ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
67 Comparison with base oils   The performance of greases in comparison to their base oils is given in Table 3. Values for the oil Castrol  815Z are taken from the SOT commissioning tests perf ormed by ESTL upon delivery of the facility. From  the values in Table 3, we see that  the performances of the MAPLUB PF PE greases are similar to those of  their base oil when considering lifetime, and that the friction coefficients for Z25 fall somewhere between  the old and new MAPLUB grease formulations. The change in gradient (demonstrated in Figure 9) is not  seen for the base oil. The performance of the Braycote  greases is also similar to their base oil Castrol  815Z.    Table 3. Friction and lifetimes of greases and oils assessed at 1.50 GPa mean contact stress  under vacuum in the SOT, 100 RPM rotation speed   (with the exception of the Z25 oil, performed at 30 RPM)   Grease Base oil Lifetime  (Orbits/μ g)Friction @  400 th orbit Friction @  half lifetime  -- Fomblin Z25 239 0.099 0.106  MAPLUB PF 100-a Fomblin Z25 184 0.100 0.112  MAPLUB PF 100-b Fomblin Z25 201 0.084 0.103  MAPLUB PF 101-a Fomblin Z25 225 0.100 0.112  MAPLUB PF 100-b Fomblin Z25 201 0.081 0.101  -- NYE 2001A 3,937 0.081 0.090  MAPLUB SH 050-a NYE 2001A 10,653 0.072 0.075  MAPLUB SH 100-b NYE 2001A 1,376 0.061 0.093  MAPLUB SH 051-a NYE 2001A 5,492 0.070 0.104  MAPLUB SH 101-b NYE 2001A 522 0.065 0.095  -- Castrol 815Z 233 0.093 0.104  Braycote 601EF  Castrol 815Z 203 0.088 0.103  Braycote 601EF Micronic  Castrol 815Z 198 0.092 0.104  -- NYE 2001 4,189 0.083 0.090  Rheolube 2000 NYE 2001 4,946 0.067 0.079    MAC MAPLUB greases were varied in comparison to their base oil NYE 2001A when considering both lifetime and friction coefficient. However, friction coeffi cients were observed to be lower for these greases  in comparison to NYE 2001A when considering the ear ly stages of the tests.  The grease based on the  NYE 2001 oil, Rheolube 2000, gave a similar lifetime, and slightly reduced friction coefficient than that of  its base oil. Similar lifetimes of NYE 2001 and R heolube 2000 (as well as Castrol 815Z and Braycote  601EF) have previously been observed when using the SOT [15].    Solid Lubricant Assessment   Tests on MoS 2 and lead displayed similar behavior to the liqui d lubricants, characterized by a long period  of low friction before a dramatic increase to failure. In all cases the running-in period to low friction was  relatively short in comparison to bearing tests, which can display high torque for many thousands of  revolutions. Summaries of lifetimes and friction coefficients for MoS 2 (as a function of coating thickness)  and lead (as a function of mean contact stress) are plotted in Figures 11 to 14. Friction coefficients were found from the mean of 100 readings around the 100,000 th orbit for MoS 2, and 1,000,000th orbit for lead. 

   NASA/CP-2010-216272
68 MoS 2  Figure 11. Friction coefficients (assessed at  ~100,000th orbit) of MoS 2 as a function of  coating thickness Figure 12. Lifetimes of MoS 2 as a function of  coating thickness  

For MoS 2 a mean friction coefficient of μ = 0.017 was found, with little variation across the assessed  coating thickness range 850 – 5,300Å (Figure 11). In addition, the lifetime of the coating is also not  observed to be dependent upon coating thickness within this range, displaying a mean of 260,000 orbits  (Figure 12). It is speculated that a greater initial thickness of MoS 2 applied to the ball merely results in  greater volumes of lubricant material being lost in t he early stages of rolling, with the ball running on a  much thinner coating for the majority of its lifetime.  X-Ray Fluorescence analysis of the post-test samples  revealed no MoS 2 remaining on the ball and loose debris pushed clear of the ball track, demonstrating  failure in these cases is caused by removal of MoS 2 by the actions of rolling, sliding and pivoting. 

The frictional performance of thin films of lead was found not to vary with increasing contact stress, with a  mean of μ = 0.046 found over the tested contact stress range (Figure 13). This friction coefficient is some  2.5x higher than the value found for MoS 2 under similar conditions, a relationship which corresponds well  with results from angular contact bearing tests [14] . Lifetimes were seen to decrease with increasing  contact stress, an expected behavior, with the lowest  stress test (1.50 GPa mean contact stress) being  stopped at +3.5million orbits  showing no indication of failure (Figure 14).  

Lead   Figure 13. Friction coefficients (assessed at  ~1,000,000th orbit) of lead as a function of mean  contact stress  Figure 14. Lifetimes of lead as a function of  mean contact stress  

Under similar conditions, the lifetimes displayed by th in lead films were extended an order of magnitude in  comparison to MoS 2. Similar extension of life when using lead lubrication is seen in angular contact  bearing tests [14]. This result demonstrates well the ab ility of the SOT to assess the performance of thin  solid film lubricants in comparison to other tribometer s, particularly pin-on-dis c (POD) tribometers, which  NASA/CP-2010-216272
69 are capable of only simulating sliding motion. Figures 15 & 16 demonstrate the difference in lifetime  gained from these two methods, with the result from  the SOT (Figure 15) being a much more accurate  model of the behavior seen in angular contact bearings.   Figure 15. Lifetimes of thin solid films (ball only)  as assessed with a SOT  Figure 16. Lifetimes of thin solid films as  assessed with a POD tribometer   XRF analysis of the samples used in the 1.50-GPa lead test (+3.5 million orbit lifetime) revealed less than 100 Å of lead remaining on the ball (Figure 17), reduced from the originally deposited 850 Å. Additionally,  low but finite lead readings were given in the ball tracks  on the flat plates, with the highest readings given  in the scrub region (Figure 18). Accurate thickness measurements could not be taken in the ball tracks as the XRF measurement area has a greater diameter than the width of the track (>1 mm). These results lead to the theory that during the test lead is re-dis tributed from the surface of  the ball onto all contact  surfaces, promoted by the ductile nature of the lead in comparison to the more friable MoS 2. It is  speculated that this process is the cause of the order of magnitude extension in life displayed by lead over MoS 2 running under similar conditions.     Figure 17. Ball from 1.50-GPa contact stress  test on lead, displaying lead mottles on  surface. XRF analysis read ~100 Å lead  remaining  Figure 18. Flat plate (top) from 1.50-GPa  contact stress test on lead, displaying lead  deposits in scrub region and ball track         NASA/CP-2010-216272
70 Conclusions 

The following conclusions are drawn in relation to the tested oils. These conclusions are specific to these  lubricants operating under boundary lubrication conditions at ambient temperatures.  • The lifetimes of the hydroc arbons NYE 2001 & 2001A outperform those of the PFPEs Fomblin  Z25 & Z60.  • The steady state friction yielded by the tested NYE oils is slightly less than that yielded by the  PFPEs at a given contact stress.  • The performance of NYE 2001 is similar to that of NYE 2001A.  • The performance of Fomblin Z60 is similar to that  of Z25, with the Z60 oil displaying slightly  shorter lifetimes and marginally greater steady-state friction coefficients.  • Increasing contact stress results in decrease in lubricant lifetime. 

The following conclusions are drawn in relation to the tested greases.   • Lifetimes are longer, and friction coefficients generally lower for MAC-based greases in  comparison to PFPE-based greases under these conditions.  • Little variation in lifetime was found for PFPE-base d greases, with all tested greases displaying  ~200 orbits/ μg  • Friction coefficients were found to be slightly lower for the new (-b) formulation PFPE-based  MAPLUB greases in comparison to the old (-a). Th is difference was observed to be largest in the  early stages of running  • Lifetimes of MAC-based MAPLUB greases were  an order of magnitude lower for the new  formulations in comparison to the old, suspecte d to be due to the higher viscosities of the new  greases.   • MAC MAPLUB greases containing MoS 2 thickener were found to display reduced lifetimes in  comparison to those with only PTFE thickener.  • Braycote 601EF and Braycote 601EF Micronic gave virtually identical lifetimes and friction  coefficients.  • The performances of the greases were generally similar to those of their base oils, with the  exception being the MAC based MAPLUB greases.   The following conclusions are drawn in relation to the tested solid lubricants.  • Lifetimes of lead coatings are greatly extended in comparison with MoS 2 of similar coating  thickness, due to the re-distribution of  the lubricant over the test surfaces.  • Lead displays a steady state friction coefficient ~2.5x greater than MoS 2, comparable to results  from ESTL bearing tests.  • The performance of MoS 2 is not dependent upon coating thickness within the range tested, with  no strong relationship when considering lifetime or friction.  • The lifetime of lead decreases with increasing load/contact stress.   The Spiral Orbit Tribometer performs well in such investigations. It enables comparative life and friction  assessments to be performed on a variety of lubricants on much shorter timescales than conventional ball bearing tests. 

   NASA/CP-2010-216272
71 Acknowledgements 

The author acknowledges Mr . Mark J. Jansen (Spiralab/ University of Toledo, OH ), Mr. William R. Jones,  Jr. (Spiralab/NASA Glenn Research  Center, OH) and Mr. Stephen V.  Pepper (Spiralab/NASA Glenn  Research Center, OH) for their cont inued guidance, knowledge and advice.   The author also acknowledges Mrs. Helen Hunter, Mr. Ian Pleavin, and Mr. Alan Swift (National Center of  Tribology/ESR Technology) for their help in the preparation of test samples.  The work described was performed under contract for the European Space Agency   References 

1. Anderson, M. J., ‘ Review and Identification of European Space-Compatible Oils ’, ESA-ESTL-TM0021 (2005)  2. Cropper, M. & Anderson, M. J., ‘ An Assessment of an Ultra-low Vapor Pressure Oil (Fomblin Z60) ’,  ESA-ESTL-TM-0043 01- (2007)  3. Sicre, J., et al. ‘ Perpetuation of MAPLUB Grease Range ’, Proc. 13 th ESMATS Symp. SP-670 (2009)  4. Jones, W. R., et al. ‘ A New Apparatus to Evaluate Lubricants for Space Applications – The Spiral  Orbit Tribometer (SOT) ’, NASA/TM-2000- 209935 (2000)  5. Price, W. B. & Cunningham, J. M., ‘ Boundary Lubricant Tester: Pr e-Failure Degradation of  Lubricants ’, ESTL-TM-146 (1994)  6. ESTL, ‘ Space Tribology Handbook, 4th Edition ’ (2007)  7. Jansen, M. J., et al. ‘ Relative Lifetimes of Several Space Liquid Lubricants Using a Vacuum Spiral  Orbit Tribometer (SOT) ’, NASA/TM-2001- 210966 (2001)  8. Pepper, S. V., & Kingsbury E. P., ‘Spiral Orbit Tribometery II – Evaluation of Three Liquid Lubricants  in Vacuum ’ Trib. Trans. (2003)   9. Jansen, M. J., et al. ‘ Evaluation of Non-Ozone-Depleting-Ch emical Cleaning Methods for Space  Mechanisms Using a Vacuum Spiral  Orbit Rolling Contact Tribometer ’, NASA/TM-2000-210050  (2000)  10. Marchetti, M., et al. ‘ Tribological Performance of Some Pennzane® based Greases for Vacuum  Applications’, NASA/TM-2002-211 373 (2002)   11. Marchetti, M., Jones, W. R. & Sicre, J., ‘ Relative Lifetimes of MAPLUB® Greases for Space  Applications ’, NASA/TM-2002-211875 (2002)  12. Lo, C. J., et al. ‘ Use of Cumulative Degradation Factor Predict ion and Life Test Result of the Thruster  Gimbal Assembly Actuator for the Dawn Flight Project ’, Proc. 12th ESMATS Symp. SP-653 (2007)  13. Watters, R. B. & Roberts, E. W., ‘ Torque Behavior of Low-Speed Greased Bearings ’, ESTL-TM-248  (2003)  14. Cunningham, J. & Swales, G., ‘ Evaluation of the Torque Performance of a Range of SpaceCompatible Greases as a Function of Temperature, Speed & Grease Quantity ’, ESTL-TM-197 (1997)  15. Marchetti, M., et al. ‘ Preliminary Evaluation of Greases for Space Mechanisms Using a Vacuum  Spiral Orbit Tribometer ’, NASA/TM-2001- 211157 (2001)            NASA/CP-2010-216272
73 Qualification of a High Accuracy Dual-Axis Antenna Deployment   and Trimming Mechanism 

Alain Gossant * and Francois Morichon ** 

Abstract   The Antenna Deployment and Trimming Mechanism Mark 2 (ADTM Mk2) has been developed to answer today's need for a generic antenna deployment and high accuracy pointing mechanism, allowing RF sensing applications and easier dual deployments configurations.  This paper presents the design and evolution from it s predecessor, the experi ence of the design team  from kick off to qualification and batch manufacture, as well as some lessons learned from ramping up "mass-production" capabilities while im plementing customer  driven changes. 

Introduction   Astrium has manufactured and flown ADTM units for the past 20 years, from an initial deployment-only  mechanism developed for the Orion program to today's Eurostar E3000 ADTM family. The Antenna ADTM Mk2 is an evolution of the original ADTM Mk1.  Although it uses Mk1 building blocks to minimize  risks associated with the development of a new produc t, it incorporates major evolutions and is the new  baseline for Astrium latest generation of Eurostar E3000 telecom satellites.  

   Figure 1: Left = ADTM Mk1 (one-axis), Right = ADTM Mk2 (two-axes) (without MLI fitted)   The ADTM Mk1 is used to deploy large antenna reflectors that are stowed on the spacecraft walls during launch. It is composed of an active trunnion (motive source for pitch deployment and trimming) and a passive trunnion (passive pitch deployment follower and motive source for roll trimming). Deployment is  ensured by a passive spring-loaded motor, while eac h trimming axis is powered by a stepper motor  gearbox (SMG).                                                    * EADS Astrium, Stevenage, UK  ** EADS Astrium, Toulouse, France 

Proceedings of the 40th Aerospace Mechanisms Symposium, NASA Kennedy Space Center, May 12-14, 2010  NASA/CP-2010-216272
74 The new ADTM Mk2 is made of two identical one-axi s Rotary Actuators (RAs ) mounted together (each  powered by a CDA Intercorp stepper motor gearbox ), with added antenna and spacecraft interfaces as  well as MLI enclosure. It has been developed to fulfil the following requirements difficult to achieve with  the previous Mk1 version:  3 Compatible with large 2.6+ m diameter antenna reflectors; 3 Compatible with sequenced deployments (complet e control of both deployment axes movements),  allowing dual antenna deployments. The Mk1 version was single-axis using a passive spring motor for deployment, only capable of a one-off deployment to a mechanical stop;  3 High resolution of 0.0025 °/step over the full ran ge of motion, continuous trimming / re-pointing over  the lifetime of the spacecraft, thus compatible with RF sensing applications and/or mission  reconfiguration, allowing steering the antenna beam on different Earth locations as requested by the mission management (relat ed to the operators needs);  3 Manufacturing process and number of parts are minimized.  At the time of writing this paper, the ADTM Mk2 Qualification Model has been manufactured and has completed a significant portion of its planned qualification testing.   Table 1: ADTM Mk2 Main Characteristics  Weight: 7.7 kg (includes 2 RAs, spacecraft in terface, antenna in terface, harness +  connectors, MLI)  Size: 240 x 240 x 290 mm (without connectors)  Axes: * 2 perpendicular RAs, used for deployment and trimming / pointing functions * 120° movement range on each RA  * 0.0025 °/step on the whole angular range  Operating temperatures: * Bulk temperatures: Operational: -25°C to +105°C, Survival: -55°C to +125°C * Antenna interface thermally decoupled from antenna due to a custom  thermal washer. Eurostar E3000 typica l operational temperature ranges are   -95°C to +115°C for antenna side and -10°C to +70°C for spacecraft side.  Power consumption: <15 W including thermal control (one axis command)  Available Output Torque: At least 20 N-m (including 200% margin, or 60 N-m without margin) across the whole operating temperature range.  Backlash: None up to a minimum of 7 N-m  Load Capacity (Min):  Axial Torque Shear Bending  Structure Interface 1600 N 100 Nm 1800 N 240 Nm  Reflector Interface 1600 N 100 Nm 1800 N 120 Nm 

Stiffness: >10,000 Nm/rad around deployment X and Y axes, >40,000 Nm/rad around remaining Z axis     Design   General Description

The Roll RA controls the deployment and trimming functions of the ADTM Mk2 in the spacecraft antenna  roll direction; the rotation axis is perpendicular to the spacecraft wall.  The Pitch RA controls the deployment and trimming f unctions of the ADTM Mk2 in the spacecraft antenna  pitch direction; the rotation axis is parallel to the spacecraft wall and adjusting the antenna pitch. This axis  is not fixed with regards to the spacecraft coordinate system but rotates during roll motion.   NASA/CP-2010-216272
75 

   Figure 2: ADTM Mk2 General Description 

Key Elements   • Each RA is composed of building blocks from the ADTM Mk1, but re-arranged and providing different  functions: 3 a stepper motor gearbox (SMG) from CDA Intercor p (Florida, USA) providing the active motion for  both deployment and trimming functions on Mk2, was used for trimming function only on Mk1,  3 a spring motor used as an anti-backlash device on Mk2, was performing the main deployment  function on Mk1,  3 a potentiometer monitoring the SMG output shaft (n o direct measurement of the RA output shaft)  • spacecraft interface,  • antenna interface, insulating the ADTM Mk2 from  the antenna thermally (low thermal distortion)  • thermal control   Anti-backlash

       Figure 3: Anti-backlash Spring Motor   The spring motor design here is one where springs are attached to a central torque drum with screws and naturally wound around the storage dr ums due to their curling treatment. When the torque drum is rotated  away from its rest position, an almost constant tor que is applied on the torque drum axis throughout the  motion range.  In the ADTM Mk2, the spring motor acts as an anti-backlash device, mo unted such that it will act in a  counter-clockwise direction. Backlash can so only occu r when the RA is subjected to a clockwise torque  exceeding the spring motor constant preload torque. The si ze of this spring motor is the limiting factor for Pitch RA  Roll RA   Antenna  InterfaceSpacecraft  interface   NASA/CP-2010-216272
76 the travel range; 120° range is enough for today's depl oyment and trimming requirements, and this range  can easily be extended by increasing the num ber of torque drum allowable rotations. 

Harness   The harness on the ADTM Mk2 has been designed to accommodate all ADTM Mk2 movements within the space allocated, so that the harness on the sate llite side is fixed. This has been accommodated by  creating a relaxation loop as shown in Figure 4, allo wing ±60° rotation around the RA1 roll axis (vertical  axis in Fig. 4).          Figure 4: Harness Relaxation Loop,  Left = Roll at reference position,     Right = Roll at -60° position   As seen in Figure 4, there are two fixing points, the harness being formed in a loop allowing the RA1 roll rotations. The top mounting point is a pivot point, mi nimizing harness stress within the whole movement  range. The lower bracket is a fixed line contact bracket, guiding the harness loop in the restricted available space.  3 Reference position (left picture on Figure 4): the harness is in its most relaxed position as set  during manufacture, and as shown on the left pictur e the harness is lightly stressing the RA1 roll  actuator vertically. This load is easily handled by the RA1 roll output shaft mounted on a duplex  pair bearing.  3 0 to +60° (not shown on Figure 4): the harness loop deploys but the loop itself is remaining similar  (similar radius all the way), the top pivot point following the harness thus minimizing torsional stress. The harness is acting similarly as in the reference position, without added torsional stress.  3 0 to -60° (right picture on Figure 4): although t he harness loop is following the movement and the  top pivot point minimizes stress, the bottom fixed li ne contact bracket is forcing the loop to stay in  the same location, which decreases the loop radius and therefore increases the torsion stress on the RA1 roll output shaft (spacecraft interface), acting clockwise.   Tests have been carried out using a rapid prototyped ADTM Mk2 with a proper harness routing but with sliding surfaces instead of a duplex pair bearing, and the worst-case torque measured was less than  0.3 N-m, which is not adding significant loads compared to the anti-backlash spring motor and load requirements.   NASA/CP-2010-216272
77 Design Board: The Key Ideas leading to Higher Stiffness   Achievement of high stiffness has been a key target  set in the ADTM Mk2 design specification. 

   Figure 5: Load Path 

Every intermediate part in the load path is another fl exibility introduced in the system, so the number of  parts used in the load path is minimized to achieve a high global stiffness goal:  3 At position 1 in Figure 5, the preliminary design of the ADTM Mk2 used an additional inter-RA  bracket, necessary because of the g eometry of the RA and the requir ed relative positions of the roll  and pitch RAs (perpendicular crossing axes). It has been determined by stress analysis that  growing the roll spring motor cover achieved the same  position function but with a higher stiffness.  3 Position 2 in Figure 5 designates the intermediate plate used as the SMG and spring motor  mounting bracket. The preliminary design was using a smaller spring motor cover, bolted to this plate, itself bolted to the main housing. Due to the required high loads going through the load path,  this roll RA intermediate plate was acting as a flex ible membrane, reducing the global stiffness. To  avoid this limitation, the spring motor cover ha s been extended to the main housing border, and  bolts are all directly screwed from the spring motor cover, through the intermediate plate, to the main housing. This resulted in much improved stiffness response, with an acceptable mass increase.   Specified stiffness was also achieved during design phase by further structural improvements:  3 Webs have been strategically placed to improve reinforcement of the structure in relatively high  load areas. Similarly, thicknesses have also been increased in the higher stress areas.  3 Position 3 in Figure 5 highlights the "legs" of the RAs main housing, used to mount the pitch RA  onto the roll RA. These are very similar to the ADTM  Mk1 active trunnion design. They are fairly far  apart to ensure good moment stiffness results around  the pitch Y axis; the right stiffness around the  other axes is achieved by adding two screws at posi tion 4 in Figure 5, better linking the pitch RA to  the stiff roll RA spring motor cover. This solution was much better in terms of manufacturing, mass and volume than for example having additional  brackets between the two actuators.  3 The preliminary design of the triangular plate interface with the spacecraft was flat, and with only 3  attachment points (to allow better accommodation on the spacecraft), it was acting as a membrane.  It has been reinforced as much as possible within the ADTM Mk2 volume and movements  constraints by adding a thick cylinder plus webs all around. It is now a very stiff bracket (see Table 1) comparing favorably with the duplex pair bearing stiffness. 

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