ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
By reviewing the lessons learned from these failures we can create more robust designs that significantly  reduce risk and are tolerant of severe conditions: high watt density, over-voltage, complex installation  and unknown stress. These lessons are consolidated in the design and testing recommendations section.  Failure History  Requirements for survival or operating heaters were determined after much of the hardware design  process, leaving this critical system as an afterthought. Flexible heaters have a long flight heritage within  the US space program. In the ISS program to date we have seen at least three failures of flexible heater  systems on-orbit: Auxiliary Power Converter Unit discharge resistor (not used for heating purposes),  Nitrogen Tank Assembly (NTA) survival heaters, and the Segment-to-Segment Attach System (SSAS)  Capture Latch Assembly survival heaters. We have also seen one failure in on ground testing, the Flex  Hose Rotary Coupler Flight Support Equipment survival heaters. The Nitrogen Tank Assembly and  Segment-to-Segment Attach System Capture Latch Assembly cases are presented in this paper.  Nitroaen Tank Assemblv [NTA)  The NTA uses two separate heater patches with layered elements. Each layer within the patches was  wired in series to form two circuits that spanned both patches (Figure 2). As a result, failing both layers in  a single patch would completely fail both strings of heaters.  String 1 Pwr  String 2 Pwr P 1- I 31 W “““I 15 W I ^A/?^^^ I I String 2 Rtn  String 1 Rtn 4-----  4  Figure 2. NTA Heater Strings  On January 17, 2003 (GMT 2003-017:16:00) both heater strings failed. The NTA heater assembly cannot  be inspected until it is replaced with a spare and returned for analysis. After the failure of the NTA  heaters, all similar ISS heaters were identified for further analysis.  SusDect Condition Action Notice [SCAN) 044  After the failure of the NTA heaters, a Suspect Condition Action Notice (SCAN) was issued across the  entire ISS program to check for similar issues. Many heaters were identified by the search criteria:  1. Multiple elements in a single patch.  2. Greater than 0.46 W/cm2 (3.0 W/in2) watt density.  Once identified, the true redundancy of the elements, software control, and possible pimpacts were  evaluated. The majority of station heaters have watt densities under 0.46 W/cm (3.0 W/in2);  unfortunately, several systems contained multiple elements in a single patch.  The critical impact of the suspect condition to ISS operations was mitigating the hazard presented by  activating both elements in a layered patch simultaneously. Simultaneous activation effectively doubles  the watt density of a heater, greatly increasing the risk of failure. Suspect software controlled heater  activation set-points were offset to reduce the possibility of this type of failure. Set-point adjustment was  not possible for thermostatically controlled heaters on orbit, and operational procedures were modified to  minimize this risk during power up.  128
PI -P3 Seament-to-Seament Attach Svstem (SSAS) Capture Latch  On May 25, 2004 (GMT 2004-146:20:01), both heaters on the Port side, segment 1-3 SSAS capture  latch failed. This heater was a single patch with dual element, layered design. Figure 3 gives a general  view of the capture latch mechanism. The heater patch wraps around the gearbox at the base and would  not be readily visible in this view. Telemetry was available at the time of failure and automated warning  alarms were activated when cold limits were approached.  Figure 3. ISS P1 Truss SSAS Capture Latch  Failure of both elements was simultaneous. The last short peak in both traces in Figure 4 represents the  failure point. The temperature of the system began to decay, and system temperatures did not begin  129
recovery until the Integrated Motor Controller Assemblies (IMCAs) were activated to provide self heating.  The IMCAs are rectangular motor assemblies that bolt to the gearbox at the base of Figure 3. They will  actuate the capture latch to berth the P3/P4 truss. Until the P3/P4 truss arrives on flight 12A, the IMCAs  will run continuously to mitigate this failure.  At the time of failure only one of the two layers was powered; however, software changes to avoid  simultaneous activation of both strings may have helped to prevent this failure. The most likely cause of  the failure is heater patch burn-through. This heater assembly is permanently installed to structure and  cannot be returned for failure investigation.  -33  -35  -37  f - d  f -39  Figure 4. Failure Temperature Traces  Proaressive Failure in a Thermofoil Heater  Progressive failure or burn-through is the result of overheating within an element. Many things cause  overheating: over-voltage, necking in the element trace, bubbles under the element, incomplete or  improper installation, or delamination within the patch. Once the overheating begins, the stages of failure  typically follow these steps:  0  0  0  0 Temperatures increase within the element, increasing resistance.  Melting temperature for the Teflon is reached and the patch delaminates. In this process the  Kapton film may char.  Once delamination occurs, there is no direct conductive path for heat transfer and radiation  becomes the dominant mode of heat transfer.  Unable to release the heat, the heater element foil continues to increase in temperature until it  melts or fractures and loses electrical continuity. At this point the heater has failed irrecoverably.  130
Conclusions  Flexible thermofoil heaters have failed several times in the ISS program. We must look at and seriously  consider the processes going into design, testing, and installation of heater patches on spacecraft.  Suitability and capability of the designs to fulfill their functions have generally been understood and  analyzed in depth; however, survivability has not always been treated with sufficient rigor. Qualification  and acceptance testing must include steps to verify heater survival.  Overestimation of the ability of the system to survive high voltages, extended use, or inadequate  installation undermines all design effort spent on the system. If the heater itself does not survive, failure of  the thermal system jeopardizes mission objectives. If no resolution can be found, the system could be a  total loss. Mitigating the impacts of these failures consumed substantial time and resources within the ISS  program. The most critical lesson learned is to do everything possible to maintain the survivability of the  mechanism via heater patch design by incorporating knowledge from previous failures during initial  design.  References  1. Bolton, Victor J “SCAN 044 Response for Structures and Mechanisms,” Boeing memo A92-J383-  STN-M-VJB-04-032, February 2004.  2. Bares, Geoffrey “PRACA 3661 : P1 Nitrogen Tank Assy (NTA) Heater Failure,” Boeing  presentation, February, 2003.  132
The BAPS concept provides a bearing cartridge structure that is of appropriate stiffness and load capacity  to enable the maximum loads (combination of preload and externally applied launch loads) for the  particular bearing size to be applied and a factor of approximately 10 highlow preload ratio. This ratio  permits at least a halving of Hertzian contact stress, with commensurate bearing torque and life benefits.  1 .E+11  Duroid/PGM-HT  lubricating Cages  MoS2 (Only on races) \  ‘h E ‘E l.E+07  Q, !e -I b 1 .E+06  1 .E+05  1 .E+04 I I I I I I I h SI \  31 \\  0 500 1000 1500 2000 2500 3000 3500 4000  Peak Hertzian Contact Stress (MPa)  Figure 1. Lifetime v Contact Stress for Bearings Lubricated by DuroidPGM-HT or MoS2  BAPS Requirements  The BAPS device requirements were determined via a comprehensive Market Survey of current or near  term applications in which more than 34 space companies were contacted. Responses to the  questionnaire indicated that over 20 applications could benefit from BAPS technology, for example  deployment hinges, steering devices, APM’s, various instruments, actuators, reactionhomentum wheels,  CMG’s, solar array drives and scan mechanisms for microwave sounding and radiometry. Design drivers  from the Survey were ranked in order or priority as Reliability, Envelope, Mass and Cost.  The Survey suggested most potential users would require a bi-stable BAPS device capable of switching  from a single high-preload (State 1) to a single low-preload (State 2), however a significant minority of  respondents also foresaw applications where the ability to continually vary preload could be valuable. A  wide range of bearing sizes was identified, from 25-190 mm bore, but the most popular size range was  50-100 mm bore. The number of operations required of a bi-stable BAPS (assumed to exclude  qualification margins) was enveloped as e20 on ground and 1 in-flight. For a variable type device, up to 2  million operations were envisaged.  Envelope constraints were also deduced from the Survey, together with thermal requirements. Excluding  the relatively infrequent cryogenic or high-temperature applications, the majority of applications were  enveloped by an operational temperature range of -40°C to +80”C.  BAPS Concept Derivation and Description  By combining the results of the Market Survey and the experience of the team, a generic requirements  specification was generated together with a number of concepts potentially capable of meeting these  requirements. The concepts were subjected to an initial filtering using “killer criteria” (Le., criteria which all  viable concepts MUST meet) after which surviving concepts were then subjected to a formal trade-off  against criteria again reflecting the aspirationskoncerns expressed in the Market Survey returns ranked  using a rigorous paired comparison method.  134
As part of the trade-off process an assessment was made of the flexibility of each candidate design for  adoption with 4 candidate design cases ranging from small (25-mm bore IS0 10 section, e.g., SNFA EX25  with State 1 preload of 200 N), to medium (50-100 mm bore, thin section) and large (>150 mm bore thin  section, e.g., Kaydon KA075AR0, with State 1 preload of 3000 N) bearing envelopes. From these design  cases, a single so-called “enveloping design case” was established as a most-challenging case for the  Breadboard Model (BBM). The enveloping case was based on the thin-section SNFA SEA55 (7CE3)  bearing, having a preload ratio of 3000 N (State 1) / 200 N (State 2) generating a peak bearing Hertzian  contact stress ratio of approximately 2000 MPa / 850 MPa.  In the trade-off winning concept, which is based on the monolithic titanium structure shown in Figure 2  below, the bearings (not shown) are located by upper (floating) and lower (interfacing) thrust rings. These  rings are linked via a series of “suspension beams” which provide the required structural stiffness  characteristics for launch. By monitoring the strains induced in these features due to changing relative  axial displacements of the two rings the bearing preload can be determined.  Rotation of a central “synchro-ring” by a few degrees adjusts the spacing of the upper and lower rings by  some microns and so modifies bearing preload. The ‘upper and lower rings remain parallel despite the  rotation of the synchro-ring because they are linked through a series of flex-struts to the upper and lower  rings preventing misalignment. These features also provide an over-center toggle action such that the  preload is stable un-powered in two synchro-ring positions. For the high preload state the central synchro-  ring is retained just away from top-dead-center (TDC) in a stable configuration (in contact with an end-  stop). To set the bearings to low preload, a “Self-Release” force is applied to displace the synchro-ring  such that the flex-struts are pushed over TDC and to a stable low-preload position used for flight. This  position is maintained by application of a “Holding Force”. For the BBM, the high (3000 N) preload state is  at a synchro-ring position of +lo0 prn, and low (200 N) preload at -600 pm as shown in Figure 3. A  maximum “Reset Force”, around 100 N on the BBM must be applied.  Figure 2. Monolithic Titanium Alloy Structure Concept and Hardware  (includes liners to interface with bearings)  135
Holding Force Displacement  Curve  1 / 50 - ielf Release Forc  v- I  -200 400 10 E  Drive Force  Curve High Preload  Posit ion -1 00 Low Preload  Position Reset Force  - IdU  Synchro-ring Displacement (micron)  Figure 3. Predicted Drive Force and Thrust Ring Displacement Curves for SEA55 BAPS  Practical Experience with Hardware  Highlights of the performance of the breadboard model (BBM) BAPS, which contains bearings lubricated  by a PFPE-based grease thickened by PTFE particles are provided below.  Dimensional inspection showed the BBM to have been manufactured to tolerances in-line with  requirements, especially when considering bearing seating tolerances. No significant misalignment of the  bearings on assembly or during BAPS operation was measurable. On assembly the BBM structure  introduced 400 pad of angular misalignment and when operated at most a further 20 prad. These values  are well within even the permissible groove-face misalignment for the bearings themselves (e.g., for  ABEC 9 grade bearings of this size, 200 pad PER BEARING is allowed). This figure is also much less  than the known face-face target misalignment for good performance of solid-lubricated bearings, which is  around 300 prad.  As expected, because the BAPS is a purely elastic system, the strain gauges used to monitor  displacement of the suspension beams were found to be very linear (sensitivity typically 20 @pn of BAPS  total axial displacement) and repeatable (within 1-2 p when BAPS is repeatedly switched from high- to  low-preload), demonstrating accurate, repeatable preload switching even after vibration.  The target preload ratio of 3000 N (State 1) / 200 N (State 2) which would have provided bearing mean  torques at room temperature of 1400 gcm and 110 gcrn respectively was not quite reached due to the  effects of elastic strain in the bearing seats. The achieved preload values were 2500 N and 300 N for  which the measured torques were around 1100 gcrn and 200 gcrn (Figure 4).  However a positive effect of this bearing seat elasticity was that because strain energy is stored in the  BAPS, bearings AND the bearing seats, on switching from high to low preload much less force is needed  to hold the device in against the low-preload end-stop position than is predicted based on a model which  excludes bearing seat effects, as can be seen in Figure 4.  136
mnn 1400  LOW PRELOAD  Measured Drive Force Predicted Drive Force  L 4 rn 0  Synchro-ring Position (microns)  Figure 4. Measured torque and Drive Force Performance of BBM BAPS Unit  Actuation Options  The BAPS can be actuated by several means depending on the application. Most commonly a direct linear  actuation will be used for which, all that is required is to provide up to around 60-70 N force and around  700 pn of total linear displacement. For the majority of applications where a typical operational  temperature range of -40 to +80°C is demanded, then actuation will be via a Shape Memory Bender or  High Output Paraffin Actuator. Both devices have the advantage of reset-ability and ease of redundant  implementation. However, the BAPS can also be entirely passively actuated, for example by exploiting  differential thermal expansion for applications that have high or low temperature requirements, or indeed  implemented in a fully-variable form. In the latter case, use of a rotary piezomotor is an attractive option.  How Designs Could Change Post-BAPS  A number of historical and more recent applications are reviewed (and summarized in Table 1). For each  we identify some of the main potential design changes and resulting engineering benefits of use of a  BAPS device in similar future applications.  Optical Pointincl and Trackina (ea. GOMOS SFA on ENVISAT)  The GOMOS Steering Front Assembly (SFA), currently flying on ENVISAT, is an optical acquisition and  tracking mechanism for stars. The mechanism includes a coarse pointing azimuth stage that provides a  wide pointing range by allowing a mirror and turntable to rotate. The bearing configuration would ideally  support the mirror and turntable during launch. However, given the need for low torque noise emission  and reliable operation the bearings required a soft, low preload, and a relatively small outside diameter.  However, high strength and stiffness were required during launch, and these conflicting demands could  only be met by use of a massive and complex motorized launch protection system (bearing offload device  or BOLD).  Using BAPS, a more elegant design using large, thin-section bearings could have been adopted, capable  of meeting the launch design case, yet also providing low and soft bearing preload on-orbit. Preliminary  analysis suggests that a pair of back-to-back thin-section bearings having a bore of 200 mm could have  met the launch load and stiffness requirements. This approach would have yielded very significant savings  in mass, cost and complexity.  137
length. The stroke capability of the new D-strut@ isolator damper assembly is in excess of * 8.13 mm  (* 0.32 in), as required. The original D-Strum Isolator is shown in Figure 1 and the new patent pending D-  Strut@ Isolator is shown in Figure 2. It needs to be noted that the mounted lengths for both of the isolators  are the same, 21.03 cm (8.28 inches).  c 1. Original D-Stru  The original D-Strut@ lsoktor shown above was used in a previous program to provide launch and on orbit  isolation for an instrument which has already flown. The environmental loads for a new payload  instrument exceed the capabilities of the existing D-Strut@ isolator.  2-  I I I I I I1  FigL. - _. . __ __ - __. ___ .__.___. . _. _____. .._.  The new D-Strut0 Isolator shown will be used to provide launch and on orbit isolation for the Laser  Communication Terminal Experiment which is under development. The launch loads for the Laser  Communication Terminal required the design changes. There is a patent pending on this new design.  Desiqn Analvsis  The loads from the derived requirements in Table 1 were used to determine the stresses in the  components that make up the new D-strut@ isolator. The analysis of the sealing bellows and  compensation bellows was performed by the bellows supplier using proprietary methods. The flexure, KB  spring, and main spring analyses were performed using finite element analysis modeling methods with  l-DEAS@ application software. There was a requirement that the factors of safety be 1.1 to the material  yield limit and 1.25 to the material ultimate limit. We calculated the loads, predicted the stresses, and  calculated the factors of safeties for all the critical structural parts in the new D-Strut@ Isolator. The limiting  part in the new D-Strut@ Isolator, from this set of analysis that needs further evaluation is the KB spring  with a minimum factor of safety to the yield strength is 1.14 and the minimum factor of safety to the  ultimate strength is 1.22. Once detailed requirements are defined, the loads in the parts will be further  analyzed.  Isolator Bipod Test  The new D-strut' isolator bipod testing was performed in August, 2005 with the isolators mounted to an  18.1 -kg (40-lb) mass. This configuration proved that the isolators were capable of meeting the extreme  load case defined for the launch quasi-static environment of 16.25 g's. The testing was performed up to  2000 Hz. The configuration is shown in Figure 3. The suspended mass was 18.1 kg (40 Ib) so that the  bipod first mode would align with the predicted system hexapod bounce mode of approximately 40 Hz.  We reconfigured the shock test setup to test with a Shock Beam methods using both the original 31.8-kg  (70-lb) mass and the new 18.1 -kg (40-lb) mass. We attached the bipod (with the masses suspended) to a  shock beam and provided the shock levels by impacting the beam with a weight on a pendulum. The  shock beam test setup is shown in Figure 4.  143
18.1-kg (40-lb) pay-  load mass shown -  Figure 3. Bipod Vibration Configuration  Electro-Dynamic shaker Input is at Bipod  Base on the left with the payload mass  placed in-line to the input. This configuration  minimized the suspended mass modes in  the measured response. Figure 4. Shock Beam Test Fixture Set-up.  This set-up was used to provide a shock to the  base of the Bipod using the impact mass  suspended on the chain. The impact mass is  allowed to swing and impact the shock beam to  produce the shock required at the bipod base on  the far end of the shock beam.  Shock Beam Test  When performing the shock beam test, we used an 18.1 -kg (40-lb) mass and a 31.8-kg (70-lb) mass. The  test setup is shown in Figure 4. Figure 4 shows the actual setup of the shock beam. There was an array  of accelerometers placed on the new D-strum isolator bipod and fixtures to measure the shock input and  the response of the isolators and payload mass. Only the in-axis input and in-axis response of the  payload mass will be discussed in this paper. The set up was the same for both of the masses. The test  results for each of the payload masses were much the same; therefore, only the results from the 18.1 -kg  (40-lb) payload mass test will be presented.  The locations of the accelerometers were selected to provide information on the isolator spring body axial  surge modes, isolator spring body lateral modes, and the mass response. The locations of the  accelerometers are shown in Figures 5. In order to ensure that the shock input to and response of the  payload mass were captured, we used shock accelerometers at the channel 1 location, and the channel 8  location.  Base input  Figure a. anocK deam Test Accelerometer Lucations  From this view, five of the accelerometers can be clearly seen. The accelerometers of interest which  indicate how each of the isolator body modes affect the payload mass are channel 1 and channel 8.  144 Channel 1 8 BiDod