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By Carlton D. Hulin

ARE we a "have" or a "have-not" nation in our domestic supply of metals and minerals? Impinging on the ears of a people weary of war and faced with the problems of reconversion to peace, the import of this question has built up with the momentum of an avalanche. Stemming from statistics cited by Pehrson and accelerated by the pessimistic and possibly biased statements of Ickes, hardly an issue of any technical periodical dealing with mining has appeared during the past year but within whose pages was to be found some comment, some discussion of the question: Are we a "have" or a "have-not" nation? We have proved commercial reserves of high-grade iron ore in the Lake Superior region sufficient for 22 years, based upon the prewar rate of consumption; commercial copper reserves to last only 34 years, one fourth being inferred ore; zinc sufficient for nineteen years, over half being inferred; lead for twelve years, two thirds of which is inferred; mercury for only three years. So read the figures. Although admitting that these figures may be extended by new discoveries, by the use of marginal and submarginal ore, and by further technological advances, are we nevertheless to assume complete pessimism in the acceptance of the meaning of these statistics; are we to expect a rapidly increasing dependence upon foreign sources of mineral supply; are we to believe that we are at or near the peak of the "metal age" as has been held by Theodore J. Hoover? Or is some more optimistic viewpoint possible?

Jan 1, 1947

By Sheldon P. Wimpfen

EVER since the Phoenicians roamed the known world in quest of metals to harden their helmets and precious metals and gems to adorn their ladies, many other nations have sought metals in the far corners of the earth. During the past 75 years American engineers have been among the foremost in seeking for and developing ore deposits. They have been responsible for many of the principal discoveries and for the application of modern methods to mines in the far comers of the earth. American mining engineers with a zest for travel and an insatiable appetite for metals have also introduced throughout the world many articles made of metal which we consider essential to everyday life. Production and consumption of these metals in many foreign countries has brought to these countries an increased standard of living and added tax revenue to their governments. But unfortunately engineers from this country are no longer given ready access to all the countries of the world. This is not only unfortunate for American engineers but also for the countries that have adopted the exclusion policies and especially so since time has proved that in countries where the American engineers have gradually been excluded there has been a significant decline in the progress of the mining industry. Removal of these barriers to professional men would aid immensely in the world-wide development of the mineral industries. New areas brought into production under the guidance of American engineers with their exceptional know-how would provide more jobs for nationals of the countries involved and a better standard of living for the nations as a whole.

Jan 1, 1946

By George C. Stone

AS most of you probably know, Australia has had many strikes. The two places that had the worst reputation were the Broken Hill mines and Port Pirie, where the smelter was located. About four or five years ago, the smelter changed hands. The new proprietors - immediately endeavored to get on better terms with the men and made such progress in that respect that, the men came to them and said, "There is no use giving us any more money. You have just raised our pay one- bob a day, but the store-keeper and the boarding-house keeper have raised their prices. The boarding housekeeper raised our board three bob a week and the storekeeper took the rest of it and we are no better off." Learning that what the men needed most was firewood wood the company bought the wood on a tract of ground across Spencer's Gulf and- sent men over to cut it. The wood was brought to the works and given to a man to cut up. He thought he had an easy job and loafed, but when the committee representing the men discovered that his loafing added to the cost of 'the wood, it promptly disciplined him and made -him work as he had never worked before. The next question that arose was the distribution of the wood. The committee got bids from several cartmen whose prices varied from about two to five shillings. It gave the work to the next to the lowest bidder. The fact that if delivering the wood cost more, the user would have to pay for it, was an entirely new idea to the men.

Jan 1, 1920

By Colin G. ink

OUTSTANDI'NG in progress among the less familiar 'metals during 1936 is the electrolytic production of 99.9 per cent manganese meta1 readily and many quantity. Strictly speaking, manganese should not' be classed as 'a rare metal, since there is 800 times as much manganese in the earth's crust as there is zinc or lead. But manganese' as a metal is rare, just as titanium metal is rare though titanium is five times as abundant in the earth's crust as manganese.

Jan 1, 1937

By W. T. lsbell

HERCULANEUM, MO., about thirty miles south of St. Louis on the Mississippi River, is the site of the lead smelter of the St. Joseph Lead Co. The lead concentrates come by rail from the Flat River district and from three properties in Madison County, Missouri; from the Balmat division in St. Lawrence County, New York; a small tonnage from the Tri-State district; and occasionally from the Aguilar property in Argentina. The average grade is 67 to 68 percent lead and 1 to 3 oz of silver. The bulk of the material is flotation concentrate, with some gravity or table concentrate. No secondary metals, outside drosses, or any materials high in bismuth, arsenic, or antimony are treated. All concentrate is received in box-caps and is unloaded directly from a trestle into the concentrate bins at the roaster plant. Flux, coke, and any miscellaneous lead-bearing material is either unloaded directly into bins or unloaded with a locomotive crane. All moving and switching of cars is done with two thirty-ton steam locomotives. Shipping facilities for the plant are provided by the Missouri-Illinois Railroad. Water for the plant and for fire protection is obtained from a dam built in a near-by creek. Drinking water both for plant and town is obtained from

Jan 1, 1947

By Franklin H. Sharp, Kenneth L. Clifford

During the last few years the New Lead Belt of Southeastern Missouri has become the main source of lead in the United States. It also produces significant amounts of zinc, copper and silver. The mines are located on a narrow, mineralized trend extending southward from Viburnum, MG. for about 35 miles. The ore occurs as replacement deposits in the Bonneterre lime- stone formation which is about 1000 ft below the surface. In most of the mines, enough water is produced to supply the milling requirements without any reuse. There are dry areas, however, and at least two mills must rely on water reclaimed from their tailings lake to augment the mine water. These mills have no choice but to utilize recycled mill water.

Jan 7, 1973

By Cyril S., Smith

WE cannot but marvel at the fact that fire is necessary for almost every operation. It takes the sands of the earth and melts them-now into glass, now into silver, minium or other lead or some substance useful to the artist or physician. By fire, minerals are disintegrated and copper produced; in fire is iron born and by fire is it subdued; by fire, gold is purified; by fire, stones are burned for the binding together of the walls of houses.... Fire is the immeasurable, uncontrollable element, concerning which it is hard to say whether it consumes more or produces more." Thus wrote Pliny in his "Natural History," that great collection of fact and error which constitutes the principal contemporary description of Roman technology and which until the seventeenth century was the authority most often quoted and requoted by those who wrote on metals, as on many other subjects.

Jan 1, 1940

By W. R. Hibbard

Evidence is presented which confirms previous findings that models of solution strengthening depending solely on lattice parameter changes are incomplete. Direct evidence for the Suzuki interaction of chemical effects was obtained from alloys with constant lattice parameter, for which solution strengthening was nearly independent of temperature from 300° to 20°K. SOLUTION hardening principles have been reviewed recently by Parker and Hazlett.' In addition, the interactions of dislocations and solute atoms were reviewed by Cottrell.' It is not proposed to review this information again here, but to refer only to those details which are pertinent to the developments which have occurred since these papers were published. Literature Survey To set the stage, it has been known for some time that solid solutions are stronger than the pure parent metals, and that the strength increases with increasing amounts of alloying element. One of the most extensive earlier investigations was that of Lacy and Gensamer,3 who reported that the solid solution strengthening of iron is a power function of the composition, that the relative strengthening is a function of the solubility limit, and that the rate of strain hardening is a function of the amount of solution hardening. Thus, the slope of the plastic portion of the stress-strain curve increases with increasing yield strength. An approach to the development of principles resulted from work by Gulyeav,4 who reported that the strengthening effect of solid solution alloys is periodic in character when plotted as a function of the atomic number of the solute. A more quantitative approach was to relate the amount of solution hardening to the lattice parameter of a solid solution. This was done quite early by Norbury,5 and later by Brick, Martin, and Angier,6 Frye and Hume-Rothery,7 and French and Hibbard. This relationship is only approximate in character. The next development was to relate the amount of solution hardening to lattice stress, as proposed by Frye and co-workers.9 They calculated what they called ionic overlap from the lattice parameter changes represented as a change in volume, using Bridgman's hydrostatic compression equation. The resulting correlation was somewhat improved as compared to the simple correlation with lattice parameter. The next step was carried out by Dorn, Pietrow-sky, and Tietz,10 who suggested an effect of valence differences in addition to the effect of lattice parameter. Their proposed parameter shows flow stress as a single valued function of the concentration and a relationship involving these two atomic variables. In addition, they contributed new data on the temperature dependence of solution hardening which indicated that it was a strong function of temperature for aluminum alloys. Allen, Schofield, and Tate11 carried the valency effect even further by reporting data indicating that with a given solvent element the stress-strain curve is a function of electron density, entirely neglecting the size factor, which varied by a factor of two. This evidence led to a quantum mechanical calculation by Cottrell, Hunter, and Nabarro,12 based on the formation of a dipole around an edge disloca-

Jan 1, 1959

By A. S. Odeh

During the last few years, several authors1-6 have advanced mathematical solutions, both exact and approximate, to the pressure behavior of a well producing from a region bounded by a circular discontinuity. The most recent solutions were reported by Carter, and by Bixel and van Poollen. Varter solved the appropriate equations analytically. Bixel and van Poollen reported a numerical solution and provided correlation charts for the interpretation of flow tests; their method uses an overlay technique and is valid for wide ranges of mobility and storage capacity ratios. The mobility ratio is defined as the ratio between k/p of the two zones, and the capacity ratio is defined as the ratio between $c of the two zones. We have developed independently an analytic solution, programmed for the CDC 1604, that agreed with Carter's.'1 The same differential equation, assumptions, boundary, and initial conditions were used. We ran 30 drawdowns and 30 buildups of a unit capacity ratio and of varied mobility ratios. We also varied the radial distances to the discontinuity, and the radii of drainage. We examined the results and obtained correlations that were simple to apply and useful in flow test interpretation. The results and the correlations reported here in graphical form allow the determination of the mobility of the two zones and the radial distance to the discontinuity. They eliminate the need to fit field data to a curve as is required in the overlay technique.6 Thus, the advantages of the present correlations are that they require less effort and they provide a unique answer. They also provide guidelines for selecting the parts of the flow test plot to be used in the interpretations. Moreover, the results obtained from the programmed analytic solution define the time intervals where the approximate solution reported by Hurstl and discussed by Larkin2 are applicable. Normal buildup and drawdown plots resulted in an early straight line, a long transition zone and a late (second) straight line. However, the second straight line will not occur if the drainage radius is not at least 10 times the radial distance to the discontinuity. The mobility of the imer and outer zones can be calculated from the slopes of the first and second straight lines, respectively, by the normal methods of the flow test analysis. We found that the radial distance to the discontinuity may be determined from the abscissa value for the point of intersection of the extensions of the two straight lines. Results similar to some of these were reported by Adams et aL7 The approximate solution compared favorably with the exact one during the first and second straight-line periods. However, during the transition zone, deviation occurred between the two solutions. Discussion of Figures Figs. 1 through 7 summarize the results of this investigation. On Figs. 3 through 7 there appears the equation of each respective plot. In this paper the mobility ratio M is defined as klp2/kzpl, where subscripts 1 and 2 refer respectively to the inner and outer zones, and dimensionless time = t, = 6.32 X In the investigation, M ranged from 0.25 to 50;

Jan 1, 1970

By W. J. Verner, J. R. Lucas

Billions of tons of bulk solid materials are processed through our industrial plants each year, and the tonnage is steadily rising. It has been estimated that for every dollar spent in industry as a whole, approxinlately 30 cents is consumed in the handling of bulk solid materials. It is in this category that tremendous savings are possible through the use of improved techniques. Since World War 11, the mining, steel, chemical, and grain industries, among others have made use of increased materials-handling automation in order to gain some of these savings. However, in a chain of automatic operations, every link must be completely reliable, for if there is a single failure, the whole system immediately becomes ineffective. Only too often the weak link in the chain proves to be the flow of bulk solids from bins. To contribute to the fundamental knowledge in this area, a research study has been continuing at The Ohio State University. The principal objective of the investigation has been to study the internal stresses existing in confined bulk solids. LITERATURE SURVEY Literature studies have revealed that many outstanding people have made contributions to the field of bulk solids. For instance, C. F. Jenkin,1 suggested: a. The pressure on a bin wall is not directly proportional to the depth beneath the surface of material in the bin. b. The vertical pressure distribution on the base of a bin is not actually uniform but is a variable. c. Arches form in bins as a result of very small deformations of the material in the bin. This deformation causes a series of gaps to appear, and consequently causes the disappearance of diagonal forces that cannot act across these gaps. The removal of the diagonal forces causes horizontal forces to be introduced which seem to be the chief cause of arching. 11. A. Janssen,4 in 1895 in connection with his experiments on bins, derived a differential equation to describe both the lateral and vertical pressure in bins with vertical sides. This represents a significant contribution to bulk solid theory, and thus its solution is reproduced as follows: L = KV = RW/U' [l-e -U'KY/R] I where L is the lateral pressure of the bulk solid in psf., V is the vertical pressure of the bulk solid in psf., K is the ratio of the lateral pressure to the vertical pressure at any point, R is the hydraulic radius of the bin in feet, W is the density of the bulk solid in lbs. per cubic feet, U' is the coefficient of friction between the stored material and the bin wall, and Y is the head of material above the point in question in feet. In 1903, Jamieson,6 in an elaborate series of experiments on full sized bins measured vertical and lateral pressures by hydraulic diaphragms. He determined many values for mediums of corn, wheat, peas, and flaxseed. In 1951, Caughey, Tooles, and scheer3 measured both lateral and vertical pressure in deep bins by using strain guages on steel bands. Values were calculated for cement, gravel, sand, soy beans, shelled corn

Jan 1, 1961

By D. Weinstein

Yieldilzg in annealed arc-cast molybdenunz in torsion was studied as a function of strain rate and tem-perature. The temperature dependence of the yield point for different strain rates was used to calculate a heat of activation for the yield point. The heat of activation was not a constant but was approximately a linear function of the stress at yield. It is proposed that the rate effect in yielding is determined in part by a site place-change of C (or other interstitial atom) in the metal. THE low-temperature embrittlement phenomenon in metals and alloys has been known to exist for a long time. Although a transition from ductile to brit-tle behavior has been observed in some close-packed hexagonal metals1 and a face-centered cubic alloy, the most common class of metals exhibiting this be-havior has a body-centered cubic lattice structure. For example, a ductile to brittle transition has been observed in iron,3 molybdenum,4 chromium,5 colum-bium,6 and tungsten.7 In the present investigation, molybdenum specimens were tested in torsion over a wide range of temperatures at three widely separated strain rates, and the temperature for onset of embrittlement was determined. Characteristic of body-centered cubic metals and closely associated with plastic deformation is the appearance of a sharp upper and lower yield point. Theoretical explanation of the yield point in iron and low-carbon steel is due to Cottrell,8 who attributes the phenomenon to the pinning of dislocations by the formation of impurity atmospheres around the dislocations. These atmospheres tend to lock the dislocations and make them more difficult to move. The stress required to move a dislocation and thereby cause plastic deformation is greatly increased, and thus an upper yield point is observed. In addition, theoretical analysis,9 as well as experimental observations, show that the upper yield point is strongly a function of temperature. The upper yield point is observed to decrease with an increase in testing temperature, and above a certain temperature (-700° K in iron) it no longer appears. This is consistent with the concept that thermal vibrations will help free a dislocation from its atmosphere, so that the external stress required to free a dislocation from its pinning atmosphere decreases with increasing temperature. Furthermore, with increasing temperature the equilibrium concentration of impurity atoms around a dislocation decreases exponentially with the result that the dislocation is less firmly pinned. In Cottrell's theory of yield point a static model is visualized in which the pinning atmospheres, particularly interstitial atoms, are immobile, and the dislocations are anchored to stationary positions in the lattice. However, experiments1,7 show the upper yield point to be time-dependent or a function of the rate of straining, a high strain rate causing a high upper yield point. Present theory does not adequately account for the rate dependence of the yield point, and the exact function of the interstitial atom is not under stood. Nevertheless, observation of the temperature dependence and rate dependence strongly suggests that the upper yield point phenomenon is due to the interaction between dislocations and pinning atoms, and that a diffusion mechanism is important in determining the yield point.

Jan 1, 1962

By G. Meyrick, H. W. Paxton

Observations on the tensile deformation of single crystals of austenitic stainless steels as a function of composition, orientation, and temperature are described and compared with relevant data for other alloys. Crystals free from second phases show sharp yield points followed by Lüders' extensions of up to 20 pct. The critical resolved shear stress is strongly temperature-dependent increasing fourfold between 423O and 77°K. A temperature-independent low rate of hardening follows the Luders' extension and is frequently associated with overshoot. Unlike pure metals where Proficse cross slip usually occurs at the end of linear hardening, this stage is terminated either by cross slip or by slip on the conjugate system, the particular node apparently depending upon initial orientation and composition. The presence of 6 ferrite or martensite has a profound effect upon the magnitude of the yield stress and the shape of the stress-strain curve. In experiments on transgranular-stress corrosion cracking of austenitic stainless steels in boiling magnesium chloride, several workers1'2 have shown that increasing amounts of nickel tend to reduce the probability of specimen failure. Reed and paxton3 have observed in single crystals that the macroscopic crack plane changes as the nickel content is increased. For 18 pct Cr-8 pct Ni, the crack plane is sensibly perpendicular to the applied tensile stress; for 20 pct Cr-20 pct Ni the crack plane is (100). The average crack-propagation speed is much lower in the 20 pct Ni crystals. Furthermore, the possible importance of stack-ing-fault energy in nucleation and growth of cracks has been suggested by a number of workers. While the general effect of variations in stack ing-fault energy appears to be quite well-correlated with observed variations in cracking characteristics, each of the theories assigns rather different detailed reasons for these effects. Because the stack ing-fault energy of austenitic steels is believed to be increased with increasing nickel content, and because some theories of trans-granular cracking postulate a mechanical stage in the crack propagation (rather than relying entirely on electrochemical solution), it appeared to be of some interest to investigate the mechanical properties of a series of different compositions of aus-tenite single crystals. These experiments were considered of intrinsic value in view of the relative scarcity of such data on fcc alloys. This paper gives the results of some experiments carried out with this aim in mind, and suggests various 'other experiments which may be helpful in future studies. The experiments are also of interest in view of the recent interest in obtaining higher-strength steels by control of the imperfection content of austenite; although the compositions are not directly comparable, the results are still probably valuable. Comparisons with experiments on other fcc alloy crystals are made where possible to see how well the general pattern of behavior is understood. EXPERIMENTAL PROCEDURE The alloys used in this investigation were vacuum-melted heats of nominally 20 pet Cr-20 pet Ni-60 pet Fe, 20pctCr-16 pct Ni-64 pct Fe, and 20 pctCr-12 pet Ni-68 pet Fe.* Analyses of the material are

Jan 1, 1964

By D. W. Fuerstenau, D. A. Sullevan

In the experimental work for this comparison of wet and dry grinding, it was found that the size distributions for wet grinding operations are characterized by a constant value of the distribution modulus, (a; whereas, for dry grinding the distribution modulus decreases slightly with increased size reduction. Wet grindability seems independent of the weight of material in the mill while dry grindability depends on weight. Experimental data and eqmtions for each process are presented for noncleavable isotropic material, for cleavable or material in which there is some form of internal stress, and for energy consumption. The equation ED/EW=BK-? (where ED and Ew are, respectively, energy for wet and dry grinding, k is the size modulus, and B and ? are constants) was found to give the ratio of energy. Dry grinding is generally considered to require more energy than wet grinding. Taggart' states that power consumption in wet grinding ranges from 60 to 90 pct of that required for dry grinding. Bond' states that dry grinding requires one third more power than wet grinding. It is interesting that Rose and Sullivan' note this same difference in power required for wet and dry grinding, but they state that on purely mechanical grounds it is difficult to see any great difference in fundamental principles between wet and dry milling, since dry milling may be regarded as wet milling with a fluid of low viscosity (i.e., air vs water). Yet it must be the viscosity of the fluid that is responsible for the differences between wet and dry grinding. Probably the most comprehensive investigation comparing wet and dry grinding has been that made by Coghill and Devaney. They compared the grind-ability of both chert and dolomite in a 19x36-in. batch ball mill under both wet and dry conditions. In these tests the weight of ore in the mill was varied from 35 to 200 lb and each grinding test was carried out for such period of time that the amount of -200-mesh material in the product was essentially the same for each test with a given material. Cog-hill and DeVaney's data for the grindability of dolomite, which are reproduced in Fig. 1, show that a comparison between wet and dry grinding is com- plex. The comminution of a material by a wet grinding method is characterized by a constant grindability over a wide range of mill loadings. Since grindability is the weight of material comminuted per unit of energy expended, it is equivalent to (w)/ (p)(t), where W is the weight of material being ground, P is the power, and t is the grinding time. In the case of wet grinding, the grindability of dolomite in the Coghill and DeVaney experiments was 0.546 * 0.009 ton per hp-hr when the load of dolomite in the mill ranged from 50 to 200 lb. On the other hand, for dry grinding no such simple relationship was found to exist between power, time, and weight of material in the mill. Dry grinding efficiency appears to depend greatly upon the amount of material within the mill. When the mill contained

Jan 1, 1961

By T. B. Massalski, Horace Pops

The occurrence and the temperature dependence of the athermal martensitic transformation in bcc Cu-Zn ß-phase alloys have been studied by cold-state microscopy, differential thermal analysis, and electrical -resistivity measurements. CinenzatograpJzic studies have shown that in the majority of alloys the transformation occurs in two stages. A thermoelastic phase, having a 'needlelike" uppearance, forms initially at the MS temperature and grows relatively slowly, chiefly in the edge (lengthwise) direction. With further cooling, an additional martensitic phase forms in rapid bursts at a lower temperature, MB, Both martensitic phases appear to have the same cry stallographic habit hut their modes of formation and growth are different. For the composition range studied, between 38.55 and 41.49 wt pct Zn, the habit planes determined are close to the (2.17 ,12) plane of the matrix, in general agreement with the prediction of the phenomenological theory of martensite formation. Thermal cycling through the MB temperature produces permanent plastic deformation in the 0 matrix and gradutally eliminates the typical burst phenomena. Thermal cycling only through the MS temperature produces no noticeable deformation effects. THE bcc Cu-Zn ß-phase alloys tend to become unstable at low temperatures and undergo a martensitic transformation,'-3 but the details of the transformation have not been completely determined. Following earlier metal log raphic and X-ray work1 Tichener and ever' determined by means of electrical-resistivity measurements the temperature range of the martensite transformation during continuous cooling. The transformation temperature was found to decrease with increasing zinc content, becoming about -130°C for an alloy containing 40.04 wt pct Zn. Hull and Garwood3 made a metallo-graphic study of Cu-Zn 0 brasses at temperatures down to -160°C, and determined the habit plane of the transformation product for an alloy containing 38.88 wt pct Zn. No metallographic studies below liquid-nitrogen temperatures have been reported. The martensitic transformation in ß-brass alloys can also be induced by deformation.1,4 The MD temperature corresponding to the appearance of the strain-induced martensite is some 300°C higher than that observed without strain, and hence the transformation can be strain-induced in alloys containing a higher percentage of zinc. Massalski and Barrett4 found that alloys containing as much as 51.89 wt pct Zn could be transformed into a faulted hcp structure by cold working at liquid-helium temperature. Because the details of the MS temperature and morphology reported in the literature are somewhat incomplete, a further study has been made of some of the above features. A design of a cryogenic stage for optical metallography5 has permitted the examination of specimens at temperatures down to 4.2oK, and the use of a cinecamera has permitted a study of the morphology on cooling and heating. The particular aim of the present work was to examine the transformation temperatures, the habit planes, and the growth morphology over a wide composition range in the ß-phase field. The effect of thermal cycling upon the subsequent transformation was also examined.

Jan 1, 1964

By A. Vignes, J. P. Sabatier

A study of ternary diffusion at 1315°C in the Fe-Co-Ni system covering the entire ternary phase diagram has been carried out. Experimental values of the four interdiffusion coefficients DFeNico,DFeCoCo, DFeCoNi were computed at 33 points of the ternary diagram. The limiting values of these coefficients on the three sides of the ternary diizgram have also been determined. This set of data is presented as a contour map on the ternary isotherm for each of the four coefficients. Assuming that the Fe-Co-Ni system can be considered as ideal, the phenomenological coefficients Lk ij (with k = Fe) have been computed and the Onsager reciprocal relutions verified. It is shown that, in the Fe-Co-Ni system, the diffusion process can also be described by only two interdiffusion coefficients DFeFe and DCoNiFe, the second cross diffusion coefficient DCoFeNi being equal to zero and the second direct difusion coefficient being given by The values of the cross interdiffusion coefficients and of the Onsager coefficients are explained in terms of atomic mobilities of the three components with the help of a kinetic calculation due to Lane and Kirkaldy.8 It is shown that all the results are explained if the tracer difusion coefficients of the three components are such that DFe > DNi ~ DCo in complete agreement with experimental data on the three binary systems. DIFFUSION in a ternary alloy system is basically and significantly different from diffusion in binary alloys. In binary diffusion, there is just one independent composition variable and only one coefficient Dij is needed to describe the interdiffusion process. In ternary diffusion, we must allow for cross interaction between two independent species, and four interdiffusion coefficients are necessary to describe the transport process. Kirkaldy et al.1,2 have determined interdiffusion coefficients in the systems, Cu-Mn-A1 and Zn-Cu-Al. Guy and Leroy4 have reported intrinsic diffusion coefficients for the r phase at 1300°C of Co-Ni-Cr alloys as a function of composition using the analysis of Philibert and Guy.3 Interdiffusion coefficients were measured at three different compositions at 850°C in the single phase copper-rich corner of the Cu-Zn-Mn system by Dayananda and race.5 Zie-bold and Ogilvie6 have determined interdiffusion coefficients in the single-phase field of the Cu-Ag-Au system at 725°C. Preliminary results on diffusion at 1136°C in the system Fe-Co-Ni have been reported.7 In this study, data are reported on a more complete determination of interdiffusion coefficients in the Fe-Co-Ni system in the entire ternary at 1315°C. There is complete solid solubility with the fcc structures at 1315ºC, and this system can be considered as nearly ideal. This approximation allows the computation of the Onsager phenomenological coefficients. These values can then be used to test a kinetic calculation of the Onsager L coefficients in ternary substitutional solid solutions due to Lane and Kirkaldy.8

Jan 1, 1970

By N. J. Themelis, A. W. Last

Autogenous grinding, broadly defined as the self-comminution of ore in a cylindrical tumbling mill without using auxiliary grinding media, was first applied in the early 1930s in the so-called Hadsel mills (Hardinge, 1955) as seen in [Fig. 1], mainly used in gold mines on the US West Coast. A number of these mills were built in sizes of 6.1 to 7.3 m (20 to 24 ft) in diameter and 1.5 to 1.8 m (5 to 6 ft) in length but were later abandoned because of operating and maintenance problems. Further development of autogenous grinding led to the Hardinge wet grinding mill, the Aerofall dry grinding mill, and others. A tabulation of all known autogenous and semiautogenous installations around the world was sent recently to the authors by G. J. Kushner of Koppers Metal Processing Division. In the size range of 7.3 to 9.8 m (24 to 32 ft) in diameter, it lists a total of thirty mills used in the processing of copper ores. An additional 120 units are listed in the same size range for the processing of iron ores and other minerals. The largest mills in existence are 11.1-m-diam X 4.6-m-long (36-ft-diam X 15-ft-long) mills in Minnesota, powered by 8950 kW (12,000 hp) motors and used in taconite beneficiation. It is indicative of the phenomenal growth of this technology that none of the above 150 mills was in operation prior to 1964. Grinding rates in autogenous mills are very sensitive to variations in ore breakage characteristics and even with fairly constant ore type can exhibit large fluctuations in throughput. Autogenous grinding also depends on availability of a fraction of the ore to be ground in the form of "competent" rock which can withstand to a certain degree the tumbling and impact action within the mill. If the ore is uniformly "soft", it will soon be crushed into small size particles and further grinding, which depends on the existence of large size rocks in the mill, becomes impossible. This apparently is the case with most copper ores, as discussed by Bachman, Last and Nabbs (1970) who described pilot testing of a large number of ore samples from Kennecott 's Chino Mines. Under circumstances where autogenous grinding is impractical, it is possible to realize most of its benefits by utilizing large grinding balls in place of the "competent" rock and practice what has become known in the mining industry as semiautogenous grinding. In semiautogenous grinding (SAG), a grate discharge tumbling mill is fed the product of a primary crusher or run-of-mine rock. Forged steel or alloy grinding balls, in proportions ranging from 3.10% of the total volume in the mill, are used to complement the self-grinding ability of the ore. Semiautogenous circuits can be either designed for single-stage closed-circuit grinding which produces a final product for recovery processes, or for two-stage operation in which the SAG mill product is fed to a second stage of grinding in a conventional ball mill.

Jan 1, 1981

By C. D. Rubert, W. J. Parton

A flotation plant was placed in operation in 1945 at the Tamaqua colliery of the Lehigh Navigation Coal Co. Inc., to recover fine anthracite which is discharged from the cleaning plant with the washery water.‡ In addition to the recovery plant, facilities were provided for the complete clarification of the washery water, recirculation of this clarified water for breaker use, and disposal of the refuse solids in a waste impounding dam. Fig I shows a view of the settling tanks and the flotation plant. Flow Diagram The flow diagram for the complete recovery project is given in Fig 2. All the waste water from the cleaning plant is collected in a sump A outside the breaker structure, from which a 7000 gpm pump B delivers it through a 16 in. line to a 75-ft diam hydroseparator C. The hydroseparator, acting as a classifier, is set to give a separation at approximately 200-mesh. The minus 200-mesh overflow is discharged by gravity into the 180-ft thickener D for final settling of the solids. The classifier underflow sludge is pumped to the flotation plant by a 6-in. quadruplex diaphragm pump E. A 10-ft diam hydroseparator F is used to make a second classification on the sludge solids at approximately 28-mesh. The underflow from this classifier containing the coarse particles is elevated by a 6-in. diaphragm pump G and is discharged on two 4- by 6-ft vibrating screens II. These screens are dressed with approximately 28-mesh cloth. Screen oversize product flows directly to the centrifuge that dewaters the cleaned froth. The screen underflow and the overflow from the 10-ft classifier flow to the flotation circuit. The flotation feed is conditioned in a 12 by 12-ft agitator tank J at a pulp density of I:2.33 (30 pct solids). Reagents are fed to the circuit by reagent feeder I. The conditioned pulp is then diluted to a pulp density of approximately I: 5 (18 pct solids) and is fed to two banks of four 100-cu ft froth flotation cells K for primary coal recovery and cleaning. Additional reagents from reagent distributor L are added to the feed weir-box of Nos. 2, 3, and 4 cells of each bank of flotation machines. Clean coal removed from these cells flows directly to the clean-coal sump. Dilute tailings from the primary cells is re-thickened in a 30-ft diam thickener M and is then retreated in a scavenger circuit for final recovery of coal. The scavenger circuit consists of a 6- by 6-ft agitator tank N and two banks of four 24-cu ft froth flotation cells 0. Coal froth from these cells flows to the dewatering equipment and the tailings flows to a vertical refuse pump which delivers it to the 180-ft diam waste thickener D. Dewatering of the coal products is accomplished in a 54- by 72-in Bird centrifugal filter P. The cake product is dropped onto a conveyor and is loaded

Jan 1, 1949

By K. Mukherjee

A quasi-thermodynamic calculation is presented for estirt~ating vacancy Parameters in fcc metals. For fcc metals the equilibrium vacancy concentration at the melting point can be approximated by a constant mole fraction. The Gibbs free energy of formation of the vacancies can then be expressed as a simple function of the absolute melting temperature alone. From this description the formation energy Ef, the vibrational entropy Sv, and the formation volume of a vacancy can be calculated in terms of known thermodynamic parameters. The calculated values are in good agree-rnent with available experimental results. IT appears that many of the physical parameters associated with lattice defects can be empirically related to thermodynamic melting parameters.'-7 Besides the implicit correlations between melting phenomenon and point defects, the melting temperature itself at atmospheric pressure can be used as a reference temperature for the purpose of comparison of vacancy concentrations in a given crystal class. The concentration at the melting point then corresponds to the maximum vacancy super saturation in the crystalline state at a fixed pressure. For pure fcc metals such as Au, Cu, Al, Ag, Pb, and Pt the mole fraction of vacancies at the melting point8 is of the order of 10*"3 to 10"1. The equation expressing the equilibrium vacancy concentration at the melting point contains three parameters: the formation energy, the vibrational entropy, and the absolute melting temperature, see Eq. [2]. Since these three parameters cover a wide range of values8 in the above-mentioned metals, it can be assumed that in other fcc metals, for which no reliable data are available, the mole fraction of vacancies at the melting point most likely lies in the range 10~3 to lO"4. This is then the basis of the calculations described in this article. Within this range the mixing entropy can be linearized and the vibrational entropy, the formation energy, and the vacancy formation volume can be calculated in terms of known thermodynamic constants. 1} LINEARIZATION OF MIXING ENTROPY The Helmholtz free energy of a crystal containing n vacancies at the melting point can be writteng as where Ef, T,, k, and S, are the formation energy of a vacancy, the absolute melting temperature, Boltz-mann's constant, and the vacancy formation entropy, respectively. After applying Stirling's approximation in Eq. [l] and minimizing with respect to n where (n/N), is the equilibrium vacancy concentration in mole fraction at the melting temperature at 1 atm pressure. If instead of setting (6~/6n) = 0 immediately after Stirling's approximation in Eq. [I], the configurational partition function is simplified, then some useful and significant results can be obtained by utilizing the experimental inforrnation After applying Stirling's approximation and rearranging is plotted as a function of (n/N) in Fig. 1. The points, not all of which are shown, are fitted by a straight line substituting from Eq. [5] into Eq. [I] and minimizing with respect to n combining Eqs. [6] and [2] It must be emphasized that Eq. [7] is only a crude approximation based on an oversimplified model of

Jan 1, 1967

By R. E. Carter, F. D. Richardson

Self-diffusion coefficients have been measured for iron in wustite (700° to 1000°C) and for cobalt in cobaltous oxide (800&apos; to 1350°C) by means of radio-isotopes. Both sectioning and decrease of surface-activity methods have been used and a particular shortcoming of the latter method has been disclosed for wustite. The effects of oxygen pressure on the composition of cobaltous oxide and on the diffusion coefficient of cobalt in the oxide have been established. The effect of the preparation conditions of the cobalt-oxide specimens on the diffusion measurements has been determined. It has been concluded that diffusion of cobalt occurs via vacant cation sites and that a cobaltous-oxide disk reaches equilibrium with its environment in a time which is very short compared with that needed for diffusion of the radio-isotope. THE availability of radioactive isotopes in recent years has permitted the measurement of many self-diffusion coefficients which were hitherto impossible to determine. With the advent of radioactivity, the method of measuring diffusion coefficients has also changed. The earlier sectioning method by which the concentration profile of the diffusing species was determined in the diffusion specimen has often been rejected in favor of the decrease of surface-activity method first developed by Von Hevesy,&apos; and extended by Steigman, Shockley, and Nix,&apos; and Raynor, Thomassen, and Rouse." With the experimental arrangement where an infinitely thin deposit of tracer is diffused into a semi-infinite cylinder, the decrease in activity of the active face due to diffusion of the tracer into the specimen is measured by counting the active face before and after diffusion. Under these conditions, the appropriate solution of the diffusion equation is R = e*(l-<K/Z) [1] where R is the ratio of the counting rates after-to-before diffusion, $& is the Gaussian error integral of \/Z, and Z is defined as Z = p2Dt, p being the linear-absorption coefficient of the detected radiation in the diffusion specimen, D the self-diffusion coefficient, and t the diffusion time. In the older sectioning method the concentration distribution of the tracer is found after diffusion by analysing thin parallel sections cut from the active face of the diffusion specimen. As seen from the solution of Fick&apos;s second law for the distribution of an infinitely thin source after diffusion into a semi-infinite cylinder the log of the concentration plotted against x&apos; will be a straight line of slope log e/4 Dt. The diffusion coefficient thus can be calculated directly. The units of the tracer concentration are completely arbitrary, since any change in units does not change the slope of the log c vs x2 lot. Although both of the above methods have been widely used, a thorough experimental comparison does not appear to have been made. Birchenall and Mehl,&apos; during an investigation of the self-diffusion rate of a and r iron by the surface method using Fe" tracer, diffused and sectioned one welded couple. The diffusion coefficient from this run agreed with those by the surface method. Buffington, Bakalar, and Cohen," during an investigation on a and y iron

Jan 1, 1955

By L. G. Austin

If S (x) is the specific rate of breakage of size x and B (x, y) (see Table 1 for Nomenclature) is the cumulative breakage distribution function, the Herbst Fuerstenau2 assumption is that Inserting this assumption into first-order grinding equations it has been shown" that there are then simple analytical solutions to the grinding equations, as shown in Table 2. In this context "plug flow" is defined as all material passing through the mill in the same length of time, the residence time t. The term "n fully-mixed" means that the mill has a residence time distribution" which can be treated as if the mill were n equal fully mixed sections in series. In addition it has been shown that the assumption of Eq. 1 also leads to simple analytical solutions for closed-circuit, steady-state grinding with an ideal classifier at the mill exit returning oversize to the mill feed. The solutions are also given in Table 2, with µ as the ideal classifier cut size, C the circulation ratio, and Q (x) the size distribution of the circuit product. The specific surface area of these size distributions is where the density-shape factor k. is assumed to be constant with size. In practice, most methods of size analysis or area [ ] measurement have a finite lower limit, so that the integration is really performed from xm, to xm, For a realistic degree of grinding the value of So is dominated by the fine sizes because of the 1/x factor. Consider batch grinding of a feed of size y to y + dy. The batch grinding solution given in Table 2 is then [ ] The fresh surface area produced is the area of the fines plus the area of the original material remaining minus the area of the original material, [ ] Little error is introduced by neglecting the last term on the right-hand side since at t = 0 it is zero, and when t is significant the first term dominates because of the fine sizes produced. In addition, the form of P (x, t) for [ ] realistic values of S (x) 2t ax° is such that the major fraction of fresh surface produced is contained within the bottom 10% of weight of broken material, that is, for sizes for which P(x, t) < 0.1 (providing y is not near x.,). Thus for surface area purposes Eq. 3 can be approximated by [ ] becomes providing again that the feed sizes are coarse relative to xmi. Then the fresh area production is and W is the mass in the mill. The rate of surface production is WK, and the specific surface is Kt. From Table 2, the specific energy is mt/W so surface production per unit energy is [ ] Surface produced per unit energy input = KW/m (8) where m is the energy input of the mill. Clearly, identical reasoning applies to open-cycle, plug-flow grinding, with t replacing t, and Eq. 8 is obtained again. Considering open-cycle, fully mixed grinding, similar reasoning gives P(x) - t S(x) for surface area purposes, and

Jan 1, 1974