ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
5 The resulting displacement of  the rotor versus the structure is then α( p1-p2). This configuration allows the  use of  larger threads, which are needed to withstand the loads , while allowing very small steps per turn  otherwise not possible.  

Magnetic clutch   A key aspect of the BSMA is then its electro magnetic clutch. This clutch was  designed based on the  following constraints:   • transmit torque during stepping sequence,   • complete loss of contact during release,   • closed when unpowered,   • perform more than 1 million operation cycles  without detrimental wear   • coil temperature within acceptable range,   • mass  allowing RSPA module to  support clutch without assistance of any launch- lock mechanism.  

The clutch design is based on an electromagnet principle.  Magnetic technology was  selected as  it allows  larger displacement compared to piezoelectric ceramics. This is needed to generate a gap which is large  enough to accommodate for manufacturing and assembly tolerances.   

Figure 6. Partial C ut View Showing Rotor, Clutch and RSPA M odule  

In practice, the clutch consists of  6 pallets driven by the central coil and magnet. Each pallet can rotate  around a flexible blade when the magnetic coil is powered.  At the end of the pallet is a friction tooth that will  interface with the rotor and transmit torque from the RSPA module. A permanent magnet generates some  flux which closes the pallets (and therefore the clutch) when the electromagnet is not powered.  

Performance of the clutch was optimized using magnetic simulation ( Figure 7) in order to minimize its size  and heating while providing enough force to ensure torque transmission.  

As the rotor turns, the distance between the tooth and rotor  contact surface can vary significantly depending  on the rotor runout  and other manufacturing tolerances. A major point of concern during the design was to  ensure that magnetic forces would not vary excessively, remaining  high enough to ensure proper torqu e  transmission.
6             Figure 7. Clutch Magnetic Simulation         Figure 8. Flexible Pivot Stress Simulation  

Specific attention was also needed for the design of the flexible blades in order to ensure a high  transmission stiffness while preserving a low flexure stress and flexure stiffness. High strength steel was  used to ensure a >99% reliability of the clamp including  the fatigue effect over lifetime.  

Prestressed Piezoelectric Actuator ( PPA)   A double PPA is placed between the lever arm and the output shaft. This PPA is composed off 2 stacked  piezoelectric components to provide redundancy.  The supplied piezo components have been validated for  space application with a LAT (Lot Acceptance Test). This LAT sequence, composed of several test group   samples, was established thanks to previous work with agencies (ESA and CNES).  

Differential Screw Wear Test Bench  

Guided lever arm   A lever arm connect s the two F SPAs and the output shaft. The resulting position of the shaft is the mean  position of the two FSPAs. Redundancy is then achieved as each FSPA can move the output shaft  independently from the other one.      

Figure 9. Schematic of BSMA for Max Position, Lever Tilt Angle is Exaggerated  

However, a major consequence of this architecture is that the lever arm tip displacement s are not straight.  The rotation of the arm induces a side displacement which must  be supported by the threaded interfaces.  

This design was selected due to  its compactness  and considering that :  • Static load is rather small (<100N) compared to allowable screw tension in static conditions ,  • Speed is very low (allowing contact heat to dissipate) ,  • Limited lifetime requirement (100 operations) .
7 Screws excessive wear   As expected, the alignment of the clutch with the rotor was  a major difficulty during  assembly. Several  assembly methods had to be tested on the first engineering models to identify the best options. During one  of the tests  and due to a mishandling, one of the screw s was damaged and had to be replaced. After  disassembly, it was found that the other screws showed excessive deformation and wear (even considering  the specific history of the model).  

Figure 10. EM1 Screws, Left Screw Shows Excessive Wear,  Right Screw Was Damaged Due t o Mishandling  

Following this result, it was decided to upgrade the design of the BSMA.   The major design improvements were to:   • replace initial soft screw material with a high strength stainless steel with space heritage,   • use both liquid and solid space qualified lubricants to minimize the risk of dry contact.  

Test bench verification   In order to validate the updated design, it was decided to build a dedicated test bench.   Two major elements were to be considered for the test bench design:   • Kinematic Accuracy: movement shall be as close as possible to the effective movement in the  BSMA,   • Load Accuracy: preload shall be the same as in the BSMA,   • Accelerated Test: rotation speed must be accelerated to get an acceptable test duration.  

A geared electric motor was used to drive the differential screw rotor. Transmission to the rotor was done  using gear wheels. Gears  would not be acceptable for normal use as their resolution is much t oo large for  the application, but this was  acceptable for the accelerated test, where each cycle was a complete turn of  the rotor.  

Figure 11. Differential Screw Test Bench
8   Figure 12. Position and T orque Measurement (400 C ycles Achieved)  

The results using the test bench validate the applied changes : Although acquisition noise reduced  measurement quality , no significant torque trend was observed during the test (see  Figure 12). Comparison  of initial and final friction coefficient showed no sign ificant trend either . Finally, an inspection of the tested  parts was performed by CNES which concluded that thread wear was  acceptable.  

Nanometric Step Size and Stability Test Bench  

Test Bench Architecture   A test bench was  designed to analyz e the output position of the motor over its full stroke, in vacuum  conditions  and under static load. A second key feature of the test bench is the ability to measure the position  stability against time of the motor at the nanometric scale over a period of more than 100 hours.  This test  bench exploits previous successful experience on long- term nanometric stability measured on PPA  piezo electric  actuators [ 5].     For this purpose, a 3- channel interferometric displacement sensor ( Figure 13a & b) was  used in vacuum to  achieve a sub- nanometer precision with high repeatability. Each channel measures the length of the laser  beam between a fiber -based sensor head  (or optical collimators)  and a mirror. The position target mirror s  show n in Figure 13 b made identically an d bonded  at the same time with the same cure process . 

A load spring wa s included as well as a force sensor to verify BSMA’s behavior while loaded with a constant  axial force.  The test bench wa s built on a baseplate with closed loop thermal control. The go al was  to  stabilize the test bench temperature within  a 0.2K peak -peak range.  

Stroke Verification   The first test aims to verify the extreme stroke of the 2 FSPA motors. Extreme stroke means reaching the  mechanical end- stops of each motor while the other motor is at maximum operational position. This test  allows to validate that, in case of a motor failure, the other motor will be able to compensate and preserve  BSMA capability to reach any position inside the operational range.  

Figure 14 shows  the output shaft position. Positions for the nominal and redundant motors are estimated  from the commands sent to the BSMA.
9   Figure 13 - BSMA Stability Test Bench  

Figure 14 - Extreme Position Verification Test in A ir     

(a) Overview of the loading assembly  (b) The 3-channel interferometric  displacement sensor without loading  assembly      (c) Inside overview of the test bench  (d) Outside overview of the test bench in  ISO7 clean room  

Glued target mirror  

                   (a) Full stroke  (b) Direction change  

End-stop verification of margins
10 Step Sizes  A major characteristic of the BMSA is its ability to generate very small steps.  Figure 15 shows analyses of  step size for nominal and redundant motor when cycling over the operational range and against its lifetime.  Performing a direction change does not impact step size nor shows any sign of backlash ( Figure 15b).   Average displacements (red points) for 3 consecutive steps are very stable for both motors and over the  complete range ( Figure 15a). Step size appears noisier for forward steps compared to backward. This result  remains to be investigated to separate effective step variations from acquisition and post -processing noise.  

Figure 15 – Step Size Analyses   

The step size distribution appears to follow a normal law with a standard deviation of 4.3 nm for the nominal  motor and 4.9 nm for the redundant motor ( Figure 15 b). In other words, the variation for more than 99% of  steps is lower than ± 10 nm. Averaging on 3 steps decreases significantly the step size scattering (less than  12% variation for 3 steps averaging).  

Lifetime Test  The lifetime qualification was still underway at the time of this document. Air lifetime is complete and shows  a slow decrease of the average step size. Less than 4 nm of step size decrease has been verified after  1 million steps  (320 times the cycle life ). The step size distribution shows a comparable standard deviation  before and after the lifetime test in air.  

Figure 16 - Step Size Against Lifetime      (a) Step size against the stroke (Flight Model  FM1)  (b) Step size for nominal & redundant motor  averaged on 5 steps (Flight Model FM1)  

   (a) Approximative step size against lifetime  (Qualific ation model  – cycles in air ) (a) Step size distribution over the operational  stroke before and after 1.2 million steps
11 Test Bench Verification for Nanoscale Stability   A major difficulty with the stability  test is that a temperature variation of 0.1K is sufficient to create a position  deviation higher than 70 nm on the test bench references.  

First, the test bench  intrinsic  stability , without BSMA , was verified . For this purpose, an aluminum board  was used instead of the BSMA to measure the drift ( Figure 17). 

Figure 17 - Test Bench Stability Verification without BSMA  

Figure 18 - Test Bench Stability without BSMA  

The center position show s a very stable beh avior while the side references were equally drifting ( Fig. 18a).  The conditions of the test show a stability of around ± 100 mK over more than 200 hours ( Figure 18b). The  phenomena explaining this drift is still under investigation. The difference between the central mirror and  the reference significantly reduce short term thermal noise.  

A second verification was to ensure that  thermal stability was below  100 mK in short -term (30 min) and  300 mK in long- term (60h).  The Figure 19 shows the temperature prof ile during the stability test presented  on Figure 20 (FM1).  

Reference measurements are then used to compensate test bench thermal expansion. This method was  found necessary to verify stability requirements. However, it remains difficult to compensate very fast or  very slow variations due to complex heat propagation in the test bench. For this reason, it remains key to  reduce temperature variations to a minimum.  

   (a) Position stability  (b) Thermal stability of the dummy plate
12    Figure 20 - Thermal Measurement During Stability of FM1  

Stability Test  One crucial objective of this test bench is to verify the stability of the BSMA position over time. The first step  of the test is to place the BSMA in stable environmental conditions (vacuum, regulated temperature) and  then wait more than 60 hours for test bench and BSMA internal components to stabilize. A single step is  then performed, and position is monitored for another 30 hours.  The stability test shows that the ± 150-nm  stability requirement over 24 hours is achieved.  

Figure 20 - FM1 Post-Processed Position during S tability Test  (Difference between output shaft position and left reference)  

Figure 20 shows the results of the stability test for the flight model FM1. A difference between the output  shaft position and references is performed in order to remove the main effects from temperature variation.  Moreover, a correction slope is removed from the test bench verification test without BSMA described in  the previous paragraph.  

The post -processed stability curved shows an exponential decrease, then the step and a rather stable  behavior. A possible explanation for the exponential effect at the beginning would be the PPA ceramics  stabilization following air -vacuum transition. It is  interesting to note that the step is not followed by a loss of  stability.  

Mechanical Environment Testing  

Random vibration testing was performed on the qualification model and flight models. Three axes were  tested. Flight models were protected against co ntamination during the test  by a plastic film.
13   Figure 21 - BSMA Random Vibration Testing  on CTEC E lectrodynamic Shaker  (Flight Models)  

The power spectral densities presented on the Figure 22 were applied on all  3 axes. The RMS levels were  gradually scaled to different amplitudes up to the maximum RMS values shown in the Table 1.  

Figure 22 - Random PSD Profiles for M echanical Environmental Testing  

The eddy current sensors embedded in the BSMA were monitored during the tests to detect any  change in  position of the rotor. Above a  given  vibration  level the BSMA rotor moved slightly  during the Y and Z axis  vibration tests  (Table 1). It is interesting to note that no damaged occurred during the tests and neither the  step size or stroke were influenced when comparing the functional tests before and after vibration.  Investigations are still ongoing for a better understanding of the rotor movement.   

Table 1 – BSMA reliability against vibration levels  

BSMA   Excited axis Random level   (g RMS)  Actuator  damaged  Step size or  performance  change  Position shift  during test   OX 5.7  No  No  No   OY 4.5  No  No  No   6  No  No  Yes   OZ 8.25  No  No  No   11  No  No  Yes
14 Conclusion  

A new patented piezoelectric motor is presented with associated experimental results  achieved on the  qualification and flight models. This motor is undergoing qualification to be used inside IASI-NG Instrument  onboard METOP -SG satellite.  

The major requirements for the BSMA we re fulfilled:   • Movement: 40- nm resolution step motion over a ± 40-µm range with up to 20- N force  • Position stability experimentally verified: less than 100nm variation over 48h,   • Launch capability  without launch- locking mechanism : No damage at all tested levels, some  movements occurred for highest levels.  

Acknowledgment  

The authors thank Fr édéric DI GESUS and Francois FAURE  from AIRBUS DS as well as Laurent  CADIERGUE from CNES, for both their technical and financial support.  

References  

1. S. Henein, P . Spanoudakis, P .Schwab, I . Kjelberg, L.  Giriens, Y . Welte, L .Dassa , R. Greger, U .Langer,  Design and Development o f The Point -Ahead Angle Mechanism For The Laser Interferometer Space  Antenna (LISA ), Proc. Conference ESMATS 2009   2. F. Barillot, K. Benoit, C. Belly, A. Guignabert, O. Freychet, Fine Stepping Piezoelectric Actuator (FSPA)  for IASI- NG, Conference ACTUATOR,  Bremen (G), 25 -27 june  2018 – Proceedings B5.5 p26   3. Stepping Piezo Actuators  (SPA) , Cedrat T echnologies , 2019, from https://www.cedrattechnologies.com/en/products/piezo- motors/stepping- piezo- actuators.html   4. C. Belly, T. Porchez , M. Bagot, F. Claeyssen,  CEDRAT TECHNOLOGIES,  Improvement of Linear and  Rotative Stepping  Piezo Actuators using design and control, B2.3 Proc ACTUATOR  2012, Pub Messe  Bremen (G), June 18- 20, 2012   5. T. Porchez, F.  Barillot, C.  Belly, Nanometric  positioning with piezo  actuator  and high stability  strain   gages , Proc. Conference ESMATS 2015
18   Figure 3. Lightweighted Aluminum Scan Mirror  

Figure 4. One of Three Mirror Mount Flexures with Associated Plate for Liquid Pinning  

Prototype testing early in the design process uncovered a high sensitivity of mirror flatness to the flexure  bolt preload torques. This finding resulted in t wo significant design changes. The first change was to remove  mirror material under the flexure installation locations  creating an undercut (Figure 3 ) such that strains could  not propagate to the opti cal surface directly. This change was effective but did not alone minimiz e errors to  acceptable levels. A  second change was implemented to result in lower bolt preload in the cold operating  environment while also maintaining preload at ambient conditions t o prevent interface gapping under launch  vibration. This was achieved using a low -CTE titanium bolt and high- CTE aluminum standoff (Figure 4).  Figure 5 provides thermal gradient sensitivity analysis results of the mirror surface.  Vibration and low  temperat ure interferometric tests confirmed the soundness of this strategy.  Figure 6 provides one of the  interferometric test results.
19  

Figure 5. Thermal Gradient Sensitivity Analysis  

Figure 6. Prototype Flexure Torque Interferometric Assessment  

Strength testing of the aforementioned liquid pin joints was performed to assess shear l oad capability of  the bonds. This testing was critical to the convergence of a design capable of withstanding launch loads.  Injection and bond geometry, surface preparation, and adhesive selection were all adjust ed as a result of  these tests. The first attempt revealed that the injection inlet and outlet size and cavity depth would only  allow very low viscosity polymers to flow , limiting selection to materials whose strength was insufficient.   Increasing the size of these features also allowed the use of larger injection needles and more manageable  injection pressure. Sanded, grit -blasted, and etched/primed surfac e preparations w ere evaluated. The grit  blasting process proved too aggressive for use with these very small samples, as the cavity containing  walls were eroded, allowing adhesive to migrate out of the cavity. There was no significant difference  between the sanded and etc hed/primed samples, however the latter approach was selected for f light  because the primed surfaces were expected to maintain good surface preparation longer than sanded  surfaces. An in -process low temperature interferometric evaluation wa s performed after  the surfaces we re  prepared but before the cavities were injected, s ince the stability of the bond surface over time is critical.   Flight design and process selection involved coupon testing, depicted in Figure 7 .
21 The flight  SMM was installed into  the TVAC  chamber shown in Figure 9 . At specific temperature plateaus,  whether at qualification or operational, the SMM dwelled for approximately f our hours in order to attain  sufficient thermal settling of the mechanism. The flex  pivots  provided a constricted  thermal path requiring  somewhat long dwells to achieve a reasonable temperature gradient that would be good enough for  validating mirror angle. Thermal cycle plots at each side of a flex  pivot and mirror are shown in Figure 1 0.  Controller Performance Chan ge Over Temperature   Controller parameters were initially optimized at ambient temperature resulting in a 30- Hz bandwidth.  Frequency response measurements were taken at various TVAC temperatures as shown in Figure 11 . It  was found that controller performanc e became less optimal as temperature decreased. As a result of  reduction in resistivity of the copper damping sleeve, damping increased with decreasing temperature,  ultimately by about a factor of 2.6 going from 25° C to -130°C. Though high damping is very beneficial in  terms of disturbance rejection, it was necessary to optimize the controller within the SMM operational  temperature range. Thus, controller parameters were chosen to provide optimal performance at operational  temperatures. A process of system identification was performed while undergoing TVAC to determine  damping as a function of temperature. Fortunately, controller performance, though not optimal at 25° C, was  adequate to perform ambient testing without the need to change controller parameters.  

         Figure 8. Flex Pivot X -Ray Cross -Section         Figure 9. TVAC Chamber before Closure
24 testing. Further experimentation by  the vendor  included measur ing the temperature change of the sensor  heads with thermocouples for full -range moves of their test fixture. Results from this testing show both that  full-range steps cause the temperature deltas predicted by the sensor head heating test, and that the time  constants of both the temperature change and position error are roughly correlated.   To minimize impact on the existing controller design, correction for this drift was handled wit hin the FPGA  firmware but outside the feedback loop. An algorithm  that predicted the sensor drift was summed with the  position command input .   The results in Figure 12  show a comparison of full -range step response utilizing the correction algorithm to  correct for the sensor drift (“Cmd Shaping” in orange) and without it (“No Shaping” in blue). The top plot  shows voltage measurements from the DPSS converted to angle, while the bottom shows angle  measurements from the autocollimator. When tu ned properly, correcting the sensor drift in this way was  able to reduce the overshoot error by a factor of 10, from 150 microradians ( µrad) down to 15 µ rad. TVAC  testing has shown that the parameters of this correction vary both with pressure and temperat ure. Analysis  of that data is ongoing.  

Figure 12 . Command Shaping Performance Improvement  

Mechanism Control Electronics  Development Strategy  

Aggressive schedule demands required a compressed MCE development schedule. Three distinct builds  were designed:  An Engineering Model (EM), an ETU, and flight . This design flow allowed early design  validation, while providing the necessary hardware for interface and controller development and testing.      The first EM  MCE  build utilized commercial equivalent parts instead of flight grade parts. Dual footprints  were incorporated into the layout to accommodate package differences between flight grade and  commercial parts. To prevent schedule delay due to a roughly 1- year electronic part lead time, an alternate  (but functionally identical) part was used for the power op- amp. A reprogrammable FPGA module was used  for development in place of the one- time-programmable FPGA that wo uld be used on the flight  boards. A  build plan change added t wo EM MCE b oards  to facilitate controller testing and development, and interface  testing. This also meant  the schematic and Printed Circuit  Board ( PCB)  layout were verified before  ETU/ flight  and allowed for a dry run of the assembly process. Board l evel electrical checkout and functional  testing procedures were also developed at this stage. A minor net -swap issue was found in the schematic  design and fixed with a white wire on the EM,  and the schematic  and layout  were updated for flight  PCBs.  

Since the SMM was not available early enough for initial EM board testing, a mechanism simulator was  built. The structure was 3- D printed, and commercial flex pivots  and a prototype aluminum mirror were used