ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
25 to closely mat ch the expected plant dynamics. Commerci al voice coil actuators  and a commercial DPSS   allowed open- loop and closed- loop testing. This rapid prototype allowed controller and FPGA development  to move forward months before the ETU SMM would be available. It was also useful as a stand in for the  ETU SMM and invaluable for testing the optical verification setup.  Views  of the mechanism si mulator are  shown in Figure 13 . 

The ETU builds utilized parts appropriate for environmental testing. The second ETU build was assembled  with the updated flight  PCB. By working out assembly and test procedures on the EM and ETU builds, it  was possible for flight  assembly and test to be completed very rapidly. Because the flight  FPGAs are not  reprogrammable, a flight  FPGA was burned and installed on an EM board first to verify the design was  successful , buying down risk.  The flight  MCE is shown in Figure 14 . 

Figure 13 . Mechanism Simulator  

Figure 14 . L’Ralph Flight Mechanism Control Electronics  

Flexibility of the Controller w as Crucial to Project Success  

As previously discussed, the  digital FPGA- based controller allowed for correction of position sensor system  non-linearity and thermal error. It also facilitated system identification at low  temperature to determine  damping by allowing adjustment  of controller parame ters. However, at the start of the project, a digital   controller was not the baseline.  

During the proposal phase, an analog controller with a simplistic digital section was chosen as the baseline.  One reason for that selection was that a digital  controller utilizing an FPGA was deemed too power hungry
27  

Figure 15. L’Ralph Flight Scan Mirror Mechanism Front  View   

Figure 16. L’Ralph Flight Scan Mirror Mechanism Rear View  

References  

1. Blumenstock, Kenneth A. , et al.  " ATLAS Beam Steering Mechanism Lessons Learned."   Proceedings of the 43rd Aerospace Mechanisms Symposium, (May 2016), pp. 1-14.   2. Donegan, Richard J.  “ Weld versus Braze. ” [Engineering considerations regarding electron beam  welded versus brazed flex pivots ].  Retrieved from http://riverhawk.com/weld -versus -braze/ , (n.d.)
30 Mechanism Overview   The piezo actuators are wired  in 2 push- pull configurations (1 per axis) to allow direct mirror rotation control.   The PAM itself is composed of the following parts:   - A bracket baseplate (in aluminum): The APA® are fixed on it with screws.   - 4 APA® (in titanium): They provide the required displacement and are fixed to the baseplate and to  the mirror support via flexural pivots. The APA® are equipped with Strain Gauge ( SG) sensors  that  are bonded in place   - 4 circular pivots ( in titanium ).  - A mirror support (in INVAR) which holds the mirror.   - A guiding blade ( in titanium ) soldered onto the central cylinder that stiffens the assembly.   - A Silicon Carbide (SiC) substrate- based mirror from Mersen OptoSiC® 

Figure 1. PAM Overview Including Piezo and O ptical Technology from Cedrat Technologies  

Strain Gauge Sensing   In order to be able to monitor the mirror angle, an indirect solution using strain gages placed on each piezo  actuator is selected, based on space heritage from other projects, especially ATLID on this matter, which  enabled an important development on SG assembly process. The initial SG redundancy requirement was  lifted, because it required an important wiring complexity (32 instead of 16 wires), amongst other additional  constraints.  

The project used constantan, 350- ohm SG. There are 2 S G per piezo stack , mounted in one full Wheatstone  bridge per rotation axis to maximize the sensitivity while minimizing thermal drift. All SG wires and printed  circuit board ( PCB) traces are the same length to limit offset drift.  

New Piezo Actuator Design   The existing CTEC actuators were either slightly too short in stroke or not stiff enough to ensure the  mechanism survival during launch. The mechanism consists of 4 APA®, deriv ed from CTEC standard  APA120S but specifically designed for the application needs.  

Based on CTEC space heritage, the APA® shell was made from Ti6Al4V titanium, allowing a theoretical   infinite fatigue lifetime in the specified operational conditions and an interesting stiffness/mass ratio. Another  benefit  of the use of titanium rather than steel as used in standard products , is the reduction  of the thermal  stroke  effect due to a better CTE match (9 ppm/K) to the piezo stack (typically - 3 to + 1ppm/K) compared to  a high performance stainless  steel for example (10- 11 ppm/K) . A total of 18 APA® were assembled and  tested, the results are indicated in Table 1.
31 Table 1. PAM Actuators Measured Results     Full stroke   (170 V pp) 1st coupled  resonant frequency    µm Hz  Average measured  149.2 6963 .6  Standard deviation  measured  2.0 60.9  Design value (worst case)  130.4 6151   Difference measureme nt/design value  14% 13% 

The design values are based on expected worst case parameters , i.e.,  low piezo gain and stiffest shell  when calculating the stroke. A conservative approach was used to ensure that the required mechanism  stroke would always be reached,  which resulted in average margins  of 14%  for the stoke and 13% for 1st  resonant frequency . 

Integrated Optics – Mirror Development  The PAM is designed around one of its core components: the mirror. For this new project, CTEC used its  experience on piezo- optical aerospace projects and collaborated with partners in order to fully integrate the  optical hardware  development, from the SiC mirror design to its integration and optical verifications.  

Mirror design   One of the main design constraints of an embedded optics mechanism is to keep the mirror surface  deformation to a minimum in order to limit the induced optical wave front error below the requirements. I n  this case, a maximum of 63- nm mirror surface flatness is the requirement . In order to ensure the  specification would be reached, CTEC developed tools and performed specific simulations in the early  design phase,  specifically includin g verification of induced surface figure error  caused by mechanical biases   and thermal deformation , as well as optimi zation of mirror shape and dimensions.  

Figure 2. Mirror Surface Deformation Simulation Meshing and Boundary Conditions  (Screw Tension and  Torque Case) (Left ) - Mirror Surface Deformation Evaluation Results ( Right ) 

The mirror design process is based on a finite element simulation of each mirror deformation contributor:  mirror clamping to its support, actuator height variation, screw tightening  and thermal operational. Each  case used  representative boundary conditions. The resulting displacement on each of the mirror surface  points is then exported. A MATLAB program is then run to post process the results from the simu lation.  Extracting the RMS reference plane from this data set, it calculates the distance of each point to this plane,  the peaks and valley of the deformed mirror and the resulting RMS deformation.  The specific contribution  of each evaluated case on the mi rror deformation is then summed up to give an estimation of the total  expected mirror surface deformation.
32 PAM development also included a regular mirror optical verification at different stages of the assembly , with  the intention of having the ability to stop the process should a mirror appear to be out of the acceptable  range or show significant  surface figure changes  from one step to another. With the recent CTEC  experience on this matter, such regular controls also allowed us to learn a lot regarding the impact of each  assembly step as well as providing a safer project assembly process . 

Figure 3.  Mirror interface (left) and surface verification with interferometer (right)  

Mirror Verification Results   A total of 4 PAM were produced and the mirror Reflected Wavefront Error, Peak -to-Valley (RWE PV,  basically 2 times mirror surface flatness  error SFE ) was measured in different phases, i.e., mirror in the  initial free condition, once clamped on its mount , and eventually after final integration in mechanism.  

Table  2. PAM Mirror Surface Flatness Measurement Results   PAM model  1.Free coated  mirror  2.Mirror clamped  on mount 3.Mirror  integrated in  final mechanism  Calculation:  Mechanism  contribution to  mirror  deformation     RWE  PV  (nm) RWE  RMS  (nm) RWE PV  (nm) RWE  RMS  (nm) RWE  PV   (nm) RWE  RMS  (nm) RWE PV  (nm) RWE  RMS  (nm)  EM 38.1 10.1 34.9 8.9 41.3 10.6 3.2 0.5  EQM  22.2 6.2 27.2 6.6 38.7 10.3 16.5 4.1  FM1 23.2 5.5 19.4 4.6 26.6 6.0 3.4 0.5  FM2 28.9 8.1 33.8 8.2 34.1 9.5 5.2 1.4                     Average  28.1 7.5 28.8 7.1 35.2 9.1 7.1 1.6 

With an average 35. 2 nm and a maximum of 41. 3 RWE PV, the <127- nm RWE specification was  reached  with significant margin.  The last column evaluates the proper mechanism contribution to mirror deformation  based on the difference between final and the initial RWE PV measurements.  

The mechanism average  contribution to the mirror deformation appears to be limited to 7. 1 nm PV and  1.6 nm RMS .
33 Eventually, we can compare these results to the expected worst -case deformations evaluated during the  design phase.  The t hermomechanical contribution could not be evaluated so only mechanism induced  deformation are considered .  

Table 3. Comparison of mirror measurements and expected design worst cases     Average measured  impact of mechanism  on mirror RWE  (nm)  Average measured  impact of mechanism  on mirror SFE  (nm)  Evaluated worst case for  mirror SFE (mechanism  integration only)  (nm)   PV 7.1 3.5 5.7  RMS  1.6 0.8 3.1 

All measured values are within the expected range of mirror deformation, indicating that the conservative  simulation approach was correct in that case.  

Mechanism Production  and Assembly  

Titanium Parts – Alpha Case Removal Process Qualification   The use of titanium for some parts, justified by its interesting mechanical properties and heritage on  previous CTEC mechanism s (for similar parts) is not without drawbacks . The main issue is due to the use  of wire electro discharge machining (WEDM) for the “flexible” parts manufacturing, i.e. , the actuator  amplification shell and the guiding blade at the center of the mechanism.  This manufacturing technique is  required due to the parts geometry.  

       Figure 4. WEDM titanium parts, guiding blade (lef t) and actuator shell (right), different scale   The WEDM locally heats the material which can induce the formation of an alpha  case on the surface, with  the adverse effect  of lifetime reduction (up to 30%  according to [6]) , and possibly  unexpected failures . This  alpha case, usually a few µm thick, can be removed through proper chemical etching.  

For this mechanism, the already qualified manufacturer was not compatible with the required schedule  (overload), hence a backup option had to be identified and qualified. The new supplier proposed an acid  etching process . The qualification process was the following:   • Machine a set of guiding blade and actuator shell, identical to flight design (supplier, process,  material batch, dimensions)   • Perform an alpha case an alysis before acid etching, parts in initial state, as well as an interstitial  hydrogen contamination measurement.   • Pass the parts through the acid etching process , adjusted to remove the required thickness   • Perform an analysis after the chemical etching: alpha case and interstitial hydrogen  contamination   • The process is qualified if analys es show no trace of alpha case after etching and no more than  150 ppm interstitial hydrogen contamination.
34       Figure 4. Alpha Case Analysis, Before ( Left) and A fter the A cid Etching ( Right)-Optical Microscope  (x1000)   No alpha case was found on that actuator shells ; on the guiding blade however , the average alpha case  thickness was 8  µm (6 to 11  µm). It was decided that a 15- µm thicknes s removal through acid etching would  be enough to remove the alpha case.  

The final analysis  after the acid etching indicates that the process efficiently removed the alpha case (see  Figure 4), no trace was left on the part. The hydrogen contamination, that can be induced during the acid  etching (penetration of hydrogen inside the material compound and local embrittlement) remained within  the boundaries (86 ppm for guiding blade and 26 ppm for actuator shells). The process is then qualified  and was used successfully to treat flight parts.  

Piezoelectric Stacks – Lot Acceptance Tests (LAT)   As one of the critical items in the mechanism (brittle, sensitive part), the piezo stacks are handled with  special care. One single piezo batch is procured for the project, with high quantity margin (at least x2).  An  LAT is then performed on 4 piezo stacks taken from this batch, prepared (SG gluing & cabling) exactly as  flight piezos.  

This LAT includes thermal cycling and lifetime tests representative to the final envir onment (this includes  mounting them in their actuator shell) . Regular basic electrical verifications (capacitance, insulation), stroke  measurements are performed before and after the tests  to detect any deviation. A final destructive physical  analysis is performed to inspect inner features of the piezo stacks mainly to detect potential voids in the  ceramic and electrode delamination. Piezo stacks are cut in two and sections are inspected.  

    Figure 5. Solder Inspection, Example of a Crack 

The destructive physical analysis reveal ed that the material is dense enough, no voids we re detected.  Electrodes were perfectly in place and no trace of delamination was found.  In compliance with CTEC
35 previous experience, small 45° cracks we re detected in the vicinity of the electrodes. The se cracks are  believed to be caused by the thermal expansion and contraction of the electrode during soldering. Since  they are covered in epoxy potting, th e cracks’  progression is contained.  

However , one unusually placed and lengthy crack was  detected. Unusual by its location, starting from piezo  edge instead of from the electrode, and by its length,  2.2 mm compared to the usual few hundred µm.  

Figure 6. Piezo N°3, Black Wire Side Unusual Crack ( Length 2 .2 mm)  Raising questions , investigations were performed but no clear specific root cause was identified. It was  found that this crack was slightly visible from the exterior and most importantly not going through the entire  stack (staying near the electrode, under the epoxy ). 

Figure 7. Exterior View of the C rack  One possible root cause for this crack is stress concentration due to a slight mispositioning inside the  actuator shell  and to the selected prestress level . Simulations representing this mispositioning allowed  CTEC to map the induced stress concentration and the result is visually similar to  the observed crack  pattern.  

   Figure 1. Photo of Piezo Mispositioning and S imulation of C orresponding Principal Deformation Map (Cut  View) Matching Crack Pattern
36 To avoid this problem, the final flight actuator shell will be larger (9 mm instead of 5 mm) than the one used  for the LAT, making it impossible for the piezo face to be in contact with an edge like this. Also,  as a  precaution, all piezo stacks were inspected before assembly in flight hardware and no exterior sign of similar  cracking  was found . 

No other unusual findings were identified , and the piezo batch wa s accepted for integration.  

Mechanism Integration   One of the main constraints  of this mechanism integration is that the support bracket (black anodized  aluminum part) has its interface surface (fixed to the L3 optical array) perpendicular to the mechanism base  plane. This geometry severely limits  access to the mechanism for torque wrench es, operator hands, pliers  etc. Hence the assembly  process had to integrate this specific constraint with the unusual approach to preassemble the actuation mechanism outside its final base plate, then transferring it onto the final support  bracket. The final mechanism once assembled is shown in Figure 9.  

Figure 9. PAM EQM P icture After Assembly   The most sensitive part of the mechanism is its SiC coated mirror  and many precautions were taken to  protect it during the assembly. A specific cover (POM- C) was designed and was used for most assembly  phases, especially to protect the mirror surface from wrenches  or pliers when needed.  In order  to ensure  additional protection during transportations phases (models are transported several times during assembly  and test phases), a transparent Plexiglas box was  designed.  

   Figure 10. Mirror Protection Cover Once Installed ( Black Part), PAM T ransportation Box
37   Cables and PCB I ntegration  A total of 16 SG sensor wires and 8 piezo power wires have to be routed from the mechanism to the  interface cables. The use of a connector on the PAM side was not possible due to size constraints.  The  selected design option was an interconnection PCB , similar to what was used in previous CTEC space  mechanisms  like the ATLID Beam Steering Mirrors(1). The multilayer PCB provides many benefits :  • Allows pre- routing of all SG bridges easily   • Easier wire handling and more reliable interconnection (from AWG36 SG wires to AWG26 pigtail  cable wire, large  number of wires)   • More control of SG trace length (and impedance) to ensure a reduction of SG offset thermal drift   • SG and piezo signals are routed through different layers, insulated and shielded from each other  (ground planes  in the middle of the PCB )  • Interconnection PCB was used in previous similar spa ce projects with positive feedback  

Figure 11. PAM PCB routing overview (merged layers for easier visualization)   Wires are routed from their starting point on the piezo actuators (SG and power), through the support  bracket via holes , up to their dedicated PCB pad. Wire s are regularly secured along their paths with epoxy  dots. One of the constraint s of SG wire is for them to have the same length within each W heatstone bridge  (one for each axis) to ensure they maintain the same total resistance within the target range.  

   Figure 12. PAM Cabling to Interconnection PCB
38 Qualification and T est Campaign  

Static Performance   The 4 mechanism performance parameters  are verified  at different steps of their acceptance tests. An initial  good health verification wa s also performed to ensure that the piezos and SG we re correctly cabled.  At the  time of this paper publication, only 3 mechanisms were tested, results are presented in Tab le 4. 

Figure 13. EQM Y A xis Full Stroke Measurement ( Mirror Angle vs P ush-Pull Voltage)  and Admittance  Sweep for EQM Y axis –  Coupled Resonance Frequency Identification  

Table 4. Static Measurement Results for PAM EM/EQM and FM2   Parameter  Required value  EM EQM  FM2  PAM total stroke at ambient ( -10/+150V)   X axis  >6 mrad  8.0 mrad  8.7 mrad  8.5 mrad   Y axis  >6 mrad  8.6 mrad  8.6 mrad  8.5 mrad   SG response at ambient - Functional tests FTM -01-B (cf, III,8)   Offset           X axis  / -0.71 mrad  -0.27 mrad  -0.46 mrad   Y axis  / -0.76 mrad  -0.35 mrad  0.11 mrad   Gain           X axis  / 1.57 mrad/V  1.49 mrad/V  1.76 mrad/V   Y axis  / 1.38 mrad/V  1.54 mrad/V  1.67 mrad/V   1st coupled resonance frequency   X axis  >800Hz  1495 Hz  1301 Hz  1257 Hz   Y axis >800Hz  1242 Hz  1272 Hz  1272 Hz   PAM mass   Total mass  / 497.2 g 497.9 g 495.1 g 

The mechanism total stroke is compliant with the specification with significant  margin,  since design  anticipated some stroke loss at  cold temperature. The total stroke is similar for all models/ax es, the slight  variation observed is expected and linked to the piezo actuators ’ stroke/stiffness variation.  

SG parameters are noted for information, only the variation of these parameters ( over temperature) matters.   Note that SG measurements are taken after bridge output conditioning (5V excitation, 337 and 352  V/V  gain for X and Y).
39 Thermal Vacuum Testing   The Qualification campaign for the EQM included both non-operational  (NOP)  and operational  (OP)  thermal  vacuum cycling. Target vacuum is <10- 5 mbar. Tests A,B,C are functional tests (stroke, SG and admittance  sweep).  

       Figure 14. NOP and OP T hermal Vacuum Cycling Test Schematic  

The main objective of the OP Thermal Vacuum Cycle is to characterize the variation of PAM stroke   (expected loss in cold)  and SG performance throughout the OP temperature range.  Only the results of this  test are detailed in this paper.  The OP Thermal Vacuum Cycle consist s of testing the mechanism (full  stroke, SG parameters) at different incremental temperatur es (every 10°C).  

Given that the PAM applied voltage was not constant during the full test (4 days  with electronic and room  temperature variations ), the stroke result s are normalized with respect to the voltage range applied, called  stroke/voltage gain. The stroke results are shown in Figure 15.  

Figure 15. X and Y Stroke/ Voltage Gain Through OP Temperature Range – EQM Results   The stroke/voltage gain plot for both axes show the expected bell shape, with a maximum at ambient  temperature and loss at  cold and hot temperatures . A 7% loss can be seen at -30°C and around 2% loss  at +60°C. A conservative 20% and 15% stroke loss was ass umed in the design phase (based on past  experience on worst cases), explaining the high  final stroke margin.  

The other noticeable feature is the thermal hysteresis , which  should not be there: it is expected for the  mechanism to have the same stroke for same temperature (the stroke vs voltage hysteresis observed  is  however nominal and expected) . The reason for this is probably a consequence of the fact PAM  temperature was not fully stabilized, tests were performed slightly too quickly after reaching target  temperatures and resulted in this thermal hysteresis.  The actual temperature data is measured on the PAM  support bracket (interface) located under the mechanism, piezo temperatures  were not monitored to avoid  any potential damage/contamination. Hence it is not possible to directly measure or control  piezo  temperature.
40   In order to confirm this thermal delay hypothesis, a simplified thermal -equivalent first order model of the  PAM was created, using measured temperature time constants (between PAM interface and the piezos)  and a bell -shaped stroke vs temperature look -up table for the piezos.  

Figure 2. Simplified PAM Thermal Equivalent Model Results (X: T emperature in °C, Y: Stroke in mrad),  Left: 10h Cycle Duration ( Close to W hat was Tested), Right: 100h Cycle Duration where Hysteresis is  Greatly Reduced.   The model can recreate this hysteretic behavior, which appear s to be greatly reduced after greatly  increasing thermal stabilization time  (x10) . However, g iven that the measurements cannot be automated,  an operator has to be there at each step (hence excluding nights and weekends ). The EQM test already  lasted 3 days and 30 days of testing is not practical . Options to improve the OP Thermal Vacuum Cycle   were discussed with the customer.  

SG parameters (offset and gain) were also measured for each temperature step, the results for EQM X - axis are shown in Figure 17.  

          Figure 17. EQM X A xis SG Gain vs T emperature ( Left), EQM X Axis SG O ffset vs T emperature ( Right)  The results are also affected by the piezo temperature delay . As anticipated, the SG gain is quite stable  with less than 1% variation through the temperature range. The offset is directly impacted by the mechanism  thermomechanical excursion, which was meas ured at 0 .4 mrad total (0. 2 to 0. 4 mrad on X and Y SG offset).  

Vibration Test  All models except EM are planned to go through vibration tests.  The tests consist of a random vibration  verification as well as a shock test, for each axis. A low-level frequency sweep is performed before and  after each test (to assess potential modal landscape changes ).