OUT-OF-PLANE MANOEUVRE CAMPAIGNS FOR METOP-A: PLANNING, MODELLING, CALIBRATION AND RECONSTRUCTION. +49(0) ,

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1 OUT-OF-PLANE MANOEUVRE CAMPAIGNS FOR METOP-A: PLANNING, MODELLING, CALIBRATION AND RECONSTRUCTION Francisco Sancho (1), David Lázaro (2), Pier Luigi Righetti (3) (1) GMV at EUMETSAT, Eumetsat-Allee 1, D Darmstadt, Germany, +49(0) , (2) Space Operations Consulting at EUMETSAT, Eumetsat-Allee 1, D Darmstadt, Germany, +49(0) , (3) EUMETSAT, Eumetsat-Allee 1, D Darmstadt, Germany, +49(0) , ABSTRACT Metop-A is the first European operational satellite for meteorology flying in a Low Earth Orbit (LEO), and the first satellite operated by EUMETSAT in this type of orbit. It is the first of a series of three satellites flying on a Sun-synchronous orbit, at an approximate altitude of 826km and with an inclination of 98.7 degrees. The repeat cycle is 29 days and 412 orbits, and the dead-band of the ground-track is ±5km. The mission Local Time of the Ascending Node (LTAN) is 21h 30min and must be maintained within 2 minutes, which derives in the need to perform active orbital inclination control. The European Space Operations Centre (ESOC) performed the first out-of-plane manoeuvre in the course of the Launch and Early Operations Phase (LEOP), on October 21 st, The spacecraft local time started its close-to-parabolic shape drift at -60 seconds, reached the maximum (+75 seconds) in August 2007 and descended to -55 seconds in April 2008, when the first out-of-plane manoeuvre campaign performed by EUMETSAT took place. The location of the thrusters on the Metop platform, part of the SPOT family, imposes the need to rotate the spacecraft around its yaw axis before performing the out-of-plane manoeuvre itself; this rotation is performed by thrusters and has a non-zero net force component that has to be considered in the overall manoeuvre effect. Besides, instrument illumination conditions impose that the satellite can only be slewed within eclipse conditions; this effectively means that the maximum time available for the manoeuvre thrust is limited and depends on the time of the year, and that more than one manoeuvre may be needed to achieve the desired inclination change (two in the case of the April 2008 campaign). Further instrument and operational constraints led to the decision, in this first manoeuvre campaign, of performing the two burns with a very small separation of around 15 hours. This resulted in a high time pressure on the Flight Dynamics operations: no reliable a-priori rotation calibration value was available for the planning of the first manoeuvre, so the impact of its mis-performance on the on-board Attitude and Orbit Control System (AOCS) telecommands and on the planning of the second manoeuvre had to be corrected in a very limited time. Analyses carried out since then have resulted in an enhanced manoeuvre modelling and calibration process: while in current operational set-up a single manoeuvre, merging the net effect of the slew and anti-slew into the main propulsion, was considered, an improved modelling of the manoeuvre as five segments (slew start and stop, main propulsion and anti-slew start and stop) has shown much better performances, permitting to obtain more reliable calibration factors.

2 1. METOP-A, MISSION AND ORBIT CONSTRAINTS Metop-A is the first European operational satellite for meteorology flying in a Low Earth Sun- Synchronous Orbit, and the first satellite operated by EUMETSAT in this type of orbit. It is the first of a series of three satellites which should ensure 15 years of continuity of mission operations. Its repeat cycle is 29 days and 412 orbits, corresponding to an approximate altitude of 826km. A deadband of ±5km around the reference ground-track and frozen eccentricity conditions must be maintained. In plane manoeuvres are thus needed for dead-band maintenance and eccentricity control. The mission Local Time of the Ascending Node (LTAN) is 21h 30min and must be maintained within 2 minutes. This constraint is imposed to keep sun-calibration geometry for the Global Ozone Monitoring Experiment (GOME) instrument and derives in the need to perform active orbital inclination control around the nominal Sun-Synchronous inclination of 98.7 degrees. The satellite Service Module (SVM) is directly derived from the one used for SPOT, ERS and ENVISAT (Fig. 1). Automatic attitude control is based on one digital sun sensor, one digital earth sensor and two bi-axial gyros. Orbit control is performed by hydrazine thrusters (two plates, each with one pair for propulsion and yaw control and two pairs for roll and pitch control). The orientation of the thrusters on the Metop platform imposes the need to rotate the spacecraft around its yaw axis (slew) before performing the out-of-plane manoeuvre itself, to align the thrust with the cross-track direction. A second rotation of opposite sign (anti-slew) is needed after the manoeuvre to return to the nominal pointing. Due to the large mounting offset in yaw of the propulsion plates with respect to the flight direction (around 10 and 20 degrees for the plates in the velocity and anti-velocity direction, respectively), the performed rotation is higher than 90 degrees. Fig. 1. Metop satellite Instrument illumination conditions impose a constraint to the manoeuvre: the satellite can only be slewed within eclipse conditions. This effectively means that the maximum time available for the manoeuvre thrust is between 7 and 10.5 minutes, depending on the time of the year, as explained in [1]. The Payload Module (PLM) of the satellite carries 13 instruments for meteorological observation. The ones of particular interest for the Flight Dynamics operations are: Global Navigation Satellite System (GNSS) Receiver for Atmospheric Sounding (GRAS): provision of precise navigation data; Advanced Scatterometer (ASCAT): constraint in repeat cycle and frozen eccentricity maintenance; GOME: constraint in LTAN maintenance.

3 2. INCLINATION EVOLUTION. OUT-OF-PLANE MANOEUVRES The secular change of the orbital inclination is mainly due to the Sun gravity field and depends on the Mean Local Solar Time of the orbit. For the Metop orbit this perturbation is close to the maximum and causes an inclination decrease of around 50mdeg per year. Due to the quite large dead-band available, it would be theoretically sufficient to perform one inclination correction manoeuvre of approximately 75mdeg every 18 months to maintain the reference conditions within limits. However, due to the operational and practical reasons discussed in [1], the long-term inclination control strategy foresees the execution of one out-ofplane manoeuvre campaign (each consisting of one or two burns) every year, preferably around the autumn longer-eclipse season. With this strategy the long-term evolution of the deviations of the LTAN and of the inclination with respect to their reference values (21:30 and deg, respectively) are as shown in Fig Inclination Deviation (mdeg) LTAN Deviation (sec) Year Fig. 2. Foreseen evolution of inclination and local time with planned manoeuvres (first 10 years of mission) In April 2008 an inclination correction of around 70mdeg, corresponding with a delta-v of the order of 10m/s, was required. Considering the quite large mass of the satellite (more than 4 tons) and the very limited thrust available (around 36 Newton) around 18 minutes of continuous propulsion would have been needed. Due to instrument illumination constraints the entire manoeuvre phase must be carried out within eclipse. Subtracting from the eclipse time (around 30 minutes) the time needed for slew and antislew (each of them of the order of 11 minutes), it became evident that the required inclination correction could not be provided by a single burn. Execution of two burns was then selected (see [1]). In order to maximise the manoeuvre efficiency, the anti-velocity plate was selected: its pitch control thrusters being fully aligned with the antivelocity direction, this plate provides better performances. A negative rotation around the satellite s Z axis is therefore applied before the manoeuvre. Fig. 3 presents an overview of an out-of-plane manoeuvre, including the rotations needed to align the main thrusters with the cross-track direction. Equator Post-DV Trajectory Anti-Slew Mounting +Ysat offset DV Out-Of-Plane Slew +Xsat +Ysat Thrust Inclination change +Xsat Pre-DV Trajectory End of Thrust Start of Thrust Earth Shadow Nominal Inclination Fig. 3. Representation of an out-of-plane manoeuvre as seen from space (angles not to scale)

4 3. OPERATIONAL MANOEUVRE MODEL EVOLUTION The initial model for the Metop-A out-of-plane manoeuvres, implemented during the Launch and Early Operations Phase (LEOP) for the first inclination-control manoeuvre of the spacecraft, performed by the European Space Operations Centre (ESOC), only took into account the effect of the commanded rotation and main burn pulses on the orbital evolution. The manoeuvre calibration showed, however, a rather large along-track component of -12.5cm/s, resulting in an angular calibration of 4.6 degrees. It was decided not to use this manoeuvre s calibration for the following manoeuvres. A-posteriori analysis of the manoeuvre showed that, contrary to what was initially expected, the net effect of the rotations start and stop phases had a significant impact on the outcome of the manoeuvre Platform slew and anti-slew characterisation The complete manoeuvre process, including the rotation of the platform to the desired attitude for the manoeuvre and back to geocentric attitude mode, is performed by the platform in the so-called Orbit Control Mode (OCM). During this mode, the attitude control of the spacecraft is performed with thrusters, as opposed to the wheels and magneto-torquers used in the nominal platform mode. In order to perform the desired rotation of the platform around its yaw axis (Z) before and after the manoeuvre, the couple of thrusters providing a torque in the needed direction is activated until the nominal rotation speed of 0.3deg/s is reached. When the rotation is about to be completed, the couple of thrusters providing a torque in the opposite direction is activated in order to stop this rotation. The activation of these thrusters can be seen in Fig. 4, based on the spacecraft telemetry taken during the LEOP out-of-plane manoeuvre, where also the main thrust and the activation of the attitude control thrusters are shown. 9 Z+/Y+ Z+/Y- Z-/Y+ Z-/Y- Y+ Y- X+ X- 8 t 1 t 2 t 3 t 4 t 5 t 6 t 7 t 8 t 9 t pulses/sample 5 4 slew start slew stop main burn anti-slew start anti-slew stop :46 18:48 18:50 18:52 18:54 18:56 18:58 19:00 19:02 19:04 19:06 19:08 19:10 Fig. 4. Thruster activations during the LEOP out-of-plane manoeuvre Fig. 5 shows more clearly the activation of the thrusters during the slew. The couple Z-/Y-, Z-/Y+ are activated for starting the commanded negative yaw rotation, while Z+/Y- and Z+/Y+ are activated to stop this rotation. Similarly, the same couples of thrusters are activated for starting and stopping the anti-slew, this time in the inverse order, being this rotation a positive one in the yaw axis.

5 Z-/Y :46 18:48 18:50 18:52 18: :46 18:48 18:50 18:52 18:54 Z-/Y- Z+/Y- Z+/Y :46 18:48 18:50 18:52 18:54 18:46 18:48 18:50 18:52 18:54 Fig. 5. Accumulated thruster pulses during slew start (above) and stop (below) From the mounting matrix of the thrusters involved in the rotations of the platform before and after the manoeuvre (see Table 1) it can be easily seen that all these vectors have a non-negligible component in the satellite s X axis, and that all these components have the same sign. Because of the spacecraft s attitude at each point, this direction corresponds to the cross-track direction in the slew start and the anti-slew stop, and to the along-track direction in the slew stop and the anti-slew start. Table 1. Thrust alignment vectors for the slew and anti-slew thrusters X Y Z Z-/Y Z-/Y Z+/Y Z+/Y Slew and anti-slew effect inclusion in manoeuvre computation In view of previous values, the spacecraft manufacturer was requested to characterise the net effect of both rotations. With the results reported in [2], summarised in Table 2 for the thrusters used in this case, the model was modified to add four new parameters to it: the net along- and cross-track delta-v caused by each the positive and negative rotation. The generation of the manoeuvre telecommand (TCH) then takes these values into account to subtract them from the commanded manoeuvre: the cross-track component is adjusted with the commanded number of pulses, and the along-track one with the commanded slew rotation. Table 2 Net effect of slew and anti-slew for a manoeuvre using the anti-velocity plate Component Delta-V (m/s) Along-track Cross-track

6 A further improvement to the model was introduced to optimise the along-track component of the manoeuvre in order to achieve a given target on the satellite s ground-track, and use this optimisation in the computation of the commanded yaw rotation used for the execution of the out-of-plane manoeuvre. These enhancements to the model, both performed before the next Metop-A out-of-plane manoeuvre in April 2008, allowed EUMETSAT to improve the predictability of the platform by: removing the cross-track effect of the rotations around the yaw axis from the commanded number of pulses, and adjusting the commanded rotation for taking into account not only the thruster mounting and the angle calibration from previous manoeuvres, as before, but also the along-track effect of the rotations and a possible in-plane target. 4. EUMETSAT MANOEUVRE CAMPAIGNS 4.1. First manoeuvre campaign In April 2008, as seen in Fig. 2, the deviation of the LTAN with respect to its reference value reached -60 seconds. The detailed deviation of the ground-track, altitude, inclination and LTAN from their nominal values around those days is shown in Fig. 6. A manoeuvre campaign was therefore planned for the 8 th and 9 th of April 2008 in order to raise the inclination and revert the LTAN drift. Following the reasoning and strategy defined in [1], it was decided to execute two burns (referred to hereafter as B1 and B2) occupying the maximum possible duration within the eclipse. The manoeuvre would be followed one or two weeks later by an in-plane touch-up manoeuvre to correct the residual drift of the ground-track and return to the frozen eccentricity conditions. Fig. 6. Deviations of ground-track (left) and altitude, inclination and LTAN (right) with respect to the reference values around the date of the first EUMETSAT out-of-plane manoeuvre campaign The target of the manoeuvre campaign, shown in Fig. 7, was then to raise the inclination as much as possible within the given constraints (around 70 millidegrees in this case), reverting the LTAN drift. At the same time, the rotation of the platform would be optimised to apply an along-track component to place Metop on the reference ground-track on the day initially foreseen for the in-plane touch-up manoeuvre, April 17 th. The orbit altitude would be increased by around 90 meters.

7 Fig. 7. Predicted evolution of the deviations of gound-track (left) and altitude, inclination and LTAN (right) with respect to the reference values across the out-of-plane manoeuvre Prior to the preparation of this manoeuvre, the Flight Dynamics software had been modified to implement the improvements to the model described in Section 3. The plate selected for performing the burns was the anti-velocity one, which, as explained above, is more efficient than the opposite one. The accumulated modulus calibration obtained from the in-plane manoeuvres performed with this plate (3.8% over-performance) was used for planning these burns. However, due to the problems found in the calibration of the LEOP out-of-plane manoeuvre and explained in Section 3, no rotation calibration was used. As a result, a considerable variability in the manoeuvre size and, above all, in its direction (due to the slew performance) was expected for the first burn. The error in the along-track component was therefore of major importance. It was assumed that it would be possible to apply to B2 calibration values obtained from B1, thus making B2 much more predictable. Simulations confirmed the validity of this assumption. A trade-off between the user needs (mission outage as short as possible) and the operational constraints coming from the various subsystems (Flight Dynamics, spacecraft including payload, Mission Planning and Ground Stations support) led to the decision of performing the two burns with a separation of only 15 hours (9 orbits). This meant that there was a very short time to calibrate the first burn and re-plan the second one before uplink of the telecommands associated to it. Operational procedures were extensively modified to cope with these constraints, and 24-hour Flight Dynamics continuous support to the operations was needed. Sensibility analyses performed on the performance of the along-track component of the manoeuvre during simulations showed that an error in the slew rotation manoeuvre of less than 4 degrees (as observed during LEOP) was equivalent to more than 30 cm/s in the along-track component. This would produce a rapid drift in the post-manoeuvre orbit, which meant that updated products for the different subsystems would have to be generated by the Flight Dynamics Facility (FDF). In particular, the most critical update needed would be the orbit-position (Position Sur l Orbit, PSO) telecommand. The calibration of the first burn should largely improve the predicted performance of the second burn in order to achieve the targets. There was a very high probability to remain into the dead-band at least during the following weeks after the manoeuvre, in time to perform the fine touch-up considering 1% errors in the execution of the second manoeuvre and in the post manoeuvre propagation.

8 The calibration of the manoeuvre was performed using S-band radiometric observations (two-way range and one-way range-rate) from EUMETSAT s Svalbard station. The used observations arc was set to two days, and the three components of one single manoeuvre (merging the main burn and the net effect of the yaw rotations) were estimated. The weight sigmas assigned to the measurements were 5m for range and 15mm/s for range-rate. Although the results of the calibration of the first burn were more reliable than the ones from the LEOP manoeuvre, the process was still rather unstable (see Section 4.3), and it took more passes than expected, plus support information from off-line Precise Orbit Determination (POD) based on GRAS GPS data (see [3] for further details), to converge to a solution with an acceptable confidence level. As it can be seen below, the applied manoeuvre model left still room for improvement, especially taking into account that the modulus and rotation calibration of the first burn had to be applied to the second one, and that the planning of this second burn had to be based on the orbit determined after the first one. The calibration factors used for the second burn were 4.4% over-performance and degrees Second manoeuvre campaign In order to apply the long-term inclination-control strategy described in [1], a single-burn manoeuvre (B3) was executed on October 23 rd, The manoeuvre model and configuration of the calibration process were the same as during the first campaign, and it was confirmed that the modelling of the manoeuvre was too simple and resulted in a slow convergence of the calibration process. This time the issue was not as critical as during the first campaign, when the calibration of the first burn and the orbit thereafter were needed for the second burn Analysis of the calibration process The evolution of the calibration during the three burns performed by EUMETSAT was very similar. The one corresponding to B3 is analysed here in detail. Table 3 shows a summary of the residuals statistics for the orbit determination and manoeuvre calibration, including 21 passes before and 7 passes after the manoeuvre (executed at 14:30). It can be clearly seen that the quality of the orbit determination decreases significantly after the manoeuvre, where a high number of measurements are rejected and the residuals are much higher than the ones obtained in the pre-manoeuvre passes. Table 3. Statistics of pre- and post-manoeuvre operational observation residuals Measurement Number of Number of measurements Type of passes RMS type passes Total Used Rejected Before manoeuvre Two-way After manoeuvre range (m) Total Before manoeuvre After manoeuvre One-way range-rate (mm/s) Total Besides, the evolution of the calibration factors throughout the calibration process, as the number of post-manoeuvre passes increased, was not stable, as can be seen from Fig. 8 and Table 4, from which the instability of the solution is evident.

9 degrees Table 4. Variation of the operational calibration factors with respect to their final values Pass Modulus Angle % % % 32.12% % 6.95% % 7.59% modulus calibration rotation calibration % 0.00% :00 16:00 18:00 20:00 22:00 0:00 2: Fig. 8. Evolution of the operational calibration factors Further to this, the orbit obtained during the calibration process seven passes after the manoeuvre has been compared with the one obtained in the off-line POD. This comparison (Fig. 9) shows differences of up to 200m in the along-track direction. The configuration and manoeuvre model for the POD is the best possible, making use of all improvements to the model described in Section 5. Also the comparison of the estimated manoeuvre size with the one reconstructed from pulse counts in telemetry shows a rather poor match (2.5% error). Fig. 9. Orbit estimated operationally with S-band data vs. off-line POD 4.4. Summary of EUMETSAT manoeuvre campaigns Table 5 shows a summary of the out-of-plane manoeuvres performed up to now, where the commanded values are derived from the target ones taking into account the net effect of slew and anti-slew, as explained above. In each case, the target rotation has been chosen in order to generate the desired inplane component for keeping the ground-track constraint. It is to be noted that the correction to be applied to this rotation varies with the manoeuvre size, which is the consequence of the slew and antislew effects being independent of this size. Table 5. Summary of EUMETSAT out-of-plane manoeuvres Date Target size Target rotation Commanded size Commanded rotation (m/s) (deg) (m/s) (deg) B1 08/04/ B2 09/04/ B3 23/10/ B4 * 17/09/ * The values given for B4, in the future at the time of writing this paper, are the foreseen ones

10 In Table 6, the calibration factors used for planning the different manoeuvres, plus the ones obtained from the calibration process described above, are summarised. As already explained, the calibration for the first manoeuvre did not include the values obtained from the LEOP out-of-plane manoeuvre, and this is why the applied rotation calibration is zero in this case. Besides, it can be seen that the values of the modulus calibration diverge, which is not very realistic. The angular calibration is more stable, though not completely satisfying. Table 6. Operational calibration factor history Applied calibration Obtained calibration Modulus Angle (deg) Modulus Angle (deg) B B B B * * TBC TBC * Efficiency that would have to be applied following current manoeuvre model The applied calibration is computed as the product of the obtained calibration values in the case of the modulus, and the opposite of the sum of the obtained calibration values in the case of the rotation. Only values obtained in manoeuvres considered as valid for calibration (larger than 20cm/s, except the LEOP out-of-plane manoeuvre) are used. Following tables give an overview of the three burns performed by EUMETSAT up to the time of writing this paper (B1, B2 and B3). Table 7 shows the delta-v reconstructed from the manoeuvre telecommand, using current manoeuvre model as implemented in the Flight Operations Manual (FOM) supplied by the spacecraft s manufacturer (with no calibration factors applied): this includes the delta- V due to slew and anti-slew rotations and disregards the attitude control thrusters. Table 7. Manoeuvre resulting from TCH reconstruction (FOM) Manoeuvre components Radial (m/s) Along (m/s) Cross (m/s) Angle (deg) B B B Taking into account these values and the estimation of the manoeuvre components obtained from the operational orbit determination process, the calibration factors resulting from each manoeuvre can be obtained independently from one another. From these results, summarised in Table 8, it can be seen that this manoeuvre model has indeed improved the angular calibration, but the calibration in modulus is still not stable enough. Table 8. Operationally estimated manoeuvre components and resulting calibration factors Estimation Calibration factors Radial (m/s) Along (m/s) Cross (m/s) Angle (deg) Modulus Angle (deg) B B B

11 5. ENHANCEMENTS TO THE MANOEUVRE MODEL, PLAN AND CALIBRATION 5.1. The new manoeuvre model: description and determination of its parameters The modification described in Section 3 in order to take into account the net effect of the slew and antislew by subtracting it from the target delta-v is equivalent to moving the delta-v due to the rotations to the time interval of the main burn. Due to the fact that the slew and anti-slew start and stop are separated in time (see Fig. 4) and spacecraft angular position in the orbit from the main burn, this model is not fully representative of the satellite s dynamics. Since an improvement to the manoeuvre model after the problems described above was needed, the next step was the better modelling of these rotations so that their effect was considered at the correct times. In view of the plots of the thrusters activations during the manoeuvre campaigns, like the one shown in Fig. 4, it seems reasonable to add one manoeuvre for each the slew start, slew stop, anti-slew start and anti-slew stop. Taking the main burn into account, this corresponds to a 5-segment manoeuvre. The calibration of the main burn represents then the pure effect of the manoeuvre telecommand. One of the problems that arise now is the definition of the times and accelerations for the auxiliary four manoeuvres. These can be extracted from the spacecraft telemetry by integration of the effect of the pulses executed during the manoeuvre once it has taken place, but this approach is not valid for the phase of the manoeuvre planning, where the splitting of the manoeuvre into 5 segments would also be of interest to improve the prediction of the post-manoeuvre orbital status. Since the torque needed for starting and stopping the rotations is independent of the manoeuvre size and the commanded rotation, so is the amount of thrusting. The net effect of the slew and anti-slew will, however, depend on the commanded rotation (since the projection of the satellite axes on the orbital frame depends on the spacecraft attitude) and the moment within the spacecraft s lifetime: the different pressure in the tanks causes the thrust per pulse to decrease with time, which means that longer intervals are needed for reaching the necessary torque. Being the commanded rotations for each manoeuvre not too different from one another, it seems reasonable to use, for a given manoeuvre, a-priori values for the times and accelerations extracted from the previous ones. It will have to be analysed during future manoeuvre campaigns, depending on the observed evolution of these parameters, whether the values from all previous manoeuvres have to be used, or just the ones from the most recent manoeuvres, to compensate for the loss of thruster level during Metop s lifetime. Once the main burn start time and duration are known, the start and stop times of the auxiliary manoeuvres can be easily determined from Table 9, where the different time intervals have been fixed in one of the three following ways: computed from the telemetry obtained during previous B1 and B2 manoeuvres (start and end of activations of Z attitude control thrusters): shown as Telemetry in the table; extracted from the FOM: shown as FOM in the table; estimated by the spacecraft s manufacturer based on the characteristics of the rotation: shown as Estimated in the table, and observed to be well in line with the telemetry obtained from the previous manoeuvres.

12 Table 9. Time intervals associated with the slew and anti-slew (nomenclature as in Fig. 4) Time Source of Description interval information Duration (s) t 2 t 1 Slew start duration Telemetry 145 t 4 t 3 Slew stop duration Telemetry 135 t 8 t 7 Anti-slew start duration Telemetry 160 t 10 t 9 Anti-slew stop duration Telemetry 140 t 5 t 1 Time between slew start and main burn start FOM 675 t 7 t 6 Time between main burn end and anti-slew start FOM 70 t 3 t 1 Time between slew start and slew stop Estimated 360 t 9 t 7 Time between anti-slew start and anti-slew stop Estimated 360 The associated delta-v values can be determined with the off-line POD based on GRAS GPS data, as there is good observability of each of the components for all segments of the manoeuvre. The results obtained as average values from B1 and B2 are given in Table 10 and are quite in agreement with the a- priori values presented in Table 2. Table 10. A-priori delta-v values for the rotations (nomenclature as in Fig. 4) Time Delta-V (m/s) Description interval Radial Along-track Cross-track t 2 t 1 Slew start t 4 t 3 Slew stop t 8 t 7 Anti-slew start t 10 t 9 Anti-slew stop N/A Total Apart from the correct modelling of each thrust within its actual time (or at least much closer to it), it can be seen that this enhanced model also adds a-priori information on the radial component of the slew and anti-slew, which was missing up to now Validation of the model In order to validate this new approach, the calibration of the B3 manoeuvre was rerun using the improved model. The set-up for the calibration process (type of measurements, observations arc, weighing assigned to each measurement type ) was left unmodified with respect to the manoeuvre calibration performed operationally. The summary of the residuals statistics before and after the manoeuvre is shown in Table 11. A significant improvement of the convergence for the postmanoeuvre passes can be seen with respect to the operational calibration, summarised in Table 3. Table 11. Statistics of pre- and post-manoeuvre observation residuals with the 5-segment model Measurement Number of Number of measurements Type of passes RMS type passes Total Used Rejected Before manoeuvre Two-way After manoeuvre range (m) Total Before manoeuvre After manoeuvre One-way range-rate (mm/s) Total The evolution of the calibration factors with the passes after the manoeuvre is shown for a 5-segment manoeuvre in Fig. 10 and Table 12. Comparison with Fig. 8 and Table 4 shows much more stability in

13 the evolution of these factors from the very first pass. This better behaviour will be of much value for future two-burn manoeuvres degrees Table 12. Variation of the calibration factors with respect to their final values (5-segment model) Pass Modulus Angle % % % -7.49% % -2.00% % -3.00% modulus calibration rotation calibration % 0.00% :00 16:00 18:00 20:00 22:00 0:00 2: Fig. 10. Evolution of the calibration factors (5-segment model) The comparison of the estimated orbit with the same off-line POD solution described in Section 4, seen in Fig. 11, also shows a big improvement with respect to the operational one (Fig. 9), with the 200-meter bias in the along-track direction disappearing. The estimated manoeuvre size now matches much better (error of the order of 0.3%) the one reconstructed from pulse counts in telemetry. Operationally, the fast convergence observed will permit to identify very early any platform misbehaviours during the manoeuvre. Fig. 11. Orbit estimated with S-band data vs. offline POD (5-segment model) 5.3. Recalibration of previous EUMETSAT manoeuvres With this new model, the reconstruction of the manoeuvre from the telecommand (Table 13) does not include the delta-v caused by the slew and anti-slew rotations any more. Table 13. Manoeuvre resulting from TCH reconstruction (5-segment model) Manoeuvre components Radial (m/s) Along (m/s) Cross (m/s) Angle (deg) B B B As in Section 4.4, comparison of the estimated manoeuvres with the reconstructed manoeuvres results in the calibration factors. The values obtained now, shown in Table 14, are much more stable.

14 Table 14. Estimated manoeuvre components and resulting calibration factors (5-segment model) Estimation Calibration factors Radial (m/s) Along (m/s) Cross (m/s) Angle (deg) Modulus Angle (deg) B B B Besides, reconstruction of the manoeuvres from the TCH using the full dynamics (i.e. including the expected attitude control thrusters, Table 15) and comparison of the estimated values against it gives extremely stable and very small values for the modulus and rotation calibrations (Table 16). This means that the calibration with the 5-segment model is properly capturing the impact of the fulldynamic behaviour on the change of size and direction caused by the foreseen activation of the attitude control thrusters during the main burn (activation confirmed with very high accuracy from telemetry). Table 15. Manoeuvre resulting from TCH reconstruction (full dynamics) Manoeuvre components Radial (m/s) Along (m/s) Cross (m/s) Angle (deg) B B B Table 16. Calibration factors obtained against the full-dynamics reconstruction Calibration factors Modulus Angle (deg) B B B It is interesting to note that the important mismatch in the radial component observed between the results presented in Table 13 and Table 14 (around 60% of the predicted thrust observed) disappears almost completely when the activation of the attitude control thrusters is considered. A third calibration scale factor is then needed to compensate for that error. 6. NEXT MANOEUVRE CAMPAIGN AND FUTURE ENHANCEMENTS The operational implementation of the enhancements mentioned above is not straight forward, since the calibration of a manoeuvre modelled as 5 segments instead of just 1 affects the interaction between several programs within the FDF (like the manoeuvre optimisation, the manoeuvre command generation or the manoeuvre calibration) and between the FDF and other subsystems like the Mission Planning Facility. All those modifications are already available on the testing platform. Next campaign will be a single-burn manoeuvre, planned to be executed on September 17 th, The modulus and rotation angle calibration factors to be applied are the ones resulting from the recalibration of the manoeuvres executed by EUMETSAT using the 5-segment model, which have been proved to be more accurate and realistic; a third calibration factor, to account for radial efficiency, is moreover used. On the testing platform the manoeuvre calibration will be performed according to the 5-segment model; the effect of the slew and anti-slew will be set to the average of the values observed in the previous manoeuvres. After analysis, from the thruster pulse telemetry, of the time interval between end of anti-slew and end of eclipse during B1, B2 and B3, and following acceptance by the Metop Spacecraft Team, it has been decided that a larger occupation of the eclipse in OCM of 60 seconds will be used in next campaign. This results in an increase in the delta-v of around 10%.

15 Besides, as part of the enhancements to the FDF software and configuration described in [4], GRAS observables will be used operationally in the manoeuvre calibration and orbit determination process. Recalibration of previous manoeuvres using such data results in orbital solutions much closer to the off-line POD, as shown by Fig. 12. For future manoeuvres, it is foreseen to have the 5-segment model fully implemented in the operational FDF software, taking the times and delta-v values of the slew and anti-slew contributions from what will be automatically computed based on previous manoeuvres telemetry. Fig. 12 Orbit obtained using S-band and GRAS observables vs. off-line POD To achieve this, the software tool used to reconstruct the provided delta-v from the thruster pulse telemetry will also have to be modified in order to: use the real spacecraft attitude at each integration step, instead of the current fix attitude; compute the rotations start and stop times from the activation of the Z attitude control thrusters. 7. CONCLUSIONS A new model has been developed by EUMETSAT for the calibration of Metop-A out-of-plane manoeuvres. Instead of the classical approach to estimate the total delta-v of the main burn and of the slew and anti-slew rotations (valid if the latter have a negligible impact), 5 propulsive segments (with fixed delta-v for the rotations) are considered. Only the main propulsive burn is then estimated. This approach leads to a much better convergence of the calibration factors already after 2 passes and to an estimated orbit that is less than 20 meters apart from the off-line POD solution. The estimated calibration factors are moreover much more stable and correlate much better with the impact of the satellite attitude control thrusters on the final delta-v. The operational benefit is evident above all in the case of double out-of-plane manoeuvres, where the early calibration of the first burn is used for the generation of the second burn telecommand. 8. REFERENCES [1] Damiano A., Righetti P.L. and Soerensen A., Operational Local Time and Eccentricity Management for Metop-A, 21 st ISSFD, Toulouse, France, [2] Batillot E., Metop Out-of-Plane OCM simulation plan, SOE.MO.TCN ASTR Issue 1 Rev. 0, EADS Astrium, Toulouse, [3] Andrés Y., Sancho F. and Marquardt C., The EUMETSAT Environment for Precise Orbit Determination, AIAA SpaceOps, Heidelberg, Germany, [4] Lázaro D., Righetti P.L., García P. and Sancho F., Automation and Enhancement of Metop-A Flight Dynamics Operations, AIAA SpaceOps, Heidelberg, Germany, BACK TO SESSION DETAILED CONTENTS BACK TO HIGHER LEVEL CONTENTS