WITH TWO GPHS MODULES 

Description ASRG 

Pu-238 

ASRG 

Am-241 

ASRG 

Sr-90 

Mission duration, yr 10 10 10 

Stirling hot end temp, °C 850 850 850 

Stirling cold end temp, °C 90 52 52 

BOM heat, W 500 114 86 

AC output, W 176 38 27 

PC&C power, W 12 9 9 

BOM DC power-nominal, W 164 29 20 

EOM DC power-nominal, W 148 29 14 

BOM daytime power (est.), W 143 25 15 

BOM night power (est.), W 172 34 22 

No. GPHS modules 2 2 2 

Mass, kg 20.2 11 9 

Specific power, W/kg 8.1 2.6 2.1 

BOM efficiency, percent 33% 25% 21% 

Alternative Isotope With Optimized ASRG System 

The previous section described the direct replacement of the Pu-238 GPHS modules with both 

Am-241 and Sr-90. Both the Am-241 and Sr-90 systems produced one-fifth the power but system mass 

was nearly identical because it was assumed that all of the ASRG components were both the same mass 

and size. If a new optimized system including an optimized convertor, smaller radiator, reduction in heat 

input/rejection flange masses, significant improvements in system mass and performance may be 

possible. Table 5 shows estimates for two GPHSs optimized for best specific power (W/kg). The two 

GPHS Am-241 systems increased its specific power over the direct ASRG replacement from 1.4 to 

2.6 W/kg and the Sr-90 specific mass improved from 1 to 2.1 W/kg. 

Controllers 

Stirling convertor controllers are used to ensure that the convertor operates within its given design 

specifications over its life. There are many control strategies that could be implemented for a Stirling 

isotope power system but emphasis in this paper is placed on control strategies that fix one parameter. 

Two control strategies are explored here to help understand the differences on system operating 

conditions with each control strategy. One control strategy is to fix the heater head temperature for the 

mission duration. In order to maintain this temperature with varying environmental conditions and the 

reduction in heat produced as the isotope decays the engine stroke is varied. By changing the stroke the 

convertor will pull more or less heat through the engine to maintain a constant heater head temperature. 

A second control scheme is to fix the heat input into the convertor. This is similar to what occurs 

when the stroke is held at a constant amplitude. The advantage of this control strategy is that only piston 

position is needed and the voltage on the alternator of the Stirling convertor is varied to achieve the 

desired stroke. Figure 19 shows a plot of heater head temperature and efficiency for a Pu-238-based 

ASRG with both the constant temperature heater head and constant heat input for the ASRG. Constant 

heat input control produces a reduction in heater head temperature as mission time increases. At 10 yr, the 

heater head temperature has dropped from 1123 K at BOM to about 800 K with a corresponding drop in 

efficiency. Figure 20 shows power output and Stirling cold-end temperature as a function of mission time. 

For the constant heater head temperature control, power output drops from 160 to 148 W while for the 

constant heat input control, power output drops from 160 to 138 W. Cold-end temperature drops slightly 

for the fixed heater head temperature control (360 to 350 K) while for the fixed heat input cold-end 

temperature rises from 360 to 367 K. Because the isotope is decaying, the total heat generated by the 

GPHS modules is falling as time passes. In this model heat into the Stirling convertor is the difference 

between heat generated by the GPHS modules at any given time and that heat lost through the insulation. 
NASA/TM—2010-216352 18 

Insulation thickness is fixed at the design point BOM from the temperature of the outer surface of the 

GPHS and the sink temperature and the difference between heat required by the engine and that generated 

by the GPHS modules. To maintain a constant heat input into the heater head the losses from the decaying 

isotope must be made up for by losing less through the insulation. Therefore the temperature of the hot 

end must decrease until the insulation loss is reduced by the amount needed to make up for the decaying 

isotope. As the hot-end temperature drops, but with a fixed thermal throughput into the engine and a fixed 

radiator area, the Carnot efficiency will drop. This leads to a rise in cold-end temperature. Other methods 

or hybrid methods were not considered for this study. 

Figure 19.—Heater head temperature and efficiency for a Pu-238 system with different control strategies. 

Figure 20.—Power output and Stirling cold end for a Pu-238 system using different control strategies. 
NASA/TM—2010-216352 19 

Figure 21.—Am-241 power output and cold-end temperature as a function of time and convertor control method. 

TABLE 6.—COMPARISON OF 160 We Am-241 AND Pu-238 SYSTEMS 

Description ASRG 

Pu-238 

Optimized ASRG 

Am-241 

Mission duration, yr 10 10 

Stirling hot end temp, °C 850 850 

Stirling cold end temp, °C 90 87 

BOM heat, W 500 569 

AC output, W 176 171 

PC&C power, W 12 11 

BOM DC power-nominal, W 164 160 

EOM DC power-nominal, W 148 157 

BOM daytime power (est.), W 143 146 

BOM night power (est.), W 172 174 

No. GPHS modules 2 10 

Mass, kg 20.2 41 

Specific power, W/kg 8.1 3.9 

BOM efficiency, percent 33% 28% 

Figure 21 shows power output as a function of time for the two controller architectures using both 

Am-241and cold-end temperatures. Because of the very long half-life of Am-241, both control schemes 

produce similar results with respect to performance and very stable power output over the entire mission 

life. Power for the constant temperature heater head falls from 29 to 28 W over the 10-yr mission while 

for the constant heat input control drops to 27 W at EOL. 

ASRG Using Pu-238 and Am-241 

Next we explore what a RPS might look like if it was required to produce 160 W BOL identical to 

what the ASRG does except with Am-241. In order for the Americium system to produce 160 W, 

10 Am-241 GPHS modules (5 per convertor side times 2 convertors = 10) must be arranged around each 

heater head to minimize the impact of conduction losses from the GPHS surface to the heater head. 

Table 6 shows mass for the system has approximately doubled due to the much greater number of GPHS 

modules, the added insulation, and its associated equipment. Cold-end mass is similar while overall 

efficiency has dropped by 5 percent due to the increased losses associated with the larger heat source area. 
NASA/TM—2010-216352 20 

Maximum Power Output of Conductively Coupled ASRG 

Integration of multiple GPHSs with a single Stirling becomes mass prohibitive at some point due to 

the increasing distance between the GPHS and the Stirling and the mass associated with integrating the 

two systems. Earlier analysis suggested over six GPHSs per convertor, the system-specific mass began to 

decrease above 1 kWe for a Pu-238 system. Because the total heat rejected is similar to what is found in 

the Pu-238 ASRG for the 5 GPHS module Am-241 system, the cold-end size and operational 

temperatures are very similar. For the Pu-238 system with five GPHSs the heat rejected has increased by 

a factor of 5 over both the 160-W Am-241 system and the ASRG. The conductively coupled design cold 

end takes the heat from the cold flange, and conducts via a conductive collar out to the ASRG casing and 

rejects the heat to space. The dramatic increase in heat rejected makes the conductively coupled cold-end 

design impractical. One method to improve heat transfer from the cold end to the radiator surface is to 

connect heat pipes to the cold end of the convertor and run these out to the surface of the radiator. For this 

five-GPHS Pu-238 design, water heat pipes are used to improve performance and spread the heat from the 

cold end of the convertor to the surface of the radiator. Details of this design can be found in 

Reference 15. Table 7 shows a comparison of both the Am-241 and Pu-238 systems conductively coupled 

systems with five GPHSs. 

TABLE 7.—COMPARISON OF FIVE GPHS Pu-238 AND Am-241 SYSTEMS 

Description 5 GPHS Pu-238 System 5 GPHS Am-241 System 

Mission duration, yr 10 10 

Stirling hot end temp, °C 850 850 

Stirling cold end temp, °C 87 87 

BOM heat, W 2500 569 

AC output, W 820 171 

PC&C power, W 20.44 11 

BOM DC power-nominal, W 800 160 

EOM DC power-nominal, W 741.7 157 

BOM daytime power (est.), W 745.2 146 

BOM night power (est.), W 885.9 174 

No. GPHS modules 10 10 

Mass, kg 111.2 41 

Specific power, W/kg 7.2 3.9 

BOM efficiency, percent 32% 28% 

Radiator area, m2 7.2 0.9 

Pumped Loop Systems 

As was discussed earlier, the GPHS modules are attached to a metal conduction plate, which makes 

up one side of the heat exchanger with liquid NaK flowing on the other. Multilayer insulation (MLI) and 

a aluminum housing covers the entire heat source assembly. Heat loss through the insulation, structure, 

and piping is set to 5 percent of the total heat generated with a fixed 2.5 percent heat loss assumed 

through the piping and structure and a 2.5 percent heat loss through the surface of the HX. The area’s 

temperatures and allowable heat loss through the pipes and HX, sizes the thickness of the MLI. Heat is 

taken from the heat source assembly via a flowing NaK loop and moved to the two Stirling convertors. It 

is necessary to ensure that the primary loop NaK does not flash to vapor and this is accomplished by 

maintaining sufficient pressure in the system with an accumulator such that the operating pressure is 

below the vapor pressure of the NaK. Figure 22 shows the vapor pressure of NaK as a function of 

temperature along with the maximum operating temperature in the primary loop for both Stirling systems 

modeled. Maximum vapor pressure occurs for the high-temperature Stirling system at 1133 K and is 

1.84 bar. The NaK is pumped around the hot side of the system with two series annular linear induction 

electromagnetic pumps (ALIPs). The series configuration is possible because the ALIPs are open channel 

pumps with external coils that provide a magnetic field. The ALIPS have efficiency of 15 percent. The 

heat rejection system for this case is assumed to be a north-south vertically oriented radiator. Water is 
NASA/TM—2010-216352 21 

Figure 22.—NaK vapor pressure versus temperature. 

used as the working fluid and two parallel impellar pumps provide its motivation. The water loop is in 

contact with water heat pipes bonded to the outside of the water pumped loop, which are embedded 

within a graphite heat rejection radiator that is discussed above. A range of Stirling cold-end temperatures 

are modeled and trades are made between minimum mass and GPHS modules. 

Power levels of 1, 2, and 5 kWe and maximum heater head temperatures of 923 and 1123 K are 

modeled for this study. Figures 23 and 24 show mass and the number of GPHS modules for a 1-kWe, 

923 K system using both a Pu-238 and Am-241 as a function of Stirling cold-end temperature. Minimum 

mass for the Pu-238 system occurs at 400 K while for the Am-241 system occurs at 360 K. Figures 25 

and 26 show the relative mass fractions for the major system components operating at a cold-end 

temperature of 360 K. Because of the much lower thermal output of the Am-241 when compared with the 

Pu-238 system, the GPHS module and heat source mass in a more significant fraction of the overall 

system mass and the minimum mass system Stirling cold-end temperature, drops to reduce the GPHS 

module count. Mass fraction of the GPHS modules for the Pu-238 system is 16 percent while for the 

Am-241 system is 36 percent. Additionally, insulation mass has increased from 3 to 5 percent due to the 

significantly higher number of GPHS modules and greater surface area. 

Figures 27 and 28 show the same mass and GPHS count plots with the exception of an 1123 K heater 

head temperature for both. Mass for the Pu-238 systems has been decreased by 10 percent with the 

change in heater head temperature while GPHS count has dropped by 16 to 12. For the Am-241 system 

mass has dropped by about 13 percent while Am-241 GPHS module count has dropped from 67 to 53. 
NASA/TM—2010-216352 22 

Figure 23.—1-kWe Pu-238 pumped loop system, 923 K heater head temperature. 

Figure 24.—1-kWe AmO2 pumped loop system, 923 K heater head temperature. 
NASA/TM—2010-216352 23 

Figure 25.—Mass fraction for a 1-kWe 1123 K heater head Pu-238 system. 

Figure 26.—Mass fraction for a 1-kWe 1123 K heater head Am-241 system. 
NASA/TM—2010-216352 24 

Figure 27.—1-kWe Pu-238 pumped loop system, 1123 K heater head temperature. 

Figure 28.—1-kWe Am-241 pumped loop system, 1123 K heater head temperature. 
NASA/TM—2010-216352 25 

Tables 8 and 9 show the 1, 2, and 5 kWe for both the 923 K heater head systems and the 1123 K 

systems with both Pu-238 and Am-241 GPHS modules. Pu-238 systems are about half the system mass of 

the Am-241 systems with this change growing with increasing power level. Heater head temperature 

reduces system mass by about 10 percent at 1 kWe and growing to about 20 percent at 5 kWe. Figure 29 

shows a comparison of the 923 and 1123 K system from 1 to 5 kWe for both Pu-238 and Am-241. 

TABLE 8.—1-, 2-, AND 5-kWe Pu-238 AND Am-241 SYSTEMS WITH 

A 923 K MAXIMUM HEATER HEAD TEMPERATURE 

Description 1 kWe Pu238 System 2 kWe Pu-238 

system 

5 kWe Pu-238 

system 

1 kWe Am241 system 2 kWe Am241 system 5 kWe Am241 system 

Isotope PuO2 PuO2 PuO2 AmO2 AmO2 AmO2

Heat source integration method Pumped 

loop 

Pumped loop Pumped loop Pumped 

loop 

Pumped 

loop 

Pumped 

loop 

Mission duration, yr 10 10 10 10 10 10 

Stirling hot end temp, °C 913 913 913 913 913 913 

Stirling cold end temp, °C 360 360 360 360 360 360 

Temperature ratio 2.54 2.54 2.54 2.54 2.54 2.54 

BOM heat, W 4000 7750 19250 3812 7625 19175 

AC output, W 1075 2151 5376 1075 2151 5376 

PC&C power, W 23 38 81 23 38 81 

BOM DC power-nominal, W 1000 2000 5000 1000 2000 5000 

EOM DC power-nominal, W 919 1844 4612 986 1971 4927 

BOM daytime power (est.), W 904 1806 4509 905 1809 4518 

BOM nightime power (est.), W 1045 2090 5226 1047 2094 5236 

No. GPHS modules 16 31 77 67 134 337 

System mass, kg 143 268 648 303 581 1426 

Specific power, W/kg 7.5 8.0 8.3 3.5 3.7 3.8 

BOM efficiency, percent 25% 26% 26% 26% 26% 26% 

TABLE 9.—1-, 2-, AND 5-kWe Pu-238 AND Am-241 SYSTEMS WITH 

A 1123 K MAXIMUM HEATER HEAD TEMPERATURE 

Description 1 kWe Pu-238 

System 

2 kWe Pu-238 

system 

5 kWe Pu238 system 1 kWe Am241 system 2 kWe Am241 system 5 kWe Am241 system 

Isotope PuO2 PuO2 PuO2 AmO2 AmO2 AmO2

Heat source integration method Pumped loop Pumped loop Pumped loop Pumped 

loop 

Pumped 

loop 

Pumped 

loop 

Mission duration, yr 10 10 10 10 10 10 

Stirling hot end temp, °C 1113 1113 1113 1113 1113 1113 

Stirling cold end temp, °C 360 360 360 360 360 360 

Temperature ratio 3.09 3.09 3.09 3.09 3.09 3.09 

BOM heat, W 3000 6000 15000 3016 5975 14965 

AC output, W 1075 2151 5376 1075 2151 5376 

PC&C power, W 23 38 81 23 38 81 

BOM DC power-nominal, W 1000 2000 5000 1000 2000 5000 

EOM DC power-nominal, W 924 1848 4616 985 1971 4926 

BOM daytime power (est.), W 931 1862 4650 931 1862 4652 

BOM nightime power (est.), W 1035 2070 5170 1034 2069 5173 

No. GPHS modules 12 24 60 53 105 263 

System mass, kg 130 231 540 264 492 1173 

Specific power, W/kg 8.3 9.3 10.0 4.1 4.4 4.6 

BOM efficiency, percent 33% 33% 33% 33% 33% 33% 
NASA/TM—2010-216352 26 

Figure 29.—Mass as a function of power level for 923 and 1123 K Pu-238 and Am-241 systems. 

Conclusions 

Substitution of Am-241 for Pu-238 into GPHS modules produces systems, which are either heavier 

for the same power or produces less power with the same GPHS count. Because of the very long half-life 

of Am-241 power output is nearly constant over the mission times considered here. Direct replacement of 

the Pu-238-based ASRG with Am-241 dropped power output from 164 to 29 W BOM and 148/28 W 

EOL dc power output, respectively. Sr-90 was also considered as another isotope option and provided 

20 W BOM but because of its much shorter half-life provided only 14 W EOM when placed into an 

ASRG. 

Two control strategies were considered for the ASRG with the various isotopes. For the Am-241 and 

Pu-238, both appear reasonable for control of the ASRG. Power production is better using the constant 

heater head temperature control for both options; however, with the very long half-life of the Am-241 

very small differences between the strategies is seen. 

As power levels grow, pumped liquid NaK on the hot side and pumped water on the cold side of the 

Stirling were used for heat transport. A recently developed low temperature (923 K), 1-kWe Stirling 

convertor was modeled along with a higher temperature future convertor (1123 K). Heater head 

temperature improvements led to an approximately 10 percent reduction in system mass at 1 kWe and 

20 percent and 5 kWe. Pu-238 systems are approximately one-half the mass of their Am-241 counterparts 

with the mass differences growing slowly as power level increases. 

Am-241 appears to be an isotopic candidate worth further consideration. When specific power is most 

important to a mission designer, Pu-238 will be the clear choice. When isotopic availability is of prime 

concern and mass a smaller driver, Am-241 systems offer an attractive alternative. 
NASA/TM—2010-216352 27 

References 

1. Pantano, D.R., Hill, D.H. “Thermal Analysis of Step 2 GPHS for Next Generation Radioisotope 

Power Source Missions,” STAIF 2005. 

2. Or, C.T., Carpenter R.T., “Steady State Temperature Predictions for General Purpose Heat Source in 

Various Environments,” Proceedings of the 33rd Intersociety Engineering Conference on Energy 

Conversion, IECEC–98–227, Colorado Springs, CO (1998). 

3. Or, C.T., Carpenter R.T., “Steady State Temperature Predictions for General Purpose Heat Source in 

Various Environments,” Proceedings of the 33rd Intersociety Engineering Conference on Energy 

Conversion, IECEC–98–227, Colorado Springs, CO (1998). 

4. http://www.ne.doe.gov/space/neSpace2g.html 

5. Or, C.T., Carpenter R.T., “Steady State Temperature Predictions for General Purpose Heat Source in 

Various Environments,” Proceedings of the 33rd Intersociety Engineering Conference on Energy 

Conversion, IECEC–98–227, Colorado Springs, CO (1998). 

6. http://www.ne.doe.gov/space/neSpace2g.html 

7. Department of Energy, Draft Environmental Impact statement for the proposed Consolidation of 

Nuclear Operations Related to Production of Radioisotopic Power Systems, DOE/EIS-0373D, p. 19. 

8. R.C. O’Brien, et al., “Safe Radioisotope Thermoelectric Generators and Heat Sources for Space 

Applications,” Journal of Nuclear Materials, Issue 377 (2008) pp. 506–521. 

9. R.C. O’Brien, et al., “Safe Radioisotope Thermoelectric Generators and Heat Sources for Space 

Applications,” Journal of Nuclear Materials, Issue 377 (2008) pp. 506–521. 

10. Wood, G., Carroll, C., Matejczyk, D., Soendker, E., Penswick, L.B., “Advanced Stirling Convertor 

Update,” STAIF 2006. 

11. Schmitz, P., Penswick, L.B., Shaltens, R., “Stirling Isotope Power Systems for Stationary and Mobile 

Lunar Applications,” 4th IECEC, San Diego, CA, 2006. 

12. Wood, G., Carroll, C., Matejczyk, D., Soendker, E., Penswick, L.B., “Advanced Stirling Convertor 

Update,” STAIF 2006. 

13. ibid. 

14. Schmitz, P., Penswick, L.B., Shaltens, R., “Stirling Isotope Power Systems for Stationary and Mobile 

Lunar Applications,” 4th IECEC, San Diego, CA, 2006. 

15. Schmitz, P., Penswick, L.B., Shaltens, R., “Modular GPHS-Stirling Power System for Lunar 

Habitation Module,” 3rd IECEC, San Francisco, CA, 2005. 

16. Bien, D.D. Guentert, D.C., “A Method for Reducing the Equivalent Sink Temperature of a Vertically 

Oriented Radiator on the Lunar Surface,” NASA TM X–1729, NASA Lewis Research Center, 

Cleveland, OH, 1969. 

17. Dallas, D., Diaguila, A.J., Saltsman, J.F., “Design Studies of the Effects of Orientation, Lunation, and 

Location on the Performance of Lunar Radiators,” NASA TM X–1846, NASA Lewis Research 

Center, Cleveland, OH, 1971. 

18. JPL presentation on lunar sink temperatures 

19. http://esto.nasa.gov/conferences/nstc2007/papers/Shaltens_Richard_D2P1_NSTC-07-0138.pdf 
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Lunar Surface Stirling Power Systems Using Isotope Heat Sources 

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13. SUPPLEMENTARY NOTES

14. ABSTRACT

For many years, NASA has used the α decay of plutonium-238 (Pu-238) (in the form of the General Purpose Heat Source (GPHS)) as a heat 

source for Radioisotope Thermoelectric Generators (RTGs), which have provided electrical power for many NASA missions. While RTGs 

have an impressive reliability record for the missions in which they have been used, their relatively low thermal to electric conversion 

efficiency and the scarcity of plutonium-238 (Pu-238) has led NASA to consider other power conversion technologies. NASA is considering 

returning both robotic and human missions to the lunar surface and, because of the long lunar nights (14.75 Earth days), isotope power 

systems are an attractive candidate to generate electrical power. NASA is currently developing the Advanced Stirling Radioisotope 

Generator (ASRG) as a candidate higher efficiency power system that produces greater than 160 W with two GPHS modules at the 

beginning of life (BOL) (~32% efficiency). The ASRG uses the same Pu-238 GPHS modules, which are used in RTG, but by coupling them 

to a Stirling convertor provides a four-fold reduction in the number of GPHS modules. This study considers the use of americium-241 (Am241) as a substitute for the Pu-238 in Stirling- convertor-based Radioisotope Power Systems (RPS) for power levels from tens of watts to 5 

kWe. The Am-241 is used as a substitute for the Pu-238 in GPHS modules. Depending on power level, different Stirling heat input and 

removal systems are modeled. It was found that substituting Am-241 GPHS modules into the ASRG reduces power output by about onefifth while maintaining approximately the same system mass. In order to obtain the nominal 160 W of electrical output of the Pu-238 ASRG 

requires 10 Am-241 GPHS modules. Higher power systems require changing from conductive coupling heat input and removal from the

Stirling convertor to either pumped loops or heat pipes. Liquid metal pumped loops are considered as the primary heat transportation on the 

hot end and water pumped loop/heat pipe radiator is considered for the heat rejection side for power levels above 1 kWe. 

15. SUBJECT TERMS

Radioisotope Thermoelectric Generator (RTG); Advanced Stirling Radioisotope Generator (ASRG); General Purpose Heat Source 

(GPHS) 

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