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By R. K. Hopkins

For a great many years fusion welding has been used in and around petroleum refineries, but it is only within six or seven years that the more important pressure vessels have been constructed by this process, because before that time the properties of fusion welds would not with-stand such rigorous service. Since the adoption of fusion welding for the construction of this equipment, the refiner has required pressure vessels suitable for operation over a very wide range of conditions. On the one hand, vessels must operate at extremely high temperatures, in the range of 950° to 1000° F., and on the other vessels must be suitable for operation at sub-zero temperatures down as low as -75° F. Also, vessels are required suitable for operation at pressures upwards of 1000 lb. per sq. in., and as low as a 29-in. vacuum. In addition, vessels are required to withstand the attack of very corrosive mediums. It has been a rather difficult task for the fabricator to meet all of these require-ments, but as a result of the extensive research carried on by steel manu-facturers and several fabricators of class I vessels, practically all of these demands have been met. In order to meet the refiners' requirements for this wide range of conditions, the fabricator has worked closely with the metallurgist. This has enabled him to choose the proper base metal for each service condition, but it has been the fabricator's problem to develop a welding rod and welding technique, as well as a heat-treatment, that would produce in the welded vessel characteristics in and around the weld comparable to those in the base metal. Where corrosion is not a serious problem, and where the operating temperatures are not too high, ordinary carbon steel has served very well. But when corrosion is serious, and/or when the temperatures are either very high or very low, it is necessary to use some form of alloy steel to give satisfactory service.

Jan 1, 1935

By D. J. DeLazaro, W. Rostoker, R. E. Riley, M. Hansen

At very low concentrations, carbon dissolves interstitially in molybdenum resulting in a linear expansion of lattice parameter with increase of carbon in solid solution. Geometrical consideration of the relative size of carbon atom to size of interstice approximately predicts the observed volume expansion. THE element carbon, occurring in molybdenum, either as an unavoidable impurity, or purposely added, markedly affects the mechanical properties of molybdenum. Several investigators1-" have noted the occurrence of molybdenum carbide at the grain boundaries of cast molybdenum, and Fischer and Rengstorff' have shown that this intergranular constituent can induce intergranular brittleness. The appreciable change, with temperature, of the solid solubility of carbon in molybdenum, found by Few and Manning," suggests other important effects upon mechanical properties. Finally the relative atom diameters of carbon and molybdenum (the ratio of atom diameters is 0.59) are indicative, by analogy with iron (ratio of carbon atom diameter to iron atom diameter is 0.65), of a possible interstitial solid solution. If carbon dissolves intersti-tially in molybdenum, anelastic effects and age-hardening effects, similar to those observed in other body-centered cubic metals, could be expected. However, from the work of Sykes, Van Horn, and Tucker," it might be inferred that carbon forms only a substitutional solid solution with molybdenum. These considerations made it desirable to study the effect of carbon on the lattice parameter of molybdenum in greater detail than heretofore. All heat treating was done in a high temperature furnace7 built specifically for heat treatment of molybdenum. The temperatures and times of heat treatment were determined from the a solubility limit for the Mo-C system." These heat treatments were carried out in argon which was purified with respect to N, and H.!0." Purification of the argon was accomplished by passing the argon over hot (750°C) titanium and then through a magnesium perchlorate drying tower. As many as six samples were run concurrently and quenched at different times without interrupting the furnace cycle. Both the purity of the furnace atmosphere and ability to quench samples from any temperature up to 4000 °F were of prime importance in the successful heat treatment of samples for the lattice parameter investigation. Carbon Analysis The Leco Combustion apparatus was used for carbon analysis. This apparatus is capable of an accuracy of & 0.003 pct for a 1 g sample in the range below 0.10 pct. However, it is felt the degree of accuracy is somewhat better than this limit in view of the consistency evident by cross checking a number of determinations. The Battelle Analytical Laboratory consistently reported check results to within k 0.001 pct C. X-ray Measurements The molybdenum wire specimens (20 mils) were carefully filed down to about 12 mils and then etched with nitric acid to 8 mils. Carbon analyses were made on specimens before reduction in size; however, since there was no evidence of surface carburization, it is probable that the carbon was uniformly distributed throughout the sample. The resultant specimens were strain free and small enough to minimize the error in the lattice parameter determinations due to X-ray absorption. Diffraction patterns were obtained with a Debye-Scherrer powder camera 114.59 mm in diam utilizing the Straumanis type asymmetrical film mounting. Nickel-filtered copper. Ka radiation and zirconium-filtered molybdenum Ka radiation were used. Line positions on the film were measured with a traveling microscope equipped with a vernier scale permitting measurement to 0.001 cm. The average of five readings was used to determine the position of each line. The lattice parameter for each specimen was determined from the data by the extrapolation method of Bradley and Jay, Y.e., the intercept at 90" of the straight line obtained by plotting the apparent lat-

Jan 1, 1953

By T. G. R. Rawlins, L. E. Brosselard

Direct platinum carbon replicas have been used to study substrates prior to growth and after initial nu-cleation of the layer. Replicas have been directly stripped and correlations have been made with other replica techniques. Mechanically polished substrates are very flat with occasional areas of "dirt", which is apparently silicon. When such substrates are etched in hydrochloric acid gas, scratches and hexagonal defects (thought to be polishing compound) are revealed, together with areas of apparent redeposition. Chemically polished substrates are much less flat although much more homogeneous in appearance. Heat treat)nents have no effect. Nuclei deposited on these surfaces, using a normal silicon tetrachloride system, have been examined. Nuclei on mechanically polished subsrates occur in random areas only, whereas on chemically polished surfaces they are more uniformly distributed. THE interest in the study of thin films and now their rapidly expanding use in many fields has increased over a period of years. Much of this current interest, particularly in the field of single-crystal or epitaxial thin films, has derived its stimulus not only from possibilities for application but also from fundamental electron-microscope studies which have indicated the manner of growth of such layers from the initial nuclei.1-5 Previous work involving study of growth processes may be divided into three categories. First, vacuum deposition, primarily of metals, on a variety of substrates has been studied by electron microscopy and electron-diffraction techniques. These have shown the three-dimensional aspect of nucleation and the remarkable liquidlike behavior of the growing nuclei which coalesce to form a complete layer. More recently this work has been expanded in two ways: by growth inside the electron microscope this has obvious advantages for kinetic studies; and by growth in ultrahigh vacuum—much of this work has been done on semiconductors.8-11 It is of interest to note that metals, specifically gold deposited on rock salt, will form only polycrystalline layers in ultrahigh vacuum unless the substrate surface is contaminated.12 In general the mode of growth for vacuum-deposited semiconductor films is the same as that for metal films. The second method of growth that has been studied is that of electrodeposition.13,14 The little that has been published suggests that the mode of growth is similar to the vacuum case, except that irregularities may perturb the electric field and disturb the growth process. The third and last primary method of growing epitaxial layers is by use of chemical methods. These techniques have generated much interest and much has been published on them. Considerable electron-microscope work has been done on the defects in the layers; however comparatively little work has been undertaken on the initial nucleation and early stages of growth. Confining ourselves now exclusively to the case of silicon, Charig et a1.15 have considered the growth of silicon on silicon substrates by hydrogen reduction of trichlorosilane, and studied the films by means of replica and transmission techniques. Subsequently they have studied the same system for growth on quartz,16 and with the addition of pyrolysis of silane on quartz and alumina.17 Also Manasevit el al. have studied deposits using silane and silicon tetrachloride on sapphire.18-14 Finally, in this volume studies of deposition on beryllia and sapphire are reported. Although Revesz and Evans20 have examined complete layers, grown by means of the widely used silicon tetrachloride reaction, no detailed examination similar to that published by Charig and Joyce21 for the trichlorosilane system has been given. Therefore this investigation has examined substrate surfaces prior to growth and then nucleation on these surfaces for the silicon tetrachloride system. EXPERIMENTAL In this work the surfaces have been studied by means of standard replica techniques. Because of the very real problems associated with artifacts introduced,22 different replicas and different cleaning

Jan 1, 1967

By P. B. Butler

IN common with others, the zinc industry found itself after the war largely over-capitalized, a condi-tion which still exists notwithstanding that Tri-State production has doubled since prewar days. This is partly due to the abnormal war demands and partly because of the war impetus given to research, a stim-ulus that because of its success has persisted (and probably will persist) constantly since then. Improved practice and metallurgy resulting from research are today responsible for an increased pro-duction in the Tri-State district of 15 per cent over the practice of five years ago, and the giant strides in this respect in the western states together with the more moderate advance in the Tri-State district has thrown out of equilibrium the more or less balanced production of before the war and production has con-sequently gained on consumption, a condition resulting in rapidly narrowing profits in the Tri-State field, where the present average cost is actually greater than the price received for our product. Despite this condition, no thinking person will de-plore the cause or means used to achieve a greater production and, as this progress is merely the economic application of intelligent thought, it follows that equal effort and thought may be expected to produce the remedy. It has been suggested that the practical application of research has resulted in less need for man-power and consequently in the lessening of employment to common labor and to engineers. I cannot think that such is the case although personally I have received more ap-plications for employment during the last two years than ever before. But when one considers the strides production of all commodities has taken, it is difficult to reconcile this immense development with the possible increase in pop-ulation especially in view of our stringent immigration laws. Whether or not improved industrial methods are responsible for the lessening of employment the ques-tion as to the wisdom of such improvement is hardly debatable. While the Tri-State district of Oklahoma, Kansas and Missouri continues to be the leading factor in the zinc¬mining industry, gradually lowering prices during 1925 and 1926 and a. decided drop during 1927, have de-creased its production almost 20 per cent, this tonnage having largely been replaced by western concentrates.

Jan 4, 1928

By J. D. Shaw, E. B. Gebert, D. O. Noel

It is, of course, expected that manufacture of the various metal powders should involve numerous methods adapted to the specific characteristics of the metals themselves. Several methods for powdering a particular metal may be available, but the physical characteristics of the powder as produced by the several methods may be quite different. In the final analysis, the application of the metal powder dictates the physical properties that the metal powder must possess, and this, to a very large extent, will determine which of several possible methods of manufacture is adopted. Sometimes powder of the same metal is produced by more than one method, each powder having specific characteristics that make it particularly suitable for its application. For purposes of classification, the various methods of powdering metals can be grouped as follows: 1. Machining. 2. Milling (by ball mills, stamps, attrition mills, etc.). 3. Shotting (pouring molten metal into water, air, etc.). 4. Granulation (by stirring molten metal while solidifying). 5. Atomizing (disintegrating by steam, compressed air, etc.). 6. Condensation of metal vapor. 7. Reduction of oxide powders. 8. Chemical precipitation. 9. Electrolytic deposition. 10. Sintering (for producing alloy powders). 11. Formation of an alloy followed by dissolving or otherwise removing one of the alloying constituents. Machining Turnings, cuttings, filings, etc., are, of course, produced as by-products in machine work of various kinds, and find certain applications as metal powders. For example, iron powder so produced is used extensively as a reducing or precipitating agent and as a constituent in certain hard-finish cements.

Jan 1, 1938

By D. O. Noel, E. B. Gebert, J. D. Shaw

It is, of course, expected that manufacture of the various metal powders should involve numerous methods adapted to the specific characteristics of the metals themselves. Several methods for powdering a particular metal may be available, but the physical characteristics of the powder as produced by the several methods may be quite different. In the final analysis, the application of the metal powder dictates the physical properties that the metal powder must possess, and this, to a very large extent, will determine which of several possible methods of manufacture is adopted. Sometimes powder of the same metal is produced by more than one method, each powder having specific characteristics that make it particularly suitable for its application. For purposes of classification, the various methods of powdering metals can be grouped as follows: 1. Machining. 2. Milling (by ball mills, stamps, attrition mills, etc.). 3. Shotting (pouring molten metal into water, air, etc.). 4. Granulation (by stirring molten metal while solidifying). 5. Atomizing (disintegrating by steam, compressed air, etc.). 6. Condensation of metal vapor. 7. Reduction of oxide powders. 8. Chemical precipitation. 9. Electrolytic deposition. 10. Sintering (for producing alloy powders). 11. Formation of an alloy followed by dissolving or otherwise removing one of the alloying constituents. Machining Turnings, cuttings, filings, etc., are, of course, produced as by-products in machine work of various kinds, and find certain applications as metal powders. For example, iron powder so produced is used extensively as a reducing or precipitating agent and as a constituent in certain hard-finish cements.

Jan 1, 1938

By A. V. Deforest, A. R. Anderson, C. W. MacGregor

ONE of the important problems in the design of structures and machine parts subjected to rapidly applied loads is the determination of the strength and ductility of the material itself under such conditions. Realizing the need for such information, many investigations have been carried out in the past. Most of these experiments have been made using notched bars in bending either with the Charpy or Izod impact machine. While yielding valuable data on the effect of notches in the energy absorption of the material and on the condition of heat-treatment, this form of test does not lend itself readily to a determination of fundamental stress-strain relations. For this and other reasons, more attention has been paid recently to the tension impact test. An extensive literature already exists on the subject and only those investigations somewhat closely related to the present paper will be mentioned. For a more comprehensive list, the reader is referred to the Symposium on Impact Testing of the A.S.T.M., 1938. The work of Mann,1 Clark and Dätwyler,2 Nádai and Manjoine,3 Winlock and Leiter,4 Deutler,5 and Brinkmann,6 however, deserve especial mention. In most of the previous investigations on the subject, when stress-strain relations were obtained under high rates of loading, it was the common procedure to determine stresses based on the original area and strains based on the original length. Stress-strain curves were then constructed by [P OL, plotting So = Ao as a function of eo = Lo] where A o and Lo refer to the original area and gauge length, respectively. At present there exists a paucity of true stress-strain data in the impact problem. The advantages of constructing so-called true stress-strain curves in order to reveal the fundamental physical properties of the material for the tension test were discussed recently by one of the authors' It was shown that the true physical behavior of the material in the tension test could best be represented [by plotting the average true stress, S = P as a function of e = q' = log o = log Log] where P, A, L, S, e, and q' refer to the load, the cross-sectioned area while the load P is applied, the actual gauge length, the average true stress, the true strain and true reduction in area, respectively. Besides giving a better physical representation, the method has certain additional advantages in that the stress-strain curve so plotted becomes a straight line from the maximum load point to fracture. The construction of such- curves, however, while comparatively simple for the slow rates of loading customarily used in the tension test, presents certain difficulties for tensile impact in that both loads and diameters are required throughout the test. The continuous measurement of test-piece

Jan 1, 1941

By J. L. Bruce

For a number of years, from an economic standpoint, the ventilation of the Colorado mine of the Davis-Daly Copper Co. has been a difficult problem. The development through the main hoisting shaft was pursued for a number of years and to a depth of about 2500 ft.( 762 m.)before large bodies of good commercial ore were opened up. This lack of early success in exploration restricted expenditures for ventilation, with the result that the present orebodies were opened up at a depth of 2500 ft., and in country rock f relatively high temperature, without adequate means f supplying the development faces and the stopes with cool or fresh air of low humidity and suitable velocity. During the progress of development, two or three connections were made at higher levels with the workings of adjoining operations, but these were of a makeshift and unsatisfactory character and afforded no means for carrying downcast air from the surface, in anything approaching a direct course, to the working faces. During the latter part of 1917, arrangements were made with. the Anaconda Copper Mining Co., whereby, during the year 1918 and the early part f the year 1919, a crosscut was driven eastwards on the 2500-ft. level of the Colorado shaft and connection was made by means f an up-raise of approximately 200 ft. to a crosscut driven westwards from the 2800-ft. level of the Belmont mine, this crosscut at the top of the raise being driven at Davis-Daly expense by the Anaconda Copper Mining Co. This connection, in conjunction with a No. 11 Sirocco fan installed near the Colorado shaft, provided the Colorado mine with a maximum of approximately 30,000 cu. ft. per min. of relatively fresh air, most of which came in a fairly direct route from the downcast Belmont shaft through the No. 11 fan to the 2500-ft. level stopes. After passing through these stopes and the stopes of overlying levels, it passed through the main crosscuts on higher levels back to the Colorado shaft, which provided the upcast for this current of air. The improvement

Jan 9, 1921

By W. A. Young

Polynomial functions of temperature were obtained for the specific heats, thermal diffusivities, and thermal conductivities of zirconium hydrides containing 4 at. pct U. Three hydrides (H/Zr atom ratios of 1.58, 1.65, and 1.70) were studied over the range a" to 900°C and a fourth (H/Zr = 1.81) was studied over the range 0° to 760°C. The specific heats were determined from enthalpy measurements which were obtained using a unique drop calorimeter specifically designed for use with materials in which high temperature phase transitions and/or high dissociation pressures occur. Thermal diffusivities were measured by the flash method using a pulsed laser. The thermal conductiuities were obtained as the product of specific heat, thermal diffusivity, and density. The specific heats agree, within 10 pct, with values derived using a theoretical model in which the hydrogen and zirconium atoms are treated as Einstein and Debye oscillators, respectively. RELIABLE values of the thermophysical properties of the fuel are required to predict the operating temperatures and temperature response of SNAP nuclear reactors. Among the most important of these properties are the thermal conductivity, specific heat, and thermal diffusivity. A considerable number of investigations1-4 have been made of these properties for the Zr-H and Zr-H-U systems.* However, little of the drides, however, this direct method cannot yield meaningful results, since the hydrogen will redistribute under the influence of the thermal gradient, thus forming a concentration gradient; hence, one has a spectrum of compositions, rather than a homogenous alloy. Although the "average" composition of the material may be identical to the initial uniform concentration, the directly measured value of conductivity will be dependent on the thickness of the specimen, due to the highly sensitive dependence of transport properties on hydrogen content. This dependence is strikingly illustrated by the work of Bickel,5 who found that the electrical conduction of zirconium hydrides ranges from primarily hole conduction to primarily electronic conduction, depending upon the hydrogen content. Fortunately, the direct measurement of thermal conductivity is unnecessary, since it can be expressed as the product of the specific heat, thermal diffusivity, and density, all of which can be directly measured with considerable accuracy. EXPERIMENTAL Specimen Preparation. The combined fuel-moderator material used in SNAP reactors is a hydrided zirconium-uranium alloy containing -10 wt pct U. The alloy used in this work was representative of that used in nuclear reactors except that normal uranium was substituted for the enriched uranium required for reactor usage. It was produced by a triple-arc-melt and double-extrusion process. All specimens were prepared from a single cylindrical extrusion which contained 10.30 pet U, 89,35 pct Zr, and 0.35 pct impurities, The specimens for each composition were hydrided simultaneously with ultrapure hydrogen (10 ppm total impurities) using standard fuel production techniques which routinely yield homogeneous, crack-free fuel with negligible increases in the impurity levels. The hydrogen content of each specimen was determined from its weight gain and the density was measured by liquid displacement, Chemical analyses yielded hydrogen concentrations which agreed with the weight gain data within ±0.02 in H/Zr atom ratio) the concentrations of all other elements agreed almost exactly with the initial values after adjustment for the added hydrogen. The specimens used for the determination of specific heat were centerless ground to 2.00 cm diam after hydriding. A thin slice was carefully removed From each end for metallographic examination. In every case, this examination revealed a uniform structure as evidenced by the appearance and distribution of the two phases present in the fuel at the hydrogen concentrations used. TWO specimens (H/Zr = 1.600 and 1.632) appeared to be entirely 6 phase with equi-axed grains; the specimen with H/Zr = 1.756 showed

Jan 1, 1970

By A. A. Smith

THE application of indium to bearings was first advocated by C. F. Smart1 in 1938, for the purpose of inhibiting corrosion of certain bearing alloys. He found that a thin layer of indium plated on the bearing surface and subsequently diffused at elevated temperatures markedly decreased the rate of corrosion in lubricating oils containing organic acids. Most of his work was done on cadmium-base alloys, although he showed that the treatment was also effective for copper-lead and other alloys. Since their introduction, indium-treated bearings have found considerable use, particularly in connection with airplane motors.2 Undoubtedly the effectiveness of the indium plate in resisting corrosion depends upon its relative concentration on the bearing surface after the diffusion treatment and after prolonged use at operating temperatures. Some diffusion experiments were made on a number of base metals that had been indium plated, to determine the depth of penetration after heating for reasonably long periods of time. As only the practical aspects were desired, no attempt was made to determine actual diffusion constants. METHOD Two different series of experiments were carried out to study the diffusion of indium into several base metals. Cast cylinders about one-inch diameter were machined to 0.750-in. diameter, so as to remove all surface imperfections, and final finishing was done with oo sandpaper. The original diameters were very carefully measured with micrometers and then a layer of indium was electrolytically deposited on the cylinders from a sulphate bath.3 In the first series of experiments, about 0.002 in. of indium was deposited and diffusion was carried out at 150°C. for a period of 500 hr. in a thermostatically con- trolled electric oven. After the diffusion period, successive layers were machined from the specimens and analyzed spectro- graphically and chemically-the spectro- graphic examination serving as a guide to determine which layers should be chemically analyzed. Apparently there was some surface oxidation of the indium plate due to the long heat-treatment, which may have had some effect on the diffusion. Also, the cadmium alloy showed some incipient melting, which probably was caused by the formation of a low-melting eutectic. To eliminate these factors, another series of experiments was carried out by heating in an electrically heated, automatically controlled, oil bath at lower temperature. The preparation of the specimens was the same as before except that a layer of indium 0.0001 in. thick, an amount commonly used commercially on bearings, was deposited on each of the cylinders. Diffusion was carried out at l20°C. for a period of 200 hours.

Jan 1, 1944

By A. Berthet

In the past two decades, of the successive steps leading from iron ore to finished steel, the production of liquid steel and its preparation for the casting operation has probably undergone the most drastic evolution. The reason for this is the development of oxygen converters (top blown and later on bottom blown) and UHP arc furnaces which are replacing the conventional processes such as the open hearth and basic Bessemer processes. These new processes have been great advances, but it would certainly be quite hazardous to assert that the situation will remain the same for the coming decades. In fact, with the advent of continuous steelmaking a new generation of processes has now arrived. The research and development programme at IRSID on continuous an a steelmaking began ten years ago has progressed systematically in the following stages : First a comparison was made on a small pilot scale of various techniques : for example, refining in a runner by top or bottom blowing, pulverizing the metal stream by oxygen, using the mixing effect of a vortex, etc. but for various reasons these processes proved unsatisfactory and were abandoned in favour of a method using the so -called "emulsion refining".

Jan 1, 1972

By J. B. Clark, A. J. McEvily, R. L. Snyder

The nonuniform distribution of precipitate particles has been recognized as a leading factor contributing to the relatively low fatigue resistance of aluminum alloys. The structure of many of these alloys is characterized by narrow precipitate-free zones adjacent to the grain boundaries. Alloys with such zones exhibit a tendency for brittle inter crystalline fracture. The interrelation between this type of structure and mechanical properties was investigated in an Al-10 wt pct Mg alloy. It was found that deformation during fatigue occurs preferentially along these zones and cracks initiate there. In Al-10wt pct Mg, the zones were found to be supersaturated even after extensive general precipitation and are due to the absence of proper precipitate nuclei in the region near the grain boundaries. Cold working the alloy prior to aging improves the mechanical properties by inducing precipitation within the zones and also by jogging of grain boundaries. The mode of fracture is changed from brittle inter crystalline to more ductile trans granular fracture. THE process of fatigue is highly structure sensitive, with the strength of the whole often dependent upon some localized discontinuity, either geometrical or metallurgical in nature. Much has been learned about the role of geometrical discontinuities, e.g., notches, in fatigue, but with the exception of the effects of inclusions or the shapes of carbides, relatively little is known about the specific effects of discontinuities in metallurgical structure such as nonuniform precipitation. In most age-hardening aluminum alloys, metallo-graphic studies have shown that the extent of precipitation adjacent to grain boundaries is much less than that which occurs in the interior of the grains. The width of these almost precipitate-free regions, which are sometimes called denuded zones, and the extent of solute depletion within them, are dependent upon the particular alloy and its aging treatment. It has been observed1 that these zones are relatively soft with the result that plastic deformation takes place preferentially within them. It has also been shown 2-4 that there exists a tendency for intercrys- talline cracking in fatigue when such zones are present. It is of interest to note that Broom et al.2,3 were able to reduce the incidence of this type of failure in an A1-4 wt pct Cu alloy by stretching the material 10 pct prior to aging. In the present study, the effects of precipitate-free regions on the fatigue properties of an A1-10 wt pct Mg alloy were studied in detail, and the effects of deformation prior to aging on the nature of the precipitation process as well as on fatigue properties were also investigated. MATERIAL AND PROCESSING An A1-10 wt pct Mg alloy was selected for this study, because it was known that well-defined precipitate-free regions along the grain boundaries are readily obtained in this alloy after aging at 200oC.5 The starting materials were 99.998 pct A1 and singly sublimed magnesium of about 99.9 pct purity. The aluminum was induction melted in a graphite crucible, and then the magnesium addition was immersed until dissolved. Chlorine gas was then bubbled through the molten alloy for 4 min to degas the melt, after which the melt was cast at a pouring temperature of 730" to 760°C into a cold, graphite-coated, tapered steel mold. Since A1-Mg alloys are difficult to homogenize,5 special care was taken to obtain a uniform composition. Two-in. cubes were cut from the ingot and heated at 446°C for 30 min. These cubes were then hot forged approximately 35 pct in each of the three cube directions and homogenized for 16 hr at 446°C. Sheet specimens were then obtained by pressing 40 pct and rolling 35 pct per pass with reheating between reduction steps to a final thickness of approximately 0.10 in. The sheet was then solution treated for 16 hr at 446°C and water quenched. The age hardening behavior of this material at 200°C was then determined, and the results are shown in Fig. 1. The age hardening of this alloy when subjected to cold work prior to aging is also shown in this figure. Preliminary work indicated that extensive deformation after quenching was required to affect drastically the precipitate-free regions in this alloy, and a rolling reduction of 50 pct was chosen. For purposes of comparison the following three conditions were studied: a) Solution treated, quenched, and aged 20 hr at 200°C

Jan 1, 1963

By W. A. Oates, J. A. Lambert, P. T. GaIIagher

Monte Carlo methods have been used to compute the arrangements of interstitial atoms dissolved in tetrahedral sites in bcc lattices. It is assumed that the presence of an interstitial atom "blocks " a certain number of neighboring sites and prevents their occupancy. Sites "blocked" by more than one filled site are allowed for. The computed values of. the mean occupation number (defined as the ratio of the total number of sites blocked to the number of solute atoms are used to calculate the configurational entropies of the solutions. These entropies are compared with those resulting from previous theoretical studies of this problem and also with available experin~ental data for the p Zr-H, Nb-H, V-H, and Ta-H systems. Evidence is also given that the "blocking" explanation of low limiting compositions in these systems, rather than this being due to initial limitations on the number of sites available, is probably correct. THE ideal partial configurational entropy of mixing of an interstitial solute in a metal is given by: where p is the number of interstitial sites per metal atom and Xi is the atomic fraction of the interstitial. For the bcc lattice. which we shall be concerned with in this paper, the interstitial positions are shown in Fig. 1. It can be seen that for the tetrahedral sites, p=6. whereas for the octahedral sites, p = 3. Different emphasis has been placed on the relative importance of energy and entropy effects in determining deviations from ideality in interstitial solid solutions. In some cases the same system, e.g., Fe-C, has been described by the contradictory regular and athermal solution models indicating that the enthalpy and entropy functions, derived from equilibrium data, are frequently not accu.rate enough to differentiate between these treatments. However, for certain metal-hydrogen solutions the equilibrium data is available over sufficiently wide ranges of temperature and composition to permit a reasonably accurate determination of the compositional variation of the heats and entropies. Hoch&apos; has attempted to interpret the results of interstitial solid solutions in terms of a regular solution model. In the case of the Ta-H system where 13 = 6, this model entails fitting the experimental relative partial entropies of solution, asH, to the equation: where ASgs is the relative partial excess entropy of solution of hydrogen. Hoch found that the results of Mallett and Koeh1 could be fitted to this equation with an approximately constant value of AF up to XH = 0.25. However, it is apparent from the solubility isotherms in this system which become asymptotic to the composition TaH that, since (Xh /6 - ~Xh ) becomes infinite only at TaH6, it is necessary that AS<&apos; tends to infinity at TaH. In other words, the low saturation composition of TaH, instead of the anticipated TaH,, eliminates the possibility of applying regular solution theory to such systems. Rather large negative excess configurational entropies must exist at higher hydrogen concentrations in order to explain the lower saturation values. To account for these low limiting compositions and excess entropies two distinctly different approaches have been followed. Rees and many others1-l2 have assumed that not all interstitial sites are crystallographically equivalent with respect to the interstitial addition; that is, in Eq. [I] p is less than the value anticipated from geometrical considerations. To describe, say, a bcc metal-hydrogen system with a limiting composition of MH by this approach one would consider that p = 1 in the first instance instead of p = 6.&apos;j3 In some cases, nonintegral values of B have been taken in order to improve the fit with the experimental data over limited ranges of composition. The other approach which has been used to explain the low saturation compositions is to assume that, although all sites are available for occupancy, strong repulsive interactions exist between the neighboring interstitial atoms, and hence occupancy of any site excludes or blocks a certain number of neighboring sites from being occupied. Earliest treatments of this concept considered the exclusion of an integral number, of nearest-neighbor sites from being occupied at all concentrations. In this case, the partial configurational entropy is given by: These early treatments failed to allow for the overlap of the blocked sites which will arise at all but the very lowest concentrations. More recently attempts have been made to calculate the effect of this decrease in the number of blocked sites on the configurational entropy. Using the quasichemical treatment of interstitial solid solutions as given by Lacher and assuming that an infinite repulsive interaction energy existed between the solute atoms. atom obtained an approximate configurational entropy applicable to the blocking with overlap case:

Jan 1, 1970

By R. A. Davidson, L. Silverman

RESIDENTS of many urban centers are becoming increasingly aware of the obscuring effect of fume and smoke discharge from power, metallurgical, chemical, and other industries; and they, as well as the legislatures of these affected cities, are agitating for cleaner air. Management&apos;s most pressing problem is to find an economical way to reduce process effluents in response to the growing pressure from population and legislative demands. The removal must be done, if possible, without handicap to the current operation, since the costs of relocating are often excessive or prohibitive. In fume recovery or disposal, an important item to consider is whether or not the material being discharged has any value. If it has commercial value, the cost of its recovery may offset or aid amortization. For this reason, in making a study of the specific problem in hand, a major factor was the nature of the material emanating from the stack: in particular, its particle size, size range, and its chemical and physical composition, as well as its potential value and utility when recovered (in either a wet or dry state). Should the product have no commercial value, it must be disposed of at minimum cost in a way to prevent recontamination. Initial studies were therefore made to determine stack concentrations and volumes of material evolved from the operations. The next phase of the study concerned the physical and chemical nature of the collected fume. The third portion of this paper describes the wet and dry collector studies undertaken to recover the fume. Cleaning Requirements for Ferroalloy Furnace Operation The basic need for any effluent collection equipment is the highest possible efficiency and the lowest tolerable resistance when the power consumption involved is considered. Since the electric furnace effluent is largely composed of fume of small size (less than 0.5u), it has high light obscuring properties, and even low concentrations will cause some loss of visibility and be evident to nearby residents. The permissible limit for fly ash emission in many cities is based on a weight value (viz, approximately 0.4 grains per cu ft), but the smoke density values are dependent upon a shade of color. In the case of the Los Angeles County code, emission is restricted to pounds per pound of material processed per hour basis (but not exceeding 40 lb per hr for any one given plant operation). If an average particle size of the fume from ferro-silicon alloy electric furnaces is assumed to be 0.4u (as shown later, this is the approximate mean size) and an average loading of 1 grain per cu ft (stp), each cubic foot of stack gas will contain approximately 75x10 10 particles (based on assumed, and confirmed, spherical shape and a standard deviation of unity). When it is realized that the air in metropolitan areas, which are also general industrial areas, contains approximately 5x108 particles, the tremendous light scattering effect of this concentration becomes apparent. Consequently, nearly 100 pct collection would be necessary to equal the average concentration. Fortunately, however, discharge from a high point above ground (50 to 100 ft) will result in at least a thousandfold dilution, or the stack concentration reaching the ground in the foregoing case might result in a ground concentration of &apos; particles. If the concentration at the source could be reduced by a factor of 100 (99 pct efficiency of collection), then a concentration of 75x10" particles would be diluted to 7.5x10&apos; which would be very satisfactory. An efficiency of 90 pct (factor of 10 decontamination) at the source would result in a discharge of 75x109 articles which upon dilution yields 75x10 which is still 15 times the general air value. Another approach to this consideration is to use the value of concentration of 0.005 grains per cu ft for the value of a visible effluent as cited by Kayse.1 To attain this value with an average loading of 1 grain per cu ft would require an efficiency of 99.5 pct. Since the foregoing value is not based on any reported size of fume particles, it is felt that the numbers&apos; approach given previously is more reliable. These calculations serve to indicate the desirability of thorough cleaning, preferably at the source, and with efficiencies well above 90 pct, preferably above 95 pct (dilution 1:20). One of the most important items in any control program is to reduce the concentrations as close to their sources as possible. The use of better furnace design, deeper coverage over the electrodes, and the prevention of blows or breaks in the surface all help to reduce dissemination; consequently, all of these improvements should be made, if possible, to cut down the effluent load. In addition, in order to minimize the volume of contaminated air that has to be cleaned, the furnace should be enclosed as much as possible. Test Arrangements Before fundamental studies with collectors were made, a furnace stack selected for the test program was sampled to determine the gas temperatures and

Jan 1, 1956

By J. P. Van Kleunen

9.3-1. Introduction. The demands of industry for handling materials at higher rates and at reduced costs have had a very significant effect on belt conveyor technology. Although the general appearance of the belt conveyor today is essentially the same as it has been for the past several decades, the similarity ends there. Each component comprising the conveyor system has been continuously improved through design and materials technology to enable present conveyor systems to operate at ever-increasing rates and reliability. For example, twenty years ago the strongest commercially available, fabric-reinforced conveyor belting had a tension rating of 50 lb per ply-in. Currently, fabric-reinforced belts are available with ratings of 350 lb per ply-in. and steel cable reinforced belting is available with tension ratings of 3000 lb per in. of belt width (3). These technological advances have therefore enabled the belt conveyor to not only attain acceptance as a preferred method for in-plant movement of bulk materials, but also to be used for long distance transportation of bulk materials, particularly in areas of adverse terrain. The information contained in this chapter is intended to assist the reader to make basic decisions regarding the use and application of a conveyor system. The design of an overland conveyor transportation system or an in-plant process conveyor is accordingly beyond the scope of this article. Many comprehensive technical works covering the design of a belt conveyor are currently available from the manufacturers of materials handling equipment (6, 7, 9, 11). The design handbook published by the Conveyor Equipment Manufacturers Association is particularly recommended (2). 9.3-2. Conveyor Capacities. One of the most important factors in determining the proper size of a conveyor system is the rate at which material must be handled. Although tables and charts are available to the conveyor designer to assist him in his selection of conveyor components, only the

Jan 1, 1968

By John Tabor, R. B. Spivey

INTRODUCTION Shafts have been sunk in a number of ways. By and large, however, most have been sunk by drill, blast, muck, and slip form methods. Most methods have provided satisfactory means of completion. The major drawbacks to the conventional methods have been: (1) un- reliable cost estimates for completion; (2) unreliable rates of progress; and (3) unreliable integrity of the shaft lining. Recently , engineering design studies have been done that strongly indicate, that, by the use of existing proven technology, shafts can be completed faster, more cheaply, safer, and with a better lining that has been possible using conventional methods. THE METHOD The alternate method utilizes the techniques of big hole drilling, shield excavator tunnelling, and pre-cast concrete lining. Big Hole Drilling - Drilled shafts have been completed for several years and I don&apos;t intend to address the technology, only to point out the use of drilled shafts is an alternate method of shaft sinking. Ideally, in a multi-shaft mine, the first shaft would be blind bored and cased to a size, whereby it can handle the muck from the excavation of subsequent shafts. It should be an 8&apos; to 10&apos; completed diameter. The cost of such a shaft in sedimentary formations compares favorably with a conventional shaft. However, it is much faster. For additional shafts, the role of the drill is to complete small holes (3&apos; to 4&apos;) to the desired horizon. This is merely a hole to allow slashing to a larger size with easier muck removal from an underground drift connected to the previously drilled shaft. The smaller holes, just as the larger one, should be cased and grouted into place so as to assure a smooth open hole. SHIELD EXCAVATOR Theory and History - Tunnelling Shields have been in common use for several decades, and steam powered shields were used in England over 100 years ago. The purpose of the shield is to provide temporary ground support, protect personnel during operation, and to house the excavating equipment. The excavating equipment approach has varied from attempts at a continuous or semi-continuous boring machine, the tunnel mole, to a mechanical or hydraulic digging arm. The shield may or may not be an active digging element. If not active, the boring tool or cutter is advanced ahead of the shield to full diameter and then the shield is moved up to the cutter, or the shield advances along with the cutter. If active, the cutter opens a conical free face, and the shield is either jacked forward so that the leading edge spalls rock to the free face or poling plates are thrust forward to spall to the free face. In either approach, various methods of muck removal are incorporated into the shield. These may be gathering arms, conveyors, or with moles, hydraulic transport, where the cuttings are sufficiently fine. The efficiency of the muck removal system is critical as excavation is rapid and the system could easily become muck-bound. Application to Shaft Sinking- Technically, the use of a shield to sink vertical shafts is no different than a tunnelling operation. It is, in fact, less complex in most respects. Muck and water removal are difficult in a blind-driven shaft. However, where an underground opening is available for handling muck and water, rapid shield sinking is possible, using a pilot hole for muck and water to fall through and be removed from below. The Tabor Mining Shield is basically a caisson approach, whereby, the ground is only seen at the face or bottom. It does have hydraulic jacks to force it downward and has segmented cutting edges which allows forepoling ahead of the shield. The hydraulic excavator allows mining in the center of the face around the previously drilled hole with the forepoling segments slabbing to this free face. First, the drilled hole is cased and grouted. The machine is a tube with an excavator. The work cycle is as follows: (1) The casing and grout are cut at predetermined intervals by shaped charges. (2) The casing is crushed by the excavator, removed from the hole and hoisted to the surface. (3) The excavator then rips out a face around the hole by digging the formation in tension and breaking the pieces

Jan 1, 1982

By R. F. Hehemann, S. T. Zegler

The w phase in zirconium alloys containing more than G pct Nb can form in a difjUsionless manner during quenching or with composition change during aging at temperatures below 550°C. The latter treatment establishes a metastable equilibrium between an w phase containing 5 pct Nb and a ß phase containing 45 pct Nb. The superconducting transition temperat~ire of quenched Z?-10 pct Nb and Zr-15 pct Nb alloys is increased by aging at 375°C and both alloys exhibit the same Tc after prolonged aging. The superconducting transition both in the quenched and in the aged alloys is relatively sharp (occurring ovev a temperature range of <0.5°K), and is interpreted in terms of the microstructure that characterizes the ß-w) mixture and proximity effects postulated for thin-film superconductors. At elevated temperatures alloys of the transition elements from groups 4 to 6 in the periodic table exhibit wide ranges of mutual solubility in the bcc structure. Except for compositions that are high in a group 4 element? the bcc structure can be retained at room temperature by quenching and several systematic studies of the influence of composition on superconductivity in bcc alloys have been reported.&apos;&apos;2 The high-temperature bcc form (ß) of the group 4 elements—titanium, zirconium, hafnium—is not retained upon quenching but instead transforms marten-sitically to the hcp structure (a). The Ms temperature can be lowered by alloying the group 4 metals with a solute from a higher group and with a critical amount of solute the martensitic transformation can be suppressed. The 3 phase, however, is not retained without modification. Rather, a metastable phase, w (in particular in titanium- and zirconium-base alloys), occurs over a fairly wide range of compositions above that required for suppression of the martensitic transformation.3,4 As in the formation of a, the transformation of ß to w occurs in a diffusionless manner during quenching and the temperature for its initiation (ws) is lowered by increasing solute content. The ws temperature can be well below room temperature and reversibility of the transformation allows w that forms on cooling to low temperatures to disappear again when the sample is returned to room temperature.5 In Zr-Nb alloys the w phase can also be formed during aging at temperatures below 550°c.5 In this treatment, a metastable equilibrium is established between a niobium-rich ß phase and a zirconium-rich w phase. The present study of Zr-Nb alloys was undertaken to examine the influence of microstructure: as controlled by composition and heat treatment, on the superconducting transition temperature, Tc. MATERIALS AND PROCEDURE The alloys were prepared from zirconium crystal bar that was spectrochemically pure except for 30 ppm of Al, 40 ppm of Cu, 100 ppm of Fe, and 1 ppm of each B and Pb. The niobium used was 99.5+ pct pure containing according to the suppliers&apos; (Stauffer Chemical Co.) analysis a maximum of 50 ppmO, 60 ppm of N, 30 ppm of C, 5 ppm of H, 200 ppm of Ta, 50 ppm of Mo, and 15 ppm each of Fe, Co, and Ni. The alloys were prepared in the form of buttons weighing approximately 4 g by arc melting on a water-cooled copper hearth in an inert-gas atmosphere. All alloys were heat-treated in the ß-phase region for 3 days at 1000°C and water-quenched. For heat treatment, the specimens, 0.5 cm in diam by 1.0 cm long, were wrapped in zirconium foil and sealed in an inert-gas atmosphere inside quartz capsules. Quenching was done by breaking the capsules under water. Subsequent aging treatments of Zr-10 pct Nb and Zr-15 pct Nb alloys were done by placing the specimens in a lead bath at 375°C and quenching in chilled brine. The phases present in the quenched and aged alloys were identified by optical and X-ray metallography. X-ray powder patterns were taken at room temperature with a 114.6-mm-diam Debye-Scherrer camera and CuK-a radiation. Superconducting transition temperatures were determined from magnetic-permeability measurements6 in a 10 oe field. The superconducting transition temperature was determined from a plot of galvanometric deflection as a function of temperature by extrapolation of the linear portion of the curve to zero deflection. In Figs. 1 and 3 the arrows indicate the temperature range of the transitions. RESULTS AND DISCUSSION Superconductivity in Quenched Alloys. In Fig. 1 Tc for alloys quenched from 1000°C is plotted against nominal composition. Indicated also in Fig. 1 are results of metallographic studies. Alloys containing less than 5 pct Nb transform martensitically to a during quenching from 1000°C and Fig. 1 demonstrates that Tc increases significantly as the niobium concentration in a is increased. On the basis of the Bardeen-Cooper-Schreiffer theory of superconductivity7 it seems most likely that this increase in Tc results from an increase in the density of electronic states as zirconium is alloyed with niobium. This increased density of states has been revealed clearly in low-temperature specific-heat measurements.8 The martensitic transformation to a is suppressed when the alloy content exceeds 5 pct Nb. In the range from approximately 6 to 30 pct Nb, quenched alloys consist of a mixture of B and w phases and the amount

Jan 1, 1967

By Irving B. Cadoff, Ernest Levine

The crack formation and propagation in the single -phase Cu-4 pct Ag alloys were studied. The alloys were loaded in mercury to various stress levels, the mercury was removed, and the specimen examined for cracks. Cracks were found to develop below the fracture stress; the frequency of such cracks increased with increasing stress level. Some cracks were nmpropagative. Fracture in mercury was found to occur by the link-up of cracks formed at various stress levels rather than by the growth and propagation of a single crack. If the mercury environment is removed prior to a critical amount of crack formation, then continued loading results in ductile fracture. The appearance of the cracks at selected grain boundaries is related to the relative orientation of the boundaries, as are the propaga-tive characteristics of the crack. The mercury interaction appears to be one of lowering the strength of the metal-metal bonds in the high-stress area of the grain boundary. GRIFFITH&apos;S microcrack theory1 proposed a critical crack size above which a crack in an elastic material grows with decreasing energy at a stress of From his theory it was proposed that the presence of a liquid tends to lower the surface energy of the microcrack faces2 leading to a decrease in the critical crack size necessary for spontaneous fracture propagation. stroh3 proposed that the stress concentration at a grain boundary due to pile-up may initiate a microcrack at the grain boundary. petch4 and Stroh5 evaluated the stress distribution at the head of a pile-up in a polycrystal-line material and deduced that the critical crack size and hence of is dependent on the grain size. Experimental verification of this dependence was found by petch6 for hydrogen embrittlement of steel. Studies in stress-corrosion cracking7 have provided a picture of fracture which shows that initial separations occur in a scattered, independent fashion in regions of high tensile stress. A minimum or threshold stress is necessary to produce a sufficient stress concentration to initiate frac- ture. These separations join up to form a crack. The extension of fracture is largely discontinuous and consists of a joining up of cracks. In recent worka evidence of this scattered crack network was found in a Cu-Ag alloy embrittled by mercury. For the Cu alloy-Hg couple, the crack path has also been found to be dependent on the orientation of adjacent grains, and with the addition of zinc to mercury a reduction in embrittlement along with a change in fracture morphology was found.9 In this present study, a mercury-dewetting method was used to observe crack initiation and fracture morphology when a Cu-4 pct Ag alloy is deformed in mercury and Hg-Zn solutions. PROCEDURE Specimens of Cu-4 pct Ag were prepared as in previous crack-path studies.&apos; The specimens were heated at 770°C for 24 hr and water-quenched Tension tests using a table-model Instron were carried out in mercury and in various concentrations of Hg-Zn. Loading was in steps up to the fracture stress, with the load being removed and the specimen examined for surface cracks at each step. The specimens were dewetted after each load to permit examination of the surface structure and rewetted prior to continued loading. The specimens were wetted by electro polish ing in phosphoric acid, rinsing in alcohol, and then immersing in a pool of mercury. Dewetting was accomplished by flame heating the specimen for 30 sec in a vacuum. Some surface contamination was found, but not enough to obscure crack configurations and grain boundaries. RESULTS Fracture Characteristics in Mercury. Fig. 1 is a stress-strain curve showing the progressive step-wise loading of the specimen. As may be seen from the graph, the first position stopped at a is at a stress 5000 psi below the expected fracture stress of 25,000 psi. Examination of the specimen after removal of mercury showed only one crack. The appearance of this crack at a stress far below the fracture stress of this alloy in mercury did not affect the stress-strain curve in any manner. The specimen was then recoated with mercury and deformation was continued (curve b, Fig. 1) raising the stress by 4000 psi, and the same procedure re~eated. The initial crack was located and appeared as in Fig. 2 (crack lb). In this figure the crack is

Jan 1, 1964

By William Hague, Robert H. Bedford

The objects of this paper are: 1. To give cost data on drilling obtained at the North Star mine in California; 2. To describe the methods of testing drills employed there, giving the results of shop tests, and the correlated data obtained underground; 3. To demonstrate that these tests form a satisfactory basis upon which to judge rock drills. The rock at this mine is either close-grained diabase in the upper levels, or tough grano-diorite in the lower levels. The vein, which has an average dip of 23", consists of about 5 ft. of " formation" lying between walls of unaltered country rock. Of the 5 ft., solid quartz and stringers make up 18 in. of pay ore. The formation, which consists of but slightly altered country rock, is almost as hard as the unaltered walls. From 4 to 44 ft. is sent to the mill. One-third of the holes are drilled in the quartz; two-thirds in the formation. The holes are about 4 1/2 ft. deep and break on the average of 1.15 tons per hole. The average number of holes per stope drill shift is at present 5.65. The number of drill shifts throughout the mine in 1913 was 18,679, and the cost of labor for drilling, power, supplies, upkeep of machines and air lines, tool sharpening, and distribution, amounted to 33 per cent. of total mining expense. To do this work there are in commission 43 No. 12A, three No. 17V and eight No. 16V Waugh stopers, three Jackhamers, two Waugh pluggers, and 17 No. 8 and two NO. 7 water Leyners. For the past three years a record has been kept of the repairs on each machine and number of shifts operated by it. The figures given in Tables I and II for the cost of repairs per drill shift have been taken from this record. The underground air pressure is about 90 lb. The air consumption per drill shift was measured by aerometers. The leakage of pipe lines between compressor and meter is not included. From the number of pieces of steel sharpened for each type of drill during the year the average used per drill shift has been ascertained. For other supplies the figures given represent a year&apos;s average.

Jan 1, 1915

By A. A. Rauth

In this paper, the author derives a series of formulas from basic principles and illustrates the application of these formulas to practical grinding charge problems. The paper establishes the nearly proportional relationship of mill power to charge level and shows how this relationship can be used to advantage in the design, manufacture and operation of grinding mills. The procedures outlined offer a rapid means of anticipating the power requirements for a given charge or determining makeup charges for a given power reading. The theories which have been developed concerning the dynamics of grinding have been motivated largely by the desire to derive workable formulas for the accurate determination of grinding mill power. This accurate determination has always been handicapped by the large number of variables to be considered, some of which are mill diameter, length, speed, liner design, grinding charge and material density, volumetric mill loading, internal friction, and charge size and gradation. In efforts to determine mill power, attempts have been made to utilize torque concepts. The formulas derived consider the grinding mass as a rigid body and utilize factors that are based on empirical data so that the movement of the charge within the mill is compensated for. These formulas are of principal interest to the mill designer since he is charged with the responsibility of analyzing all of the variables and from them predicting a mill size and a motor size which are compatible both with each other and with the grinding job for which they were intended. The mill operator, on the other hand, has an entirely different problem. The mill dimensions are already established and so is the driving motor. The operator&apos;s main concern is to obtain maximum performance from the mill he already has. For him, the many variables which perplex the mill designer are fixed values. Mill diameter, length and speed are set and, to a large extent, so are the conditions inside the mill from which the torque formula advocates derive their empirical data. It is conceivable that if the major mill dimensions which affect power demand are held constant, and if the variables influencing empirical factors are held fairly constant, as they are in an operating mill, then the only major variable remaining which affects power is the level of the charge itself. Under this condition.

Jan 1, 1970