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By J. O. Lewis, Carl H. Beal

Introduction ............................... 492 Theories Relating to the Recovery of Oil.................. 493 Methods Commonly Employed for the Estimation of Future Production..... 495 Saturation Method........................... 495 Production Curve Method....................... 496 Production per Acre Method.....................: 496 Appraisal Curves.............................. 498 Derivation and Construction...................... 498 Comparison of Appraisal Curves of the Osage and Nowata Fields...... 505 Use of Appraisal Curves........................ 506 Accuracy of the Apprais?l Curves and Their Sources of Error........ 508 Relation Between Future Productions of Wells of Equal Output....... 510 Use of Appraisal Curves for the Derivation of Generalized Future Decline Curves. 513 Theory of Generalized Curves Showing the Decline in Production...... 513 Suggested Methods for Making Closer Estimates.............. 517 Use of Logarithmic Coordinate Paper.................... 518 Summary................................ 520 Introduction Oil wells usually reach their maximum daily output shortly after they are completed. From that time they decline in production, the rapidity of decline depending on the output of the wells and on other factork governing their productivity. The production curve of a well shows the amount of oil produced per unit of time for several consecutive periods; if the conditions affecting the rate of production are not changed by outside influences, the curve will be fairly regular, and, if projected, will furnish useful knowledge as to the future production of the well. By the aid of this knowledge the value of a property maybe judged, and proper depletion and depreciation charges may be made on' the books of the

Jan 1, 1918

By H. R. Cooper, T. S. Mika

T. S. Mika (Department of Mineral Technology, University of California, Berkeley, Calif.) - Dr. Cooper's attempt to establish a correlation between process behavior and operational variables on the basis of a statistical analysis after imposing a reasonable process model is a very commendable improvement on the use of standard regression techniques. However, it must be recognized that the imposition of a model has the potential of yielding a poorer representation if its basic assumptions or mathematical formulation are invalid. It appears that at least two aspects of his treatment require some comment. First, the limitations on the kinetic law where xta represents a hypothetical terminal floatable solids concentration (cf. Bushell1), should be mentioned. Most current investigations2-9 appear to utilize the concept of a distribution of rate constants rather than a single unique value, k, to describe flotation kinetics. A distributed rate constant is certainly a more physically meaningful concept than that of a terminal concentration. The study of Jowett and safvi10 strongly indicates that xta is merely an empirical parameter, whose actual behavior does not correspond to that expected from a true terminal concentration. Rather than being a strictly mineralogical variable, as Dr. Cooper's treatment implies, it apparently represents the hydromechanical nature of the test cell as well as the flotation chemistry. The extension of batch cell kinetic results to full-scale continuous cell operation is a suspect procedure if the effect of such nonmineralogical influences on x,, remain unevaluated. There is evidence that introduction of a terminal concentration is necessitated by the inherent errors which arise in batch testing and are eliminated by continuous testing methods.' Possible lack of validity of the author's use of Eq. 1 is indicated by two unexpected results of the statistical analysis of his batch data. The first is the apparent corroboration of the assumption that the rate constant, k, is independent of particle size, i.e., of changes in the size distribution of floatable material. This assumption directly contradicts numerous results 2,4,11-l8 for cases where first order kinetics prevailed and ignores the phenomenological basis for the analysis of flotation in terms of a distribution of k's. It must be recognized that, if the rate constant is size dependent, the lumped over-all k would be time dependent; Eq. 1 would then no longer be valid. Cooper's x,, is determined by batch flotation of a distribution of sizes for an arbitrary period of time. If the size dependence of k is artificially suppressed, x,, will become a function of the experimental flotation time used in its determination. Upon reviewing the rather extensive literature concerning batch flotation kinetics, there appear to be few instances where constant k and x,, adequately adsorb variations in floatability due to particle size. The second surprising result is the low values of the distribution modulus, n, determined. Contrary to Cooper's assertion, most batch grinding (ball or rod mill) products yield values of n > 0.6, which increase as the material becomes harder.'' It is likely that the values of n = 0.25 and n = 0.42 for Trials 1 and 2, respectively, are completely unreasonable, and even the value n = 0.54 obtained for Trial 3 is unexpectedly low. Possibly, this indicates inherent flaws in the three trial models considered, in particular the assumed particle size independence of the rate constant, k. The above does not necessitate that Eq. 1 (and the terminal concentration concept) is invalid; it could constitute a good first approximation. However, the qualitative arguments used by Dr. Cooper in its justification are somewhat frail and require verification, particularly since much of the flotation kinetics literature is in opposition. Apparently, no effort was made to test these hypotheses on the actual data; in fact, since they pertain to a single batch test time, his data cannot be utilized to evaluate the kinetics of flotation. To evolve a control algorithm on the basis of this infirm foundation seems a questionable procedure. Another difficulty in his analysis arises in consideration of the froth concentrating process. As Bushel1 ' notes, for Eq. 1 to be valid it is necessary that the rate of recycle from the froth be directly proportional (independent of particle size) to the rate of flotation transport from the pulp to the froth, a restrictive condition." Harris suggests that it is more realistic to assume that depletion occurs in proportion to the amount of floatable material in the pertinent froth phase volume (treating that volume as perfectly mixed).12,21,22 The physical implications of

Jan 1, 1968

By I. Fatt, D. N. Saraf

A method is described for measuring two- and three-phase relative permeabilities in sandstones or sand packs using a nuclear magnetic resonance (NMR) technique to determine fluid saturations. Two- and three-phase relative permeabilities have been determined on Boise sandstone using the NMR technique of saturation measurement. Three-phase relative permeability to water was found to depend only on the water saturation, whereas three-phase permeability to oil depended on both the water and oil saturations. Relative permeability to gas in three-phase flow was found to depend only on the total liquid saturation. INTRODUCTION Three-phase relative permeabilities are extremely useful in calculating field performance for reservoirs being produced by simultaneous water and gas drives. Three-phase relative permeability data are also needed for analyzing solution gas-drive reservoirs which are partially depleted and are being produced by water drive. Some thermal recovery processes involve three-phase flow which require three-phase relative permeability data for predicting reservoir behavior. Unfortunately, three-phase relative permeability measurements have rarely been made. Also, because of the scarcity of three-phase data, it has not been possible to date to relate other measured rock characteristics to the relative permeabilities with a great certainty. Leverett and Lewis,l Reid2 and Snell3 have reported three-phase relative permeability data on unconsolidated sands. Leverett and Lewis used ring electrodes spaced along the length of the sand sample to measure the resistivity of the sample which was assumed to be monotonically related to brine saturation. Gas saturation was determined from pressure-volume measurements. Oil saturation was obtained by material balance on the cell containing the sand sample. This method is involved and time consuming. Another difficulty arises from the fact that the resistivity of the sand is a function not only of saturation of brine but also of the distribution and saturation history of the brine in the pore spaces. Reid2 used a gamma ray absorption technique for measuring liquid saturation. This method has the disadvantage that total liquid saturation rather than oil or brine saturation is all that can be measured and still another method is required to determine the saturations of individual components. snel13 used a neutron bombardment method which also required a separate determination of the individual component saturations. Caudle et al.4 measured three-phase relative permeability on consolidated sandstones using vacuum distillation for determining fluid saturations. Distillation after each reading makes this technique very lengthy and time consuming. Corey et a1. 5 and Naar and Wygal 6 measured three-phase relative permeability on sandstones by the capillary-pressure method. Semipermeable diaphragm assemblies were used at each end of the core specimen to keep the water phase in the core. Gravimetric methods were used to determine fluid saturations. Sarem7 recently reported an unsteady-state method for measuring three-phase relative permeability on sandstones. This method is an extension of Welge's method8 for measuring two-phase relative permeability. Although Sarem's method is simple and comparatively fast, the assumptions involved may oversimplify the problem. Sarem's assumption, that in all rocks relative permeability to each fluid will depend only on the saturation of that fluid, seems to be rather unrealistic. Neglecting capillary effects at the end of the core is also a weak assumption. Donaldson and dean9 measured three-phase relative permeability using a method similar to Sarem's.

By R. S. Shoemaker, D. L. Harris

PST investigations into the possibility of concentrating the low-grade iron ores of the Gogebic Range in Wisconsin have been hampered by the complex association of the constituent minerals. In part the problem arises from the fact that the iron occurs essentially as two minerals, hematite and goethite, having nearly the same chemical composition but exhibiting different physical properties. Individual grains of these minerals are usually so small that the ore must be ground to a size finer than 325 mesh if substantial liberation of the minerals is to be secured. Appraisal of beneficiation tests on this iron ore requires, therefore, some method of determining quantitatively the contents of constituent minerals in the various products of the beneficiation. Because of the chemical similarity of the two iron minerals, chemical analyses do not provide a means of differentiation. Although the minerals are different physically, they are both opaque and friable, and as the ore must be ground to a very fine size for liberation to be secured, quantitative microscopic methods of analysis are extremely difficult at best. In an effort to overcome these analytical problems, a quantitative X-ray diffraction analysis method has been developed at the University of Wisconsin for the research program on iron ore beneficiation. The X-ray diffraction analysis of chemical compounds of crystalline form has proved of inestimable value for structure studies and rapid qualitative analytical purposes in a number of fields.' In some cases X-ray diffraction analysis has been found suitable for quantitative determinations. H. P. Klug2 found that quantitative procedures for quartz powders when mixed with calcium fluoride were reproducible within ±1.0 pct. Talvenheimo and White3 reported quantitative analyses of clay minerals to be accurate within 10 pct and also detected 1 pct of bentonite in 99 pct illite. In other investigations,'" powders of quartz and mica have been analyzed quantitatively with accuracies from 1 to 5 pct. A powder of tungsten carbide was analyzed similarly by Redmond7 and Rooksby8 reported the determination of small amounts (0.11 pct) of calcium oxide in magnesium oxide. In the most generally practiced use of X-ray diffraction, traces of the diffracted beam are recorded as lines on a film much as lines of elements are recorded on an optical spectrograph film. As with the spectrograph, the darkness of the lines denotes the amount of the constituent in the sample. If the film is replaced by a Geiger tube assembly and a graphic recorder or a scaling unit, the intensity of the diffracted beam can be registered as a peak on a chart or energy impulses per unit of time on the scaler. Several methods of quantitative measurement have been investigated and reported in literature. Klug2 used manual operation of the goniometer and a counting technique to determine the difference in counting rates at the diffracted angle and for the background. In another case3 the area under the recorded diffracted peak was the basis for the determination. In other instances the height of the recorded diffraction peak above background, measured while the goniometer was moving, has been the basis for the measurement. The sizes of particles most suitable for diffraction techniques are reported in the literature as follows: for tungsten carbide a maximum of 40 microns7 and for quartz a maximum of 5 microns.' Very fine sizes of crystals are known to cause a broadening of the diffused beams; hence some workers have set a lower limit of 0.1 micron for quantitative results. Guinier10 has indicated that broadening of the diffracted X-ray beam is not appreciable for crystallites larger than 0.02 to 0.03

Jan 1, 1956

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

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

Jan 1, 1983

By T. P. Papazoglou, N. A. D. Parlee, W. C. Phelps

HE high vapor pressures of metals such as lead, calcium, lithium, bismuth, and magnesium at steel-making temperatures present experimental problems which have thus far rendered it almost impossible to obtain really reliable data for their solubilities in liquid iron. Many investigators have attempted to equilibrate liquid iron in the same crucible with the undissolved liquid phase of such volatile metals. The published results,l-6 although useful, are often questionable on the grounds of not being obtained under true equilibrium conditions. Such information is not always of practical value in any case since the equilibrium vapor pressure is often several atmospheres and steel is not currently made at these pressures. Solubility measurements at lower partial pressures (< 1 atm) appeared to be more practical and easier to make. These data could be used to calculate solubilities at other pressures where Henry&apos;s Law is obeyed, and to estimate solubilities under other conditions of temperature and pressure with useful accuracy. This note outlines the development of a general technique for the measurement of solubilities in systems of this kind. The method involves equilibration of the high-boiling metal at known temperatures with the vapors of a low-boiling metal at known partial pressures. Analyses of quenched samples provide vapor pressure-solubility data and make calculations of alloy thermodynamics possible. The principle involved is similar to the closed-system diffusion-transport method which is difficult to apply with most high-melting liquid metals like iron, due to closure problems. A two-zone horizontal (molybdenum wire element) tube furnace, Fig. 1, provides two constant-temperature zones in a reaction tuve (3/4 in. ID by 1 in. OD by 36 in.). A high-boiling metal sample, 9, introduced in the higher-temperature zone, is maintained at temperature T1, and a volatile metal sample, 7, in the lower-temperature zone, is held at T2. A stream of argon with 10 pct hydrogen is passed via flowmeters, 1, and mixing-drying towers, 2 and 3. through the reaction tube at a favorable flowrate and acts as a carrier gas which entrains vapor from the volatile liquid metal at a specific constant partial pressure. The gas mixture passes through a system of constrictions, 6 and 8, to the higher-temperature zone, establishes intimate contact with the high-boiling liquid sample, 9, and exits through a small condenser tube. The entire system is at (nearly) atmospheric pressure and the volume of carrier gas passed is measured with a wet gas meter. When the high-boiling sample attains equilibrium with the vapor phase, both samples are removed and quenched. Analysis of the high-boiling sample for solute content reveals its solubility for the low-boiling metal at a vapor pressure calculated from carrier-gas volume and either a) weight loss of volatile metal or 6) weight gain of high-boiling sample plus condenser (which collects the undissolved vapors). To test the capabilities of the method, liquid silver at 1403&apos;C was equilibrated with lead vapor at various partial pressures. The relatively simple Pb-Ag system was particularly suitable for uncovering experimental problems and providing the first step toward investigations with more volatile metals and more difficult combinations.

Jan 1, 1965

By William Schriever

The flow of a gas-free liquid through a spherical-grain sand has been investigated by S1ichter.l By theoretical considerations involving a rather large number of approximations he arrives at the following flow-formula: V = CPd2At NLB{1 - m) where V is the volume of liquid delivered in time t through a cylindrical column of sand of length L and cross-sectional area A by a pressure-drop P when the sand grains of diameter d are packed to a porosity m and the coefficient of viscosity of the liquid is n. C and B are constants which were determined from geometrical relations; B was in reality a function of . Values of B(l — m) for various values of m were given in a table. In the same Annual Report of the U. S. Geological Survey, King reported the results of his experiments on the flow of water through porous media. He found that there was a departure of about 86 per cent. from a linear relationship between rate of flow and pressure, for a specimen of Dunnville sandstone for pressure drops not exceeding 60 cm. of mercury. The specimen was approximately 4 cm. in diameter and 5.1 cm. in length. Over the same range of pressure he found a departure of 45 per cent. for a specimen of Madison sandstone. For unconsolidated sands he found departures as great as 49 per cent. for pressure drops of 70 cm. of mercury; for pressure drops not exceeding a few inches of water he obtained a close approximation to a linear relation. He also investi-

Jan 1, 1930

By B. E. Eakin, A. L. Lee

Atmospheric pressure viscosity values For light hydrocarbons were calculated by a simple mixing rule based on the Sutherland constants for pure components. This method was further extended so that the atmospheric pressure viscosity of a mixture may be calculated based on the molecular weight information only. INTRODUCTION A proven method for estimating the gas-phase viscosity of a mixture is the one by Herning and zipperer.l This mixing rule has the following form: One method for calculating the pure component viscosity i is the Sutherland equation,2 which has the following form: The Sutherland constants B and S are determined from experimental data. A set of Sutherland constants for light hydrocarbons based on the data of several investigators3-5 is presented in Table 1. APPLICATION The Herning-Zipperer equation and the Sutherland equation were used to obtain an equation for calculating the gas-phase viscosity of light hydrocarbons. This equation has the form, Eq. 3 was used successfully in the study of methane-n-butane mixtures.6 Further analysis indicated that the Sutherland constants are a function of molecular weight.7 The gas-phase viscosity may be calculated based on molecular weight information only. The following equation was developed: The atmospheric pressure viscosity of ethane, propane, n-butaneand four methane-propane mixtures calculated on the basis of Eq. 6 are compared with the data of other investigators in Tables 2, 3, 4 and 5, respectively. A plot of atmospheric pressure viscosity vs molecular weight based on Eq. 6 is presented in Fig. 1.

Jan 1, 1965

By Charles Stein

The effects of several different prior treatments an the creep behavior of 99.9995 pct aluminum at 260°C and 1000 psi canstant stress are compared with annealed specimens. Quenching from 538oC, irradiation with 2 mev electrons, and tension-compression room temperature fatigue damage were used to change the substructure of the specimens prior to their creep testing. The quenched and the irradiated specimens showed a larger primary and transient stage contribution to the creep curve than those specimens which were only annealed prim to creep testing. The specimen receiving prior fatigue damage at room temperature showed no first stage or transient creep when tested under identical conditions as the above specimens and had an average creep rate seven orders of magnitude lower than that of either the annealed, the quenched or the irradiated creep specimens. In a previous electron transmission microscopy investigationl of the substructure developed during the creep of pure aluminum at 0.57 Tm, it was noted that a large concentration of vacancy loops and dislocation loops were present in and near the subboundaries while the interior of the subgrains contained few of these defects and had a low dislocation density, see Fig. 1. Yim and Grant2 and Hazlett3 have shown that the presence in nickel of a substructure developed by prior cold work effectively reduces the primary and transient contributions to the creep curve. Hultgren,4 McLean,5 Gervais, Norton, and Grant,6 Chang and Grant,7 and others have shown the same effect with the higher stacking fault energy material, aluminum. However, the specific defect responsible for the change in the creep behavior developed by the prior cold work was not established. Specifically, what effectively interfered with dislocation motion in these materials at elevated temperatures? Was it jogs on dislocations produced by their interaction with other dislocations or with vacancy or dipole loops, see Fig. 1, or did the subgrain walls determine the mean free path for glissile dislocations? This would be a realistic possibility only if subgrain boundaries are effective barriers to dislocation motion. The ability of a subgrain boundary to act effectively against glissile dislocations depends on the number and the arrangement of dislocations in the subboundary,8 which in turn is a function of the creep strain.&apos; This paper compares the creep rate of specimens possessing a large vacancy concentration to that of annealed specimens and with specimens having a stable subboundary wall containing Lomer locks produced by prior fatigue damage. EXPERIMENTAL PROCEDURE Vacancy Loops. Aluminum creep specimens, 99.9995 pct pure, were machined from zone refined rod to a 2-in. gage length and a cross-sectional area of 0.025 in. These specimens were annealed at 538oC, 1000°F, for 10 min and quenched into ice water. They were then reheated to the creep test temperature of 260°C, 500°F, and held at this temperature for 18 min to al-low for vacancy condensation and loop formation. The specimens were subsequently creep tested at 1000 psi constant stress at 260°C. The temperature along the gage length was monitored by three Pt-Pt, 13 pct Rh thermocouples, two embedded in the upper and lower shoulders of the specimen and one attached to the mid-point of the gage length- The temperature gradient was held to ± l°C or less throughout the Creep test. Elongations were measured by two LVDT&apos;s &apos;connected in an averaging &apos;On-figuration, with quartz extension arms placed against the specimen at the extremes of the 2-in. gage length. i K 3J?--&apos; * *H - >^> - &apos; ^L r 77 . *y J OJL, J Fig. 1—Transmission electron micrograph of pure aluminum strained 8.2 pct at 260°C and 1000 psi constant stress. Note the presence of numerous loops near the light dislocation tangle

Jan 1, 1970

By A. H. Larson, R. J. McClincy

Laboratory-scale decopperizing experiments with multiple sulfur addifions were conducted at 330°C on ternary Pb-Cu alloys containing, as the third elenlent, Sn, Ag, As, Sb, Bi, Zn, and Au, common impurities in lead blast-furnace bullion. For silver and tin, an increased rate and extent of &apos;cofifier removal was obsert3ed. The elements As, Sb, Zn, Au, and Bi had no effect or less effect as compared to sulfur additions with no i)npurily additions. THE production of primary lead in the blast furnace yields an impure lead frequently containing such impurities as copper. antimony. arsenic. tin, gold, silver iron, oxygen. and sulfur. By cooling this lead to a temperature near its melting point. most of the iron, sulfur, and oxygen and part of the other impurities are removed in the form of a dross. With incipient solidification of the lead, the copper concentration wil have been reduced to 0.02 to 0.05 pct. depending upon the concentration of the other impurities. according to Davey.&apos; Since copper interferes with the treatment of silver after the desilverizing process, it is desirable to decrease the copper content of the lead still fur-ther before the lead is desilvered. The decopperizing of the lead is accomplished by stirring a small quantity. approximately 0.1 pct. of elemental sulfur into the lead at a temperature near its melting point, 330" to 360°C. The copper is removed as a copper sulfide which constitutes a small fraction of a voluminous dross consisting mostly of lead sulfide and entrained metallic lead. The residual copper concentration following the decopperizing operation is frequently as low as 0.001 to 0.005 pct. Thi fact has aroused considerable interest because the equilibrium copper concentration of lead in contact with solid PbS and solid Cu2S is at least an order of magnitude greater, 0.05 pct Cu at 330C. 1, 2 Most investigators have suggested that various impurities in the lead bullion are responsible for the very low copper concentrations frequently encountered in practice. There is little agreement, however? as to which of the impurities are helpful and which are not.3"11 Also. few investigators have sought to explain the mechanisms responsible for the removal of copper to very low concentrations. Willis and Blanks9 have proposed that a nonstoichiometric copper-deficient cuprous sulfide forms in place of the supposed Cu2S. Being copper-deficient, this sulfide phase would possess a low copper activity, and the diffusion of copper dissolved in the liquid lead into this phase would be greatly facilitated. Pin and wagner2 have investigated the removal of copper from liquid lead by studying the effect of impurity-doped lead sulfide on the decopperizing of pure Pb-Cu alloys. Samples of the doped PbS were held in contact with copper-saturated lead for 1 week at 33&apos;7°C. They reported a beneficial effect on decopperizing with bismuth and antimony and no effect with tin or silver. which is directly opposite to the results observed in practice and those reported by Davey 3 and this studv. The purpose of this paper is to describe the effects of certain additive elements on the extent to which copper can be removed fro111 liquid lead by successive additions of sulfur. The impurity elements were added individually to prepared Pb-Cu alloys. The resulting ternary alloys as well as a binary Pb-Cu alloy were then decopperized with repeated additions of sulfur. EXPERIMENTAL Materials. Granulated test lead with a purity of 99.999 pct and the additive elements Cu. Ag. Sb. Bi. Zn. Sn. and Au with purities of 99.99 pct were American Smelting and Refining Co. research-grade materials. The major impurities in the lead were 1 ppm each of iron and copper. all others being less than 1 ppm. The arsenic used was a technical-grade arsenic of 98+ pct purity. Reagent-grade flowers of sulfur were melted under argon to provide small pieces free of fines. Apparatus. The decopperizing experiments were carried out in a 25-mm-OD by 375-mm-long Pyrex tube sealed at one end. The tube was mounted vertically in a resistance-heated. hinge-type tube furnace controlled to within ±lcC. Temperature measurement was accomplished by means of a standardized chromel-alumel thermocouple sealed into the base of a silica. paddle-type stirring rod. All decopperizing experiments were carried out under an argon atmosphere. Procedure. A Pb-Cu starting alloy containing 0.05 pet Cu was prepared under carbon and poured into cold tap water to produce shot. The ternary alloys were prepared by melting together 100 g of the starting alloy and a sufficient amount of the impurity element to yield the desired concentration. The resulting alloy was then homogenized in a Pyrex tube at 450C with continuous stirring. The furnace temperature was then lowered to the operating temperature of 330°C. When thermal equilibrium had been obtained at the operating temperature, individual additions of 0.2 pct (0.2 g) of solid sulfur were added to the melt and stirred in. Stirring was continued for a period of 3 min. discontinued for 5 min. and resumed for the remaining 2 min of a 10-min cycle. This cycle was repeated for as many sulfur additions as desired. When the decopperizing experiment had been completed the lead bullion was quenched and samples of the bullion and dross phases were taken for analysis. Results. The results obtained in the decopperizing

Jan 1, 1970

By A. G. Taylor, H. P. Ehrlinger, U. C. Tainton

FOR the last 15 years lead has been the standard material for anodes in electrolytic zinc production and it has been generally accepted that this lead should be as free as possible from impurities. Laistl says, "Next in importance to the purity of solution is the purity of the electrodes and tank lining material. The purest lead obtainable should be used for anodes." The same conclusion is expressed by Ingalls2 and by two German investigators who say that experiments have demonstrated that the purer the lead, the greater is its resistance to corrosion during electrolysis.3 Pure lead, however, while the best material hitherto available for anodes, is not entirely free from objection. The plain lead anode gradually disintegrates under electrolysis, so that the life of an anode, say 1/4 in. thick, is frequently not more than two years. Part of the lead from the anode finds its way to the cathode, lowering the purity of the deposited zinc and decreasing the hydrogen overvoltage. The rest of the lead that comes from the disintegration of the anode goes into the manganese dioxide which is precipitated in the cells, rendering unsaleable what might otherwise be a valuable by-product. Yet a further disadvantage of plain lead anodes comes from their tendency to bend or buckle during electrolysis, as a result of intercrystalline oxidation. This makes it necessary to use a fairly aide spacing between anode and cathode to avoid short circuits, and thus leads to a higher power consurnption than would otherwise be necessary. The power consumption is also affected by the high decomposition voltage at a lead peroxide surface.

Jan 1, 1929

By T. O. Breitling

Two significant problems In mill design are environmental control and cost reduction. Environmental control is predicted to be 5-10% of total plant costs in some parts of the U.S. Operating and design companies will assign this new "unit operation" a place in the early flow sheets. More advanced environmental control will add to both capital and operating costs, so mill designers will strive to reduce other costs.

Jan 1, 1970

By David A. Keys

The actual procedure in estimating depth of overburden from magnetic observations made on the surface will vary with the form of the deposit and any theoretical discussion will apply only so far as the form of the deposit conforms to the mathematical premises on which the theory is based. Since the shapes of natural deposits in the field cannot be expressed rigorously in mathematical form, the best that can be done is to assume some simple form and deduce results. Fortunately many small variations in form produce only slight changes in the estimated depth of overburden, therefore the methods have a definite practical value. For the theoretical treatment, magnetic ore deposits may be divided roughly into three classes, depending upon whether the magnetic anomalies may be considered to arise from (1) a single magnetic pole, (2) a dike or (3) a magnetic doublet. Particular cases, falling under one or other of these classes, have been discussed by different authors.1"4 In general, it may be stated that when the magnetic effects as measured on the surface are due only to one pole of the body, the other pole being so remote that its effect may be neglected, the resultant anomaly F at any point will vary as the inverse square of the distance r of the point from the pole, which may be expressed mathematically by F = k/r2, where k is a constant. When the deposit occurs in the form of a long, thin dike, the resultant anomaly will be of the form F = k&apos;/r and when the deposit is in the shape of a lens, so that both poles have an appreciable effect on the surface, the result will be similar to that caused by a magnetic doublet, and the form will be F = k&apos;/r3. The particular class under which a deposit falls may be deduced from the results of the magnetic survey. The second class is readily found, but there may be difficulty in deciding between the first and third. Isody-namic contours of both the vertical and horizontal anomalies will often be a good guide in making a correct decision. The correctness of the theoretical calculations may be checked in two ways. The first is to make laboratory models of various known shapes and pole distributions, to calculate the anomalies to be expected

Jan 1, 1940

By David A. Keys

The actual procedure in estimating depth of overburden from magnetic observations made on the surface will vary with the form of the deposit and any theoretical discussion will apply only so far as the form of the deposit conforms to the mathematical premises on which the theory is based. Since the shapes of natural deposits in the field cannot be expressed rigorously in mathematical form, the best that can be done is to assume some simple form and deduce results. Fortunately many small variations in form produce only slight changes in the estimated depth of overburden, therefore the methods have a definite practical value. For the theoretical treatment, magnetic ore deposits may be divided roughly into three classes, depending upon whether the magnetic anomalies may be considered to arise from (1) a single magnetic pole, (2) a dike or (3) a magnetic doublet. Particular cases, falling under one or other of these classes, have been discussed by different authors.1"4 In general, it may be stated that when the magnetic effects as measured on the surface are due only to one pole of the body, the other pole being so remote that its effect may be neglected, the resultant anomaly F at any point will vary as the inverse square of the distance r of the point from the pole, which may be expressed mathematically by F = k/r2, where k is a constant. When the deposit occurs in the form of a long, thin dike, the resultant anomaly will be of the form F = k&apos;/r and when the deposit is in the shape of a lens, so that both poles have an appreciable effect on the surface, the result will be similar to that caused by a magnetic doublet, and the form will be F = k&apos;/r3. The particular class under which a deposit falls may be deduced from the results of the magnetic survey. The second class is readily found, but there may be difficulty in deciding between the first and third. Isody-namic contours of both the vertical and horizontal anomalies will often be a good guide in making a correct decision. The correctness of the theoretical calculations may be checked in two ways. The first is to make laboratory models of various known shapes and pole distributions, to calculate the anomalies to be expected

Jan 1, 1940

By Kenyon Burch

ONE of the advantages presented by electric drive in many classes of work is the ease with which the electric motor can be controlled automatically. In a large number of cases certain features of the control are automatic-for example, the rate of acceleration may be limited automatically or the equipment may be stopped automatically at the limit of travel-but the equipment is started and ordinarily is stopped by hand. . In other cases the motion of the machinery is utilized to start, control the: speed, and stop the motor automatically, independently of any operator. A considerable proportion of the large mine hoists now in use have certain automatic features, particularly protective devices against over-winding, and, in some classes of electric hoists, devices for preventing excessive acceleration or retardation. The large automatic hoists discussed in this paper, however, are completely automatic, i.e., capable of snaking their trips without the presence of an operator at the control levers. According to circumstances, various advantages may be obtained by automatic control, chief of which are decreased power consumption&apos;, increased precision and safety of operation, and decreased, cost of attendance. The first step in the analysis of a prospective automatic mine hoist is to determine whether automatic operation is feasible at all. If men are to be hoisted, or levels changed, the attention of an operator is required for these purposes; but tinder some conditions it may be entirely practicable and advantageous to build the equipment so that, while provision is made for hoisting mien or changing levels, ore can be hoisted automatically from, any one level. If, however, an operator&apos;s attention is required every few minutes for changing levels, handling men or drills, or other work requiring hand control, it is obvious that automatic operation between times will not be of any, practical benefit.

Jan 3, 1916

By H. W. Schadler

The superplastic behavior of low carbon and manganese bearing steels has been evaluated. The results of elevated-temperature stress-strain rate and elongation tests are reported which indicate that high strain rate sensitivity (>0.5) and adequate elongations are achievable in ultrafine-grained steels 1 to 2 p, but at strain rates which are not commercially attractive. It has been demonstrated that the fine grain size is retained after long times and high strains if the temperature is kept within the range of the two-phase (a + ?) field. INTEREST in superplasticity has been stimulated by the desire to: 1) understand the origin of, and mechanisms responsible for, the large uniform elongations observed, and 2) exploit this property for commercial advantage in metal forming operations. Although the Zn-22 wt pct Al alloy presently being studied as a model material1-5 may find application in sheet forming and extrusion, the potential of super-plasticity should best be realized in large tonnage materials, such as steel and aluminum alloys. The degree of superplasticity in low-carbon and manganese-bearing steels has been evaluated. The results of elevated-temperature stress-strain rate and elongation tests on ultrafine grain size material are reported. Avery and Backofen6 and Hart7 have shown that geometrically stable flow leading to extensive uniform elongation in the tension test is associated with high values of the strain rate sensitivity, m, defined: Lozinsky had reported that titanium and zirconium experience permanent deformation if subjected to a constant load and cyclic heating through the temperature range of the phase transformation. Although strain rate sensitivities in excess of about 0.2 had not been reported previously for steel, permanent deformation had been observed11-13 in a wide variety of steels subjected to a constant load and cyclic heating through the temperature range of the a-? phase transformation. Thus by analogy, steel could be expected to exhibit high strain rate sensitivity but only in the a + ? condition. High strain rate sensitivity has been observed at 650°C (a + Fe3C), but not reported as such, by Bailey, Dickenson, and pearson14 in 1931 at a strain rate of about 10-9 per min. It then remained to determine whether high tensile elongations would be observed in a rate-controlled test at constant temperature and at what strain rate strain rate sensitivity values greater than 0.5 would be observed. Since previous experience8 had indicated the importance of fine grain size to realizing high m at reasonable strain rates, it was first necessary to produce an ultrafine grain size and then keep the grains from growing during the test. The results of this investigation show that fine grain size can be readily produced16 and maintained by restricting the temperature of testing to below the ? transus. Further, superplasticity (high m and large uniform elongation) is observed. However, the strain rate range is only marginally useful for commercial forming operations with the finest grain size produced. MATERIAL The material investigated initially was a 1.9 wt pct Mn, 0.42 wt pct C, hot-rolled bar previously used by Low and Turka1015 and available in the laboratory. The 0.500-in. round was cold-rolled to a 0.090-in. flat, annealed for 4 hr at 850°C in argon, air-cooled, and cold-finished to 0.050 in. Subsequently, four additional steels of the compositions given in Table I were investigated to explore the effects of manganese, carbon, and test temperature on the observed stress-strain rate behavior. These steels were melted under argon, cast to 0.75 by 2 by 5 in. slabs, hot-rolled at 850°C, surface-finished to 0.13 in., and cold-finished to 0.050 in. Sheet tensile specimens 0.200 in. x thickness with 1- or 2-in. gage length were cut parallel to the rolling direction. Table I also includes the nominal transformation temperatures for the AISI 1340 and the four experimental steels. EXPERIMENTAL PROCEDURE Production of Fine Grain Size. Ultrafine grain size (1 to 5 p) was considered essential for this studv. Grange16 has described two techniques for producing

Jan 1, 1969

By W. C. Leslie, R. M. Fisher, K. G. Carroll

A nitride, believed to be Si3N, has been separated from three nitrided silicon steels. Germanium nitride, Ge3N4, has been prepared from pure germanium. Comparison of the diffraction patterns indicates that the two nitrides are isomorphous; on orthorhombic structure is suggested in place of the rhombohedral structure previously reported for Ge3N4. THE possibility that a nitride of silicon may, under appropriate conditions, precipitate in silicon steels or in steels killed with silicon makes it desirable to have some positive means of identifying such a compound. One such means is the X-ray or electron diffraction pattern of the nitride. A review of the meager data in the literature indicates that the nitride most likely to form is Si3N4, a conclusion supported by the results of a study of nitrided silicon steels to be published shortly by L. S. Darken and R. P. Smith, of this Laboratory. Unfortunately, there is available no diffraction pattern for this nitride. Data have been reported, however, for Ge3N4 which, judging from the similarity between germanium and silicon, might be expected to be isomorphous with Si3N4. An effort was made, therefore, to form such a silicon nitride, to determine its composition and diffraction pattern and, if possible, its structure. To this end a series of three silicon steels was nitrided under controlled conditions with the resultant formation of nitride particles which yielded an electron diffraction pattern in situ. The particles were then extracted from the steel and an attempt was made to determine their chemical composition. X-ray and electron diffraction patterns were also obtained from the extracted particles, which indicate that the nitride is isomorphous with Ge3N4, although a complete determination of the structure has not been possible. These results show that a silicon nitride with a well-defined diffraction pattern can form in silicon steels, and they suggest that this nitride is Si3N4. Materials and Procedures The sillicon steels investigated were In the form of thin sheet or strip and had the composition shown in Table I. The 0.58 and 1.21 pct Si steels were nitrided by holding them at 1110°F in an H2-NH3 atmosphere containing 3 pct ammonia for 48 hr. The nitride particle size was increased by subsequent heating at 1500°F for 13 1/2 hr in helium. The resulting microstructure is shown in Fig. 1. The steel containing 3.20 pct Si was nitrided at 1200°F for 16 1/2 hr after which it was held in helium at 1500°F for 4 hr. Its structure as seen under the electron microscope is illustrated by Fig. 2. As would be expected from the higher silicon content and the shorter holding time at 1500 °F, the nitride particles are smaller and more numerous than those in the 1.21 pct Si steel. The ammonia-hydrogen treatment reduced the carbon content of the steels to a very low level, so no interference was encountered from carbon or carbides. In the case of the 3.2 pct Si steel, the carbon was reduced to 0.003 pct before nitriding by heating in dry hydrogen. Attempts to obtain an X-ray diffraction pattern from polished and etched surfaces of the 1.21 and the 3.20 pct Si steels were unsuccessful. However, an electron diffraction pattern was obtained from the surface of the steels. The interplanar spacings obtained from these patterns arc shown in Table 11, col. 5. The nitride particles were then extracted from all three steels by dissolving the ferrite matrix in bromine-methyl acetate, the solution used in the Beeghly method for the extraction of aluminum nitride from steel. X-ray diffraction patterns of these residues, obtained by means of a spectrometer and by a 57 mm Debye-Scherrer powder camera using filtered cobalt or chromium radiation, are given in Table II along with the pattern obtained by

Jan 1, 1953

By S. Stopek, E. D. Hinkley, R. H. Rediker

Epitaxial heterojunctions have been prepared by melting the lower -melting-point semiconductor of the interface between two dijferent semiconductors. when the temperature is reduced, the melted material recrystallizes, having alloyed into the higher -melting-point semiconditctor. The electrical and elcctro-optical properties of such single-crystal heterojunctions between GaAs and Gash and between p-type InAs and n-type Gash are the subject of this paper. The forward current varies as exp (AV), where A is substantially independent of temperatuve. For Gds-Gash heterojunctions at temperatures above, 370°K, if the current-voltage relationship were to he expressed as exp (qV/nkT), then n would he less than unity. The injection luminescence associated with forward current is, for the most part, characteristic. of the lower bandgap semicondutctor. These results can he explained by carrier injection into the lower bandgap semiconductor by tunneling through a barrier at the interface. The photovoltaic effect measured for incident photons having energies in the range between the bandgaps of the two semiconductors is much smaller than that produced by higher-energy photon The smallness of this between -the-gap photovoltaic response can he explained by the low probability for penetration of the barrier by the carriers produced in the smaller -bandgap semiconductor. THE technique of interface alloying has been used to produce single-crystal junctions between dissimilar semiconductors.&apos; Oriented wafers are placed on a carbon heater strip, Fig. 1. so that semiconductor S1, which has the lower melting point, is supported by semiconductor S2. Electrical current passed through the heater strip produces a temperature gradient such that S2 is at a higher temperature than S1. As the temperature is raised the lower face of S1 begins to melt. Before the entire wafer can melt, however, the heater-strip current is turned off and, as illustrated in Fig. l. the melted portion recrystallizes, having alloyed into S2. Junctions have been fabricated by the above procedure between GaAs and germanium, between GaAs and GaSb, and between InAs and GaSb. Mroczkowski, Lavine. and Gatos have described the metallurgical and chemical aspects of the GaAs-Ge junction.2 The transition from GaAs to germanium is not monotonic and a portion of the recrystallized region consists of the GaAs-Ge eutectic. Since gallium is an acceptor and arsenic is a donor in germanium, since germanium dopes GaAs, and since the electrical properties of the GaAs-Ge eutectic have not been investigated, any interpretation of the electrical characteristics in terms of simple heterojunction theory would be incorrect. That the rectification of GaAs-Ge heterojunctions is not a property of the impurity doping of the GaAs or the germanium, but is most probably due to the impurity distribution in the recrystallized region, is clear from the fact that forward conduction occurred for all the GaAs-Ge interface-alloy junctions (whether they be n-n,n-p. p-n. or p-p) when the germanium was biased positively. The electrical characteristics of these GaAs-Ge heterojunctions will not be discussed further in this paper. Electron-beam microprobe analysis of GaAs-GaSb heterojunctions showed that the transition from arsenic to antimony atoms was without structure. and that the transition occurred within a 2 to 3 region.&apos; In this paper we will describe the electrical properties of the GaAs-GaSb heterojunctions as well as electrical and electro-optical properties of the InAs-GaSb heterojunctions. A band model for the junction will be proposed which can explain these properties. ELECTRICAL PROPERTIES OF GaAs-GaSb HETEROJUNCTIONS After interface alloying. in preparation for the electrical and electro-optical experiments, ohmic contacts were made by conventional means. For example. Kovar tabs clad with tin were alloyed to n-GaAs or similar tabs clad with Au-Zn were alloyed to p-GaAs. The units were mounted and then etched. All four combinations of conductivity types

Jan 1, 1965

By Walter Rose

In a recent technical note, Owen Thornton&apos; suggests that wetting liquid relative permeability may be derived from the relationship: where Pd/Pc is the ratio of displacement pressure to capillary pressure at the wetting liquid satdration Sw, and I is the resistivity index at this saturation. Thornton shows that this expression gives values for relative permeability in good agreement with those experimentally determined by Leverett. We have recently developed another expression for krw which seems preferable to Equation 1 since it requires fewer experimental data for its verification. This equation may be derived in the following manner. Making use of the analogy between mass transfer of fluid in a porous medium and electrical conductivity through the fluid in the same medium, we can postulate a hydraulic formation resistivity factor analogous to an electrical formation resistivity factor. From Poiseuille&apos;s Law the resistivity to flow in a tube of radius R is of the form 8u/R², where is the viscosity of the fluid. Similarly7 the resistivity to flow in a porous medium of the same dimensions as the tube is given by D&apos;Arcy&apos;s Law as where k is the permeability of the medium. The hydraulic formation resistivity factor, Ff, is thus the quotient of these two resistivities, or: But permeability is defined by the KO-zeny equation as: where: r is the average pore radius of the porous medium, L. is the average tortuous length of the average pore, + is the porosity, and L is the bed length. This gives Fr as a function of the (La/L) .-.-.tortuosity ratio and porosity, or in explicit form: where: T = (La/L)&apos;, and N is the number of pores of radius, r, which will be found in any cross sectional area, x R&apos;. That is, N = R²/r². It will be clear from the foregoing considerations that the hydraulic formation fat. tor should be dependent on the value of R which in every case must be arbitrarily selected. By analogy, the hydrauiic formation factor, Fcf, characterizing the porous medium at Sw<1,will be given by: R² NT. where ke is the effective wetting liquid permeability obtaining at Sw<l, and Te is the square of the (Las/L) ratio defined by Thornton. Therefore, the hydraulic resistivity index, If, is: However, Thrnton gives I as: and it follows that: Equation 2, when checked against the data of wyckoff and Botset² and some of the Morse et al data³, gives comparisons as shown at the bottom of this page. Computations based on Leverett&apos;s data, quoted by Thorntaon, gives computed krw values somewhat lower than the experimental values. Moreover, other instances can be cited where we find that neither Equation 1 nor Equa tion 2 is checked by experimental data. The work of Botset&apos; on the Nichols Buff sandstone and Morse et a1 on oil wetted Bradford sandstone are such instances. A reason for these discrepancies may lie in the fact that if both Equations 1 and 2 are presumed always to give accurate values for krw, it follows that PC = Pd/Sw". Such an expression, it is well known, is too simple to describe accurately the capillary pressure behavior of all porous media. However, it is believed that a particular advantage resides on the use of Equation 2, since it is not dependent for its utility on a knowledge of the capillary pressures obtaining in dynamic flow svsterns. No practical technique for measuring the capillary pressures characterizing the fluid distributions in dynamic flow systems has yet been proposed. Acknowledgment is given to Dr. Paul D. Foote, executive vice-president of the Gulf Research and Development Company, for permission to publish this note. 1. Owen F. Thornton, "A Note on the Valuation of Relative Permeability," J. Pet. Tech., 1:7, Section 1, July, 1949. 2. R. D. Wyckoff arid H. G. Botset, "The Flow of Gas-Liquid Mixtures through Porous Media," Physics, 7:325-345, 1936. 3. R. A. Morse, P: L. Terwilliger, .and S. T. Yuster, "Relative Permeability Measurements of Small Core Samples," Oil and Gas J., 46:109-125, Aug. 23, 1947. 4. H. G. Botset, "Flow of Gas-Liquid Mixtures Through Consolidated Sand," Trans. AIME, 136:91, 1940.

Jan 1, 1949

By Walter Rose

In a recent technical note, Owen Thornton&apos; suggests that wetting liquid relative permeability may be derived from the relationship: where Pd/Pc is the ratio of displacement pressure to capillary pressure at the wetting liquid satdration Sw, and I is the resistivity index at this saturation. Thornton shows that this expression gives values for relative permeability in good agreement with those experimentally determined by Leverett. We have recently developed another expression for krw which seems preferable to Equation 1 since it requires fewer experimental data for its verification. This equation may be derived in the following manner. Making use of the analogy between mass transfer of fluid in a porous medium and electrical conductivity through the fluid in the same medium, we can postulate a hydraulic formation resistivity factor analogous to an electrical formation resistivity factor. From Poiseuille&apos;s Law the resistivity to flow in a tube of radius R is of the form 8u/R², where is the viscosity of the fluid. Similarly7 the resistivity to flow in a porous medium of the same dimensions as the tube is given by D&apos;Arcy&apos;s Law as where k is the permeability of the medium. The hydraulic formation resistivity factor, Ff, is thus the quotient of these two resistivities, or: But permeability is defined by the KO-zeny equation as: where: r is the average pore radius of the porous medium, L. is the average tortuous length of the average pore, + is the porosity, and L is the bed length. This gives Fr as a function of the (La/L) .-.-.tortuosity ratio and porosity, or in explicit form: where: T = (La/L)&apos;, and N is the number of pores of radius, r, which will be found in any cross sectional area, x R&apos;. That is, N = R²/r². It will be clear from the foregoing considerations that the hydraulic formation fat. tor should be dependent on the value of R which in every case must be arbitrarily selected. By analogy, the hydrauiic formation factor, Fcf, characterizing the porous medium at Sw<1,will be given by: R² NT. where ke is the effective wetting liquid permeability obtaining at Sw<l, and Te is the square of the (Las/L) ratio defined by Thornton. Therefore, the hydraulic resistivity index, If, is: However, Thrnton gives I as: and it follows that: Equation 2, when checked against the data of wyckoff and Botset² and some of the Morse et al data³, gives comparisons as shown at the bottom of this page. Computations based on Leverett&apos;s data, quoted by Thorntaon, gives computed krw values somewhat lower than the experimental values. Moreover, other instances can be cited where we find that neither Equation 1 nor Equa tion 2 is checked by experimental data. The work of Botset&apos; on the Nichols Buff sandstone and Morse et a1 on oil wetted Bradford sandstone are such instances. A reason for these discrepancies may lie in the fact that if both Equations 1 and 2 are presumed always to give accurate values for krw, it follows that PC = Pd/Sw". Such an expression, it is well known, is too simple to describe accurately the capillary pressure behavior of all porous media. However, it is believed that a particular advantage resides on the use of Equation 2, since it is not dependent for its utility on a knowledge of the capillary pressures obtaining in dynamic flow svsterns. No practical technique for measuring the capillary pressures characterizing the fluid distributions in dynamic flow systems has yet been proposed. Acknowledgment is given to Dr. Paul D. Foote, executive vice-president of the Gulf Research and Development Company, for permission to publish this note. 1. Owen F. Thornton, "A Note on the Valuation of Relative Permeability," J. Pet. Tech., 1:7, Section 1, July, 1949. 2. R. D. Wyckoff arid H. G. Botset, "The Flow of Gas-Liquid Mixtures through Porous Media," Physics, 7:325-345, 1936. 3. R. A. Morse, P: L. Terwilliger, .and S. T. Yuster, "Relative Permeability Measurements of Small Core Samples," Oil and Gas J., 46:109-125, Aug. 23, 1947. 4. H. G. Botset, "Flow of Gas-Liquid Mixtures Through Consolidated Sand," Trans. AIME, 136:91, 1940.

Jan 1, 1949