Anna E. Thomas

Georgia Institute of Technology, Atlanta, Georgia

Nikita T. Saxena

Tufts University, Medford, Massachusetts

Dale T. Shouse and Craig Neuroth

Air Force Research Laboratory, Wright-Patterson Air Force Base, Wright-Patterson AFB, Ohio

Robert C. Hendricks

Glenn Research Center, Cleveland, Ohio

Amy Lynch, Charles W. Frayne, Jeffrey S. Stutrud, Edwin Corporan, and Terry Hankins

Air Force Research Laboratory, Wright-Patterson Air Force Base, Wright-Patterson AFB, Ohio

Alternate-Fueled Combustor-Sector Performance

NASA/TP—2013-217735

September 2013
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Anna E. Thomas

Georgia Institute of Technology, Atlanta, Georgia

Nikita T. Saxena

Tufts University, Medford, Massachusetts

Dale T. Shouse and Craig Neuroth

Air Force Research Laboratory, Wright-Patterson Air Force Base, Wright-Patterson AFB, Ohio

Robert C. Hendricks

Glenn Research Center, Cleveland, Ohio

Amy Lynch, Charles W. Frayne, Jeffrey S. Stutrud, Edwin Corporan, and Terry Hankins

Air Force Research Laboratory, Wright-Patterson Air Force Base, Wright-Patterson AFB, Ohio

Alternate-Fueled Combustor-Sector Performance

NASA/TP—2013-217735

September 2013

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Prepared for the

14th International Symposium on Transport Phenomena and Dynamics of Rotating Machinery

(ISROMAC–14)

sponsored by the Fluid Systems Technology (FST) of Technische Universitat Darmstadt

Honolulu, Hawaii, February 27–March 2, 2012
Acknowledgments

The authors would like to thank the NASA GRC technical editing and publishing staff

for their diligent efforts in completing our paper.

Available from

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Available electronically at http://www.sti.nasa.gov

Trade names and trademarks are used in this report for identifi cation 

only. Their usage does not constitute an offi cial endorsement, 

either expressed or implied, by the National Aeronautics and 

Space Administration.

This work was sponsored by the Fundamental Aeronautics Program 

at the NASA Glenn Research Center.

Level of Review: This material has been technically reviewed by expert reviewer(s). 
NASA/TP—2013-217735 1 

Alternate-Fueled Combustor-Sector Performance 

Anna E. Thomas

Georgia Institute of Technology 

Atlanta, Georgia 30332 

Nikita T. Saxena

Tufts University 

Medford, Massachusetts 02155 

Dale T. Shouse and Craig Neuroth

Air Force Research Laboratory 

Wright-Patterson Air Force Base 

Wright-Patterson AFB, Ohio 45433 

Robert C. Hendricks

National Aeronautics and Space Administration 

Glenn Research Center 

Cleveland, Ohio 44135 

Amy Lynch, Charles W. Frayne, Jeffrey S. Stutrud, Edwin Corporan, and Terry Hankins 

Air Force Research Laboratory 

Wright-Patterson Air Force Base 

Wright-Patterson AFB, Ohio 45433 

Summary 

In order to realize alternative fueling for military and 

commercial use, the industry has set forth guidelines that must 

be met by each fuel. These aviation fueling requirements are 

outlined in MIL–DTL–83133F(2008) or ASTM D 7566 

Annex (2011) standards, and are classified as “drop-in” fuel 

replacements. This report provides combustor performance 

data for synthetic-paraffinic-kerosene- (SPK-) type (FischerTropsch (FT)) fuel and blends with JP–8+100, relative to JP–

8+100 as baseline fueling. Data were taken at various nominal 

inlet conditions: 75 psia (0.52 MPa) at 500 F (533 K), 

125 psia (0.86 MPa) at 625 F (603 K), 175 psia (1.21 MPa) at 

725 F (658 K), and 225 psia (1.55 MPa) at 790 F (694 K). 

Combustor performance analysis assessments were made for 

the change in flame temperatures, combustor efficiency, wall 

temperatures, and exhaust plane temperatures at 3, 4, and 

5 percent combustor pressure drop (P) for fuel:air ratios 

(F/As) ranging from 0.010 to 0.025. Significant general trends 

show lower liner temperatures and higher flame and 

combustor outlet temperatures with increases in FT fueling 

relative to JP–8+100 fueling. The latter affects both turbine 

efficiency and blade and vane lives. 

Introduction 

Finding an alternative fuel for aviation application requires 

a fuel feedstock with sustainable supply at a low cost with low 

or no negative environmental impact. The requirements for 

these “drop-in” fuel replacements are outlined in the MIL–

DTL–83133F(2008) or ASTM D 7566 Annex (2011), 

approved standards for military and civil use, respectively. 

Alternate jet fuels need to be compatible with current engines 

and aircraft fuel handling systems in order to reduce the need 

for new systems to accommodate new fuels that may perform 

differently than the currently used petroleum fuels. 

Even proven alternate fuels face tough issues such as 

secure, sustainable productivity at competitive pricing. 

Recently, the Federal Aviation Administration (FAA) 

announced support of eight companies conducting research 

into commercial jet fuel alternatives that conform to ASTM D 

7566 and are based on resources readily available in the 

United States (Ref. 1). One of the ideas being explored is to 

produce aviation fuels from carbon monoxide given off by 

industrial waste gases that would otherwise add to 

atmospheric pollution. Another idea is to explore the 

conversion from cellulosic and conventional plant sugars to 

fuels. Others involve the development of catalysts to convert 

different carbon sources into fuels in small-scale reactors to 

serve as distributed fuel production sources. Currently, the 

biofuels used by the U.S. Navy cost about $26/gal ($6.87/L) 

(Ref. 2). The money that the FAA is funneling into these 

projects could boost the production of cost-effective fuels 

from biomass and waste feedstocks to enable the affordability 

for commercial and military aviation fueling. Yet to date, the 

alternate fuels industry competitive costs and productivity 

have not responded to feedstock restraints, incentives, 

subsides, or mandates, and compliance taxes are passed to 

consumers (Ref. 3). 

 
NASA/TP—2013-217735 2 

Adopting alternate fuels and fuel blends requires the use of 

fuel-flexible systems (combustors and engines) without 

sacrificing performance requirements. For military aviation, 

an alternative for traditional fuel for gas turbine and diesel 

systems is required. However, in many proposed alternates, 

the lack of sufficient amounts of aromatics that swell some 

fuel system seals and sulfur, which provides fuel injection 

pump lubricity, have the potential to reduce design component 

useful life (Ref. 4). For these fuels, additives are needed to 

increase useful component life while maintaining the 

performance. It is thought that Fischer-Tropsch- (FT)-type 

fuels can support gas turbine engines at similar levels as well 

as have potential use for diesel systems. 

This report provides preliminary combustor performance 

data for SPK-type FT fuel and blends relative to pure JP–

8+100 (herein referred to as “JP–8”), the currently used 

aviation fuel. Data for “Combustor A,” a three-cup sector 

representative of current engine combustor technology (see 

computational fluid dynamics (CFD) images, Ref. 6, 

proprietary details withheld), were taken at 0, 50, and 

100 percent FT fueling (denoted as JP–8, 50:50, and FT, 

respectively) with varied parameters of fuel:air ratio (F/A), 

percent combustor pressure drop (%ΔP), and absolute 

pressure. The data collected show that higher combustor 

operating temperatures have the potential to enhance system 

efficiency, but also take a toll on component life, as they have 

a greater impact on the oxidation and failure of the materials 

within the combustor and turbine. A small temperature 

difference of combustor gas entering the turbine can both be 

critical to turbine life and affect efficiency; there is a need for 

a good balance. “Bleed air,” used to cool the combustor, case, 

turbine blades, vanes, and nozzles, could be increased to 

compensate for the enhanced turbine inlet temperature; this 

parasitic air decreases the system efficiency but helps to 

maintain a reasonable turbine life. 

Fuel specifications; the test facility conditions, operations, 

schematic, and fueling system; estimates of measurement 

errors; and combustor thermal data and postprocessing 

parameters of Combustor A are given in the Shouse et al. 

paper presented at ISROMAC 13 (Ref. 5). The CFD analysis 

and figures of combustor geometry and flow are in the Ryder 

et al. paper (Ref. 6), also presented there. Compositional 

examination of the synthetic-paraffinic kerosene with the 

compositional-explicit distillation curve method is discussed in 

the Bruno and Bailbourine article (Ref. 7), who makes a useful 

comparison for heats of combustion based on molecular weight, 

volume, and mass. 

Combustor Thermal Performance 

The proprietary-geometry combustor, Combustor A, used 

for data collection represents a three-cup sector of a current 

engine combustor technology. Figure 1 shows a numerical 

model representation of Combustor A (Ref. 6). The Shouse et 

al. paper (Ref. 5) outlines the results given for the 225-psia 

(1.55-MPa) inlet condition, for three fueling compositions and 

three F/A values at 3 percent ΔP, and this report continues to 

analyze all four inlet combustor conditions at all three 

combustor ΔP values tested for the FT fuel blends of 0, 50, 

and 100 percent with JP–8 fueling. Data were taken at various 

nominal inlet conditions as follows: 

(1) FT fuel composition: 0, 50 (±5 percent), and 100 percent 

(2) Pressure (P) and temperature (T): 

 75 psia (0.52 MPa) and 500 F (533 K) 

 125 psia (0.86 MPa) and 625 F (603 K) 

 175 psia (1.21 MPa) and 725 F (658 K) 

 225 psia (1.55 MPa) and 790 F (694 K) 

(3) Combustor pressure drop (ΔP): 3, 4, and 5 percent 

(4) F/A: 0.010, 0.015, 0.020, and 0.025 

Combustor performance analysis assessments were made 

for the change in flame temperatures (Tflame), combustor 

efficiency, wall temperatures, and exhaust plane temperatures 

(Tplane). 

The combustion efficiencies do not provide enough insight 

for determining significant combustor changes, yet they do 

show a trend to decrease with increased %ΔP; thus, other 

aforementioned (Tflame and Tplane) parameters will also be 

investigated. At 75 psia (0.52 MPa) the combustion efficiency 

of all fuel blends are outlined in Table I (Ref. 5). 

Surface Thermal Measurements 

Thermocouples and pressure taps were placed throughout 

the combustion chamber to record temperature and pressure. 

The details of pressure drop measurements will not be 

presented in this paper, but it should be noted that no 

inconsistent pressure measurements were found. The 

convection-cooled liner and wall surface temperature 

measurements are noted as either “sidewall” or “liner” (which 

has inside and outside faces). 
NASA/TP—2013-217735 3 

TABLE I.—COMBUSTOR EFFICIENCY DATA SUMMARY 

[(P, T)inlet = (75 psia (0.52 MPa), 500 F (533 K)).] 

Fuel %P

3 4 5 

F/A F/A F/A

0.010 to 0.020 0.025 0.010 to 0.020 0.025 0.010 to 0.020 0.025 

Efficiency, percent 

JP–8 99.61 99.32 99.56 99.10 99.46 99.20 

50:50 99.65 99.24 99.56 99.09 99.46 98.90 

FT 99.66 99.20 99.58 98.96 99.48 98.87 

Average 99.64 99.25 99.56 99.05 99.47 98.99 

Standard 

deviation 0.024 0.058 0.011 0.080 0.013 0.181 

Sidewalls 

The forward, middle, and aft axial positions of the 

thermocouples along the sector combustor sidewalls are 

represented as “FWD,” “MID,” and “AFT.” Figure 2 

represents the sidewall temperature data obtained from the 

(P,T)inlet = [75 psia (0.52 MPa), 500 F (533 K)] and the 

(P,T)inlet = [125 psia (0.86 MPa), 625 F (603 K)] runs at 

4 percent P. This figure adequately represents the sidewall 

temperature trends shown for all inlet pressures and %P. 

Sidewall temperature profiles illustrate a decrease in 

temperature from the FWD to MID sections and an increase in 

temperatures from the MID to AFT along the combustor, 

where the temperatures are highest. The temperatures also 

increase as the F/A increases, the only exception being the F/A

of 0.025 at 75 psia (0.52 MPa) inlet pressure, which slightly 

decreases in temperature relative to the F/A of 0.020. The 

75 psia (0.52 MPa) data set is the only one that includes F/A

of 0.025. Data for 75 psia (0.52 MPa) at 3, 4, and 5 percent P 

at the F/A of 0.025 are consistent with this trend, with 

insufficient data to conclude whether this is a peak in 

combustor temperatures around F/A = 0.020. The sidewall 

temperatures depend strongly on F/A and weakly on the fuel 

blend composition. 

Unwrapped Combustor Liner 

Figure 3 is a representative plot of the unwrapped liner 

surface temperatures for F/A of 0.010 at (P,T)inlet = [75 psia 

(0.52 MPa), 500 F (533 K)], The term “unwrapped” refers to 

the normalized outside liner surface circumference (0 to 1) 

along with the normalized inner liner circumferential surface 

(1 to 2) as a continuous loop mapped onto a plane. Table II 

provides the combustor liner normalized circumferential 

and axial X coordinates along with the thermal data. The 

unwrapped combustor liner temperature profile shows a peak 

temperature increase measured by the thermocouple as %P

increases as well as an increase with F/A. 

The absolute inlet pressure and temperature of the system 

also shows an effect on the peak liner temperature (Fig. 3). 

Overall, the peaks in the temperature profile become more 

pronounced as F/A increases. They also become more 

pronounced as %P increases, but F/A appears to have the 

greater effect. 
NASA/TP—2013-217735 4 

TABLE II.—COMBUSTOR OUTER AND INNER LINER TEMPERATURES, THERMAL DIFFERENCE DATA, 

AND NUMBER AND NAME OF THERMOCOUPLES (TCs) USED TO PLOT THERMAL PROFILES.a,b,c

[From Ref. 5.] 

 F/A = 0.010 F/A = 0.015 F/A = 0.020 

TC Unwrappedd JP–8 FT 50:50 JP–8 FT 50:50 JP–8 FT 50:50 

TC 

no. 

X  Col C Col G Col F Col O Col K Col M Col Q Col V Col T 

Outer liner 

 0.00 

TOLAL 22 0.94 0.20 817 825 825 852 850 846 882 873 881 

TOLFL 20 0.00 0.22 900 904 907 968 935 946 1030 971 1002 

TOLML 21 0.67 0.26 861 870 869 927 924 925 987 957 985 

TOLMWA 24 0.19 0.32 870 875 878 962 917 944 1027 960 1002 

TOLFM 27 0.00 0.34 899 902 908 983 944 956 1060 985 1024 

TOLCA 25 1.00 0.52 862 872 872 920 910 912 986 952 974 

TOLAM 28 0.94 0.58 814 825 825 852 851 850 888 874 886 

TOLMR 36 0.67 0.62 873 887 884 953 918 941 1013 955 996 

TOLFR 34 0.00 0.74 887 888 891 965 924 944 1016 956 1001 

TOLMWI 23 0.33 0.79 879 879 884 933 905 914 992 937 969 

TOLAR 37 0.84 0.86 831 843 843 877 872 874 917 906 922 

TSWFD 41 0.22 0.97 1218 1288 1280 1438 1469 1444 1544 1565 1581 

TSWFT 30 0.78 1.00 814 820 820 847 836 841 854 842 850 

Inner liner 

TILMWI 38 0.50 1.20 897 908 906 994 969 976 1091 1023 1066 

TILFR 35 0.00 1.23 890 917 919 995 959 967 1075 1006 1050 

TILMWO 26 0.50 1.33 902 908 909 988 962 968 1096 1025 1071 

TILCA 29 1.00 1.50 1056 1065 1058 1229 1217 1219 1411 1326 1348 

TILFL 39 0.00 1.54 913 917 916 1002 970 981 1098 1030 1069 

 2.00 

a

Geometric position accuracy of thermocouple position coordinates is estimated at 1.5 percent. b

For nominal inlet pressure 225 psia (1.55 MPa), 800 F (700 K), and 3 percent combustor pressure drop. c

Note: TC no. and Col C, G, F, K, M, Q, V, and T are reference data set location parameters. 

d

X = x/L (which varies from 0 to 1), where x is the TC position measured from the liner inlet and L is the overall liner length. 

 = circumferential TC position measured over the liner outside y/Lθ (0 to 1) and continuing back along the inside liner (1 to 

2), where Lθ is half the unwrapped liner “width.” The normalized unwrapped coordinate (X, ) is the TC location (x, y). 
NASA/TP—2013-217735 5 

TABLE III.—AVERAGE UNWRAPPED LINER TEMPERATURE DIFFERENCES T:

a

 

FT FUEL AND 50:50 BLEND WITH RESPECT TO JP–8+100 

Inlet pressure and 

temperature, 

(P, T)inlet

%P, 

percent 

T = Tfuel blend – TJP–8,

b

 

F

F/A = 0.010 F/A = 0015 F/A = 0.020 F/A = 0.025

JP–8 FT 50:50 JP–8 FT 50:50 JP–8 FT 50:50 JP–8 FT 50:50 

[75 psia (0.52 MPa), 

500 F (533 K)] 

3 0 –4 –5 0 1 –6 0 –15 –11 0 –12 –16 

4 0 –2 0 0 –2 –2 0 –9 –7 0 –12 –8 

5 0 –3 –2 0 –1 –2 0 –2 –3 0 –3 1 

[125 psia (0.86 MPa), 

625 F (603 K)]

3 0 –6 –6 0 –20 –15 0 –28 –18 

4 0 1 –1 0 –6 –6 0 –29 –16 

5 0 6 2 0 –10 –5 0 –23 –11 

[175 psia (1.21 MPa), 

725 F (658 K)]

3 0 7 2 0 –25 –6 0 –39 –18 

4 0 1 0 0 –13 –10 0 –22 –14 

5 0 2 1 0 –11 –7 0 –26 –17 

[225 psia (1.55 MPa), 

790 F (694 K)]

3 0 12 12 0 –20 –13 0 –46 –16 

4 0 18 17 0 –9 –4 0 –31 –14 

5 0 6 3 0 –9 1 0 –39 –14 a

ΔT < 0 implies cooler liner temperature. b

T (F) = 1.8T (K). 

Using Figure 3, it is difficult to differentiate between the 

temperature differences of the varied fuel blends. By 

calculating the difference in temperatures read by the 

thermocouples at each location on the combustor for the blend 

and the 100 percent FT fuel relative to those recorded for 

JP–8, T = Tfuel blend – TJP–8 (where “fuel blend” refers to a 0 to 

100 percent FT fuel blend with JP–8), a better sense for each 

fuel’s performance may be obtained. Figure 4 illustrates the 

trends seen for the average temperature differences in relation 

to F/A at (P,T)inlet = [125 psia (0.86 MPa), 625 F (603 K)] and 

[225 psia (1.55 MPa) and 790 F (694 K)] at 3 percent P and 

gives a better picture for the heat performance of the fuels. 

Table III outlines all the average temperature differences 

(convective and radiation cooling) for all testing conditions. 

The overall trend shows that at F/A = 0.010, both the 

blend and the FT fuel generally run at higher temperatures 

than JP–8 for all pressure values. As the F/A increases, the 

50:50 blend and the FT fuel temperatures decrease, ending 

with cooler operating temperatures relative to JP–8 at the 

0.020 ratio. With respect to the increasing F/A, there is a 
NASA/TP—2013-217735 6 

larger deviation in the temperature performance of the FT and 

50:50 fuels. At the higher F/A, the FT fuel runs at 

temperatures cooler than both the JP–8 and the blend, 

illustrating that at these higher F/A values, the impact of the 

FT performance within the blend decreases and JP–8 

performance dominates. This would mean that at high F/A, 

the FT fuel would decrease liner temperatures relative to the 

JP–8 and the 50:50 blend and could increase component life 

within the combustor and yet not the turbine. Also, greater 

temperature differences are shown for higher inlet pressures 

and temperatures. 

Combustion Exhaust Rake Temperature 

The exhaust plane temperature trends are illustrated in 

Figure 5. These temperature profiles represent data-averaged 

temperature values collected through use of a temperature 

probe (rake) placed in the exhaust plane of the combustor. In 

these data, the signal from the top thermocouple was lost. For 

all data sets, there is a monotonic increase in the exhaust plane 

temperature as F/A is increased. The increase in %P also 

creates an increase in the temperature. At higher %P and F/A

values, the FT fuel tends to run at higher exhaust temperatures 

compared to JP–8 and the blend, which also gives slightly 

higher temperatures. Upon further analysis, as F/A increases, 

the FT fuel at higher exhaust plane temperature generally has 

less effect on the performance of the blend compared to that of 

JP–8. There is not a large temperature difference between the 

fuels at the higher %P and F/A (T = Tfuel blend – TJP–8 = 20 to 

50 °F (11 to 27 °C)) where there are larger differences at higher 

inlet temperatures and pressures; however, a small change in 

temperature can have major impact on the turbine life and 

efficiency, so these effects must be taken into consideration 

when selecting an alternative turbine engine fuel. 

Plotting the combustor exhaust rake temperature differences 

versus the inlet pressure and F/A (Fig. 6) displays a minimum, 

above which the variables have a positive effect on the 

combustor performance of alternate fuels over JP–8 fuels. 

Performing the same analysis for all %P conditions for 

both the 100 percent and 50:50 FT fuel temperature difference 

data, it is clear that increasing the inlet pressure increases the 

temperature differences of the FT fuel compared to JP–8. 

There also appears to be a peak in performance around 3 to 

4 percent P for the alternate fuel in relation to the 

performance of JP–8 (illustrated in Fig. 7). 

Calculated flame temperature data are outlined in Figure 8 

for 225 psia (1.55 MPa) at 3 percent P. As the inlet pressure 

and temperature is increased, there is a small increase in the 

flame temperature. The same trend is displayed with 

increasing %P, although %P does not seem to affect the 

temperature differences between the fuels to a significant 

extent. The %ΔP trend is more pronounced than that of the 

changing inlet pressure and temperature, especially when 

increasing F/A. At higher F/A, there is a greater difference in 

flame temperatures between the fuels. The FT generally had 

higher flame temperatures than the JP–8, and the 50:50 blend 

temperatures fell to temperatures between the FT and the JP–8 

fuel. These trends are also displayed in Table IV, which 

contains the flame temperature differences between JP–8 and 

the FT fuels. No significant trend determining whether the FT 

or JP–8 performance had the dominant role in the flame 

temperature performance of the 50:50 blend was found. 

Bester and Yates (Ref. 8) compared SPK to Jet-A1 fueling 

in an RR-Allison T63–A–700 Model 250–C18 B gas turbine 

engine and also cite approved FT-blended fuel use in civil 

aviation flying from South Africa based on UK Ministry of 

Defense DEFSTAN 91–91, Issue 3, 1999. It is gratifying to 

note that using a modified combustor engine system similar to 

that of Corporan et al. (Ref. 9), they found a 1.2-percent gain 

in turbine efficiency with SPK fueling over that of JetA-1 

along with more favorable emissions. Although the present 

report provides a more comprehensive and complete 

comparisons of SPK to JetA-1 combustor performance for 

current commercial aircraft (Fig. 1), it lacks the turbine engine 

performance afforded by Bester and Yates including the cited 

increase in turbine efficiency. 
NASA/TP—2013-217735 7 
NASA/TP—2013-217735 8 
NASA/TP—2013-217735 9 

TABLE IV.—FLAME TEMPERATURE DIFFERENCES T FOR FT FUEL AND 

50:50 BLEND WITH RESPECT TO JP–8+100 

[Combustion at (P,T)inlet = [225 psia (1.55 MPa), 790 F (694 K)].] 

%P, 

percent 

T = Tfuel blend – TJP–8, 

F (K) 

F/A = 0.010 F/A = 0.010 F/A = 0.010

Fuel blend Fuel blend Fuel blend 

FT 50:50 FT 50:50 FT 50:50 

3 73 (41) 55 (31) 94 (52) 43 (24) 34 (19) –3 (–2) 

4 107 (59) 90 (50) 84 (47) 41 (23) 105 (58) 42 (23) 

5 71 (39) 33 (18) 63 (35) 52 (29) 90 (50) 70 (39) 

Conclusions 

The data and analysis of combustor sector alternate fuel 

performance of Combustor A show lower average liner 

temperatures and higher flame and average combustor outlet 

temperatures with increasing Fischer-Tropsch (FT) relative to 

JP–8+100 fueling. The combustor outlet temperature affects 

turbine efficiency and blade and vane lives and may be due to 

the higher heat of combustion of the FT fuel per unit mass. 

Thus, the engine would not be running at higher fuel:air ratio

(F/A). A more accurate way to assess how these outlet 

temperatures will affect the turbine life would be to look at the 

pattern factor of the exhaust temperatures. 

Sidewall temperatures depend mainly on F/A for 

performance decreasing in temperature from the FWD to MID 

and increasing in temperature from the MID to AFT. The 

unwrapped liner temperature data show that the blend and the 

FT fuel run hotter than JP–8 at lower F/A, but cooler at F/A

above ~0.015. At the higher F/A values the FT fuel 

temperature differs more so from the JP–8 than the blend. 

Peak liner temperatures also increase with increasing F/A and 

%P but seem unaffected by the type of fuel blend to a 

significant extent. Lower liner temperatures result from 

decreased radiative heat transfer from reduced aromatic 

content. 

The 100 percent FT fuel tends to run at higher exhaust 

temperatures compared to the JP–8 and 50:50 blend fuels, 

with a similar trend for flame temperature. Overall, increasing 

F/A and %P increases the thermal performance of the 

combustor, which will almost always occur unless there is a 

decrease in the efficiency of combustion. 

Glenn Research Center 

National Aeronautics and Space Administration 

Cleveland, Ohio, September 10, 2013. 
NASA/TP—2013-217735 10 

References 

1. Federal Aviation Administration: Press Release—FAA 

Announces Alternative Energy Investments For Cleaner, More 

Sustainable Jet Fuel. 2011. 

http://www.faa.gov/news/press_releases/news_story.cfm?newsId=

13254 Accessed Sept. 24, 2012. 

2. DiMascio, Jen: U.S. Navy Buys 450,000 Gallons Of Biofuel. 

Aerospace Daily & Defense Report, McGraw-Hill Companies 

Inc., New York, NY, 2011. 

3. Wall Street Journal: The Cellulosic Ethanol Debacle. Review & 

Outlook, 2011. 

http://online.wsj.com/article/SB10001424052970204012004577

072470158115782.html?mod=WSJ_Opinion_AboveLEFTTop 

Accessed Sept. 24, 2012. 

4. Batt, R.J.; McMillan, J.A.; and Bradbury, I.P.: Lubricity 

Additives—Performance and No-Harm Effects in Low Sulfur 

Fuels. SAE Technical Paper 961943, 1996. 

5. Shouse, D.T., et al.: Alternate-Fueled Combustor-Sector 

Performance: Part A: Combustor Performance: Part B: 

Combustor Emissions. ISROMAC13‒2010‒49, 2010. 

http://ntrs.nasa.gov/archive/nasa/casi.ntrs.nasa.gov/2010001741

3_2010018362.pdf Accessed Sept. 24, 2012. 

6. Ryder, R., et al.: Computational Analysis of Dynamic SPK(S8)‒

JP‒8 Fueled Combustor-Sector Performance. ISROMAC13‒

2010‒61, 2010. 

http://ntrs.nasa.gov/archive/nasa/casi.ntrs.nasa.gov/2010001728

2_2010018363.pdf Accessed Sept. 24, 2012. 

7. Bruno, Thomas J.; and Baibourine, E.: Comparison of BiomassDerived Turbine Fuels with the Composition-Explicit Distillation 

Curve Method. Energy Fuels, vol. 25, 2011, pp. 1847‒1858. 

8. Bester, Nigel; and Yates, Andy: Assessment of the Operational 

Performance of Fischer-Tropsch Synthetic-Paraffinic Kerosene 

in a T63 Gas Turbine Compared to Conventional Jet A‒1 Fuel. 

GT2009‒60333, vol. 2, 2009, pp. 1063‒1077. 

9. Corporan, Edwin, et al.: Emissions Characteristics of a Turbine 

Engine and Research Combustor Burning a Fischer‒Tropsch Jet 

Fuel. Energy Fuels, vol. 21, 2007, pp. 2615‒2626. 

http://pubs.acs.org/doi/pdfplus/10.1021/ef070015j Accessed 

July 31, 2013. 

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Alternate-Fueled Combustor-Sector Performance 

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6. AUTHOR(S)

Thomas, Anna, E.; Saxena, Nikita, T.; Shouse, Dale, T.; Neuroth, Craig; Hendricks, Robert, 

C.; Lynch, Amy; Frayne, Charles, W.; Stutrud, Jeffrey, S.; Corporan, Edwin; Hankins, Terry 

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WBS 561581.02.08.03.16.03 

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14. ABSTRACT

In order to realize alternative fueling for military and commercial use, the industry has set forth guidelines that must be met by each fuel. 

These aviation fueling requirements are outlined in MIL-DTL-83133F(2008) or ASTM D 7566 Annex (2011) standards, and are classified 

as “drop-in” fuel replacements. This report provides combustor performance data for synthetic-paraffinic-kerosene- (SPK-) type (FischerTropsch (FT)) fuel and blends with JP-8+100, relative to JP-8+100 as baseline fueling. Data were taken at various nominal inlet conditions: 

75 psia (0.52 MPa) at 500 °F (533 K), 125 psia (0.86 MPa) at 625 °F (603 K), 175 psia (1.21 MPa) at 725 °F (658 K), and 225 psia (1.55 

MPa) at 790 °F (694 K). Combustor performance analysis assessments were made for the change in flame temperatures, combustor 

efficiency, wall temperatures, and exhaust plane temperatures at 3, 4, and 5 percent combustor pressure drop (DP) for fuel:air ratios (F/A) 

ranging from 0.010 to 0.025. Significant general trends show lower liner temperatures and higher flame and combustor outlet temperatures 

with increases in FT fueling relative to JP-8+100 fueling. The latter affects both turbine efficiency and blade and vane lives.

15. SUBJECT TERMS

Green aviation; Fueling; Energy; Particulates; Pollution control; Air pollution; Combustion product; Combustor design 

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