EUROPEAN CONFERENCE FOR AEROSPACE SCIENCES (EUCASS)

Introduction

The Automated Transfer Vehicle (ATV) is a supply cargo for the International Space Station (ISS), constituted of two elements. The first one is the SCA (Space-Craft Subassembly) equipped with a propulsion bay (propulsion tanks and thrusters) and an avionics bay (batteries, gyroscopes and harness). The second part is the Integrated Cargo Carrier (ICC), containing the equipped external bay (water and gas tanks, web structure) and the equipped pressurized module (containers, cargo and the attitude control thrusters). A sketch of the ATV is shown in Figure 1.

At the end of its mission to ISS, during its re-entry into Earth atmosphere, ATV is exposed to a hypersonic flow associated with high heat-loads leading to the structural heating and the fragmentation of the vehicle. It has been concluded that residual hypergolic propellants may ignite and explode upon exposure to the hot and reactive flow

environment. Depending on the location and type of explosion, this could lead to a sudden destruction of ATV at higher altitude, modifying significantly the impact footprint of vehicle fragments on Earth, with potential implications for the safety of the population on the ground.

This study proposes a refined assessment of the explosion potential during ATV re-entry, taking into account a perforation of the spacecraft body resulting in internal flows an in temperature and pressure increases within the vehicle. In a first step, a detailed trajectory analysis [1] of ATV re-entry has been performed and several trajectory points have been selected. The flow-field around the vehicle has been simulated for the selected trajectory points using the TAU code [2] developed by DLR. Accounting for chemical non-equilibrium, these numerical simulations provide the heat-flux distribution over the surface of the vehicle.

The perforation of the body has been assumed to occur at the location of maximum heat-flux. In a second step, the computational

COMPUTATIONAL ANALYSIS OF ATV RE-ENTRY FLOW AND EXPLOSION ASSESSMENT

R. Schmehl, Ph. Reynier, AOES – Leiden, The Netherlands

D. -E. Boutamine, J. Steelant & L. Marraffa ESA-ESTEC, Noordwijk, The Netherlands

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domain was extended to the interior of the vehicle, including the connecting gap in the vehicle body. The corresponding CFD simulations provide the internal temperature, pressure and species distributions accounting for dissociation and recombination reactions.

Figure 1: Sketch of the ATV.

In a second step, the potential of propellant ignition and explosion is assessed by means of qualitative and quantitative data from literature [3-5]. Correlations for minimum auto-ignition pressure and temperature as well as for ignition delay times are derived for various reaction sets used in conjunction with the CFD results and potential scenarios for propellant leakage in order to construct an analytical framework for the description of the explosion potential. Finally, for the set of trajectory points selected, the results obtained with the numerical simulations are analysed and compared to the characteristic parameters in order to assess the ignition risk along the trajectory.

Approach

Trajectory Analysis

The re-entry trajectory of the ATV has been analysed for several cases. Four options were investigated [1], for nominal and non-nominal ATV and for a controlled and non-controlled entry. The four cases are the following combinations:

- Non-nominal and uncontrolled;

- Non-nominal and controlled;

- Nominal and uncontrolled;

- Nominal and controlled.

The nominal re-entry is related to the ATV filled with ISS waste and propellant residuals, while the non-nominal case refers to an uncompleted docking; i.e. the down cargo (during the re-entry) is the up-cargo, which includes a significant amount of unused propellants. In the case of non-nominal vehicle, ATV carries about 2000 kg more mass than in the nominal one. As a consequence, the centre of mass and the moment of inertia are different which influences significantly the flight path during re-entry.

The non-controlled entry is without any correction of trajectory. For the controlled entry, the ATV is oriented to perform a ballistic re-entry and the trajectory is controlled by generating an angular velocity with a pitch rate of 10 degrees per seconds provoking a tumbling. The scenario retained for the ATV end of mission is the non-controlled option without tumbling. Therefore, in this study, the analysis is restricted for a nominal ATV performing a non-controlled entry.

Table 1

Free stream parameters used for the computations at the two selected trajectory points

Altitude (km) 75 80

Pressure (Pa) 2.35 1.05

Density (kg/m3) 3.9 10-5 1.85 10-5

Temperature (K) 208 198.7

Velocity (m/s) 7175 7202

ATV destruction has been investigated in [1] using engineering methods. According to

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this work, ATV looses its solar arrays at 92 km of altitude, the main fragmentation is sup-posed to occur at 75 km [6]. In an attempt to refine this prediction, the explosion risk of ATV is assessed in the current work for two trajectory points corresponding to the alti-tudes of 75 and 80 km. The main characteris-tics of these two trajectory points, computed at ESTEC using the trajectory input data available in [2,6], are reported in Table 1.

Numerical Solver

The numerical simulations of the flow-field for the ATV with and without a perfo-rated structure have been performed with the code TAU [2]. This tool solves the three-dimensional Navier-Stokes equations, written in a conservative form, using a finite volume approach. It can handle structured, unstruc-tured and hybrid meshes built with prisms, pyramids, tetrahedra and hexaedra. Dissipa-tion terms are also added to damp high fre-quency oscillations. For the current study, the AUSM-DV scheme, a second order upwind scheme, has been used. The integration in time is carried out through an explicit multi-step Runge-Kuttta multistep scheme (without preconditioning). A module for grid adapta-tion is available and has been used to check the grid independence of the results.

Explosion Analysis

An explosion of residual propellants is af-fected by a variety of parameters, such as the mode and rate of propellant release (valve leakage, pipe rupture or tank bursting), the flow environment (composition, temperature, pressure, velocity and presence of hot metallic surfaces), the ignition characteristics of fuel/oxidizer combinations (minimum pres-sure and temperature for auto-ignition, delay times) as well as the coupling effects (impact of blast wave or fragments on piping and tanks). Fundamental ignition properties of

hydrazine derivates, MMH and UDMH, both on board of ATV during its re-entry are listed in Table 2.

Table 2

Flammability data of liquid hydrazine in pure air at atmospheric pressure [4]

MMH UDMH

Flash point T [K] 294 258

Auto-ignition Temperature

T [K] 467 523

Minimum ignition pressure

P [Pa] 3.14 4

Range of flammability

X [%] 2.5 – 98 2 - 95

Figure 2: Minimum auto-ignition temperatures of liquid hydrazine in air/NO2-mixtures at atmospheric pressure [4].

A contamination of the air with strong oxidizers, either from leaking propellant or from dissociation of oxygen, drastically decreases the auto-ignition limits. A quantitative assessment of the decrease of the

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auto-ignition temperature due to NO2 contamination of the air is illustrated in Figure 2. NO2 is a major dissociation product of evaporated NTO (Nitrogen Tetra-Oxide, N2O4), which is an oxidizer on board of ATV.

Taking into account the various scenarios of propellant release into the flow during re-entry and the availability of validated experimental data, the following gas phase reactions are analysed in more details:

MMH + O2 (1)

MMH + N2O4/NO2 (2)

UDMH + N2O4/ NO2 (3)

Figure 3: Mesh used for the computations of the exte-rior flow field at 75 km altitude.

Numerical Simulations

Computational Conditions

Two points of the re-entry trajectory, cor-responding to the altitudes of 75 and 80 km, have been computed. The corresponding flow conditions in terms of velocity, pressure, den-sity and temperatures for the two points are reported in Table 1. No angle of attack has been accounted for; therefore axisymmetric computations have been performed. Since the ATV has a complicated shape, as shown in Figure 1, the simplified geometry depicted in

Figure 3 has been used for the numerical simu-lations. In a first step, computations have been performed for the nominal ATV without struc-tural perforation. The most probable location of a perforation has been derived from the peak in the calculated heat-flux distribution on the ve-hicle surface. Then, new computations have been performed accounting for structure perfo-ration. For all the calculations, an isothermal catalytic wall at 800 K has been applied as boundary condition. The chemical non-equilibrium is accounted for through a 5 spe-cies and 17 equations model [7,8].

Figure 4: Mesh used for the computations with perforated vehicle body at 75 km of altitude.

The meshes have been generated using CENTAUR [6], this software provides a com-plete grid generation package that can be used in all phases of computational engineering. The grid generator produces hybrid meshes built, for two-dimensional cases, of triangles and quadrilaterals. For both cases (intact and perfo-rated vehicle body) an initial mesh has been made and then adapted with the adaptation module of TAU by adding 10 % more points at each adaptation loop. This process has been stopped when the grid convergence was reached. For the cases of intact vehicle body, after seven adaptations the drag was quasi the same with a variation of about 0.05%. The mesh of the boundary layer has been adapted to have a value of y+ lower than 1. The final meshes for the geometry without intact and with perforated vehicle body are 40000 and

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42000 elements respectively. They are plotted in Figures 3 and 4. The computations have been performed with a CFL number of 1 and the convergence was reached in approxi-mately 50000 iterations. For the cases of per-forated spacecraft body, calculations were terminated as soon as a residual flow through the gap was reached.

Figure 5: Mach number distribution for an altitude of 75 km (no angle of attack).

Figure 6: Temperature distribution for an altitude of 75 km (no angle of attack).

External flow-field

Computations have been performed for the ATV with intact body at the altitudes of 75 and 80 km. In Figures 5 and 6, the Mach number and temperature distributions ob-tained for an altitude of 75 km around the

ATV are plotted. The shock standoff is located at 44 cm from the stagnation point. The tem-perature reaches a maximum value close to 15900 K in the shock layer. Figure 7 represents the surface pressure distribution along the ATV. The numbers between brackets in the figure correspond to the locations over the ATV indicated by the same numbers in the ATV sketch of the same figure. The maximum pressure is located at the stagnation point with a value of 1950 Pa. It decreases progressively along the surface and after the first edge the value is close to 1050 Pa. Downstream of the second edge the pressure is very low.

Figure 7: Surface pressure distribution over the ATV.

Figure 8: Heat-flux distribution over the ATV.

The heat flux distribution is shown in Fig-ure 8. As for the pressure, the numbers repre-sents the locations over the surface indicated in the sketch plotted in the same figure. The maximum of heat flux is 360 kW/m2. It is lo-cated at the first edge of the vehicle and not at the stagnation point; this is a consequence of the geometry of ATV, with a wall perpendicu-lar to the flow and a sharp edge. Since the

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maximum of heat-flux is located at this edge in the next computations the gap has been lo-cated there.

Figure 9: Pressure distribution (in Pa) inside and out-side ATV for an altitude of 75 km.

Figure 10: Distribution of partial pressure (in Pa) of atomic oxygen for an altitude of 75 km.

Coupled external and internal flow fields

The flow field around and inside the per-forated vehicle has been computed for the al-titudes of 75 and 80 km. As illustrated in Fig-ure 4, a 20 cm wide gap in the front wall has been considered. The total time to fill up ATV with air from the exterior is around 0.1 s. In a first approximation, the unsteady effects due to the variations of the exterior pressure and density fields can be neglected and the steady approach retained for the CFD is valid. The

converged solution shows only a residual inter-nal flow with a Mach number of 0.01 and an in-ternal pressure nearly equal to the pressure of the shock layer with a value of 1790 Pa. This is highlighted in Figure 9 representing the pres-sure distribution around and inside ATV.

Figure 11: Distribution of partial pressure (in Pa) of dia-tomic oxygen for an altitude of 75 km.

The temperature distribution inside the ATV is quasi uniform with an average value around 1300 K. The dissociation of molecular oxygen starts at a temperature level of 2500 K. As a consequence, with the selected approach, there is little atomic oxygen inside the ATV. The partial pressure of atomic oxygen corre-sponding to these conditions is 50 Pa. This quantity reaches a level of 630 Pa but at the stagnation point. This is highlighted in Figure 10 showing the distribution of atomic oxygen inside and outside ATV. In fact, the atomic oxygen coming inside the ATV recombines as shown in Figure 11 where the partial pressure of diatomic oxygen is displayed.

The same computations have been done for an altitude of 80 km. The flow pattern is similar to the one observed at 75 km. In Table 3, the main results obtained for both altitudes are summarized. Note that there is a factor of two between the pressures for a difference of 5 km in altitude. This is due to the dependency of the dynamic pressure on the atmosphere evolution in terms of pressure and density. These results

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have been exploited to assess the risk of ex-

plosion at these altitudes.

Explosion Analysis

Auto-ignition characteristics

Considering the reactions (1) - (3) of hy-drazine derivates MMH and UDMH, either with oxygen or with the oxidizer NTO, Figure 12 summarizes experimental auto-ignition data and correlations (adapted from [5] and [6]).

Figure 12: Minimum auto-ignition pressure of different

reaction sets. Experimental data from [5] and [6], dashed lines extrapolated behaviour.

Due to a lack of data for higher temperatures, the functional dependency of the minimum auto-ignition pressure has to be extrapolated. The analytical form of the correlations is as follows,

pign,min = T2 exp(5115.5/T-13.425) (4)

pign,min = 70.0 T2 exp(2100.0/T-15.8000) (5)

pign,min=133.3 T2 exp(1811.6/T-16.3136) (6)

where the pressure is in Pa and correlation (4) describes the reaction (1), correlation (5) the reaction (2) and correlation (6) the reaction (3). For the later two reactions the ignition delay is given by the following correlations [10],

τign p2/T4=3.04 10-20 exp(4550/T) (7)

τign p2/T4=1.95 10-20 exp(3260/T) (8)

Table 3

Summary of the main results for the two selected trajectory points; Max is for the maximum value

Altitude Z=80 km Z= 75 km

External Internal External Internal

Max Velocity [m/s] 7200 7 7175 7

Max Temperature [K] 15720 1100 15390 1300

Max pressure [Pa] 930 840 1930 1790

Max O partial pressure [Pa] 300 20 630 50

Max O2 partial pressure [Pa] 188 174 390 370

Max heat flux [kW/m2] 230 330 360 500

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where τign is in [s], p is in [atm] and T is in [K]. With respect to the reaction (1), the minimum auto-ignition pressure is not af-fected by the presence of nitrogen [4].

Coupling with CFD

Based on the computed pressure and temperature distributions in the exterior and interior flow-field, the correlations for the minimum auto-ignition pressure can be employed to determine whether a release of fuels MMH or UDMH alone or with additional release of oxidiser NTO will lead to a local ignition of the mixture. Since the pressure and temperature distributions inside ATV are nearly homogeneous, the analysis can be drastically simplified. Figure 13 shows how the various characteristic pressure values are evolving for decreasing altitude.

Figure 13: Pressure data from trajectory analysis, CFD simulation and auto-ignition limits.

Down to an altitude of about h=65 km, the maximum pressure resulting from the trajectory analysis is not affected by the ballistic coefficient. This pressure evolution somewhat overestimates the maximum external and internal pressure resulting from the CFD simulation performed at altitudes of 75 and 80 km. Making use of the computed maximum temperatures summarised in Table 3, the minimum auto-ignition pressure is

evaluated and included in Figure 12. It is obvious, that pign,,min is roughly one order of magnitude below the maximum pressure level at the two altitudes. Thus, a leak of MMH or UDMH in the interior of ATV will ignite in contact with molecular oxygen, which will potentially trigger an explosion of the vehicle. A simultaneous leak of NTO intensifies this effect as well as residual atomic oxygen.

Conclusions

From the analysis of the re-entry trajectory of ATV, two points have been selected in order to assess the risk of explosion at a higher altitude than expected. For these two points numerical simulations of the flow around the ATV have predicted the part of the structure, which will sustain the maximum heat-flux and heat-load during the re-entry. In order to assess the explosion risk numerical predictions of the flow with a perforated structure have been undertaken. The post processing of the results show a high risk of explosion if the propellants enter in contact with the internal flow.

This shows that a destruction of ATV at higher altitude than expected is possible. However, this study is only a first step to investigate with a refined analysis the ATV re-entry. To have a more accurate predictions phenomena like the fluid-structure aspect should have to be accounted for in order to predict the perforation of the external structure of the ATV but also of the propellant tanks embarked on ATV.

Acknowledgements

The authors wish to thank the ATV development teams working at ESTEC and CNES for their support.

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