Search Documents

By Warren Weaver

Some of the advantages of the inductive method of electrical prospecting were emphasized in a paper by Dr. Max Mason. Since this emphasis was misunderstood by some to indicate a too exclusive interest in inductive methods, we are especially happy to help correct this impression by reporting on some of our work on surface potential methods. Earth Currents and Surface Potentials When two electrodes, A and B, are placed on the surface of the ground a distance d apart, and a 100-volt battery is connected across the elec- trodes, current flows from the positive electrode A to the negative electrode B. The current does not pass directly from A to B along the straight line joining these two points. In fact, a negligible amount of the total current chooses this direct path, and the flow actually takes place

Jan 1, 1929

EXECUTIVE COMMITTEES OF LOCAL SECTIONS New York L. W. FRANCIS, Chairman, WILLARD S. MORSE, Vice-Chairman. THOMAS -T. READ, Secretary, Woolworth Bldg., New York: N. Y. P. A. MOSMAN, Treasurer. LOUIS D: HUNTOON. Boston HENRY L. SMYTH, Chairman. ALFRED C. LANE, Vice-Chairman. AUGUSTUS H. EUSTIS, Secretary-Treasurer, 131 State St., Boston, Mass. TIMOTHY W. SPRAGUE, HENRY A. WENTWORTH. Columbia F. A. THOMSON, Chairman. GEORGE W. RODDEWIG, Vice-Chairman. LYNDON K. ARMSTRONG, Secretary-Treasurer, P. O. Drawer 2154, Spokane, Wash. R. S. McCAFFERY, S. H. RICHARDSON.

Jan 10, 1914

By R. I. Jaffee, F. C. Holden, H. R. Ogden

Ti-Mn alloys were studied in order to determine the factors affecting the mechanical properties of &stabilized titanium alloys. The principal compositional factors have been found to be solid-solution strengthening, the martensitic transformation, and instability of the P phase. Structural factors, such as grain size and shape, were found to have more influence on ductility and toughness than on strength. THE alloys of titanium with the &stabilizing elements offer the chief hope for developing useful heat treatments. Conversely, the heat-treatment response possible with the p-stabilizing elements causes difficulties such as weld embrittlement and thermal instability during service at elevated temperature. Therefore, it is of considerable technical importance to understand the factors that govern the heat-treatment responses in these alloys, so as to be able to soften the alloys if they are embrittled by an adventitious heat treatment, to render them stable in service, or to harden them while maintaining adequate ductility and toughness. The Ti-Mn system is a good example with which to demonstrate the factors that govern strength and heat-treatment response in. a P-stabilized system. The region of the a-j3 transformation, after the work of Maykuth, Ogden, and Jaffee,' is shown in Fig. 1. The a solubilities are low. This means that, in the two-phase a-p field, partition of manganese between the two phases occurs practically only to the P phase. The solid-solution effects are therefore chiefly the result of solution in the ,8 phase. The eutectoid occurs at 20 pct Mn and 550°C, and proceeds very sluggishly. It can be ignored in heat treatments in hypo-eutectoid alloys, except for long-time storage at low temperatures below the eutectoid temperature. Even here, there is some question that diffusion proceeds far enough to involve eutectoid decomposition. Three compositional factors bear on the mechanical properties and structure of titanium a-P alloys. These are: 1—solid-solution strengthening, 2—mar-tensite transformation, and 3—P instability. Combined with the purely structural factors of grain size and shape, these govern the properties of the alloys. Perhaps the most important factor is solid-solution strengthening. The facts that manganese dissolves in a titanium to a maximum of only 0.5 pct and most of the manganese partitions to the p phase dominate the structure and properties of the a-/3 alloys. Because of the overwhelming preference of manganese for the ,3 phase, the P phase is harder and the a phase is softer. The relative amounts of both phases may be modified by heat treatment in the a-P field as application of the lever rule in the two-phase field clearly shows. A critical temperature in the Ti-Mn diagram is 800°C. This is the P transus temperature of the 6.4 pct Mn alloy, which has the lowest manganese content which will form all retained-P phase on quenching to room temperature. Any quench performed from below 800°C will produce a structure consisting of only a, a plus retained P, or retained p, for alloys of the compositions used in this work. For alloys containing less than 6.4 pct Mn, any quench from above 800°C will always result in structures

Jan 1, 1955

By A. Rosenzweig, M. M. Karnowsky

The generalized binary diagram of the Au-Sn-Pb system is presented. The chief characteristic is the pseudobinary between lead and AuSn. On the high-gold side of the pseudobinary, four reaction isotherms were found at 363°, 274°, 225°, and 204°C. On the low-gold side of the pseudobinary, three reaction isotherms at 277°, 205°, and 176°C were found. A vertical section from the eutectic solder composition to gold was developed from the isothermal sections generated. METALLURGICAL problems arising from the intentional or accidental formation of Au-Sn-Pb alloys in soldering applications have indicated a need for information on the phase equilibria in this ternary system. The literature indicates that a complete diagram presenting this information has not been prepared. Even the recent excellent work by prince' entitled "The Au-Pb-Sn Ternary System" actually extends only to the AuSn or 50 at. pct Au range. There are more events in the 50 to 100 at. pct Au than in the less than 50 at. pct range. This article by Prince furnished much guidance in preparing this report. Initial efforts with the isopleth from the eutectic solder composition 61.9 wt pct Sn-38.1 wt pct Pb and 0 to 90 wt pct Au indicated that extensions of the work to the whole of the diagram (even if in less detail) were necessary to understand this system. In undertaking the construction of such a diagram, this study used differential thermal analyses (DTA) and metallography with verification and identification of phases by electron microprobe analyses. EXPERIMENTAL PROCEDURE The procedure began with the development of a cell which would be unreactive with lead, tin, and gold, retain its strength to 1100°C, exclude oxygen, and also furnish a suitable well for insertion and protection of a thermocouple to monitor temperature changes in the material. Such a cell would make it possible to reuse the sample should a repetition of determination be desired. Fig. 1 is a photograph of two such quartz cells as finally developed. The specimens have been run many times. The cell is made up of an inner sample bucket and an outer capsule with a thermocouple well which the sample surrounds. The DTA unit was an assembly devised from available instrumentation. The sample holder was a 3-in.-long cylinder of nickel, 2 1/2 in. in diam, fitted securely in a vertical tube furnace. Two holes 7/8 in. in diam were symmetrically located to receive the quartz cells. A dummy cell with nickel powder was used as the reference. The absolute temperature was recorded with a Pt/Pt 10-pct Rh thermocouple with an automatic Kaye ice bath compensator. (Calibration indicated the compensator to be accurate to better than ±0.5° over the 100° to 1000°C range.) The absolute temperature was measured in the sample and recorded on an L&N mul-tirange recorder with a full-scale display of 1 mv and automatic ranging of 99 mv. The differential signal between the sample temperature and the reference temperature was amplified with an L&N microvolt amplifier and recorded on an L&N Speedomax G recorder. Since two charts were obtained, the chart speeds were equalized and each event was noted on the differential chart, and the temperature at that time was obtained from the temperature record.

Jan 1, 1969

By H. Warlimont

Precipitation of cementite plates from carbon-rich austenite gives rise to localized formation of martensite in the matrix next to the precipitate upon quenching. The growth of this martensite is urmsually restricted. A distinct lattice relationship is found between this martensite and the cementite phase. It is concluded that the martensite is nacleated in the region of lowest carbon concentration at the austenite-cementite integace, with low-energy matching to the cementite lattice. The restriction of the growth of the martensite phase to the vicinity of the cementite plates can be explained by misfit in the austenite -martensite glissile interface and the uphill carbon gradient in the austenite. The investigation was carried out by light- and electron-microscopy and by electron-diffraction of thin foils. THE precipitation of cementite from carbon-rich austenite gives rise to a localized formation of martensite in the matrix next to the precipitate upon quenching. This behavior has been described previously by A. B. Greninger and A. R. Troiano1 for a 1.78 pct C steel ("pseudomorphic martensite"), and by A. Hultgren2 for a Fe-0.99 pct C—5.20 pct Mn steel. By measuring the axial ratio of the martensite phase,' it was found that the carbon content of this martensite was lower than that of the martensite which formed without prior partial transformation to cementite. Therefore, it was concluded that lowering of the carbon concentration in the matrix next to the precipitate raised the M, temperature locally and thus led to nucleation of martensite at the ce-mentite-austenite interface. However, the restriction of the growth of the martensite phase to the vicinity of the cementite plates cannot be explained by localized carbon impoverishment alone. Under ordinary conditions a martensite plate which has been nucleated in a favorable site can continue to grow until it encounters a physical barrier, or until it encounters a compositional barrier in the matrix, where the total free-energy change for the reaction is not negative, since it has been shown by numerous experiments that the activation energy for growth of ordinary martensite is virtually zero.' An investigation was conducted by transmission electron microscopy and electron diffraction to study this extraordinary martensite. It was expected that crystallographic in addition to compo- sitional conditions caused this abnormal mode of martensite formation. MATERIAL AND PROCEDURE An Fe-C-Mn alloy containing 1.15 pct C and 5.1 pct Mn was used for most of the experiments. For metallographic work, specimens 15 by 15 by 3 mm were used. For preparation of thin films, the initial specimen size was 15 by 15 by 0.7 mm. The heat treatment of each specimen consisted of austenitizing for 1/2hr at 1100'~ in an argon atmosphere, subsequent isothermal holding in a lead bath at one of a series of different temperatures, and, finally, quenching into brine at room temperature. For metallographic examination, the specimens were mounted in bakelite (slight tempering effects resulted from the heat in the mounting press) ground, polished, and etched in nital. Thin films for transmission electron microscopy and electron diffraction were obtained by grinding the specimens to approximately 0.2 mm thickness and subsequent electrolytical thinning. The orientation relationships were evaluated from sets of electron micrographs and selected-area electron diffraction patterns of single units of the transformation products and the adjacent matrix. RESULTS In the alloy which was used for the experiments, proeutectoid carbide precipitates over a wide tem-

Jan 1, 1962

By K. A. Svanstrom, W. R. Opie

Problems associated with deposition of titanium infused chloride baths using TiCl4 as a feed material are reviewed. A potentially workable cell design using Alumdum diaphragms is discussed. Problems inherent to diaphragm-type cells are emphasized and experiments described which led to the development of a cell without a diaphragm. INVESTIGATORS assigned the problem of developing a commercial electrolytic process for making titanium metal have found that they are faced with the difficulty of overtaking rapidly improving magnesium and -sodium reduction processes. The ultimate goal must be a simple process that can be conducted with low-cost raw materials in equipment that is inexpensive to build and operate. This paper deals with some work of a preliminary nature conducted by personnel of the Research Laboratories, National Lead Co., Titanium Division. Many others have reported on the general subject of depositing titanium from fused salt electrolytes. A partial bibliography is attached. PRIMARY CONCEPTS Anyone attempting to deposit titanium metal from fused chloride electrolytes using TiC14 as a source material is faced with at least the following problems: 1) Getting the titanium into solution in the bath. 2) Controlling oxidation and reduction reactions associated with multiple valency of titanium, which result in low electrical efficiency. a) Prevention of dichloride oxidizing to trichloride and trichloride to tetrachloride at the anode. b) Prevention of trichloride from being reduced only to dichloride at the cathode. 3) Prevention of metal formation away from the cathode thru disproportionation of titanium dichloride to satisfy the equilibrium. 3 TiCl2 == 2 TiCl3 Ti [I] 4) Protection of reduced chlorides in the bath from reaction with oxygen from the bath container or air. 5) Removing from the bath the deposited metal formed without allowing it to become contaminated with oxygen or nitrogen. 6) Selecting cell materials of construction which will not contaminate the metal product or cause cell inefficiencies. EARLY EXPERIMENTS The initial attempts made to electrodeposit titanium from fused chloride baths using TiCl4 as a source material indicated that TiCl4 was essentially insoluble in molten chlorides of alkali and alkaline earth metals. If, however, the TiCl4 was introduced close to a nickel cathode some metal could be formed and appreciable quantities of reduced chlorides would be present. These reduced chlorides reacted with any oxygen present in the atmosphere above the bath and would also diffuse quickly to the anode (graphite) where they would react with chlorine and be evolved as TiCl4. The net electrolytic efficiency of such a cell is essentially nil. The TiCl4 evolves from the anode at almost the same rate that it is introduced in the vicinity of the cathode although a small amount of titanium metal can be formed on the walls of the cathodic pipe. It becomes apparent from the operation of such a cell that: metal can be deposited on the cathode, although of small particle size and a nonadherent nature, and when reduced chlorides are present in the electrolyte, TiCl4 can be absorbed, probably by the following reactions: TiCl4 + TiC12 - 2 TiCl3 [21 2 TiCl3 + 2 faradays - 2 TiC12 + C12 [31 DIAPHRAGM CELLS These facts encouraged the development of cells containing diaphragms to prevent migration of reduced chlorides from the cathode area to the anode. Small-scale cells for process development are described by Alpert, Schultz, and Sullivan.20 A larger brick-lined cell incorporating internal heating by passing ac through the bath is shown in

Jan 1, 1960

(Separates of all the Pamphlets published within the last three years (starting with No. 1469) are available, with few exceptions, at Institute headquarters. Separates of all the Technical Publications are also available. All the papers are on file in public, university and technical libraries, and when so indicated in the list, may be found either in the TRANSACTIONS or in the special volumes issued by the Institute.] METAL MINING TRANS. PAMPHLET VOLUME NUMBER TITLE AUTHOR MEETING AND PAGE'S 1033 Relation of Air Pressure to Drilling H. W. Seamon New York 66, 743, 756 Speeds of Hammer Drills Feb., 1921 1034 Time Factor in Depletion of Mines J. W. Roberts New York 66,711 Feb., 1921 1035 Calculation of Ore Tonnage and Grade J. E. Harding New York 66, 117, 122 from Drill-hole Samples Feb., 1921 1038 Discussion of Papers Presented in 1920 1041 Run-off and Mine Drainage H. N. Eavenson New York 66, 624,630 Feb., 1921 1043 The Health of the Underground Worker A. J. Lanza New York Feb., 1021 1045 Dust-ventilation Studies in Metal D. Harrington New York 66, 272, 289Mines Feb., 1921 1046 Discussion of Papers Presented at New York Meeting, 1021 1072 Breakage and Heat Treatment of Rock- B. F. Tillson New York 66,1758 drill Steel Feb., 1921 1073 Discussion of Papers Presented at New York Meeting, 1921 1076-M Analysis of Some Drill-steel Testa F. B. Foley New York 66, 789 Feb., 1921 1077-M Application of Magnetic. Analysis to C. W. Burrows New York 66, 796 Rock Drills Feb. 1921 1078-M Heat Treatment of Rock-drill Steel G.H. Gilman New York 66, 779_ Feb., 1921 1080 What is the Ideal Drill Steel? F. H. Kingdon New York 66,807 Feb., 1921 1051 Rock Drill Steels, etc. - W. T. Ober New York 66,814 Feb., 1921 1082-C Underground Mine Development, Its J. B. Dilworth Wilkes-Barre 66,715, 7.21 Definition and Valuation Sept., 1921 1084 Drill Steel from Hollow Ingots P. A. E. Armstrong New York, 66, 809 Feb., 1921 1088 Authors' Replies to Recent Discussion 1092 Heat Treatment of Rock Drill Steel A. E. Perkins New York 66,811 Feb., 1921 1096-M Octagonal Ventilation Shaft. of Davis- J. L. Bruce Wilkes-Barre 66, 252, 268 Daly Copper Co. Sept., 1921 1100-C Determination of Electrical Equip- G. Bright Wilkes-Barre 66, 580, 605 ment for a Mine Hoist Sept. 1921

Jan 1, 1927

By Raymond Mancha

TO determine the optimum inside dimension of an airshaft, it is necessary to strike the proper balance between the cost of power for air friction and turbulence losses within the airshaft, on the one hand, and the cost of sinking the airshaft, on the other. An airshaft larger than necessary is extravagant because the increased amortization rate exceeds the reduction in power cost. Conversely, an airshaft smaller than is needed wastes more power than can be justified by the corresponding reduction in amortization rate. The airshaft power is a function of the air volume Q cu. ft. per min. and the air-pressure drop p in. water gauge across the airshaft. † The airshaft pressure drop is calculated by the formula: R X l X O X Q2 P= 5-2 X A3 wherein R is the air-friction coefficient for the portion of the airshaft 1 ft. deep, 0 ft. in perimeter and A sq. ft. in area, when passing an air volume of Q cu. ft. per minute. The airpower w kw. dissipated within the airshaft is computed by the formula: w= 0.746Q X P 6345 With a combined over-all unit efficiency of the fan and drive of E per cent, the air-shaft power requirements become W = 100w There are 8760 hours per year, so with energy at I cent per kilowatt-hour the annual cost per kilowatt for continuous operation is $87.60; therefore, with energy at C¢ per kw-hr. the annual airshaft power cost MI dollars per year becomes: M1 = 87.60W X C which, expressed by the component parts, becomes: M1 V 5.2X4' / XQ Z 0.746 X 100 X 87.60 X C V 60 X E ) The amortization rate a dollars per year per dollar invested includes both depreciation and interest on the investment. These in turn involve the expected useful life of the airshaft and available molletary interest rate. For the purpose of this discussion the cost of airshaft sinking is assumed to comprise a fixed move-in charge Y dollars, which covers the cost of transporting, erecting and removing the shaft-sinking equipment and is independent of the air-shaft size. Also, it is assumed that the cost of shaft lining, if any, is fixed at $X per cu. yd. of lining material and that excava-

Jan 1, 1947

By Raymond Mancha

TO determine the optimum inside dimension of an airshaft, it is necessary to strike the proper balance between the cost of power for air friction and turbulence losses within the airshaft, on the one hand, and the cost of sinking the airshaft, on the other. An airshaft larger than necessary is extravagant because the increased amortization rate exceeds the reduction in power cost. Conversely, an airshaft smaller than is needed wastes more power than can be justified by the corresponding reduction in amortization rate. The airshaft power is a function of the air volume Q cu. ft. per min. and the air-pressure drop p in. water gauge across the airshaft. † The airshaft pressure drop is calculated by the formula: R X l X O X Q2 P= 5-2 X A3 wherein R is the air-friction coefficient for the portion of the airshaft 1 ft. deep, 0 ft. in perimeter and A sq. ft. in area, when passing an air volume of Q cu. ft. per minute. The airpower w kw. dissipated within the airshaft is computed by the formula: w= 0.746Q X P 6345 With a combined over-all unit efficiency of the fan and drive of E per cent, the air-shaft power requirements become W = 100w There are 8760 hours per year, so with energy at I cent per kilowatt-hour the annual cost per kilowatt for continuous operation is $87.60; therefore, with energy at C¢ per kw-hr. the annual airshaft power cost MI dollars per year becomes: M1 = 87.60W X C which, expressed by the component parts, becomes: M1 V 5.2X4' / XQ Z 0.746 X 100 X 87.60 X C V 60 X E ) The amortization rate a dollars per year per dollar invested includes both depreciation and interest on the investment. These in turn involve the expected useful life of the airshaft and available molletary interest rate. For the purpose of this discussion the cost of airshaft sinking is assumed to comprise a fixed move-in charge Y dollars, which covers the cost of transporting, erecting and removing the shaft-sinking equipment and is independent of the air-shaft size. Also, it is assumed that the cost of shaft lining, if any, is fixed at $X per cu. yd. of lining material and that excava-

Jan 1, 1947

[¦Abad, Leopoldo F., College of Min., Univ. of California Berkeley, Cal. '23 ¦Abarquez, Ramon F., Met., Bureau of Science Manila, P. 1. '24 ¦Abbey, Robert Graham, Student, Case School of Applied Science Cleveland, Ohio. '21 Abbott, A. N., Mines Supt., Mazapil Copper Co., Ltd ...Concepcion del Oro, Zacatecas, Mexico. '23 Abbott, Clarence B., Gen'l Supt., Iron Mines Div., Tenn. Coal, Iron & R. R. Co., 1405 Minnesota Ave., Bessemer, Ala.' '04] [ ]

Jan 1, 1925

By A. Blake Caldwell

The continental margins and ocean basins represent by far the largest unexplored frontier for discovery and development of mineral values. In April 1971, Deepsea Ventures, Inc., at a press conference in its Gloucester Point, Va. headquarters, announced what may become the first commercial deepsea venture--15,000 ft below the ocean surface. During the presentation, Deepsea's president, John E. Flipse, announced that studies made by company personnel had proven the feasibility of mining, transporting, and milling L million tpy of manganese nodules taken from the ocean floor, and that, based on current conditions and prevailing construction costs, a venture of this magnitude would cost about $200 million split evenly between mining and processing facilities. A subsidiary of Tenneco, Deepsea has tentatively reserved 1976 as the year in which it plans to mine, process and market the nodules.

Jan 1, 1971

By J. H. Bowden

On the night of January 23d, 1889, Erich C. Schaufuss, a member of the Institute, while engaged in making the regular mine-surveys in No. 4 Slope of the Susquehanna Coal Co, at Nanticoke, Pa., was killed by an explosion of fire-damp. The catastrophe, which cost another life also, is attributed to the carelessness of the " fire-boss" leading the party, who advanced with a naked lamp into a chamber

Jan 1, 1889

By Alfred E. Hunt, Geo. H. Clapp

The great advantage to he gained by a careful inspection of all materials used in construction, in regard to safety and permanency, and to accuracy of workmanship, has been longer recognized in Great Britain and on the Continent than in America. This is probably owing to the way in which great enterprises have been " rushed" in the United States. Time and cheap first cost, in too many instances, have been considered of more importance than absolute safety and durability; the structures being afterwards strengthened if this was found necessary. This has been most noticeable in railroads built through " new " country, where the object in view has been to reach the terminal point and get to work. These " loose " methods are now being very largely done away with, and a large proportion of the structures lately built are not only carefully inspected during the process of

Jan 1, 1891

By Leon Dominian

I. INTRODUCTION APART from local needs in the country's development, the interest attached to Turkish fuels grows chiefly out of the importance of the railway industry in Asiatic Turkey. Geographical conditions make of the tracks in this region the most essential link in a chain of world-girdling lines. By its position Asia Minor can be likened to a bridge in the through routes destined to connect European factories with Asiatic and African markets. An uninterrupted right-of-way from points in central Europe to populous Indian cities must necessarily pass through the valleys A Anatolia. Crossing thence the Taurus Mountains its natural passage is indicated as the Mesopotamian lowland and the shores of the Indian Ocean. Similarly land connection between Europe and Africa can be obtained only by passage through Asia Minor, Syria and Palestine. The occurrence of fuel within an area of such significance from the

Jan 6, 1916

By D. E. Thomas

Oxidation rate constants were determined for Cu-Pd and Cu-Pt alloys as a function of alloy composition and temperature. Reaction products were identified. Relationship between oxidation rate constants and diffusion constants in the reaction zones is discussed. THE mechanism of oxidation of alloys is consider--'¦ ably more complex and not so well understood as that of pure metals. The oxidation of a one-phase binary alloy of which one component is a noble metal should logically be the simplest case of alloy oxidation, for in this case one is concerned presumably only with the oxidation of the base metal and the diffusion process by which it reaches the reaction site. This had led a number of investigators to study such systems. Probably the earliest work was that of Tammann and Rienacker' on the tarnishing of Cu-Au alloys at temperatures up to 300°C. These alloys oxidize according to the logarithmic law, that is, the film thickness is proportional to the logarithm of the time. The rate of film formation was found to decrease with increasing gold content. Raub and Engel² investigated the oxidation of Au-Cu, Cu-Pd, Au-Ag-Cu, Au-Pd-Cu, and Ag-Pd-Cu alloys at 750°C in air. These alloy systems form a continuous series of solid solutions at the temperature of oxidation with the exception of those containing silver. Au-Cu alloys oxidize parabolically, the rate constant decreasing slowly with increasing gold content up to about 15 atomic pct* Au, and more rapidly thereafter. An alloy with 14 pct Cu forms only a very thin layer of CuO on the surface and probably also a subscale.† Alloys containing more than 14 pct Cu have scales consisting of CuO and subscales of Cu,O embedded in nearly pure gold. The Cu-Pd alloys obey the parabolic law except for slight deviations for alloys containing 53 and 84 pct Cu. The overall oxidation rate decreases rather smoothly with increasing palladium content. A scale consisting of a thin layer of CuO and a thick layer of Cu²O and a subscale of Cu²O embedded in nearly pure palladium was observed for alloys containing 84, 53, and 30 pct Cu. An alloy with 19 pct Cu formed only a subscale of CuO embedded in nearly pure palladium, and an 8 pct Cu alloy formed only a thin film which was presumed to be PdO. The ternary alloys considered by Raub and Engel oxidized in much the same way as did the binaries. Wagner and Grunewald³ found that Ni-Au alloys, oxidized at 900°C in oxygen, formed a scale of NiO and a subscale containing NiO and gold. The overall oxidation rate increases with increasing gold content, passes through a maximum and then decreases to zero for pure gold. The fact that gold increases the oxidation rate was attributed to "pores" in the scale. Kubaschewski' investigated the air oxidation of alloys of platinum or gold with up to 10 pct Cu or Ni at temperatures in the range 300" to 1550°C. Parabolic behavior was noted for all alloys except Au-Ni, with rates decreasing in the order Au-Cu, Au-Ni, Pt-Ni, Pt-Cu. The diffusion coefficients for the latter three systems are known, and they decrease in the same order. However, activation energies for the oxidation of these alloys do not agree with the activation energies for diffusion in the respective alloy systems. Kubaschewski and Goldbeck7 nvestigated the oxidation of Ni-Pt alloys in air at temperatures of 600"

Jan 1, 1952

By R. W. Harper, M. C. Fuerstenau, J. D. Miller

Data were obtained with hematite with octyl hydroxamate and oleate as collectors to determine the mechanism of collector adsorption and also to establish the roles that conditioning time and temperature assume in this system. The presence of hydroxy complexes of iron are apparently necessary for chemisorption of collector to occur, and it appears as if extended conditioning times and elevated temperature promote greater dissolution of the mineral and hence greater concentrations of hydroxy complexes. Data were also collected with two natural hematitic ores. With one ore that was ground to 70% - 15u, an addition of 0.4 lb per ton hydrox-amate resulted in a final concentrate recovery of 86% at a grade of 64% iron. Satisfactory concentrate grade could not be obtained with fatty acid under these conditions. The effect of conditioning time prior to the addition of collector was also examined with another ore. When the ore was conditioned for 3 min, 19% of the iron was recovered in the concentrate at a grade of 67% Fe. With a conditioning time of 7 min, 73% of the iron reported to the concentrate at a grade of 62% Fe. These products were obtained with an addition of 0.2 lb per ton hydroxamate. This ore also responded well to flotation with fatty acid. The use of potassium octyl hydroxamate as a collector for iron oxide has been the subject of an earlier study.1 From this initial study, it was suspected that hydrolysis of surface ferric iron was involved in collector adsorption. More recent work in other systems has shown that slight mineral dissolution, followed by hydrolysis of cations common to the mineral lattice, is apparently necessary, and that adsorption of these hydroxy com- plexes probably occurs prior to the adsorption of collector. In view of these facts, work has been continued on the hydroxamate-hematite and oleate-hematite systems to gain better insight into the mechanisms of collector adsorption on iron oxide. Microflotation experiments in conjunction with infrared analysis were used in this regard. In addition it was felt necessary to establish more clearly the responses that can be obtained from natural ores when hydroxamate is used as collector as compared to those obtained when fatty acid is used. Experimental Materials and Techniques Pure potassium oleate, potassium octyl hydroxamate, synthesized from hydroxylamine hydrochloride and methyl octanoate, and a distilled Tall Oil were employed as collectors. Conductivity water was used in the

Jan 1, 1971

By J. H. Jackson, P. D. Frost, C. H. Lorig, A. C. Loonam

THIS paper describes (1) the effect of sodium on the tensile ductility of magnesium-lithium base alloys, and (2) the precautions necessary to avoid sodium contamination. Effect of Sodium on Properties The harmful effect of sodium in magnesium-lithium base alloys has been recognized by other investigators.1,2 However, quantitative data on its effect have not hitherto been published. The effect of sodium on the ductility of an 87 pct Mg-9 pct Li-4 pct Zn alloy is shown in fig. 1. The sodium was introduced by adding sodium chloride to the flux, which consisted of 50 pct LiC1-50 pct KC1. Moderately high ductility was obtained at an indicated sodium content of about 0.05 pct. Actually, the value probably was less than 0.05 pct, be-cause analyses of the lithium, which were more accurate than those of the alloy, indicated that the highest ductility was obtained when its sodium content was 0.02 pct or less. In fact, in alloys having somewhat greater strength than this particular composition, the sodium content of the lithium must be under 0.01 pct to avoid embrittlement. The effect of sodium on the ductility of several alloys having fairly low strength levels is shown in fig. 2. These alloys were melted with a flux containing 75 pct LiC1-25 pct LiF, to which various quantities of sodium chloride had been added. Sodium had no significant effect in pure magnesium but markedly decreased the ductility of the magnesium-lithium-zinc alloy, even though the alloy contained only 3 pct Li. As the lithium content of the binary alloys was increased, the ductility of the sodium-free alloys also increased. However, the effect of sodium on ductility became progressively more pronounced. Sodium raised the minimum hot-rolling temperatures, particularly those of the high-strength alloys, but its effect, in this respect, was less critical than its effect on tensile ductility. For optimum cold roll-ability, a low-sodium content was required. No deleterious effect of sodium on the stability of properties at slightly elevated temperatures was found. Control of Sodium in the Alloys Optimum ductility is obtained in the magnesium-lithium base alloys when sodium-free materials are used. Commercial lithium was formerly the chief source of contamination. Now purified lithium is usually available. When lithium of ordinary purity is used, the effects of sodium can be minimized by using a flux consisting of 75 pct LiC1-25 pct LiF. Bubbling nitrogen through the melt also helps remove sodium. The damaging influence of sodium on the ductility of the magnesium-lithium base alloys was observed very early in the alloy development program. The discovery came about as a result of chemical analyses made on the lithium used in the alloys. Ductile alloys had been made from lithium containing about 0.2 pct Na, whereas the lithium used in the brittle alloys contained about 0.9 pct Na. The first magnesium-lithium alloys were melted under commercial fluxes ordinarily used in melting magnesium alloys. These fluxes were unsuited for scavenging the contaminating sodium, and the resulting alloys were not ductile. Next, a flux containing 50 pct KC1-50 pct LiCl was tried. This was somewhat more effective but still was not satisfactory. After a number of different experimental fluxes had been investigated, one containing 75 pct LiC1-25 pct LiF was adopted. This flux made it possible to obtain good ductility in the alloys. The

Jan 1, 1951

By Robert F. Brindisi

There are approximately thirty to forty operating mills in the United States which are currently producing uranium yellow cake. This figure includes a significant number of in situ and by-product operations however the conventional acid leach flowsheet depicted in Figure 1 is the most common process used by domes- tic yellow cake producers. Ground ore is leached with sulphuric acid and the solids are separated from the pregnant liquor and washed in a series of counter- current decantation thickeners. After clarification of the pregnant liquor, uranium is recovered from the solvent extraction/stripping circuit by the addition of ammonia with precipitation of ammonium diuranate, or simply yellow cake. The yellow cake precipitate is then thickened in a conventional thickener before re- porting to the dewatering/washing unit operation. De- watered solids subsequently report to thermal drying, and clarified liquor from the dewatering/washing step is recycled to the yellow cake thickener. The yellow cake dewatering/washing unit operation is generally accomplished using either a solid bowl centrifuge or a rotary vacuum drum filter. In most cases the underflow from the yellow cake thickener is pumped to the dewatering/washing unit operation at approximately fifteen to twenty percent solids. For the purpose of this discussion ammonium diuranate, the product of ammonia precipitation from a conventional acid-leach uranium mill flow sheet, as discussed earlier, will be considered. With only slight adjustment this discussion will be equally applicable to sodium diuranate or uranium perioxide precipitates and to either in situ or by-product operations where such precipitates are generated. Some of the domestic uranium mills which have selected solid bowl centrifuges instead of vacuum drum filters for the dewatering/washing unit operations are listed in Figure 2. Why have these uranium mills, as well as others not listed, chosen to use the solid bowl centrifuge for this particular application? It is simply a matter of performance, reliability and economics. In the remainder of this paper the performance, initial cost, installation cost, maintenance and power consumption of both the solid bowl centrifuge and the rotary vacuum drum filter will be com- pared. In order to provide a consistent basis for comparison the appropriate sized (centrifuge or filter) unit for the dewatering/washing of 400 to 500 pounds per hour of yellow cake solids will be considered. Such a flow will usually be generated by a typical 3,000 ton per day (acid-leach) uranium mill.

Jan 1, 1980

By J. H. Keeler

QUESTIONS raised during earlier investigations of the preferred orientation in titanium1 and zirco-nium sheet led to the examination of the rolling and annealing textures of beryllium and hafnium sheet. Since the c/a ratios are 1.568 and 1.584, less than that of ideal close packing, 1.633, it would be expected that the preferred orientations would be similar to those exhibited by zirconium and titanium. This was found to be true. However, in several regards the orientations observed for beryllium differed from those reported by Smigels-kas and Barrett4 and so are given here. The okientations for hafnium are presented since only the cold-rolling texture has been reported in the unclassified literature. The pole figures were determined by transmission using an X-ray spectrogoniometer with copper Ka radiation and the automatic pole-figure recorder designed by Geisler described previously.6 The intensity units in the pole figures in this case are only arbitrary since calibration against a random sample of the same material was not considered necessary for the original purpose of the examination. The beryllium sheet, cold-rolled from 0.036 in. to 0.006 in., a reduction of about 84 pct, exhibited a split texture with the [1010] direction in the rolling direction as shown in Fig. 1. The same specimen annealed in vacuo for 1 hr at 1000°C showed no change in orientation. These textures

Jan 1, 1959

By Behram H. Wadia, Robert S. Lehto

A new metallurgical treatment for capturing as much of the values as possible from zinc refinery feed has been devised by personnel in Cerro de Pasco's Peruvian operation at La Oroya. Results from a pilot test illustrate that 50 metric tons of throughput can be maintained whereby substantially all the silver and copper report in the sponge iron while zinc and lead with indium and cadmium volatilize in the fume. This new technique named the Zileret (zinc leach residue treatment) process complements the electrolytic system used in the refinery so that all values in a zinc concentrate are economically recovered. In the system, dried zinc leach residue is pelletized with soluble zinc sulfate which already exists in the leach residue and a finely divided reducing agent such as anthracite coal. Upon heat addition, the loose anthracite in the charge maintains a reducing atmosphere while the free space above the charge has an oxidizing nature. When the charge reaches the reaction zone at about 1100°C, metals such as zinc, lead, cadmium and indium volatilize from the pellets, are oxidized in the upper kiln atmosphere, and are then carried in suspension as an oxide fume to the cooler and baghouse. The copper and silver content in their reduced forms, along with the iron, discharge from the kiln as a calcined product which is then cooled to below 130°C before being exposed to the air.

Jan 9, 1969