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By O. Jay Myers

The war effort has furnished the necessary impetus for better magnesium foundry practice. Four or five years ago, there were but a few formulas in general use for core-sand mixtures for magnesium castings. Certain sands, binders, and inhibitors were specified and the proportions of these materials in each mixture were maintained between very narrow limits. As more and more war industries entered the field, much research was carried out and methods and practices were improved. Magnesium founding is still in a state of flux; therefore it is not uncommon to find one foundry spraying inhibitors on its cores, and another foundry using sand mixtures with incorporated protective agents. The basic core or dry-sand mold mixture for magnesium in use today is compounded with sand, core oil or resin binder, cereal, and moisture. Core oil or resin binder are added for baked strength while the cereal binder and moisture are for green strength. The normal mixtures (based on the weight of the sand) contain from o.5 to 1.5 per cent of both cereal binder and core oil, and the moisture content variesbetween 1.5 and 5 per cent. The ratio of core oil to cereal binder varies from 60:40 to 40:60, while the correct amount of moisture depends upon the method of coremaking. Protective Agents In all magnesium practice, protective agents are used in conjunction with the core and molding sands to prevent oxidation of the metal. The use of these inhibitors is not novel, but publications covering this field are not entirely adequate. Beck1 says: Magnesium alloys decompose water to form magnesium oxide, hydrogen being liberated in the process (Mg + H2O ? H2 ?+ MgO). This reaction results in the blackening of the skin and the appearance of areas of porosity on the casting surface, the cavities of which are filled with a grey oxide powder (these are termed 'burns'). Although sand cores for magnesium castings are thoroughly dried during the baking cycle, burning will take place on the casting5 unless the metal is protected properly. This type of burning is more difficult to define than the "water-magnesium" reaction cited by Beckl A possible explanation may lie in the following equation: 2Mg + O2—> 2MgO Here, the source of the oxygen is not water, but the air in the mold cavity, together with the chemically combined oxygen in the core-binding media. Generally, two methods are used to prevent this reaction from taking place.

Jan 1, 1945

By E. S. Messer

A knowledge of the magnitude of the irreducible inter.;titial water in a porous medium is so important to petroleum engineering that its determination has become routine in core analyses. The method of determination, being a production problem, should encompass the basic requirements of simplicity in technique and calculations, with reproducible results obtainable in a short interval of time. The results of the evaluation tests outlined in this report indicate that the evaporation method for determining the irreducible water is a technique which meets the requirements. The procedure consists, as the name implies. of permitting the saturant in the pore spaces to evaporate until only an irreducible volume remains. The determination of this volume can be made either graphically or by a mathematical comparison of fluid flows; the time required for each determination being dependent on the fluid used. When fluids other than those having reservoir characteristics were used, a volume factor had to be calculated which was based on the relative volume of various liquids adsorbed on grain surfaces and retained in pores. This factor made possible the calculation of an irreducible water volume when more volatile fluids such as toluene and benzene were used as the saturants. Also presented is the theoretical discussion necessary for the calculation of the capillary pressure as determined from the evaporation curve. A comparison is made between the calculated values and those obtained by experimental means. INTRODUCTION In all geological formations there exists, in the pore spaces of the rock structure, water that is held in a state of equilibrium between capillary and hydrostatic forces. "Interstitial water" is the term given to this water and is defined as that water coexisting in the pore space with the oil prior to exploitation. The term ''connate water" has often been used synonymously with this term; however, this can be true only by a specific definition since, geologically, it means the water in place at the time the rock structure was formed. The quantity of the interstitial water is a variable factor in any formation, since it depends on the hydrostatic forces present in any multiple-phase system. These forces may become unbalanced by the introduction of an extraneous force such as the raising or lowering of the "water table" or the migration of oil into a water-filled formation. Any unbalanced force results in a change in the interstitial water. There exists, however, an irreducible interstitial water. for a particular sand, that is the fraction of the pore space occupied by water when the capillary pressure at the particular point in question is at an equilibrium with the hydrostatic head of the oil sand in the reservoir. For this discussion the term "irreducible water saturation" will be used in place of "irreducible interstitial water saturation" for the sake of brevity; however, they are understood to be identical. A great amount of work has been devoted to the theory and methods for studying the irreducible water saturation and its related capillary pressure. As a result of the publications of Leverett;' Hassler, Brunner and Deahl;2 Calhoun and Lewis;3 and others, the role of capillary pressure studies is being accepted by the industry as a tool for studying suhsurface phenomena. Many techniques have been developed and published for determining the capillary pressure and irreducible water. In general, these techniques may be grouped into three classifications. One of the first was the capillary pressure method described by Leverett1 and expanded by Bruce and Welge.4 The experimental results were compared with water saturation of cores obtained using oil-base mud. Thornton and Marshall compared the irreducible water saturation of core samples determined by the capillary pressure method and by salinity and reported good agreement between the two methods. The second classification for determining the irreducible water and capillary pressure may be referred to as the "centrifugal force method." The general technique is similar to the capillary pressure method except that the force driving the reservoir fluid from the sample is of a centrifugal nature. A complete description of this method was presented by J. J. . McCullough and F. W. Albaugh.6 A process, the reverse of the capillary pressure method, was presented by W. R. Purcell.7 Mercury under pressure is driven into the pores of the rock and the saturation of the core determined at each applied pressure. The resulting capillary pressure curve is used to evaluate the irreducible water saturation. The techniques mentioned are singular in their approach to the irreducible water saturation. In all cases. an external force was applied to the core. The forces employed in the evaporation method are the vapor pressure of the liquid causing evaporation, the kinetic diffusion forces. adsorptive forces and. to a lesser degree, the viscous forces resisting flow to the surface. The basic definition of irreducible water is that water held in a state of equilibrium between capillary and hydrostatic forces This water has been described by previous investigators as being held in the microcapillaries too small to support fluid flow. Actually, this fluid volume is made up of the water in the microcapillaries and as a film adhering to the surface of the crystals. All capillaries. therefore, possess some liquid as a film, the thickness of the film being dependent on the properties of the fluid and solid. A discussion of experiments with references pertaining to the measurement of this immobile layer next to the solid surface can be found in the text by J. J. Bikerman.8 Eversole and Lahr calculated the thickness of this layer to be in the order of 10 ' to 10' cm for aqueous solutions and glass. Between two quartz surfaces they found the thickness to be 2 x 10 cm. The work of Volkova, on the capillary movement of water and toluene in quartz grains, indicated the thickness of the Immobile layers to be near 10' cm. Since any measurement is an average value, it is easy to understand that an absolute value would depend on the roughness of the surfaces involved and the complexity of the system. A calculated effective pore radius of 2 x 10 cm is obtained at the, irreducible saturation of a porous media in a water-air system when a capillary pressure of 100 psi is applied. Since the separation of the sand grains is of the same approximate magnitude as the immobile layer.

Jan 1, 1951

By E. S. Messer

A knowledge of the magnitude of the irreducible inter.;titial water in a porous medium is so important to petroleum engineering that its determination has become routine in core analyses. The method of determination, being a production problem, should encompass the basic requirements of simplicity in technique and calculations, with reproducible results obtainable in a short interval of time. The results of the evaluation tests outlined in this report indicate that the evaporation method for determining the irreducible water is a technique which meets the requirements. The procedure consists, as the name implies. of permitting the saturant in the pore spaces to evaporate until only an irreducible volume remains. The determination of this volume can be made either graphically or by a mathematical comparison of fluid flows; the time required for each determination being dependent on the fluid used. When fluids other than those having reservoir characteristics were used, a volume factor had to be calculated which was based on the relative volume of various liquids adsorbed on grain surfaces and retained in pores. This factor made possible the calculation of an irreducible water volume when more volatile fluids such as toluene and benzene were used as the saturants. Also presented is the theoretical discussion necessary for the calculation of the capillary pressure as determined from the evaporation curve. A comparison is made between the calculated values and those obtained by experimental means. INTRODUCTION In all geological formations there exists, in the pore spaces of the rock structure, water that is held in a state of equilibrium between capillary and hydrostatic forces. "Interstitial water" is the term given to this water and is defined as that water coexisting in the pore space with the oil prior to exploitation. The term ''connate water" has often been used synonymously with this term; however, this can be true only by a specific definition since, geologically, it means the water in place at the time the rock structure was formed. The quantity of the interstitial water is a variable factor in any formation, since it depends on the hydrostatic forces present in any multiple-phase system. These forces may become unbalanced by the introduction of an extraneous force such as the raising or lowering of the "water table" or the migration of oil into a water-filled formation. Any unbalanced force results in a change in the interstitial water. There exists, however, an irreducible interstitial water. for a particular sand, that is the fraction of the pore space occupied by water when the capillary pressure at the particular point in question is at an equilibrium with the hydrostatic head of the oil sand in the reservoir. For this discussion the term "irreducible water saturation" will be used in place of "irreducible interstitial water saturation" for the sake of brevity; however, they are understood to be identical. A great amount of work has been devoted to the theory and methods for studying the irreducible water saturation and its related capillary pressure. As a result of the publications of Leverett;' Hassler, Brunner and Deahl;2 Calhoun and Lewis;3 and others, the role of capillary pressure studies is being accepted by the industry as a tool for studying suhsurface phenomena. Many techniques have been developed and published for determining the capillary pressure and irreducible water. In general, these techniques may be grouped into three classifications. One of the first was the capillary pressure method described by Leverett1 and expanded by Bruce and Welge.4 The experimental results were compared with water saturation of cores obtained using oil-base mud. Thornton and Marshall compared the irreducible water saturation of core samples determined by the capillary pressure method and by salinity and reported good agreement between the two methods. The second classification for determining the irreducible water and capillary pressure may be referred to as the "centrifugal force method." The general technique is similar to the capillary pressure method except that the force driving the reservoir fluid from the sample is of a centrifugal nature. A complete description of this method was presented by J. J. . McCullough and F. W. Albaugh.6 A process, the reverse of the capillary pressure method, was presented by W. R. Purcell.7 Mercury under pressure is driven into the pores of the rock and the saturation of the core determined at each applied pressure. The resulting capillary pressure curve is used to evaluate the irreducible water saturation. The techniques mentioned are singular in their approach to the irreducible water saturation. In all cases. an external force was applied to the core. The forces employed in the evaporation method are the vapor pressure of the liquid causing evaporation, the kinetic diffusion forces. adsorptive forces and. to a lesser degree, the viscous forces resisting flow to the surface. The basic definition of irreducible water is that water held in a state of equilibrium between capillary and hydrostatic forces This water has been described by previous investigators as being held in the microcapillaries too small to support fluid flow. Actually, this fluid volume is made up of the water in the microcapillaries and as a film adhering to the surface of the crystals. All capillaries. therefore, possess some liquid as a film, the thickness of the film being dependent on the properties of the fluid and solid. A discussion of experiments with references pertaining to the measurement of this immobile layer next to the solid surface can be found in the text by J. J. Bikerman.8 Eversole and Lahr calculated the thickness of this layer to be in the order of 10 ' to 10' cm for aqueous solutions and glass. Between two quartz surfaces they found the thickness to be 2 x 10 cm. The work of Volkova, on the capillary movement of water and toluene in quartz grains, indicated the thickness of the Immobile layers to be near 10' cm. Since any measurement is an average value, it is easy to understand that an absolute value would depend on the roughness of the surfaces involved and the complexity of the system. A calculated effective pore radius of 2 x 10 cm is obtained at the, irreducible saturation of a porous media in a water-air system when a capillary pressure of 100 psi is applied. Since the separation of the sand grains is of the same approximate magnitude as the immobile layer.

Jan 1, 1951

By John W. Vanderwilt

THE Climax district is in northeastern Lake County, Colorado, on Fremont Pass (elevation 11,320 ft.) where the continental divide runs east-west joining high peaks of the Mosquito Range with the Sawatch Range, prominent ranges of the central Rocky Mountains. The mining towns of Leadville and Kokomo are thirteen miles to the south and seven miles to the northwest, respectively. Good highways lead to Denver (100 miles) by way of Kokomo, Loveland Pass, and Idaho Springs, and to Colorado Springs by way of Leadville. The ore deposit and mine are about one mile east of Fremont Pass and 600 ft. higher on the southwestern flank of Bartlett Mountain, a peak in the Mosquito Range. The molybdenite deposit lies within the Colorado mineral belt extending across the Rocky Mountains from Jamestown and Boulder to Silverton and Telluride. (See Fig. 1.) Fremont Pass is a broad open divide separating the drainage of the Tenmile River (Pacific) from that of the East Fork of the Arkansas River (Atlantic) ; but the mountains to the east and west are extremely rugged; altitudes in the vicinity vary from 11,000 to 14,000 ft. A conspicuous topographic feature within the district is the Tenmile Amphitheater, which lies east of Climax and

Jan 1, 1946

By Jr. Ensign C. O., J. W. Trammell

The two principal methods of room-and-pillar mining practiced at White Pine make it important to predict variations in the thickness and rock types of a stratum called the upper sandstone. In full column mining, the nearly barren upper sandstone occurs between two ore horizons (called upper shale and parting shale, respectively), and is mined as a part of the ore column. In general, full column mining is not practiced where excessive upper sandstone thickness causes full column ore grade to be substantially lower than parting shale ore. In parting shale mining, the upper sandstone forms the mine roof. Difficulty in predicting upper sandstone character arose because the extent and directions of its variations were not apparent in the property's diamond drilling, in which holes are spaced on 1000-ft centers. A study of sedimentary features, undertaken to improve the predictability of trends in the upper sandstone, led to an interpretation of the sedimentary environment in which the sandstone was deposited. Ripple marks, mud cracks, and cross-bedding, as well as other, less well known features, such as channel casts, flute casts, and current crescent casts, were mapped or recorded. These data, coupled with the knowledge of regional facies changes gained from studying drill core, show that the upper sandstone was deposited by a series of streams flowing northeastward over the underlying parting shale. Deposition of the lowermost bed of the upper sandstone, near the ancient shoreline, was locally preceded by erosion of the parting shale, and the greatest thicknesses of sandstone are found in channels scoured out of the parting shale. Awareness of the rather strong linear trends in the upper sandstone makes it possible to project continuous areas of thick upper sandstone through apparently isolated "highs" in the upper sandstone thickness contour map, which is based on drill-hole information. Since local exceptions to general stratigraphic trends exist, a method was also needed for esti- mating, in detail, the thickness and degree of shali-ness of upper sandstone forming the roof in active parting shale mining areas. Because locally the parting shale was thinned by erosion during the deposition of upper sandstone, parting shale thickness is inversely proportional to upper sandstone thickness. Utilizing information taken from short drill holes into the roof, curves were constructed for the correlation of parting shale thickness (measurable in the mine) with the upper sandstone total thickness, as well as the thickness of its basal member. The White Pine mine is situated in Ontonagon County, approximately 6 miles south of Lake Superior, in the Upper Peninsula of Michigan (Fig. 1). The orebody mined by White Pine Copper Co. occurs in the lowermost 20 to 25 ft of the Nonesuch formation, a series of middle- to upper-Keweenawan shales, siltstones, and sandstones. This paper discusses the solution of problems in ore grade control and mine planning arising from the presence of an essentially barren sandstone stratum within the orebody. THE OREBODY AND MINING TYPES A brief description of the orebody and mining methods is necessary to show how the upper sandstone affects mine planning and grade control. In the mine, the cupriferous zone of the Nonesuch shale is divided into three major portions, as shown in Fig. 2. Lying conformably on top of the Copper Harbor formation (or "lower sandstone") is the lower part of the orebody, the parting shale. The parting shale is overlain by the upper sandstone. On top of the upper sandstone, the basal portion (8 ft) of the upper shale is quite similar to the parting shale. Both the upper shale and parting shale are divided into a number of smaller stratigraphic units, each having a characteristic copper content, as shown by the histogram of Fig. 2. Other authors1'3 have dealt in detail with the stratigraphy and mineralogy of the orebody. However, in this article, we are concerned with the major subdivisions only, and it will suffice to note the low copper content of the upper sandstone, shown in Fig. 2. At present, two types of room-and-pillar openings are created in the orebody. In one type, called parting

Jan 1, 1964

By A. W. Bethune, L. M. Pidgeon

The equilibrium vapor pressure of zinc has been determined over the systems: ZnS(s) + Fe(s) = FeS(s) + Zn (vapor) and ZnS(s) + 2Cu(s) = Cu2S(s) + Zn (vapor) by reacting the components in an evacuated tube containing a thin copper fiber, and allowing equilibrium to be established with respect to the brass formed and the zinc in the vapor phase. The composition of the brass formed was determined and equilibrium vapor pressure values obtained from existing data. Between 850' and 1000°C, values for the first reaction ranged from 8 to 58 mm Hg; for the second reaction from 1.8 to 22 mm Hg. THE direct reduction of metallic sulphides may be indicated by an equation of the form MS + X = XS + M [1] where X is some suitable reducing agent. In the case where M is a metal of relatively high volatility, the use of vacuum will displace the equilibrium to the right at comparatively low temperatures in the face of the usually unfavorable thermochemistry. For this technique to be feasible, MS, X, and XS, must of course be nonvolatile compared to M. In this regard, zinc becomes an obvious metal both from the point of view of its volatility and the fact that most commercial ore deposits of this metal occur as the sulphide, and its direct reduction would avoid the expensive roasting process necessary prior to present day reduction methods. Two reducing agents are of interest, namely copper and iron, with the former having the added attraction of being relatively easily regenerated by oxidation in the copper converter. The possibility of these reactions has been considered before. Imbert1 patented a process whereby zinc sulphide was reduced by either copper or iron. The process as described took place at atmospheric pressure and, at the high temperatures required, the system was molten, resulting in the formation of a matte with the consequent reduction in the activities of the reactants. Peterson2 did further experimental work on these reactions and, although promising results were obtained on a small scale, attempts to increase the size of the furnace led to excessive "blue powder" formation. The early workers failed to realize the large reduction in temperature afforded by the use of vacuum and its desirable results. More recently, the direct reduction of zinc sulphide by iron has been the subject of a thermody-namic investigation by Kelley,3 and Gross and War-rington' have studied the kinetics of the reaction in the temperature range 900" to 1000°C, their experiments being conducted using vacuum and indicating that the reaction would proceed practically to completion in 1 to 3 hr, depending on the temperature. A qualitative examination of the system ZnS-Cu carried out at this University indicated that 98 pct of the theoretical zinc in the charge could be recovered in 1 hr at 1000°C when operating under reduced pressure. The purpose of this research program was to study the equilibrium of the two reactions. During the investigation, reference to the experimental work of Schenk5 was found but, due to lack of detail in the report, it was felt that it would be desirable to continue the equilibrium measurements. Selection of Method The measurement of vapor pressure of metals presents obvious problems resulting from the high temperatures which must be employed. In the present case, the production of the "reaction pressure" introduces a further difficulty since a chemical reaction must proceed before any metal vapor pressure is produced, resulting in depletion of the re-

Jan 1, 1954

By T. P. Chen, F. W. Bowdish

Studies made with a captive bubble apparatus on the sulfidization and collection by amyl xanthate of true chrysocolla specimens have defined the ranges of pH value and sulfide concentration which permit contact between the bubble and the mineral surface. Titanium compounds were the most effective of the materials found to activate the sulfidization of chrysocolla. With titanium activation, the contact angles and the ranges of pH value and sulfide wncentration giving bubble contact were all increased. Chrysocolla ores were concentrated by flotation. Chrysocolla ores occur at many localities in grade and quantity sufficient to make mining and millin feasible, but no satisfactory method of concentratio has been found. Although chrysocolla may be leached with acid, only those ores without acid-consuming gangue may be leached economically. Because of its potential importance, a study of the conditions nece sary for flotation of chrysocolla has been carried ou The literature contains a few references to flotation of chrysocolla. Two methods were developed by the U. S. Bureau of Mines.1,2 The first consisted of a fatty acid soap and a high xanthate as collectors of chrysocolla from a synthetic ore, while the second involved the use of hydrogen sulfide and xanthate. Ludt and DeWitt3 demonstrated the difference in adsorptive powers of chrysocolla and quartz for bas triphenyl methane dyes and suggested the use of butyl, hexyl or octyl-substituted malachite green as collector. Jackel4 emphasized the effects of combin tions of reagents such as Aerofloat 31, pine oil, and Reagents 404 and 425 with sodium sulfide and zinc hydrosulfite as conditioning agents. Although he reported recoveries of 89% from a synthetic ore and 98% from a natural ore containing azurite, malachite, chalcopyrite and chrysocolla, careful application of Jackel's method to chrysocolla from Tyrone, N.M., failed to give a high recovery. MATERIALS AND TECHNIQUE Samples from Inspiration, Ariz., and Tyrone and Magdalena, N. M., were used for experimentation and verified as true chrysocolla by leaching tests, specific gravity tests and X-ray diffraction. Chrysocolla does not dissolve at pH 4, although malachite and azurite do. Chrysocolla is about half as dense as the copper carbonates. X-ray diffraction analyses by the powder camera method confirmed the samples as true chrysocolla. A captive bubble apparatus, which cast an enlarged image of the air bubble and the mineral surface upon a screen, was used to check on the character of the surfaces. The specimens were prepared by grinding a flat surface on a glass plate using fine abrasive; then they were washed and kept in distilled water until they were to be treated with reagents. Before each reagent treatment, the specimen was carefully checked for cleanliness in the captive bubble apparatus. It was assumed that the surface was clean if, after fine grinding and washing of the specimen, the bubble would not stick. Specimens were handled with glass forceps, and precautions were taken to avoid contamination of the mineral surfaces. Contact angle measurements were carefully made several times on each treated specimen to obtain reliable average values. EFFECT OF pH VALUE AND SODIUM SULFIDE CONCENTRATION In each experiment, a specimen with a freshly ground surface was immersed for 10 min in a solution of sodium sulfide, washed and immersed for 15 min in a solution containing 30 mg per 1 of potassium amyl xanthate. The specimen was then washed again in distilled water and tested for contact angle in the captive bubble apparatus while submerged in distilled water. In this series of experiments, the pH of the sulfidizing solution was varied from 3 to 7, and the concentration of sodium sulfide, containing 60% Na2S, was varied from 50 to 650 mg per 1. Many combinations of pH value and sulfide concentration resulted in no contact between the bubble and the surface, but over a limited range of conditions, contact angles varying from 24ºto 52ºwere obtained. The data in Fig. 1 show sulfidization conditions that lead to bubble contact and those that do not. The region of contact is surprisingly small, which may indicate why flotation of chrysocolla involving sulfidization has proven so difficult in practice. Several features of the system are illustrated in Fig. 1. In the region between pH values of 4 and 6 with sodium sulfide concentrations below about 350

Jan 1, 1963

By F. H. Callaway, W. O. Keller

The various theories as to the well spacing-recovery relationship are reviewed in considerable detail and these theories analyzed in terms of their consistency with modern reservoir engineering concepts. It is concluded that the well spacing problem must be analyzed in terms of recovery efficiency and that a positive answer to the relation between well density and recovery efficiency is not available from direct comparisons of the production histories of wells and fields. The results of an engineering analysis designed to permit approximate calculation of recovery efficiencies as a function of well spacing in a depletion type reservoir from basic reservoir data is presented. Results of this type analysis indicate that the effect of well spacing on recovery efficiency in depletion type reservoirs can be expected to be very small. Limitations of this approach are pointed out, particularly with respect to its application in lenticular reservoirs. Testing techniques are outlined which should indicate whether or not a reservoir is continuous between wells and whether or not satisfactory drainage is being obtained with present spacings. A mass of data of this type indicates continuity to exist in most fields. INTRODUCTION The purpose of this paper is to review critically the engineering aspects of the well spacing problem, both from the standpoint of certain concepts and from the standpoint of reservoir mechanics. The well spacing problem is primarily an economic problem in which the optimum well density for a particular field is that density which will yield the greatest oil recovery consistent with justifiable development costs. The well spacing answer in terms of economic conditions, however, is extremely sensitive to the variation in recovery efficiency with well density. The variation in recovery efficiency with 'References given at end of paper. Manuscript received at the office of the Petroleum Branch October 2, 1949. Paper presented at the Petroleum Branch Meeting in San Antonio. Texas, October 5-7. 1949. well density is properly an engineering problem. Different opinions as to the correct answer to this engineering problem is the basis for most of the wide difference in opinion among various members of the industry as to optimum well spacing. This paper will be confined to the engineering problem of the relation between well density and ultimate oil recovery; economic considerations necessary for the evaluation of optimum well density for any particular field will not be discussed. The paper can be logically divided into two parts. The first part deals with a critical examination of the background and logic of the Cutler Rule and of similar studies by other authors and of related well spacing concepts. It is indicated that the variations in recoveries with well density* observed by Cutler and. others can be logically attributed to regional migration. Theoretical justification of the Cutler type relation wherein observed variations in recovery with well density in the same field is attributed to variations in recovery effiiency, in terms of energy relations, is refuted. The second part consists of a review of concepts of reservoir mechanics with regard to the well density-recovery relation. It is indicated that little variation of recovery efficiency with well density can be expected in a depletion type reservoir, unless lenticular conditions prevent communication between wells. The significance of field test data with regard to the existence or non-existence of lenticular conditions is pointed out. THE CUTLER RULE The first published article concerning the engineering aspects of the well spacing problem to receive wide attention was the work of W. W. Cutler' of the U. S. Bureau of Mines, published in 1924. In a study of decline curves from a large *The expression "well density" as utilized herein refers to the number of acres attributed to each well in a uniform well pattern, and is expressed in acres per well. The term "well spacing" Is utilized to denote the average distance between adjacent wells in a uniform well pattern. †Although the Cutler rule is treated critically in this discussion the authors intend no personal criticism of W. W. Cutler or of his hark The currently accepted concepts of reservoir mechanics were no": existent at the time of Cutler's work, as were even the moat basic tools (such as the bottom hole pressure bomb) for observing reservoir behavior.

Jan 1, 1950

By Vernon Read, John K. Hayes

Generally speaking, drilling equipment and techniques presently in use are not capable of meeting the complex technical-and economic-requirements of deep core drilling. Drilling to great depths involves increasing risk, time, and money. And it is especially true that costs increase with depth when conventional core drilling methods are used. Introduction of wire-line drilling techniques has been a great boon to deep core drilling, and is rapidly gaining general acceptance. But, in spite of recent progress, there remains a great need to improve productive drilling time and penetration rates, and still maintain high core recoveries.

Jan 1, 1959

By James F. Lewis

The hematite iron ore mines east of the Hudson River are confined to a strip of country ten to fifteen miles wide, commencing on the south, near Fishkill, running northeast through Dutchess County, and striking the southeast corner of Columbia County, extending into Litchfield County, Conn., known there as the Salisbury mines, cropping out again at and near West Stockbridge, Lanesboro, and Cheshire, Mass., also in Bennington County, Vermont. The ores are found in beds or deposits and in veins, with a limestone ledge on one side and slate on the other, from six to thirty feet from the surface. Portions of it are quite solid, called rock ore, but

Jan 7, 1950

By J. W. Snavely

THE high tension belt conveyor is introducing a new and tremendously expanded era of low cost bulk material handling. High tension belt conveyors are generally those installations involving very long centers, high lifts, or drops, in which the belts are stressed up to their maximum tension values, and further, where the belt construction provides tension capacity far beyond what is possible with conventional belt constructions. With these high tension installations, the magnitude of the forces involved demands careful refinement of accepted design practice in order to achieve optimum balance of all factors. No attempt will be made to evaluate the relative merits of belt conveyor haulage with other means of transportation. For present purposes, it is assumed this has already been done in favor of belt conveyor. Neither will any attempt be made to evaluate the various conveyor belt constructions now available or to balance the advantages of various types of mechanical equipment. It is also assumed that the basic haulage information on which the conveyor design is based is accurate and complete. A sustained maximum, uniform load on the belt at all times must be achieved through proper feed control and the use of adequate surge storage to level the peaks and valleys of any varying demand for the material being handled. General Belt Capacity Considerations The belt conveyor capacity tables published by various belting and conveyor equipment manufacturers vary to a considerable degree, and the ratings given are quite conservative. Of necessity, these published ratings are based on the handling of average materials under average conditions. In applying a high tension belt, all possible capacity from the belt must be obtained in order to hold its width to a minimum and thereby limit the initial cost. Two factors are involved, loading to maximum cross section area and traveling at a maximum practical speed. Belt Loading: Proper treatment of the loading of the belt will result in maximum cross section to the load, and published capacity ratings can be exceeded, sometimes by appreciable margins. On the 10-mile conveyor haul used in the construction of Shasta Dam, California, although the rated capacity of the belt line was 1100 tons per hr, at times the system handled peak loads of 1400 tons per hr, almost 25 pct better than the rated capacity. One of the large coal companies has been able to exceed rated capacity by as much as 50 pct. Loading conditions which must be controlled are: 1. Large lumps must be scalped off and rejected or the load must be primary crushed before being placed on the belt. 2. The material weight per cubic foot must be accurate, must be known for all the materials being handled, and must be known for the complete range of conditions of the individual material being handled. Long centers and high lifts magnify small differences into serious proportions. 3. Uniform feeding to the belt is most important. Various types of feeders are available, which can be used to place a constant predetermined volume of material on the belt, or, where an appreciable range of material weight exists, through electrical control actuated by current demand, to place a predetermined uniform tonnage on the belt. One long slope belt in a coal mine in Pennsylvania is being fed at three separate stations with the controls so arranged that whenever the maximum load is going onto the belt from the first station, the other two stations automatically cut out. Whenever the load from the first station drops back, the other two stations again automatically cut in. 4. Careful design of the chutes and skirts is necessary to get the load centered on the belt with a minimum of free margin along each edge. Some free margin at the edge of the belt is necessary to prevent spillage, but if the load can be kept accurately centered, this free margin area can be reduced, and more material can be carried on the belt. What can be accomplished in this respect will vary widely, depending on the nature of the material being hauled. The chute and skirt design must also protect the belt. 5. The design of chutes and skirts should also get the load traveling in the same direction and close to belt speed, so that the load comes to rest on the belt as quickly as possible. The design of the chutes and skirts is worthy of careful study, and after a system is put into operation it should be experimented with to get the best results. Belt Speed: High belt speeds should be used in high tension work. Obviously, high belt speeds enable haulage on a narrower belt, reducing initial cost. The major portion of belt wear takes place at the loading point and around the terminal pulleys. The

Jan 1, 1952

By C. S. Smith

The solubility of tantalum at 1100°C is 0.025 pct in pure copper, 1.2 pct with 20 pct Ni, and 2.7 pct with -30 pct Ni. The solubility decreases with temperature, and the alloys are precipitation hardenable. A 79/20/1 Cu-Ni-Ta alloy reaches maximum hardness after aging at about 750°C and, if cold worked, does not recrystal-lize below that temperature. The alloys have good tensile properties at moderately elevated temperatures and, since they can be hot and cold worked nearly as easily as cupronickel, they are suggested for service at temperatures above the usual limits for copper alloys. MASSIVE tantalum dissolves extremely slowly in copper-rich alloys, and tantalum powder forms refractory surface layers which prevent its solution unless special precautions are taken. After many trials, an 80 pit recovery was consistently obtained by using a mixture of a finely divided tantalum powder with two or three times its volume of potassium tantalum fluoride (K2TaF1,-melting point about 750°C) poured in a slow stream directly on to the surface of the molten copper alloy, so that each particle would remain coated with flux and be immediately and individually wetted by the molten metal. A 50-50 nickel-tantalum alloy has a melting point of about 1400°C and small lumps of it slowly but satisfactorily dissolve in molten copper-nickel alloys as does commerical "ferro-tantalum." The high affinity of tantalum for carbon makes it necessary to avoid carbonaceous crucibles. COPPER-TANTALUM ALLOYS Castings with good shrinkage resulted when 0.1 pct Ta or more was added to copper melted under charcoal and cast in air. Copper in two-pound melts to which 0.1, 0.15, and 0.2 pct Ta had been added retained residual amounts, by analysis, of 0.025, 0.023 and 0.026 pct, respectively. The solubility of tantalum in molten copper at about 1200 °C is therefore about 0.025 pct. The electrical conductivity of these three samples was 100.48, 100.00 and 100.20 pct IACS in the annealed condition, and 98.24, 97.98 and 98.08 in the cold-drawn condition. Wires 0.080 in. in diam withstood from thirteen to nineteen reversed bends after annealing for 30 min in hydrogen at 850°C and the metal was therefore completely deoxidized. The annealing temperature corresponding to 50 pct loss of work hardness in 1 hr in a strip cold rolled 77 pct reduction was 175C, compared to 230°C for equivalent undeoxidized material. It is unusual for a decrease of recrystallization temperature to result from an alloying addition, and the tantalum probably combines with and removes some minor impurity in the original copper that restrained its recrystalliza- tion. No difficulty whatever was encountered in hot or cold rolling or drawing these ingots, and were tantalum not so expensive it would make an excellent deoxidizer for copper. COPPER-NIcKEL-TANTALUM ALLOYS Studies were made of the solubility of tantalum in a large number of copper-rich alloys, but of these significant solubilities were found only in the case of an iron alloy (which, however, separated into two immiscible liquid layers) and alloys of copper and nickel, in which tantalum is freely soluble and which were found to have interesting properties. The latter alloys are the subject of U. S. Patent No. 2,430,306 (1947). They are not at present available commercially. The Ternary Constitution Diagram—The copper-rich corner of the constitution diagram was constructed on the basis of microscopic examination of samples of various composition and treatments. The alloys were hot rolled to 0.25 in. from castings 0.625 in. thick, then annealed for 1 hr at 1000°C, quenched, cold rolled to 0.10 in. and finally annealed for various periods of time at temperatures between 800" and 1100°C in a nonoxidizing atmosphere and quenched. When the solid-solubility limit had been transgressed, particles of a compound believed to be Ni2Ta*were

Jan 1, 1960

By S. Floreen, R. M. Pilliar, H. W. Hayden

The fracture behavior of a 50 pct Cu-50 pct Fe mi-croduplex alloy, laminated composites of copper and iron and an extruded 50-50 Cu-Fe elemental powder composite was studied. Very low ductile-brittle transition temperatures were achieved in all cases, but for different reasons. In the microduplex alloy both the initiation and also the propagation of cleavage fractures appeared retarded by the very small in-terphase distances. In the composites, crack propagation through the sumples was prevented in most cases by delamination fractures perpendicular to the advancing cracks. These delaminations occurred at different regions and by different mechanisms in the various composites. In the extruded powder composite, de-lamination appeared to take place along preexisting flaws. In the crack arrest geometry of the laminated plates, delamination took place by localized shear fractures within the copper near the Fe-Cu interfaces. In this case delamination was enhanced by thicker laminate layers, and by having the resistance to shear failure of the copper sufficiently low compared to the toughness of the iron. BRITTLE fracture in engineering materials has long been a problem, and many different ways of preventing it have been considered. One method that has been of growing interest lately is to prevent crack propagation by the introduction of mechanical discontinuities into the structure. These discontinuities may act in several ways. They may simply act as crack stoppers. They may introduce secondary fractures such as de-laminations that deflect the initial crack into new, less damaging directions. Alternatively, they may subdivide a fairly large bulk sample that would have been loaded in plane strain, for example, into a number of subunits that are individually loaded in plane stress and thus are more resistant to fracture. Other mechanisms, or combinations of mechanisms, are also feasible. A number of methods exist for introducing mechanical discontinuities into a structure. Composites by their nature have discontinuities in structure, and numerous studies have shown that fracture propagation in materials of this type can be radically changed by suitable control of the composite parameters. Of particular significance to the present work are recent investigations of layered composites made by joining high strength steel sheets by various means.'-4 These studies have shown that through proper control of the mechanical properties of the bonds joining the sheets it was possible to introduce delamination fractures that markedly improved the overall toughness of the composites and in some cases completely prevented through-the-thickness fractures. Another technique for introducing structural discontinuities is simply to use a two-phase alloy. It has been recognized for many years that a small amount of a second phase may improve toughness either by homogenizing plastic flow and thus preventing localized stress concentrations that nucleate fracture, or by interacting with an advancing crack. In most of these studies of two-phase materials, the decreases in ductile-brittle transition temperatures produced by the second phase were relatively small. More recently, work on two-phase stainless steels having a very fine grain microduplex structure has shown that the presence of on the order of 40 to 50 pct of a tougher second phase may lower the ductile-brittle transition temperature of the brittle phase by approximately 300°F. 5-7 In these alloys delaminations were seldom observed. The tougher second phase appeared to minimize the ease of both the initiation and the propagation of cleavage fractures. These results show that both the composite approach and the microduplex alloy approach are effective methods of preventing brittle fracture. Therefore, it was of interest to compare the fracture behavior of a microduplex alloy with composites made from the two-phases that were present in the alloy. To simplify this comparison the 50 pct Cu-50 pct Fe system was selected for study. At low temperatures the equilibrium tie line phases in this system are essentially pure ferrite and pure copper. A 50-50 alloy was cast and hot worked to produce a microduplex structure. Two types of composites were studied; laminated structures prepared by roll bonding iron and copper sheets of the tie line compositions, and an extruded powder composite made from high purity elemental powders. The fracture behavior of these materials was then compared. EXPERIMENTAL PROCEDURE Alloy Preparation. The 50-50 Fe-Cu alloy and the components for the roll bonded composites were prepared by vacuum induction melting 30-lb heats using electrolytic grades of iron and copper as charge materials. A carbon boil was used to deoxidize the melts. Small additions of copper and iron were made to the iron and copper heats, respectively, to approximate

Jan 1, 1970

By F. L. Toy, C. E. Sims

PNEUMATIC processes for converting molten pig iron to steel were the major producers of steel during the latter half of the 19th Century and until shortly after the turn of the century, when these processes began losing ground to their rival—the basic open hearth process. The first pneumatic process for converting molten pig iron to steel was introduced almost simultaneously by Sir Henry Bessemer in England and by William Kelley in the United States during the decade 1850-1860, but it was Bessemer who pushed it to a successful commercial operation. The process itself is so basically simple that it has not lent itself to many extensive or revolutionary improvements since its inception. During the first 30 yr, numerous' designs of vessel and many modifications of practice were tried before the industry standardized on something resembling current practice. The evolution in construction has resulted in proportions and arrangements based more on mechanical expedience than on metallurgical requirements. Many of the early converters were stationary and had to be tapped similarly to the cupola or open hearth furnace. Such types were used for many years in Sweden and Germany, but the rotating or tilting types were favored in England and the United States. There was evident trouble with bottom tuyeres, becAuse about half of the early designs were constructed for side blowing with tuyeres submerged below the metal-bath surface. Even portable tuyeres, introduced from above, were tried. The principal advance, which made bottom blowing practical, was the invention, in 1863, by Henry Bessemer, and subsequent improvement by A. L. Holley, of a detachable bottom so that the tuyere section could be renewed. In the meantime, it was the demonstration, by Robert Mushet, of the value of manganese in overcoming hot-shortness in converter steel, that made the process metallurgically feasible. At first, only the acid practice was used, and Swedish pig iron, because of its low sulphur and phosphorus, was in great demand. The development of a suitable basic lining, by Sidney G. Thomas and Percy G. Gilchrist in 1876-1879, made the process applicable to the high-phosphorus ores of Europe. From the beginning of the use of the Thomas-Gilchrist process, trouble was experienced in obtaining sufficiently high temperatures in the blown metal, particularly in small heats. It is necessary, of course, to have the low-carbon blown metal at a considerably higher temperature for handling from vessel to ladle to molds than is needed for the molten pig iron. All of the increase in temperature in a pneumatic process must be obtained from the heat of combustion of the elements contained in the pig iron plus that from silicon when ferrosilicon is added. In bottom blowing practice, carbon is almost completely burned, but inasmuch as it cannot be oxidized to a higher state of oxidation than to CO when in contact with liquid iron, only a minor portion of the heat-producing capacity is utilized. Under such conditions, it is not considered to be an efficient source of heat. In the acid practice, therefore, silicon and manganese are the large heat producers, are quickly oxidized, and are efficient fuels. Too much manganese, however, thins the slag greatly and tends to promote excessive slopping or throwing of slag and metal from the converter during the blow. Most operators, therefore, like to keep the manganese content of the iron below 0.60 pct, and thus the main source of heat is the silicon, which must be at least 1 pct and sometimes 2 to 3 pct for small blows such as for castings. Ferrosilicon is sometimes added during the blow to obtain an adequate temperature level. In the basic (Thomas) practice, even more heat is required, because cold lime is added to flux the phosphorus after it is oxidized, and the basic slag must be maintained at a proper temperature. Silicon

Jan 1, 1951

By Jack A. Caldwell

Jan 1, 1985

By L. P. Teas

The main facts brought out by the study of the blende-bearing ores examined are as follows: 1. Chalcopyrite as minute triangular or rectangular dots, or as

Jan 1, 1918

By Norikazu Hashimoto

SILICIDE coatings on refractory metals are often applied by transporting a silicon halide vapor to a hot metal surface . The coatings are usually grown by solid-state diffusion reaction. The diffusion is often slower than the vapor transport and hence may become rate-controlling.

Jan 1, 1968

By Ralph B. Stevens

INTRODUCTION Underground fires continue to be one of the most serious hazards to life and property in the mining industry. Although underground mines are analogous to high-rise buildings where persons are isolated from immediate escape or rescue, application of technology to locate and control fire hazards while still in their controllable state is slow to be implemented in underground mines. Even in large surface structures such as hotels, often only fire protection systems which meet minimal laws are implemented due to the high cost of adding extensive extinguishing systems, isolation barriers, alternate ventilation, escape routes and alarm systems. Incomplete and ineffective protection occasionally is evidenced where costs would not seem to be a factor, such as the $211 million MGM Grand Hotel fire November 21, 19801. Paramount in increasing fire safety and decreasing the threat of serious fire is early warning followed by proper decision analysis to perform the correct action. However, very complex fire situations can be produced in structures such as high-rise buildings and underground mines simply because of the distances between the numerous fire-potential locations and fire safe areas. Other complexities arise when normal activities occur that emit products of combustion signaling a fire condition to a sensitive fire/smoke sensor. For example, the operation of diesel equipment or the performance of regular blasting can produce combustion products that reach the sensitive alarm points of many sensors2. Smoke detectors for surface installations provide fire warning when occupants are at a distant location or when sleeping, thus greatly reducing injuries and property damage. However, when installed in the harsh environments of underground mines, fire and smoke detection equipment soon becomes inoperative, unreliable, or requires excessive maintenance. The U.S. Bureau of Mines has performed many studies and tests to improve fire and smoke protection for underground mine workers3. This paper describes several USBM safety programs which included in-mine testing with mine fire and smoke sensors, telemetry and instrumentation to develop recommendations for improving mine fire safety. It is hoped that the technology developed during these programs can be added to other programs to provide the mining industry with the necessary fire safety facts. By recognizing fire potentials and being provided with cost-effective, proven components that will perform reliably under the poor environmental conditions of mining, mine operators can provide protection for their working life and property equal to that which they provide for themselves and their families at home. The basis of this report is two USBM programs for fire protection in metal and nonmetal mines4,5 and one coal program6. The data was collected beginning in May 1974 and continuing through the present with underground tests of a South African fire system installed at Magma Mine in Superior, Arizona, and a computer-assisted, experimental system at Peabody Coal Mine in Pawnee, Illinois. The conduct of each program was as follows: • Define the problem and its magnitude in the industry • Develop concepts to solve or diminish the problem • Review available hardware or systems approaches to fit the concepts • Install and demonstrate the performance of a prototype system through fire tests in an operating mine. MINE FIRE FACTS Whether in coal or metal and nonmetal mines, the potential severity of fire hazard is directly related to location. As shown in Figure 1, fire in intake air at zones A, B, C or D can cause contamined air to route throughout the mine quickly if not detected, isolated or rerouted. Causes and location of former metal and nonmetal fires are represented in Table 1; the cause and location of fatalities and injuries is shown in Table 2. Coal-related fires and their impact on deaths and injuries are graphed in Figure 2; their locations are described in Table 37. Significantly the table shows that the hazard to personnel was three times greater for fires occurring in shaft or slope areas, and the percentage of deaths and injuries was four times that of other areas. Number of Persons Affected A 129-mine sample indicated that from 8 to 479 employees per shift work in underground metal and nonmetal mines, and that deeper mines have larger populations, as shown in Figure 3. Coal mining relates similar employment, and a 16-state sample of 670 mines employing at least 25 persons shows the distribution in Figure 4. Drift mines accounted for 58 percent of the sample but employ only 45 percent of the underground workers.

Jan 1, 1982

By Craig S. Hartley

An approximate expression is obtained for the energy of a closed dislocation loop in equi1ibriu)n with a constant net stress. The result obtained is valid for loops in isotropic or anisotropzc materials provided that they are suJficiently large that the energy per unit length of a segment of the loop can be approximated by that of an infinite straight dislocation tangent to the loop. It is shown that this approximation leads to very close agreement with a more rigorous calculation of the elastic energy of a circular glide loop. The Gibbs-Wulff Form, GWF, of a dislocation is the closed planar loop which has the smallest elastic self-energy of all possible loops having the same Burgers vector and enclosing a fixed area, A.' The energy of such a loop is related to the net resolved shear stress* required to expand the loop and to the stress required to activate a Frank-Read source.223 In the following sectiorls the problem of determining the form of the GWF is discussed and an approximate method for calculating its elastic self-energy is presented. It is demonstrated that the approximations employed lead to no serious errors when applied to a calculation of the elastic energy of a circular glide loop. This method is then used to obtain a closed form expression for the energy of GWFs in isotropic and anisotropic materials. THEORY Burton, Frank, and cabrera4 have proved that the relationship of the equilibrium shape of a two-dimensional array of atoms under the influence of the Gibbs free energy associated with unit length of its boundary, G(O), is that the polar plot of G(0) vs 0 is proportional to the pedal of the GWF.* The angle 0 is measured "The pedal of the polar graph ofG(0) vs0 is the envelope of tangents to the eraph.relative to some crystallographic reference direction. The difficulty in applying this result to a closed dislocation loop arises from the self-interaction of the loop. For a dislocation the energy analogous to G(0) is a function of the total configurati~n.~ Consequently the relation which determines the GWF is an integro-differential equation rather than the simple differen- tial equation which results when G(8) is a function of 0 alone. Mitchell and smialek3 and Brown~ have used the self-stress concept introduced by ~rown' to calculate the shapes of dislocations in equilibrium with an applied stress. In this approach the glide force on an element of the dislocation loop due to the interaction of the element with the rest of the loop is equated to the glide force exerted by the local applied stress. The shape of the loop is then adjusted so that the two forces above are equal at all points on the loop. It is possible to calculate the energies of such loops by noting that, for equilibrium with an applied stress, the energy is equal to pijbiAj (summation convention) where bi is the Burgers vector, p.. is the local net stress tensor, and Ai is a vector directed perpendicular to the plane of thd loop with magnitude equal to the area of the loop. Also Brown' has calculated the energy of a hypothetical polygonal GWF using the above technique and anisotropic elasticity. However, his indicated solution for the energy in the general case of an arbitrary GWF is only slightly less involved than an iterative solution of the integrodifferential equation referred to earlier. In the present work the approximation employed by DeWit and Koehler' is used to calculate the energy of a closed loop in equilibrium with an applied stress. That is, the energy of a loop segment, ds, is approximated by the product of ds and the energy per unit length of an infinite, straight dislocation in a cylinder coaxial with the tangent to the loop at the angular position of the segment. This is known as the "line tension" approximation. The inner cutoff radius of the elastic solution defines the core radius, while the outer cutoff radius is determined by some characteristic dimension of the loop. Actually, both of these radii vary with the edge-screw character of the segment. The effective core radius changes because of the orientation dependence of the Peierls width of a dislocation,8 and the outer radius should be the radial distance from the circumference of the loop to the center of symmetry of the area enclosed by the loop.g However, since the energy varies logarithmically with the ratio of these radii while depending directly on the effective elastic constants, only the effect of the latter is considered. This approximation also neglects the self-interaction of the loop segments. For small loops this will doubtless be extremely important, but for large glide loops produced by plastic deformation the self-interaction is not nearly so important in determining the energy of the loop. This point is illustrated by the following calculation of the energy of a circular loop. Consider a circular loop of radius R which lies in the XI - x, plane of an infinite isotropic continuum and whose Burgers vector makes an angle $ with xs. The first-order solution for the elastic self-energy is:'

Jan 1, 1970