ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
419 not a critical parameter as long as the static load capacity exceeds that of the baseline superalloy and it  does (Table II).  Based upon the results from the mechanical loads and stresses review and the preliminary corrosion studies, it appears that a 60NiTi bearing utilizing the very same geometry and dimensions as the cobalt-hybrid baseline design is a good initial baseline design to consider for manufacturability. Before a manufacturing pathfinder trial can begin, a computer-generated model of the target bearing was developed. This model helps the manufacturing engineer visualize the bearing geometry and devise a manufacturing plan. The bearing model is shown in Figure 6.  

Figure 6.  Computer model of DA bearing.   The manufacture of 60NiTi balls is a fully commercialized process and many standard ball sizes are available. The steps needed to make precise raceways from 60NiTi to a desired geometry are not yet fully developed. Decades of experience with various steels has resulted in numerous rules-of-thumb (ROT) for their manufacture. Preferred annealing, rough machining, hard turning, heat treatment, grinding, polishing, acidic surface passivation, normalizing and other processes for steels are well understood and accepted by the manufacturing community. For 60NiTi these ROT’s remain undefined. The following paragraphs attempt to layout an acceptable, by not necessarily an optimal, processing path driven by 60NiTi’s several unique properties and characteristics, namely its low elastic modulus, resistance to metal cutting and deformation (i.e., low ductility), and its intrinsically brittle tensile behavior.  The basic manufacturing steps employed to go from ingots to finished bearings are as follows: rough machining to near final shape and size using a combination of wire electro discharge machining (EDM) and carbide tool based machining, heat treating to develop high hardness, finish grinding and polishing. Because 60NiTi exhibits limited ductility and is susceptible to brittle tensile behavior, heavy machining operations such as drilling, high removal turning and milling should be avoided. Instead, we have successfully employed a combination of plunge and wire EDM to generate simple blanks from larger ingots and these are then rough machined to near final shape using more conventional machining methods. Figure 7 shows a slice of 60NiTi made by wire EDM and subsequently cut into cylindrical (donut
420 shaped) race blanks. The ingot, made by powder metallurgy has a rind and center core of steel with  60NiTi filling the annular space. This is part of the PM process and enables more efficient use of the relatively high cost 60NiTi materials.  

Figure 7.  60NiTi race blanks cu t from ingot using EDM processes.   The resulting ring blanks are then shaped into race rings using properly shaped carbide tooling through a plunge turning operation. With this operation, the ring inside diameter is first turned or ground to ensure it is uniform and close to finish dimension (within ~0.01 mm). The ring is then mounted on an expanding type mandrel and care is taken to avoid excessive clamping forces. Because the elastic modulus is half  that of steel, excessive clamping force on the inside diameter can result in considerable stretching of the ID leading to unpredictable dimensions after turned race is removed from the mandrel. Once mounted, the OD is plunge turned using a carbide tool previously shaped to match the desired race ring profile. In our tests, we used wire EDM to manufacture the plunge tool from a standard carbide lathe bit. Figure 8 shows the process for turning an inner race ring prior to the heat treatment step.
421  

Figure 8.  Plunge turning of 60NiTi race rings using a carbide tool profiled to match the desired  race geometry using wire EDM. 

The final steps include heat treatment and finish grinding and lapping. The heat treatment has been  described in detail in reference 6 and briefly includes a solution treatment at 1000 C followed by a rapid  water quenching. A vacuum or inert atmosphere is used to prevent excessive surface oxidation. Figure 9  shows the microstructure of 60NiTi prior to and after heat treatment. The heat treatment dissolves undesirable higher order phases resulting in a largely homogenous NiTi phase structure with hardness in the range of 58 to 63 on the Rockwell C scale.
422  

Figure 9.  Microstructure of 60NiTi before (left) and after heat treatment (right). 

 After heat treatment, standard industry grinding and polishing are employed to yield a finished bearing. Care must be taken to avoid excessive tensile stresses during bearing assembly. Unlike steel that can endure high tensile deformation without fracture, 60NiTi is brittle in tension and outer races, in particular, can break during bearing assembly when hoop stresses are high. To overcome this problem, either separable bearing designs or the use of differential inner ring cooling and outer ring heating are recommended. Beyond that, no other special considerations are needed. The bearings behave and perform as one expects based upon the geometry and lubricant selected. Future efforts to fully characterize the rolling contact fatigue stress-life relationships and gain a better understanding of the tensile strength and toughness behavior of 60 NiTi are planned to guide more mechanically demanding bearing applications.   Summary Remarks 

60NiTi offers a viable path towards rolling element bearings that have exemplary corrosion and wear resistance. Though sensitive to brittle fracture in tension, research suggests that extreme static load capacity and resilience imparted by its superelastic behavior is possible. The reduced elastic modulus, compared to steel, may impact highly loaded bearing internal geometry but for the DA bearing application being considered, a direct 60NiTi replacement is reasonable. Lastly, manufacturing research underway suggests that with proper care, modern manufacturing methods are capable of producing high precision ball bearings from 60NiTi.   Acknowledgements 

The authors wish to acknowledge the support of the NASA Engineering and Safety Center (NESC) and the NASA Fundamental Aeronautics Program (Subsonic Rotary Wing Project). Their continued funding and technical and programmatic guidance has contributed greatly to bringing this new technology from the lab and closer to commercialization.
423 References 

1.  C. DellaCorte, R.D. Noebe, M.K. Stanford, and S.A. Padula: “Resilient and Corrosion-Proof Rolling  Element Bearings Made From Superelastic Ni-Ti Alloys for Aerospace Mechanism Applications,”  NASA TM-2011-217105, September 2012.  2.  C. DellaCorte, S. V. Pepper, R. D. Noebe, D.R. Hull, and G. Glennon: “Intermetallic Nickel-Titanium  Alloys for Oil-Lubrication Bearing Applications,” NASA TM-2009-215646, March 2009.  3. S. V. Pepper, C. DellaCorte, R. D. Noebe, D.R. Hull, and G. Glennon: “Nitinol 60 as a Material for  Spacecraft Triboelements,” ESMATS 13 Conference, Vienna, AUstria, September 2009.  4. A. Palmgen: ”Ball and Roller Bearing Engineering, Chapter 8, ”Bearing Failures,” pages 217-225, 3rd  Edition, SKF Industries, Philadelphia, PA, 1959.  5. M.D. McNeese, D.C. Lagoudas, and T.C. Pollock: ”Processing of TiNi from Elemental Powders by  Hot Isostatic Pressing,” Materials Science and Engineering, A280 (2000), pages 334-348.  6.  M.K. Stanford, F. Thomas, and C. DellaCorte: ”Processing issues for Preliminary Melts of the  Intermetallic Compound 60-NiTiNOL, NASA TM- 2011-2 xxxx, A ugust 2011.  7. D.L. Carter: ”Status of the Regenrative ECLSS Water Recovery System, SAE 2009-01-2352, 39th  International Conference on Environmental Systems,  July 12-16, Savannah, GA, 2009.  8. C.D. Wingard: ”Compatibility Testing of Polymeric Materials for the Urine Processor Assembly (UPA)  of International Space Station (ISS), NASA report 20030106448-2003127090, NASA Marshall Space Flight Center, Huntsville, AL, 2003.  9. D.W. Holder and C.F. Hutchens: ”Development Status of the International Space Station Urine  Processor Assembly,” SAE 2003-01-2690, 33rd International Conference on Environmental Systems, July 7-10, Vancouver British Columbia, 2003.  10. T.L. Krantz: ”On Calculation Methods and Results for Straight Cylindrical Roller Bearing Deflection,  Stiffness and Stress,” Proceedings of the ASME 2011 International Design Engineering Technical Conferences & Computers and Information in Engineering Conference IDETC/CIE 2011, August 29-31, 2011, Washington, D.C., paper number DET2011-47930.
427 prevalent remaining constituent in the chamber during testing is water vapor as determined by residual  gas analyzer [3]. Figure 5 provides a simplified schematic of the test rollers labeled with some of the nomenclature used herein. 

Figure 3 – Vacuum roller rig. 

A set of sensors on the test apparatus monitors the test conditions. The outputs of the analog sensors were digitized and stored via a data acquisition system at a rate of up to 0.66 hertz. Each of the sensors will be described in turn. 

The misalignment of the driving roller shaft and driven roller shaft is depicted in an exaggerated manner  in Figure 4(b). The misalignment is measured via a linear variable differential transformer (LVDT). The transducer housing is attached to the bedplate, and the translating, spring-loaded transducer tip contacts against a mechanical stop on the turntable. To establish the aligned condition, special tooling blocks were machined to locate the roller-mounting surfaces of the two shafts as parallel. With the shafts aligned by the tooling blocks, the transducer circuit balance was adjusted to provide an output of zero. The precision of this method for aligning the shafts was limited by the dimensions of the roller mounting surfaces used as the reference planes. From analysis of the test rig drawing tolerances and geometry, the alignment procedure using the tooling blocks to define the zero-degree position has a precision within 0.08 degrees. Rotation of the turntable from the aligned position moves the LVDT sensor. The angular displacement of the turntable was determined by mounting a laser light source on the moving shaft at the roller mounting location and directing the light onto a paper placed at a known radial distance from the center of the turntable. The movement of the laser light was marked on the paper and distance between the points measured and used to relate the sensor output to the angular motion of the turntable.  The torque on the output shaft is monitored by a strain-gage type torquemeter of 22 N-m (200 in-lb) torque capacity. Calibration was done in place using deadweights acting on a torque arm of known length attached at the test roller position and reacting the output shaft to ground.  The load pressing the rollers together is termed herein the “normal load” (Figure 5). The normal load is applied via an air-pressure actuated piston. The air piston acts through a load cell against the drive motor plate that is gimbal-mounted relative to the test chamber (Figure 4(a)). In this way the air piston moves the roller on the input shaft in an arc motion toward the test roller. The arc motion of the input shaft table is measured by an LVDT. Once the rollers are in contact, additional force commanded to the air piston increases the normal load between the test rollers. Careful calibration processes allow calculating the normal load on the test roller based on the sensor outputs from the load cell and the LVDT that measures the input shaft position.
429  (a) 

 (b) 

Figure 4 – Schematic views of the vacuum roller rig. (a)  Schematic, side view.  (b)  Schematic, overhead view with shaft misalignment depicted and exaggerated. 

Figure 5 – Simplified schematic view including some of the important sensed data.  (a)  Schematic, front view. (b) Schematic, side view.   ferrofluid seals and bearingstorque/force sensor brakepivot point for drive motor  plate relative to bed plate drive motor bedplate hydraulic cylinder“normal load”  force sensorvac chamber turntable bedplatebellows ferrofluid seal and bearingtorque/force sensor brakedrive motor bedplateturntable on bearingload cell for “tangential” force, mounted to turntable normal load braking torquetangential force normal loadnormal load braking torquetangential force normal load normal load axial force normal loadinput shaft output shaft normal load axial force normal loadinput shaftinput shaft output shaft (a) (b)normal load braking torquetangential force normal loadnormal load braking torquetangential force normal load normal load axial force normal loadinput shaft output shaft normal load axial force normal loadinput shaftinput shaft output shaft (a) (b)
431  

Figure 6 – Typical surface roughness of test rollers, profile traces obtained using a 2-micrometer  radius tip conisphere stylus traci ng in the direction of rolling.  (a) Ti6Al4V roller.  (b) 440F roller.  (c)  440C rollers. 

Contact analyses were completed to select roller normal loads, speeds, and alignment angle and to relate  those choices to the operating conditions of the qualification test unit rollers. The contact conditions for the qualification test were studied assuming that the contact pressure distribution was affected by the moment produced by the axial load caused by misalignment. To estimate this effect, it was assumed that the ratio of axial load to normal load was 0.6 and such load produced an overturning moment by acting through the ball bearing center. It was also assumed that the supporting structure provided 90 percent of the reaction to the overturning moment while the pressure distribution provided 10 percent of the reaction. The pressure distribution for such a loading condition while accounting for the roller profile having a crown but modified with a flat section was calculated using the method of Vijayakar [9-10]. The contact condition for a perfectly aligned roller having zero axial force was also solved. The predicted contact pressure distributions are provided in Fig. 7. The misalignment of the roller axis will cause a shift of the pressure from the central flat section onto the crowned region, and a maximum contact pressure of about 450 MPa occurs near the transition from flat to crown geometry. Because of test rig limitations, the laboratory test conditions could not match these conditions exactly. The laboratory testing was done at maximum contact pressure of about 770 MPa. The laboratory tests were used to study trends and fundamental qualitative wear behavior. 

The test rig speed was selected using the idea that the “contact time” of the test unit and laboratory rollers  should be matched. Here the “contact time” is the time required for a point on the roller to pass through the Hertz contact region. The contact time on the qualification test unit for condition of misaligned rollers was 0.015 second. To match this condition the test rig was operated at 1.6 rad/s (15 rpm) for the majority of laboratory testing.  -3-2-10123 (a)height (micrometer) -3-2-10123 (b) profile position (mm)02468 1 0 1 2-3-2-10123 (c)
434 (a) (b) (c) (d)   588 cycles 5,965 cycles 25,158 cycles 77,357 cycles  (e) (f) (g) (h)   402cycles 6,140 cycles 24,468 cycles 96,193 cycles  (i) (j)   (k)  (l)   500 cycles 93,854 cycles 250,196 cycles    347,925 cycles  Figure 8 - Progression of wear during 3 tests. The cycle count is denoted below each photo.  (a-d) 440F at 1.4° misalignment. (e-h) 440C at 1.5° misalignment. (i-l) 440F at 0.0° misalignment.
435 (a) (b)  (c) (d)  Figure 9 – Scanning electron inspections of the wear track on the titanium test roller for test #3.   (a) SEM image. (b) Spectrum for region noted as 1 in 9(a).  (c) Spectrum for region noted as 2 in  Fig. 9(a).  (d) Spectrum for region noted as 3 in Fig. 9(a). 

The roller masses were determined using a scale with digital readout to 0.0001 gram. Masses of the rollers were measured after cleaning and just before installation in the test rig. The roller masses were also measured just after removal from the test chamber. The change in mass at the end of the test quantifies the net transfer of material by adhesive wear. The sum of the mass change for the two rollers provides a calculated value of debris lost from the roller pair. Because of practical test considerations, the test durations were not the same for each test. To provide a method for direct quantitative comparison, the change in mass was divided by the total number of revolutions of the input shaft (that is the total number of contact stress passes) to quantify the wear rate. The mass change data and calculated wear rates are summarized in Table 2.  Often wear is modeled as proportional to the sliding distance. In these roller tests, to a first order effect the sliding distance is proportional to the misalignment angle. The rate of mass change was plotted as a function of misalignment angle (Fig. 10). Note that in the net, the steel rollers lost mass while the titanium alloy rollers gained mass. We also note that the 440F material had a higher rate of wear compared to the 440C in all cases. Rabinowicz [8] has noted that the experimental evidence for wear rate being proportional to sliding distance (as often assumed true) is mixed. He states that usually the relationship is not perfectly obeyed but the proportional relationship represents experimental data “reasonably well”. This observation matches the behavior from this investigation in that the misalignment angle (and thereby the sliding distance) correlates to the wear rate, but the linear relationship is not exact. Not only the wear rate but also the wear behavior differed for tests with 440F steel vs. 440C steel. Adhesive wear tends to progress toward an “equilibrium surface” [8] meaning the surface roughness changes with running, and
436 the wear process can either roughen or smoothen the surfaces depending on the materials and starting  conditions. The equilibrium surfaces for the tests with 440F vs. 440C differed as depicted in Fig. 11. Tests with the 440C material resulted in smoother running surfaces (Fig 11(b)), and the photos suggest that the particle size of wear debris is likely smaller, in general, for the case of 440C. The Ti6Al4V rollers when mated with the 440F steel tended to gain material by adhesive wear in a more uniform fashion, that is, the coverage of the surface was more significant and uniform in appearance. It is possible that the hardened 440C behavior differed from that of the annealed 440F not only because of the change of the elastic limit but also because of an alteration of the metallurgical structure and alteration of the adhesion compatibility with the Ti6Al4V [8].   Table 2 – Summary of Roller Wear Test Results    

   Figure 10 – Change of mass per cycle (shaft revo lution) as a function of misalignment angle. (a)  Mass loss of 440 steel rollers. (b) Mass increase of Ti6Al4V rollers, symbols denoting the mating material.     test durationmass change;  Ti6Al4V mass change;   440 steel calculated  mass liberated calculated rate of  mass liberated test # test IDrevolutions of  input shaftgrams grams milligramsmicrograms per  cycle 1 1 26,080 0.0049 -0.0054 0.500 0.019 2 2 60,228 0.0981 -0.1258 27.700 0.460 3 3 77,326 0.1350 -0.2072 72.200 0.934 4 4 347,925 -0.0006 -0.0052 5.800 0.017 5 5 55,411 0.0817 -0.1052 23.500 0.424 6 6 60,850 0.0151 -0.0210 5.900 0.0977 7 124,900 0.1314 -0.1664 35.000 0.280 8 8(a) 93,027 -0.0005 -0.0038 4.300 0.046 9 8(b) 24,001 0.0012 -0.0012 0.000 0.000 10 8(c) 24,000 0.0017 -0.0018 0.100 0.00411 8(d) 24,095 0.0025 -0.0027 0.200 0.00812 8(e) 93,117 0.0464 -0.0501 3.700 0.040 13 9 96,193 0.0682 -0.0698 1.600 0.017 14 10 41,055 0.0200 -0.0210 1.000 0.024 15 11 90,004 0.0672 -0.0688 1.600 0.018 misalignment angle (degree)-1.6 -1.2 -0.8 -0.4 0.0mass change  (microgram / cycle) -3.0-2.5-2.0-1.5-1.0-0.50.0 440F 440C (a) misalignment angle (degree)-1.6 -1.2 -0.8 -0.4 0.0mass change  (microgram / cycle) 0.00.51.01.52.02.53.0440F 440C (b)
437 (a)  (b)  (c)  (d)  Figure 11 – Condition of rollers at end of test showing differing behavior of 440F vs. 440C.  (a) 440F, test ID 5.  (b) 440C, test ID 9. (c) Ti6Al4V test ID 5. (d) Ti6Al4V, test ID 9.  The test ID numbers refer to Table 1. 

Wear debris.  During testing of the qualification test unit, some loose debris was formed. The scope of the NESC  assessment included consideration of mission risk from the loose debris. The laboratory testing revealed aspects of the loose debris as follows. From Table 2, note that loose debris was generated at rates on the order of 0.5 micrograms per cycle for large roller misalignment. The wear rates were greatly reduced for the case of zero roller misalignment. However, the adhesive wear was not completely eliminated with aligned rollers, and some debris was formed. Figure 12 (a-b) provides images of the largest sized particles collected from the two tests operated with zero roller misalignment. Figure 12 (c) is a typical example of the large number and widely distributed debris that occurred for extensive run time and the largest misalignment angles tested. A portion of the debris pan was not within direct line of sight of the rollers, but that region still contained significant numbers of particles.  

Summary 

Laboratory testing of 440 steel rollers in contact with Ti6Al4V in vacuum was completed to study the wear  behavior and to quantify wear rates. The tests also assessed the influence of material condition and roller misalignment. The wear rate was found to be roughly proportional to the roller misalignment angle. The rate of loose particle mass created was on the order of 0.5 micrograms per cycle. The adhesive wear was
438 a two-way phenomenon with titanium alloy material found adhered to the steel and vice versa. Loose  particles with linear dimensions on the order of 200 micrometer were created even for the condition of zero roller misalignment. For large misalignment angles, large numbers of loose particles could be created, and debris was found in locations that were not within a direct line of sight. The wear behavior differed for the case of 440F and 440C steel rollers. With the 440C rollers, the adhesive wear rate was reduced, and the final surface textures of the 440C rollers were smoother than the 440F rollers.    Figure 12 - Examples of loose particle debris.  (a) Test ID 4 with zero roller misalignment. (b) Test  ID 8(a) with zero roller misalignment. (c) Test ID 5, rollers misalignment -1.4 degrees.  The test ID numbers refer to Table 1. 

Acknowledgements 

This research was supported by the NASA Engineering Safety Center. The experimental work was done  with assistance from Mr. Richard Manco, Sierra Lobo Inc. Dr. Iqbal Shareef’s contributions were supported by the NASA GRC Summer Faculty Fellowship Program. 

References 

1. McClendon, M., “NIRSpec MSS Magnet Actuator Life Test Unit Wear Particle Evaluation”,  obtained by Krantz, T., May 16, 2011  2. Authors unstated, “Micro Shutter Subsystem (MSS) Qualification Unit Test Report”, JWST-RPT013819, Rev. A, June 2010.  3. Pepper, S., “Research Note-Characterization of the Test Environment of JWST Roller Wear  Evaluation at NASA-GRC”, Aug. 1, 2011.  4. Johnson, K.L., Contact Mechanics , Cambridge University Press, 1985.  5. Kalker, J.J., “Rolling contact phenomena: linear elasticity”, Rolling Contact Phenomena CISM  Courses and Lectures, Issue 411, Spinger-Verlag, 2000.  6. McGinness, H., “Lateral forces induced by a misaligned roller”, DSN Progress Report 42-45,  March and April 1978, Jet Propulsion Laboratory, Pasadena, CA, 1978.  7. Krantz, T., DellaCorte, C., Dube, M., “Experimental Investigation of Forces Produced by  Misaligned Steel Rollers”, proceedings of the 40 th Aerospace Mechanisms Symposium,  NASA/CP-2010-216272, also NASA/TM-2010-216741, 2010.    8. Rabinowicz, E., Friction and Wear of Materials , 2nd edition, Wiley-Interscience, 1995.  9.  Vijayakar, S., “A combined surface integral and finite element solution for a three-dimensional  contact problem”, International J. of Numerical Methods for Engineering, vol. 31, 1991.  10. Vijayakar, S.  “Multi-body Dynamic Contact Analysis Tool for Transmission Design – SBIR  Phase II Final Report”, Army Research Laboratory  ARL-CR-487, 2003.  (a) (b) (c)