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By G. S. Brooks

WITH the gradual depletion of the natural-gas pools of the Kansas district, together with the uncertainty of further cheap fuel developments, some. of the western zinc companies turned to the coal fields, attracted by their permanency and the 'success of the Matthiessen & Hegeler Co. and the Illinois Zinc Co. Being then confronted with the problem of using a relatively high-priced fuel, although the ultimate' supply was very great, the fuel economies of the kilns and furnaces became of far more importance, and the more modern equipments, as a result, embodied distinct gads in this phase of the business. In addition, there have been some pronounced advances in the actual making of gas from these coals. Increased efficiency in gas making has contributed its share to lowering costs and those plants located in the higher priced fuel districts have probably been obliged to study these economies somewhat more closely. Naturally, many beneficial changes have resulted in both practice and apparatus in these sections. On, certain western coals, for instance, the actual saving in fuel consumption has been 50 per cent. By this is meant that, where about 1.75 tons were burned to smelt 1 ton of oxidized ore, 0.7 to 0.9 ton is now required. While comparisons by W. R. Ingalls1 of fuel economies in smelting in this country and Germany before the war may, in general, be correct, they do not appear to-do justice to the more progressive. American practice. It is doubtful whether German plants were substantially bettering these fuel ratios. Data furnished by Bender and Frams do not afford records averaging less than 0.9 ton coal per ton ore on materials from 40 per cent. to 50 per cent. zinc.

Jan 9, 1919

By D. D. Porter, H. G. Carlevato

Blasting to prepare orebodies for in situ leaching usually involves extraordinary conditions which sometimes require special and innovative techniques. To date several unusually large blasts have been made for leaching with results that are very encouraging. Among the factors considered when designing these large blasts are orebody shape and depth, rock properties, vibration and damage potential, loading logistics, and safety. All of these factors, and others, are discussed with an effort made to put each of them into its proper perspective.

Jan 1, 1974

By T. S. Jr. Sligh

OF all the qualities that are essential in a good cigar tobacco none is quite so important as the burn. This term is general and includes many points, the most important of which are evenness of burn, color of ash, firmness and coherence of ash, and fire-holding capacity. The fire-holding capacity refers to the length of time the leaf or cigar will continue to, glow after ignition. Chlorides tend to prevent complete combustion and products are formed thereby that are injurious to the flavor and aroma. On the other hand, the carbonates of the alkalies, particularly of potassium, aid the combustion and increase the fire-holding capacity. Barth1 thought that the harmful effect of the chlorides was due to their fusing and coating the tobacco, thereby preventing complete combustion. Schlosing,2 Nessler,3 and Garner,4 suggested widely different theories to account for the favorable action of potassium salts. In order to study carefully the action of the various salts upon the course of combustion of the cigar a knowledge of the temperature of the burning cigar is necessary. Lehmann5 who seems to be the only investigator who has made any attempt to determine the temperature of a burning cigar gives 480° C. as the maximum temperature which he recorded. It is evident, however, that his measurements were in error, probably on account of the method and apparatus used, since the lowest visible red corresponds to a temperature of about 575° C.

Jan 9, 1919

By A. H. Geisler, R. F. Mehl, C. S. Barrett

The mechanism by which the supersaturated solid solution of 20 per cent silver in aluminum decomposes has been reported in two previous publications.1'2 The analysis of streaks in Laue photograms showed that thin platelike particles (the so-called Guinier-Preston aggregates) form upon matrix planes in the early state of decomposition.1 The size of these particles is larger, the higher the temperature at which they form in accordance with nucleation theory;³ at all temperatures they increase in size with aging.2 When they have grown thick enough to afford three-dimensional diffraction, powder X-ray photograms show that the structure is not that of the equilibrium precipitate but that of a transition lattice . Sincc this is also platelike in shape and also lies upon the (III) matrix planes with the [IIO] direction in the matrix parallel to the [II.O] direction in it seems likely that the original platelet particles are merely in an early stage of formation, and thus finely dispersed. The atomic spacing and pattern on the (00.1) plane in y' are identical with those on the (111) matrix plane to which it lies parallel, and since comparable directions in the two planes are parallel, the lattices are thus coherent at the interface; an analogous coherence has been found in aluminum-copper alloys.3 The structures y' and y are both close-packed hexagonal and are oriented alike; they differ only in interatomic distances. Consequently, y' is regarded as y strained from its ordinary dimensions by the lattice coherency between it and the matrix.2 The lattice movements that transform the matrix to the equilibrium precipitate appear to be very simple.2 The mechanism of precipitation in this system is simpler and is more precisely known than that in any other system hence the system should be a profitable one for the difficult task of correlating property changes with structural alterations. The present report provides additional studies on powder X-ray diffraction from alloys of 10 per cent, 20 per cent and 30 per cent Ag, and presents the correlation of crystallographic changes with changes in microstructurc, Brinell hardness and electrical resistivity of the three alloys. Preparation of Samples and Experimental Procedure The preparation of the 20.2 per cent Ag alloy has been described2 Alloys containing 9.9 and 29.7 per cent Ag were prepared from 99.97 per cent A1 and 99.99+ per cent Ag using a similar procedure. Rolled

Jan 1, 1943

By R. F. Mehl, C. S. Barrett, A. H. Geisler

The mechanism by which the supersaturated solid solution of 20 per cent silver in aluminum decomposes has been reported in two previous publications.1'2 The analysis of streaks in Laue photograms showed that thin platelike particles (the so-called Guinier-Preston aggregates) form upon matrix planes in the early state of decomposition.1 The size of these particles is larger, the higher the temperature at which they form in accordance with nucleation theory;³ at all temperatures they increase in size with aging.2 When they have grown thick enough to afford three-dimensional diffraction, powder X-ray photograms show that the structure is not that of the equilibrium precipitate but that of a transition lattice . Sincc this is also platelike in shape and also lies upon the (III) matrix planes with the [IIO] direction in the matrix parallel to the [II.O] direction in it seems likely that the original platelet particles are merely in an early stage of formation, and thus finely dispersed. The atomic spacing and pattern on the (00.1) plane in y' are identical with those on the (111) matrix plane to which it lies parallel, and since comparable directions in the two planes are parallel, the lattices are thus coherent at the interface; an analogous coherence has been found in aluminum-copper alloys.3 The structures y' and y are both close-packed hexagonal and are oriented alike; they differ only in interatomic distances. Consequently, y' is regarded as y strained from its ordinary dimensions by the lattice coherency between it and the matrix.2 The lattice movements that transform the matrix to the equilibrium precipitate appear to be very simple.2 The mechanism of precipitation in this system is simpler and is more precisely known than that in any other system hence the system should be a profitable one for the difficult task of correlating property changes with structural alterations. The present report provides additional studies on powder X-ray diffraction from alloys of 10 per cent, 20 per cent and 30 per cent Ag, and presents the correlation of crystallographic changes with changes in microstructurc, Brinell hardness and electrical resistivity of the three alloys. Preparation of Samples and Experimental Procedure The preparation of the 20.2 per cent Ag alloy has been described2 Alloys containing 9.9 and 29.7 per cent Ag were prepared from 99.97 per cent A1 and 99.99+ per cent Ag using a similar procedure. Rolled

Jan 1, 1943

By M. F. Hawkes, R. F. Mehl

The rate of isothermal transformation of austenite to pearlite depends upon the rate of nucleation, N, and the rate of growth, G, of pearlite in austenite. Values of N are given in terms of the number of nuclei forming in unit time per unit grain boundary area of unreacted austenite; values of C are given in terms of the linear rate of radial growth. Variations in the isothermal rate with temperature result from variation of N and G with temperature. Depth of hardening of a steel depends fundamentally upon values of N and G; changes in carbon and alloy content effect changes in N and G and thus affect depth of hardening. Quantitative studies are available on N and G for the formation of pearlite in plain carbon eutectoid steels.3 Systematic studies in this field are necessary if the phenomena are to be well enough known to furnish a proper basis for full rationalization and for theory. Accordingly, the studies are continuing, both with respect to the effect of alloying elements upon the values of N and G in the formation of pearlite and to the values of N and G for the formation of ferrite in hypoeutectoid steels. The present paper is the first relating to the effect of alloying elements upon the values of N and G for pearlite. It offers special interest since cobalt is known to have an anomalous effect upon the rate of formation of pearlite, accelerating it, whereas all other elements retard it.4 It will be shown that cobalt increases both N and G for pearlite and that this effect is inherent in the Fe-Co-C system and not dependent upon factors relating to austenite heterogeneity nor upon any recognizable adventitious factor. Inasmuch as the rate of diffusion of carbon in austenite and the interlamellar spacing of pearlite are variables related to G and possibly to N (in a manner not yet certain), measurements of the effect of cobalt upon them are also reported here. In the course of the work, measurements were made on the effect of cobalt upon the marten-site temperature and these are given. COBALT AS AN ALLOYING ELEMENT IN STEEL Properties of Cobalt Cobalt resembles iron more closely than does any other chemical element. Its atomic number is 27, hence it is the element immediately following iron in the periodic system. The two differ in atomic structure only in the fact that cobalt contains seven electrons instead of six in the third, or transition, level of .the third shell. The next shell, which is the

Jan 1, 1948

By M. F. Hawkes, R. F. Mehl

The rate of isothermal transformation of austenite to pearlite depends upon the rate of nucleation, N, and the rate of growth, G, of pearlite in austenite. Values of N are given in terms of the number of nuclei forming in unit time per unit grain boundary area of unreacted austenite; values of C are given in terms of the linear rate of radial growth. Variations in the isothermal rate with temperature result from variation of N and G with temperature. Depth of hardening of a steel depends fundamentally upon values of N and G; changes in carbon and alloy content effect changes in N and G and thus affect depth of hardening. Quantitative studies are available on N and G for the formation of pearlite in plain carbon eutectoid steels.3 Systematic studies in this field are necessary if the phenomena are to be well enough known to furnish a proper basis for full rationalization and for theory. Accordingly, the studies are continuing, both with respect to the effect of alloying elements upon the values of N and G in the formation of pearlite and to the values of N and G for the formation of ferrite in hypoeutectoid steels. The present paper is the first relating to the effect of alloying elements upon the values of N and G for pearlite. It offers special interest since cobalt is known to have an anomalous effect upon the rate of formation of pearlite, accelerating it, whereas all other elements retard it.4 It will be shown that cobalt increases both N and G for pearlite and that this effect is inherent in the Fe-Co-C system and not dependent upon factors relating to austenite heterogeneity nor upon any recognizable adventitious factor. Inasmuch as the rate of diffusion of carbon in austenite and the interlamellar spacing of pearlite are variables related to G and possibly to N (in a manner not yet certain), measurements of the effect of cobalt upon them are also reported here. In the course of the work, measurements were made on the effect of cobalt upon the marten-site temperature and these are given. COBALT AS AN ALLOYING ELEMENT IN STEEL Properties of Cobalt Cobalt resembles iron more closely than does any other chemical element. Its atomic number is 27, hence it is the element immediately following iron in the periodic system. The two differ in atomic structure only in the fact that cobalt contains seven electrons instead of six in the third, or transition, level of .the third shell. The next shell, which is the

Jan 1, 1948

By R. D. Reiswig, J. M. Dickinson

The 0s-Ir equilibrium diagram was determined. The diagram is of the simple peritectic type, with a peritectic temperature of about 2660°C. The solid miscibility gap is narrower than previously reported, the nearly vertical solvus lines occurring at about 42 and 63 wt pct 0s. Widmanstaetten figures resembling martensite in some of the two-phase alloys account for a previously suspected transformation. In view of the long history of use of such alloys, it is surprising that the 0s-Ir equilibrium diagram has not yet been completely determined. The boundaries of the miscibility gap in this system have been estimated to be between 24 and 39 wt pct 0s and 79 and 100 wt pct OS. It has been predicted3 that alloys of ruthenium, osmium, and rhenium with the fcc platinum metals should solidify peritectically in their two-phase regions. The metallographic examination of natural osmiridium alloys has been said4 to indicate, by the presence of a martensitic structure, the existence of a transformation or a phase separation. On the basis of superconducting transition temperature measurements, a 70 at. pct 0s alloy was found5 to be two-phase prior to annealing. EXPERIMENTAL The elemental components used in this work were obtained from several suppliers as sponge. The purities, stated by each supplier to be in excess of 99.8 pct, were confirmed in each case by a semiquantitative spectrographic analysis. The pure components were premelted in an inert-gas tungsten arc furnace with the evolution of some fume, crushed, and pickled for use as alloy melting stock. Upon arc melting to the final compositions, no fume evolution was observed. Because weight losses of the order of 15 mg or less were observed in preparing alloy buttons weighing 15 g, "synthetic analyses" were used in determining the compositions of the alloys. Alloys used in this investigation contained 15, 20, 33, 40, 42, 43, 45, 50, 55, 60, 63, 64, 66, 70, 79, and 80 wt pct 0s. Heat treatments of the alloy buttons and pieces thereof were performed in a previously described6 tungsten tube furnace under pressures of 105 torr or less. It was found that homogenization treatments on the specimens were essential to remove the severe coring which was present, a 3-hr treatment above 2300°C removing all the metallographi-cally visible evidence of coring. Following the various heat treatments, the specimens were "vacuum-quenched" by melting a tungsten fuse wire and allowing them to drop onto a water-cooled copper plate. In a few instances, specimens were quenched into molten tin for comparison, but no difference was apparent. Vacuum-quenched specimens cooled from 2400°C to below a red heat in less than 60 sec. Typical equilibration times varied from 3 hr above 2300°C to 88 hr near 1100°C. The utilization of successively longer equilibration times and the approach to equilibrium from both directions indicated that the times used were sufficient to give a close approach to true equilibrium. Heat-treatment temperatures were measured with a calibrated Leeds and Northrup optical pyrometer through a 0.120-in. lateral sight hole in the tungsten heater tube. It was known that the assumption of black-body conditions inside the 1-in.-diam by 10-in.-long tube was unwarranted. The empirical results of work with thermocouples and melting

Jan 1, 1964

By W. F. Hosford

W. F. Hosford, (Massachusetts Institute of Technology) —evelopment of methods for predicting the plastic anisotropy of textured metals is an important step toward the improvement of properties by texture control. The engineering value of the present analysis is indicated by the good agreement between prediction and experimental findings. The purpose of this discussion, however, is to point out the weakness of the assumptions on which the analysis is based. The authors state that "the relative width and thickness strains are believed to depend only on the slip direction" and accordingly they disregard the orientation of the active slip plane. where E,, cW, and et are the tensile, width, and thickness strains, A,, A,, and At are the angles between the tensile axis x, the width direction w, and the thickness direction t, respectively; and G,, GW, and 4 are the angles between the slip-plane normal and x, w, and t, respectively. Therefore, the ratio of width to thickness strains resulting from slip on a single system would be These relations hold regardless of whether the slip plane is a crystallographic plane of the {110}, (1121, or (123) type or whether noncrystallographic slip occurs. If noncrystallographic slip occurs, it might be reasonable to assume that the effective slip plane is the most highly stressed plane containing the slip direction. In this case the slip direction, S.D., the slip plane normal, S.P.N., and the tensile axis, x, are coplanar (Fig. 19) and the angles are related by following expressions: where 9 is the angle between t and the projection of the slip direction in the w-t plane (igs. 1 and 2). The normal strains, Eq. [9], become

Jan 1, 1963

By John W. Brauns, David H. Orr

9.1-1. Description and Basic Function. A conventional railroad haulage system for a surface mine performs the function of transporting the ore and/or waste out of the pit to a crushing plant or dump. Trains consisting of a locomotive and gondola or side-dump care suitable for loading by power shovels are used. Gondola cars are restricted to hauling ore as a rotary dumper is required to unload them. Side-dump cars are more versa- tile, being satisfactory for carrying both ore and waste; but because of their additional dumping mechanism they are likely to be more expensive to maintain (12). Also they weigh up to two times as much as a flat-bottom car of the same capacity, and, therefore, fewer cars can be hauled per train. Standard-gauge track is most commonly used. Track systems incorporate temporary track of low quality in areas where moves are frequent and permanent track of high quality on main lines. Rail haulage enabled open pit mining to establish its reputation, for efficiency. Prior to World War II it was the principal type of transportation used in large pits. Other successful surface mine haulage methods have been developed since that time. These have served to supplement railroad transportation, permitting selection of a method that is best adapted to the peculiar requirements of each mining project. 9.1-2. General Applicability and Controlling Physical Conditions. For handling rough rock, the shovel-train system is unexcelled (6). The ideal application is in a physically large, but not excessively deep, open pit mine from which the coarsely blasted ore and waste must be transported distances in excess of three miles. If, in addition, a high production rate is to be maintained, railroad haulage will probably afford a lower cost per ton transported than will any other haulage method. Relatively gentle

Jan 1, 1968

Secondary minerals are the result of a process of concentration and enrichment and are commonly richer than the primary minerals of the same deposit. Secondary ores that contain abundant sulphides are commonly distinguishable in the field from primary ores; this distinction is more difficult with ores that carry a small quantity only of sulphide minerals, and microscopical examination of the ore in thin section is advisable in all doubtful cases. It is frequently difficult, and in some cases impossible, to determine the primary or secondary character of quartzose gold ores. The safest guide is the presence or absence of gaseous or fluid inclusions in association with the ore minerals; such primary association coupled with a lack of traces of oxidation is indicative of primary origin. Any ore that carries traces of oxidation, such as iron stains, endritic oxides of manganese, kaolin and so forth must be considered to have been acted upon by surface agencies, and secondary enrichment must be suspected. Further examination, however, may indicate that the contained value is primary, and, therefore, persistent in its vertical distribution. The manner of occurrence of secondary minerals is frequently characteristic; secondary minerals, being due to the action of surface waters, are commonly found in cracks through earlier minerals, or as crusts lining vugs or surrounding nucleal masses, or in general, in some distribution later than and not conforming to the arrangement of the primary minerals. Secondary sulphides are often found as soft black pulverulent masses or coatings, although they frequently occur well crystallized or as

Jan 1, 1932

By A. Hansson, G. E. Tardiff

WHILE it is generally known that cold-worked low-and medium-carbon steels exhibit substantial increases in tensile rupture ductility with increased deformation velocity172 (up to the von Karman limit), the observations for annealed or normalized steel do not permit simple generalization. Various investigators, for example, have reported both small increases1,3 and small decreases4,5' in rupture ductility with increased rates of deformation of normalized low-carbon steel. The purpose of the present note is to report on the dependence observed for strain-aged steel since preliminary work here indicates that strain aging results in the development of a consistent type of behavior that is understandable in the light of current knowledge of the lower yield dynamics for steels which deform inhomogeneously.6 This realization throws additional light on the influences responsible for the behavior of normalized steel. The results presented here were obtained for a 0.21 pct C, 0.012 pct N sheet (0.012 in. thick) steel which was received in a temper-rolled condition (approximately 3 pct reduction in thickness after process annealing); the as-received material had a spherodized structure with a ferritic grain size of ASTM 8. The material was selected for testing after 6 months room temperature aging, at which time the static lower yield and ultimate tensile strength measured 79 and 82 kpsi, respectively. ASTM 2-in. reduced section, 1/2-in.-wide specimens were used in smooth bar tests and 3/4-in.-wide specimens containing two 60-deg opposed edge notches (root radii = 0.002 in.) were used in notched bar tests. To facilitate strain measurement, the smooth bars were blued with ink and marked at 0.1-in. intervals along the gage length; the notched

Jan 1, 1968

By Stanley J. LeFond

Jan 1, 1985

By E. V. Potter, E. T. Hayes, H. C. Lukens

LARGE volumes of hydrogen are liberated at the cathode during electrolytic precipitation of manganese. Most of the gas escapes from the electrolyte, but a considerable amount may be entrapped in the manganese plate. Gas has been observed to escape from manganese melts in puffs that ignite at the surface. It has been reported that castings made with electrolytic manganese tend to rise in the molds; for this reason the manganese is often treated to remove hydrogen before the melt is made or it is premelted and used as cast manganese in making the alloys. The effect of hydrogen on the properties of manganese and its alloys is not known, but its effect on iron and iron alloys has been studied extensively.1.2 The close relationship between iron and man¬ganese suggests that hydrogen may have a similar effect on the properties of manganese and its alloys. The purpose of this investigation was to determine the amount of hydrogen or other gases in typical commercial electrolytic manganese sheet and to investigate ways of removing it most conveniently. APPARATUS The apparatus used is shown schematically in Fig. I. It consists of a chamber A in which the metal sample is placed, a vacuum or pressure gauge B, volume gauge C, vacuum pump D, hydrogen container E, auxiliary gas chamber F, and furnace G. The specimen chamber was made of two concentric quartz tubes. The outer one, a, connects with the closed gas system and is surrounded by the furnace G. A second tube, b, fits closely inside a and has a small end that projects into the space for the specimen and holds the thermocouple for measuring the temperature of the sample. The inner tube is joined to the outer tube outside the furnace, with rubber tape and sealing wax, to prevent gas leakage into or out of the closed system. This type of specimen container is loaded very easily and can be used several times if molten metal does not come in contact with the quartz. An ordinary mercury manometer B is used to measure the pressure in the system. The volume gauge C is essentially another mercury manometer, but the volume of the column is much larger than normal and the height of the mercury can be varied by raising or lowering the well H. In this apparatus the tube had a volume of I00 c.c. for a 60-cm. length. For larger changes in gas volumes; auxiliary containers F were used to increase the volume of the system by known amounts. Valves were arranged so that these auxiliary containers could be filled from the system and evacuated by the pump as often as necessary to accommodate the increases in the volume of gas in the system. The furnace G, used to heat the samples, was equipped with an automatic temperature controller, which maintained the desired temperatures within ± 5°C.

Jan 1, 1945

By Lloyd R. Jackson

IN the past 15 years there has been a great deal of interest in the fundamentals of plastic flow and rupture in metals and a number of papers have presented substantial advances toward a fundamental insight into the mechanism of these two processes. This paper is an appraisal of the present position with respect to the following: I. The relative importance of shear and hydrostatic stresses in controlling the plastic flow of metals. 2. The relative importance of shear and hydrostatic stresses in determining the amount of flow before rupture in metals. 3. The use of generalid strains in determining the effect of plastic flow on the work-hardening of metals. FUNDAMENTALS INVOLVED IN PLASTIC FLOW In attempting to devise a quantitative description of plastic flow in metals, it has been customary to assume that metals are homogeneous and isotropic. The fact that neither of these assumptions is correct complicates the experimental appraisal of relations developed. Never- the less, as will be shown, metals are close enough to being isotropic and homogeneous so that considerable insight into the mechanism of plastic flow can be obtained. The assumptions of .isotropy and homogeneity allow both the elastic and plastic behavior of metals to be described in terms of 12 variables-six being stress variables, and six being strain variables. ) From this viewpoint, it should be possible to represent the stress-strain relations in the plastic range for metals by a surface in 12 dimensions. One boundary of this surface would be the surface of transition from elastic to plastic behavior; the other would be the surface of rupture. These two surfaces would be joined by the surface of plastic flow. The shape of this composite surface would depend on the test temperature and speed of straining. This paper is concerned only with results obtained at constant temperature and constant speed of straining. While such a representation is conceivable, the demonstration of such a surface would present almost insurmountable obstacles. The most fruitful simplification of this concept has been the recognition that the behavior of metals in the plastic region is primarily a result of shear stresses and strains. Based on this viewpoint, it is possible to utilize stress and strain invariants that are independent of hydrostatic components to define an effective stress and strain. In effect, this amounts to reducing the task of representing stresses and strains in 12 dimensions to one of representation in two dimensions. The two invariants (one for stresses . and one for strains) that have received most attention are the following: [ ]

Jan 1, 1946

By R. W. Smith, K. H. Lee

Bond type simulated closed-circuit grindability tests and batch grindability tests were run on six pure minerals plus a limestone and a taconite ore. The simulated closed-circuit tests were made at five different meshes of grind. The minerals tested varied widely in hardness and breakage pattern. For all simulated closed-circuit tests both Bond and Holmes work indices were calculated and compared. The Bond grindability testlm6 has been widely and quite successfully used for predictions of such items as ball mill power requirements and as an aid in plant comminution equipment selection. The test is interesting in that although run as a series of batch tests it is over-all a simulated closed-circuit test with a circulating load of 250%. The simulated closed-circuit feature is probably the primary reason for the notable success of the test. Perhaps the major criticism that can be brought against the test is that it is time consuming, since it normally takes a trained technician about an eight-hour shift to perform it.4 For many purposes this is an acceptable time requirement. However, if a grindability test is to be used for mill control purposes it should be performed in as short a time interval as possible, preferably under an hour. Sometimes in industry strictly batch tumbling mill tests are run and an attempt is made to correlate such batch tests with plant mill operation. Then an attempt is made to adjust operating tumbling mills based on data from the tests. The tests, however, frequently do not correlate well with operating data. The purpose of the present work was to explore some of the similarities and differences between a simulated closed-circui.t grindability test and a batch type test. Six different minerals plus a limestone and a taconite ore were tested in the experimentation. These materials varied widely in hardness, breakage characteristics, etc. Standard dry Bond grindability tests were mn on all materials at 48, 65, 100, 150 and 200 meshes of grind. Feed material was prepared by stage cmshing to -6 mesh. Initial feed volume was 700 cc. Care was taken to maintain the standard constant weight (about 20125 gm) and number of balls (283) during all the tests. Rotational speed was 70 rpm. Table I lists the materials studied along with pertinent data regarding the materials and Bond grindabilities obtained. In addition, dry batch tests were performed on all materials. These tests were made in the same mill as was used for the Bond tests. An identical ball charge and an identical feed volume were used. After rotating the mill for a predetermined number of revolutions at the same rotational speed as for the Bond test, the material was removed and separated from the ball charge. A screen analysis was performed on this product. Knowing the original size distribution of the feed and the size distribution of the product, it was possible to calculate the grindability or grams of a particular mesh produced per revolution of the mill. After the screen analysis of the product, the various fractions were recombined and placed back in the mill for further grinding. After grinding for additional revolutions of the mill the procedure was repeated. Thus, it was possible to obtain grindabilities for each material at a number of meshes of grind and for various revolutions of the mill. Grind-abilities obtained from batch tests are shown in Table 11. COMPARISON OF BOND AND BATCH GRINDABILITIES Figs. 1A through 1D show relationships between batch and Bond grindabilities. Figs. lA, lB, 1C and 1D show respectively graphs indicating the relationships for batch tests run for 100, 300, 500 and 700 revolutions of the mill. For a particular material on each graph, points are shown for 200, 150, 100, 65, and 48 meshes of grind except for several missing coarse meshes of grind for galena which did not fall conveniently on the figures. For each material and graph the point nearest the origin is for 200 mesh, the next for 150 mesh, the next for 100 mesh, etc. Considering Fig. lA, it is interesting to note that all relationships shown are essentially straight lines. There are, however, considerable differences

Jan 1, 1969

By W. Philippoff

ALTHOUGH Gaudin1 and more recently Sutherland2 have calculated the probability of collision of a falling mineral particle with a rising bubble, there is no published information concerning the details of the mechanism of attachment of a collector-coated particle to a bubble. During the past year the writer has developed a theory for the mechanism of attachment, which has been substantiated experimentally. Funds for the investigation and for some of the equipment used have been supplied by the Mines Experiment Station of the University of Minnesota. Motion picture studies of the phenomena involved in the collision between mineral particles and bubbles, such as those of Spedden and Hannan," show that the contact can be completed within 0.3 millisec. Formulas developed for rigid bodies have hitherto been used' for the calculation of the motion of bubble and particle, but it is obvious that a bubble cannot be regarded as a rigid body. On the contrary, Spedden and Hannan's pictures show a great degree of deformation during the collision. The time of attachment was calculated as the time the particle drifted past the bubble. Time of Collision The theory presented in this paper enables calculation of the time of collision, using the concept that the bubble, or more generally, a liquid-air interface, acts as an elastic body. The elasticity, defined as the restoring force on a mechanical deformation, is caused by the surface tension and is the result of the principle of the minimum of free surface energy. It is well known that an elasticity together with a mass determines a frequency of vibration. The vibrations of jets and drops caused by the elasticity of the interface are known to comply exactly with the classical theory of capillarity.' However, the vibrations of isolated bubbles, as distinct from foams, have not been investigated previously. The following equation, presented elsewhere,' has been deduced for these frequencies: 3_____________________ fB = 9.20.vV.vn. (n-1).(n+2)/8 [I] in which fB is the frequency of a harmonic of the bubble in cycles per second, V the volume of the bubble in cc, n a number determining the order of the harmonic, and n = 2 the basic vibration. The first (basic) harmonic describes a change of the spherical bubble to an ellipsoidal bubble. The higher harmonics are more complicated, for the circumference of the bubble is divided approximately into as many parts as the order of the harmonic. As an example, Spedden and Hannan's published motion picture of a vibrating bubble corresponds to the sixth harmonic. Eq 1 shows that only the first and third harmonics are simple multiples (1 and 3), all the others being irrational fractions of the basic frequency. This means that the shape of the vibration can change with time and is in general unsym-metric in respect to the time axis. Such conditions prevail when there is a distributed elasticity or mass, as in the case of vibrating membranes or rods. The constant 9.20 is valid for water at room temperature, but a general solution involving the physical constants of the liquid has not been found. The case of the floating particle is much easier to treat than that of the bubble. It can be assumed that the elasticity is caused exclusively by the interface and that the mass is concentrated in the particle together with some adhering water. The following expression for the frequency of a system of one degree of freedom can be applied: fP = -1/2-vE/m [2] Here fp is the frequency of the particle vibration in cycles per second, E the elasticity in dynes per cm, and m the mass in grams. The classical theory of impact phenomena gives the time of collision during the striking of a spring (in this case the surface of the bubble) by a mass, as: t, = 2/f = nvm/E [3] It is now possible to develop an expression for the elasticity of a floating cylindrical particle. The force equilibrium of a cylinder floating end on at the air-liquid interface is given by the well-known equation (Poisson7 1831) P = /4 D2-pL-g-h + D.y sin a [4] which accounts for the buoyancy and the action of the surface tension where P is the force acting on the particle in dynes (weight-buoyancy), D the diameter of the cylinder in cm, PL the density of the liquid in grams per cc, g the acceleration of gravity = 981 cm per sec2, h the depression of the cylinder below the surface of the liquid in cm, y the surface tension in dynes per cm and a the supporting angles or the one required to insure equilibrium, a being smaller than the contact angle 0. Although demonstrated by Poisson, it has not

Jan 1, 1953

By W. Philippoff

ALTHOUGH Gaudin1 and more recently Sutherland2 have calculated the probability of collision of a falling mineral particle with a rising bubble, there is no published information concerning the details of the mechanism of attachment of a collector-coated particle to a bubble. During the past year the writer has developed a theory for the mechanism of attachment, which has been substantiated experimentally. Funds for the investigation and for some of the equipment used have been supplied by the Mines Experiment Station of the University of Minnesota. Motion picture studies of the phenomena involved in the collision between mineral particles and bubbles, such as those of Spedden and Hannan," show that the contact can be completed within 0.3 millisec. Formulas developed for rigid bodies have hitherto been used' for the calculation of the motion of bubble and particle, but it is obvious that a bubble cannot be regarded as a rigid body. On the contrary, Spedden and Hannan's pictures show a great degree of deformation during the collision. The time of attachment was calculated as the time the particle drifted past the bubble. Time of Collision The theory presented in this paper enables calculation of the time of collision, using the concept that the bubble, or more generally, a liquid-air interface, acts as an elastic body. The elasticity, defined as the restoring force on a mechanical deformation, is caused by the surface tension and is the result of the principle of the minimum of free surface energy. It is well known that an elasticity together with a mass determines a frequency of vibration. The vibrations of jets and drops caused by the elasticity of the interface are known to comply exactly with the classical theory of capillarity.' However, the vibrations of isolated bubbles, as distinct from foams, have not been investigated previously. The following equation, presented elsewhere,' has been deduced for these frequencies: 3_____________________ fB = 9.20.vV.vn. (n-1).(n+2)/8 [I] in which fB is the frequency of a harmonic of the bubble in cycles per second, V the volume of the bubble in cc, n a number determining the order of the harmonic, and n = 2 the basic vibration. The first (basic) harmonic describes a change of the spherical bubble to an ellipsoidal bubble. The higher harmonics are more complicated, for the circumference of the bubble is divided approximately into as many parts as the order of the harmonic. As an example, Spedden and Hannan's published motion picture of a vibrating bubble corresponds to the sixth harmonic. Eq 1 shows that only the first and third harmonics are simple multiples (1 and 3), all the others being irrational fractions of the basic frequency. This means that the shape of the vibration can change with time and is in general unsym-metric in respect to the time axis. Such conditions prevail when there is a distributed elasticity or mass, as in the case of vibrating membranes or rods. The constant 9.20 is valid for water at room temperature, but a general solution involving the physical constants of the liquid has not been found. The case of the floating particle is much easier to treat than that of the bubble. It can be assumed that the elasticity is caused exclusively by the interface and that the mass is concentrated in the particle together with some adhering water. The following expression for the frequency of a system of one degree of freedom can be applied: fP = -1/2-vE/m [2] Here fp is the frequency of the particle vibration in cycles per second, E the elasticity in dynes per cm, and m the mass in grams. The classical theory of impact phenomena gives the time of collision during the striking of a spring (in this case the surface of the bubble) by a mass, as: t, = 2/f = nvm/E [3] It is now possible to develop an expression for the elasticity of a floating cylindrical particle. The force equilibrium of a cylinder floating end on at the air-liquid interface is given by the well-known equation (Poisson7 1831) P = /4 D2-pL-g-h + D.y sin a [4] which accounts for the buoyancy and the action of the surface tension where P is the force acting on the particle in dynes (weight-buoyancy), D the diameter of the cylinder in cm, PL the density of the liquid in grams per cc, g the acceleration of gravity = 981 cm per sec2, h the depression of the cylinder below the surface of the liquid in cm, y the surface tension in dynes per cm and a the supporting angles or the one required to insure equilibrium, a being smaller than the contact angle 0. Although demonstrated by Poisson, it has not

Jan 1, 1953

By Campbell Hunter

PETROLEUM is found in almost every province in Persia. On the northern frontier, along the southern shore of the Caspian Sea, it is found near Anzelli and Shakhtesar and gas at Khoremabad. Oil is also found at Gumush Tepe, northwest of Astrabad, on the southeastern shore of the Caspian Sea. Further inland, to the south of Astrabad, oil is found at Schahkuh-i-balae, also on the margin of the Khorasan desert at Semnan, 115 mi. (35 m.) east of Teheran. Along the western frontier, from northwest to southeast, oil is encountered at Ouschachi, north of Lake Urumieh; in the province of Azerbaijan, at Zohab, Khanikin, and other places in the district of Kormanshah in the province, of Ardelan. Further south, in the province of Luristan, oil is found east of Mendeli and in the Pusht-i-kuh districts. Considerable quantities of oil are also obtained from Schuster, Maidan-i-Naphtun, Ram-Hormuz, Beheban, which are almost on a straight line, running northwest and southeast along the foothills of the Bakhtiari Mountains. At Ahwaz, in the province of Arabistan, oil has been found along another range of hills whose axis also lies in a northwest-southeast direction. In the Fars province, boring for oil has taken place at Daliki, and indications of oil are found at Kheri, Fasa, Darab, and other places. A gas show is also recorded at Kuhi-Sung-Atush in this province, 30 mi. (48 km.) east of Darab. In the south of Persia, oil is encountered on. Qishm Island, also at Ahmedi and other places north of Bunder Abbas. On the southeastern frontier, oil is found on the Sarhad range of hills. Thus, oil indications have been noted over a distance of approximately 1100 mi. along the western and 700 mi. along the northern frontiers of Persia.

Jan 2, 1920

By W. Bonfield

The temperature dependence of the microscopic yield stress (the stress to produce a plastic strain of 2 x 10-6 in. per in.) and the stress-plastic strain curve of polycrystalline Cu 1.9 wt pct Be have been measured for the solution treated condition, an intermediate condition containing G.P. zones and ?' precipitate and the overaged ? precipitate condition, in the range from -58° to 200° C. A transition in micro -yield behavior and a large temperature dependence were noted for the intermediate condition, which are interpreted in terms of the interaction of glide dislocations with two differently sized zones. In comparison the microscopic yield stresses of the solution treated and overaged conditions were less sensitive to temperature variations and are satisfied by the Mott-Nabarro and dislocation bowing theories, respectively. A determination of the temperature dependence of the yield stress of a precipitation hardening alloy has provided a powerful tool for evaluation of the operative deformation mechanism. There is a marked contrast between the effect of temperature on the yield behavior of a metal containing coherent zones or intermediate precipitates, which can be "cut through" by mobile dislocations, and a metal containing a dispersion of noncoherent particles, through which dislocation "bowing out" is the dominant role of deformation.' These studies have in general been confined to single crystals, as it was considered that similar experiments on polycrystalline material did not produce good data because of the lack of sensitivity with which the yield stress could be determined. However, this objection has been removed by the introduction of mi-crostrain techniques, with which the yield stress in polycrystalline materials can be measured to a strain sensitivity of 10-6. Such measurements have not only shown that the deformation of polycrystalline precipitation hardening alloys can be examined with the same detail as single crystals, but also that some unexpected results are obtained.' In this paper the results obtained from a study of the temperature dependence of the microscopic yield stress (the stress to produce a plastic strain of 2 x 10-6 in. per in.) and the stress-plastic strain curve of a polycrystalline Cu 1.9 wt pct Be precipitation hardening alloy (Berylco 25) are discussed. The temperature dependence of the alloy was measured for three different conditions: 1) The solution treated condition (a supersaturated solid solution of a containing ~12 at. pct Be3) which is obtained by water quenching the alloy from 800° C. 2) The condition of y' intermediate precipitate, to- gether with some G.P. zones,' which is produced after an aging treatment of 2 hr at 315°C from the solution treated condition. (The alloy was cold rolled to 40 pct reduction prior to aging to minimize grain boundary precipitation effects.)4 3) The condition with equilibrium ? precipitate structure2 which is developed after an aging treatment of 24 hr at 425° C. EXPERIMENTAL PROCEDURE Tensile specimens of gage length 1 in. and with rectangular cross section of 0.18 by 0.06 in. were prepared from the solution treated, cold rolled alloy and were either resolution treated for 1 hr at 800°C, followed by water quenching, or aged for 2 hr at 315°C and 24 hr at 425° C to produce the desired precipitate structures. The microstrain characteristics of the aged specimens were determined at temperatures from —58" to 200° C and those of the solution treated specimens from -58° to 30° C. Each temperature was controlled to ± 0.2°C, which was a level of stability sufficient to eliminate thermal expansion effects from the measurements (~1.2°C temperature increase produced an extension of 2 x 10-6 in.). The microplastic behavior of the specimens in the temperature range below 82" C was measured with a standard Tuckerman strain gage,5 while at temperatures above 82°C a modified Tuckerman gage with a reduced strain sensitivity (4 x10-6 in. per- in.) was used. A load-unload technique was used to establish values of the microscopic yield stress. The specimen was strained at a constant cross head speed of 2 x 10-2 in. per min to a given stress level, at which the total strain was measured. Then the specimen was immediately unloaded at the same rate and any residual plastic strain determined. This procedure was repeated for an increasing series of stress levels until the microscopic yield stress was established by a direct measure of the stress to produce a residual plastic strain of 2 x 10-6 in. per in. (It should be noted that, as reversible dislocation motion occurs at stresses less than the microscopic yield stress,2 the plastic strain rate at this level was not constant.) In an ideal test, the microscopic yield stress would be determined from a continuous stress-strain measurement, rather than from a load-unload sequence, in order to eliminate mechanical recovery effects.6 However, it was found experimentally that mechanical recovery was negligible in Cu 1.9 wt pct Be at small plastic strains for all the temperatures investigated, as the microscopic yield stress was independent of the number of load-unload cycles employed (i.e., the values measured for specimens subjected to different numbers of cycles was within the experimental scatter determined for specimens tested in an identical manner). Therefore, it is reasonable to consider the microscopic yield stress determined in the load-unload

Jan 1, 1969