ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
386 Hinge Geometry and Design  

The geo metry of the hinge cutout in the foldable tube system was  derived to both maximize torque output  and reduce the high stress concentration areas that result from complex curvatures near the fold. By  creating a narrow slit in the boom,  the cross section effectively becomes two tape springs, which are  commonly used in deployable structures for solar array applications. At the ends of the slit, however, the  transition area must accommodate deformations to go from flattened tape springs to  a full tube cross  section. To avoid large strain areas due to the complexity of this transition area, a circular cutout was   introduced.  This cutout simply removes this difficult transition area.  

Figure 5: Foldable Hinge Geometry  

In the hinge cutout, there are several dimensions and material properties that effect the overall  performance. These dimensions include the hinge length “L”, slot width “w”, end cutout diameter “d”, tube  diameter “D”. The material properties of thickness, density, and modulus (viscoelastic and elastic  contributions) must be accounted for to fully characterize hinge performance.  

Verification of Solid -State Hinge  

Design of the solid- state hinge for the ROC -PDS required verification testing and analysis of material  allowable limits, hinge torque output, and torsional stiffness. For this, several testing methods had to be  developed by Roccor to investigate the properties and effects of interest such as; material bending limits,  hinge performance, and performance degradation due to long term holds.  

Material Testing for Thin laminate Bending   As detailed in  reference 7 the column bend test was developed to combine the attributes of both the platen  (Ref. 8) and LD -FPBT (Ref. 9) into a simple test method intended to measure the moment -curvature  relationship of a given HSC laminate. Testing is done by fixing a laminate sample into two identical “arms”  that are then pinned into clevis mounts attached to a load frame. When the load frame compresses the  fixture, the arms rotate about the pins requiring the sample to curve between the two arms. Using geometry,  the moment -curvature relationships can be found using methods and equations detailed in Ref. 7. 

From this testing we can obtain the allowable bending curvatures, and relative strains, for a given laminate.  For the ROC -PDS the laminate architecture selected contained a combination of Astroquartz and carbon  fibers cured in a thermoset epoxy resin matri x. Testing was conducted on coupons cut at four orientations  (0°, 45°, 90°, and 135°), and testing was conducted to obtain both the quasi -static load failure as well as  the long- term hold failure curvatures ( Figure 7).  

Results from ultimate testing were p lotted on a polar chart for visualization of “allowable failure”. The  average failure of each laminate orientation is utilized to create a polar curve fit “allowable failure” value to  be used in analytical failure investigations of the stowed and deploying  boom structure. Results from
388 Where the curve fit to the equations is found using E quation 1 with allowable values s hown in Table 1. 

𝐾𝐾𝜃𝜃𝑓𝑓=𝐾𝐾𝜃𝜃𝑓𝑓cos(𝜃𝜃)2+𝐾𝐾90𝑓𝑓cos(𝜃𝜃)2+�4𝐾𝐾45𝑓𝑓-2(𝐾𝐾0𝑓𝑓+𝐾𝐾90𝑓𝑓)�[sin(𝜃𝜃)cos (𝜃𝜃)]2                       ( 1)  Table 1 Instantaneous failure curvature results   Orientation   (degrees)  Curvature   (rad/m)  (rad/m) (rad/m)  0 329.70*  0.73 N/A  90 328.82*  0.01 N/A  -45 275.2  8.03 13.54   45 306.44  15.13  25.5  *Maximum curvature observed due to column bend test fixture limitations. No  physical failure observed.  

Hinge Element Failure Evaluation   The folded boom structure was evaluated for failure simply by taking  the element level curvature values  and converting them into the principal values of K 1, K2, and θ using Moore’s Law equations (equation 2- 4).  For this effort, the principal curvature values were taken and re- plotted on the undeformed structure to more  clearly illustrate the critical elements within the structure (Figure 9). 

𝐾𝐾1= 𝜅𝜅1+𝜅𝜅2 2+��𝜅𝜅2−𝜅𝜅1 2�2 +�𝜅𝜅3 2�2               (2)  𝐾𝐾2= 𝜅𝜅1+𝜅𝜅2 2−��𝜅𝜅2−𝜅𝜅1 2�2 +�𝜅𝜅3 2�2               (3)  𝜃𝜃 =0.5∗atan�𝜅𝜅3 𝜅𝜅1−𝜅𝜅2 �                               (4) 

Figure 9: Converting local element curvatures to principal values using Moore's Law conversion  equations.  

Hinge geometry for the ROC -PDS was initially selected to satisfy failure criteria . However , it was soon  discovered that the driving design requirements would drive the hinge design.  All hinge geometry trade  study values were subsequently based upon the cri tical limits of the material testing.
389 Hinge Torsional Stiffness Trade Study   Torsional stability of the hinge was shown to be a driving design factor of the ROC -PDS. To address what  design parameters would affect the torsional stability of the hinge most a model was developed. The model  that was used to evaluate the hinges torsiona l stiffness was conducted using ABAQUS explicit dynamic  solver, where the hinge section was fixed in all directions at one end and a pure moment was applied at  the tip of the boom about the tube axis ( Figure 10). 

Figure 10: Torsional Stiffness trade study model setup  

It was expected that the diameter of the end circle cutout  controlled by the “d” variable shown in Figure 5  would be the primary driver of torsional stability due to it producing the smallest effective torsional cross  section. This evaluation was  conducted to evaluate the effects of both length and diametric cutout of the  end cutouts. The curvature results were also compared to the material allowable values were then  evaluated using the methods described earlier .  

Figure 11: Initial torsional stiffness evaluation based on hinge geometry and material allowable prediction  

Results of this investigation showed both that the risk of material failure due to torsional inputs was very  low, and that the primary concern was geometric buckling behaviors. However, these buckling behaviors  occur at least an order of magnitude beyond any operational load cases, so this risk was retired. The design  output from this investigation was that minimizing the end cutout geometry and using a short slot cutout  length result in the stiffest hinge geometry.  

Hinge Torque Output Trade Study   Based on standard practice at Roccor with deployable structures, it is required that any deployment hinge  provide enough torque to deploy the system,  from any static position within the deployment , with a factor
390 of safety of two. The analysis model from the torsional stiffness investigation was utilized again for the  torque output investigation, with the application of tip load changed and a small perturbation  force added  on the compressed hinge tape to promote a controlled buckling of the hinge ( Figure 12). 

Figure 12: Hinge torque trade study model setup  

To maximize the torque output of the hinge the “w”  term from Figure 5 was minimized in order to maximize  the potential energy of the hinge. Verification of this design consisted of first evaluating the material  survivability by modeling the folded hinge with finite element s and using the evaluation methods described  earlier.  

Figure 13: Initial torque output capabilities based on hinge geometry and material allowable prediction  

Results of th e analytical hinge studies were to select hinges with minimalized “d” and “L” terms for the  hinge, however further investigations using a longer boom which allowed for more continuous deformation  of the tube cross section resulted in the hinge getting deformed far passed its material allowable limits. The  reason this was not seen in the initial investigation was due to the end conditions creating a rigid circle  relatively close to the hinge cutout area, resulting in artificially positive margins for the area around the  circular cutout. To address this issue the hinges were re- evaluated using a longer tube section ( Figure 14).  This allowed for more natural curvature regions to form around the hinges.
391   Figure 14: Lengthened hinge geometry trade analysis model  

This effect was exacerbated by the ability for the full ROC -PDS system to seat itself into a system low  energy state ( Figure 15), where the local curvatures around the hinge often resulted in negative margin.  Analysis results did not conclude with a geometric result which closed for this case, however a design  alteration to the overall system was devised to add rigidity to the loc al area and hold the hinge open. This  design feature is known as a “power band” and is the topic of another study paper currently in progress at  Roccor.  

Figure 15: (Left) High curvatures in full system analysis; (Right) hinge shift into low energy state.  

Hinge Torque and Long- term Storage Testing   Full-scale hinged boom testing was used to validate the analysis results and determine the properties of  the hinges where analytical techniques were unable to predict  the behavior  of this complex  system. The  primary test used was hinge torque testing under ambient temperatures and operational temperature  extremes.  This was  coupled with long- term storage testing to evaluate material survivability  and viscoelastic  effect s on strain energy relaxation.   

A torque test fixture was developed and qualified to accurately measure the hinge output torque. This fixture  is shown in Figure 16; the fixture uses torque cells to measure the output torque of each hinge as the boom  is slowly allowed to deploy, resulting in a semi -static torque measurement. This fixture was qualified and  validated by measuring the torque output of an easily modelled metallic leaf spring.
392   Figure 16: Structural testing of integrated solid- state HSC hinge  

The primary resistive torque in this design was due to a flat, flexible electrical harness  inside the boom. The  behavior of this harness depends intimately on the boundary conditions supplied by the boom, so testing  of the harness alone proved impossible. Instead, the hinge torque was measured with and without the  harness , and the harness torque was derived by comparing these two torque curves. These values were  then used to derive and show deployment torque margin.  A typical torque curve derived from this testing is  shown below in Figure 17.
393 Figure 17 Typical hinge torque of proximal hinge (left) and distal hinge (right).  

Later tests at operational temperature extremes were enabled by constructing a thermal chamber around  the torque test hardware. This testing was both used to directly verify the torque margin and validate the  analytical predictions.  

The impacts of the viscoelasticity on the deployment kinematics were also a key risk, and due to a lack of  standardized test methods  and stress relaxation failure models  these impacts are very difficult to analytically  model. Long- term storage testing was therefore used to sim ulate the impact of storage before deployment.  Full-length booms were stowed in the anticipated storage shape, as shown in Figure 18, and t he timetemper ature superposition relationship was used to reduce the storage time required from years to days .  

Figure 18 Long- term storage testing  

During and after storage the boom was inspected for damage. The corners of each hinge w ere at particular  risk of stress rupture during long- term storage, so these areas were subjected to extra scrutiny. Detailed  images of these areas are shown in Figure 19. After storage at elevated temperature the kick -off force,  hinge torque, and final boom geometry was measured. These measured values were fed into the analytical  model to better predict the deployment of the hinged boom after storage.
394   Figure 19 Inspection of the proximal hinge (left) and distal hinge (right).  

The results of this testing were then used to validate the analytical model , show compliance with  requirements, and provide knock -down factors for post -storage deployments. In particular , the  requirements for long- term storage survival and deployment torque margins were verified directly through  test. 

Final Boom Laminate Architecture Features  

At the culmination of the development program,  there were several  design features resulting from detailed  analysis and testing efforts done at Roccor. First,  the hinges were strengthened by “power bands” in order  to keep the slit from collapsing when the s ystem is required to be compressed beyond the material limit s of  the hinge deformation. Next, the “razor backs”  were added which are localized areas with adjusted  thickness used to tune the required torque output of the hinge tapes to develop the required strain energy  balance for the deployment. Last, the “lateral lines”12 were added in the boom, which  are a localized  laminate change which is utilized to tune the kickoff energy of the system. All these features were developed  using the methods described in the previous section for material allowable calculations , as well as physical  testing for validation of concepts and  verification against flight qualification environments.  

Figure 20: ROC -PDS boom design features
© 2020 Sierra Nevada Corporation   397 Mars 2020 Motor Bearing Failure, Investigation  and Response  

David Suffern*, Jeff Mobley* and Stephen Smith*  

Abstract  

The prime mover s in each joint of the external robotic arm of NASA Jet Propulsion Laboratory’s (JPL) Mars  2020 rover are planetary gearmotors containing a brushless DC motor  and brake assembly , designated as  M45S. Qualification l ife testing was performed at multiple levels of assembly in an effort to retire risk to t he  program. A  significant failure occurred within the motor  in the process of executing a life test  at the  gearmotor level of assembly . The combination of bearing retainer , lubrication scheme,  test temperatures ,  and long life led to a stall  failure of the M45S front bearing at 78 million revolutions, not satisfying the life  test requirement of 105.2  million motor rev olution s, which is double the expected life. Ultimately, increasing  the minimum life test temperature from -70°C to -55°C allowed for succe ssful qualification of the robotic  arm joints  with the baseline bearing configuration.  Details of the requirements, design,  life test, test failure  investigation, response , and lessons learned will be presented.  

Application and Requirements  

Reliable performance of the two- meter -long robotic arm carried by the Mars 2020 rover, shown in Figure 1,  is central to the success of this mission. Each  joint of this robotic arm is driven by  a M45S motor  and brake  assembly.  The M45S was incorporated into three different planetary gear motor  designs covering multiple   applications: s houlder and elbow (ShEl), w rist and t urret (WAT),  and feed. The ShEl  and WAT gearmotors  were then further integrated into harmonic drive mechanisms in the robotic arm. Each application has  unique load and life requirements based on the estimated robotic arm usage and travel.  Of these, the ShEl  application provides the enve loping requirement for mission life at  52.6 million motor revolutions , resulting  in a margined life test requirement of 105.2 million motor revolutions . 

Figure 1 – Robotic Arm of the Mars 2020 Rover  (Courtesy of NASA/JPL- Caltech)  

* Sierra Nevada Corporation, Durham, NC ; dave.suffern@sncorp.com  

Proceedings of the 45th Aerospace Mechanisms Symposium, NASA Johnson Space Center, 2020
© 2020 Sierra Nevada Corporation   400 on a separate Mars 2020 cont ract to sup ply gearboxes. This testing was intended to demonstrate bearing  life capability in the low cycle count , high load regime [ 2]. Unfortunately, i t was not possible to adapt that  test setup to encompass the speeds or cycle count required for the M45S motor  design.  Consequently,  demonstration of M45S bearing life was deferred until its design verification (DV)  life testing . 

Performance and Life Test Results  

M45S Performance   The goal to minimize torque variation over temperature was a significant program concern and drove the  decision not to implement a higher bearing lubricant fill . Performance testing was completed and proved  that the chosen motor design, including bearing lubrication,  performed consis tently across the entire  operational temperature range. The net motor torque constant , including viscous losses from lubrication,  across a temperature range of  -70°C to +70 °C resulted in a torque uncertainty of ±0.005  N•m at the 0.200  N•m output torque lev el at constant speed. This value was within  the original program goal of ±0.009 N•m .  With respect to torque uncertainty concerns, t he chosen design and lubrication scheme proved highly  effective.  

Life Testing and Results   Identified early in the program as an area of risk, the lubricant life w as ultimately tested during the M45S  DV life test, which exposed the unit to twice  the expected life, or 105.2 million motor revolutions. This life  test, completed in early  2019, had been delayed due to other  M45S design and integration challenges .  These delays postponed the retirement of the bearing life risk. As a result, within six weeks of finishing the  M45S DV life test,  two of the M45S gearmotor life tests ( ShEl  DV and WAT DV ) were completed or  stopped  due to anomalous conditions.  

This timing afforded the opportunity for near -concurrent visual inspection of motor bearings experiencing  the same operational temperature environment, but different total motor revolutions with various speed and  load combinati ons. The revolution count and condition of the M45S motor bearings from these three initial  life tests  were compared for similarities and differences, shown in Table 1. 

Table 1 – Results of Life Tests  Ending in Early 2019   Life Test  Motor Revs /  % of 2X Life Temperature  Distribution  Post -Life  Performance  Visual Bearing  Condition   M45S Motor -only 105M / 100%  25% min. @ - 70°C  25% min. @ - 55°C  25% min. @ +70 °C Acceptable  Poor   WAT  Gearmotor  66M / 100%  Acceptable  Marginal   ShEl Gearmotor  78M / 74%  N/A – Stalled  Very Poor / Damaged  

Figure 3 provides a visual comparison of the condition of the front bearing from the M45S mot or of each of  these three units, with the following observations:   • Motor -Only: Slight ly rough  rotation;  dry, powdery, dark wear debris; retainer ball pockets worn   • WAT : Rough rotation;  clumped, dark wear debris   • ShEl : Very rough rotation; retainer periodically  caught under balls; nearly empty of  any lubricant or  wear debris; d amage to retainer vis ible
© 2020 Sierra Nevada Corporation   401 Figure 3 – Post-Life Test Condition of Front M45S Motor Bearings  

Details of the ShEl Life Test  and Termination   The program required that life tests perform twice  the expected number of Integration and Test (I&T) and  Mission revolutions over at least 10 thermal cycles, each consisting of temperature plateaus at +70° C,   -70°C, -55°C, and +20° C. Additionally, the gearmotor output revolutions were required to be approximately  equally distributed across the thermal cycles with a minimum of 25%  of the revolutions  to be performed at  each of the +70° C, -70°C, -55°C plateaus . The balance of revolutions  could be performed at any  temperature between +70° C and - 70°C, but were mostly completed around +20 °C.  

By early 2019, nearly all of the gearmotor DV life tests had been completed without issue. However, the  ShEl life test unit unexpectedly  stopped operating  during the - 55°C plateau (just after the - 70°C plateau) of   thermal cycle 6 of 10 in early 2019. Figure 4 provides an overview of the data acquired during this plateau,  including the three stall events that occurred prior to halting operation for evaluation. The stall events are  indicated by the t hree sharp rises in the motor current and temperature followed by recovery . 

Figure 4 – Output Torque and Motor Current Data Acquired During Thermal Cycle #6, - 55°C Plateau
© 2020 Sierra Nevada Corporation   402 Figure 5 provides detailed data of the first stall event as i t occurred, with the following explanation:   • Per the automated profile, the unit bega n execution of 80 seconds of operation in the CW direction.   • After approximately 15 seconds of operation against the prescribed torque of approximately 6.6  N•m, the measured output speed of the unit  unexpectedly  dropped to 0 rpm and the motor current  jumped to the 7 A current limit.   • The unit remained  in this state for approximately 60 seconds , until the  automated profile removed  power from the motor . The motor  remain ed off for 15 to 20 seconds, as planned.   • The automated profile initiate d operation in the CCW direction.   • Operation then proceed ed nominally for the next 80 to 90 minutes until a second stall event was  encountered with the same signature as the first (reference Figure 4). Note, some initial torque  noise is visible on the graph and is associated with t he cogging effects of the hastily stopped  hysteresis brake dynamometer .  All three stall events were readily identified by the sharp increase in motor housing temperature associated  with Joule heating from the motor current being at its 7 A limit.  

Figure 5 – Detailed Data Acquired during First Stall Event  of ShEl DV in Life Cycle 6 of 10  

Test Failure Investigation  

The team first needed to determine whether the stall was due to an anomaly within the unit under test (UUT)  or the mechanical or electrical ground support e quipment (MGSE / EGSE) . The life test setup was more  complicated than most due to the requirements to  apply resistive torque and axial load simultaneously,  while support ing the UUT in the thermal chamber and recording data as noted.  The complexity of the test  setup presented multiple opportunities for the root cause to be found in the MGSE / EGSE.  A block  diagram  of the complicated ShEl life test setup is shown in Figure 6, with data acquired from components identified  in the orange boxes .
© 2020 Sierra Nevada Corporation   403   Figure 6 – Block Diagram of Complex ShEl DV Life Test Setup  

The following describes  the sequence of the investigation to determine the source of  the ShEl DV unit life  test anomaly :  1. To determine if the EGSE was responsible, t he team attempted to recreat e the stall event while  recording  the current of each motor phase and the voltage of each Hall sensor  with an oscilloscope.  Result: stall was recreated and oscilloscope data indicated EGSE was performing nominally.   2. The performance of the motor’s  brake was characterized to ensure it was not inadvertently  engaging and stopping the motor. Result: motor brake performance was verified to be nominal.   3. Thermal chamber was taken to ambient temperature and a recreation of the stall event w as again  attempted in a manner that would allow a determination of the source of the anomaly: MGSE or  UUT. Result: stall was recreated and there was no wind- up or tension in the MGSE test setup to  indicate  that it was causing any issues.   4. UUT was disconnect ed from all MGSE and a Startup Sensitivity health check was attempted to  verify the motor was able to initiate rotation at a current value in family with prior health checks .  The UUT failed to rotate with up to 2.8 A applied, when far less current  was gene rally needed to  initiate rotation. This provided convincing evidence that the source of the stall was within the UUT  itself.   5. JPL performed a computed tomography (CT) scan on the motor and discovered  that the source of  the stall condition was within th e front motor bearing; the crown steel bearing retainer jammed  under a ball , as shown in Figure 7. 

Figure 7 – CT Scan of Front M45S Bearing within ShEl Life Test Unit (Courtesy of NASA/JPL- Caltech)