ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
180  

Figure 1. The STIG mechanism, spindle actuator, and Core Break Lockout (CBLO) actuator are shown in  the context of the Mars  2020 corer.  

Figure 2. The output interfaces of the dual -output gearhead , part of the spindle actuator . The large, outer  internal spline provides high torque. The smaller, external spline undergoes fewer planetary  gearing stages  and outputs at a  faster speed. The high speed output is recessed deeper into the actuator than the high  torque output . This difference in depth is essential to the function of STIG .
181  

Figure 3. The STIG input interfaces . The large, three -tooth external spline interfaces with the actuator’s  high torque output  when the selector is retracted into STIG . The smaller, ten- toothed spline interfaces with  the high speed output when the selector is extended further into the actuator . The gear in view is invol ved  in positioning the selector, but it is not part of the torque- carrying path between actuator and drill bit. 

Figure 4. Completed STIG  flight assembly . In this image, the left side of the mechanism outputs torque via  the large gear to the spindle gear train, and ultimately to the drill bit . The right side of the image interfaces  to the dual -output spindle actuator as well as to a separate gear train that determines gear shifts . Also  visible in the image are a heater, two platinu m resistance thermometers, and their associated cabling .  Overall STIG  dimens ions are approximately 120 mm in length and  60 mm in diameter, with some flanges  extending further to serve as covers for other corer components.
182 Rather than having an additional, dedicated actuator that determines the axial position of the selector, STIG  references the position of a separate mechanism that is positioned according to whether drilling or core  break operations are occurring. This aforementioned  mechanism is named c ore break lockout (CBLO). The  CBLO  mechanism’s actuator  output  is attached to a single piece part containing both a lead screw and a  spur gear . More detail about this actuator can be found in [2]. The lead screw/ gear is shown in Figure 5.  

Figure 5. CBLO actuator and combined lead  screw/spur gear output piece part . The lead screw drives  a  mechanism that interact s with  the sample tube. The gear (via an idler) connects to STIG, allowing the  selector to interface to the proper dual -output actuator interf ace. 

The CBLO lead nut is extended during core break operations, allowing the CBLO mechanism to interact  with the sample tube and fix it rotationally, enabling shear of the rock core sample as the drill bit rotates  with high torque. During other operations such as drilling, CBLO is retracted to avoid contact with the sample  tube. As the lead screw operates during CBLO motion, the spur gear on the same piece part  mounted to  the actuator output  rotates, which is an input to STIG. Figure 6 provides an overall  schematic for the STIG  architecture . This gear does not provide torque to the drill bit, rather, it acts in translating the STIG selector  axially so that the selector interfaces with the desired output of the spindle actuator. The task of STIG is to  use knowledge of the CBLO mechanism’s position to place the selector at the appropriate output of the  spindle actuator, in order to provide the appropriate torque and speed to the bit .
183  

Figure 6. STIG utilizes  the positi on of the CBLO actuator to deter mine which output of the spindle is  connected to the coring bit . Other bits may be used as well . During core break, CBLO extends into the  sample tube, fixing it rotationally . Core break torque is applied t o the bit, creating eccentric motion and  shear forc es in the rock core.  

Mechanism Function  

Figure 7 shows the STIG design in detail, with several piece parts and features labeled. STIG’s lead screw  is not directly attached to the selector. Rather, the lead screw and selector  are connected by a spring that  is preloaded in a state of compression. This use of this pr eloaded spring has several advantages . First  when the selector is moved to a given position, there is 37 N of force holding the selector in place. This  force is sufficient to prevent  relative motion between the selector  and other  drill components during the   roughly 6 G accelerations produced by rotary -percussive coring and abrading, keeping the selector fully  engaged with the intended spindle actuator output . More detail about the corer’s percus sion mechanism is  available in [3]. This preload also allows the selector to be held in place while in the launch stow  configuration.  

In addition, the spring can be further compressed past its preload once the selector reaches the end of its  travel in either direction. This allows CBLO to be commandable to four different positions, with two positions  mapping to each of the two spindle actuator outputs. For example, CBLO can be posit ioned in the drilling  state, which positions the selector in the high speed mode. But if CBLO continues to retract in order to  perform a homing operation, whereby a non- rotating part of the CBLO mechanism seats into a groove in a  spindle piece part and locks out motion, the STIG lead screw can be further retracted. The STIG spring will  be further compressed, but the selector remains in place and mated with the high speed actuator output .  Conversely, while the sel ector remains  fully en gaged with the spindle actuator high torque interface, CBLO  can either be fully extended to interface with the sample tube during core break operations, or only partially  extended in order to provide breakoff torque to the bit without  locking out the tube, which could be useful in  recovering from a situation where a bit becomes stuck in the rock.
184  

Figure 7. STIG parts in detail . Red parts translate axiall y (and in some cases, rotate) . Magenta parts rotate   without translating. The selector  (dark blue)  is supported on its right side by the spindle actuator (not shown)  and on its left side by a bushing mounted to corer structure (not shown) . The right end of the output  shaft  is supported by splines tha t ride on the outside of the sel ector and on its left side by corer structure (not  shown).  

Finally, the spring provides the ability to shift gears while being agnostic to particular clockings of the  selector and actua tor outputs. As the STIG lead sc rew retracts, in the example of a shift from high torque  to high speed, the selector initially tracks the motion of the lead screw due to the preloaded spring. This  simultaneous motion continues until the selector’s high speed splines contact the splines of  the actuator  high speed output. Due to the differences in acutat or output gearing ratios and differences in backlash, no  assumptions are made about the relative clocking of the splines of the two output interfaces before a shift.  The lead screw continues to retract while the selector is prevented from moving, further compressing the  spring. Once the lead screw reaches its intended positon, the spindle is slowly rotated until the splines  align, allowing the selector to translate under the force of the sprin g preload into the fully -mated state.  

It is possible that friction between the contacting axial faces of the actuator and selector is capable of  sustaining  torque, so the selector may rotate with the actuator instead of remaining in place rotationally and  waiting for spline alignment . In order to guarantee successful mating, a previous benefit of the STIG spring  is utilized, namely, the ability for CBLO to travel to two locations for a given STIG gear. By driving CBLO  into its homing state and locking the STIG/spindle gear train, relative motion can be enforced between STIG  and the actuator. When the splines align and the selector moves axially into the mated position, a stall will  be detected by the actuator as it attempt s to drive a locked gear train. At that point, the gear shift is  guaranteed to be complete and CBLO can be extended to unlock the gear train without moving STIG out  of the high speed gear .  

The motion of the spring and a shifting example are shown step- by-step in Figure 8. A similar sequence of  events is used to attain high torque gearing, whereby CBLO locks the tube rotationally to initiate a stall  between the sp indle actuator and the gear train.
185   

Figure 8. STIG is shown shifting from high torque to high speed.     

A: STIG in in high torque mode, with  the the selector interfacing with the  spindle actuator’s high torque output .  Note that some corer components that  provide support to stig components are  not shown in the images.  

B: The lead screw and attached  components begin retracting (moving  right) . The selector follows.  

C: The lead screw and attached  components continue to their final  position, but are agnostic to the fact  that the selector had bottomed out on  the high speed spindle output and  stopped translating. A spring is  compressed during this difference in  axial travel between the lead screw  and selector.  

D: The actuator is rotated. At some  point, the high speed splines align. The  dotted line represents the extent of the  splines in this cross section before  actuator rotation.  

E: The selector is immediately pushed  toward, and interfaces with, the high  speed output . The spring partially  relaxes . The corer is now in the homing  configuration and STIG is in high  speed.
187 a highly loaded planet, a double- row ball bearing or  porous bronze bushing, would have resulted in an  unacceptable increase to the gearbox volume and mass .  

Figure 9. Cross Section of Dual -Output Actuator . Credit: Sierra Nevada Corporation  

While the hybrid ceramic needle roller was show n to be  capable of withstanding the demanding load  requirements, another concern early in the design phase was the roller bearing’s impact on drag torque. Its  location in the second  stage of the gear train required rotation at approximately 1000  rpm during drilling  operations, at a temperature extreme of - 70°C. Braycote 600EF grease was selected as the baseline  grease lubricant for the gearbox due to its performance at low temperatures and widespread space  heritage. In order to validate the bearing selection and lubrication scheme, a bearing torque test over  temperature was necessary. A single bearing expedited from the vendor and lubricated in- house with a   10-15% fill by volume of grease was tested. The breakaway torque as well as the running torque at  maximum speed of 1000 rpm at - 67 °C measured approximately 17.6 mNm. This test was performed such  that the bearing was operated at the hot operational temperature extreme of 70 °C to distribute the grease  and the n taken to the cold temperature extreme in a static condition to avoid disturbing the grease. Then,  the breakaway torque was measured prior to measuring running torque.  

Gear  train drag torques reflected back to the motor were calculated and found to be wi thin the bounds of  the motor performance specification. Key to meeting the bearing performance was the use of advanced  materials in the bearing design. The roller bearings utilized Cronidur X30 raceways, PEEK cages , and  Silicon Nitride rollers. Sierra Nevada Corporation  had successfully implemented these material  combinations  on a prior program but they lacked widespread space heritage, especially given the operating  temperature extremes. As such, the results of the life test and outcome of the teardown ins pection were of  great interest.  

The loads in the bearings resulted in several operating points that were within the high risk region for Bray  grease. The cumulative degradation factor (CDF) however was <1.0 for each of the three roller bearings  used in the design [4]. Figure 10 plots the cumulative stress cycles for a 2X life of each bearing relative to  the CDF factors of 2 and 8 in addition against guidelines established by NASA CP- 3062  [5]. For the  purposes of this figure, a stress cycle is calculated si milar to a ‘ball pass’ in a ball bearing.
188  

Figure 10: Mean Compressive Stress Cycles . Credit: Sierra Nevada Corporation  

Folloiwng the qualif ication test program, the actuator  met all  performance specifications. The testing  included static loading, which subjected the roller bearing to an estimated 3476 MPa ( 504 ksi) max hertzian  contact stress prior to any operation, random vibration testing, shock testing, and conclud ed with a 2X lif e  test.  

The life test consisted of 10 thermal cycles, the first five cycles tested the high- speed output, while the next  five cycles tested the low -speed output. The number of revolutions at each torque level were distributed  evenly among four  different temperatures ( -70°C, -55°C, +25°C , +70°C ). 

Figure 11: (Left) Inner r ace of roller bearing (X30) post life test. (Right)  cage and rollers, post life test.  Credit: Sierra Nevada Corporation
190  

Figure 12 . Torque and current data from the corer in the high -speed  gear at drilling speed (top) and in the  high torque gear measured statically against a load cell (bottom), for positive and negative directions of  spindle rotation. Data was collected at and between - 70 °C and +50 °C, the protoflight temperature  requireme nts of the corer body based on the Jezero Crater landing site.
191  

Figure  13. The absorber spring has a height of about 25 mm and a diameter of about 19 mm . The bottom  end is threaded for installation into the output shaft that surrounds the selector . The top end includes a  feature  where a two- piece split collar and retaining ring are placed . Contact with the selector is made at this  collar during impact . Regions of highest stresses are found at the round holes that terminate each slot.  

When designing a transmission that utilizes spring preload to drive gear changes, it was found to be  essential to study all off -nominal releases of energy and to plan for shift s occurring prior to planned  state  changes. For example, lead- in chamfers on the selector, designed to aid in engagement of splines during  shifting, made it possible for the selector to rotate and backdrive the spindle actuator  due to axial forces  sustaine d during shifting. Were backdriving of the selector to occur before the STIG lead screw reached its  end of travel, the selector could be accelerated and small par ts used to load the STIG spring w ould see  impact loads against each other.  These parts demanded a great deal of structural analysis and design  iteration  before this load case was deemed to be safe.  

When shifting in the high torque direction, the selector impacts corer structure. Although the compliance of  that structure is less than that of the absorber spring, the direct path of the load to structure means that  acceptable loads were substantially hi gher. Therefore, designing parts for impact here required less  iteration than that of the absorber spring.  Shunting loads to structure is preferable to a load path that  includes small  or fragile mechanism components whose designs are driven by other requir ements.  

Force Uncertainty in Coupled Lead Screws   An additional lesson learned was the need to plan for large force uncertainty when using two lead screws  geared together and driven by one actuator . Due to large differences in lead screw diameter (6 mm vs   31 mm, for the CBLO and STIG lead screw, respectively), the linear force output to input torque relations  of the lead screws are quite different. Potential small changes in observed actuator current could be due to  noise;  overcoming a small amount of unpl anned drag in the large, low -output force ratio lead screw; or  forces of several kilonewtons in the smaller, high- output force ratio lead screw. While testing a prototype
192 concept, a piece of ground support equipment that did not provide clearance for the f ull range of travel of  the high output CBLO  lead screw nearly led to the generation of forces capabl e of severely damaging the  corer . This fault was prevented by an operator E -stop press. The late addition of a spring in the CBLO  mechanism allowed for these high forces to be eliminated in a specific jamming fault case of concern.  

Conclusions  

When coupled to a dual -output actuator, STIG provides a compact approach to gear shifting. In the case  of the corer, by leveraging the position of a different mechanism, STIG does not require the addition of an  additional actuator specifically tasked with gear selection. STIG, through the use of spindle actuator stalls,  provides feedback that gear shifts have completed and splines are in full contact. No additional sensors are  required. STIG retains the ability to provide motion in either direction  in eithe r gear, by avoiding the use of  ratchets  and impact drivers . When designing mechanisms similar to STIG that use a spring to allow gear  shifts to be agnostic to actuator output spline positions, care must be taken to plan for energy dissipation.  Coupled lead screws of different sizes are an additional challenge,  due to their tendency  to obscure large  force generation. Testing of STIG, along with the core break lockout mechanism continues in the  qualification model test program at JPL. Launch is pl anned for July of 2020.  

Acknowledgement  

Except where noted otherwise, this  work was carried out at the Jet Propulsion Laboratory, California  Institute of Technology, under a contract with the National Aeronautics and Spac e Administration  (80NM0018D004) . © 2020. All rights reversed.  

References  

1. Myrick., T., US Patent No. 6,550,549, "Core Break off Mechanism”   2. LoSchiavo et. al . “Mars 2020 maxon Commercial Motor Development from Commercial -Off-the-Shelf  to Flight -Qualified Motors, Gearboxes, and Detent Brakes: Overcoming Issues and Lessons Learned” .   Proce edings of the 45th Aerospace Mechanisms Symposium. 2020.   3. Chrystal, Kyle . “Percussion Mechanism for the Mars  2020 Coring Drill” . Proce edings of the 45th  Aerospace Mechanisms Symposium. 2020.   4. Lo, C.J., et al., “Use of Cumulative Degradation Factor Prediction and Life Test Result of the Thruster  Gimbal Assembly Actuator for the Dawn Flight Project”, NASA CR -2009- 215681   5. Conley, P.L. and Bohner, J.J., “Experience with Synthetic Fluorinated Fluid Lubricants”, Proc. 24th  Aerospace Mech. Symp.,  NASA CP- 3062, Kennedy Space Center, FL (1990)
194  

Figure 1: NEA BCIS 8030 Family  

Design  Development Background   At the 2019 European Space Mechanisms And Tribology Symposi um (ESMATS), a summary of lessons  learned during Qualification testing of the heritage design was presented. The presentation discussed how  the BCIS had failed to actuate properly , and the details can be found in the ESMATS archive (Titled “ Bypass  Switches – A Case Study in Test Like You Fly Merits ”). To summarize the results, t he failure investigation  found that there were multiple causes for this anomaly.   1. A low actuation current (less than 3 amps) leads to an undesirable  condition that could affect the  performance of the BCIS.   2. At temperatures less than ambient, the likelihood of an anomaly increases.  

The reason that these two causes led to a failure was because they led to a potential negative force margin   on the split spool actuation device within the BCIS . Force margin  is defined as the ratio between  the  separating force and the resistive force. In this case, it is specifically the ratio between the restraining wire  separating force and the fuse wire res istive force, including all frictional forces . If the restraining wire  separating force is greater than the resistive forces caused by the interaction between the restraining wire   loop and fuse wire, then the restraining wire  can unwind resulting in successful actuation.  

The heritage design used a restraining wire with a 0.014- in (0.36- mm) diameter , which  corresponded to a  separation force of approxi mately 0.44 lbf (2 N) . In most cases, this force was sufficient to separate the  restraining wire from the fuse wire, leading to a successful actuation, but the causes 1 and 2 listed above,  combined with production lot variation, led to a possible condition where the restraining wire would not  separate from the fuse wire  or the separation would be delayed.  This paper discusses the design iterations  and lessons learned during the design and qualificatio n of the improved BCIS design, specifically within the  release mechanism portion of the BCIS . 

Design Development  

The development of the improved BCIS started with a set of basic requirements:   1. The design improvements  should not affect the envelope of the existing desig n  2. The improved design should be able to withstand the most extreme environments and conditions   that have been requested by our customers   3. The electrical performance of the BCIS should remain unchanged  

Spool  Restraining Wire  Fuse wire  Plunger  High Current  Carrying  Components  T2 T3 T1  Release  Mechanism  T4 & T5
195 Knowing the limitations of the existing design, the app roach was to improve the force margin  to such an  extent that low current, environmental impacts and assembly variation would not affect the performance of  the BCIS.  There were two initial designs that were considered.  

The first was to  reduce resistive  forces by  remov ing some of the insulating material around the fuse wire  and restraining wire. This would eliminate frictional forces and would guarantee a positive force margin.  However, a Design Failure Mode and Effects Analysis ( DFMEA)  was held and multiple potential risks were  identified. Removing material had the potential of creating pinch points, snagging possibilities  and other  potential risks that eventually led to the elimination of this option.  A rigorous vetting process was abl e to  detect these issues  before significant resources were expended on this design.  

The second option was to increase the separation force. The easiest approach was to increase the  restraining wire diameter, there by increasing the spring back force of the wire and yield positive margin.  Initially, two different sizes were considered, a 0.016- in (0.41- mm) diameter wire and an 0.018 -in (0.46- mm)  diameter wire. A nalysis and prototype testing were performed to determine the best choice.  

Proto type Testing   The spring back force of each size restraining wire was calculated through a mathematical model, and both  sizes were expected to provide significant margin over  the heritage design.  Since the envelope  of the design   was critical, an assessment of the clearance was one of the first tests performed. Multiple units were  functioned, and the unwinding  diameter  of the restraining wire was measured.  This w as important, as a  large unwinding diameter could indicate that the restraining wire may not have sufficient clearance to  unwind properly . In addition to post -test inspection,  functional testing was performed with the use of a hi ghspeed camera. It was clear that the 0.016- in (0.41- mm) restraining wire unwound smoothly and fit well  inside  the BCIS, while the 0.018- in (0.46 -mm) restraining wire unwound to a considerably larger diameter  and resulted in a condition where the restraining wire could potentially become jammed against the wall of  the housing.  Therefore, the 0.016- in (0.41- mm) restraining wire was chosen for the final design. The  analytical model was validated by test and shown to be in good agreement with the test results, with the  restraining wire loads found to be normally distributed about the mean.  It was also shown to be a si gnificant  improvement to the heritage design. Figure 2 is a plot of the restraining wire force of both the 0.014- in  (0.36- mm) and 0.016- in (0.41- mm) restraining wire.  The data shows that with the 0.014- in (0.36- mm)  restraining wire there exists  a very low probability of not providing enough force to separate from the fuse  wire. For the 0.016- in (0.41 -mm) restraining wire, t he force margin was always positive when using a  +4 sigma for the frictional force and a -4 sigma for the separation force.   

Figure 2: NEA Ultimate Load Test