ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
231   Figure 13  – Axial force as a ratio of the normal load as a function of braking torque applied to the  output shaft for a bare 440C roller and a nitrided 15-5 roller operating in vacuum at 15 rpm. 

Summary   A set of experiments were done using roller pairs to understand and quantify the forces that can develop for a variety of test conditions. The materials and test conditions were selected to help understand the behavior of the SARJ mechanism. Tests were done usin g hardened 440C rollers mated with 15-5 rollers,  the 15-5 rollers being both in nitrided and bare (not nitrided) condition. Tests were done with no lubrication, solid gold-film lubrication, and grease lubr ication. Of great significanc e to the operation of the  SARJ mechanism is the magnitude of the axial force that will develop because of shaft misalignment. The  experiments demonstrate the critical role of good lubrication for the SARJ mechanism. The following  specific results were obtained:    1. For the case of a bare (un-plat ed) 440C roller against a nitrided 15-5 roller without lubrication, the  axial force can be as great as 0.4 times the normal load for a shaft angle of 0.5 degree.  Experiments and analysis done by others have shown that such a magnitude of force on a SARJ  roller would cause roller tipping and contact pressures much greater than anticipated by the designers.    2. The axial force for the case of bare (un-plated) 440C vs. un-nitrided 15-5 with no grease is  somewhat less than can be obtained using the nitrid ed 15-5 surfaces. Still, with no lubrication the  axial force can be high, an undesirable condition for the SARJ mechanism.    3. For the case of a gold-plated 440C roller agains t a nitrided 15-5 roller without grease, the axial  force does not exceed about 0.2 times the normal  load even for the largest misalignment angles  tested. 

4. For the case of a bare (un-plated) 440C ro ller against a nitrided 15-5 roller with grease  lubrication, the axial force does not exceed about  0.15 times the normal load even for the largest  misalignment angles tested.   braking torque (N-m) 0123456axial force / normal load 0.00.10.20.30.40.50.6 NASA/CP-2010-216272
232 5. The experiments and data demonstrate the critical role of good lubrication. The gold films  provided good lubrication for the short duration testing done and reported herein. Grease  lubrication limited the magnitude of the axial fo rce to even smaller magnitudes than was achieved  with the gold-film lubrication. 

6. For the case of a bare (un-plat ed) 440C roller against a nitrided 15-5 roller without lubrication, the  maximum attainable axial force was slightly less when testing in air as opposed to testing in  vacuum. 

Acknowledgements 

This research was supported by the NASA Engineeri ng Safety Center. The experimental work was done  with assistance from Mr. Justin Elchert of Case Western Reserve Un iversity as part of the NASA GRC  LeRCIP Summer Student Program.     References 

                                                  1. Loewenthal, S. and Schuller, F., “Feasibility study of a discrete bearing/roller drive rotary joint for the  Space Station”, NASA Techni cal Memorandum 88800, 1986.   2. Johnson, K.L., Contact Mechanics, Cambridge University Press, 1985.   3. Kalker, J.J., “Rolling contact phenomena: linear  elasticity”, Rolling Contact Phenomena  CISM  Courses and Lectures , Issue: 411, Springer-Verlag, 2000.   4. Woody, D. and Lamb, W., "A design for a precis ion 10-m sub-millimeter antenna", ALMA Memo  241,Mar 1999. http://www.mma.nrao.edu/memos/html-memos/alma241/memo241a.pdf ,  accessed  Feb. 26, 2009.  5.  McGinness, H., "Antenna azimuth bearing model experiment," DSN Progress Report 42-53, July and  August 1979, Jet Propulsion Labo ratory, Pasadena, Calif., 1979.   6.  McGinness, H., "Lateral forces induced by a misaligned roller," DSN Progress Report 42-45, March  and April 1978, Jet Propulsion Laboratory, Pasadena, Calif., 1978.   7.  Hamrock, B., and Brewe, D., “Simplified Equat ion for Stresses and Deformations,” ASME J. Lubr.  Technol., 105, pp. 171–177, 1983.   NASA/CP-2010-216272
234 SA/SPaH also contains other hardware that supports the overall sampli ng investigation (Figures 3 and 4).  These include two spare bits for the Drill in indivi dual Bit Boxes, an Organic Check Material (OCM), an  Observation Tray, and Inlet Cover mechanisms that ar e placed over the SAM and Chemin solid sample  

Figure 1.  SA/SPaH is mounted on the front of the Rover and is shown in a configuration suitable  for driving. 

Figure 2.  The approximately 600-mm-diameter Turret consists of 5 devices.   CHIMRA  MAHLIAPXS Drill  DRT  NASA/CP-2010-216272
235  

Figure 3.  Top View of the stowed SA/SPaH on the Rover 

Figure 4.  Front View of the stowed SA/SPaH on the Rover  

Robotic Arm and Turret  Organic Check  Material  Observation  Tray  Extra Bits in Bit Boxes  Extra Drill  Bits  Observation  Tray Robotic Arm and  Turret  Chemin and SAM  Inlet Covers  Organic Check  Material  NASA/CP-2010-216272
239 respect to gravity is used only to bias the dire ction of sample movement but nowhere within the  subsystem is gravity relied upon as the sole prime mover.   Sample Paths

The functionality described in the previous section supports the following five minimum sample pathways  required for the scientific investigation on Mars.   1. Acquire rock sample with the Drill and deliver up to six 45-65 mm3 portions of particles less  than 150 µm to SAM and Chemin.  2. Acquire regolith sample with the CHIMRA Scoop and deliver up to six 45-65 mm3 portions of  particles less than 150 µm to SAM and Chemin.  3. Deliver the remaining bulk processed sub-150 µm  sample (rock or regolith), after portioning to  SAM and Chemin, to the Observation Tray.  4. Acquire regolith sample with the CHIMRA Scoop and deliver one 45-130 mm3 portion of  particles less than 1 mm to SAM.  5. Acquire OCM sample with the Drill and deliver up to six 45-65 mm3 portions of particles less  than 150 µm to SAM and Chemin.   Figure 5 illustrates three representative Robotic Arm poses used during sample pathway 1. In Figure 5a,  the Arm has positioned and preloaded the Drill on the tar get rock. The Drill then uses its translation,  percussion, and rotation degrees-of-freedom to powder the rock and auger the powder into a collection chamber in its Drill bit assembly. Sample moves from the Drill bit assembly into CHIMRA through a  Sample Transfer Tube connecting the two devices. Part icle motion is accomplished by using the Arm to  bias the particle flow with respect to gravit y and alternately turning on Drill percussion and CHIMRA 

Figure 7.  CHIMRA has three required sample pa thways: Drill to sub-150 µm portion, Scoop to  sub-150 µm portion, and Scoop to sub-1 mm portion. The Drill to sub-1 mm portion pathway is also physically possible. [Reference 2]   NASA/CP-2010-216272
241  

               Figure 8.  Flowchart of the Drill, CHIMRA, and Inlet Cover functions comprising the 5 sample  pathways and the individual degrees of freedom required to perform the function.  Note: Arm degrees of freedom are not listed on this  chart but are required for all sampling functionality  1  3 4 2  1, 2, 5 1, 5 1, 5  1, 2, 5 5  1, 2, 5 1, 2, 5 4 4 4 Drill Translation  Drill Percussion  Drill Rotation Acquire powder from  rocks with Drill   Drill Translation  Drill Percussion  Drill RotationAcquire powder from  OCM with Drill  

CHIMRA Scoop   Acquire loose regolith  with CHIMRA scoop   CHIMRA Vibe  CHIMRA Thwack  CHIMRA Portion Move sample to 150 µm portion  box, create, and deliver portion  Chemin Cove rOpen Chemin  Inlet Cover and  receive portion CHIMRA Vibe  CHIMRA Portion Deliver sample to  Observation Tray 

SAM Cover 1Open SAM  Inlet Cover 1 and  receive portion  SAM  Cover 2Open SAM   Inlet Cover 2 and  receive portion CHIMRA Vibe  CHIMRA Scoop Move sample to 1 mm portion box,  create, and deliver portion CHIMRA Vibe  CHIMRA Thwac k Move sample to 150 µm sieve   and sieve sample   CHIMRA Vibe  CHIMRA Scoo p Move sample to 1 mm sieve   and sieve sample  Drill Percussion  Drill Rotation Move sample through Drill Bit   CHIMRA Vibe Move sample through Sample  Transfer Tube into CHIMRA  Legend   Sample path # Required degrees of  freedom Function   NASA/CP-2010-216272
243 the parallel blade flexure. The vibration level is adjusted by changing the amount  of the eccentric mass  and the mass is chosen to produce between 4 and 10 G at the 150-µm sieve. Development testing  showed that 150-µm sieving required the highest le vel of vibration among all the CHIMRA sample  functions. 4 G at the 150-µm sieve provided performanc e with margin for sieving the most difficult  material that was tested. 10 G was used as the design limit. An initial prediction of the CHIMRA vibration  mode was about 85 Hz but the range of 70-100 Hz was re served to account for uncertainties. The Turret  design keeps the non-CHIMRA Turret modes out of the 70-100 Hz range to minimize the coupling  between CHIMRA modes and other Turret modes through frequency separation. The Instrument Isolators  are required to limit the Instrument  response at the Instrument mounti ng interface to: 1) 4 G during steady  state operation of the CHIMRA and the Drill, and 2) 6 G during trans ient operation of the CHIMRA and  the Drill (such as startup).  The initial Isolator design concept was composed of six linear compression spring struts in a hexapod arrangement (Figure 9). The isolation performance of  this concept can be accurately modeled so the  spring stiffness can be chosen analytically and then c onfirmed by test without  numerous iterations, a  clear advantage over the final wire r ope design. However, this design did not fit within the severe volume  constraints on the Turret. Also the part count and mec hanical complexity of the spring struts was higher  than the wire rope Isolator design.  Our wire rope Isolator design (Figure 9) is based on a commercial product used for disturbance attenuation for equipment mounted on aircraft and in shipping containers. It is a mechanically simple component but it is highly nonlinear and not easily modeled. Testing showed that as the disturbance  amplitude increases, the Isolator  frequency decreases. Sizing our Is olator design required numerous  development test iterations. After the desired loop l ength was chosen by testing the commercial versions,  an Isolator suitable for flight was created using alum inum caps with stainless steel wire rope bonded into  holes tailored to provide a good bonded joint. Further  development testing determined the appropriate  number of wires needed for each Instrument Isolator to meet its performance requirements while minimizing the instances of the Isolator bottomi ng out during random vibe testing. Response limiting  during random vibe testing is also being used to prevent this issue.    

                 Figure 9.  Two Isolator designs were considered before the wire rope design was chosen. The  APXS instrument is shown on its wire rope Isolator. One of the wire rope Isolators is shown deforming under the PF level launch environment during an Instrument Isolator development test.   Wire Rope  Isolator   (1 of 3 ) Turret  mounting interface Wire Rope Isolator Assembly APXS Instrument Contact Sensor Assembl y Linear spring / hexapod design  NASA/CP-2010-216272
245 the eccentric mass. Data were taken with the Arm in  the sieve pose with the Turret at different angles and  the Drill in the surface stow position.  The Drill surface stow position ali gns the sample exit on the Drill bit  with the Sample Transfer Tube. In the nominal oper ating position (Turret at 180 degrees), the results  show the Instrument response meet s the 4 G requirement during steady  state CHIMRA operation with the  steady state response at CHIMRA set near 8 G in t he X direction. The CHIMRA response in the Y and Z  directions is less than 4 G. No higher response was obs erved during startup. It is interesting to note that  by changing the Turret position, the CHIMRA respons e can be increased or decreased (over the range of  7 to 10.5 G in the X direction) while still meeting t he Instrument requirements. This may prove to be a  useful feature if difficulties are encountered during testing or on Mars. 

The EM Turret is about to begin testing with Mars analog materials in a low-pressure environment which  is called EM Dirty Testing to highlight the fact  that rocks and regolith are being processed. Although  extensive development testing was done, this is the first testing wi th Mars analog materials on flight-like  hardware in an assembled flight-like configuration.  

               Figure 10.  The assembled EM Turret with the Drill, CHIMRA, and Instrument Isolators installed.  Mass models are used for the DRT, APXS, and MAHLI. 

            Figure 11.  The EM Turret / Robotic Arm during dy namic characterization testing: primary sieve  pose with the Turret at 180 degrees (left) and depositing a 150-µm portion (right).  NASA/CP-2010-216272
246                          Figure 12.  The Turret is shown installed on the test manipulator and drilling a rock in the sandbox  (left) and inside the chamber (right). 

 Figure 12 illustrates the Turret installed on the test manipulator and in the test chamber. The 4 degree-offreedom test manipulator was developed as a substitu te for the Arm during both the EM and QM Dirty  Test programs for two reasons. One is that no Arm is available to support these tests since the EM Arm is  fully subscribed in the testbed and there is no QM Arm. Two is the test manipulator enables the testing to  occur in a chamber that is more manageable in size  than the one that would accommodate the Arm. The  more manageable size comes at the cost of anot her dynamic characterization program. Dynamic  characterization of the EM Turret instrumented with  accelerometers on the test manipulator will be done  in the same manner as when it was installed on the Arm. For the sampling functional performance results  on the test manipulator to be valid, the Turret response needs to reproduce or be bounded by the Turret  response on the Arm so the functional performance c an be linked to the operational dynamic environment  that the flight configuration produc es. Adjustments in the test manipulator after characterization may be  required to accomplish this.  Lessons

Characterize the operational dynamic response in har dware configurations other than the nominal ones.  Although this time will be difficult to find in the typically  oversubscribed test schedule, it can result in ways  to alter the behavior of the system  that may prove useful when diffi culties are encountered both in the  Earth-based sampling testing and in service on Mars.  The additional complexity added to the test program  when using intentional dynamic environments to  process sample cannot be overstated. Care must be taken to ensure that the relevant environment is  produced when the hardware test  configuration changes.      ~ 1.13 m  ~ 2.04 m  Turret  Sandbox Test Manipulator  Reachable Works pace NASA/CP-2010-216272
247 Conclusions 

The MSL SA/SPaH has been designed and implemented, with flight-like EM hardware about to start Dirty  Testing and the remaining hardware (QM, FM) soon to follow. Along the way some lessons were learned  in subsystem configuration, faul t tolerance, intentional dynamic environments, and special testing.  Additional lessons are still to come as the EM, QM, and FM test programs are completed. 

Acknowledgements  The work described in this paper was performed by t he Jet Propulsion Laboratory, California Institute of  Technology, under contract with the Nati onal Aeronautics and Space Administration.   Creating something like the SA/SPaH and producing it for flight is an enormous undertaking that relies on substantial contributions from far too many people to name effectively. So I will acknowledge a set of JPL  people while recognizing that they like me are just t he most visible representat ives of the group. The  hardware cognizant engineers and their teams have worked tirelessly for years to reach this point: Kyle  Brown (Bit Box), Brett Kennedy (Robot ic Arm), Dave Levine (Turret Conf iguration, Isolators), Avi Okon  (Drill), and Dan Sunshine (CHIMRA). Mark Balzer, Richard Barela, Ke vin Burke, Joe Melko, Suparna  Mukherjee, Matt Orzewalla, Don Sevilla, Lori Shir aishi, Jeff Umland, and Chris Voorhees have made  substantial contributions, also for years. Special thanks go to Dave Putnam of Lockheed Martin,  Sunnyvale, who spent a great deal of time with the team when the Drill and CHIMRA mechanisms were  in their infancy.   References 

1. Okon, Avi B. “Mars Science Laboratory Drill.” Proceedings of the 40 th Aerospace Mechanisms  Symposium , (May 2010).  2. Sunshine, Daniel. “Mars Science Laboratory CHI MRA: A Device for Processing Powdered Martian  Samples.” Proceedings of the 40th Aerospace Mechanisms Symposium , (May 2010). 

NASA/CP-2010-216272
251 Overview of CHIMRA 

The majority of CHIMRA’s sample processing func tionality is accomplished inherently with the overall  design configuration. The internal structure is a labyrinth (Figure 5) that contains two primary  passageways used to flow sample into a central pr ocessing reservoir from two independent sources: bulk  regolith, scooped from the Martian soil by CHIMRA,  or powdered sample collected by the drill and  transferred via the sample transfer tube. From the c entral processing reservoir, sample can be diverted  through either a 150-micron sieve or a 1-millimeter sieve into two independent portioning chambers. A  schematic of the CHIMRA sample processi ng paths is shown in Figure 4.   

Figure 4. CHIMRA Sample Processing Paths 

        Figure 5. Internal CHIMRA Labyrinths       Figure 6. One-way valves to control flow  NASA/CP-2010-216272
258 Good Vibrations: Development of the Vibration Mechanism 

The heart and soul of CHIMRA’s samp le processing ability, the Vibratio n Mechanism (VM) (Figure 13), is  required to generate the dynamic env ironment CHIMRA uses to flow, sort and portion sample over an  operational life of two years and 200 million revolutions. The load generating component of the VM is an eccentric tungsten mass that is  supported by two back-to-back, spring-preloaded, angular contact  bearings. The motor-side bearing is pressed onto the shaft as well as into the housing; whereas the outboard bearing only contains a press fit on the outer race. The inner race is pressed onto a hollow shaft (green in Figure 13) that slips over the primary sha ft of the mechanism. The sliding inner race is then  preloaded with a wave spring, reducing the preload s ensitivity to assembly tolerances. The central shaft  and eccentric mass is driven by the vibration motor through a flexible helical-coupler at speeds between 4000 and 6000 revolutions per minute.                      Figure 13. Vibration Mechanism   The torque available from the largest MSL motor was limited when operating at 6000 rotations per minute. This power limitation forced the VM to balance the competing requirements of bearing lubrication drag and long mechanism life. This balance was struck  through development testing in the following way.  The bearing and lubrication design evolution dict ated finding a bearing configuration that produced  contact stresses less than 1.38 GPa (200 ksi), convergi ng through test on the maximum lubrication fill  factor that afforded acceptable torque margin, and testing the mechanism across the significant life requirements of the device. A development test program was conducted that compared the final  lubrication drag across a temperatur e range of -70 to +70 degrees Celsius for two bearing sizes (10-mm  OD and 30-mm OD), two different radial loads (150 and 250 N), three test speeds (4000, 5000 and 6000 rotations per minute) and increasing f ill factors working up to 30% fill factor by volume. The bearings were  initially greaseplated with Braycote and additional f ill factors were generated by inserting a slurry of  Braycote grease and Brayco oil. By fabricating and  testing three different mechanisms simultaneously,  each of the previous variables could be individually va ried. When the optimal configuration was found, it  was successfully tested past 200 million cycles across a fluctuating temperature profile that represented  the flight environment (Figure 14).    A striking result from this testing was that the co mbination of the parameters unique to this application  resulted in visible transitions across three different lubrication regimes: mixed boundary/  elastohydrodynamic , purely elastohydrodynamic and hydrodynamic. For discussion purposes, the  10-mm OD bearing did not display significant changes in performance across the parameter design space and therefore the discussion will focus only on the 30-mm OD bearing. In the vibration mechanism,  the load generated from the eccentric mass is proportional to the operating speed squared. However,  NASA/CP-2010-216272
259 testing on both the development unit and the flight unit indicated that the recorded lubrication drag was  insensitive to changes in the load applied. This wa s true even if the load was increased from 100 to   250 N. It is therefore assumed that substantial changes in recorded drag  at different operating speeds are  primarily a result of the change in speed. Figure 15  illustrates how changes in speed and chamber test  temperature affect the lubrication drag within the me chanism. Predicted trends of an inverse relationship  between lubrication drag and operating speed were obs erved. This suggests that increased operating  speed is driving the lubrication regime of the bearing from an elastohydrodynamic regime to more  hydrodynamic than was expected.                        Figure 14. 200-million cycle bearing life test    The transition through different lubrication regimes is  also obvious where the expected inverse relation  between chamber temperature and lubr ication drag was not observed. Specifically, across all three  operating speeds, the lubrication drag  recorded at -25 degrees Celsiu s was less than at ambient (20  degrees) and -65 degree Celsius. The low lubrication viscosity at ambient temperatures resulted in a  mixed boundary/ elastohydrodynamic operating regime. However, as the temperature decreased to -25 degrees Celsius and viscosity increased accordingly,  the regime changed to be pure elastohydrodynamic  one and a corresponding lower lubrication drag was recorded. Further decreases  in temperature (-65  degrees Celsius) were sufficient to thicken the visco sity to a level where the operating regime of the  bearing was hydrodynamic and the corresponding lubr ication drag increased beyond that of the  elastohydrodynamic regime. This same transition acro ss lubrication regimes was recorded when testing  two operating speed across the temperature range on t he flight VM (Figure 16) and is consistent with  Stribeck Curve Theory (3). A signifi cant increase in drag can be seen when the lubrication fill factor is  increased from 15% to 20%. The final  parameters selected for the flight  configuration were a 30-mm OD  bearing with an additional 15% fill factor by volume th at is operated at a cont act stress of 862 MPa (125  ksi). Once assembled, an infant mortality test was successfully per formed that exercised the flight  mechanism to 20 million revolutions.     Reflections on Development Testing to Support the Design of a Sample Processing Device 

To mature the design of the CHIMRA sample proc essing system, an extensive suite of development tests  were conducted to form a basis of particle flow mech anics. A quick review of the relevant results that the  CHIMRA designing principles were ba sed on will provide future designe rs a more advanced platform to  start from when creating a similar design. The results fr om a series of tests aimed at sample flow across  surfaces revealed that both the surface finish and method of surface preparation had a strong effect on  NASA/CP-2010-216272