ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
68 The strong difference of behavior between the test s with a mechanical release system and the tests  done  with a fuse spool indicate that the issue was related to the spool and not to the locking ring and stage  mechanism. This behavior was then discussed wit h Eaton, the NEA supplier, which identified t wo potential  root causes:   

• The actuation current applied is under the required minimum. Actuation is still possible however  the fuse wire tends to melt to slowly which has  sometimes  resulted into spool getting stuck closed .  When the right current is applied, the fuse wire burns  very fast (flashing behavior) and most of it  sublimates . This issue has  also been seen on other development s [5]. However,  such a root cause  does not explain how after some time the spool released . In similar cases , the spool remain ed  stuck indefinitely .  

• The resetting procedure used could wear the surface of the spool . Such a wear could create barbs  of plastic which prevent opening. After some time under the pressure of the plunger, the plastic  barbs creep and the spool finally releases .  

Table 4: Unlocking test duration and temperatures   Test description  Unlocking  duration (s)  Test  temperature   Thermal vacuum chamber  test EQM  3592  [-59°C; 21°C]   Thermal vacuum chamber  test FM  74 [-56°C; -49°C]   Mechanical unlock 1  0.78 -56°C   Mechanical unlock 2  0.49 -56°C   Mechanical unlock 3  0.55 -56°C   Mechanical unlock 4  0.20 -57°C   Mechanical unlock 5  0.28 -71°C   Spool unlock 1  3117  -57°C   Spool unlock 2  0.42 -56°C   Spool unlock 3  25 -57°C   Spool unlock 4  4457  -57°C   Spool unlock 5 with new reset procedure  0.42 -57°C   Spool unlock 6 with new reset procedure  3.48 -57°C   Spool unlock 7 with new reset procedure  250 -56°C   Spool unlock 8 with new reset procedure  3.20 -57°C   Spool unlock 9 with new reset procedure  106 -57°C  

In order to solve this issue, the approach implemented considering the very tight schedule and the quite  advance status of the development was to update the launch lock resetting procedure. In this update,  tooling was developed to avoid any wear of the spo ol during its installation into the NEA. With this solution  implemented further testing was performed. It can be seen in Table 4 that the durations  obtained with the  updated procedure are significantly better than the ones before.  

While the implemented solution did not fully solve the issue, considering the mission schedule and the issue  associated risk, the deviation from expected behavior was accepted. The main argument backing the  acceptance is that despite the sometimes  very long delay, the launch lock  did not ever get stuck: i t always  released.  

Lessons Learnt for F uture Development  

Looking retrospectively at the CLUPI focus mechanism development, it seems that many issues could have  been averted and that better solutions could have been implemented to solve encountered issues . This  observation is to be put into perspective with the mission major constrains whi ch were: mass, cost and
69 schedule. Furthermore, these solutions may appear straightforward, now that the development is complete,  but they were certainly not as obvious at the beginning of the project. That is precisely why it is important  to look back to t he now completed development and identify important lessons that would help future  development s.  

These lessons learned presented here are two- fold. There are specific lessons about what could be done  from the beginning if a very similar mechanism were to  be developed. Then there are more generic lessons  about aspects which had an important impact on the CLUPI Focus Mechanism but are also applicable to a  lot of  other  mechanism developments .  

Improvement for a F uture Focus Mechanism   Solutions were found to answer the issues and address the challenges of the CLUPI Focus Mechanism .  However, these solutions had to be compatible with the very tight mission schedule. For the development  of a future focus mechanism such solutions could be implemented at earlier s tage in the development and  thus be better integrated in the design. In some case, alternative more elegant solution s could be  implemented. All these solutions were presented in detail in [ 6] and are summarized hereafter:  

1. Improvement of mechanical interfaces:  It would allow a reduc tion of the amplification and  transmission of the mechanical load to the focus mechanism.  

Figure 12: Improvement of bracket design to improve mechanical interface  

2. Transverse Blades Integrated to the Guiding  Plate Structure:  A more symmetrical design of the  guiding plate would allow a better optimization of the flexure parameter s to increase resilience  without impacting the guiding movement stiffness.     

Figure 13: Transverse blades integrated to the guiding plate structure
70 3. Coalignment of the Actuation:  Centering the actuation would enable  full symmetry on the design  and drastically improve the guiding alignment stability performance.  

Figure 14: Actuation and s ensing centered around  the optical axis  

4. Implementation of a M odel-Based Control -Law: With a model -based control law, the residual LVDT  friction could be modelled as a known perturbation (measured in calibration or on- line) to improve  the positioning accur acy 

5. Open Locking Ring to Reduce Friction During Unlocking:  Such a solution maintains  high friction  loads in the locked position to dissipate vibration energy  during launch  but minimize s the friction  during unlocking.  

Figure 15: Launch Locking System with Open Locking Ring 

Lessons Learnt for F uture Mechanism Development   Apart from potential design improvement, the CLUPI focus mechanism gave important insight about general  development of mechanism. These lessons are numbered hereafter along with a brief description about  how it became an issue within the CFM development.  

1. Resilience of C ompliant Mechanism is Always Critical:  Discuss the mechanical environment  requirements  with the mission and your customer . Mechanical requirements must be backed up by  strong rational es and all must be done to minimize the loading applied to the mechanism.
74 Figure 1. LP-PWI overview  (CAD model deployed configuration – left; CAD model stowed configuration –   right; as built model stowed configuration -  bottom)  

The boom has  two HDRMs: Base HDRM and HDRM#2, which hold the CFRP tubes  in a stowed  configuration during the launch. The Base  HDRM is placed close to the Langmuir probe while HDRM#2 is  placed close to the Central Hinge . The position of HDRM#2 was optimized to maximize  the first natural   frequency of the boom.  Both mechanisms are identical in terms of functionality ; the difference between  them lies in the structure:  HDRM#2 is standalone while Base HDRM shares the structure with the Base  Hinge.  

In a stowed configuration the boom’s length is under  1.6 m; it deploys to the length of 3 m. The LP -PWI’s  angular position must be kept within ±0.5 deg cone with a tip in Base H inge. The length of the boom must  be kept within  the tolerance of  ±5 mm. During the mission the LP -PWI will be exposed to  highly demanding  environmental conditions . There are four identical LP -PWI booms located on three different corners of the  S/C. As consequence,  the level of Sun illumination is different for each boom . The S/C trajectory comprises  a close Sun approach with high Sun illumination,  as well as long time spent in the shadow during Jovian  Tour with Sun eclipse.  This leads  to a wide range of temperatures  on the LP, ranging from -220°C up to  200°C . However , the main driver behind  LP-PWI’s architecture was a limited mass budget  – in order to  comply with requirements, each LP -PWI needed to have a mass under 1.3 kg.  

The low mass of the unit  did not allow for using hold and release mechanisms  available on the market  due  to the mass requirement . The HDRM also needed to be part of the  boom structure. In the LP-PWI, the  HDRMs are also the  main support of the boom as well as the interface with the S/C. Each HDRM contains  mechanical and electrical interfaces  with the S/C.  

The design concept of the HDRM  was based on the Hold Down and Release Mechanism from MUPUS  instrument on Rosseta mission [ 4]. A similar design was later also used in  the DRAGON -8U Nanosatellite  Orbital Deployer [ 5]. However , this specific application ( for LP -PWI) forced  a major tailoring of the  mechanism.
75 HDRM System overview  

The core of the HDRM  is a Preload Jaw which provides the preload to the CFRP  tubes  and keeps  them in  the stowed configuration (Figure 2). The Prel oad Jaw consist s of a stainless -steel  V-shape spring with two  Tension Clamp s attached to each arm of the spring.  The Tension Clamps press the Tube Interface , which  is attached to the upper 2nd section of the boom.  The upper section lay s down on the separators , which are  a part of the lower 1st section of the boom. The pressing on the T ube Interface is transfer red through  separators to the lower section and it holds both of the CFRP tubes . The Tube I nterface is equipped with  overlay  made of Vespel SP1. The T ension Clamps are made of nitride d titanium , which is essential for  further investigation in this article. The materials  are a good and stable friction pair  with friction coefficient  0.21-0.24 in vacuum, in room and low temperature ( -80°C) [6]. Vespel SP1 also has high strength,  which  does not degrade significantly in the higher temperatures (tensile strength: 86.2 MPa at 25° C; 41.4 MPa at  260°C). These features make Vespel SP1 the best choice for this application.   

Figure 2. HDRM design  

In a closed position of  the HDRM , the Preload Jaw is bent and supported by two levers. The levers are  connected to each other above the CFRP  tubes with a Vectran string. The string is wrapped around the  resis tors placed in the lever. The resistors play a role of thermal knives that cut the Vect ran string . In a  single HDRM, there are two resistors (one on each lever): primary and redundant. The geometry of the  lever  combined with the inclination angle of the T ension Clamp  (α) lowers the loads seen by V ectran string  with a ratio 28.5.  

The opening of the HDRM ( Figure 3) is initiated by a weakening of the Vectran string  by heat from  deployment resistors . The levers are motorized by t heir own torsion springs and they are released after  cutting  of the Vectran string . The constrain resulting  from the closed levers disappear s, and the V-shape  spring comes  back to its open position. The T ension Clamp s stop pressing the T ube Interface and the  boom’s sections are ready to be opened.
76 Figure 3. Closed HDRM #2 - left; Open HDRM #2 - right 

HDRM preloading   The project entailed the delivery of 4 identical LP -PWIs, hence all of the eight HDRMs shall be similarly  preloaded.  The preload cannot be determined only by its hardware.  The manufacturing tolerances did not  allow for reaching the required repeatability in the preloading. Naturally, it cannot rely only on the hardware  dimensioning and needs to be adjusted during integration process ( Figure 4). The nominal preload for one  HDRM was selected to a range of 160-170 N. A higher loading of HDRM could lead to the cumulation of  too high a stress in the structure. The preload is applied to the Preload Jaw by the T ension Tube. During  integration,  the Tension Tube is loaded and tightens the V -shape spring when HDRM stays in a closed  position. When all parts are preloaded,  the tube is blocked by 6 blocking  screws tightened to the HDRM  walls. On the one hand,  this solution releases  the mechanism from manufacturing inaccuracy , but on the  other the final preload depends on the correctness of the integration process . 

Figure 4. Application of the preload in HDRM  

Qualification  Campaign  – Failure to Open after the Vibration  Test 

Qualification c ampaign overview   The full qualification model (QM- 1) of the LP -PWI was subjected to several  functional and environmental  tests during the qualification test campaign. The functional  tests were focused on the deployment reliability.  The boom was designed to be deployed in micro -gravity.  On ground , the LP-PWI can only be opened in a  horizontal pos ition. For the deployment test , the boom requires  a 3x4 m flat table and ball support s attached  to the CFRP tubes . During the deployment , the ball  supports move on the table and offload the boom from  gravity force. The dynamic opening of the boom excluded any other  offloading methods.  Due to the  necessity of opening on a large  ground support , the deployment following  the environmental test w as  replaced with a limited release. The limited release was restricted to the opening of the HDRMs  in vertical  position without opening the hinges.   Blocking   screws
77 The environmental test included the thermal -vacuum (TVAC) cycling , shock and vibration tests. The TVAC  test was performed on a reduced qualification model (QM-2), which was limited to one hinge, one HDRM,   and shorter CFRP tubes  with a dummy mass . The unit pass ed the thermal cycling in the range of - 180°C  and 100°C , and was successfully deployed in cold conditions  (-50°C). After the test,  the QM- 2 parts  were   used in a full model QM-1. 

The QM-1 was subjected to vibration and shock test s. The v ibration  test included resonance search, sine  and random vibrations , and was performed on qualification  level.  The boom was subjected to vibr ations  in  all 3 axes  (Figure 5). The model passed  the vibrations , but the limited release after the test was not fully  successful .  Figure 5. Configuration of LP -PWI on the shaker  

It is worth mentioning that the qualification test campaign was preceded by a test campaign of the  engineering model  in Phase B . The design of the HDRM had not been modified significantly as compared  to the  previous phase. The Breadboard  (Engineering)  Model (BBM)  pass ed over 50 deployments and  random vibrations  at acceptance level  (-3 dB from qualification level) . The HDRM  was considered well  tested on BBM level and the failure on QM came very unexpected.  

HDRM opening failure   After the vibration test on QM- 1 (in all 3 axes) followed by limited deployment test , the Base HDRM did not  open. The Clamp  got stuck on the Tube Interface after the opening of the levers ( Figure 6) - the Clamp s  were block ed. In order to release them , several finger taps onto the V-shape s pring were needed . At the  same time HDRM#2 opened without any problems.   

Figure 6. Failure to open of HDRM  

Even  before the issue with the actual deployment, t he first sign indicating that something changed in the  Base HDRM was the  modal  response recorded by  an accelerometer placed on the levers  (Figure 7). The  resonance searches in Z axis  before and after random vibration  show large differences between these  stages . These response changes could not be  linked with the behavior of the remaining subsystems of the  boom.  The f irst mode was (at 500 Hz) increased by over 200 Hz. The mode around 1700  Hz moved close  to 2000 Hz. Main mode recorded at 1200 Hz totally changed  its shape.  Generally, f or the higher frequencies   (above 800 Hz) the modal responses start to highly deviate from each other.
78 The increase and the major change of modal response s turned out to be the first symptom  that something  was blocked in the HDRM . After  the unsuccessful HDRM opening it became  clear  that t he blocked Clamp   transfers  the loads from the CFRP tubes to the levers  differently . This  caused the change in the modal  response.  

Figure 7. Resonance response of the lever in Base HDRM in Z -axis before and after  vibration of QM- 1 

Root Cause of the F ailure  

After the  failure of the Base HDRM the most critical task was to find the root cause of the problem. At the  beginning , blocking of the Tension Clamp  seemed very unlikely . The  design  of the HDRM was well known  and had been used in two others  flight instruments . However, a deeper  analysis revealed 4 potential  root  causes of this failure. All of them affect each other , therefore it  is impossible to clearly identify which problem  was the most critical  for the system.  

Preload change during vibration   As mentioned above,  the QM-1 was not the first model that was teste d. The qualification phase was  proceeded by the breadboard (engineering model)  testing phase. The HDRMs in BBM  and in QM- 1 were  practically identical in terms of the Preload Jaw design. The B BM passed similar  tests during the  breadboard test campaign. However, the BBM was subjected to random vibrations with a level reduced to  -3 dB (from qualification level ) due to the limitations  of the available shaker. Hence,  the BBM  had never  seen full qualification  loads before the QM vibe test.  

The preload in LP -PWI was between 160 N and 170 N. This value was established  by scaling  the preload   form the MUPUS instrument. The vibration load s in the Preload Jaw  were estimated at  1120 N. It was clear  that the preload was not  high enough to avoid gapping during v ibrations. Unfortunately,  a higher preload  could  not be applied due to the strength limitation of  the HDRM structure.  Strengthening  the structure would  affect the mass budget and therefore could not be applied . There was also  a strong conviction that if the  BBM survived  the vibrations , then the QM-1 should also pass the test. However, t he qualification  level of   random vibrations appeared to be critical. The gapping between the Tension Clamp  and the Tube Interface  caused micro movement of the Clamps (Figure 8), which led  to a higher tension of the V -shape spring and  leveraged the preload.
79 Figure 8. Position change of Tension Clamp  on Vespel overlay  

Stacking in the softer material   The design of the HDRM  assumed the protruding of the Tube I nterface from the Tension Clamp . The edges   of the harder Clamp touch the relatively  soft interface surface. In case of gapp ing during the vibration,  the  Tension Clamp moved  and start ed to press one edge into the Vespel SP1 surface (Figure 8). This led to  an uncontrolled increase of contact pressure between the parts . As a consequence , the Clamp was blocked   on the Vespel SP1 overlay .  

Mechanical analysis –  low margin.   The mechanical analysis of the HDRM’s  performance is  relatively complex . The behavior of the Preload  Jaw is very hard to model. The V -shape spring is bent  and preloaded at the same time. At first , the  motorization margin was not calculated for the Prel oad Jaw , due to the Clamp rotation theory  described  above:  the preload force pulls the  Tension Clamp  at a certain  distance from the contact area with the Tube   Interface.  This creates  the torque that rotates  the Clamp. The rotated Clamp loses  contact with the T ube  Interface,  hence the friction between them can be neglected. The mechanism was considered safe from  being blocked. The previous experience from BBM (over 50 successful deployment s) seemed  to confirm  this theory . The movement of the Clamp s during the opening is very dynamic and it was very difficult  to  observe the behavior of the mechanism in real time.   

Finally, the QM -1 vibration tests showed that the Preload Jaw  can be blocked on the T ube Interface. This  situation was later repeated  manually in a lab. The failure of the HDRM opening sparked a reconsideration  of the mathematical model and the behavior of the V-shape springs. The approac h was changed -  if a more  accurate model cannot be built then the most conservative one  should be used. The new mathematical  model excludes  the rotation of the Clamp during the opening. Now the Clamp slides from the interface and  maintains  the contact with the T ube Interface.  

The Preload Jaw is  fully symmetrical, so the model is presented only for one Clamp. Figure 9 presents the  forces acting on the Clamp after the release of the levers. All forces ’ values are presented for  one Clamp .
80   Figure 9. Force relation in Tension Clamp   

The Clamp  slides  down from the interface driven by an opening force. The opening force is a summary of  the force from V -shape spring’s  bending  (FB) and the sliding force (S). The sliding force is a result of the  initial preload  (P) and its change introduced by residual force after vibrations  (Fv). The contact force (Q) is  also a result of the residual vibration force and the preload force. The contact force is additionally reduced  by a part of  the bending force (F B’). The friction force (F F) simply  result s from the contact force lowered by  a part of the bending force (F B’’) and multipli ed by the friction coefficient (μ) between the Clamp  and Tube  Interface. The success criteria for the HDRM opening was presented in Eq.1. The opening force must be  higher than the friction force.  The Eq.  2 presents the dependences between the forces and the angle of the  Clamp (α). 

(𝑄𝑄−𝐹𝐹𝐵𝐵′′)𝜇𝜇< 𝐹𝐹𝐵𝐵′+𝑆𝑆     Equation 1  

((𝑃𝑃+𝐹𝐹𝑉𝑉)𝑐𝑐𝑐𝑐𝑐𝑐𝑐𝑐−𝐹𝐹𝐵𝐵𝑐𝑐𝑠𝑠𝑠𝑠𝑐𝑐)𝜇𝜇< 𝐹𝐹𝐵𝐵𝑐𝑐𝑐𝑐𝑐𝑐𝑐𝑐 +(𝑃𝑃+Fv)𝑐𝑐𝑠𝑠𝑠𝑠𝑐𝑐   Equation 2  

The QM design was confronted with the new mathematical model . The calculation showed that the QM had  a barely positive motorization.  The nominal working point of the HDRM was located below but still very  close to the motorization line (Figure 10). The change of the preload combined with a slight increase of the   friction coefficient  could block the HDRM,  presented by the movement of the working point into the red area  on the plot . The conclusion was that t he margin of safety  for the HDRM working point  was too low to  be  accept ed. 

The friction coefficient can increase when the contact surface is worn out or when the parts were not  properly finished. There is no possibility to check the friction coefficient on the unit itself. The contact force  can be locally increased by a wedging of the Clamp into the interface. Both effects lead to the increase of  the friction force and blocking of the mechanism.
81   Figure 10. Motorization of Preload Jaw – QM-1 (α=10 deg; FB=6N; μ=0.22 ) 

Preload application   Another  cause of the preload change  following the vibration  test could be a wrong method of its application.    The l ower preload cause s a higher  gaping  than it was expected.  Generally,  the preload is adjusted during  the integration process by  the loading and movement of  the tension tube (Figure 4). In BBM the preload  was applied in a vertically positioned boom.  The Tension Tube was loaded with an adjusted mass.  In QM- 1  the method was modified. The preload was applied through the force meter with a stiff hook attached to the  tension tube.  The HDRMs walls were constrained and could not move due to the horizontal positioning of  the boom. In this configuration,  the wedging of the tube between the walls was highly possible  (Figure 11).  The force meter indicated the correct value of the preload,  but this force was not necessary  fully transferred  to the  Preload Jaw . It could have been consumed by  the block ing of  Tension Tube between  the HDRM   walls. This way of the preload adjusting did not allow for the full control of the preload in HDR M. As a   consequence,  the final pr eload could have been randomly lower, which  led to a higher gapping during  the  vibration test. The final position of the tube as well as its  tilting were not measured. Additionally, t he tilted  tube tighten ed the Preload Jaw asymmetrically . The asymmetrical position of the V -shape spring causes   higher pressure of one side of the Clamps , which increases the tendency  of pressing the edge into the soft  Vespel SP1  overlay . The modification between the BBM and the QM-1, which  was intended to be an  improvement , introduced an additional uncertainty to t he mechanism.