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By E. S. Lenker, H. J. Van Hook

A complete swies of solid solutions has been found between the compounds InAs and GaAs Below the solidus. The melting relations determined by differential thermal and static quenching technzques, indicate that the liquidus and solidus are of the simple ascendant type. No evidence of unmixing of the homogeneous solid solutions was observed at sub-solidus temperatures. A quenching method for locnting the solidus, using X-ray diffraction techniques, is described in some detail. THE semi-conducting properties of compounds formed by various combinations of the less familiar intermetallic elements have created an increasing interest in the phase relations in systems involving these elements. The complete phase diagram of such a system, containing a compound of special interest, will often provide information as to the most favorable conditions for preparing large crystals from a melt, as well as specifying the temperature stability of the crystalline phase or phases produced. In view of the current importance of gallium arsenide in semiconductor research and of the structural similarities of indium arsenide, it was thought that the complete phase relations of the system InAs-GaAs should be investigated. Previous work on the system Ga-As by Koster and homa' and on In-As by Liu and peretti2 has shown that GaAs and InAs are the only compounds formed, having melting points of 1238° and 942°C, respectively. In a survey of various systems of elements from groups IIIb and Vb, spengler3 also re- Mines Ltd., for his continuing encouragement and for permission to publish the results of this work. ports complete miscibility of gallium, indium, and arsenic in the liquid state. In a partial investigation of the InAs-GaAs system, Woolley and smith4 found extensive solid solution throughout the alloy range. Their subsolidus work, however, was inconclusive at temperatures below 900" and they did not determine the liquidus relationships of the system. PRESENT INVESTIGATION The starting materials used were high-purity gallium, indium, and arsenic. The method of preparation of the compounds follows that devised by M. Shafer and is described in detail in a paper by Shafer and eiser.' The end-member compounds, GaAs and InAs, were first prepared by reaction of the elements in the desired proportions in evacuated silica glass vials. The vials were rotated in a furnace as the temperature was slowly raised to insure rapid and complete reaction at low temperatures. Thereafter the starting materials, GaAs and InAs, were melted by heating the sealed vials to 1300" and 1000°C, respectively; the vials in this case were filled as completely as possible to keep the vapor space small and thereby minimize the loss of arsenic to the vapor phase. Ten binary mixtures of these compounds were then prepared by reaction of the powdered material in the solid state just below solidus temperatures. At 900C preparations containing 3, 5, 8, 10, 25, 35, 50, and 75 mol pct GaAs were found to be single homogeneous phases, indicating a complete series of solid solutions between GaAs and InAs. Moreover, it was discovered that the homogeneity of any crystalline binary preparation could be very accurately checked by X-ray diffraction because of the large differences in unit cell dimensions of the end membe5s. The cubic lattice constant ofor GaAs is a = 5.652A and that of InAs is a = 6.061A

Jan 1, 1963

By F. L. Vogel

Densities and distributions of dislocations in plastically bent germanium crystals before and after annealing were studied. In the bent and annealed crystals, the theoretical relationship between radius of curvature and density of dislocations is confirmed. Before annealing, however, more dislocations are present than required, and these are distributed with a minimum at the neutral axis and maxima at the top and bottom surfaces. On annealing, three significant changes occur in the bent bars: 1) the average dislocation density is reduced, presumably by the annihilation of opposite signs; 2) dislocations migrate from the high density outside regions toward the low density neutral axis, thus producing the equilibrium distribution of dislocations; and 3) a polygonized structure is formed by movement of the dislocations into walls normal to the slip plane. ETCHING to reveal edge dislocations'-" offers a unique opportunity to study, in germanium single crystals, the functioning of dislocations in plastic deformation. Plastic bending is ideally suited to this purpose because it produces a static array of edge dislocations in an ideally plastic material which is predictable from the theory." " Nye" has given a rather complete treatment of the geometry of bent crystals in terms of dislocations. The density of excess dislocations is related to the radius of curvature of the slip plane of a bent crystal, as shown in Fig. 1, assuming the absence of macroscopic elastic stresses. The curvature of any plane in the zone of the bend axis is found by resolving the Burgers vector into that plane, so that for the usual case where the slip plane is inclined at some angle 0 to the neutral plane of the bar, the equation applies, with T equal to the radius of curvature of the neutral plane of the bent crystal. Two predictions which are relevant to this study may be made from Eq. 1. First, in crystals bent to various radii, the average excess dislocation density is inversely proportional to the radius and second, for a bar bent to a radius which is large compared with its thick-

Jan 1, 1957

THE strike and dip of bedded formations is readily determined by trigonometry if some marker horizon is intercepted by three holes. If no marker can be identified and if orientation of the cores is unknown, as is the usual case in diamond drilling, the determination still can be made readily by the use of a stereonet if the bearing, inclination, and angle the bedding makes with the core axis is known for each of the three holes. Such determination, valid only if the bedding orientation is fairly uniform within the area drilled, is considered simpler, faster, and more effective than any solution previously published. Method: The steps of this procedure are as follows: 1) Mark north and south poles on tracing paper placed over stereonet.* 2) Plot the three diamond drillholes by drawing a line from the center of the net in the direction of the bearings of the holes. Then rotate the paper until each line, in turn, lies along the equatorial diameter of the net; the end of each line is determined by the inclination of the hole measured in degrees from the periphery toward the center of the net along the equatorial diameter. Fig. 1 illustrates a vertical hole and a hole inclined at —23" due east. 3) With the projection of one diamond drillhole along the equatorial diameter, measure from the end of the hole, to the east and then to the west, a dis- tance in degrees equal to the core axis-bedding plane angle. Find two more points by measuring the same number of degrees north then south along the great circle from the end of the drillhole. Draw a circle, the center somewhere along the direction of the projection of the drillhole, to connect these four points. Draw circles for the other two diamond drillholes in the same manner. See Fig. 1 for two circles representing core axis-bedding plane angles of 35" and 23" and Fig. 2 for these plus a third circle representing a core axis-bedding angle of 42" in a hole bearing S 50 W inclined at —52". 4) Rotate the tracing paper until one great circle on the stereonet is found which is tangent to all three of the circles or parts of circles derived in the previous step. Trace this great circle on the paper and join its ends with a straight line passing through the center of the net. The bearing of this line is the strike of the bedding. In Fig. 2 the bedding strikes N 77 E. 5) Rotate the paper until the strike is along the polar diameter of the net. The dip of the bedding is determined by measuring along the equatorial diameter the distance in degrees from the periphery of the net to the great circle derived in step 4. The dip of plane illustrated in Fig. 2 is 55" to the N 13 W. Fig. 3 indicates the solution for three holes whose bearings and inclinations are N 20 W and —43o, due E and —30°, S 30 W and 23" with core axis-bedding angles of 30°, 50°, and 43" respectively. The bedding (stippled area) strikes N 44 W and dips 70" to the southwest.

Jan 1, 1956

By Warren L. Howes

THIS paper deals primarily with an investigation of ball-mill liners that was conducted by the writer over a period of six years at the Mammoth mill in Arizona. The investigation covered a wide variety of designs of liners in both primary and secondary mills. The comparative results are presented herein as evidence of the extent to which design influenced liner life and mill performance. FLOWSHEET The feed to the Mammoth mill was a complex, oxidized, gold-silver-lead-molybdenum-vanadium ore with rhyolite, andesite and granite as gangue. Although the ore varied between stopes, the long-term average in character remained substantially the same. Capacity was 560 tons per day. [t] Fig. I shows an abbreviated flowsheet of the grinding section of the mill. PRIMARY BALL MILL A 3-ft. Symons shorthead crusher in open circuit crushed to ¼ -in. maximum one-way dimension the feed to the 5 by 10-ft. Marcy open-end ball mill. The mill rotated at 80 per cent of critical speed and was charged with 4-in. forged steel balls. Cast balls were used for a few months at a slight reduction in ball cost, but breakage of the large balls reduced the ball size and grinding capacity to a point where the over-all costs suffered out of proportion to the saving in cost of balls. For metallurgical reasons, the finished product of the mill was maintained at minus 8-mesh. This coarse product, plus the very abrasive ore and the absence of fines in the feed, caused abnormally rapid wear of the mill liners. The resultant high operating cost led to the most extensive phase of the investigation. [ ] In order to clarify the difference in design of the various types of liners used, Fig. 2 shows the profiles with the arrangement in which grinding balls of maximum size would lie on the face of the liners. The results with each type of shell liner in the primary mill are shown in Table I.

Jan 1, 1943

By Donald L. Katz

Charts for predicting the pressure to which natural gases may be expanded without hydrate formation have been prepared for gases of even gravity. Pressure-temperature curves for hydrate formation were established for gases having gravities from 0.6 to 1.0. These curves and the thermal behavior of the gases during free and adiabatic expansion were used to prepare the charts for estimating the permissible expansion of natural gases without hydrate formation. The problem was solved by L. F, Albright, W. T. Boyd, J. J. McKetta, G. Martin, P. Poettman. and A. P. Snyder.-, who are first-year graduate students in the department of Chemical and Metallurgical Engineering at the University of Michigan. This paper is essentially a summary of their results. Prediction of Conditions for Hydrate Formation in Natural Gases Each natural gas under a given pressure will form solid hydrates at a corresponding temperature provided sufficient water is present.1-5 The temperature of natural gases below about 5000 lb. per sq. in. decreases when the gases are expanded freely.6,7 This decrease in temperature may cause the expanding gas to enter the region of temperature and pressure at which hydrates will form' The final pressure to which a natural gas may 'be expanded without hydrate formation depends upon the initial temperature and pressure and the gas composition. This paper presents charts that give the final pressures to which gases of gravity 0.6 to 1.0 at given initial temperatures and pressures may be expanded without formation of hydrate. Hydrate-forming Temperatures and Pressures as a Function OF Gas -Gravity Experimental data of pressure versus temperature at which solid hydrate will form, provided sufficient water is present, are available for a series of gases.2,3 A method of predicting the pressure-temperature curve for a given gas composition has been reported.' The solution to this problem requires that these curves relating pressure to temperature be established for gases of given gravities.

Jan 1, 1945

By Frank Firmstone

The piece of spiegeleisen, the analysis of which is given below, was collected by my father, together with various other specimens, while he was manager of the Glendon Iron Works. It bears a label stating that it was taken from a building in the town of Newton, New Jersey, erected in 1765 and torn down in 1863. The building belonged, in 1863, to T. N. McCarter, Esq. The label is in the hand-writing of Mr. McCarter, now deceased, well known as a member of the bar of northern New Jersey. The specimen was, no doubt, given by him to illy father, with whom he had business in 1863 and 1864. It is now in the metallurgical collection of Lehigh University. I know nothing further as to the history of the specimen, but there is no reasonable doubt that it was made from ore from the Andover mine, at the old charcoal-furnace which once stood at the village of Andover, about 5 miles south of Newton, on the Sussex railroad. It was probably broken from one end of a bearer over an open fire-place, is about 6 in. wide by 2.5 in. thick, has been cast in open sand and shows the charcteristic fracture of spiegeleisen rather low in manganese, as shown in Table I. One edge has been white-washed. Table I.—Analyses of Spiegeleisen Made from Andover Ore. 1. 11. Per Cent. Per Cent Graphite,.0.101 0.005 Combined carbon,... 4.529 4.212 Silicon,..0.037 0.294 Manganese,..... 5.760 3.750 Sulphur,.0.003 0.031 Phosphorus,.0.060 0.072 Iron,........ 89.440 91.570 99.920 99.934 I. Spiegeleisen made at Andorer (charcoal) furnace not later than 1765. IT. Spiegeleisen made from Andover ore at Phillipsburg, N. J., anthracire-furnwe, 185-. Both analyses were made by Dr. P. W. Shimer.

Jan 1, 1907

By G. C. Wallick, W. R. Foster, J. G. Savins

The theory of the differentiation method for the Couette flow experiment is reviewed. Particular attention is given to the requirements on data analyses in the case of the class of non-Newtonian materials described as viscoplastics, i.e., materials characterized by a yield point or yield stress. Here changes in boundary conditions arise when the shearing stress attains a critical value with the result that the form of the basic integral equation for Couette flow is determined by the flow conditions existing during the measurement. INTRODUCTION In the preceding papers in this series,l,2 the salient features of the differentiation method of rheological analysis in Poiseuille-type flow were discussed. It was shown that a dual differentiation-integration method analysis of the Poiseuille flow of idealized generalized Newtonian and visco-plastic models could be used to develop a spectrum of highly sensitive response patterns in terms of certain characteristic derivative functions. These functions were shown to optimize the selection of the most appropriate functional relationship between f(p) and p from the Poiseuille flow experiment. The present paper reviews the theory of the differentiation method as applied to the equally important Couette flow experiment. We will also discuss the range of variables over which the basic integral equation for Couette flow is applicable when the non-Newtonian material is of the viscoplastic type, i.e., characterized by a yield point or yield stress. THEORY Having described the application of the differentiation method to Poiseuille-type flow in the preceding papers, we now proceed to the case where the test liquid is confined to the annular space between coaxial cylinders of length L, one of which is in motion, i.e., Couette flow, formulating the basic integral equation after the method of Mooney.3 The observed kinematical and dynamical quantities are the angular velocity O and the torque T. Here, the one nonvanishing component of the shear-rate tensor is

ST. LOUIS SECTION EUGENE MCAULIFFE, Chairman, J. N. HOUSER, Vice-chairman, VICTOR RAKOWSKY, Vice-chairman., H. G. WASHBURN, Vice-Chairman, W. E. MCCOURT, Sec'y-Treas., Washington University, St. Louis, Mo. H. A. BUEHLER, G. H. COx, L. C. GLENN, E. A. HOLBROOK, W. E. NEWNAM. The eighth meeting of the St. Louis Section was held at the Mercantile Club on Apr. 15, 1918. ,, The meeting was preceded by a dinner, at which 51 members and guests were present, and at which T. T. Brewster presided. Bradley Stoughton, Secretary of the Institute, and Major J. R. Fordyce were the guests of the Section. A number' of the Joplin members were obliged to leave before the close of the dinner, in order to make train connections. The Vice, chairman, therefore, called on Victor Rakowsky, who brought a message of greeting from the Southwest district. At the close of the dinner, the Vice-chairman, T. T. Brewster, introduced Bradley Stoughton, who talked on various subjects of interest to members of the Institute, especially the affiliation of local sections with those of other national engineering societies. The minutes of previous meetings were read and approved-and-the Treasurer presented a financial report covering the previous year.

Jan 6, 1918

LIMA (PERU) Established May, 1952 (Official year begins in June) Luncheon meetings third Wednesday monthly Ernesto A Baertl, Chairman Arthur C Hall, Chairman-elect Raymond R Beard, 1st Vice-Chairman Edgardo Portaro, 2nd Vice-Chairman Walter E King, Secretary-Treasurer Av Santa Cruz 331 San Isidro, Lima, Peru Alberto E Brazzini, Asst Treasurer Daniel M Rodriguez, Asst Secretary Robert P Koenig, '56, Section Delegate Arthur C Hall, Alternate Section Delegate

Jan 1, 1955

LIST OF THE MEETINGS OF THE INSTITUTE AND THEIR LOCALITIES FROM ITS ORGANIZATION Transactions Number Place Date Vol Page 1 Wilkes-Barre, Pa May, ?71 1 3 2 Bethlehem, Pa August, ?71 1 10 3 Troy, N Y November, ?71 1 13 4 Philadelphia, Pa February, ?72 1 17 5 New York, N Y May, ?72 1 20 6 Pittsburg, Pa October, ?72 1 25 7 Boston, Mass February, ?73 1 28 8 Philadelphia, Pa May, ?73 2 3 9 Easton, Pa October, ?73 2 7 10 New York, N Y February, ?74 211 11 St Louis, Mo May, ?74 3 3 12 Hazleton, Pa October, ?74 3 8 13 New Haven, Conn February, ?75 3 15 14 Dover, N J May, ?75 4 3 15 Cleveland, 0 October, ?75 4 9 16 Washington, D C February, ?76 4 18 17 Philadelphia, Pat June, ?76, 5 3 18 Philadelphia, Pa October, ?76 5 9 19 New York, N Y February, ?77 5 27 20 Wilkes-Barre, Pa May, ?77 6 3 21 Amenia, N Y October, ?77 6 10 22 Philadelphia, Pa. May, ?78 6 18 23 Chattanooga, Tenn May ?78 7 3 24 Lake George, N.Y. Oct. ?78 7 25 Baltimore, Md. Feb. ?79 7 26 Pittsburgh Pa, May ?79 8 27 Montreal, Canada Sept, ?79 8 28 New York, N.Y. Feb, ?80 9 29 Lake Superior, Mich Aug, ?80 9 30 Philadelphia, Pa. Feb, 81 9

Jan 1, 1923

By C. N. Garman

Homestake-New Mexico Partners operate a 750-tpd carbonate leach uranium concentrate mill near Grants, N.M. The highly mineralized water available as process water leaves much to be desired. The 628 ppm as CaCO 3 makes the use of raw water very troublesome in pipes and on filter cloths. However, the residual sodium carbonate in the final filter cake going to tails makes an ideal softening agent. To take advantage of this fact, all makeup water used in the mill is first used as tailing slurry dilution water and comes to the mill from the tailings pond. The 5-acre tailings pond serves as a thickener and 100 to 150 gpm of nearly clear solution is decanted to a pump to be returned to the mill. Since this tailings water has small quantities of uranium in the solution an ion exchange scavenger unit was installed to remove as much uranium as possible. The ion exchange raffinate is then used as final filter wash ahead of the tailings slurrying step. In spite of the large settling area this return water is not clean enough for ion exchange feed. The solids present are very fine and composed of approximately 15 pct (by weight) burnable carbonaceous material common to the sandstone uranium ores in the area, 40 pct SiOz plus 45 pct CaC03. Laboratory work showed that this material responds very well to flotation. Before deciding to use flotation, various clarifying systems such as pressure leaf filters, sand filters, and continuous vacuum pre-coat filters, were considered. Each of these could have solved the problem but with much more operating labor, more reagents and greater installation costs than the flotation step. About 100 to 150 gpm of fouled water is fed to two 66-in. Fagergren cells, in series. Reagents used at the beginning were Arquad 2HT75 and Arquad C50, at the rate of about 1% lb per 8-hr shift, or about 0.0053 lb each per ton of ore. This did not completely remove the solids but does an acceptable job. Approximately 75 pct of the slimes are a size that can be caught on a 41-Whatman Paper are removed. Removal of these slimes also allows much better settling of the coarse nonfloatable material. Advantage is also taken of this fact in a small settling tank ahead of precipitation. Removal of this amount of the slimes makes the ion t:xchange feasible. PREGNANT SOLUTION CIRCUIT The carbonate? leach-caustic precipitation method of uranium concertration does not provide for any process purification step ahead of precipitation. Therefore, any fine solids getting into the pregnant solution through the filter cloth show up in the final concentrate. This, of course, lowers the grade, and, at times, the slimy nature of these very fine solids rendered final filtration of the concentrate difficult if not impossible At Homestake-New Mexico Partners a 75-ft thickener was available for gravity clarification of 100 to 120 gpm of this pregnant solution. However this did not sufficiently remove the slimes. Laboratory investigation of the whole range of flocculants that were suggested by literature, salesmen, and friends failed to turn up anything of consequence. A continuous vacuum pre-coat filter would do the job and was investigated. The capital cost and the operating labor and materials made this a last chance choice. Following work done in the metallurgical laboratory on the tailings return water, it was found that some changes in the reagent strengths and combinations made a very definite decrease in the solids in the pregnant solution. Concentrate grade improved about 5 pct anti the final product after drying had an appreciably greater bulk density. Compared to a cost of about 2.2e per ton for pre-coat filter opelation for cleaning just one circuit, flotation costs less than 1.0 per ton of ore for cleaning two circuits. While a pre-coat filter would do a more thorough job, the flotation does all that is required for either circuit. Gravity causes the froth produced to run back into the leach circuit. This does not appear to result in a build-up of objectionable slime. No extra manpower is required; the operators in the separate areas can observe the operation of the cells and mix the small quantities of reagents as needed. Normally the 66-in. Fagergren cell requires 15 hp per cell, but this very dilute slurry needs only 10 hp for both cells. Originally, a combination of the two Arquads mentioned previously served as frothers and promoters. As further testing

Jan 1, 1962

By R. B. Russell

SINCE Cahn's¹ paper on plastic deformation in orthorhombic (a) uranium an increasing amount of work is to be expected on orientation of single crystals and preferred orientation in this metal. It would seem desirable to have a table of angles between crystallographic planes (hkl) and (HKL). Table I shows these angles as distances from (loo), (0l0), and (00l), and in Fig. 1 as a standard (001) stereographic projection. Only the locations of poles whose diffracting planes produce Cu Ka reflections up to 0 = 70.5" (i.e. (154) ) are given." The relative reflection of these planes are represented by circle diameters. It is hoped that the projection may also be used in making inverse pole figures, originally suggested by Harris.' Values of 4 were calculated from the expression hH kK 1L cos = hH/a²+kK/b²-lL/c² where the lattice parameters a, b, c, and the spac-ings d were taken from the work of Jacob and Warren." (They report a = 2.852 kX (2.8578 A), b = 5.865 kX (5.8768 A), c = 4.945 kX (4.9550 A).) The parameters reported by them may be subject to revision,4,5 accordingly the interplanar angles are listed to only the nearest tenth of a degree. References ¹R. W. Cahn: Acta Metallurgica (1953) 1, p. 49. ¹G. B. Harris: Philosophical Magazine (1952) 43, p. 113. 3C. W. Jacob and B. E. Warren: Journal ACS (1937) 59, p. 2588. 4J. S. Lukesh: Acta Crystallographica (1949) 2, p. 420. ³ J. Thewlis: Nature (1951) 168, p. 198.

Jan 1, 1954

By L. R. Jackson

THE idea presented in this paper is that the capacity of a metal for plastic flow before rupture is dependent on the type of stress system applied-that is, the absolute magnitude of the stresses are unimportant provided that they are large enough to produce continued flow; the important factor is the "type" of stress system. A term called the "triaxial stress ratio" is defined in such a way that this ratio can be used to characterize the "type" of system. The paper is divided into three sections. The first section discusses means of representing the effect of complex stresses on plastic flow and defines "triaxial stress ratio"; the second section discusses various methods of' obtaining complex stress distribution. The third section gives some experimental results showing the effect of types of complex stress on the capacity of various metals for flow. An appendix discusses some details on methods of obtaining complex stresses. I. REPRESENTATION OF COMPLEX STRESS SYSTEMS AND THEIR EFFECT ON PLASTIC FLOW Representation of Stresses It is convenient to have a consistent system of notation that will allow the representation of the entire range of complex stresses from hydrostatic compression to hydrostatic tension on plastic flow in a single diagram. Such a system is provided by making use of the fact that any complex stress system can be resolved into a hydrostatic tension or compression and a pure shear. Consider a stress system having the principal stresses [S1 S2 5 S3]. This system can be resolved into a new system consisting of the three components [ ] and three pure shears. If Sm is positive, the net stress effect is hydrostatic tension; if it is negative, the effect is hydrostatic compression. In order to produce a systematic notation, a quantity that will be referred to as the "triaxial ratio" will be defined by the equation [ ] Sm is given by Eq. I and S1 is the absolute magnitude of the largest principal stress without regard to sign. Table I lists the triaxial ratios for various common systems. [ ] Representation of Strains Strains will be plotted as "true" strains according to the system proposed by Ludwik.1

Jan 1, 1944

LIST OF THE MEETINGS OF THE INSTITUTE AND THEIR LOCALITIES FROM ITS ORGANIZATION Transactions Number Place Date Vol Page 1 Wilkes-Barre, Pa May, ?71 1 3 2 Bethlehem, Pa August, ?71 1 10 3 Troy, N Y November, ?71 1 13 4 Philadelphia, Pa February, ?72 1 17 5 New York, N Y May, ?72 1 20 6 Pittsburg, Pa October, ?72 1 25 7 Boston, Mass February, ?73 1 28 8 Philadelphia, Pa May, ?73 2 3 9 Easton, Pa October, ?73 2 7 10 New York, N Y February, ?74 211 11 St Louis, Mo May, ?74 3 3 12 Hazleton, Pa October, ?74 3 8 13 New Haven, Conn February, ?75 3 15 14 Dover, N J May, ?75 4 3 15 Cleveland, 0 October, ?75 4 9 16 Washington, D C February, ?76 4 18 17 Philadelphia, Pat June, ?76, 5 3 18 Philadelphia, Pa October, ?76 5 9 19 New York, N Y February, ?77 5 27 20 Wilkes-Barre, Pa May, ?77 6 3 21 Amenia, N Y October, ?77 6 10 22 Philadelphia, Pa. May, ?78 6 18 23 Chattanooga, Tenn May ?78 7 3 24 Lake George, N.Y. Oct. ?78 7 25 Baltimore, Md. Feb. ?79 7 26 Pittsburgh Pa, May ?79 8 27 Montreal, Canada Sept, ?79 8 28 New York, N.Y. Feb, ?80 9 29 Lake Superior, Mich Aug, ?80 9 30 Philadelphia, Pa. Feb, 81 9 31 Staunton, Va May, 81 10

Jan 1, 1923

By William T. Folwell

The Lucky Friday mine east of Mullan, Idaho, is an outstanding example of a property in the Coeur d’Alene district where a small and insignificant- appearing silver-lead-zinc vein at the surface has changed at depth into a large vein of great importance. The Lucky Friday vein has little if any surface expression and above the 1200 level the ore shoots are small and discontinuous. Between the 1200 level and 2450 level, the lowest developed level, the main ore shoot has shown remarkable improvement on each succeeding lower level, and today the mine is one of the major lead-silver producers in the Coeur d'Alene district. History: The Lucky Friday property is on the north side of the South Fork of the Coeur d'Alene River in sections 25, 26 and 35, T. 48 N., R. 5 E., Hunter mining district, Shoshone County, Idaho. This is about one mile east of the town of Mullan, which serves the eastern portion of the Coeur d'Alene district. The southern part of the property is crossed by a branch line of the Northern Pacific Ry. and by U. S. Highway 10. This highway is the principal road crossing the panhandle of Idaho and connects the district with Spokane, Wash., on the west and Missoula, Mont., on the east. The main portal and surface plant of the Lucky Friday mine is at an elevation of 3365 ft, only a short distance above the valley floor and a few hundred feet from U. S. Highway 10, so the mine is readily accessible for year-round operation.

Jan 12, 1958

LIST OF THE MEETINGS OF THE INSTITUTE AND THEIR LOCALITIES FROM ITS ORGANIZATION Transactions Number Place Date Vol Page 1 Wilkes-Barre, Pa May, ?71 1 3 2 Bethlehem, Pa August, ?71 1 10 3 Troy, N Y November, ?71 1 13 4 Philadelphia, Pa February, ?72 1 17 5 New York, N Y May, ?72 1 20 6 Pittsburg, Pa October, ?72 1 25 7 Boston, Mass February, ?73 1 28 8 Philadelphia, Pa May, ?73 2 3 9 Easton, Pa October, ?73 2 7 10 New York, N Y February, ?74 211 11 St Louis, Mo May, ?74 3 3 12 Hazleton, Pa October, ?74 3 8 13 New Haven, Conn February, ?75 3 15 14 Dover, N J May, ?75 4 3 15 Cleveland, 0 October, ?75 4 9 16 Washington, D C February, ?76 4 18 17 Philadelphia, Pat June, ?76, 5 3 18 Philadelphia, Pa October, ?76 5 9 19 New York, N Y February, ?77 5 27 20 Wilkes-Barre, Pa May, ?77 6 3 21 Amenia, N Y October, ?77 6 10 22 Philadelphia, Pa. May, ?78 6 18 23 Chattanooga, Tenn May ?78 7 3 24 Lake George, N.Y. Oct. ?78 7 25 Baltimore, Md. Feb. ?79 7 26 Pittsburgh Pa, May ?79 8 27 Montreal, Canada Sept, ?79 8 28 New York, N.Y. Feb, ?80 9 29 Lake Superior, Mich Aug, ?80

Jan 1, 1923

By T. Smales, P. L. Wright, K. Barron

The results of field measurements using an air injection technique to investigate the integrity of three pillars in the Mclntyre Mines Ltd. 114 seam at Crande Cache, Alberta, are reported. These pillars are shown to be stable and an estimate is made of the strength of these pillars. Using these results, together with Hustrulid's pillar strength formulae, a general pillar strength equation for the 114 seam is developed. This equation takes into account the probability of failure that may be deemed tolerable within the design requirements. The variation of the minimum pillar dimensions with depth below surface for uniform pillar layouts is then examined for a range of tolerable failure probabilities. A general equation of the form B = 2.60 +B Z relates the minimum pillar design width, B metres, with the overburden depth, Z metres, where B is a parameter related to the tolerable failure probability. This equation only applies for 3.05m high pillars with 6.lm wide entries. Finally, the question of what failure probability is tolerable within the design of mine pillars is discussed. This will depend on the support role required of that pillar and the anticipated period over which the pillar must perform that role. Five categories of coal mine pillars are defined, each with a somewhat different role to perform. Tentative design criteria in the form of "tolerable probabilities of failure" are suggested for four of these five categories of pillar. The fifth category of pillar has no reliable support function to perform and support must be obtained by other means.

Jan 1, 1983

By Arnt Solbu, Jomar Thonstad

The reaction between CO, and aluminum in cryolite-alumina melts in contact with aluminum has been studied by passing CO2 over the melt. In unstirred melts a homogeneous reaction between dissolved metal and dissolved CO2 was observed. In stirred melts in which convection was induced by bubbling argon through the melt, the dissolved metal apparently reacted mainly with gaseous CO2. The rate of formation of CO increased slightly with increasing depth of the melt, and it did not depend on whether CO2 was passed over or bubbled through the melt. The rate of formation of CO increased with increasing area of the metal/melt interface and with the application of anodic current to the metal. It is concluded that the dissolution of metal into the melt is the rate-determining reaction. THE current efficiency in aluminum electrolysis is determined by the rate of the recombination reaction between the anode gas and the metal: 2A1 + 3CO2—A12O3 + 3CO [1] as originally stated by Pearson and waddington.1 The occurrence of this reaction in cryolite-alumina melts in contact with aluminum was first verified experimentally by Schadinger.2 Thonstad3 has shown that the reaction may proceed further to give free carbon: 2A1 + 3CO— A12O3 + 3C [2] Normally only a few percent of the CO formed undergoes such reduction. The mechanism of these reactions has not yet been clarified. Aluminum, as well as CO,, is soluble in the melt. The solubility of aluminum in cryolite-alumina melts at around 1000°C corresponds to 75 x 10- 6 mole A1 per cu cm,4 while that of CO2 is only 3 x 10-6 mole CO, per cu cm.5 Taking into account the stoichiometry of Reaction [I], the ratio between dissolved aluminum and dissolved CO2 available for the reaction in a saturated melt is about 40. Therefore, as will be shown in the following, the reaction probably mainly occurs between gaseous COa and dissolved aluminum. The dissolved aluminum presumably consists of subvalent ions of aluminum and sodium.4'6 Since the interpretation of the present results is not dependent upon the nature of this solution, the dissolved metal will be designated solely as Al+ in the following. The reaction can then be divided into four steps: A) dissolution of metal, e.g., 2A1 + Al3 — 3A1+ [3] B) diffusion of dissolved metal through a boundary layer; C) transport of dissolved metal through the bulk of the melt; D) Reaction [1]. If dissolved CO, takes part in the reaction, three additional steps embodying the dissolution and transport of CO2 must be added. schadinger2 observed, when bubbling CO2 through the melt, that the rate of formation of CO (in the following designated rfco) did not depend on the distance from the metal surface. The results also indicate that the rate of bubbling did not affect the rfco. When passing CO, over the melt, Revazyan7 found that the loss of metal did not depend on the depth of the melt above the metal or on the flow rate of CO2, and concluded that Step A is rate-determining. In an unstirred melt, however, Gjerstad and welch8 found that the rfCo decreased with increasing depth of the melt, indicating that step C was rate-determining. It thus appears that the rate control of the process depends on the experimental conditions, particularly on the convection. In the present measurements the reaction has been studied in unstirred as well as in stirred melts. EXPERIMENTAL AND RESULTS The experiments were carried out at 1000°C in a Kanthal furnace with a 10-cm uniform temperature zone (±0.l°C). The melts were made up of "super purity" aluminum (99.998 pct), hand-picked natural cryolite, and reagent-grade alumina. In experiments where alumina crucibles were used, the alumina content in the melt was close to saturation (13.5 wt pct9); otherwise it was 4 wt pct. Pure Co2 (99.85 pct) was passed over the melt, and the exit gas was analyzed for CO2 and CO by the conventional absorption method.3 From the weighed amount of CO (as CO2) the rfco was calculated as the number of moles of CO formed per min per sq cm of the surface area of the melt. The amount of carbon formed by Reaction [2] was not determined. As already indicated the rfco is much higher than the rfC, by Reaction [2]. Since the rfC probably is proportional to the rfco, the measured rfco should then the proportional to, but slightly lower than, the total rate of Reactions [I] and 121. In general the scatter of results obtained in duplicate measurements was ±5 to 10 pct, while within a given run a precision of ±3 to 5 pct was obtained. The various crucible assemblies that were used will be described below. Measurements in Unstirred Melts. When carrying out aluminum electrolysis in small alumina crucibles. Tuset10 observed that after solidification the lower part of the electrolyte was gray and contained free metal, while the upper part near the anode was white and contained no metal. One may test for the presence of free metal by treating with dilute hydrochlorid acid.

Jan 1, 1969

By Roy Moore

THE Tom Reed Extension and Big Jim mines of United Eastern Mining Co. are situated in the Oatman district, Mohave County, Ariz., about 28 miles southwest of Kingman, the nearest railway point. Articles descriptive of the mine surface plant, mill construction, and plant operations have been written by Otto Wartenweiler1, the designing engineer; Wheeler 0. North,2 when he was assistant general manager for the company, and Earle M. Bagley, mill superintendent.3 The last article covers average costs for the full time the mill was operated. The geology of the mines is described by F. L. Ransome4 of the United States Geological Survey. Production was started in January,. 1917, and the known orebodies were exhausted in May, 1925, showing a life of seven years and five months. On account of the comparatively short life and the fact that stoping operations were confined to two orebodies, little experimenting was done with mining methods. The two orebodies were in what is generally called the Tom Reed-United Eastern vein, which occupies a fault fissure having a normal displacement of several hundred feet. Seven or eight very productive ore shoots were mined along this vein in a distance of 1 ½ miles, the two worked by the United Eastern being 1 mile apart. The maximum dimensions of the Tom Reed Extension ore shoot were height, 750. ft.; length, 950 ft.; thickness, 48 ft. The vein structure varied from a solid quartz-calcite filling to a series of parallel stringers separated by andesite. From this body, 511,976 tons of ore were mined. (Fig. 1.)

Jan 1, 1928

By P. R. Sperry, P. A. Beck, J. Towers

Dahl and Pawlek1 found that electrolytic copper develops extremely coarse grains at 1000°C after about 90 pct reduction by rolling. This coarsening occurs only under conditions of penultimate grain size, deformation, and alloying which lead to the "cube" recrystallization texture.l,2,3,5 The peculiar angular shapes and straight grain boundaries of the coarse grains were noted by several investigator.1,4,5 On the other hand, coarsening in Fe-containing aluminum or in AI-Mn alloys8 does not depend on a "cube" (or any well developed) recrystallization texture. It is true that increasing deformation by rolling, and, therefore, an increasingly well developed re-crystallization texture, are associated with decreasing incubation periods of coarsening.6-7-8 Nevertheless, coarsening readily develops in aluminum even after only 30 pct reduction by rolling, where the recrystallization texture is very weak.6,8 Also, coarsening was observed by Jeffries9 many years ago in sintered thoriated tungsten, which presumably has no preferred orientation. In all these cases coarsening is associated with grain growth inhibition by a dispersed second phase.8,9 The annealing temperature has to be suficiently high to overcome the inhibition at a few locations. But if it is too high, growth starts at many points, and the resulting grain size becomes much smaller.9 Normally, the coarse grains are more or less equiaxed, and the boundaries have a typical ragged appearance.6.8 Cook and Macquarie4 demonstrated that, in addition to the texture-dependent coarsening previously found at 1000°C,l electrolytic tough pitch copper may also coarsen at 800°C after 50 pct reduction by cross rolling. The coarse grains formed under such conditions have rounded shapes and ragged boundaries, like those in aluminum. When the annealing temperature is higher, the tendency for their formation decreases. All these observations suggest that the coarsening at 800°C is associated with inhibition by a second phase. Actually, coarsening at 800°C after 50 pct reduction by cross rolling was observed only in tough pitch copper,4 which contains Cu2O particles. On the other hand, the texture-dependent 1000°C coarsening occurs in both tough pitch and oxygen-free copper;4 it does not appear to depend on the presence of a dispersed second phase. However, the interpretation of the 800°C coarsening in Cu after 50 pct rolling as an inhibition-dependent process, similar to the coarsening in A1-Mn alloys, is somewhat weakened by the fact that this coarsening was reported4 to occur only after cross rolling, and not after straight rolling. It was, therefore, decided to re-examine this question. A 1 in. diam electrolytic tough pitch copper rod, No. 2 hard drawn, was annealed for 20 min at 700°C, rolled to 0.5 in., annealed 10 min at 700°C, and straight rolled to 0.064 in. It was then given a penultimate anneal of 20 min at 500°C and it was cut into four sections, which were given final reductions by straight rolling as follows: A 30 pct reduction of area B 50 pct reduction of area C 70 pct reduction of area D 90 pct reduction of area Specimens cut from the four sections were finally annealed at 800°C in an oxidizing atmosphere. Strip A remained fine grained up to 10 hr, but the specimen annealed 12 hr consisted of only 2 large grains. Strip B had a few scattered large (1/2 to 3/4 mm) grains after 1 min, although the balance of the specimen consisted of fine grains of about 0.02 mm. After 5 min there were several 10 to 15 mm grains present, and after 1 hr strip B was completely coarsened. The coarse grains had the same characteristics (see Fig 1) as those obtained by Cook and Macquarie at 800°C after cross rolling. Strip C had several grains of 0.05 to 1 mm after 1 min, but it was still largely fine grained after 12 hr. After 48 hr it consisted entirely of grains of about 0.5 to 4 mm, with an extraordinarily large number of twin bands. Strip D remained com- pletely fine grained after 4 hr at 800°C. These results indicate that, in the deformation range of 30 to 70 pct reduction, the incubation period for coarsening as well as the rate of growth and the final size of the coarse grains decreases with increasing deformation. Similar

Jan 1, 1950