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By K. Tangri, M. Chaturvedi

The addition of 0.5 wt pct Cu to Zr-2.5 Cb alloy increases the as -quenched hardness of the hexagonal martensitic a' phase, produced by water-quenching bccß-Zr phase, by about 35 pct. This strengthening has been attributed to the solid -solution hardening of the matrix. On aging ternary martensite, a' phase reverts to equilibrium a and Zr2Cu and ß-Cb precipitate out, mainly at the twin and grain boundaries, causing a secondary hardening of the matrix. COLD-worked Zircaloy-2 pressure tubes have been in use in power reactors for a considerable period of time. The search for a better material led to the development of Zr-2.5 wt pct Cb alloy which in the quench-aged condition develops 50 pct more strength than that of cold-worked Zircaloy-2, however, its corrosion resistance in water and steam in the temperature range of 316" to 400°C, in absence of neutron flux, is inferior to that of zircaloy-2.' Work carried out by Ells et al.1 and Dalgaard2 has shown that the corrosion properties of Zr-2.5 wt pct Cb alloy can be considerably improved by the ternary addition of 0.5 wt pct Cu. This paper is concerned with the effect of 0.5 wt pct Cu on the formation of martensitic a and its aging characteristics in a Zr-2.5 wt pct Cb alloy. MATERIALS AND EXPERIMENTAL TECHNIQUES Zr-2.5 Cb-0.5 Cu (referred to as the ternary alloy) and Zr-2.5 Cb (referred to as the binary alloy) alloys, supplied by the Chalk River Nuclear Laboratories of the AECL were used. The detailed chemical analysis is given in Table I. Cold rolling and swagging with frequent intermediate anneal of 1000°C were used for the initial fabrication of the alloys. All the heat treatments were carried out after the specimens were wrapped in zirconium foils and encapsulated in silica tubes under a vacuum of 5 x 10-6 mm of Hg. For optical metallography and hardness measurements specimens were mechanically and then chemically polished in a 45 pct HNOj, 45 pct HzO, and 10 pct HF solution. Hardness was measured on a Vickers hardness tester using a 10-kg load. For each specimen at least fifteen indentations were made in order to obtain a representative value. The phase identification and structural analysis were carried out using X-rays and electron diffraction techniques. Wires of 1.5 mm diam reduced to 0.12 mm diam by chemical etching were used for making Debye-Scherrer powder patterns using Cu Ka radiation in a 114.6 mm diam camera. Carbon extraction replicas were prepared by etching the specimens, after depositing a layer of carbon on the metallographic specimen, in one part HF and thirty parts ethyl alcohol. Thin films were prepared by electropolishing heat-treated 3/4 by 1/2 by 0.005 in. thick strips using a modified Bollman-Window technique. The 10 pct perchloric acid-90 pct methyl alcohol bath was kept at -50°C and polishing was done at 5 to 10 V. The thinned specimens were washed in ethyl alcohol at -30º to -40°C and dried between filter papers. Replicas and thin films were examined in a Phillips 300 G electron microscope. For resistivity measurements thin strip specimens 0.02 by 0.3 by 10.0 cm long were used. The potential leads were spot welded to the specimens in order to maintain a fixed length for the initial and the final resistivity measurements. The resistivity was measured by a Kelvin bridge in a temperature controlled room. The temperature was maintained at 72º ±1°F and the accuracy of the resistivity measurements was 0.03 µa-cm. RESULTS As-Quenched Structures. In order to produce a homogeneous matrix to study the precipitation reaction the solution-treatments of both the alloys were carried out in the -field region. From the Zr-Cb phase diagram due to Lundin and cox3 ß/a + ß phase boundary for Zr-2.5 wt pct Cb alloy is 820°C. Ells et al.1 have reported this boundary for Zr-2.5 Cb alloy containing 1100 ppm 0 to be at 920°C. Also, the addition of 0.5 wt pct Cu reduces this temperature by 50°C. Therefore, the solution-treatments were carried out at 1000°C to ensure that the alloys were in ß-phase region. The soaking time was 1 hr and the specimens were water-quenched. The as-quenched hardness of the binary alloy was 245 Vpn whereas, that of the ternary alloy was 330 Vpn. The X-ray diffraction studies indicated that the as-quenched structure of both the alloys consists of martensitic hexagonal phase a', with a c/a ratio of 1.591, and some retained ß-Zr. The presence of a' phase was further confirmed by thin film electron microscopy. Electron micrographs of typical ß-quenched structures of the ternary and the binary alloys are shown in Figs. 1 and 2, respectively. Fig. 3 shows the diffraction pattern from an area similar to that shown in Fig. 1. Although, the as-quenched hardness of the ternary alloy is about 35 pct greater than that of the binary alloy, the structure of both the alloys seems to be the same. The matrix of both alloys is heavily twinned and shows very few dislocations. Furthermore, there is no evidence of any precipitation taking place in either of the two specimens during quenching from the solution-treatment temperature. Aging Behavior of Martensitic a'. The aging kinetics of the ternary alloy were followed by resistivity and hardness measurements. The as-quenched values

Jan 1, 1970

By William P. Blake

Postscript to the paper read by Prof. William P. Blake at the Lake Superior meeting, September, 1904. POSTSCRIPT.*-Since the publication of my paper upon the blackening of the surface of rocks in desert regions, I have noted that similar views of the origin of superficial discoloration in the case of certain rocks from the Salt Lake basin of Utah have been expressed by Prof. George P. Merrill. In the pamphlet, he described the brown-coated weathered boulders from Tooele county, reaching the conclusion that the discoloration was due to the solution of the manganese and iron compounds 'in the interior of the boulder, while they were in the water, and the gradual bringing of this material to the surface through capillarity and its oxidation when exposed.

Mar 1, 1905

By G. C. Kuczynski

Two particles in mutual contact form a system which is not in thermo-dynamical equilibrium, because its total surface free energy is not a minimum. If such a system is left for a certain period of time, the bonding of the two particles will take place in order to decrease the total surface area, even though the temperature is lower than the melting point. This phenomenon of bonding of two or more particles with the application of heat only and at temperatures below melting point of any component of the system will be called sintering, although the powder metallurgists use this term in a broader sense, including the presence of molten phase and pressure. It is the objective of this paper to study this process and the mechanisms involved in it. This problem is of utmost importance to powder metallurgy and powder ceramics, and its technological aspects have been studied for a good many years. However, the powder metallurgical operations are too complex and include too many superimposing mechanisms, and too many variables for a direct study. It was therefore advisable for the purposes of this study to reduce the variables to a minimum. In this work the radius of the interface formed during bonding in a simple system composed of a spherical particle and a large block of the same metal was studied as a function of time and temperature. It is believed that the mechanism involved in this simple process is fundamental to any sintering operations. Previous Work J. Frenkell was the first to make a serious attempt to develop a theory of sintering. He assumed that the process consists of a slow deformation of crystalline particles under the influence of surface tension which reduces to a viscous flow where the coefficient of viscosity n is related to the self-diffusion coefficient D by the following equation 1 _ D D6kT [1] where 6 is interatomic distance, k the Boltzman constant, and T the absolute temperature. This type of viscous flow of a crystalline substance is essentially different from the ordinary plastic flow. The latter is a specific propcrty of crystals and cannot take place in amorphous bodies. According to Frenkel this viscous type of flow is due to the diffusion of the holes or vacancies arising in the lattice. He was able to derive an equation relating the growth of the interface between two spherical crystalline particles or between a particle and a semi-infinite crystal (Fig 1) to time t at constant temperature. This relationship can be written as follows: x²— 3/2 a/nt [2] where x is the radius of the interface assumed to be circular, a the original radius of the sphere and a surface tension of the material. The other assump- x tions are that x is less than 0.3 and that during the period 1 the original radius, a, of the metallic particle did not change appreciably. A. J. Shaler and J. Wulff2 followed closely the ideas of Frenkel in their theory of sintering of a mass of metallic powder. Neither Frenkel nor Shaler has validated his theoretical speculations with conclusive experimental data. Two measurements reported by

Jan 1, 1950

By G. R. Speich, D. J. Mack

The transformation of was studied by isothermal methods. At all temperatures, the ß transforms quickly to fine grained ß" which develops silver-rich striations. At higher temperatures the striations disappear, the final structure being Widmanstatten a in ß'. At lower temperatures, the striated ß" is consumed by pearlite nodules of a + ß' which in turn form the Widmanstatten a in ß'. The transformation is unlike any previously reported for a eutectoid. AS part of a general program of study on the eutectoid reaction, the Ag-Cd eutectoid seemed particularly attractive because the ß phase which undergoes the reaction is another of the typical "electron compounds," AgCd, having the disordered body-centered cubic structure. This work deals with the eutectoid which occurs at 50.5 wt pct Ag, and not with the eutectoid which has been reported at 42.9 pct Ag.' There has been no work reported specifically on the Ag-Cd eutectoid which was studied. The only information available on this eutectoid is that work done in the determination of the phase diagram. For the most part, the terminal portions of the diagram, Fig. 1, are well established; however, the region from 40 to 60 pct Cd is still somewhat in doubt. The X-ray evidence'..' in this region is in conflict with the results of thermal analysis. X-ray analysis3 does not show a polymorphic transformation at about 240°C as indicated by thermal analysis.1,5 Even the X-ray evidence, within itself, is confusing in this region of the diagram. One investigator2 reports the ß phase as hexagonal close-packed and has found three phases in the one-phase ß field. Other X-ray investigators"' ' report the ß phase as body-centered cubic. The authors believe that much of the confusion resulted from incomplete homogenization of the alloys. It was found that it takes almost a week to produce a completely homogeneous two-phase structure at 370°C. Another possible source of error is the volatilization of cadmium from the surface and cracks of the specimen during annealing with the possible formation of a cadmium-poor phase. The only metallographic data available on this eutectoid alloy are those of Durrant1 and Fraenkel and Wolff.5 his information is very limited. The phase diagram chosen as most nearly correct is that in the Metals Handbook," which is almost identical with the phase diagram suggested by Durrant.1 Materials and Procedure The alloy was prepared from high purity silver (999.5 plus fine) and high purity cadmium (0.02 Pb, 0.005 Cu, trace Fe). The metals were melted in carbon under a nitrogen atmosphere. Some cadmium was lost by volatilization but by remelting several times and adding cadmium a 200 g ingot of approximately eutectoid composition was obtained. Because the CdO fumes are quite toxic, the exit gas from the furnace was led through a water trap to condense the CdO. The ingot was homogenized at 650°C for 24 hr, no attempt was made to prevent decadmiumization. This resulted in a narrow decadmiumized rim on the ingot which was removed. Filings for analysis were obtained from cross sections of the ingot. Silver was determined by the Volhard thiocyanate method,? cadmium by difference. The ingot analyzed 50.8 pct Ag at the top of the ingot and 50.7 pct at the middle and at the bottom of the ingot. This alloy was only slightly hypo-eutectoid compared to the accepted value of 50.5 pct Ag and was used for the isothermal studies. It was felt that further attempts to adjust the composition might be futile because of the volatility of the cadmium. The ingot was sectioned to specimens approximately 3/4 x 3/4 x 1/8 in. These specimens were "austeni-tized" for 3 hr in air at 650°C to yield homogeneous ß and then quenched into a salt bath held at constant subeutectoid temperatures where the transformation

Jan 1, 1954

By P. T. Bail, J. E. Bobek

In reservoirs containing soft shales interbedded with oil-producing sands, it is possible to have water movement from the shales into the sands during reservoir depletion. This paper presents the results of scaled flow-model tests performed to investigate the efficiency of oil displacement by this vertical influx of water into oil-producing sands. In these model tests, the effects of rate of influx and rock wettability were considered. Oil recoveries by shale-wafer influx and by water flooding were cornpared after 0.4 hydrocarbon volume of water entered the producing sand. This comparison was made for water-wet and oil-wet systems. In water-wet sand, shale-water influx produces slightly less oil than conventional water flooding; in oil-wet sand, shale-water influx produces about one-half as much oil as conventional water flooding. In another type of comparison, it was found that shale-water influx followed by water flooding recovers less oil than water flooding alone. This is true for both water-wet and oil-wet sands. A comparison of tests run at different rates of shale-water influx shows that oil recovery is insensitive to rate in water-wet sand; in oil-wet sand, oil recovery decreases with decreasing rate of influx. INTRODUCTION Localized subsidence of the earth's surface has accompanied the depletion of some oil reservoirs, particularly in California,1, 2 along the Texas Gulf Coast,3 4 and in the Bolivar Coastal fields of Venezuela. Subsidence occurs when fluid withdrawals reduce the pressure in the reservoir. The reasons for subsidence accompanying a reduction in reservoir pressure are not known with certainty.10° However, there is considerable evidence that compaction of soft shale beds adjacent to producing oil sands can be a major cause of this subsi-dence.3, 6, 10-12 When the shales compact, a limited amount of water moves from the shales into the producing sands. A schematic illustration of this oil-producing mechanism is shown in Fig. 1. As the pressure in the sand declines due to oil production, a predominantly vertical pressure gradient develops at the sand-shale boundary. This causes water to move from the shale into the sand as indicated by the solid arrows. This paper reports the results of scaled flow-model tests to investigate the efficiency of oil displacement by the vertical influx of shale water into producing sands. The model tests were not performed with the aim of duplicating the recovery behavior of any specific reservoir. The primary purpose of the tests was to obtain a better understanding of oil displacement by shale-water influx. A knowledge of the oil- and water-production characteristics of even a simplified reservoir system should be of value to engineers estimating the performance of reservoirs where shale-water influx is occurring in conjunction with other production mechanisms. The following limitations and assumptions apply to the model work. 1. The sand of the prototype reservoir is homogeneous and horizontal. 2. The shale-water influx is uniformly distributed along the top and bottom of the sand member being studied. 3. No effects of edge-water drive or dissolved-gas drive that might accompany shale-water influx are considered. PROTOTYPE RESERVOIRS Scaled-model tests were performed to study oil displacement from three prototype reservoir sections. The reservoir sections modeled consist of single sand layers out of a sequence of -sand and -shale layers, as illustrated in Fig. 1. Each prototype represents a homogeneous, high-permeability, unconsolidated sand layer. In the field, the prototype is oriented in such a manner that one end is at the radius of drainage of a producing well and the other end is pointed toward the producing well (Fig. 2). The sand and fluid properties and the dimensions of the three prototypes are given in Table 1. The capillary pressure relationships for the three prototypes are presented in Figs. 3, 4 and 5. The oil-water per-

By Paul B. Crawford, Melvin B. Burton

A procedure is given which describes how the gelatin model may be-used to study the effect of mobility ratio on areas swept before and after breakthrough in fluid injection programs. Differences in mobility ratio are achieved through the proper selection of electrolytes for the invading and displaced fluids. The progress of the front is followed visually and photographed or traced onto plates (it desirrd positions before and after breakthrough. Example applications to standard flooding pattrrrls, vertically and horizontally fractured reservoirs and the in-situ combustion recovery process ore presentrtl. INTRODUCTION In water flooding, gas cycling and other fluid injection programs, the sweep efficiencies and swept areas after breakthrough are known to depend on the mobility ratio. Previous investigators have described the application of the X-ray technique, stepwise use of the poten-tiometric model and numerical procedures for studying these problems.1,3,4,6 Others have described the application of the electrolytic and gelatin models for studying sweep efficiencies at a mobility ratio near one.2,5,7,8,9 It is a purpose of this paper to describe a procedure for using the gelatin model to study swept areas at mobility ratios other than one. PROCEDURE The technique developed for gelatin model studies is composed of four phases: the construction of the model, the selection and preparation of electrolytes, the running procedure, and the recording of the swept areas. Each phase is discussed independently in order that the over-all procedure will not be confused. Construction of the Model Since some of the advantages of the gelatin model are its economy, and speed, it is desirable to have as small a model as is possihle without creating error through the inability to follow visually the progress of the colored front. For this work, it was found that a distance of 4 to 6 cm between salt bridges is convenient to use. Since the size is relatively small, it follows that the gelatin layer must be thin. For this study, 1/8 in. was used as the gelatin thickness in each case. The model was constructed on a flat 1/2-in. plastic (Lucite) base with boundaries constructed from 1/8-in. plastic strips. The well pattern to be studied was outlined with the 1/8-in. plastic strips on the base with narrow apertures (- - 2 mm) in the boundary to represent the wells. Behind each well opening, a rectangular area of some 8 to 12 sq cm was outlined by additional plastic strips. The plastic strips were fastened to the base by any suitable non-conducting cement so that the boundary system was liquid tight. The electrodes were constructed from sheet copper in a fashion that would allow easy removal. In this work, it was found that by simply forming a large U-shaped clip from a strip of the copper that would slip over the boundary was satisfactory and easily removed. The electrodes should be as wide as one of the sides of the rectangular area behind the well opening. If a fractured pattern is studied, fractures inside the well pattern may be represented by strips of copper of the desired shape mounted in some easily removable fashion. The reason that apertures (or salt bridges) were used for wells instead of metallic electrodes is that the plating of the electroactive ion at the cathode and the introduction of an electroactive ion at the anode has a detrimental effect on the swept area outline if these operations \rere permitted to be carried on within the well pattern itself. The anode may be used within the well pattern

Jan 1, 1957

By J. Lohrenz, R. J. Francis, G. C. Clark

Compositional material balance methods have been extended to solution gas-gas cap-water drive reservoirs. Partial or complete pressure maintenance as well as reservoir repressuring by gas and/or water injection may be considered. Increrments of reservoir withdrawals and injections equivalent to a given time increment are used as the basic iterative step to allow the extension. An example field study shows how the reservoir volumes. extent of the water drive, and relative permeability ratio function may be estimated. Knowing these, the methods can be used to predict reservoir behavior. INTRODUCTION The compositional material balance is similar to conventional material balances in many aspects. The assumption of a pressure-homogenous reservoir with no transients is common to all conventional material balances1-1 although numerical methods are different. The compositional material balance also contains this assumption. Inclusion of phase equilibria directly in reservoir calculations is the key difference and improvement in the compositional material balance. A continuous account of the compositions of the reservoir fluids and effluents is determined. Calculations of fluid phases and properties can be based on these compositions. Excepting fluid phases and properties determined in this manner, the compositional material balance uses procedures exactly analogous to conventional methods. Brinkman and Weinaug5 and Jacoby and Berry6 have given compositional material balances for simple solution gas drive reservoirs; subsequently Jacoby and Berry' extended their procedure to include pressure maintenance by gas injection at constant reservoir pressure. This paper extends compositional material balance techniques to a more general combination drive reservoir with various types of pressure maintenance operations. The procedure for this extended compositional material balance and an example of its field application are presented. PROCEDURE The numerical calculations of this compositional mate- rial balance are performed on a high-speed digital computer. The most comprehensive case the present compositional material balance program can handle is a solution gas-gas cap-water drive reservoir with (1) water production, (2) water influx, (3) separate or commingled production from the oil and gas zones, (4) any surface processing arrangement including LPG and gasoline production, and (5) gas injection into the gas and/or oil zones." This description of the procedure applies to this most comprehensive case. Appropriate deletions in the procedure can be made for more simple mechanisms. ASSUMPTIONS In common with conventional material balances, the compositional material balance makes the following assumptions: (1) the total reservoir is at a uniform pressure at a given time; (2) the volumetric ratio of gas to liquid withdrawal at reservoir conditions is directly proportional to their mobilities; and (3) instead of assuming a pressure-dependent total gas-oil equilibrium, the cornpbsitional material balance makes a sianificantly different and more rigorous assumption as follows. Thermodynamic equilibrium exists for each component in the liquid-gas mixtures in the reservoir and in surface processing. These three assumptions are fundamental in the compositional material balance procedure. Other assumptions involve arbitrary simplifications of the reservoir mechanisms. This compositional material balance uses the following simplifying assumptions: (1) no bulk mixing occurs between the gas and oil zone; (2) total mixing occurs between any injected gas and the contents of the injection zone; (3) no in situ hydrocarbon liquid is produced from the gas zone; (4) no reservoir gas is retained behind any net bottom-water influx; (5) the residual saturation of hydrocarbon liquid at reservoir conditions behind any net bottom-water influx is constant; and (6) fluids retained behind any net bottom-water influx exert no further influence on the reservoir. These six assumptions are not fundamental. Modifications of the present procedure can be made for different assumptions of the reservoir mechanisms. THE TIME INCREMENT Previous compositional material balances have used a pressure increment for simple solution gas drive reservoirs"" and an increment of injected gas for pressure maintenance.' These increments are suitable for these simple cases. They are, however, inadequate for reservoirs with time dependent drive mechanisms including water drive. A time increment removes this inadequacy and is

By I. Zaplatynskyj, G. C. Kuczynski

SINCE the excellent studies of metal powder compaction executed by Kamm, Steinberg, and Wulff, no work on the subject has appeared in the technical literature. Kamm et al. were the first to investigate quantitatively the distribution of density in a powder compact, as a function of pressure, height and diameter of the die, powder particle size, etc. For this purpose they devised a very ingenious method. They inserted a thin lead grid into a die and then filled the die with powder. During cold pressing, the lead grid -moved downward. Due to its deformability it followed closely the downward descent of the adjacent powder. The die was then radiographed and, from the deformations of the grid openings measured on the radiographs, the density in the vicinity of each opening was evaluated. Their findings led them to the conclusion that the friction on the side walls of the die is a major controlling factor of the density distribution inside the cold pressed powder compacts. However, it is difficult to conduct tests similar to those of Kamm, Steinberg, and Wulff in a processing plant laboratory, which usually lacks the necessary expensive radiographic equipment. The purpose of this note is to describe a method which makes it possible to determine the density distribution in green powder compacts with the same accuracy as the lead-grid method but which is far simpler and faster and requires only standard laboratory equipment. Instead of using a lead grid, the hardness at selected points of the sectioned compact is determined. If the relationship between hardness and density is known the density distribution is readily obtained. This relationship is easily found from measuring hardness and densities of thin compacts. In such compacts the density, for all practical purposes, is constant throughout. The method was tested on a Plast nickel powder, grade FlA-A20, kindly supplied by Plast Metals Co. Thin compacts (height to diameter ratio 0.2) were pressed in a single action cylindrica1 die at various pressures. The density of the compacts was determined by weighing and by measuring the height and diameter with a micrometer. Subsequently the Rh hardness was determined. Up to ten readings were taken on each specimen. The difference in hardness readings on each compact was never greater than three corresponding to about 0.1 g per cu cm uncertainty in density. By varying the compaction pressure of the samples, a whole scale of densities was obtained allowing construction of a set of density-hardness calibration curves. To test the method several compacts of different height were pressed under varying pressures. Die walls were always well lubricated. After pressing, the compacts were mounted in bakelite and halved by grinding, and hardness readings taken along the diameters of the specimens. The distance between adjacent indentations was about 2.5 mm. After hardness readings were translated into densities, contour maps of densities of the compacts were prepared. One of them is reproduced in Fig. 1. The areas encompassed by the curves are labeled with numbers representing the density in g per cu cm of those regions. In Fig. 2 the same results are re-plotted to emphasize the radial density changes at various depths. The results thus obtained confirm the findings of Kamm, Steinberg, and Wulff kentioned previously. They are recapitulated here briefly. 1—The densest part of a compact pressed from one side is at the outer circumference at the top and the least dense part at the bottom, at the outer circumference. Figs. 1 and 2 show this clearly. 2—The density near the cylindrical surfaces of the compact decreases uniformly with height from top to bottom. 3—The density variation increases with the height to diameter ratio. References R. Kamm, M. A. Steinberg, and J. Wulff: Trans. AIME (1941) 171,'R. p. 439 ;MetaLs Technology (February 1941). R.p. Kamm, M. A. Steinberg, and J. Wulff: Trans. AlME (1949) 180, 2R.P. 694; MetalS Technology (December 1948).

Jan 1, 1957

By B. Ancker, E. R. Parker

The small-angle dislocation-boundary density of nickel and some of its alloys was investigated as a function of strength. It was found that the strength is a linear function of the density for pure nickel and for Ni-Ti alloys, but not for Ni-Co alloys. The results are interpreted in terms of dislocation theory. THE processes by which metals deform are still poorly understood, even though there has been a concentrated research effort in this field during the past fifteen years. Single crystals and individual grains in polycrystalline aggregates often develop a subgrain structure when deformed. Substructure formation occurs as a result of local plastic bending and its development is augmented by annealing at moderate temperatures.' Investigations made to date, largely qualitative in nature, have shown that the substructure is composed of a network of small-angle boundaries. The purpose of the present paper is to present some quantitative data on substructure (obtained by X-ray techniques) and to correlate the results with the strength of the metal in which the substructure existed. Nickel and some of its alloys were used for the experiments. The presence of certain kinds of foreign atoms greatly alters the formation and effect of the substructure. Polycrystalline tensile specimens were prepared from high purity nickel, two Ni-Ti alloys, and two Ni-Co alloys; the composition of these materials is given in Table I. Tensile specimens having a reduced section 1/4 in. diam by 3 in. long were machined from materials which were refined, cast, forged, and swaged in the authors' laboratories. The specimens were annealed for 1/2 hr at 1150°C, a treatment which produced grains having a diameter of approximately 0.5 mm. This rather large grain size was used to facilitate later examination of the substructure. Duplicate specimens of pure nickel and of each alloy were strained various controlled amounts in tension at room temperature and recovered for 1 hr at 800°C. (This strain induced prior to recovery will be referred to hereafter as the "prestrain.") One of each of the recovered specimens was pulled in tension at room temperature to obtain the stress-strain curves. The other bar was cross-sectioned in the gage length, carefully polished, and deeply etched to remove all cold-worked material. This specimen was then examined by the X-ray microscopic technique2 to determine the nature and magnitude of the substructure. The use of duplicate specimens permitted a direct correlation between tensile properties and substructure characteristics. X-ray microscopic studies were made with an apparatus constructed to have an average resolving power of about 4 microns. The monochromatic radiation reflected from grains which were oriented favorably was recorded on fine-grain photographic plates. These were examined at Xl00 which represents the true magnification, since there was essentially none in the X-ray system. This permitted detection of angular differences between subgrains of as little as 10 min. A very large number (over 100) of grains for each specimen was examined at this magnification and the number of substructure boundaries per grain was tabulated. An accurate relative measure was obtained under fixed experimental conditions, e.g., same grain size, resolving power, magnification, etc. The greatest source of error lay in counting the subboundaries, but this proved to be surprisingly small when repeated counts were made by different investigators. Furthermore, as the subsequent analysis indicates, whatever human error was introduced was not significant. Analysis of the Substructure Data When the number of substructure boundaries per grain was plotted against the frequency, skew (asymmetrical) distribution curves of the type shown in Fig. 1 were obtained. These are smoothed curves for the 1 atomic pct Ti in nickel alloy subjected to various prestrains and are representative of the curves obtained with all the metals. Since skew-distribution curves cannot be analyzed by elementary statistical methods, detailed investigations were undertaken in order to 1—test the consistency

Jan 1, 1955

By C. C. Harris

A critical literature review leads to a descriptive model of tumbling mill operation based upon energy partition concepts and the necessary requirements for particle fracture. As grinding proceeds into the ultra-fine region, conditions which are of little significance during normal operations gradually become controlling. These involve increasing resistance to fracture and an increasing tendency to aggregate as particles become smaller, and a growing fraction of the power input being consumed fruitlessly. The result is that fineness asymptotes to a maximum value as comminution proceeds indefinitely; thus where t is time, p and n are experimentally determined parameters, and Ø is a measure of fineness, which is presumed to be proportional to specific surface area, or proportional to the reciprocal of the size modulus (provided that the distribution modulus remains constant with time), or both. Ø, is the asymptotic value of Ø and is a measure of the grind limit. A method for its graphical determination is given. In a previous paper1 the physical and mathematical principles requisite to a study of the role of energy in comminution were reviewed. The micro-process of comminution — or the comminution event — can be investigated in terms of (a) the magnitude of the force-field localized around the individual particles which gives rise to the stresses generated in the particles, which in turn induces their fracture, and (b) the distribution of fragment sizes, should the event prove fruitful. The entire fragmentation operation can be considered to be the totality of its individual events, and in order to obtain a single independent variable of the process it is necessary to transform and sum the local stresses in terms of energy.1 This energy, however, will account for only a fraction of the energy input to the comminution process, the actual magnitude depending upon the kind of mill and several related factors. From the comminution viewpoint the primary function of a mill is that of stressing as many as possible of the individual particles of the charge to failure, with the maximum economy of energy expenditure. This is difficult indeed to achieve, for stressing a particle does not always break it, while the overwhelming portion of the energy input is involved with various internal mill processes which — almost incidentally — determine the magnitude, frequency, and manner in which the forces are applied to the particles. Essentially, these processes depend upon the mill dimensions, loading and speed, and they may be little affected whether fracture occurs or not. These processes must be delineated if a complete description of the role of energy in comminution is to be given; the energy account sheet1 (Fig. 1) in showing how the energy is partitioned for the requirements of the several functions it performs, provides a framework for the development of a model of the comminution process. Actual numerical values for the variables appearing in Fig. 1 will depend upon the mill, materials, and operating conditions. A possible general relationship between the variables will be suggested in this paper. Before proceeding further, a cautionary note must be issued. It was pointed out in an earlier paper' that while energy input is not changed by increasing the power and decreasing the time in the same proportion, and vice versa, the results of the process in terms of particle fineness may be different. Thus, although energy is chosen as the independent variable of a comminution process, it is not absolutely independent, and care must be exercised to ensure that it is used within its range of valid application. Time has the necessary independence, and it will be used in place of energy where it is appropriate. The residence time in a mill is usually of the order of minutes. Under these conditions the particles are subjected to "... a relatively constant fracture producing environment".2 However, under the more rigorous conditions of greatly extended time, which pertain in ultra-fine grinding, indications are that the fracture producing environment no longer remains constant; this is the major subject of this paper. At this extreme limit the mill is probably taxed to its utmost, and at the same time the bulk material properties are becoming closer to ideal, while the physico chemical effects associated with surfaces, edges and corners are multiplying. (The edge length per unit volume is proportional to the square of the specific

Jan 1, 1968

By P. W. Rainier

THE Chivor emerald field is situated on the eastern slope of the Andes in the Department of Boyacá, at an elevation of about 8000 ft. above sea level. It overlooks the Llanos (plains) of the Orinoco and is watered by many streams flowing into the Rio Meta, the Orinoco's largest tributary. The Rio Meta is navigable from the Orinoco to a point within about 50 miles of the Chivor district and may eventually prove to be the reasonable outlet for the future development of this part of the country. The topography of the district is extremely rugged and mountainous (Fig. 1), the difference in elevation between two points on a mining claim 1 km. square being sometimes as much as 4000 ft. The ground is covered with dense forest growth and the available timber is suitable and sufficient for mining purposes, but it is becoming gradually depleted by Indian clearings as the country around the mines becomes settled. The many streams combined with the rough topography give good possi-bilities for hydroelectric power. The center of population nearest to Chivor is Bogota, the capital of the republic and a city of some 300,000 inhabitants; the distance is two and one-half days' travel. The first stage of the journey from Bogota, is accomplished by rail to Chocontá, some 4 hr. travel. From Chocontá the journey is made in the saddle, one night being spent in Guateque. The latter is a small town and the nearest post office to the mines, being some 10 hr. ride distant. The roads are mere bridle paths, in parts hazardous and almost impassable in the wet season. All supplies brought to the mines, other than those produced locally, are carried by pack mule; mule freight from Chocontá costs 3 centavos1 per pound. The climate of the Chivor district is healthful and bracing. The wet season begins in March and concludes in November, the remaining months being dry. The annual precipitation is about 150 in. The temperature throughout the year is nearly constant, varying from about 45° F. to about 65° F. shade temperature. The rains are cold and unpleasant, and for the greater part of each day in the wet season the country is shrouded in dense mist. A fire can with comfort be kept burning in a living room all the year round. There are a few slight

Jan 1, 1929

BARBOUR, PERCY E., Deputy Supt., (Captain), New York State Troopers; Captain, 22nd Regiment, N. Y. N. G. BARLING, 1i. B., 1st Lieutenant, Engineer Officers' Reserve Corps. CHAPMAN, R. H., Major, Engineer Officers' Reserve Corps. DUNN, THEODORE S., 1st Lieutenant, Engineer Officers' Reserve Corps. DWIGHT; ARTHUR S., Major, Engineer Officers' Reserve Corps. FORBES, CARROLL R., Captain, Engineer Officers' Reserve Corps. GOODSPEED,. G. E., JR., 1st Lieutenant, Engineer Officers' Reserve Corps.

Jan 7, 1917

The following is a report of the papers submitted to the Papers 'and Publications Committee during 1928: The Committee held nine regular meetings during the year 1928. During that period there were presented to the Institute 214 papers which were acted upon as follows: Accepted as Technical Pub1ications............................ 85 Declined for publication...................................... 16 Action pending, Dec. 31...................................... 25 Manuscripts held for annual volumes, Dec. 31................... 7 Withdrawn temporarily...................................... 1 Accepted as Preprints........................................ 20 Printed for first time in volumes.............................. 35 Published in Mining and Metallurgy........................ 12 Papers in Technical Publication No. 156 (Coal Classification) and not accounted for elsewhere in this table..................... 10 Mimeographed only*................................... 3 Total.................................................. 214 * A considerable number of papers were mimeographed for the annual meeting. Of the 85 papers accepted as Technical Publications 8 were reprinted in the Transactions, 7 in Proceedings of the Institute of Metals Division (1928), 7 in Iron and Steel Technology in 1928, 13 in Geophysical Prospecting, and 50 were in Technical Publication form on Dec. 31. Of the 20 papers accepted as Preprints 1 was reprinted in the Transactions, 13 in Petroleum Development and Technology in 1927, 2 in Geophysical Prospecting, and 4 were in Preprint form on Dec. 31. Of the 35 papers printed for the first time in the various volumes 8 were included in the Transactions, 2 in Proceedings of the Institute of Metals Division (1928), 1 in Iron and Steel Technology in 1928, and 24 in Petroleum Development and Technology in 1927. Of the 12 papers accepted for Mining and Metallurgy 1 was, accepted for Geophysical Prospecting and 1 for the Transactions. There were published in 1928 113 Technical Publications containing a total of 1928 pages, as against 88 in 1927, totaling 1528 pages. The Technical Publications published in 1928 were classified as follows: Group A, Metal Mining, 17 papers (361 pages); B, Milling and Concentration, 4 papers (39 pages); C, Iron and Steel, 15 papers (256 pages); D, Non-ferrous Metallurgy, 13 papers (191 pages); E, Institute of Metals, 36 papers (588 pages); F, Coal and Coke, 15 papers (392 pages); G, Petroleum and Gas, 10 papers (246 pages); H, Non-metallic Minerals, 3 papers (33 pages); I, Mining Geology, 18 papers (356 pages); K, Mining Administration, etc., 3 papers (47 pages); L, Geophysical Prospecting, 21 papers (430 pages).

Jan 1, 1929

By M. N. Pathasarathi, P. A. Beck, T. J. Koppenaal

EARLIER work1 indicated that in rolled and annealed copper the volume fraction of the cube-texture component may increase on continued isothermal annealing. Merlini found2 that in rolled copper the well-known increase in the amount of cube-oriented material with increasing annealing temperature3 is concurrent with grain growth, which takes place after recrystallization is complete. It was found that both in copper2 and in aluminum4 the texture resulting from primary recrystallization, even if this is followed by some grain growth, may in general comprise, in addition to the cube-texture component, four symmetrical texture components of the general type (123)[412], i.e., components not very far from the ideal orientations of the rolling texture. The volume fraction of the cube-texture component increases on further annealing at the expense of these four components.' The present work was undertaken to study quantitatively the increase in the volume fraction of the cube-texture component in the course of normal grain growth, under varying conditions. Since the grain-gr3wth behavior of A1-Mn alloys has been already investigated in some detaiL5 an A1 + 0.8 pct Mn alloy was chosen for this work, in which the extent of normal grain growth can be effectively controlled by the inhibiting effect of a dispersed second phase. EXPERIMENTAL PROCEDURE High-purity aluminum of the following composition was used: A1 99.997 pct, Cu 0.0004 pct, Fe 0.0005 pct, and Si 0.001 pct. A high-purity A1-Mn master alloy was prepared by dehydrating manganese chloride at 500°C and reacting it with molten high-purity aluminum at 750°C in a graphite crucible.5 The Mn metal reduced from the salt by the molten aluminum goes into solution in the latter. The master alloy? containing 13 pct Mn, was solidified under conditions suitable for minimizing segregation. An ingot of the final alloy, to be used in the experiments, was per pared by melting high-purity aluminum and the necessary amount of master alloy in a graphite crucible, and by pouring the molten alloy into a graphite mold preheated to 700°C through a box containing graphite baffles designed to intercept most of the aluminum oxide films floating in the melt. The melt was then solidified by placing the mold on a water-cooled copper plate and by keeping the top of the mold hot, in order to minimize piping. The 6 1/2 in. long, 11/4-in.diam ingot consisted largely of a single crystal. Chemical analysis showed that there was little segregation from the bottom (0.83 pct Mn) to the top (0.81 pct Mn). Spectrographic analysis of the ingot showed that the impurity content remained essentially the same as that of the aluminum used. The ingot was processed so as to obtain a uniformly small grain size. without banding. The ini-tial steps of rolling in the form of a flat bar and of annealing, down through a thickness of 0.4 in., are shown in Table I. At this point the bar was cut into four sections, each approximately 3 in. long, which were processed separately as shown in Table If. The table also gives the penultimate grain sizes obtained in each section by annealing prior to the final rolling of 90 pct RA. During the final rolling the strips were reversed end to end between passes. Annealing treatments were given in a salt pot furnace, with the temperature controlled to 1°C. All annealing of section "6s" was carried out above the solvus point to avoid the formation of second-phase precipitate. The penultimate annealing was sufficiently short to give a relatively fine grain size in spite of the high temperature. Section "6L" was processed in a similar manner, except that the penultimate annealing was much longer, so as to allow considerable grain growth. Section "4" was given the two final

Jan 1, 1961

By E. E. Underwood, L. L. Marsh

Aluminum single crystals, alloyed with 0.042 atomic pet Cu and 0.11 and 1.1 atomic pct Mg, were subjected to constant stress creep tests, tensile tests, and hot hardness measurements within a temperature range of 300° to 866OK. Calculations based on Dorn's temperature-compensated time parameter, 6, gave a value of DH, = 27,000 cal per mol for the activation energy of early creep in aluminum single crystals. Correlations have been obtained for aluminum alloy single crystals with the parameter E for solid solution strengthening, as well as with the parameter F for solid solution hardening, by using a valence of three for aluminum. Limited measurements on tensile specimens show that the slip band density tends to decrease with increasing temperature and with decreasing solute concentration. INCREASING interest is being shown in the mechanisms of plastic deformation in single crystals during tensile and creep testing. The complexity of deformational processes in polycrystalline materials has led to a search for simpler experimental conditions. Numerous creep and tensile investigations have been conducted with pure, metallic single crystals. To a lesser degree, the effects of alloying elements on the tensile properties of single crystals have been determined. However: the literature dealing with the effects of alloying additions in single crystals under creep conditions is vanishingly small. This paper represents an attempt to narrow this gap in the knowledge of the subject. Previous investigations of creep behavior at the Battelle Memorial Institute1-1 have been of great value in the analysis of the present single crystal data. It is equally desirable, however, to ascertain the extent of correspondence between the behavior of single crystal and polycrystalline materials. For this purpose, correlations, similar to those developed for polycrystalline aluminum alloys by Dorn and co-workers, have been made with the single crystal data. The results from this study have tended to confirm and extend those correlations to the case of alloyed single crystals. Materials and Procedures Three dilute binary aluminum alloys were prepared for this investigation from 99.99 + wt pct Al, 99.8 wt pct Mg, and 99.92 wt pct (electrolytic) Cu. The nominal compositions of the alloys were 0.042 atomic pct Cu, and 0.11 and 1.1 atomic pet Mg. Precautions were taken to avoid contamination of the stock during melting and casting of the alloys. After a heat treatment of 8 hr at 925°F the alloys were extruded, then machined into) threaded tensile specimens with a 3 in. reduced section and a 0.505 in. diam. Spectrographic examination showed less than 47 ppm metallic impurities in each alloy. The single crystals were grown by the strain-anneal method, with a critical strain of about 11/4 pet giving the optimum results. In general, 3 in. crystals were obtained with the Al-Cu alloy, but smaller crystals in the magnesium alloys necessitated the use of 2 in. and 1 in. gage lengths with the low and high magnesium alloy specimens, respectively. After an electrolytic polish, the orientation of that portion of the specimen containing the largest single crystal was determined from Laue back-reflection photographs. Tensile tests were conducted at a constant load rate of about 2 lb per min. Creep runs were made in a constant temperature room, under constant stress at the higher creep temperatures, and constant load at the lower temperatures. The eloneation was measured to within ±5 microin. by a specially designed capacitance extensometer. The ex-tensometer arms were attached to the 3 in. specimens at the shoulders of the test piece or, where the crystals were smaller than 3 in., by knife-edge grips.

Jan 1, 1957

By D. L. Luffel

This paper presents two main developments: (1) a short, constantly accelerated centrifuge method of measuring drainage capillary pressure properties of cores, and (2) a new analytic method of treatment of centrifuge measurements to obtain saturation-capillary pressure relations. Toward the first objective, a clever experimental method is presented which is certainly more rapid than previous centrifuge methods. Toward the second objective, a new analytic method is derived to eliminate the approximate method of Hassler and runner. The derivation of this new analytic method, however, appears to be in error. The new analytic solution is developed as Eq. 5 in Appendix A. At the beginning of this derivation, the average wetting phase saturation of a core is expressed as, where is the centrifuge angular velocity. Eq. D1.2 makes explicit the dependence of S and S upon The second equation shown in Appendix A expresses the relation between the three variables Since for a given porous medium the saturation is uniquely related to capillary pressure, the functional dependence of S upon P is trivial, i.e., Pc.ri does not actually appear. Thus, the function 5 (r, ) may be expressed as a new function, S*, of PC, , alone The author transforms the variable of integration in Eq. Dl through use of Eq. D2 and writes

Jan 1, 1965

By E. A. Barnard

ARSENIC is a very old substance. The ancients speak of it in their writings, and its use has developed very little until recent years. The ancients used it in making pigments, in medicine, and for poison. They also tell of the difficulties of handling it-an experience which is very similar to our own. Until about 10 years ago, arsenic was very cheap and was a drug on the market. About that time the boll weevil began making serious inroads on the American cotton crops, and after much investigation the Depart-ment of Agriculture developed a very successful poison for this pest, consisting of calcium arsenate. The weevil pest has continued, and as the cotton planters have become convinced that calcium arsenate is the best material to use against the weevil, the demands for it have been steadily increasing, resulting not only in a rising price of arsenic but in a very serious shortage so far as our country is concerned. So serious has this shortage become, and so vital is it to the cotton growing industry, that Congress has appointed a committee to investigate and report on the situation. A meeting of arsenic producers and consumers was held in New York last December to take stock of the situation and to see what could be done. The extreme increase in, the market price of arsenic-from 6 c. to 15 c. per lb. in less than a year-has been very harmful to the producers as well as to the consumers. Some very interesting' data were brought out at this meeting, as follows: CONSUMPTION PER YEAR, SHORT TONS Arsenate of lead 2,000 Paris green 1,200 Proprietary mixtures 750 Sheep dips : 1,000 Weed killers 1,500 Glass : 2,000 8,450 1922 Consumption as calcium arsenate 4,000 12,450 1923 Estimated consumption as calcium arsenate... 12,000 Estimated 1923 consumption 20,450

Jan 8, 1923

By P. de Vitry, Oliver Bowles

THE conventional wire saw, introduced in the slate district of Pennsylvania by the Bureau of Mines in 1927, and used thereafter with remarkable success, consists of a three-strand steel cable having a diameter of 3/16 or 1/4 in. The equipment, methods used, and accomplishments have been described in previous reports.1-7 Similar wire saws have been used extensively in Europe, in quarries and in shaping blocks and slabs. During the period 1932-1937, competition in European marble quarries was keen, prices were declining, and orders were not abundant. When orders were received, there was generally an urgent call for speedy delivery. Because of the diversification of demands, it was impossible to maintain in stock marbles of the color, size, and thickness called for, therefore the material generally had to be fabricated after the order was received. The wire saw enabled stone mills to fill such orders promptly. With skillful handling, 7/8-in. slabs could be cut accurately and rapidly with this equipment. It was found, however, that, as the three-strand wire became worn, the cutting speed dropped rapidly and the cut became less accurate. To overcome these difficulties the junior author of this paper, after much experimentation, developed a new type of wire consisting of a single ribbon-shaped strand, twisted as shown in Fig. 1. To ensure straight sawing, the twist of the ribbon is reversed from right to left and vice versa every 25 ft. The reversal of the [ ] twist is shown in Fig. 2. This is not a new feature; it has been used to advantage with the three-strand wire both in Europe and America for several years. However, some difficulties arise in splicing the three-strand reverse-twist wire, and the claim is made that cutting is less accurate as the wire becomes worn. To make an endless belt, the ends of the single-strand wire are brazed together with

Jan 1, 1941

By Arthur F. Daily

8.5-1. Bucket Line Dredge. INTRODUCTION. The modern bucket-line dredge for mining placer deposits is a combined floating excavating machine and a gravity concentration system. Material is excavated at the forward end by the buckets, elevated, washed and screened, concentrated for recovery of valuable mineral, and discharged off the stern as coarse and small fractions. The end product is a concentrate consisting either of the valuable mineral(s) and associated heavy minerals, which are taken to a treatment plant on shore; or in the case of gold, a clean amalgam ready for retorting. The bucket-line dredge may be adapted for purposes other than mining for valuable mineral by installing, in lieu of a concentrating system, machinery for pumping all of the solids ashore in a pulp, or for producing washed gravel in selected size fractions of which all or a selected fraction is pumped to disposal areas on shore or into side barges. The two principal types of bucket-line dredges designed for mining are: 1) those having an inclined belt conveyor stacker to discharge the coarse fraction of washed tailing well astern of the hull, and sluices to discharge the sand fraction (Fig. 8.5-1), and 2) those having only sluices for discharge of all tailing astern. Use of the latter type has been confined, with few exceptions, to mining cassiterite-bearing deposits in Malaysia and at sea off Indonesia. Small dredges having only flumes for washing and concentrating all material dug by the buckets have been used for mining shallow stream valley deposits. GENERAL ARRANGEMENT. The general arrangement of a modern bucket-line dredge for mining tin, with a primary concentrating plant on board, is shown in Fig. 8.5-2. The hull and superstructure of most dredges are fabricated of structural steel; wood also has been used for both small and large dredges with complete satisfaction. All-riveted or partially welded unit hulls have been widely used and are desirable for exceptionally severe service. The hull

Jan 1, 1968

By A. B. Kinzel

MUCH work has been done recently in an attempt to analyze the physicochemical mechanism involved in the production of steel by the open-hearth process. . This has resulted in reducing the process to a series of interrelated phenomena obeying the usual laws of mass action and diffusion. Of these, the reaction of the iron oxide and carbon in the molten steel is of most interest from a technical point of view as it con-stitutes the main action in the usual steelmaking processes. A. L. Feild1 has formulated a basic differential equation expressing the open-hearth reactions and Hefty has done considerable experimental work on the determination of a number of constants involved in the basic open-hearth reaction. In considering the application of the law of mass action. to the reaction FeO +C ?? CO +Fe, the iron will be taken as constant, as is the case in the open hearth and the carbon monoxide pressure will be assumed' constant at one atmosphere. From this xy = m at equilibrium, where x is the percentage weight of carbon, y percentage weight of iron oxide and m the weight-concentration equilibrium constant. While it is customary for physical chemists to express equilibrium constants in terms of molar concentrations, the constant is here expressed as the product of the percentage-weight concentrations for ease of comparison with the constants quoted by Herty and by Feild, and for more ready applicability to the open-hearth reactions. Herty and his associates have attempted to calculate the value of this constant from data on open-hearth heats. So many variables affect these results that an adequate calculation presents very great difficulties. A value of 4 X 10-2 was first arrived at, but this was later changed to 1.6 X 10-2. Feild, in order to illustrate the operation of the differential equation of the open hearth, has assumed a value considerably lower than Herty's, namely, 1.0 X 10-2, the largest value consistent with certain theoretical

Jan 1, 1929