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By J. Chipman, N. J. Grant, H. L. Bishop

In a recent review1 of the iron-oxide activity of simple open-hearth type slags containing lime, magnesia, silica, and iron oxide, it was established that activity values were lacking in the range of slag composition most interesting to open-hearth operators. That is, laboratory experimental results were not available for simple slag compositions having basicity ratios ranging from about 1.5 to 4.0 and iron-oxide concentrations ranging from about 10 to 30 pct. However, in order to estimate iron-oxide activities for the range of slag compositions of interest to open-hearth operators, activity values were calculated using ionic concepts of slags by following the treatment of Flood and Grjotheim.2 A study of these data indicated that the desired slag compositions were probably not obtained due to the limited solubility of the basic constituents lime and magnesia. Thus, it was decided to make additional heats equilibrating molten carbon-free iron with slags containing MnO, P2O5, and A12O3 in percentages commonly found in open-hearth refining slags, and to examine the liquid slag field over a greater composition range than that of the simple slags studied previously. The ranges of the constituents used in the present study are summarized as follows: Metal Pct 0 0.070-0.201 PctMn 0.005-0.142 Slag Pet CaO 24.45-41.90 Pct MgO 3.23- 9.21 Pct MnO 1.13- 9.62 Pct FeO 20.08-48.7 Pet Fe2o3 2.23-10.14 Pct SiO, 3.51-28.50 Pct P2O5 0.23- 3.65 Pct A12O3 0.43- 3.38 The experimental procedure consisted of equilibrating the slag and metal in an induction furnace provided with a carbon-arc slag heater as described previously by Winkler and chipman.3 The data obtained were used to establish both the iron-oxide activity values and the activity coefficients for FeO and MnO in the open-hearth type slags. RESULTS AND DISCUSSION I) Iron-Oxide Activity of Open-Hearth Type Slags—The iron-oxide activities of various iron-oxide contents of the open-hearth type slags, as calculated from the oxygen content of the metal by the method of Taylor and Chipman, 4 are plotted against the molar ratio of bases to acids in Fig. 1. The data represent slag-metal studies carried out in the temperature range 1530" to 1700°C. For the purpose of this study, the constituents CaO, MgO, and MnO are assumed equal on a molar basis in their effect on the iron-oxide activity; and, PO2,5 is calculated as an equivalent to SiO2 and A12O3 on a molar basis. The total iron content of the slag is calculated to FeO and is represented by the symbol (FeO) The points presented in Fig. 1 represent the results of this investigation for iron-oxide slags containing between 1 and 8 mol pct MnO and less than about 2 mol pct each of A12o3 and PO2,5 The curved lines shown in Fig. 1 represent the results of the summary1 of the iron-oxide activities of simple slags of the system (CaO + MgO) - SiOz - (FeO),. The solid portions of the curved lines represent experimental data, and the broken portions represent iron-oxide activities calculated from ionic concepts of slags. The iron-oxide values of the open-hearth type slags of this study are in good agreement with the values of the simple iron-oxide slag system. Only a slight modification is indicated. The results of this study suggest that the line for 20 mol pct (FeO), should be lowered slightly in the basicity range 1.5 to about 3.

Jan 1, 1959

By F. G. Tahmoush, N. J. Grant, E. P. Abrahamson

The effect of grain size on the brittle -ductile transition temperature of pure iron and three dilute Fe-W alloys has been investigated by slow bend tests. The brittle-ductile transition temperature for pure iron increased with increasing grain sizes from 8000 to 500 grains per sq mm where the transition temperature then decreased with further increase in grain size. The dilute additions of tungsten lowered the transition temperature for fine grain sizes, suppressed a reversal of the curve in the coarse grained portion, and raised the transition temperature for this region. The frequency of twin occurrence increased with increasing grain size and decreasing temperature. The object of the present investigation was to re-examine the effect of grain size on the brittle -ductile transition of pure iron and several dilute Fe-W alloys. Current literature normally presents a simple relationship between grain size and the brittle-ductile transition temperature, namely, that the transition temperature increases with increasing grain size. Heslop and etch' have studied the transition temperature of mild steel by V-notch Charpy impact tests for the grain size range between 200 and 6000 grains per sq mm. The result was a linear plot of In 1-1/2 vs brittle-ductile transition temperature, where 1 is the grain diameter. Recent experiments by sipos2 revealed the possibility of a reversal in the slope of the grain size vs transition temperature plot for pure iron; however, his investigation was limited in scope, and the data obtained were insufficient to permit a definitive curve. The present investigation was undertaken in an effort to confirm or extend the grain size effect observed in iron by Sipos. PROCEDURE All alloys were vacuum induction melted, using 99.95 pct Fe containing 0.03 pct 0 and 0.004 pct C. Four ingots were made, ranging from pure iron to iron with 0.035 pct W. The chemical composition of each alloy, after processing, is given in Table I. The four alloys were prepared as 45 lb heats and cast into 4 in. diam by 16 in. molds. They were then extruded into 3/4 in. rods at 1800oF. Specimens for chemical analysis, taken from various positions along the rods, revealed no variation in tungsten content along the length. The extruded rods were reduced to 0.156 in. diam stock by a series of cold swaging operations involving intermediate vacuum annealing treatments. This procedure yielded a final cold reduction in area ranging from 10 to 90 pct. The swaged rods were chemically cleaned, cold pressed to a cross section of 0.10 by 0.18 by 1.0 in. long (2.54 by 4.73 by 25 mm). The specimens were wrapped in iron foil and vacuum sealed (10-4 mm Hg) in quartz. Various heat treatments at temperatures from 700" to 1300oC were utilized to produce grain sizes from 8000 to 1 grain per sq mm. Grain sizes were determined by measuring the average area of grains in a polished section. In specimens with grain size greater than 20 grains per sq mm, the total number of grains in the cross section was counted and then divided by the area to yield grains per sq mm. Grain size was measured both for longitudinal and transverse cross sections. The average of the two measurements was plotted vs transition temperature. An Instron Table Model testing machine was used to determine the transition temperature of specimens tested in three point horizontal bending. The specimen was supported by two 1/4 in. dowels one in. apart. Bending was accomplished with a push rod, the end of which had an 1/8 in. radius. The head speed was 50 in. per min. Temperatures were established through the use

Jan 1, 1963

INTRODUCTION The meaning of cash flow has already been defined as after-tax earnings available for use by the corporation or business. It is therefore apparent that a significant part of mineral economic evaluation is concerned with correct and accurate computation of taxable income under existing tax laws, and the proper treatment of tax allowances and deductions. The various methods of depreciation, depletion, and amortization are all means whereby the investor can earn back or recover, before taxes are imposed, the investments made for purposes of doing business or for the production of income. Each of these three types of tax deduction are applied to different investment items such as property, equipment, buildings, research, etc. The cost of buildings, equipment, machinery, and transportation equipment are representative of capitalized business costs that must be recovered by depreciation over the life of the asset. The property investment costs of acquiring natural resources such as minerals, oil and gas, and standing timber are examples of investment costs that may be recovered by depletion. Several other capitalized business costs, such as research expenses and trademark expenses, may be recovered by amortization. The term "capitalized" means that the in- vestment cost is not deducted as an operating expense in the year that it is incurred, but is spread out over a period of years through depreciation, depletion, or amortization deductions according to lawful schedules. Of course, any capitalized cost can only be deducted once. The other costs of doing business which are not capitalized, such as labor and operating costs, are said to be "expensed," i.e., they are written off for tax purposes in the year they occur. DEPRECIATION Depreciation, according to US tax law, is a deduction from taxable income permitted as a "reasonable allowance for the exhaustion, wear and tear of property out of its use or employment in a business." The original cost of the property is the usual basis for determining depreciation. Since depreciation write-off is applied over the useful life of the asset, the first step is to determine the estimated useful life or its allowable depreciable life. These two lives may or may not be the same. No average life is applicable to all types of depreciable assets under different conditions or in different types of business. It is the taxpayer's responsibility to determine the useful life of the asset according to his own operating conditions and experience, within the guidelines set by the US Internal Revenue Service (IRS). The Class Life Depreciation Range (ADR) System permits the taxpayer to choose from a range of depreciable lives at 20% below or above the present guideline life. The second step required in calculating depreciation is to determine an estimated salvage value for the asset. The salvage value is an estimate, made at the time of acquisition, of the amount of money which can be realized by the sale of the asset at the end of its useful life. The four depreciation methods most widely used in the US are: (1) the straight-line method, (2) the declining balance method (double declining balance), (3) sum of the years - digits method, and (4) units of production method. Straight-Line Depreciation Method The straight-line depreciation method is the simplest for calculating depreciation. The depreciation for each year is determined by subtracting the estimated salvage value from the first cost of the asset and dividing the result by the years of life. The method gives equal annual deduction over the depreciable life. Salvage value equal to or less than 10% of the first cost may be ignored. [first cost - salvage Straight line = years of life ] Declining Balance Depreciation Method Declining balance depreciation, or double declining balance (DDB) as most frequently applied, is used for accelerating depreciation in the early years of the asset and is the fastest depreciation method allowed by the US Internal Revenue Service. The declining balance

Jan 1, 1980

By G. R. Smithson, John E. Hanway

This paper, the first of two papers covering the development of a metallurgical process for the treatment of a complex ore, describes the bench-scale experimental Program undertaken as a basis for initial process design. The results of using various potentially applicable process steps are detailed and the basis for selecting certain of these steps for Pilot-plant development is outlined. A process consisting of the steps of sulfate roasting, leaching, copper recovery, solution purification, zinc electrolysis, and brine leaching for lead recovery was recommended for further study. THIS paper describes the bench-scale development of a sulfate roasting process designed to recover copper, lead, and zinc from a complex ore. The extrapolation of the bench-scale results to pilot-plant design and the results of pilot-plant operation will be described by Lund, et al.' in a subsequent paper. Several years ago, the research divisions of the St. Joseph Lead Co. became responsible for the development of a metallurgical procedure to treat a complex sulfide ore containing copper, zinc, and lead sulfides. The primary objective was to devise a procedure which would permit economic recovery of the nonferrous values. As could logically be expected, the study of conventional flotation procedures was undertaken; in addition, however, because the physical nature of the mineral association in the complex ore indicated that physical methods of ore beneficiation might not prove to be entirely satisfactory, a program also was established to investigate alternative methods of direct metallurgical treatment for recovery of the metal values. Exploratory test work by the St. Joseph Lead staff led to the conclusion that a sulfate roasting technique represented the most attractive approach for a direct metallurgical process. Sulfate roasting would provide for the selective conversion of the nonferrous values to forms soluble in conventiona1.1 leaching agents and would permit their efficient separation from the insoluble iron and other gangue materials. It should allow efficient utilization of the thermal energy available from oxidation of the sulfide raw material, and might provide an opportunity for recovering byproducts such as iron and sulfur. The decision to evaluate a selective sulfation procedure as a basic processing step also was influenced by the excellent results obtained by the St. Joseph Lead Co. in using this procedure for the recovery of cadmium from zinc sulfide concentrates. The choice of selective sulfation as a major processing step served as a basis for an extended experimental investigation of the process with bench- and pilot-scale equipment to prove its feasibility. St Joseph Lead Co., aware that Battelle Memorial 1nstitute had participated in the development of several fluidized-bed sulfate roasting procedures, asked Battelle to undertake the bench-scale development of a suitable sulfation process. This paper describes the procedure used and the results obtained at Battelle in the treatment of the complex sulfide ore. EXPERIMENTAL WORK Nature of the Raw Material. The sulfide ore deposits of interest are essentially a massive sulfide, with the indivudual ore minerals being extremely fine grained and intimately intergrown with each other and with the associated gangue. The zinc in the ore is present mainly as sphalerite (ZnS), the lead as galena (PbS), and the iron either as pyrite (Fes2) or as pyrrhotite (FeS). Table I shows the chemical analysis of various individual samples of the ore. These analyses illustrate the variability of the ore body and impose a rather severe requirement on any type of metallurgical process—that of being able to tolerate probable wide variations in the assay of the head ore. The tentative flowsheet designed to serve as a guide for the initial experimental investigation of

Jan 1, 1962

By G. K. Manning, W. E. Few

T has been known for some time that both'inter-granular carbide and intergranular oxide phases cause brittleness in molybdenum. Parke and Ham' indicated that 0.0025 pct 0 present in molybdenum after solidification from the melt was sufficient to induce intergranular brittleness during forging. However, material containing 0.06 pct C could still be forged. However, adding this amount of carbon to insure deoxidation causes the formation of carbides on freezing. These carbides are precipitated at grain boundaries, and Rengstorff and Fischer' have shown they have an unfavorable effect on the room-temperature ductility of molybdenum. Take? in 1928 proposed that the solubility of carbon was a constant value of 0.30 pct from room temperature up to 1800°C. In 1935, Sykes, Van Horn, and Tucker' obtained X-ray data suggesting that the solid solubility of carbon in molybdenum was between 0.07 and 0.09 pct at temperatures from 1500" to 2100°C. The a-solubility line for carbon was included in the Mo-C diagram given in the Metals Handbook as a tentative line based on the data of Sykes, Van Horn, and Tucker. In order to determine the partial phase diagrams for these systems, it was necessary to construct a high-temperature furnace suitable for heat treating molybdenum at temperatures up to 4000°F. A molybdenum resistance furnace was built for this purpose in which samples could be heat treated in a purified atmosphere and quenched from any desired temperature.' Experimental Work Solubility of Carbon: Arc-cast molybdenum supplied by the Climax Molybdenum Co. and 18-gage molybdenum wire samples produced by powder metallurgy at the Fansteel Metallurgical Co. were used in determining the a-solubility line for the MO-C system. The cast molybdenum contained 0.032 pct C initially. Two lots of 18-gage wire were used, one containing 0.011 pct C and the other 0.005 pct C initially. All material contained a carbide phase as received. Fig. 1 illustrates the carbides found in the untreated material. Heat treatment in a purified hydrogen atmosphere was found to reduce the carbon content in the as-cast molybdenum. Purified argonT also decarburized molybdenum samples at temperatures above 3500°F; however, the carbon content remained constant during heat treatment in purified argon at lower tem- peratures. The argon was purified by passing it over titanium at 750 °C and then through a magnesium perchlorate drying tower. The hydrogen was purified by passing it through a commercial hydrogen deoxidizer, followed by a drying tower. Specimens having different initial carbon contents were heat treated in argon at 3500°F for periods ranging from 5 min to 5 hr and were analyzed for carbon. Samples of the same final carbon content had the same microstructures regardless of time at temperature. Furthermore, a gradient between center and surface in the amount of undissolved carbide particles present was never observed. It was concluded that at these low carbon levels, and at the high temperatures involved, diffusion proceeds so rapidly that the specimen is always substantially homogeneous in carbon content, in spite of the gradual loss in carbon. Based on the above findings, the following procedure was used to determine the carbon-solubility limit at 3000°, 3500°, and 4000°F. At each of these temperatures, a number of samples were heat treated, as shown in Table I, and rapidly cooled to room temperature. Since the samples lost carbon in proportion to the length of time heated, a series of samples were obtained of varying carbon contents. By use of a rapid cooling rate, the phases present at the high temperature were retained at room temperature. All samples were examined microscopically. Those found to contain either no excess carbide phase or only a very small amount were then analyzed for carbon. The samples actually analyzed for carbon

Jan 1, 1953

By K. D. Jacob, W. L. Hill

Fluorine recovery in the United States has been restricted chiefly to manufacture of ordinary superphosphate and wet-process phosphoric acid. However, there is an expanding use of fluorine by industry in the form of hydrofluoric acid. Improved methods for recovering fluorine from phosphate rock would yield an adequate supply of fluosilicates and a substantial contribution to the hydrofluoric acid requirement. THE bulk of natural phosphates is comprised of calcium phosphates, which are usually apatites;' calcium aluminum phosphates such as pseudowavel-1ite; and aluminum phosphates, which occur in extensive deposits, notably on Grand Connetable Island." Nearly all deposits of current commercial value are apatite, often with admixed material of other classes. In the trade coarsely crystalline material, mainly fluorapatite of igneous origin, moves as apatite, whereas the fine-grained material of sedimentary origin is marketed as phosphate rock. The latter classification includes the phosphate shales in western North America. Although several genetic varieties of phosphate rock are recognized, most known reserves are of marine origin. Calcium phosphate is concentrated from Florida land pebble and Florida hard rock deposits and from Tennessee brown rock deposits. Phosphate Rock as a Fluorine Carrier: More than two decades ago Reynolds and Jacob surveyed the fluorine content of commercial phosphate rock from various parts of the world. Later these samples were re-analyzed by a new and more reliable method of determining flourine. The revised results with interim additions of new samples, published a few years ago in a brief summary," are shown in Figs. 1-3. Through the points representing the respective varieties of rock medial straight lines are drawn. In some instances the lines are well defined by the points; in others, where scattering renders the position of the best line uncertain, the greater weight is given points near the end of the covered range. Wherever distinct varieties are not recognized among regional deposits the custom is to identify the sample with the geographical location of the deposit. Although all rock varieties cannot be classified with certainty, the plotted results exhibit three distinct types of rock in which the F to P2O5 ratio diminishes, remains sensibly constant, and increases, respectively, as the grade of rock increases. Existence of two types of Florida land pebble, Fig. 1, is supported by additional results recently supplied by a producer." Generally speaking, phosphate rock from continental deposits carries considerably more fluorine, and apatite and insular phosphates less, in proportion to the phosphorus content than is required by the formula for fluorapatite. Exceptions to this are apatite from Canada and certain samples of Florida waste-pond phosphate. The manner in which excess fluorine is held and the nature of the carbonate in carbonate apatite are discussed in recent articles. High-grade calcium phosphates with F to P2O5 ratios less than 0.05 have been found in Curacao, Christmas Island," Mexic,13 and elsewhere. The calcium aluminum phosphates and aluminum phosphates aften carry notable amounts of fluorine. The F to P2O, ratio for one sample of pseudowavel-lite concentrate was 0.05." Aluminum phosphate from Grand Connetable Island is very low in fluorine, whereas some aluminum phosphate minerals" carry as much fluorine as highly fluorinated phosphate rock. Fluorine in Domestic Phosphate: The curves shown in the figures permit interpolation of the F to P2O5 ratio corresponding to the midpoint of selected ranges in grade of any one of the several varieties of phosphate. Multiplication of the interpolated ratio by the mid-range grade (P2O5) yields a figure for the percentage of fluorine in this grade. Then, with the use of the mid-range grade, the amount of fluorine held in phosphate reserves can be estimated, see Table I. Accordingly, on Jan. 1, 1950, the fluorine complement of known domestic reserves exceeded 420,000,000 long tons of the element. A previous figure," deduced in 1940 by a simpler procedure

Jan 1, 1955

By F. W. Jessen, C. A. Johnson

The effect of treatment with ferrochrome lignosul-fonate on both sodium and calcium bentonites has been examined. In the early stages of treatment it appears that some base exchange of iron and chromium for sodium may occur with the sodium bentonite. Such exchange appears specific for exchange of sodium ions when the calcium ion concentration represents 40-5.5 per cent of the total exchangeable cations, and when additions of lignosulfonate are less than 1 lb per bbl. Adsorption appears entirely physical above 1 lb per bbl treatment level for all concentrations of calcium studied. Gel strength and plastic viscosity of both calcium and sodium bentonite systems were reduced with treatments of ferrochrome lignosulfonate up to 2 Ib per bbl, the greatest reduction occurring within a range of calcium concentration of 40-5.5 per cent of the total exchange capacity of the clay. INTRODUCTION The problem of control of viscosity and gel strength of drilling fluids by the action of thinners or dispersants has been studied by many investigators.1-7 In addition, much work has been done on formation of gel structure in clay suspensions. However, gel formation and structure still are incompletely understood, and the use of the various dispersing chemicals is based, to a large extent, on data almost entirely empirical in nature. In recent years the lignosulfonates have become increasingly important in the treatment and control of drilling fluids. More specifically, ferrochrome lignosulfonate has found rather wide acceptance, particularly in calcium systems.5-7 Lignosulfonates are refined lignin products made from spent sulfite liquor from which the fermentable sugars have been removed.15 The exact structure of lignin is not known; however, it is believed to be a phenolic propane type polymer which may exist as a branched chain or cross linked structure. The molecular weights of lignosulfonates probably range from 300 to 100,000 or more.15 A better understanding of the mechanism of adsorption of dispersants is essential before improvement and development of characteristics of drilling fluids and drilling fluid control is possible. This investigation was undertaken to determine the action of ferrochrome lignosulfonate on sodium and calcium montmOrillonite' LABORATORY INVESTIGATION A Stormer viscosimeter and Fann V-G meter were used to measure viscosity and gel strength. A Zeiss spectrophotometer was used for analysis of Fe and Cr and the flame attachment was employed in the analysis of Na, Ca and Mg. A General Electric X-ray diffraction unit utilizing Copper Ka radiation and a Ni-filter was utilized in the X-ray studies. PROCEDURE Two systems were studied. Predominately sodium clay, obtained from Wyoming, and a predominately calcium clay from Texas were used. These particular samples were chosen because of the extensive work of characterization completed by Mungan. Ionic base exchange capacity determined by flame photometry and total exchange capacity by the Kjehldahl method were obtained experimentally. A 7-per cent suspension of the sodium clay and a 20-per cent suspension of the calcium clay in distilled water were prepared by mixing with a high-speed mixer. The suspensions were allowed to stand for two weeks for complete hydration. Calcium chloride was added to portions of the sodium and calcium cla suspensions to obtain suspensions varying in Ca++ concentration as shown in Table I. A series of treatments was run with 500-ml portions of the original suspensions using a ferrochrome lignosulfonate. Treatment was carried out in the same way with each of the suspensions of varying Ca concentration. Viscosity and gel strength were measured and a filtrate obtained at each treatment level. Ten ml of filtrate were diluted to 50 ml, and aliquot portions of the 50 ml were taken for the Na, Ca, and Mg analysis by the flame photometer. The pH of these filtrates ranged from 7 to 7.3. Standard

By A. L. Lee, M. H. Gonzalez, R. F. Bukacek

Experimental viscosity data for methane are presented for temperatures from 100 to 340F and pressures from 200 to 8,000 psia. A summary is given of the available data for methane, and a comparison is presented for that data common to the experimental range reported in this paper. Correlation of the data is presented, and predicted values are given for temperatures up to 460F and pressures up to 10,000 psia. INTRODUCTION The increasing ranges of temperature and pressure at which fluids are produced, transferred and processed in the petroleum and chemical industries stress the need for accurate information on physical properties, both for engineering calculation and for improving the methods used to estimate physical properties. A survey of the literature reveals that disagreements between published data on the viscosity of methane are common and that most investigations have been conducted over restricted temperature and pressure ranges. This situation made it desirable to undertake an investigation with a scope common to most of the literature data available. APPARATUS AND MATERIALS The apparatus used in this investigation was the capillary tube viscometer'described by Dolan,10 with modifications introduced in the general design and operation of the instrument. It has an effective pressure range from 14.7 to 10,000 psia and a temperature range from room temperature to 400F. The design of the viscometer is based on the establishment of a manometric head between two vessels containing the test fluid and a volume of mercury. The reservoirs are connected by a capillary tube through which the test fluid flows and a tube through which mercury flows. A pressure gradient is established by elevating one of the vessels above the other, the resulting flow of mercury displacing the test fluid through the capillary. The schematic diagram of the system (Fig. 1) shows the arrangement of the equipment auxiliary to the viscometer. The density cells assembly (E, Fig. 1) was not used in this investigation since reliable data are available on methane. 23 The assembly consists of a bank of eight 316 stainless steel pycnometers. Experimental values of the density of a natural gas19 and i-butane14 samples have been obtained for pressures up to 8,000 psia and temperatures up to 340F. The methane used was obtained from the Southern California Gas Co. Mass spectrometric analysis showed a composition of 99.8 percent CH4, 0.1 percent C2H6 and 0.1 percent C3H8. EXPERIMENTAL VALUES AND ACCURACY OF DATA Determinations were carried out at several temperatures. With the temperature in the cabinet controlled at a given level, a series of runs were made at various pressures. To establish pressures lower than that obtained from the methane cylinder, it was necessary to vent the gas slowly to the atmosphere until the desired pressure was reached. The consistency of the test run time for a particular condition can be used as a criterion for the accuracy of the data. A further check is obtained by comparison of the viscosity values obtained using different driving heads. To establish the instrument's reproducibility, a large number of runs were made at three completely different conditions of temperature and pressure. Data obtained were analyzed statistically to determine the most reliable value for the variance of the experimental data. It was found that the variances of the data taken at different experimental conditions were not statistically equivalent, but they were small. The largest variance was found at 8,000 psia and 100F, where the standard deviation was 0.385 micropoise. Therefore, 99.9 percent of the data can be expected to fall within 1.15 micropoises of the population mean, except for the intervention of systematic errors in operation.

By A. S. Michaels, M. C. Porter, A. Stancell

Previous laboratory studies have demonstrated that the injection of small quantities of reverse wetting agents during water displacement can increase oil recovery from unconsolidated porous media.9 In the present investigation, an attempt has been made to determine more fully the effects of reverse wetting treatments and to clarify the mechanism by which increased oil recovery is effected. Water-oil displacements were performed in beds of 140-200 mesh silica sand. Hexylamine slugs (injected after 0.25 pore volume of water throughput), when adequate in size and concentration, were effective in promoting additional oil recovery. Their effectiveness increased with the quantity of amine injected. However, slugs of sufficient size and concentration to stimulate oil production at water flow rates of 34 ft/day did not do so at 4 ft/day. Visual studies in a glass grid micromodel have shown that the stimulation of oil production, via ayueous hexylamine, is a result of transient changes in the oil wettability of the pore walls. if the amine slug is of sufficient size and concentration to induce significant changes in the adhesion-tension, large continuous oil masses will be formed. if the superficial water velocity is high enough to result in rapid desorption of the amine, a favorable "wettobility gradient" may be established across the masses; under such conditions, high oil mobility is observed, and increased oil recovery results. INTRODUCTION It is generally agreed that the efficiency of oil displacement by water in porous media is limited in part by capillary forces which cause the retention of isolated masses of oil — resulting in the so-called "irreducible minimum oil saturation". Recent estimates10 indicate that there are about 220 billion bbl of petroleum in United States reservoirs which are not economically recoverable with present techniques (such as water flooding). This amounts to almost five times the known recoverable reserves. It has been recognized for some time that a suitable alteration in the water-oil interfacial tension and/or the contact angle, as measured between the water-oil interface and the solid surface, should result in better displacement efficiency. Surface active agents can be used as interfacial tension depressants to accomplish this objective,1,2,7 but unfortunately, the additional oil recovery is seldom commensurate with the treatment cost. In contrast to interfacial tension depressants, the effect of contact angle alterations on water-oil displacements has received relatively little attention in the literature. It is known that the wettability affects the displacement process.3,4,11 Displacements in water-wet systems generally result in lower residual oil saturations than those in oil-wet systems. The effect of "'transient" wettability alterations concurrent with the displacement process have been investigated by Wagner, Leach and coworkers,4,11 wherein it has been demonstrated that the establishment of water-wet conditions during water flooding of oil-wet, oil-saturated porous media is accompanied by significant increase in oil displacement efficiency. Michaels and Timminss studied the effects of transient contact angle alterations resulting from chromatographic transport of reverse wetting agents through unconsolidated sand. It was demonstrated that chromatographic transport of short-chain (C4 through C8) primary aliphatic amines can improve oil recovery and that the recovery increases with the quantity of amine injected (i.e., with either the amine concentration or the volume of the slug injected). Circumstantial evidence indicated that the increased displacement efficiency resulted primarily from transient changes in wettability of the porous medium. In the Present investigation. additional information has been obtained on the effects of reverse wetting treatments and the mechanism by which increased oil recovery is accomplished. Hexylamine slugs of varying size and concentration were chromato-

Jan 1, 1965

By A. Styka, G. Fischer, S. Tour

Described herein are the results obtained during research work involving coating titanium alloys with molybdenum by vapor-deposition methods. Results show that the method can be used successfully to deposit hard adherent wear-resistant coating of methodmetallic molybdenum on titanium or titanium alloys without changing the microstructure of the titanium base. PRODUCTION and use of titanium and its alloys have increased tremendously in the past few years. The physical, mechanical, and chemical properties of this engineering metal have caused considerable enthusiasm among metallurgists and engineers. However, as with all materials, there are some drawbacks. For example, titanium is susceptible to seizing and galling on wearing surfaces. In order to produce a good wear-resistant surface, many investigations have been conducted evaluating oxidizing, carburizing, nitriding, carbonitriding, sili-conizing, and chromizing of titanium. All of these processes require high temperatures for rather extensive periods of time. It has been possible to produce hard surfaces on titanium and its alloys by a number of these processes at the expense of a deterioration of the physical properties of the base material. A practical solution to the problem of providing titanium and titanium alloy parts with a satisfactory wear-resistant nongalling surface must be one that does not involve excessive temperatures for long periods of time. Sam Tour & Co. Inc. has completed a research and development contract for the United States Army Ordnance Corps on vapor deposition of molybdenum on titanium and titanium alloys. On the basis of results obtained from the research program, titanium alloys can be coated successfully with molybdenum to provide a hard wear-resistant surface without harm to their base mechanical properties. The remainder of this article is a description of the techniques used to apply the coating and the results obtained from wear tests on molybdenum-coated titanium. Theory Basically, the technique used was the application of molybdenum by a vapor-deposition method, thermal decomposition of molybdenum hexacarbonyl [Mo(CO)] with hydrogen as the carrier gas. The selection of this method was based chiefly on the following considerations: 1—low plating temperature which would minimize excessive grain growth, 2— low decomposition temperature of the carbonyl, 3— possibilities of controlling the hardness of the plate by adjusting the carbon content using the plating temperature and pressure as the primary variables, 4—use of high vacuum, thereby avoiding unnecessary oxidation and other complex reactions with the atmosphere. Free energy data1 show the strong tendency for titanium to combine with O2 N2 or carbon from normal temperatures to temperatures above its melting point. To maintain the titanium in a film free condition, the use of as high a vacuum as possible is indicated. The literature' indicates that the reaction product of hydrogen with titanium is unstable in vacuum at around 400°C. 5—Although it is indicated from the thermodynamic standpoint that there is little possibility of removing carbon from either titanium or molybdenum, the work of Lander and Germer -ndicates that the use of a suitable H,/CO ratio is somewhat effective in minimizing carbon deposition. 6—Availability of molybdenum carbonyl. In the simplest form, the conditions of equilibrium3 in chemical reactions involving decomposition of carbonyl are as follows: Mo (CO) ? Mo + 6CO 2CO * C + CO,. The plating rate is proportional to the rate of evolution of CO from the decomposed carbonyl, i.e., it is proportional to the pressure of the gaseous decomposition products. Carbon released from CO gas is the significant constituent for the variability of the physical properties of the product. High carbon deposition is favored by increasing pressure. Limitations are imposed in the selection of plating temperature and pressure. Excessive temperature re-

Jan 1, 1956

INTRODUCTION Mineral explorationists range throughout the world looking for mineral deposits which can subsequently be made into mines. Mines are not found! Mines are made, usually at great cost and not often without significant risk. The process of making a mineral deposit into a mine involves a great deal more than a complete appraisal of the deposit. The making of a mine demands detailed feasibility studies of many factors, any one of which, if shown to be negative, could prevent the deposit from becoming a successful mine. Total feasibility analysis involves at least three major fields of study: (1) economic analysis, (2) financial analysis, and (3) political and environmental analysis. Economic analysis is essential for proper decision making at many stages in the evaluation of a mineral deposit. It is the yardstick by which the proposed mining project and alter- native investment opportunities are measured. Economic analysis is concerned with costs and potential profits and can be applied at various stages in the evaluation of a mineral deposit to determine if the project merits the additional funds necessary to bring about each succeeding stage. The nearer any deposit approaches to becoming a mine, the more carefully the economic analysis should be made. Financial analysis has reference to how, where, and at what cost the investment funds can be obtained to finance the proposed mine project. Political and environmental analysis has reference to the many intangible factors that can affect the making of a successful mine. These factors loom larger today than ever before, due to the profusion of government regulations and various environmental constraints now imposed upon the mineral industry both by law and by the weight of public opinion. Ultimately, many of these intangible factors can be resolved back into the economic analysis as an added cost to mine development, but more and more often they have become insurmountable obstacles to any development of certain deposits. The role of financial analysis and the analysis of intangible factors must never be underestimated when making a feasibility study upon which to base investment decisions; however, the emphasis of this text will be upon economic factors. THE PARAMETERS OF MINERAL ECONOMIC ANALYSIS The two major elements of economic analysis are: 1) The potential earnings or annual generated cash flow of the project. 2) The investment costs necessary to realize such earnings. If the difference between the two elements is large enough, i.e., if the return on investment from the cash flow is high enough to attract investors, the project can be said to be financially feasible. The mineral economic analysis must take into consideration many parameters, some of which are known or knowable within certain limits, and others of which are unknown or unpredictable within certain limits and for which estimates must be made. Because both known and estimated parameters enter into the same analytical computations, it is important for the analyst to realize the nature of each when considering the results of an analysis. Some parameters such as the grade of mining block, although estimated, are said to be known within mathematically defined limits. Other parameters, such as the future price for minerals produced, etc., can only be estimated or "guestimated" within reasonable limits for which no precision can be assigned. Known or Knowable Mineral Economic Parameters Known or knowable parameters are: 1) Tonnage and grade of valuable minerals for the total mineral body and for individual mining blocks. 2) Metallurgical recovery of valuable mineral. 3) Minimum operating life of mineral de- posit at any specified production rate. 4) Royalty schedule, property payments, etc. 5) Sales-purchasing agreements, smelter schedules, marketing contracts, etc.

Jan 1, 1980

By S. D. Dahlgren

ORDERED G.P. zones having the cesium chloride structure were found to exist in the bcc iron-rich grains of an Fe-Au-Cu alloy that had been supersaturated with approximately equal atomic percentages of gold and copper, and aged at 400° and 350°C. Streaks and superlattice spots in electron diffraction patterns revealed the presence of the ordered zones in the aged material. The ordered Ll2 structure is commonly found for precipitates in fcc metals (e.g., Ni3A1 in nickelL7') but it is uncommon to find precipitates in bcc metals having the cesium chloride structure. The formation of the G.P. zones caused age hardening of the grains. Samples containing 3.8 at. pct each of gold and copper (with the remainder being iron) were annealed at 95@ or 1125°C in quartz capsules and were quenched by breaking the capsules under water. X-ray diffraction patterns revealed two solid solutions in the quenched samples. One solid solution was bcc and the other was fcc. The microstructure consisted of iron-rich grains of 0.3 mm diam surrounded by a yellow gold-copper-rich grain boundary network. The Au-Cu phase diagram %hows continuous solid solubility of gold and copper with ordered phases forming below 410°C. None of the ordered Au-Cu phases have the cesium chloride structure. Both Fe-Au and Fe-Cu phase diagrams3 show extended miscibility gaps and only terminal solid-solution phases. Consequently, it is reasonably certain the iron-rich grains consisted of an a-phase terminal solid solution, and the grain boundary network consisted of a gold-copper-rich terminal solid solution. Lattice parameters of the quenched solid solutions were consistent with the suggestion that the quenched phases were terminal solutions. The lattice parameter of the bcc phase (2.879A) obtained upon quenching from 950°C was higher than that (2.866A) for pure iron. The fcc phase in this sample had a lower lattice parameter than that for a 50 at. pct Au and 50 at. pci Cu solid solution, i.e., 3.866 compared to 3.877A. The lattice parameter for the 50 pct Au and 50 pct Cu solid solution was interpolated from data for Cu-Au solid solutions.4 More gold and copper were soluble in iron at 1125°C than at 950°C as indicated by a 0.012A increase in the bcc lattice parameter for a sample quenched from 1125°C. After the alloy had been quenched from 950° and 1125°C, the bcc lattice parameters were 2.879 and 2.891A1 respectively G.P. zones having the cesium chloride structure-which formed during low-temperature aging—were detected by electron diffraction and observed with electron transmission microscopy. In Fig. 1, the asymmetrical streaks parallel to ( 100) directions indicate that platelike G.P. zones had formed parallel to (100) planes of the matrix. The direction of streak asymmetry shows that the bcc unit cell of the G.P. zones was larger than the bcc cell of the matrix. The ordered cesium chloride structure is indicated by low-intensity superlattice spots at 010 positions which are nonallowed for bcc structures. The superlattice spots are streaked; therefore they are from the G.P. zones and are not caused by rel-rod effects of a thin matrix phase. The appearance and size of the zones observed in electron transmission microscopy and the appearance of the asymmetrical streaks in the electron diffraction patterns for the Fe-Au-Cu alloy were the same as those observed for a similarly treated Fe-Au alloy.' Ordered G.P. zones were detected also for an Fe-Au-Cu sample aged only at 400°C. The production of the ordered cesium chloride structure zones by low-temperature aging of super-

Jan 1, 1968

By C. R. King

Most of the acid volcanic glasses such as obsidian, perlite, pitchstone, pumice, and pumicite (volcanic ash) are susceptible to some expansion if suddenly subjected to a suitably high temperature in a properly designed furnace, and subsequently suddenly cooled. Expansion results from the action of chemically combined or dissolved water expanding as a vapor within the softened heated rock particles. This expansion forms countless minute bubbles within the pasty mass of the natural glass particles. If the particles are cooled before these bubbles can escape, the resulting products are cellular, more or less spherulitized glass froths or hollow glass spheres of many types. Chemically, they are essentially aluminum silicates containing about 70 to 75 pct silica, 10 to 16 pct alumina, and 2 to 7 pct sodium plus potassium oxides. They are relatively resistant to attack by acids or alkalies, are resistant to heat up to over 1000°F, and are completely inert to insect or fungus attack. These products are of great potential value in many branches of industry as insulation materials, aggregates, fillers, abrasives, and additives. By proper selection of the perlite rock type and control of processing conditions, expanded perlite of a bulk density varying from about 2 to 20 Ib per cu ft can be • made. The physical properties of the particles comprising a product of a given bulk density also may be varied within wide limits of porosity, water absorption, and crushing strength. In general, crushing strength of expanded perlite increases in direct relation to increase in bulk density, but insulating value varies in inverse ratio to bulk density. The selection of a particular grade of expanded perlite for a given use as an aggregate is therefore always a compromise between insulating value, lightness, and friability. If porosity and water absorption are factors affecting the use of the material, factors other than bulk density are important. Some types of perlite rock tend to decrepitate excessively during expansion, resulting in a highly porous or even hygroscopic product of high friability relative to the bulk density of material produced. In general, treatment of any good perlite rock below the temperature necessary for maximum expansion, or insufficient time in the hot zone of the furnace for effective heat penetration to the core of the rock particles will result in some decrepitation and high porosity in the expanded product. On the other hand, exposure of the material to the temperature required for optimum expansion for a period of time slightly longer than that necessary for proper heat absorption will result in a slightly heavier product of extremely low porosity, i. e., a glazed surface. It is beyond the scope of this paper to describe in detail the rock types and processing methods used to produce the many types of expanded perlite in present use, or that may be produced to fill the present and future needs of industry. A few phases of the economics involved in processing and selling products made from this abundant and cheap mineral raw material, plentifully distributed in large and relatively accessible deposits in most western states and in other areas within economic shipping distance of large centers of consumption, will be treated with particular reference to the Los Angeles area. Relation of Balk Density to Use and Cost of Production Perlite differs from most other lightweight aggregates and fillers in that the raw material (perlite rock) may be quarried much as rock arid sand is mined, and then shipped into consuming areas as crude rock in open gondola cars at correspondingly low freight rates and handling costs. At processing plants in or near large consuming areas the rock is converted into relatively fragile, very bulky products which in most cases are used within a short haul of the processing plant or enter into a manufactured product made at the site of the perlite processing plant.

Jan 1, 1950

By W. Rostoker, H. Nichols

It has been demonstrated in previous work1'2 that wetting of aluminum alloys by liquid mercury can cause fracture to occur with substantial suppression of prior plastic flow. This has been interpreted as the action of a critical liquid species in reducing the cohesive strength of a stressed solid at sites of potential crack nucleation. In this way the critical transverse stress for fracture nucleation is reached at much lower stresses than is normally the case. The stresses for both fracture nucleation and fracture propagation are depressed by the same mechanism. In so doing the plastic-energy component of fracture energy is reduced to a level which is of the same order of magnitude as the surface, interfacial, or cleavage energy. Recent work2 has illustrated the influence of dispersed precipitates, cold work, and combinations of these on brittle fracture associated with plastic strains of the microyield order of magnitude. This has been rationalized in terms of the contribution of internal microstress fields to the stress magnifications at points of obstruction of active slip bands in a stressed metal. The presently reported studies on an Al-4-1/2 pct Mg alloy amplify this subject by introducing thermal-mechanical conditions which can produce strain aging. As amply demonstrated by Phillips, Swain, and borall, the tensile behavior of aluminum alloys containing magnesium reveals a yield-point elongation and successive discontinuous elongations over the whole plastic-strain range. This gives a stepped or serrated character to the stress-strain diagram. The yield-point elongation is not temperature-dependent. On the other hand, the discontinuous extensions in the plastic-strain range are gradually suppressed by lower temperatures of testing. At -78"C, only the yield-point elongation persists. The yield-point elongation itself can be suppressed by rapid quenching to room temperature from an annealing temperature of 500°C. The argument presented by Phillips et 1. and Cottre114 is that the yield-point elongation derives from surmounting at a critical stress the restraint to slip by grain boundaries which have been reinforced by appreciable magnesium-atom segregation. The serrated character of the plastic-strain range is the result of rapid strain aging during the test. This is made possible by the inevitable proximity of magnesium atoms to any dislocation line. In the Al-4-1/2 pct Mg alloy there is one magnesium atom for every twenty-two aluminum atoms and, with a coordination number of 12, a magnesium atom is on the average only about two atomic distances from any given lattice site. Rapid strain aging is therefore a reasonable expectation. A study of mercury embrittlement of the Al-4-1/2 pct Mg alloy (commercial designation 5083) provides an opportunity to relate the interplay of grain size, grain boundary restraint to slip and strain aging in the brittle-fracture process. One lot of commercial sheet material of 0.125-in. thickness was used for all of the tests. The specimen shape and test procedures used were the same as those described in a previous publication.2 A range of grain sizes between 0.0057 and 0.0125 mm was produced by various combinations of plastic prestrain and annealing. The coarsest grain was somewhat elongated in shape. In this case the measurement was taken in the direction transverse to the axis of tension as more meaningful to a brittle-fracture process. In the annealed condition, specimens wetted with mercury failed with zero permanent elongation at stresses slightly below the normal yield point, i.e., stresses at 0.05 to 0.2 pct strain. Since this occurs before yield-point elongation can begin, fracture derives from slip contained within individual grains and restrained from propagating across grain boundaries. No significant difference in wetted fracture stress was observed between specimens which had been water-quenched or furnace-cooled from the annealing temperatures. This signifies that the degree of magnesium segregation at grain boundaries

Jan 1, 1964

By B. L. Averbach, M. Cohen, C. H. Shih

The isothermal formation of martensite is studied in Fe-Ni-Mn and Fe-Mn-C alloys under conditions where the athermal transformation is completely avoided, there being no martensite present at the beginning of the isothermal reaction. The initial rate of martensite formation is extremely low under these conditions, but it increases greatly as soon as some martensite is formed. It appears that true initial characteristics of the isothermal reaction can be studied only in samples wherein the formation of martensite on cooling is completely eliminated. THE phenomenon of isothermal martensite formation is well established1 -" but the observations have been made in the presence of varying quantities of athermal martensite. Complete avoidance of athermal martensite and the formation of martensite under pure isothermal conditions have been reported by Kurdjumov and Maksimova2 in an Fe-Ni-Mn alloy (23 pct Ni, 3.4 pct Mn, and the balance Fe) and have been confirmed by recent experiments of Cech and Hollomon in a similar alloy. In each case, the isothermal transformation rate was a maximum at the beginning of the transformation and decreased as the reaction proceeded. These initial transformation rates were used by Fisher7 for the confirmation of calculated nucleation rates based on classical nucleation theory extended to include the effect of elastic strain energy. Cech and Hollomon have admitted, however, that up to 3 pct martensite formed on cooling in their samples and thus was present at the initiation of their isothermal runs. In this paper, experiments on similar alloys are described, except that care was taken to remove the incidental martensite developed at the surface prior to the isothermal reaction. It is shown here that the reaction rate at the beginning of the strictly isothermal reaction is very low. As soon as some detectable amount of martensite is produced, the transformation rate increases rapidly but ultimately decreases as the reaction proceeds. When martensite is allowed to form isothermally in the presence of about 2 pct of prior martensite, the apparent initial reaction rate is a maximum and corresponds to the observations of Kurdjumov and Maksimova and Cech and Hollomon. It is evident that the cooperative nature of the martensitic reaction results in a significant stimulating or autocatalytic effect on the isothermal reaction rate. Experimental Procedure The chemical compositions of the alloys used in this investigation are listed in Table I. The Fe-Ni-Mn alloys A and B were melted in an induction furnace, cast into ingots 2 in. diam by 4 in. long, and forged into % in. diam rods. The manganese steel (alloy C) was melted in an arc furnace and cast into ceramic molds 1/4 in. diam by 5 in. long. The rods were sealed in evacuated Vycor tubes and homogenized at 1175" to 1200°C for 50 hr. The long homogenization resulted in some manganese loss from the surface, producing a thin layer of transformed material. This surface layer was removed by grinding and the specimens were then swaged to 0.050 in. diam wire. Measurements of electrical resistance were made with a Kelvin double bridge to follow the course of the martensitic transformation. A small permanent magnet was also useful in scanning an entire specimen for small localized quantities of martensite. This test is almost as sensitive as the electrical resistance method in detecting the presence of martensite (which is ferromagnetic in these alloys) and it was especially useful in deciding whether a specimen was completely free of martensite. The wire specimens (0.050 in. diam by 2.5 in. long) were sealed in evacuated tubes and austenitized for 5 min at 1150°C. The specimens were quenched into water and, unless otherwise stated, stress-relieved for 1 hr at 650 °C in order to induce greater uniformity in the subsequent martensitic transformation. As dem-

Jan 1, 1956

By Milton F. Goodrich

The idea of using water-only cyclones to scalp clean coal from the feed to other devices has recently been gaining in popularity. 1,2,6 An indication of this popularity is that water-only cyclone scalping is already employed in a few commercial plants. Water-only cyclones scalp the feed to a heavy media vessel in the Warwick plant,3 to Deister tables in the Rowland plant,4 and to a heavy media cyclone circuit in the Amherst Coal Company's McGregor plant.5 Despite this trend, we found surprisingly little information in the literature on the performance of water-only cyclone scalping. Therefore, we undertook a study of its technical merits, limiting our investigation to a water-only cyclone scalper in conjunction with heavy-media cyclones. We felt this to be a reasonable combination of equipment, since both processes operate well on the same size coal. First, we developed simple mathematical relationships revealing general characteristics of the system. Next, we compared the clean coal yield and quality obtainable by the system on four different coals with that obtainable by heavymedia cyclone circuit alone and water-only cyclones middlings recirculation circuit alone. This was done by using computer program to simulate the operation of the three systems. Finally, we considered the effect that implementation of the system would have on a typical preparation plant circuit. The System The system chosen for study [(shown schematically in Fig. 1)] comprises two stages: a primary stage (PI) consisting of water-only cyclones, and a secondary stage (P2) consisting of heavy media cyclones. The water-only cyclones separate raw coal at a low specific gravity, producing a relatively small amount of high quality clean coal and a refuse containing much good coal. This refuse is then recleaned in the secondary stage, where the sharper separation of heavy media cyclones, operated at a higher specific gravity, reclaims additional clean coal. Finally, the clean coal from both stages is combined as the system's clean coal, while the refuse from the secondary stage becomes the system's refuse output. Theoretical Analysis Two assumptions lie behind the following theoretical analysis. The first assumption is that partition curves can represent not only the performance of each stage but also the performance of the system as a whole. These curves show partition number as a function of specific gravity, where partition number is either the percent or fraction of a raw coal with a given specific gravity reporting to clean coal. The second assumption is that these curves are independent of the raw coal characteristics. At any specific gravity the system's partition number is solely a function of the partition numbers of each stage. For convenience, we use the following variables in describing this relationship: g = specific gravity, PS(g) = the partition number of the system at specific gravity g, P1(g) = the partition number of the primary stage at specific gravity g, P2(g) = the partition number of the secondary stage at a specific gravity g. For this system, we can show algebraically that the relationship is: [Ps (g) = PI (g) + (1- P1(g)) P2 (9)(1)] With this equation the system's partition curve can be determined if the partition curves of each stage are known. By definition, the partition numbers P1(g) and P2(g) must each be less than, or equal to, one. Placing these constraints on Eq. (1), we arrive at the relations: [Ps(g) % P1(g) and Ps(g) % P2(g)(2)] Thus, at any specific gravity the system's partition number is always greater than that of both stages. [Figure 2] shows typical partition curves. Curve I represents the primary stage, which consists of water-only cyclones operating at a separating gravity of 1.3. (The separating gravity or the

Jan 1, 1979

By S. Weinig, E. S. Machlin

IN a previous study of the grain boundary stress relaxation phenomenon,' the authors had arrived at the conclusion that two successive steps were involved in the complete relaxation of stress at a grain boundary. These two steps were grain boundary migration and grain boundary slip. It was further suggested that the slower of these two processes is rate controlling. In an attempt to ascertain the validity of the above, a study of grain growth kinetics on the same alloys used for the previous study was undertaken. Separate wire specimens, 0.030 in. diam, in the same as-deformed state (99 pct reduction in area) as in the previous investigation were heated to the various temperatures for a sequence of times. The compositions used and the temperatures of testing are summarized in Table I. A statistical determination of grain size was then made on each specimen, using the average length of traverse between boundaries as a measure of the grain size. Examples of typical results obtained are shown in Figs. '1 and 2. The grain growth law D = ktn was found to fit most of the data. Using this relation, values of the grain growth exponent n were calculated and the inter -esting dependence of n on solute content shown in Figs. 3 and 4 was then obtained. Also, it was found that the constant K which was obtained from extrapolation is roughly independent of solute content. (K is uncertain to within ±25 pct.) An estimate of the activation energy for grain growth was made using the technique of Burke? This method utilizes the plot of the reciprocal of time necessary for the grain size to increase from a definite starting size Do, to a final size D, for each isothermal heat treatment. The slope of this curve corresponds to the heat of activation for the process. In Fig. 5 the rate of grain growth is plotted against the reciprocal of absolute temperature. The activation energies were obtained from these curves and are summarized in Fig. 6. Discussion It is interesting to note that the same saturation in the variability with concentration has been found for the grain growth exponent n in this research as was found in the previous researches on both grain boundary stress relaxationa and internal friction of dislocation origin." The significant feature is that saturation in the variation of these properties has been found to occur at about the same level of solute concentration, namely, between 0.1 and 0.5 atomic pct. The only characteristic common to these different quantities is that solute adsorption at grain boundaries and dislocation can take place and affect these values in a similar fashion. Thus, it is concluded that the grain growth exponent n depends upon solute adsorption at the grain boundaries. It appears, therefore, that theoretical considerations leading to the contrary conclusions5,6 have not treated the real process by which solute atoms exert their effect on grain growth. Prior to the analysis of the observed activation energies, certain of the aspects of the aforementioned grain boundary stress relaxation investigation will be briefly reviewed. It was found that with the addition of solute atoms the grain boundary stress relaxation peak (internal friction vs temperature) decreased in intensity. Simultaneously a second peak was observed which increased with increasing solute atoms. Both of these peaks were shown to be the result of grain boundary stress relaxation. They have been termed, respectively, poire peak and solute peak. It was suggested that the solute peak was dependent upon grain boundary migration, whereas the pure peak was a manifestation of grain boundry slip. Activation energies for both peaks were obtained as a function of Solute content. The value of the activation energy in the temperature range corresponding to the solute peak was

Jan 1, 1958

By H. L. Overton, J. B. Lipson

The concept of sodium single-ion equivalent activity as developed by Gondouin, Tixier and Simard,' was used to determine the filtrate resistivity-activity relationships for 150 laboratory and 49 field drilling muds. With the exception of the native and phosphate high-resistivity muds and the calcium-treated muds, the filtrate resistivity-activity relationships were found to be within 20 per cent of that for pure sodium-chloride solutions. Gondouin, et al,* previously suggested a method of treating the high-resistivity muds in quantitative electrical log analysis. A new treatment is presented which uses the P-alkalinity of the filtrate as a correlating parameter in handling the calcium-treated muds in quantitative electrical log analysis. INTRODUCTION Gondouin, Tixier, and Simard in 1957 introduced new interpretive procedures for analyzing the self-potential curve of electric logging. They re-emphasized the previous work of the Schlumberger brothers, Wyllie and others, and re-defined the activity concept in terms of a number which they termed the effective resistivity. Their procedures are particuarly effective when the connate-water resistivities are less than 0.08 ohm-m or greater than 0.3 ohm-m. For gyp muds where the calcium content is known, a procedure was presented which allows calculation of the connate-water resistivity with improved accuracy. These procedures assume that one of the fluids in the system is a pure sodium-chloride solution, but they are not limited by the assumptions. For the fresh and saline connate waters, for instance, it is assumed that the drilling mud filtrate is a pure sodium-chloride solution. For the gyp mud interpretive procedure, it is assumed that the connate water is a pure sodium-chloride solution. Brown,' on the other hand, has suggested that it would be extremely fortuitous if drilling mud filtrates could all be treated electrochemically as sodium-chloride solutions, and he cites some examples where discrepancies exist. A study has been made of the deviations of sodium single-ion equivalent activities, an of drilling mud filtrates from that of pure sodium-chloride solutions. Drilling fluids in common use on the Texas and Louisiana Gulf Coast and laboratory-controlled muds were investigated. Using a cell patterned after Gondouin, et al, (Fig. I), the equivalent activities were measured. It should be remembered here that the activity measured is merely an index which uses numbers that represent the sodium single-ion equivalent activity a,., of the mud filtrate. The membrane potential developed across the cell was determined with a null-type potentiometer. The resistivities of the solutions were measured by the four-electrode method. The resistivity system was calibrated by using saturated KC1 solution as a standard. Chemical properties of the drilling fluids studied were measured with standard API-approved equipment for the particular property in question. For the activity measurements, two shale discs were chosen to serve as cationic membranes. These two discs were both 1 1/4 in. in diameter and, respectively, 1/4 and 1/8 in. in thickness. The membrane potential, Em, was determined using the shale discs as cationic selective membranes separating two standard sodium-chloride solutions of known mean activity. Then by comparing the constant Km in the following equation, with the theoretical value of — 59 mv at 75OF, the perfection of the membrane was determined. The variation of Km for the two membranes used in this experimental work, as well as one of the many unsuitable membranes, is shown in Fig. 2. The measurement of the activity of mud filtrate involves the balancing of its combined individual activities against a standard solution of known mean activity. At the point where there exists zero potential across the membrane cell, the activity effects may be assumed equal for all practical purposes. The mechanics of making a null reading of this sort requires having a large number of different mean activity solutions. A more practical method involves

By P. J. Root, J. E. Warren

An idealized model has been developed for the purpose of studying the characteristic behavior of a permeable medium which contains regions which contribute significantly to the pore volume of the system but contribute negligibly to the flow capacity; e.g., a naturally fractured or vugular reservoir. Unsteady-state flow in this model reservoir has been investigated analytically. The pressure build-up performance has been examined in some detail; and, a technique for analyzing the build-up data to evaluate the desired parameters has been suggested. The use of this approach in the interpretation of field data has been discussed. As a result of this study, the following general conclusions can be drawn: 1. Two parameters are sufficient to characterize the deviation of the behavior of a medium with "double porositym from that of a homogeneously porous medium. 2. These parameters can be evaluated by the proper analysis of pressure build-up data obtained from adequately designed tests. 3. Since the build-up curve associated with this type of porous system is similar to that obtained from a stratified reservoir, an unambiguous interpretation is not possible without additional information. 4. Differencing methods which utilize pressure data from the final stages of a build-up test should be used with extreme caution. INTRODUCTION In order to plan a sound exploitation program or a successful secondary-recovery project, sufficient reliable information concerning the nature of the reservoir-fluid system must be available. Since it is evident that an adequate description of the reservoir rock is necessary if this condition is to be fulfilled, the present investigation was undertaken for the purpose of improving the fluid-flow characterization, based on normally available data, of a particular porous medium. DISCUSSION OF THE PROBLEM For many years it was widely assumed that, for the purpose of making engineering studies, two param- eters were sufficient to describe the single-phase flow properties of a producing formation, i.e., the absolute permeability and the effective porosity. It later became evident that the concept of directional permeability was of more than academic interest; consequently, the degree of permeability anisotropy and the orientation of the principal axes of permeability were accepted as basic parameters governing reservoir performance. 1,2 More recently, 3-6 it was recognized that at least one additional parameter was required to depict the behavior of a porous system containing regions which contributed significantly to the pore volume but contributed negligibly to the flow capacity. Microscopically, these regions could be "dead-end" or "storage" pores or, macroscopi-cally, they could be discrete volumes of low-permeability matrix rock combined with natural fissures in a reservoir. It is obvious that some provision for the inclusion of all the indicated parameters, as well as their spatial variations, must be made if a truly useful, conceptual model of a reservoir is to be developed. A dichotomy of the internal voids of reservoir rocks has been suggested.7,8 These two classes of porosity can be described as follows: a. Primary porosity is intergranular and controlled by deposition and lithification. It is highly interconnected and usually can be correlated with permeability since it is largely dependent on the geometry, size distribution and spatial distribution of the grains. The void systems of sands, sandstones and oolitic limestones are typical of this type. b. Secondary porosity is foramenular and is controlled by fracturing, jointing and/or solution in circulating water although it may be modified by infilling as a result of precipitation. It is not highly interconnected and usually cannot be correlated with permeability. Solution channels or vugular voids developed during weathering or burial in sedimentary basins are indigenous to carbonate rocks such as limestones or dolomites. Joints or fissures which occur in massive, extensive formations composed of shale, siltstone, schist, limestone or dolomite are generally vertical, and they are ascribed to tensional failure during mechanical deformation (the permeability associated with this type of void system is often anisotropic). Shrinkage cracks are the result

By W. S. Hyler, W. L. Bruckart

The powder metallurgy and arc-cast types of wrought molybdenum stock were studied in rotating beam fatigue. Endurance ratios of unnotched specimens after 5x10 cycles were found to be 0.74 and 0.81, respectively. Fatigue notch factors (K,) of 1.36 and 2.05, respectively, were obtained on V-notched specimens having a theoretical stress-concentration factor Kt = 3.1. MOLYBDENUM and its alloys are rapidly gaining importance as engineering materials of construction for high temperature applications. AS a wrought material, molybdenum and its alloys are outstanding above 1800°F. It is necessary, however, with molybdenum to provide elevated temperature oxidation protection. This is currently being studied by a number of investigators. Thus far, there have been no published data on the fatigue behavior of molybdenum either at room temperature or at elevated temperature. The knowledge of fatigue behavior, particularly notched fatigue behavior, becomes increasingly important as the engineering importance of a material increases, especially from the design standpoint. The work described here was directed toward determination of the notched and unnotched rotating beam fatigue characteristics of molybdenum at room temperature. Also, it was desired to learn whether any difference exists between wrought molybdenum of the powder metallurgy and the are-cast types. Fabrication History of Bar Stock For this work, bar stock was procured from Climax Molybdenum Co. (are-cast type) and from Westing-house Electric Corp. (powder metallurgy type). The processing histories of these materials are listed in Table I. The first lot of are-cast molybdenum was deoxidized by aluminum additions. The second lot was deoxidized by carbon." Both lots of the powder Microstructure A microscopic examination of these materials, as received, revealed the powder metallurgy molybdenum to be free of inclusions and to be finely fibered. On the other hand, the are-cast C molybdenum possessed numerous carbide inclusions, and the arc-cast A appeared to have coarse grained fibers. These are shown in Fig. 1. Owing to the large apparent difference in fineness of fiber between the are-cast A and the powder product and are-cast C, a check was made of the recrystallized grain sizes to determine whether these differences persisted. A series of recrystallization anneals revealed mixed grain sizes in both the are-cast A and the powder product. Variations in grain diameter of as much as 50 times were found after full recrystallization at 1200°C for 30 min. There was no distinction between the two. Arc-cast C was not checked. Room Temperature Mechanical Properties Prior to subjecting the materials to fatigue tests, room temperature mechanical properties were determined. These data, together with the tensil: strength of the notched bars, data,togetherare listed in Table tensil: There were no significant differences in the tensile strengths of the two types of molybdenum. However, the powder metallurgy material was generally more ductile and yielded at a somewhat lower stress than the arc-cast material. Tensile strengths of the notched specimens indicated a slightly greater notch strengthening effect in the are-cast C than in the powder product, lot 2. Room Temperature Fatigue Properties Sample Preparation: Specimens were prepared for use in R. R. Moore rotating beam fatigue machines. These specimens had a continuous radius test section with a minimum diameter of 0.260 in. The specimens were prepared by grinding to shape and polishing to produce a smooth unworked surface. Polishing consisted of rotating a disk in the continuous radius section on which were attached, successively, emery belts of 240, 400, and 600 grit. The specimens were slowly rotated during the operation. The belt speed was high, and rotation was such that polishing scratches were longitudinal. Unnotched Fatigue Properties: Unnotched fatigue data were obtained from a number of specimens of each of the four lots of molybdenum. The results of the tests are presented in Fig. 2. From this figure,

Jan 1, 1956