Friction, Lubrication, and Wear Technology (1997) Part 5 pps

calomel electrode. This system used a recycled slurry that was replenished with fresh slurry periodically to minimize the effects that were due to slurry particle rounding. At jetting speeds ranging from 9.3 to 21.5 m/s (30.5 to 70.5 ft/s), the T (C,W) from both a high-carbon steel and a high- chrome cast iron impinged by a quartz slurry were mainly influenced by the S' (C,W) . This synergistic effect accounted for 8.3 to 54.9% of the total material losses, with the greatest effect occurring for larger particles and higher jet velocities. Closed-Loop Pipeline Experiments. Postlethwaite et al. (Ref 26) measured the electrochemical corrosion rate of steels in a closed-loop pipeline experimental apparatus. He mounted electrode specimens flush with the pipe wall. The test section of pipe was a 38 mm (1.5 in.) diam vertical leg of the test loop. The vertical orientation of the pipe eliminated any effects of gravity on the solids concentration in the pipe. The value for T (C,W) was obtained by weight losses, whereas C w was obtained by polarization resistance techniques. The values obtained for W 0 (during cathodic protection of the electrode) were less than 5% of the total material losses, and S' (C,W) varied between 2.8 and 40.9% of the total material losses. Most of the material losses from a commercial carbon steel pipe were due to electrochemical corrosion. Depending on the slurry flow rate, C w varied from 56 to 92% of the T (C,W) . It was shown that the erosive effect prevented the formation of a rust film that normally stifles the diffusion of oxygen to the corroding surface. Grinding Wear Systems The abrasive wear in grinding is different from that experienced in slurry particle impingement systems, because the abrasive particles from the crushed ore participate in three-body abrasion between two solid surfaces. Rotating Cylinder/Anvil Experiments. Kotlyar, Pitt, and Wadsworth (Ref 36) used a rotating cylinder/anvil apparatus to measure the corrosive wear of a HCLA steel. The apparatus consisted of a rotating cylindrical specimen that rotated between two opposing abrasive anvils that were applied to the cylinder to a controlled load. The specimen served as a working electrode in a three-electrode system. Values of T (C,W) were determined by weighing the specimens before and after an experiment. To estimate W 0 , cathodic protection was applied to the sample by fixing the voltage at -1.7 V versus a saturated calomel electrode (SCE). The electrochemical corrosion rate was determined using a linear polarization technique. In experiments using HCLA steel rotating cylinders in a 15 wt% quartz slurry (pH = 9), total material losses were higher by a factor of four when quartz, rather than HCLA, anvils were used. As shown in Fig. 11, the increase in total material loss, T, was linear with applied load; corrosion, C w , was nearly constant with load; the purely mechanical wear, W 0 , increased and then leveled off as a function of load; and S' (C,W) was large and variable (Ref 37). The synergistic contribution of corrosion to abrasion increased markedly with loads higher than 65 N (6.6 kgf). This may result from the initiation and propagation of microcracks, assisted by anodic corrosion. The surface cracking caused by pitting and subsurface cracks (Ref 38) also has been observed after wet grinding in laboratory ball mills. The synergistic effect was found to decrease linearly with respect to pH in the range from 7 to 10. Above a pH of 10, pitting is retarded, and the metal loss that results from synergism decreases. Fig. 11 Total wear rate, T (C,W) ; mechanical wear rate, W; corrosion rate during wear, C w ; and the synergistic term, S' (C,W) , for HCLA steel using quartz anvils in a 15% quartz slurry at pH 9.0 as a function of load. Source: Ref 37 When localized corrosion is present, the corrosion component is small, whereas the synergistic component can be as high as 50% of the total material losses. The large synergism is accounted for by macroscopic removal of metal, resulting from localized corrosion. Rotating Ball-on-Electrode Experiments. In investigations of the electrochemistry of complex sulfide-grinding systems, the galvanic currents of pyrite, pyrrhotite, and mild steel electrodes were determined in short-circuit connections under abrasive and nonabrasive conditions (Ref 39, 40). The device shown in Fig. 12was developed for the tests performed under abrasive conditions. It consisted of a porcelain ball rotating against three fixed electrodes of pyrite, pyrrhotite, and mild steel. Quartz slurries prepared with mild steel balls were used, and the galvanic current for each electrode was separately and alternately monitored via a switch with a zero-resistance ammeter. Fig. 12 Experimental ball-on-electrode device to measure galvanic cur rents between minerals and alloys. Source: Ref 38 Figure 13 depicts an electrochemical model for a two-sulfide/grinding media system according to the corrosion theory for a multielectrode galvanic cell (Ref 41). In this system, the noblest electrode is generally a sulfide mineral that is cathodic. The grinding media is usually more active than the sulfide minerals and is anodic. The other sulfide mineral develops an intermediate cathodic or anodic behavior, depending on its rest potential, the electrochemical characteristics of the electrodes, its surface area, and the material that it contacts. Fig. 13 Model of galvanic interactions among two minerals and grinding media. Source: Ref 39 Major reactions that are thought to take place at the surface of the respective mineral or steel electrode are (Ref 39): O 2 + H 2 O + 2e - = 2OH - (at cathodic mineral) (Eq 5) (Eq 6) Fe = Fe 2+ + 2e - (corrosion of steel media) (Eq 7) (Eq 8) Dissociation of the minerals may also occur, but this is usually a slow process. The precipitation of other metal oxides is possible when dissolved metal ions from the minerals combine with hydroxide ions in a manner similar to Eq 8. By using electrochemical and surface analysis techniques, Adam et al. (Ref 42) studied the pyrrhotite-grinding media system. They found a coating of hydroxide and/or sulfate species of iron on the pyrrhotite surface as a result of galvanic interactions (Eq 5, 7, 8) between the mineral and the grinding media. Similar findings were reported by Learmont and Iwasaki (Ref 43) for the galena-grinding media combination. The iron hydroxide coatings on pyrite produced by reactions (Eq 5, 7, 8) can impair floatability, whereas local pH changes that are due to acid formation via oxidation of elemental sulfur to sulfate and hydrogen ions at the pyrrhotite surface (Eq 6) can prevent the formation of the detrimental coating. Corrosion currents that are due to the galvanic coupling between the sulfide minerals can be determined by obtaining polarization curves of individual electrodes under abrasive conditions, and by determining the point at which the cathodic curve crosses the anodic curve. These curves must be adjusted for the cathode/anode surface areas (Ref 44). Estimates based on this technique have shown good agreement with estimates from marked-ball water test data (Ref 9). Mechanisms of Wear/Corrosion Synergism in Abrasive and Impact Wear At stated earlier, Dunn (Ref 1) has indicated a number of mechanisms that account for wear/corrosion synergism in mineral processing systems. The following discussion applies the proposed mechanisms to various materials and handling systems. Abrasion. The plastic deformation by high-stress metal-mineral contact causes strain hardening and susceptibility to chemical attack. This was shown by Kotlyar and Wadsworth (Ref 27), who demonstrated the existence of localized electrochemical cells between strained and unstrained regions of alloys, which promoted pitting. The abrasion mechanism itself can: • Remove protective metal oxides and passive films to expose unoxidized metal to a corrosive environment. This is true for high-speed slurry wear of stainless steels (Ref 33 ). The corrosion rate of 316 stainless steel was shown to be as high as mild steel in the presence of only 2 wt% silica sand slurries • Form microscopic grooves and dents that serve as sites for concentration cell corrosion. This can happen to many alloys, including HCLA steels, in which the formation of microgrooves and partial removal of oxide films lead to the initiation of pits that affect the total material losses (Ref 27) • Increase the microscopic surface area exposed to corrosion, thus increasing the corrosion current • Remove strain-hardened surface layers that are caused by repetitive impact. These layers are often more brittle than the underlying metal and do not adhere to the surface as well • Crack brittle metal constituents, forming sites for impact hydraulic splitting. As corrosive liquids are trapped in cracks, forces on the metal surface seal the crack containing the liquid, and then hydraulically split adjacent metal to propagate the crack. Corrosion, as a mechanism, can: • Produce pits that can be precursors to microcracking, which invites hydraulic splitting during impact • Roughen the surface, thus leading to lower energy requirements to abrade the metal • Produce hydrogen, which can embrittle the metal and lead to cracking (Ref 45) • Selectively attack grain boundaries and less- noble phases of multiphase microstructure, weakening adjacent metal. This is especially true of white cast irons, where corrosion attacks the chromium- depleted zones that surround the hard carbides in the structure. This corrosion eventually promotes the premature removal of the harder carbide phases, thus accelerating material losses Impact. Plastic deformation can make some constituents more susceptible to corrosion. The impact mechanism itself can: • Crack brittle constituents and tear apart ductile constituents to form sites for crevice corrosion and hydraulic splitting • Supply the kinetic energy to drive the abrasion mechanism, thus acce lerating the abrasive component of wear. In ore processing, some minerals are driven into the softer steel grinding media, promoting the formation of galvanic corrosion cells • Pressurize mill water to cause splitting, cavitation, and jet erosion of metal a nd protective oxidized material • Pressurize mill water and gases to produce localized temperatures and phase changes in the liquid • Heat grinding media, mill liners, ore, and fluids to increase corrosive effects Means for Combating Corrosive Wear Research and engineering efforts in wear, corrosion, and materials science are being pursued to combat corrosive wear in aqueous environments. The areas of investigation described below include materials selection, surface treatments, and handling environment modifications. Materials Selection. The selection of the right material for a particular corrosive wear environment can lead to extended life of component parts, less costly downtime, and other economic advantages. One approach is to place a number of materials in actual service and to compare the material losses of each over a given time. This technique can give the best solution to materials selection, but is time consuming. It also limits the number of materials that can be tried, and does not often result in the application of scientific principles to obtain the most cost- effective material for the given life of a process. However, actual field service should be used as a concluding step before final choices of materials are made. Another approach is to adapt laboratory tests to field situations. Isaacson et al. (Ref 46) used a method in which electrodes were mounted at the ends of mill liner bolts, and polarization curves of the media material electrodes were obtained by a telemetry-radio system. Although the data were often noisy, correlations with laboratory corrosive wear tests were made. Madsen (Ref 47, 48) used a test developed in the laboratory in a field experiment. He converted a laboratory slurry pot test to a portable model that could be transported to the field for use with slurries there. This approach enabled the testing of a large number of candidate materials in a relatively short time, because this slurry pot design allowed for the simultaneous comparison of 16 specimens. The equipment could be useful in screening materials for further evaluation in actual service, thus shortening the time needed to determine the better material while offering a wider selection than actual field use could. This test is most useful when the hydrodynamic and geometric parameters in the test chamber are similar to those experienced in the actual service conditions. A more fundamental and scientific approach is to study the wear and corrosion characteristics of various materials in the laboratory and extrapolate these results to actual field use. A great deal of knowledge and modeling of wear, corrosion, 0and wear-corrosion synergism is needed, along with a detailed model of the hydrodynamics. This research is still in its infancy, but is making progress as the number of researchers who can recognize wear-corrosion synergism increases. Surface Treatment. Rather than use the more-corrosive-resistant material to make the entire component, one may choose to use a thinner layer of a costly material on the surface of the more economical substrate material. Claddings, surface treatments such as hard facings and patching with welds, or replaceable liners have all been used. In pipelines where corrosion is a problem, it is sometimes economical to line the pipes with a smaller-diameter polymer that is more corrosion resistant, but lacks the strength to withstand high pressures on its own. Modification of the materials handling environment can be effective in controlling the corrosive wear of component parts. Solution conditioning, such as adjusting the pH and deaeration, can reduce the amount of material losses in a corrosive-wear environment. Slurry conditioning is not economical for the short slurry lines used in mining operations unless some method of water separation and recirculation is used (Ref 19). Use of Corrosion Inhibitors. Oxidizing inhibitors, such as chromates and nitrites, have been used to raise the potential of an alloy into passive regions and to lower their corrosion rates (Ref 49). These inhibitors form a surface film when used in high concentrations. Chromates also act as effective inhibitors when used at concentrations much lower than those required to produce active-passive transitions. At these lower concentrations, chromates act as cathodic inhibitors in neutral solutions. However, because chromates are toxic, alternative nontoxic inhibitors with a similar action should be considered. Many governments restrict the effluent limit to levels as low as 0.05 ppm chromate, which makes the choice of chromate inhibitors unacceptable. Progress in developing nonchromate inhibitors has been slow (Ref 19). Slurry Parameters. Reduction of the slurry velocity is a major factor in controlling the rate of material losses if the mechanical wear, W, is important, because the wear rate generally varies with the velocity raised to exponents of 2, 3, or 4. For slurry pipelines that carry 20% silica sand (30 × 50 mesh), very little abrasive wear occurs below a velocity of 6 m/s (20 ft/s) (threshold value). Above this velocity, mechanical wear becomes more dominant as a means of material degradation. Particle size can also be a factor in the wear of slurry handling equipment. The mechanical wear, W, will not be a problem if the particle size is sufficiently reduced so that the particles are fine enough to follow the streamlines of the solution, rather than impact the walls of the containment part. Postlethwaite points out, however, that fine slurry particles and low velocities may result in conditions mild enough to permit the growth of rust films and scale, which can lead to pitting (Ref 19). The particle size and velocity combination should be maintained below the threshold value, where W is not a factor, but with a particle size that causes enough abrasion to maintain a rust-and scale-free pipe surface. In this condition, however, the pipe should be protected from corrosion by inhibitors and/or deaeration. This would eliminate the need for unnecessary size reduction. Cathodic protection is a useful method for protecting short sections of slurry pipelines, pumps, elbows, and other equipment (Ref 50). However, the length of the protection must be kept sufficiently short to prevent overprotection and subsequent hydrogen blistering of the protected surface. The throwing power of cathodic protection, either by impressed current or galvanic anodes, is insufficient for this method to be used to protect the inside of slurry pipelines. Sacrificial electrodes, such as zinc plugs in water or slurry pumps, have been used to cathodically protect the pump casting. This means of protection may be applicable to other instances where it is economically feasible. Design of Pumps, Valves, Elbows. Research by Roco et al. (Ref 51, 52, 53, 54) and Ahmad et al. (Ref 55) has shown that computer modeling is becoming a tool for the design of wear-resistant slurry pumps. A computer code is used to carry out the numerical flow simulations within the pump channels, and an energy-based, wear-predictive model has allowed for the prediction of wear rates for various geometries. The use of these types of models can extend the wear life of pumps by altering the geometry of the interior of the pump. This effectively removes the areas of high wear rates, and spreads the energy absorbed by the pump head evenly around its interior. This same type of modeling has been done by Postlethwaite (Ref 56) and Nesic and Postlethwaite (Ref 57, 58, 59) for areas of disturbed flow. Other means of controlling corrosive wear are available, including: • Increasing the pipe diameter in order to decrease the slurry velocity and help ensure laminar flow • Increasing the thickness of materials in critical areas • Inserting impingement plates or baffles to shield critical areas from high wear • Directing the inlet flow of materials to avoid high particle velocities at the wall of containment vessels References 1. D.J. Dunn, Metal Removal Mechanisms Comprising Wear in Mineral Processing, Wear of Materials, K.C. Ludema, Ed., American Society of Mechanical Engineers, 1985, p 501-508 2. A.M.F. Carter and D. Howarth, "A Literature Review of the Factors that Influence the Wear of Slurry- Handling Systems," Report M319, Council for Mineral Technology, Sept 1987 3. F.H. Stott, J.E. Breakell, and R.C. Newman, The Corrosive Wear of Cast Iron Under Potentiostatically- Controlled Conditions in Sulfuric Acid Solutions, Corros. Sci., Vol 30, 1990, p 813-830 4. R.L. Pozzo and I. Iwasaki, Pyrite- Pyrrhotite Grinding Media Interactions and Their Effects on Media Wear and Flotation, J. Electrochem. Soc., Vol 136 (No. 6), 1989, p 1734-1740 5. National Materials Advisory Board, "Comminution and Energy Consumption," NMAB- 364, National Academy Press, 1981 6. K. Adam, K.A. Natarajan, S.C. Riemer, and I. Iwasaki, Electrochemical Aspects of Grinding Media- Mineral Interaction in Sulfide Ore Grinding, Corrosion, Vol 42 (No. 8), 1986, p 440-446 7. R.L. Pozzo and I. Iwasaki, Effect of Pyrite and Pyrrhotite on the Corrosive Wear of Grinding Media, Miner. Metall. Process., Aug 1987, p 166-171 8. K.A. Natarajan, S.C. Riemer, and I. Iwasaki, Influence of Pyrrhotite on the Corrosive Wear of Grinding Balls in Magnetite Ore Grinding, Int. J. Miner. Process., Vol 13, 1984, p 73-81 9. R.L. Pozzo and I. Iwasaki, An Electrochemical Study of Pyrrhotite- Grinding Media Interaction Under Abrasive Conditions, Corrosion, Vol 43 (No. 3), 1987, p 159-169 10. A.V. Levy and Y. Man, "Erosion-Corrosion Mechanisms and Rates in Fe- Cr Steels," Paper III, Corrosion 86 (Houston), National Association of Corrosion Engineers, 1986 11. S. Agarwal and M.A.H. Howes, Erosion/ Corrosion of Materials in High- Temperature Environments, Proceedings of AIME Conference on High Temperature Corrosion in Energy Systems (Detroit), American Institute of Mining, Metallurgical, and Petroleum Engineers, Sept 1984 12. T. Foley and A.V. Levy, The Erosion of Heat Treated Steels, Wear, Vol 48 (No. 1), 1983, p 181 13. B.W. Madsen and R. Blickensderfer, A New Flow-Through Slurry Erosion Wear Test, Slurry Erosion: Uses, Applications, and Test Methods, STP 946, J.E. Miller and F.E. Schmidt, Jr., Ed., ASTM, 1987, p 160- 184 14. B.W. Madsen, A Study of Parameters Using a New Constant-Wear-Rate Slurry Test, Wear of Materials 1985, K.C. Ludema, Ed., American Society of Mechanical Engineers, 1985, p 345 15. R. Blickensderfer, J.H. Tylczak, and B.W. Madsen, "Laboratory Wear Testing Capabilities of the Bureau of Mines," IC 9001, Bureau of Mines, 1985 16. J.D.A. Bitter, A Study of Erosion Phenomena Part II, Wear, Vol 6, 1963, p 169-190 17. R.S. Lynn, K.K. Wong, and H. Mcl. Clark, On the Particle Size Effect in Slurry Erosion, in Wear of Materials 1991, K.C. Ludema, Ed., American Society of Mechanical Engineers, 1991, p 77-82 18. H.Mcl. Clark, Slurry Erosion, Proceedings of Conference on Corrosion-Erosion- Wear of Material at Elevated Temperatures (Berkeley), Electric Power Research Institute/ National Association of Corrosion Engineers, Jan 1990 19. J. Postlethwaite, The Control of Erosion-Corrosion in Slurry Pipelines, Mater. Perform., Dec 1987, p 41-45 20. L.D.A. Jackson, Slurry Abrasion, Trans. Can. Inst. Min. Metall., Vol 70, 1967, p 219-224 21. H. Hojo, K. Tsuda, and T. Yabu, Erosion Damage of Polymeric Material by Slurry, Wear, Vol 112, 1986, p 17-28 22. H. Hojo, K. Tsuda, and T. Yabu, Erosion Behavior of Plastics by Slurry Jet Method, Kagaku Ronbunshu, Vol 14, 1988, p 161-166 23. W. Blatt, T. Kohley, U. Lotz, and E. Heitz, The Influence of Hydrodynamics on Erosion-Corrosion in Two- Phase Liquid-Particle Flow, Corrosion, Vol 45 (No. 10), 1989, p 793-804 24. S. Nesic and J. Postlethwaite, Relationship Between the Structure of Disturbed Flow and Erosion- Corrosion, Corrosion, Vol 46 (No. 11), 1990, p 874-880 25. U. Lotz and J. Postlethwaite, Erosion-Corrosion in Disturbed Two Phase Liquid/ Particle Flow, Corros. Sci., Vol 30 (No. 1), 1990, p 95-106 26. J. Postlethwaite, M.H. Dobbin, and K. Bergevin, The Role of Oxygen Mass Transfer in the Erosion- Corrosion of Slurry Pipelines, Corrosion, Vol 42 (No. 9), 1986, p 514-521 27. D. Kotlyar and M.E . Wadsworth, "The Role of Localized Corrosion on Corrosive Abrasive Wear of High Carbon Low Alloy Steel Grinding Media," Paper 246, Corrosion 88 (Houston), National Association of Corrosion Engineers, 1988 28. J.D. Watson, P.J. Mutton, and I.R. Sare, Abrasive Wear of White Cast Irons, Met. Forum, Vol 3 (No. 1), 1980 29. R. Perez and J.J. Moore, The Influence of Grinding Ball Composition and Wet Grinding Conditions on Metal Wear, Wear of Materials 1983, American Society of Mechanical Engineers, 1983, p 67-78 30. K.A. Natarajan and I. Iwasaki, Electrochemical Aspects of Grinding Media- Mineral Interactions in Magnetite Ore Grinding, Int. J. Miner. Process., Vol 13, 1984, p 53-71 31. J.W. Jang, I. Iwasaki, and J.J. Moore, The Effect of Galvanic Interaction B etween Martensite and Ferrite in Grinding Media Wear, Corrosion, Vol 45 (No. 5), 1989, p 402-407 32. B.W. Madsen, Measurement of Wear and Corrosion Rates Using a Novel Slurry Wear Test Apparatus, Mater. Perform., Vol 26 (No. 1), 1987, p 21 33. B.W. Madsen, Measurement of Erosion-Corrosion Synergism with a Slurry Wear Test Apparatus, Wear, Vol 123, 1988, p 127-142 34. C.H. Pitt and Y.M. Chang, Electrochemical Determination of Erosive Wear of High Carbon Steel Grinding Balls, Miner. Metall. Process., Aug 1985, p 166 35. Y.M. Chang and C.H. Pitt, Corrosive- Erosive Wear of Grinding Ball Metals at High Jet Velocities, Corrosion, Vol 43 (No. 10), 1987, p 599-605 36. D. Kotlyar, C.H. Pitt, and M.E. Wadsworth, Simultaneous Corrosion and Abrasion Measurements Un der Grinding Conditions, Corrosion, Vol 44 (No. 5), 1988, p 221-228 37. C.H. Pitt, Y.M. Chang, M.E. Wadsworth, and D. Kotlyar, Laboratory Abrasion and Electrochemical Test Methods as a Means of Determining Mechanism and Rates of Corrosion and Wear in Ball Mills, Int. J. Miner. Process., Vol 22, 1988, p 361-380 38. A.K. Gangopadhyay and J.J. Moore, "An Assessment of Wear Mechanisms in Grinding Media," presented at SME-AIME fall meeting, Society of Mining Engineers/ AIME, 1984 39. R.L. Pozzo and I. Iwasaki, Pyrite- Pyrrhotite Grinding Media Interactions and Their Effects on Media Wear and Flotation, J. Electrochem. Soc., Vol 136 (No. 6), 1989, p 1734-1739 40. R.L. Pozzo, A.S. Malicsi, and I. Iwasaki, Pyrite-Pyrrhotite Grinding Media Contact and Its Effect o n Flotation, Miner. Metall. Process., Feb 1990, p 16-21 41. N.D. Tomashov, Theory of Corrosion and Protection of Metals, MacMillan, p 509-527 42. K. Adam, K.A. Natarajan, and I. Iwasaki, Grinding Media Wear and Its Effect on the Flotation of Sulfide Minerals, Int. J. Miner. Process., Vol 12, 1984, p 39-54 43. M.E. Learmont and I. Iwasaki, Effect of Grinding Media on Galena Flotation, Miner. Metall. Process., Aug 1984, p 136-144 44. I. Iwasaki, R.L. Pozzo, K.A. Natarajan, K. Adam, and J.N. Orlich, Nature of Corrosive and Abrasive Wear in Ball Mill Grinding, Int. J. Miner. Process., Vol 22, 1988, p 345-360 45. E. Wandke and M. Moser, The Influence of Corrosion and Hydrogen Cracking on Blast Wear in Wet Media, Wear, Vol 121, 1988, p 15-26 46. A.E. Isaacso n, P.J. McDonough, and J.H. Maysilles, "Corrosion Rate Determination in Industrial Ore Grinding Environments," Paper 232, Corrosion 88 (Houston), National Association of Corrosion Engineers, 1988 47. B.W. Madsen, A Portable Slurry Wear Test for the Field, Trans. ASME J. Fluids Eng., Vol 111, 1989, p 324-340 48. B.W. Madsen, A Comparison of the Wear of Polymers and Metal Alloys in Laboratory and Field Slurries, Wear, Vol 134, 1989, p 59-79 49. G.R. Hoey, W. Dingley, and C. Freeman, Corrosion Inhibitors R educe Ball Wear in Grinding Sulfide Ore, CIM Bull., Vol 68, 1975, p 120-123 50. J. Postlethwaite, B.J. Brady, M.W. Hawrylak, and E.B. Tinker, Effects of Corrosion on the Wear Patterns in Horizontal Slurry Pipelines, Corrosion, Vol 34 (No. 7), 1978, p 245 51. M.C. Roco, P. Nair, and G.R. Addie, Test Approach for Dense Slurry Erosion, Slurry Erosion: Uses, Applications, and Test Methods, STP 946, J.E. Miller and F.E. Schmidt, Jr., Ed., ASTM, 1987, p 185-210 52. M.C. Roco, "Wear Patterns in Centrifugal Slur ry Pumps," Paper 224, Corrosion 88 (Houston), National Association of Corrosion Engineers, 1988 53. M.C. Roco, Optimum Wearing High Efficiency Design of Phosphate Slurry Pumps, Proceedings of the 11th International Conference of Slurry Technology, March 1986, p 277-285 54. M.C. Roco and P. Nair, Erosion of Concentrated Slurries in Turbulent Flow, J. Pipelines, Vol 4, 1984, p 213-221 55. K. Ahmad, R.C. Baker, and A. Goulas, Computation and Experimental Results of Wear in a Slurry Pump Impeller, Proceedings of the Institution of Mechanical Engineers, Vol 200 (No. C6), 1986 56. J. Postlethwaite, "Influence of Flow System Geometry on Erosion- Corrosion," Corrosion 91 (Houston), National Association of Corrosion Engineers, 1991 57. S. Nesic and J. Postlethwaite, Hydrodynamics of Disturbed Flow and Erosion-Corrosion, Part I A Single- Phase Flow Study, Can. J. Chem. Eng., Vol 69, 1991, p 698-703 58. S. Nesic and J. Postlethwaite, Hydrodynamics of Disturbed Flow and Erosion-Corrosion, Part II A Two- Phase Flow Study, Can. J. Chem. Eng., Vol 69, 1991, p 704 59. S. Nesic and J. Postlethwaite, A Predictive Model for Localized Erosion-Corrosion, Corrosion, Vol 47 (No. 8), 1991, p 582-589 [...]... Parameter (a) 3 4.07 (13.4) 29 .55 (6.62) 13. 05 (2.92) 86.0 (187) 56 .5 (134) 34.8 ( 95) 2.49 53 .11 (127.6) 0.0468 99.12 (210.4) 2.769 4 4.07 (13.4) 39.40 (8.83) 19.38 (4.34) 112 .5 (2 35) 52 .2 (126) 42 .5 (109) 4 .55 78.88 (174.0) 0. 057 7 137.0 (278.6) 3.119 5 4.07 (13.4) 54 .17 (12.13) 25. 44 (5. 70) 122.0 ( 252 ) 87.0 (189) 52 .0 (126) 3.09 103 .54 (218.4) 0.0299 138.27 (280.9) 6.172 1 49 50 0 0 37 63 0 1 30 69 0 Proportions... Caplan and Cohen (Ref 14) Table 4 Using FINDAP on results of Table 1 Parameter Wear debris analysis DE DTH TH, m ( in.) N TF, °C (°F) A, m ( in.) APT APT QP, × 103 Experiment number 1 2 Only Fe2O3 Only Fe2O3 0. 054 16 0. 050 45 0. 054 16 0. 050 45 2.0 (80) 2 .5 (100) 398 54 6 440 (8 25) 4 15 (780) 0.98 (39) 1.33 (53 .2) 2.92E12 9.33E12 1 .5 × 106 1 .5 × 106 208 208 3 Fe2O3 and Fe3O4 0. 059 06 0. 059 06 3. 25 (130) 1 95 518... 54 6 PRINT "THEORETICAL DIVISION OF HEAT";DTH 54 8 PRINT "ACTIVATION ENERGY";QP 54 9 PRINT "ARRHENIUS CONSTANT";AP 55 0 PRINT "VIRTUAL PIN TEMP (TP)";TP 55 1 PRINT "VIRTUAL DISC TEMP (TD)";TD 55 2 PRINT "RADIUS OF CONTACT (a)";A 55 3 PRINT 55 4 PRINT 55 5 NEXT TH 55 6 NEXT K 56 0 STOP 57 0 END These pairs of values of N and TH are then inserted into the mild oxidational wear equation to eventually provide a value... (96) 27 .5 (82) 28.7 (84) 0 .50 0.90 9.18 17.87 0. 054 2 0. 050 5 39.78 (103.6) 51 . 85 (1 25. 3) 2.063 4.030 3 1.97 (6.27) 19.70 (4.41) 14.22 (3.18) 67.2 (1 35) 46 .5 (116) 34.2 (94) 1. 65 28.01 0. 059 1 76.10 (169) 3.003 4 1.97 (6.27) 29 .55 (6.62) 13.98 (3.13) 74.0 (1 65) 53 .0 (127) 33.0 (91) 1.78 27 .54 0.0648 83 .54 (182.4) 4.418 5 1.97 (6.27) 39.4 (8.83) 17.66 (3.96) 149.0 (300) 94.0 (201) 48.0 (118) 4 .58 34.79... 499 LET RH =5. 7E3 50 0 REM RH IS THE DENSITY OF FEO 50 1 LET APD=WR(K)*U(K)*P*TH*TH*F*F*RH*RH 50 2 LET APN=W(K)*(2.0)*A*EXP((-QP)/(R*(TF+273))) 50 4 LET AP=APD/APN 50 7 GOTO 54 0 50 8 REM LINES 51 0 TO 51 8 ARE A REPEAT 50 9 REM OF LINES 494 TO 50 2 FOR FE2O3 51 0 LET QP=208000.0 51 1 LET F=0.3006 51 2 LET RH =5. 24E3 51 3 LET APD=WR(K)*U(K)*P*TH*TH*F*F*RH*RH 51 4 LET APN=W(K)*(2.0)*A*EXP((-QP)/(R*(TF+273))) 51 6 LET AP=APD/APN... 51 6 LET AP=APD/APN 51 8 GOTO 54 0 51 9 REM LINES 53 0 TO 53 6 ARE A REPEAT 52 0 REM OF LINES 494 TO 50 2 FOR FE304 53 0 LET QP=96000.0 53 2 LET F=0.28 85 534 LET RH =5. 21E3 53 6 LET APD=WR(K)*U(K)*P*TH*TH*F*F*RH*RH 53 7 LET APN=W(K)*(2.0)*A*EXP((-QP)/(R*(TF+273))) 53 8 LET AP=APD/APN 54 0 PRINT "OXIDE THICKNESS(M)";TH 54 2 PRINT "NUMBER OF CONTACTS (N)";N 54 4 PRINT "OXIDATION TEMPERATURE(TO)";TF 54 6 PRINT "THEORETICAL... REM LINES 449 TO 457 ARE REPEAT OF LINES 432 448 REM TO 443 WITH N=N2 449 LET A2=SQR(W(K)/(PI*N2*P)) 450 LET L2=U(K)*A2/(2.0*XK) 451 LET TP2=HT(K)/(4.0*A2*KS*N2)+HT(K)*TH/(PI*A2*A2*(KO*N2) 452 IF L2< =5. 0 THEN 453 453 LET TD2=(-0.1021*L2+0.86 05) *HT(K)/4.0*A2*N2*KS) 454 IF L2 >5. 0 THEN 455 455 LET TD2=0. 85* HT(K)/(4.0*A2*N2*KS*SQR(L2)) 456 LET DTH2=TD2/(TD2+TP2) 457 LET FN2=(DTH2-DE(K)) 458 REM LINES 460... CONTACT RADIUS (A) 53 6 REM WHERE RD IS THE Rth VALUE (FOR STARTING VALUE - SEE 53 7 REM 400) AND IS GIVEN BY EQUATION (A2) OF PAPER BY 53 8 REM QUINN, ROWSON AND SULLIVAN (REF 2) 54 0 IF ABS(R1-RD) . TEMP (TP)";TP 55 1 PRINT "VIRTUAL DISC TEMP (TD)";TD 55 2 PRINT "RADIUS OF CONTACT (a)";A 55 3 PRINT 55 4 PRINT 55 5 NEXT TH 55 6 NEXT K 56 0 STOP 57 0 END These pairs. 51 6 LET AP=APD/APN 51 8 GOTO 54 0 51 9 REM LINES 53 0 TO 53 6 ARE A REPEAT 52 0 REM OF LINES 494 TO 50 2 FOR FE304 53 0 LET QP=96000.0 53 2 LET F=0.28 85 53 4 LET RH =5. 21E3 53 6 LET APD=WR(K)*U(K)*P*TH*TH*F*F*RH*RH. L2=U(K)*A2/(2.0*XK) 451 LET TP2=HT(K)/(4.0*A2*KS*N2)+HT(K)*TH/(PI*A2*A2*(KO*N2) 452 IF L2< =5. 0 THEN 453 453 LET TD2=(-0.1021*L2+0.86 05) *HT(K)/4.0*A2*N2*KS) 454 IF L2> ;5. 0 THEN 455 455 LET TD2=0. 85* HT(K)/(4.0*A2*N2*KS*SQR(L2))