FINAL CFD DESIGN OF SCIROCCO ARC-JET TEST CONDITIONS FOR IXV TPS INTERFACES

Transcription

1 FINAL CFD DESIGN OF SCIROCCO ARC-JET TEST CONDITIONS FOR IXV TPS INTERFACES Davide Cinquegrana and Eduardo Trifoni CIRA, The Italian Aerospace Research Centre, Capua, Italy ABSTRACT Starting from the results of the study carried out during the phase C2/D of the ESA Intermediate experimental Vehicle (IXV) project, leaded by TAS-I as prime contractor, CIRA was in charge of the final design of the test conditions of the two qualification tests foreseen in SCIROCCO arc-jet facility on two different test articles made by CMC and ablative materials and representative of IXV TPS interfaces. The test conditions have been designed to reproduce on certain TPS interface locations of the test articles the maximum heat flux value predicted on the steepest IXV trajectory at a pressure level conservative with respect to the flight conditions, as per TAS- I requirements, for a duration equivalent to the shallow trajectory in terms of integral heat load. Two different test conditions have been defined to match the heat flux and pressure requirements on the two test articles with 3D CFD simulations performed with the CIRA thermal and chemical non-equilibrium CFD code. The test design loop was started preliminarily with 2D CFD simulations scaled to 3D with empirical coefficients to rapidly converge to the SCIROCCO operating point to be used as input for the more complex 3D CFD simulations at radiative equilibrium, fully catalytic, non catalytic and finite-rate catalytic wall boundary conditions. Based on finite-rate catalytic wall predictions, the hottest test article sections made of CMC materials were plotted in wall temperature and pressure coordinates to check the occurrence of undesired passive-to-active oxidation transitions. Key words: IXV; CFD; Plasma Wind Tunnel. 1. INTRODUCTION - IXV HEAT SHIELD MATE- RIALS AND REQUIREMENTS The ESA (European Space Agency) IXV (Intermediate experimental Vehicle) project, leaded by TAS-I (Thales Alenia Space-Italy) as prime contractor, aims to demonstrate European capabilities in hypersonic un-powered manoeuvring re-entry flight of a lifting configuration. The vehicle is also intended to serve as a test-bed for in-flight qualification of vehicle subsystems and systems, and to provide another source of data on fundamental Figure 1: IXV model with TPS. hypersonic aerothermodynamic phenomena for validation of tools, databases, and design processes. In this frame, a key role is played by the definition of the vehicle thermal protection system (TPS), since the success of the mission depends on the correct dimensioning of the TPS assemblies and right choice of TPS materials. In fact, the main function provided by TPS/HS is to protect cold structure of IXV from the heat loads experienced during the atmospheric re-entry and, furthermore, provide the stiffness to withstand the dynamic pressure and preserve the aeroshape on trajectory. The interface between assemblies of TPS subsystem, designed by TAS-I, could be the most complex and problematic part when solicited by high level of heat flux. From here the necessity to qualify materials for space flight through a test campaign to be conducted at SCIROCCO plasma wind tunnel facility. Fig. 1 resume, on the IXV vehicle, the TPS&HS subsystem that consist of Ceramic Nose and Windward Assembly (designed by Herakles), Ceramic Body Flap Assembly (designed by MT Aerospace) and Ablative TPS Leeward, Lateral and Base Assemblies (design by AVIO). The Fig. 2 shows the CAD reproduction of the two interfaces (labeled as 1 and 2 ) to be tested, indicated as IF#1 and IF#2, respectively, chosen because the worst heat flux is foreseen for those configurations. IF#1 is the interface between CMC Shingle and Hinge TPS (see details in Fig. 2), while IF#2 is the interface between Shingle, Hinge TPS and ablative P50 (zoom in Fig. 3). The Fig. 4 shows the real final configuration of the Test Article assembly to be tested for IF#2 test interface. In order to test IF#1 interface at required heat flux level, the ablative P50 panel will be replaced by a CMC one. This work follow the preparatory activities held before 2011 [BT12], and since that time some requirements are changed. IXV was launched from Korou in French

2 Figure 5: Heat and Pressure Load History. Table 1: Test Requirements. Figure 2: Zoom of IF#1 interface. IF#n q target [kw/m 2 ] p target [P a] Time/Heat Load IF#1 514±5% 3250±10% 710s/365[MJ/m 2 ] IF#2 170±5% 1755±10% 706s/120[MJ/m 2 ] Figure 3: Zoom of IF#2 interface. Guyana by VEGA and, after reach LEO (400 km), it reentry as a lifting body through atmosphere until to the splashdown in Pacific Ocean. Fig. 5 shows the heat and pressure load experienced by IXV from the atmospheric entry to the splashdown, so as foreseen during the design phase. In particular, the Figure shows thermo-mechanical loads, belonging to different control points monitored in the wall heat flux and pressure calculations. The loads regarding the interface of interest are the one highlighted by the circles. In Tab. 1 are reported the exact amount of heat flux and pressure that the TAA has to withstand, and also the test duration time to reach the required amount of heat and pressure load. Those values can be considered conservative, in fact for the heat fluxes it was considered the steepest descent trajectory, i.e. the worst case for the peak of heat fluxes, while for the heat loads, the shallow one [And89]. The value of heat flux required are to be reached on specific location of the Test Article that are different according to the IF1 and IF2. The hypothesis of wall condition are Fully Catalytic behaviour and in radiative thermal equilibrium, with a constant emissivity value supposed equal to ɛ = 0.8. This kind of approach is surely conservative if compared to a finite rate hypothesis. 2. TEST DESIGN LOOP Figure 4: IF#2 Test Article configuration. The Fig. 6 resume the various steps followed to design each test condition aimed to reach the flight conditions or customer requirements. Summarizing the graph, a first attempt on the reservoir conditions, in terms of total pressure and enthalpy (P 0, H 0 ), are argued by considering both the feasibility in the Facility and of a tools based on CFD and experimental data from previous tests. From those, axis-symmetric CFD simulation of nozzle are conducted and the condition in the test section are obtained,

3 Figure 6: Loop for the experimental test design. for the Nozzle D configuration. The test section conditions foreseen by CFD are treated as free-stream boundary conditions for the preliminary bi-dimensional CFD of the Test Article OML, evaluating heat flux and pressure level effectively reached. This is a first trial and error stage, where also TA attitude is investigated. During this inner loop, the tailoring of the reservoir conditions can be made employing empirical correlations, reported in Eq. 1, and derived from Fay-Riddel and isentropic relations for stagnation pressure [And89]: ( ) p0 p 0 = p st,t H 0,T = p st Q st Q st,t p0 p st (H 0 c p T st ) + c p T st,t (1) here the subscript T is related to the target values and indicates the new PWT test condition in terms of total enthalpy and total pressure. The current values, signed with subscript st are the CFD data coming from 2D simulation and scaled with coefficients that correct it to obtain 3D values. This empirical scaling is based on the local ratio obtained comparing a large DB of 3D and 2D CFD runs. Fig. 7 shows those coefficients relative to different zones of the TA geometry: in particular was evidenced the symmetry plane and the slice passing on the final part of the TPS near the TA corner, and the latter represents the worst condition for the wall heat flux. After reached the target value 3D simulation were conducted, and, if necessary, further tailoring to the reservoir condition can be applied. Finally, an axis-symmetric CFD simulation of the Calibration probe at the final reservoir conditions is needed in order to provide pressure and heat flux at the probe stagnation point. This calculations will be useful during the test to check if in the test chamber the stagnation conditions are reached comparing the value read by the probe sensor with the numerical one. 3. NUMERICAL APPROACH Numerical CFD simulations conducted to design the test at SCIROCCO have been carried out with the CIRA code CAST [RB06], that solves, on a multi-block structured grid, the Reynolds Averaged Navier-Stokes equations in Figure 7: Scalar coefficients employed for 2D to 3D scaling of Heat Fluxes ( q ) and Pressure ( p ). a density-based finite volume approach, with a cell centred, Flux Difference Splitting at second order, ENO-like, upwind scheme for the convective terms. The code, in parallel version, has been run both on the CIRA Cluster Jet equipped with Linux Redhat OS, 14 nodes with Intel XEON E GHz - 8 core processors, 64 GB of RAM memory for each node and a storage capability of 48 TB. The physical modelling available inside the numerical code include non-viscous (Eulerian) fluid flow and viscous laminar/transitional/turbulent flow with air modelled as an ideal gas or in both thermo-chemical non-equilibrium and equilibrium, this latter option being a good compromise between accuracy of predicted results and CPU-time consumption. While the 2D simulation was conducted on single processor, the 3D run was parallelized on 16 cores by a METIS algorithm Computational Grids The computational grids designed for the numerical computations of PWT nozzle, Test Article and calibration probe, was designed with ICEMCFD software. All grid was structured, with quad and hexahedral cells, stretched across the boundary layer with the heigh of the first cells from the wall that ensure an y + = 1 for the correct treating of the boundary layer and then for wall flux estimations too. The Fig. 8 shows the structured, 2D multiblock grid that was employed to simulate the symmetric OML of the Test Article, composed by 8 blocks and 28k cells. In Fig. 9 is showed the structured 3D multiblock for the complete TA CFD simulations, composed by 58 block and 1.32M cells. The Fig. 10 and 11 shows some details of the step of the TPS thickness over the TA assembly. 4. TEST ARTICLE ATTITUDE As depicted in the design loop of Fig. 6, in the stage concerning preliminary 2D CFD analysis, a study of TA attitude is foreseen. Also during the IXV tests design a

4 Figure 8: 2D computational Grid. Figure 12: Wall Heat Flux distributions of 2D simulations varying AoA. Figure 9: 3D computational Grid. Figure 13: Wall Pressure loads distribution of 2D simulations varying AoA. Figure 10: Zoom of 3D computational Grid. sensitivity studies of the Attitude of the test article on the heat flux and pressure loads distributions was conducted. This investigation has highlighted that on the wall heat fluxes distributions, negligible effects are foreseen in the region of IF#1 and IF#2 interfaces, as showed in Fig. 12, where the heat flux are calculated with the Global efficiency approach. The reservoir conditions are relative to an intermediate steps through the PWT settings design loop. Fig. 13 refers to the wall pressure loads distributions: here the loads on the flat plate is moving away from the uniform distribution increasing AoA. For these loads features and also for a technical motivation due to the frontal section of TA assembly, that exceed to a value not allowed in the PWT test section, the 45[deg] setting was chosen as the final one. 5. Figure 11: Zoom of 3D computational Grid. WALL CATALYSIS TREATMENT As stated in the beginning of test design phase, the target to be reached in terms of heat flux was to be calculated guessing a wall Finite rate hypothesis [BT12]. As a

5 Figure 14: Catalytic efficiency coefficient as empirically derived at VKI. Figure 16: Test Campaign (FLPP program) adopted to define Global catalytic efficiency parameter. Table 2: Comparison between numerical and experimental wall Temperature and heat flux at G2 T w [k] q w [kw/m 2 ] CFD NC CFD FC exp G RESULTS Figure 15: Heat Flux loads over 2D article CFD run with different Catalytic Surfaces hypotheses (IF#1). model to be implemented in the CIRA code, experimental data obtained on a small sample by VKI was available in a tabular form, showed in the graph of 14. During the iterative phase, this model had shown a very low catalytic behaviour of both the IXV CMC materials (see Fig. 15). With experience coming from recent Campaign on FLPP Shingle tested in SCIROCCO PWT[BRMT08], a global approach was adopted [TBC15], defining λ w coefficient of TPS, derived numerically from data showed in Tab. 2. Follows that: q exp = q NC + λ w ( q F C q NC ) (2) λ w = ( q exp q NC ) ( q F C q NC ) (3) Considering data in Tab. 2 that refers to the point labelled as G2 in Fig., follows the numerical value: λ w = (4) Even if the coefficient λ w is coming from a punctual measurement, the author of the research [TBC15] shows that also in different points, with different flows behaviour, the ratio is very close to the value adopted in this work. This section shows the CFD results of 3D numerical simulations, obtained after that the preliminary loop was conducted on the 2D geometry based on the OML of the TA Assembly. Starting from the preliminary reservoir conditions of the two tests, Fully-catalytic and Non-catalytic 3D Navier-Stokes with thermochemical non-equilibrium chemistry was conducted. Fig. 17 shows the results in terms of heat flux and pressure load over the TA, with a Fully Catalytic wall in radiative equilibrium (ɛ = 0.8). The IF#1 Control point is located with an offset of 7.8 cm from the symmetry axis above the curved shingle. There, with a H 0 = 20.7[MJ/kg] and P 0 = 7.0[bar], a wall heat flux of 512[kW/m 2 ] and a wall pressure of 2946[P a] is foreseen, well within the range of the target of Tab. 1. Fig. 18 instead, shows the results in terms of wall heat flux and pressure load over the TA, in the same wall conditions, at IF#2 settings. Here, H 0 = 10.0[MJ/kg] and P 0 = 4.5[bar] are necessary to reach a wall heat flux of 161[kW/m 2 ] and a pressure load of 1739[P a] at the end of curved shingle on the symmetry plane. To a better identification of target for IF2, Fig. 19 shows in green the regions where the level of heat flux and pressure that are within the range of the requirements. While the wall pressure shows an uniform distributions over the plane region until the curved shingle, the wall heat flux shows a different behaviour with a narrowed re-

6 Figure 19: Requirements interval highlighted for Heat Flux and pressure in IF2 case. Figure 17: Heat Flux and Pressure Contour over TA for IF1 with FC wall hypothesis. Table 3: Numerical Assessment of Requirements. Fully Catalytic wall T.A. Numerical PWT Settings loc. q fc [kw/m 2 ] p fc [P a] H 0 [MJ/kg] P 0 [bar] IF IF Figure 18: Heat Flux and Pressure Contour over TA for IF2 with FC wall hypothesis. gion in the IF2 zone of interest. The Table 3 resume the SCIROCCO reservoir conditions obtained at the end of the design loop described before, for both the interface requirements. Both the final settings allows to met the requirements, considering the tolerances of Heat flux and pressure, reported in Table 1. The wall heat flux attained shows differences that are below the 5% tolerance, especially for IF1 (< 1%). For both the test conditions, 3D solutions are also obtained with non-catalytic behaviour imposed as boundary conditions at wall. With this results, and applying the global approach for the catalysis mentioned before, with the relations in Eq. 2, it was possible to foreseen the heat flux and the wall temperature for the actual case of finite rate catalysis at wall. Fig. 20 shows the wall temperatures contour (ɛ = 0.8) for the IF#2 conditions, obtained applying FC condition for the plain plate, FRC for curved shingle. Here it is possible to appreciate a catalytic jump at the interface between the two plates (CMC-P50 materials) where temperature of the order of 1900[K] are foreseen. In terms of heat flux and for the worst scenario (i.e. Non-Catalytic/Fully Catalytic transition), the foreseen overshoot is quite near the 300% at interface, if compared to the uniform Fully catalytic heat flux distribution. Based on finite-rate catalytic wall predictions, the hottest test article sections made of CMC materials were plotted in wall temperature and pressure coordinates to check the occurrence of undesired passiveto-active oxidation transitions. In particular, for IF#1 test

7 Table 4: Numerical Assessment of Requirements. FRC hypothesis T.A. Numerical PWT Settings loc. q frc [kw/m 2 ] p frc [P a] H 0 [MJ/kg] P 0 [bar] IF IF Figure 20: Temperature distribution with Catalytic jump for IF2. Figure 22: CFD results in terms of Mach and Pressure Contour for test IF1 Probe. (a) IF#1 conditions, Fig. 21a shows the point relative to the slices passing through IF#1 points (z = 0.078[m]) and on the corner side (z = 0.2[m]), valued with global efficiency hypothesis. The hottest section at corner proximity reach a zone that is just before the limit of active oxidation. Fig. 21b shows the results for IF#2 test conditions with slices passing through the symmetry plane and on the corner side: for this case the results shows a safer conditions if compared with IF#1, far from the active oxidation limits. The values of heat fluxes and pressure foreseen with the Global Efficiency model are resumed in Tab. 4. Defined the final PWT operating conditions, numerical simulation over the hemispherical probe was conduced. The stagnation heat flux and pressure foreseen at Fully Catalytic, cold wall condition (T w = 300[K]), will be employed as a reference conditions. In fact, those values, to be measured during the test, ensure the achievement of the desired operating condition (H 0, P 0 ) in test chamber. Numerical values of the stagnation heat flux and pressure are reported in Tab 5. Finally, Fig. 22 and 23, show the contours of Mach number and pressure surrounding the calibration Probe realized for the two test conditions, respectively, for IF#1 and IF#2. (b) IF#2 Figure 21: Heat Flux slices on the Oxidation Map for IF#1 and IF#2 with FRC wall hypothesis. 7. CONCLUSIONS This work has showed the numerical activities carried out to support the SCIROCCO Plasma Wind Tunnel assess-

8 Figure 23: CFD results in terms of Mach and Pressure Contour for test IF2 Probe. Table 5: Numerical Assessment of PWT Probe, FC-cw hyp. on a large tps demonstrator. Journal, InTech Open [BT12] S. Di Benedetto and E. Trifoni. Aerothermodynamic design of scirocco plasma wind tunnel tests for ixv tps interfaces. In International Astronautic Conference 2012, Naples, September [RB06] G. Ranuzzi and S. Borreca. H3ns: Code development verification and validation. Technical report, CIRA, [TBC15] E. Trifoni, S. Di Benedetto, and M. Di Clemente. A new global approach to evaluate the surface catalytic efficiency of tps materials based on arc jet experiments and numerical predictions. In 8th European Symposium on Aerothermodynamics for Space Vehicles, March PWT Settings CFD Probe loc. H 0 [MJ/kg] P 0 [bar] q s [MW/m 2 ] p s [kp a] IF# IF# ment, foreseen for the test of IXV TPS interface, at two different set points. The activities led to a pre-test DE- SIGN, where TA attitude and finite rate catalysis of the TA surfaces was defined by means of 2D CFD numerical analysis, and successively followed by 3D CFD simulations to define in details the final PWT settings of the two different conditions. Once achieved the final PWT settings for both IF tests, wall heat flux with Finite rate catalysis, valued with Global Efficiency approach, are also conducted. This data was employed to check if the passive-active oxidation transition could happens at this conditions, but this is not the case for both. ACKNOWLEDGMENTS We wish to acknowledge the IXV Teams: ESA, MT Aerospace, Safran Herakles and TAS-I. REFERENCES [And89] J. Jr. Anderson. Hypersonic and High Temperature Gas Dynamics. McGraw-Hill Book Company, New York, [BRMT08] Sara Di Benedetto, Giuseppe C. Rufolo, Marco Marini, and Eduardo Trifoni. Rebuilding and analysis of a scirocco pwt test