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Progress Report on Preliminary Design of the LE-X Turbopump
Makoto.Kojima1 and Akihide.Kurosu2 Japan Aerospace Exploration Agency, Tsukuba, Ibaraki, Japan, 305-8505
Masaharu.Uchiumi Satoshi.Takada and Keiichiro.Noda Japan Aerospace Exploration Agency, Kakuda, Miyagi, Japan, 981-1525
and
Tsutomu.Mizuno3 IHI Corporation, Mizuho, Tokyo, Japan,190-1297
JAXA has been conducting technology demonstration of the next generation booster engine called LE-X. The LE-X engine is a new cryogenic booster engine which has higher thrust, higher reliability and lower cost compared to the existing engines, and will be applied to the Japan’s next primary launch system. Its engine system adopts the expander-bleed-cycle run on the LH2/LOX. The LE-X turbopumps, which are key components in the engine, have technical issues of developing rotordynamics with huge turbine, high turbine efficiency, high reliable rotor system and low-cost manufacturing technique. In order to clear the issues and to avoid the risks at preliminary design phase, several component tests and analysis were conducted aiming to validate the preliminary design. It appears that preliminary design of the LE-X components is successfully conducted. Feasibility of the components designs and the manufacturing process will also be confirmed in the preliminary design phase. Based on these studies, we will continue the turbopumps design of the prototype LE-X engine.
I. Introduction apan Aerospace Exploration Agency(JAXA) has been conducting technology demonstration of the Japan’s next generation booster engine called LE-X. 1-4 The LE-X engine is a new cryogenic booster engine which is capable
of making a higher thrust, higher reliability and lower cost compared to the existing engines, and will be applied to the Japan’s next primary launch system. This engine driven by LH2/LOX adopts the expander-bleed-cycle as an engine cycle system.
It is generally difficult to apply the expander-bleed-cycle to a booster engine because such engine requires turbopumps to have high turbopump efficiency and large shaft power. In order to achieve the high turbopump efficiency, the LE-X turbopumps have huge turbine. Therefore as compared with Japanese traditional rocket engine turbopumps, the LE-X turbopumps have to overcome technical issues of rotordynamics with huge turbine, high turbine efficiency, high reliable rotor system and low-cost manufacturing technique. The preliminary design of the LE-X engine system was performed in the past several years. In order to attain an optimal balance among required performance, high reliability and low cost, the LE-X turbopumps are being studied using advanced simulation technologies. At first, we listed technical issues, supposed failure modes and risks of turbopump system, pump, turbine, bearing and seal for each turbopumps. In order to clear the issues and avoid the risks at preliminary design phase, several component tests and analysis were conducted to validate the preliminary design. This paper reports our progress of a preliminary design and analysis for the LE-X turbopump.
1 Engineer, Space Transportation Propulsion Research and Development Center, [email protected] 2 Associate Senior Engineer, Space Transportation Propulsion Research and Development Center, AIAA Member 3 Professional Engineer, Space technology Group, Reseach & Engineering Division.
J
47th AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit31 July - 03 August 2011, San Diego, California
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II. Specification of the LE-X turbopumps The LE-X engine adopts chamber expander-bleed-cycle system and realize throttling operation. The fuel
turbopump is driven by regeneratively heated hydrogen gas with a temperature of about 550K. As the energy of turbine driving gas depends on the combustion chamber heat load, temperature of the gas is not so high compared with that of the staged combustion cycle or the gas generator cycle which utilizes the combustion energy. Therefore a technical challenge is to generate the required shaft power of the turbopump by the turbine driving gas with low enthalpy and low flow rate. Figure 1 illustrates the reference characteristics and schematic diagram of the LE-X. The characteristics of the LE-X turbopump is summarized in Table.1. The 100% vacuum thrust of the LE-X engine is set at about 1,500 kN. The shaft powers of the fuel turbopump (FTP) and the oxygen turbopump (OTP) are approximately 15,000 kW and 6,000 kW, respectively. The key design of each component is described in more detail in the following section.
III. Fuel turbopump Figure 2 presents the three-dimensional model of the LE-X FTP. The LE-X
FTP suctions about 50 kg/s of liquid hydrogen (LH2) from a fuel tank and pressurizes it from 0.5 MPa to 18 MPa so as to inject it into the high pressure combustion chamber. In order for the turbopumps of the expander-bleed-cycle engine to achieve high large shaft power required for booster engine, high turbine performance is required.
A. Rotor system High turbine efficiency and large shaft power of turbopumps are required
with low-enthalpy turbine driving fluid. Therefore, one of the characteristic aspects of the FTP rotor system is its huge turbine pitch circle diameter. Since such a rotor system causes a large overhang, a large amplitude shaft vibration is induced without an appropriate damping or an adequate shaft stiffness. In addition, in general, the rocket turbopumps are forced to be operated at high speed conditions for smaller and lighter body. This leads us to spend much time and energy trying to reduce the rotor vibration in the presence of unsteady hydrodynamic forces in the rotating impeller and inducer, and destabilizing forces of the seals and Thomas forces of turbines. Higher bearing stiffness causes lower damping. However, lower bearing stiffness causes lower margin for avoiding critical speed. Since the FTP is designed to be operated at higher speed than the first critical speed, the shaft diameter, shaft length and bearing of the FTP is designed in order to attain an optimal balance among internal damping, overhang, bearing stiffness and critical speed.
B. Inducer A two-stage inducer is adopted to achieve high pump efficiency, high pump-head and short-length disk-shaft for
the FTP. Hydrodynamic characteristics, cavitation instability and rotor dynamics of the two-stage inducer are evaluated by using computational fluid dynamics (CFD) analyses. Based on the results, the number of rotor blade was decided for each stage. In addition, the design parameters which affect the inducer performance and instability were extracted. To determine the optimum inducer design, each design parameters were allocated to the orthogonal array. A relationship between the evaluation functions and the control parameters are visualized by the factor effect
Table 1. LE-X turbopump characteristics.
Unit FTP OTP
Pump mass flow rate kg/s 49.7 293.3
Turbine mass flow rate kg/s 7.9 6.8
Pump pressure rise MPa 17.6 19.2
Turbine expansion ratio - 9.3 1.4
Pump efficiency - 0.75 0.75
Turbine efficiency - 0.5 0.7
Power kW 16148 6277 Figure 1. LE-X’s reference characteristics and diagram.
Figure 2. Three-dimensional model of the LE-X FTP.
LH2
LOX
MOV
MFV
CCV
MIX
FTP
OTP
MCC
NE
TCV
LH2
LOX
MOV
MFV
CCV
MIX
FTP
OTP
MCC
NE
TCV
Engine cycle:Chamber expander bleed
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diagram. Using the factor effect analysis, response surface of each evaluation function is calculated. Based on the calculated results, the optimum inducer design was determined. In addition, experimental water flow test was separately performed to investigate the performance of the inducer. Pump-head and suction performance are in agreement with the analysis results.
Presently, JAXA is conducting benchmark activities with CNES in France to mutually improve cryogenic inducer design/evaluation technologies. After designing and fabricating each inducer, the benchmark tests were performed at each test facility. The tests were performed with water and liquid nitrogen (LN2) by JAXA, and with water and Freon by CNES. Figure 3 shows a photograph of water tunnel test of JAXA inducer. Suction performance, cavitation performance, cavitaion instability and thermodynamic effect were evaluated in these tests. In the performance evaluation, comparison of CFD results with test results of both JAXA inducer and CNES inducer indicated good agreement. Furthermore, the performance of the designed inducer met the requirements of LE-X reference characteristics.
In the cavitation performance, significant difference in the head break point between CFD and experimental comparison with LN2 water and Freon case is seen in Figure 5. In addition, there are significant differences in the head break point among each flow rate as seen in Figure 6. When the flow rate at the inducer decreases, the differences in the head break point widen toward lower cavitation number. That is because of the well-known thermo-dynamic effects. In the cavitation instability, AC (Attached Cavitation) and RC (Rotating Cavitation) is observed in some tests of both of JAXA inducer and CNES inducer. Figure 4 displays the trial result of unsteady CFD performed to estimate cavity shape. JAXA and CNES exchanged both experimental test and CFD results each other to improve each analyzing technique.
C. Impeller As mentioned before, because the LE-X engine adopts expander-bleed-cycle, its pump discharging pressure is
not so high as that of the staged combustion cycle. Figure 7 shows pump-head and efficiency of LE-X compared with LE-7A that adopts staged combustion cycle. It is possible to achieve the reference characteristics of pump-head of the LE-X with a two-stage inducer and a single-stage impeller. A single-stage impeller is adopted to achieve high pump efficiency and short-length disk-shaft. In order to realize the robust throttling operation, the shaft length is an
comparison
Figure 4. Cavity shape of JAXA’s inducer predicted by unsteady CFD.
Hea
d co
ef. (-
)
CP
Exp J2 water@6000rpm, 90%Q with step casingExp J3 LN2@9000rpm, 90%Q without step casingExp J2 Water@6000rpm, 100%Q with step casingExp J3 LN2@9000rpm, 100%Q without step casingExp J2 Water@6000rpm, 110%Q with step casingExp J3 LN2@9000rpm, 110%Q without step casing
Figure 5. Cavitation performance of JAXA’s inducer from CFD analysis and experiment results.
Figure 3. Water tunnel test of JAXA’s inducer.
Figure 6. Cavitation performance of JAXA’s inducer in each flow rate.
0.00
0.20
0.40
0.60
0.80
1.00
1.20
Hea
d ra
tio (
H/H
ref)
Cavitation Parameter
CFD J3 LN2@9000rpm without step casing
Exp J3 LN2@9000rpm without step casing
CFD J3 Water@5000rpm without step casign
Exp. J2 water@6000rpm with step casing
CFD J3 Freon@5000rpm without step casign
Exp. J1 Freon@5000rpm with step casing
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important parameter for avoiding the critical speeds. The simple configuration of single-stage impeller allows a working operation of the rotor under the second critical rotor speed, and also contributes to accomplish the cost reduction and high reliability.
We consider the adoption of open impeller to single-stage impeller of FTP to reduce the cost. To confirm the feasibility of open impeller design, an experimental open impeller was designed based on the closed impeller of LE-7A FTP. Hydrodynamic characteristics and structural strength of the open impeller are evaluated by combining CFD and finite element method (FEM). The FEM result presented in Figure 8 shows the high stress at the center of the impeller and the base of the blade. In order to validate the FEM result, the rotational speed of impeller will be increased to the point of burst in the impeller spin test which is planed to be carried out this year. Then, the optimum design is determined by a method that each design parameters were allocated to the orthogonal array and using the factor effect analysis and response surface similar to that for the inducer. Based on the results, experimental water flow test was performed to investigate the validity of the design approach for the open impeller. The three-dimensional model of the FTP impeller is shown in Figure 9.
The optimum design of the diffuser and swirl breaker were also determined by similar method for the inducer. In the design process of the swirl breaker, a CFD analysis was performed to estimate the pressure distribution and the swirl speed ratio. An example of the CFD result is shown in Figure 10. As seen in Figure 10, the flowfield is complicated in the swirl breaker. Based on these CFD results, the relationship between the evaluation functions and the control parameters are visualized by factor effect diagram. It was confirmed by these CFD results that the groove of the swirl breaker made pressure variation. Therefore, the number of grooves were designed properly to avoid the natural frequency of the impeller.
Volute and guide vane were designed to meet the optimum design of impeller and inducer. Figure 11 shows a CFD result of the volute to estimate pressure distribution in the volute. This confirms that separated flow causes the flow velocity bias was seen in the initial design. The velocity causes efficiency loss. Then the design of impeller was modified. Thus, the impeller design was improved and the flow velocity bias was reduced in the CFD result. Experimental water flow test with inducer are planned to validate the preliminary pump design.
Figure 8. FEM result of FTP impeller.
Figure 9. Prototype of FTP impeller.
Figure 10. An example of CFD result of swirl breaker.
Figure 11. Flow velocity distribution CFD result of volute. left: initial design. right: improved design.
Figure 7. Pump head and efficiency of LE-X and LE-7A.
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D. Turbine The FTP is recognized as a potentially high risk component because the turbine efficiency is one of the main
drivers of engine system performance. Therefore, it is important to estimate turbine efficiency precisely. In addition, the expander bleed cycle needs high-pressure ratio turbine for FTP due to its configuration. 2stage impulse turbine is adopted for FTP. The performance and structural strength of the stator vane and rotor vane of supersonic turbine were estimated by CFD / FEM. Then, the optimum design was determined by using the orthogonal array similar to inducer.
As other design approach, we use genetic algorithm and unsteady CFD to achieve efficiency optimum. Figure 12 shows an example of unsteady CFD result. It was confirmed that the efficiency target can be achieved in each analysis. However, reflection of shock wave at 1st stator and separated flow at 2nd stator were confirmed in these analysis. To prevent these issues, minor change was conducted at these blades design. Figure 13 shows the CFD result of the modified design. In order to verify the CFD analysis, turbine rig test is underway conducted. In the test, turbine efficiency is confirmed.
In addition, turbine manifold was designed to meet the optimum design of turbine. Figure 14 shows CFD result of pressure distribution in the turbine manifold.
E. Bearing and Shaft seal The bearing and shaft seal are key components for high reliability. The bearing and shaft seal for the FTP of the
LE-X engine are required high reliability compared with that of the LE-7A despite the condition of high load and high rotational speed. A hybrid ceramic ball bearing is applied to the FTP. A lot of fundamental tests of the bearing were conducted in Japan. The hybrid ceramic ball bearing is expected to reduce amount of heat generation in the bearing. Since a rotor system of the FTP which has huge turbine causes a large overhang, high stiffness bearing is required. Taking into account the damping of the casing, a large-diameter bearing was adopted to achieve the target of the system stiffness. Similarly, the number of ball, the contact angle and raceway curvature were determined to achieve the target of contact pressure and radial clearance. On the other hand, various seal types were compared in
Figure 12. An example of unsteady CFD result of turbine.
Figure 14. CFD result of FTP turbine manifold.
Figure 13. CFD result of the modified design.
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terms of performance, cost and development risk for the FTP shaft seal. As the result of the trade-off, mechanical seal type was selected. However, each part has its own technical challenges. Because of the thick shaft, contact area of the seal is large. High rotating speed and differential pressure are unfavorable conditions for the seal. Additionally, seal space should be as reduced as possible to reduce the overhang. The bearing and shaft seal system test will be conducted to mitigate the considered risks within a year.
IV. Oxygen turbopump The oxygen turbopump consists of a single-stage impeller with a traditional single-inducer, and a two-stage
turbine. A three-dimensional model of the OTP is shown in Figure 15. The OTP turbine efficiency is also an important parameter as well as the FTP because the turbine efficiency gives a large impact on determining the turbine diameter. The inlet temperature and pressure of OTP turbine are low compared to the FTP, therefore relatively large turbine is required to generate sufficient energy. Similar to the results in the FTP, the performance and structural strength of the turbine were estimated by CFD / FEM. Then, the optimum design was determined by use of the orthogonal array similar to that for the FTP. A rigid rotor is usually adopted for stabilization, but the high rotational speed of a rocket turbopump blocks a working operation of the rotor under the first critical rotor speed for a general rotor configuration. The huge turbine will cause a large amplitude shaft vibration. The turbine side bearing is located in the back of the turbine disks to realize the rigid rotor. We considered the adoption of oil-lubrication type to the bearing. A photograph of the oil-lubrication bearing is shown in Figure 16. However, each part has some technical issues. One issue is operational temperature range of oil-lubrication bearing. The ambient temperature of the bearing is expected to drop to about 160 K during the pre-cooling, and to rise to about 460 K during operation. Therefore, usable temperature range of oil-lubrication bearing needs to be clarified. Candidate oil with wide temperature range was selected from synthetic, JET, fluorine and cyclopentane oil. The rotational tests were performed to investigate the torque of each selected oil in low ambient temperature. Figure 17 shows torque characteristics of JET oil as an example of the rotational test result. As shown in the result, rapid increase of torque was confirmed at about -80℃. When the starting temperature decreases, the starting torque rises. From these results, operational temperature range of each oil was evaluated.
The other issue is reactivity with hydrogen. Because the LE-X engine which adopts the expander-bleed-cycle is run on LH2/LOX propellant, the ambient gas of the bearing is hydrogen. Thus, oil reactivity with hydrogen needs to be clarified. We conducted further hydrogen reactivity tests to examine corrosion of several engine materials. The hydrogen reactivity test base material is summarized in Table.2. The high temperature rotational test will be conducted to investigate amount of heat generation of the bearing. The pump side slinger was designed, and the rotational test is underway conducted to examine whether it is necessary or not. Figure 18 is photograph of the rotational test of the slinger. Finally, a bearing and shaft seal system test with a dummy inducer, impeller and turbine will be conducted to mitigate the risks within two years.
Figure 16. Oil-lubrication bearing.
Oil-lubrication bearing
Slinger
Figure 15. A 3D model of the OTP.
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V. Conclusion At the present stage, it appears that the preliminary design of the LE-X components is successfully conducted.
As a next step, the feasibility of the components designs and the manufacturing process will be examined. Based on these studies, we will continue to design the turbopumps for the prototype LE-X engine.
Acknowledgments The authors wish to acknowledge member of the inducer cooperation between CNES and JAXA, including
Mr.B.Pouffary and Mr.M.Illig.
References Papers
1 Uchiumi, H., Kojima, M., Okita, K. and Mizuno, T.: A preliminary design study of rotor system for LE-X turbopump, AJCPP2010-069.
2 Kojima, M., Sunakawa, H., Kurosu, A., Uchiumi, M., Okita, K., Ogawara, A., Onga, T., “Preliminary Design and Analysis for the LE-X Engine Components,” AIAA2009-5485.
3 Kurosu, A., Sunakawa, H., Yamanishi, N., Okita, K., Ogawara, A. and Onga, T., “LE-X –Japanese Next Liquid Booster Engine–,“ AIAA2008-4665.
4 Kurosu, A., Sunakawa, H., Kojima, M., Yamanishi, N., Noda, K., Ogawara, A., Tamura, T., Mizuno, T. and Kobayashi, S., “Progress on the LE-X Cryogenic Booster Engine,“ 4th EUCASS.
