ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
231  

Figure 6. Image of spring mechanism during thermal testing with  evidence of  hardware failure circled      After the failure, the qualification spring mechanism was disassembled. The r oot cause was identified to be  invalid analytical  model simplification. The analytical model simplified the torsion spring geometry as a  cylinder with uniform inner diameter. The real rectangular cross section torsion springs had slight variations  in the i nner diameter between coils, with sharp cutting edges presented to the Teflon bushings during  cycling. Therefore, the contact stress in the real hardware at the cutting edge was substantially higher than  in the idealized analytical model.  In addition, the sleeve bearing had been designed with a helical cut along  the axis of the bearing, designed to allow radial compliance as the torsion spring inner diameter changes  during winding/unwinding operations . However, the helical cut also drastically reduced axial  stiffness of the   part. As such, when the spring coils moved axially during winding, the edges of the bushing began to contact  each other and plastically deform.  These failures can be seen in F igure 7. 

This failure resulted in the opening of a JPL Problem Failure Report. As such, a technical team was  assembled to oversee the investigation and resolution of the failure. Because of the multi -mission  applicability of the hardware design, the team was compos ed of representatives from both the SWOT and  NISAR projects. Any resulting actions needed to be approved by both missions.  There was programmatic  motivation to utilize as much of the existing hardware as possible.  

Figure 7. Images of Teflon  sleeve bear ings after thermal testing  

Mechanisms Modification   Ultimately, a solution was developed that replaced the Teflon  sleeve bearing with a grease- plated 440C  stainless steel sleeve bearing with modified geometry. Over 400 functional cycles and 20 disassembly  procedures  of the qualification mechanisms were carried out during the hardware failure. Table 1  summarizes the test campaign that was conducted to find a new sleeve bearing design.  The fundamental  approach was  to change one parameter at a time from the original bushing design and evaluate its effect  on the health of the component and test performance until an  acceptable solution was found.  An ac ceptable  solution had to simultaneously meet mechanism  torque performance needs  as well as avoid significant  FOD generation or damage to the bushing through three times the planned number of flight unit life cycles .
232 Table 1: Summary of hardware failure investigation testing  

Index  Material  Bearing  Design  Test Type  Test Result  Notes   1 Teflon   Helical  Cut Vibrational, Thermal, Cycle  Life Fail -Initial Failure   2 6061  Torque Characterization  Fail -Torque requirement failure   3 304C  Torque Characterization  Fail - Noise from mechanism   -Torque requirement failure  - Helical cut deemed unacceptable   4 Bronze  Solid  Sleeve  Bearing  Torque Characterization  Fail - FOD found   5 440C  Torque Characterization  Fail - FOD Found   6 Copper   Extended,  solid  sleeve  bearing Torque Characterization  Fail - FOD found   - Torque requirement failure   7 440C  Torque Characterization, Cold  Pass - Noise witnessed   - Good Torque   8 Bronze  Torque Characterization  Fail - FOD found   - Good torque   9 440C  Vibe, Cold, Hot, Cycle Life  Pass - Full Life Test   - Good Torque  - Noise witnessed   10 440C  Torque Characterization  Pass  - Confirm lubricant  alleviates noise  

As described previously, torque t esting of the mechanism included using a  transducer and encoder to  measure torque and rotational position, respectively. This torque testing was carried out for each potential  bushing design. If the torque was deemed acceptable, the unit was then disassembled and inspected for  any FOD or other potential failures. Figure 8 displays the torque performance of the spring mechanism for  the final bushing design in ambient  and cold  conditions  during cycle life testing  (defined as at least three  times the expected number of mechanism cycles) . Torque performance is seen to degrade up to 1.1 N -m  (10 in- lb) over the course of 30 cycles at ambient . Further, torque performance degrades at cold  temperature about 1.1 N -m (10 in- lb). Despite  performance degradation , torque never violated the 3.2 N -m  (28 in-lb) torque requirement . Also notable,  the unwinding torque at cold temperature is seen to be nearly  constant. This differs from the analytical model of linearly decreasing torque. The cause for near -constant  torque is suspected to be internal  mechanism friction caused by  migration and degradation of lubricant on  the bushing as it is cycled . 

Toward the end of the Problem Failure Report investigation, during final life cycle testing of the hardware,  audible sound was observed from the hardware. This sound triggered further investigation and res ulted in  the development of an assembly -level relubrication process for the mechanisms. This procedure eliminated  the s ource of the concerning noise.  The relubrication process seeks to augment lubrication in areas on the  sleeve bearing where lubricant may have worn aw ay during mechanical cycling.
233  

Figure 8. Torque profile of mechanism as it is wound and unwound at ambient and cold temperature. Blue  indicates initial cycles, yellow  intermediate cycles, and red represents later cycles.  

Conclusion  

Following resolution of the hardware failure, all flight spring assemblies  have been updated, passed flight  acceptance environmental testing, and  have been integrated into both SWOT and NISAR flight masts . The  mechanisms have successfully been tested at higher levels of assembly  and performance is consistent .  The spring mechanisms require re lubricat ion every 8 cycles of ground testing  based on the process  developed in the hardware failure investigation, wh ich is achievable at the integrated level of assembly .  

Key l essons learned from the development of these torsion springs :  • Avoid rectangular cross section springs unless volume limitations necessitate their use. Round wire  springs have greater geometric a nd performance consistency and are simpler to analytically model.   • Beware of analytical model simplifications that may oversimplify  and invalidate the results.   • Rectangular cross section springs will twist about the axis of the wire when wound. This twisting  needs to be considered when designing any hardware  coming in contact  with the spring.   • Consider both the wound and unwound geometry of the spring during design of the mechanism.   • There is  high value in a complete mechanical cycle life test at the prototype stage  of development .  

References  

1. Paul Lytal  and Marcel Renson. “Spacecraft Common Deployable Boom Hinge Deploy and Latching  Mechanisms.” Proceedings of the 44th Aerospace Mechanism Symposium, (May 2017), 403- 416.     Acknowledgement  

The research described above was carried out at the Jet Propulsion Laboratory, California Institute of  Technology under a contract with the National Aeronautics and Space Administration. © 2020. California  Institute of Technology. Government sponsorship acknowledged.
235 NSI Performance Improvement through the use of Automation  

Jason Kozmic*, Bill Gratzl* and Hobin Lee * 

Abstract  

Chemring Energetic Devices has transferred the manufacture of the NASA standard initiator ( NSI) from its  Torrance, CA facility to its Downers Grove, IL facility  and performed re- qualification of the manufacturing  process. As part of the transfer effort, the use of automation has been introduced into the process to  eliminate operations historically performed by hand which has led to reduced cryogenic function time  variability, increased manufacturing throughput, and safer handling of the energetics.  

Introduction  

The NASA Standard Initiator (NSI)  along with Chemring’s commercial equival ent (PC -23) is a two (2) pin  electrically activated, hot -wire, electro- explosive device which provides a source of pyrotechnic energy used  to initiate a variety of space mechanisms for use on both satellite and launch vehicle applications.  Mechanisms include pyrotechnic valves, separation nuts/ bolts, cable/ bolt cutters, pin pullers and many  others.  The reliable initiation of these one- time use mechanisms is often mission critical leading , to stringent  test requirements being levied upon the manufacture and acceptance of the NSI . 

Electrically activated initiators are conceptually simple devices but  possess  manufacturing  sensitivities that  can have significant effects on the final performance.  A typical electric initiator is depicted in Figure 1 and  consists of a glass to metal sealed header (with receptacle), a bridgewire welded across the header pins,  energetic ignition mix (ZPP typical) consolidated onto the bridgewire, and a welded closure output.  When  an electrical stimulus is applied to the head er pins (5 a mps, 20 ms typical) the current heats the thin  bridgewire which in turn heats the consolidated ignition mix.  Once the ignition mix reaches it s auto- ignition  temperature the energetics undergo a self -sustaining reaction which produces heat, gas,  and hot particles.   These thermal outputs are used to ignite secondary energetics in an energetic train/cartridge or can be  used perform work in a device without any additional booster . 

Figure 1.  Cross Section of NASA Standard  Initiator  (NSI)  

* Chemring Energetic Devices, Downers Grove, IL;   jkozmic@ced.us.com  

Proceedings of the 45th Aerospace Mechanisms Symposium, NASA Johnson  Space Center, 20 20  Glass to Metal Seal w/ Receptacle   Electrical Bridgewire   Energetic Mix (ZPP)   Welded Closure
236 Chemring Energetic Devices  (CED)  is currently NASA ’s only  qualified  manufacturer of the NSI . The all -fire/  no-fire stimulus, function time (application of power to first pressure) , and peak pressure are all critical  performance requirements of an electrical initiator and the NSI requires stringent control of these  performance characteristics.  All can be sensitive to the manufacturing process  with the cryogenic (22 K )  firing time , unique to the NSI , being particularly susceptible.  

Process Automation  

Historically , the NSI had been produced out of CED’s recently closed Hi -Shear facility located in Torrance,  CA. The manufacturing process involved performing critical operations by hand with consistency limited by  the manufacturing technician.  These critical operations include blending of the energetic material batches,  welding the bridgewire,  applying wet energetic mix (slurry ) onto the bridgewire , and loading and  consolidating powder zirconium potassium perchlorate ( ZPP) into the output charge cup.  Performing these  critical tasks by hand has limited manufacturing throughput and product consistency, resulting in cryogenic  firing performance with variability that put the acceptance of an NSI lot at risk.  

With the transfer of NSI manufacturing to CED ’s Downers Grove, IL facility , the use of automation has been  introduced into the manufacture of the NSI leading to less variability in cryogenic firing performance and   increased manufacturing bandwidth.  To date, three  critical aspects of the manufacturing process have been  automated at CED;  bridgewire welding,  application of slurry to the bridgewire, and mix and blend of the  energetics (Z PP). A photograph of CED ’s automated bridgewire welder  is provided in Figure  2. 

Figure 2.  Automated Bridgewire Welder   The automated bridgewire welder  reduces operat or involvement in the process of  loading the prepared  glass to metal seal headers into the machine and starting the process . The automatic welder spools the   <0.005-in (0.13- mm) diameter bridgewire onto the header pins then positions the weld electrodes using a  vision system.  Once positi oned , the weld force is contro lled with feedback control and once welded, the  system severs/discard s the remaining wire pigtails.  Using the automated bridgewire welder has  shown to  consistently  produce welds meeting the 0.95 – 1.15 ohms bridg ewire resistance requirements with an  average of 1.08 ohms and a standard deviation of 0.018 ohm across four lots as shown in Figure 3.  In
237 contrast, the previous hand welding operation involved a technician cutting the wire to length and  positioning the wire for  welding and securing in place by hand using a hold- down fixture.  The extra touch  time introduces the potential  to damage the thin bridgewire ( nicks, dings, etc.) which can affect bridge  resistance and create hot spots that can cause no- fire failures when 1A/1W is applied for 5 minutes.   Additionally,  this process can result in unit to unit variability in the configuration of the welded bridgewire,  not present in automated process, which can influence how energetics interface with the bridgewire and  ultimately affect cryogenic function times (see Figure 5 ). 

While the end quality of the hand bridgewire weld ( resistance, all- fire/ no- fire performance) is similar to that  of the automated process , the automated process significantly increases the first pass yield reducing the  need to rework or scrap in- process hardware.  

In conjunction with bridgewire welding, the interface between the bridgewire and the energetics as well as  the amount of energeti c material applied to the bridgewire have significant effects on the cryogenic function  time.  Historically the application of wetted energetics (“slurry”) directly to the bridgewire has been performed  by hand and been difficult to precisely control.  The Do wners Grove manufacturing line has implemented an  automated application process which controls the rate and duration of slurry application, tightly controlling  the amount and condition of the s lurry applied to the bridgewire.  The process has averaged 5.2 m g with   0.6 mg standard deviation across four lots as shown in Figure 3.  

   Figure 3.  Automated Process Consistency (Mean w/ Standard Deviation Error Bars )  Bridgewire welding and application of slurry to the bridgewire are critical operations controlling the all -fire/  no-fire current, output pressure performance of the initiator between –162°C  and + 149°C as well as the  function time performance at cryogenic temperatures as low as 22 K . The automated processes have   demonstrated significant improvement in the function time variability at the cryogenic  conditions.  

A comparison illustrating the effects  of process automation is presented in Figure 4.  A single blended lot of  ZPP (lot 13 -44352) was manufactured using four methods ; hand welded with a hand slurry application   using the documented Torrance method, hand welded with a hand slurry application using a Downers  Grove developed method, hand welded with automated slurry application, and both automated welding and  slurry application.  These groups were then tested at both 77 K  and 22 K. 

The data demonstrates higher average function times and wider variability when loaded with the heritage  hand operations.  By implementing either an improved hand or automated slurry application process to the 00.20.40.60.811.2 DJA DJB DJC MTWBRIDGEWIRE RESISTANCE, OHMSBridge Resistance Consistency 00.0010.0020.0030.0040.0050.006 DJA DJB DJC MTWAPPLIED SLURRY, GRAMSSlurry Consistency
238 heritage hand welding process, the average function times and the variability were shown to decrease,  which indicates the inherent variability of the hand process. These batches still produced flyers which  resulted in standard deviations similar to the hand process, which were attributable to inconsistency in the  hand welding operation.  By adding the aut omated bridgewire weld process to the automated slurry  application process, the flyers were eliminated and both average and standard deviation cryogenic function  times was significantly improved.  

Figure 4.  Effect of Process Automation on ZPP L/N 13- 44352   The ZPP lot (13- 4432) depicted in Figure 4 was deemed not suitable for loading into production NSIs during  batch acceptance testing due to function time variability which is required to have an average of l ess than  1 ms and a 3- sigma limit less than 3  ms. Historically when a blended ZPP lot failed to meet batch  acceptance testing, the ZPP lot was scrapped and a new lot blended.  The re sults presented in Figure 4,   indicate  the root cause of the cryogenic funct ion time failure was not the raw material but the process.  The  test units built with the same ZPP lot, but with the automated processes produced acceptable results,  indicating the lot to be suitable for production.  

Cryogenic function times of NSI test lots spanning 2013 through 2017, manufactured at both CED’s  Torrance facility (using hand operations) and CED’s Downers Grove facility (with automated bridgewire  welding and slurry application) is presented in Figure 5.  

Data shows th at lots loaded at CED’ s Downers Grove facility ( DG) had function times ( 0.43 ms and 0. 44 ms  mean with 0. 19 and 0. 23 ms standard deviation) which were faster and tighter than the most consistent  lot  ever fabricated at CED’s Torrance Facility (0.59 ms mean with 0. 25 ms standard deviation manufactured
239 in 2013)  and had significantly tighter performance when compared to routine production lots loaded in 2016  and 2017.  

Figure 5. Cryogenic Function Time Comparison   In addition to improved cryogenic firing performance,  the implementation of  automation has  allowed CED  to increase manufacturing throughput.  The automated bridgewire welder allows CED to bridge 100 initiator  headers in roughly an hour and a half which previously took  two days when this operation was performed  by han d. Automated slurry application has increase capacity from approximately 200 pieces per day to  200 pieces in 30 minutes  of application time.  This increased capability allows CED to reduce lead time of  the NSI and other initiators . 

Automation has also helped CED to manufacture initiators in a safer manner than previously possible at its   Torrance facility.  The advantage of ZPP as an initiating material, its  high sensitivity and rapid self -sustaining  reaction, also makes i t dangerous to blend as a raw material.  The blending of the ZPP raw materials (fine  granular zirconiu m and potassium perchlorate) had historically been done by hand at a remote  facility by  specific, highly trained technicians.  The hand blending operation previously put  the technician in close  proximity to the volatile raw ingredients  and exposes  them to a lethal amount of energetics.  Chemring is  committed to improving safety and has developed automated energetic blending capability , depicted in  Figure 6 . It ensures  no personnel are present in proximity to the raw materials during blending.  Once a  blend is completed the equipment di spenses and seals the ZPP into 7- gram velostat pucks for safe  handling.  ZPP batches blended using the automated energetic blend  equipment have passed required  differential scanning calorimetry /thermogravimetry analysis (DSC/ TGA) and heat of explosion testing  (HOE) required to accept a batch as well as meet cryogenic function time requirements required to be  suitable  for producti on NSI manufacture.  

Along wi th automating the blend process,  CED is currently working to load NSIs using a robotic manipulator.   The robotic manipulator will remotely load ZPP into the NSI charge cup, consolidate it to a feedback - controlled pressure, then confirm and log ZPP charge weights on a piece part basis.  Qualification of this 00.511.522.533.544.5 0 20 40 60 80 100 120 140Function Time (ms)-420F Function Time Results for ZPP for test units built at TR and DG TR, Dec 2016" TR May 2017" DG July 2017" TR circa 2013" TR May 2017" DG July 2017"
240 effort is set to begin in 2020 and once complete, in conjunction with the automated blend equipment will  allow CED to blend and load ZPP without personnel coming into contact w ith the raw materials.  

   Figure 6. Automated ZPP Blend Machine (left) and Robotic Manipulator (right)   By implementing key automation processes, CED has improved consistency of the already reliable NSI and  tightened performance, increased the manufacturing throughput, and markedly improved the safety by  minimizing the exposure of employees to potentially hazardous conditions.  Upcoming work with the robotic  manipulator will further increase these benefits.   

More consistent performance, shorter leads times, and reduced scrap will allow Chemring to better service  NASA by providing NSIs . In addition to the NSI, CED offers the PC -23 initiator produced on the same  manufacturing line as a commercial equivalent.  The PC -23 is currently in use on a number of space  platforms for many of the prime contractors and those customers are set to realize the benefits implemented  on the NSI.
241 Microvibrations Modelling and Measurement on Sentinel 4 UVN Calibration  Assembly using a Piezoelectric 6 Component Force Dynamometer   

Benoît Marquet*, Jean- Yves Plesseria *, Jérôme Jacobs * and Christof Sonderegger ** 

Abstract  

The Sentinel 4 (S4) Ultraviolet Visible Near -infrared (UVN) instrument is a high- resolution spectrometer  embarked on the MTG -S platform. The radiometric accuracy of the UVN instrument relies on regular inorbit re- calibrations using the C alibration Assembly (C AA). This CAA mechanism is comprised  of a multi - functional wheel containing calibrated diffusers. Positioning is via a stepper motor and,  as the instrument  is mounted on the MTG -S platform, i t is imperative that  the resultant microvibrations and torques  are  controlled as the MTG IRS instrument is incredibly sensitive to them.  

To this end, CSL has developed a Simulink model that inputs a motion profile to simulate  the real-world  mechanisms, generat ing the torques and microvibrations at the satellite s center of gravity. To validate this   prediction, a specially designed  Kistler dynamometer allows the actual forces and torques exported by the  CAA qualification model to be measured.  

Measuring with Piezoelectric Dynamometers  

Microvibrations consist of extremely small accelerations with very low intensities . Measuring them is a  very challenging task and methods for properly doing so have only become available in the past few  years. Piezoelectric force sensors and dynamometers are ideally suited for this purpose. Their incredibly  high span- to-resolution ratio of greater than 100000 is a particular advantage. Thus,  it is possible to  measure dynamic force changes down to 0.01 N, even when the object being measured weighs more  than 10 kg. The static weight can be “eliminated” by resetting the charge amplifier ( this acts like a tar e  function,  effectively re-zeroing the scale).  In addition, the very high stiffness of piezoelectric force sensors  permits natural frequencies  of 1000 Hz or more.  

To optimize the system,  the measurement setup itself is  also highly impor tant. As can be seen in Fig. 1,  the dynamometer needs to be mounted on a vibration- isolated table to prevent structure- borne sound and  external vibrations  affecting the results . These external influences can severely distort measurements as  the dynamometer , and the mass of the object being measured,  can generate a large acceleration   themselves . With a correct set -up, the interference signal can reach levels below 0.01 N or 0.003 Nm  (RMS; 3…350 Hz).  

* Centre Spatial de Liège,  Université de Liège,  Angleur, Belgium;    jjacobs@uliege.be   ** Kistler Instrumente AG, Winterthur, Switzerland ;  christof.sonderegger@kistler.com 

Proceedings of the 45th Aerospace Mechanisms Symposium, NASA Johnson Space Center, 2020