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By Francis B. Foley

Fundamentally, slag control in acid open-hearth steelmaking cannot be considered from any different viewpoint than can that in basic open-hearth melting. The same laws of mass action and distribution apply in both cases, so that while most, if not all, of the recent intensive study of slag-metal reactions have been carried on in basic furnaces the principles evolved are nevertheless applicable to acid open-hearth steelmaking and it is therefore not our intention to discuss the various possible chemical reactions in the slag and between slag and metal. The points we have to carry in mind, in considering acid melting practice, are the differences in the type of steels produced and the difference in the chemistry of both the steel and the slag. The acid furnace is used to produce, let us not say "quality" steel, for, after all, a "quality" steel for one purpose is not a "quality" steel for another, but steels that require acid melting practice to be of proper "quality." Steels that are required to have high physical properties in tests taken transversely from forgings are peculiarly the product of acid furnaces. In this class may be placed various ordnance materials, notably gun forgings and projectiles and large commercial forgings requiring high transverse ductility. Forgings requiring an extremely hard surface, polished and free from imperfections, are also typical acid-steel products. Of course, such forgings can be made of basic open-hearth steel, but not with the same assurance of success. At the start, the preparation of the slag in the acid and basic furnaces differs. Lime is charged in the beginning of a basic heat to form a slag, but in the acid furnace nothing is added for the special purpose of producing a slag when the heat is melted down. The result is that the acid furnace has much less volume of slag than the basic. The slag formed upon melting down is made up almost entirely of the oxides of elements contained in the metallic charge, notably oxides of iron, manganese and silicon, the amount of silica being augmented somewhat by erosion of the furnace lining. A typical acid open-hearth slag, when the bath is completely melted, will have an analysis about as follows: CaO, 3.40 per cent; SiO2, 51.9; Fe2O3, 1.25 ; FeO, 28.86; Al2o3, 1.82; MnO, 12.22. The high iron oxide content of the slag gives it a black appearance. The metal is high in carbon and some manganese and silicon will be present, depending on the amount in the original charge, the amount of

Jan 1, 1935

By Francis B. Foley

Fundamentally, slag control in acid open-hearth steelmaking cannot be considered from any different viewpoint than can that in basic open-hearth melting. The same laws of mass action and distribution apply in both cases, so that while most, if not all, of the recent intensive study of slag-metal reactions have been carried on in basic furnaces the principles evolved are nevertheless applicable to acid open-hearth steelmaking and it is therefore not our intention to discuss the various possible chemical reactions in the slag and between slag and metal. The points we have to carry in mind, in considering acid melting practice, are the differences in the type of steels produced and the difference in the chemistry of both the steel and the slag. The acid furnace is used to produce, let us not say "quality" steel, for, after all, a "quality" steel for one purpose is not a "quality" steel for another, but steels that require acid melting practice to be of proper "quality." Steels that are required to have high physical properties in tests taken transversely from forgings are peculiarly the product of acid furnaces. In this class may be placed various ordnance materials, notably gun forgings and projectiles and large commercial forgings requiring high transverse ductility. Forgings requiring an extremely hard surface, polished and free from imperfections, are also typical acid-steel products. Of course, such forgings can be made of basic open-hearth steel, but not with the same assurance of success. At the start, the preparation of the slag in the acid and basic furnaces differs. Lime is charged in the beginning of a basic heat to form a slag, but in the acid furnace nothing is added for the special purpose of producing a slag when the heat is melted down. The result is that the acid furnace has much less volume of slag than the basic. The slag formed upon melting down is made up almost entirely of the oxides of elements contained in the metallic charge, notably oxides of iron, manganese and silicon, the amount of silica being augmented somewhat by erosion of the furnace lining. A typical acid open-hearth slag, when the bath is completely melted, will have an analysis about as follows: CaO, 3.40 per cent; SiO2, 51.9; Fe2O3, 1.25 ; FeO, 28.86; Al2o3, 1.82; MnO, 12.22. The high iron oxide content of the slag gives it a black appearance. The metal is high in carbon and some manganese and silicon will be present, depending on the amount in the original charge, the amount of

Jan 1, 1935

By I. D. Shah

A roasting procedure for synthetic copper sulfide which results in a reproducible transient melting at 450° and 470°C is described in this paper. Possible causes for this melting phenomenon are suggested, and also a procedure to avoid it during the roasting. This melting is believed to be responsible for the occasional occurrence of stickiness in the commercial roasting of copper sulfide concentrate. Many workers have studied the roasting of copper sulfides, but only Lewis et al. have reported that fusion and, in some cases, fusion with "boiling" occurred in the neighborhood of 450°C. The DTA and TGA apparatus, the materials used in this investigation, and the discussion of chemical reactions occurring during the CuS roasting are described in an earlier publication. CuS, in the form of either pellet or powder, was heated at 280° to 330°C in a flowing air or 30% 02: 70% N2 atmosphere. Heating was continued at this temperature until no further weight increase due to sulfates formation was observed on the TGA thermogram. This is believed to be due to the inability of SO, to diffuse out from the sulfate layer on the surface of the pellet, thus halting the sulfate-forming reactions. Hereafter, the sample oxidized in this manner will be referred to as "partially oxidized CuS." X-ray analysis and microscopic examination of a polished cross section of a partially oxidized CuS pellet showed the following order of layers starting from the center of the pellet: Cu2S, Cu2O, CuO • CUSO4, and CuSO4. A bare thermocouple inserted in the center of a partially oxidized CuS pellet (2.5 g) did not indicate any noticeable change in the temperature of the pellet when melted at 450° to 470°C in an inert or oxidizing atmosphere. During the preparation of a partially oxidized pellet, the temperature of a CuS pellet increased by 50°C due to the exothermic formation of sulfates in air at 320°C. The endothermic peak associated with the melting phenomenon at 450°C is shown in [Fig. 1.] A CuS pellet (0.2 g) was oxidized in flowing air (150 cc per min) at 10°C per min up to 320°C and held there for about 2 hr. Then, N2 replaced air and heating was continued up to 500°C. The exothermic peak at 320°C represents an overall formation of sulfates, and an endothermic peak at 450° to 470°C represents the melting phenomenon. To observe the melting visually, a small (50 mg), partially oxidized CuS pellet was heated (in static air) in a hot-stage microscope. The melting phenomenon was observed on an auxiliary TV screen. The sample

Jan 1, 1976

By Howard Clarke, Cyril Stanley Smith, Clarence Zener

THE importance of cold-working in the fabrication of metals is in marked contrast with our ignorance as to the nature of the accompanying internal changes. The failure of the usual methods of testing to provide a consistent picture of these internal changes has given the incentive for the present program of investigating the effect of cold-working upon the internal friction of metals measured at small strain amplitudes. It has previously been shown1,2,3, that the internal friction introduced by cold-work is removed by annealing at comparatively low temperatures. In the present paper the following additional factors are investigated: (I) frequency of measurement, (2) amount of cold-work, (3) temperature of measurement. Special care was taken to minimize the internal friction due to causes other than cold-work. The present paper is purely experimental, the authors believing that any conclusions drawn from this work as to the nature of the internal changes accompanying cold-work would be premature." PREPARATIONS OF SPECIMENS In this study of cold-work it was desirable to eliminate as far as possible all other sources of internal friction. One disturbing source is the lack of temperature equilibrium between adjacent crystallites of a vibrating specimen.4 Temperature equilibrium is aided by small grain size and by the use of low-frequency vibrations. Unless unduly long specimens are used, low frequencies can be obtained only in transverse vibrations. But in transverse vibrations a second source of internal frictions is present, that arising from transverse thermal currents.5 The effect of these may be reduced by increasing the transverse dimensions. The grain size and transverse dimensions of the specimens were thus so chosen as to render the vibrations simultaneously nearly isothermal with respect to intercrystalline thermal currents, and nearly adiabatic with respect to transverse thermal currents. The material for this investigation was a 70-30 brass, part of the same lot of material that was used for some earlier studies on the effect of strain on proportional limit.10 It had the following composition: copper, 70.59 per cent; zinc, 29.39; lead, 0.01; iron, 0.01. This material was received from the brass mill in the form of 3/4-in. rod, cold-drawn to final size by a 65 per cent reduction of area from material that had been annealed to give a grain size of 0.060 mm. Rods of appropriate length were annealed for 2 hr. at 400°C. and cooled in air. The annealed rods (grain size, 0.012 mm.) were then stretched in a tensiontesting machine, using special grips10 to ensure axial application of the load. The rate

Jan 1, 1941

By E. H. Crabtree

THE heavy-media separation plant at the Central mill of the Eagle-Picher Mining and Smelting Co., near Picher, Okla., was started in February 1939. Since that time twenty-four million tons of lead-zinc ore has been treated in the mill, of which approximately 70 pct has been treated by heavy-media separation. Until January 1945, galena concentrate from the flotation plant was used as the sink-float medium. At that time minus 100-mesh ferrosilicon was substituted for the galena, and its use has been continued. The purpose of this paper is to compare the operating results and the costs of operation of the two media. For the purpose of this comparison the period Jan. 1, 1943, to Jan. 29, 1945, has been taken to represent the operation using galena medium; the period Jan. 29, 1945, to Sept. 1, 1946, to represent the ferrosilicon medium. During the first of 'these periods about seven and one-half million tons was milled; during the latter, or ferrosilicon period, more than five million tons was milled. It is believed that these tonnages are large enough to give representative data. FLOWSHEET The Central mill has a capacity of 15,000 tons per day. Briefly, the treatment used consists of crushing the ore through 11/2-in. and wet-screening out the minus 3/16in. portion. The minus 1 1/2 plus 3/16in. washed product is treated by the heavy medium; the minus 3/16-in. product is deslimed and jigged. The heavy-media cone tailing is discarded and the cone .concentrate is recrushed and jigged for the production of a coarse lead concentrate. The cone concentrate, impoverished of lead, is then ground in ball mills and treated by differential flotation. The minus 3/16-in. jig product is similarly treated. HEAVY-MEDIA PLANT The heavy-media cone plant treating the minus 1 ½ plus 3/16-in. product consists of one open-top cone 20 ft in diameter by 20 ft deep, equipped with a central airlift. Concentrate is discharged from the top of the airlift to a 3 by 14-ft low-head Allis-Chalmers drainage screen. Medium drains back into the cone, from the first half of this screen. The last half of the screen is provided with washing sprays for further removal of medium. Cone tailing is similarly drained and washed. Undersize from the washing screen is thickened and then cleaned in the medium-cleaning plant. When galena was used as a medium, cleaning was done in Fagergren and Denver flotation machines. The ferrosilicon is cleaned in three 36-in. and one 24-in. Crockett magnetic separators. A flowsheet of the cone plant is shown in Fig 1.

Jan 1, 1947

By R. J. Sherman, M. R. Achter

IT has often been reported that coatings may strengthen single crystals and polycrystalline specimens of coarse grain size. This note reports the effect of a surface layer of fine grains on the creep-rupture properties of coarse grained nickel 99.8 pct pure. Sheet specimens were machined from cold-rolled material, finished by surface grinding and annealed at 1230°C in vacuum for 4 hr. Half of them were re-ground on faces and sides and annealed at 925"C for 1 1/4 hr. The double grind-anneal treatment produced a structure consisting of a surface layer of fine grains of 0.02 mm diam on a coarse-grained substrate of grains of 0.40 mm diam. Comparison with specimens which received only the single grind showed that the second grind-anneal did not alter the structure of the substrate. Several modifications were made in specimen design to eliminate erratic rupture behavior outside of the gage length. In the final design the gage section, oriented parallel to the rolling direction, was 1.00 in. long, 0.4 in. wide, and 0.0625 in. thick. End tabs were pinned in place. Although other configurations (welded end tabs, transverse orientation, and wider gage section) showed similar effects of a layer of fine grains, because of the erratic rupture behavior, only the data from the final design are reported here. Creep tests were performed at 650°C in air and in a vacuum of 10~5 mm Hg. Creep rates were measured from the movements of the lever arm. Since the broken sheet specimens could not be fitted together, elongations at fracture measured by means of gage marks, do not include the necked portions adjacent to the fracture. Results of the creep tests with and without the fine grained surface are given in Table I. An explanation for the lower strength in air than in vacuum has been proposed previously in terms of the reduction of crack propagation energy by gas adsorption.l In both environments the fine-grained layer produced a small decrease in minimum creep rate and, except for the shortest test in air, a substantial increase in rupture life. The elongation values at rupture, as they are measured here, do not demonstrate an appreciable effect of the grain structure but do show much smaller ductilities in air than in vacuum. Greater elongations maybe associated with the finegrained layer, however, especially in vacuum, in creep curves such as those in Fig. 1. Duplicate tests were not run to determine the re-producibility of the data. It is felt, however, that, in view of the consistent behavior, the prolongation of rupture life may be attributed to the effect of the fine-grained surface. In the data not reported be-

Jan 1, 1962

By E. H. Crabtree

The heavy-media separation plant at the Central mill of the Eagle-Picher Mining and Smelting Co., near Picher, Okla., was started in February 1939. Since that time twenty-four million tons of lead-zinc ore has been treated in the mill, of which approximately 70 pct has been treated by heavy-media separation. Until January 1945, galena concentrate from the flotation plant was used as the sink-float medium. At that time minus 100-mesh ferrosilicon was substituted for the galena, and its use has been continued. The purpose of this paper is to compare the operating results and the costs of operation of the two media. For the purpose of this comparison the period Jan. I, 1943, to Jan. 29, 1945, has been taken to represent the operation using galena medium; the period Jan. 29, 1945, to Sept. I, 1946, to represent the ferrosilicon medium. During the first of these periods about seven and one-half million tons was milled; during the latter, or ferrosilicon period, more than five million tons was milled. It is believed that these tonnages are large enough to give representative data. Flowsheet The Central mill has a capacity of 15,000 tons per day. Briefly, the treatment used consists of crushing the ore through I½- in. and wet-screening out the minus 3/16-in. portion. The minus 136 plus 3/16-in. washed product is treated by the heavy medium; the minus 3/16-in. product is deslimed and jigged. The heavy-media cone tailing is discarded and the cone concentrate is recrushed and jigged for the production of a coarse lead concentrate. The cone concentrate, impoverished of lead, is then ground in ball mills and treated by differential flotation. The minus 3/16-in. jig product is similarly treated. Heavy-media Plant The heavy-media cone plant treating the minus I½ plus 3/16-in. product consists of one open-top cone 20 ft in diameter by 20 ft deep, equipped with a central airlift. Concentrate is discharged from the top of the airlift to a 3 by 14-ft low-head Allis-Chalmers drainage screen. Medium drains back into the cone, from the first half of this screen. The last half of the screen is provided with washing sprays for further removal of medium. Cone tailing is similarly drained and washed. Undersize from the washing screen is thickened and then cleaned in the medium-cleaning plant. When galena was used as a medium, cleaning was done in Fagergren and Denver flotation machines. The ferrosilicon is cleaned in three 36-in. and one 24-in. Crockett magnetic separators. A flowsheet of the cone plant is shown in Fig I.

Jan 1, 1949

By Clarence Zener, Howard Clarke, Cyril Stanley Smith

The importance of cold-working in the fabrication of metals is in marked contrast with our ignorance as to the nature of the accompanying internal changes. The failure of the usual methods of testing to provide a consistent picture of these internal changes has given the incentive for the present program of investigating the effect of cold-working upon the internal friction of metals measured at small strain amplitudes. It has previously been shown1,2,3, that the internal friction introduced by cold-work is removed by annealing at comparatively low temperatures. In the present paper the following additional factors are investigated: (I) frequency of measurement, (2) amount of cold-work, (3) temperature of measurement. Special care was taken to minimize the internal friction due to causes other than cold-work. The present paper is purely experimental, the authors believing that any conclusions drawn from this work as to the nature of the internal changes accompanying cold-work would be premature.'l PrEpaRations oF Specimens In this study of cold-work it was desirable to eliminate as far as possible all other sources of internal friction. One disturbing source is the lack of temperature equilibrium between adjacent crystallites of a vibrating specimen.4 Temperature equilibrium is aided by small grain size and by the use of low-frequency vibrations. Unless unduly long specimens are used, low frequencies can be obtained only in transverse vibrations. But in transverse vibrations a second source of internal friction is present, that arising from transverse thermal currents.5 The effect of these may be reduced by increasing the transverse dimensions. The grain size and transverse dimensions of the specimens were thus so chosen as to render the vibrations simultaneously nearly isothermal with respect to intercrystalline thermal currents, and nearly adiabatic with respect to transverse thermal currents. The material for this investigation was a 70-30 brass, part of the same lot of material that was used for some earlier studies on the effect of strain on proportional limit.10 It had the following composition: copper, 70.59 per cent; zinc, 29.39; lead, 0.01; iron, 0.01. This material was received from the brass mill in the form of 3/4-in. rod, cold-drawn to final size by a 65 per cent reduction of area from material that had been annealed to give a grain size of 0.060 mm, Rods of appropriate length were annealed for 2 hr. at 400°C. and cooled in air. The annealed rods (grain size, 0.012 mm.) were then stretched in a tension-testing machine, using special grips1' to ensure axial application of the load. The rate

Jan 1, 1942

By Tell Ertl, Ernest E. Burgh

The Bureau of Mines is operating an oil-shale mine 10 miles west of Rifle, Colo., as part of its Synthetic Liquid Fuels program. The purpose of operating this mine is twofold: First, to supply oil shale to an experimental Processing plant; and second, to develop practices and select methods for mining oil shale on a commercial scale at the lowest practicable cost. The second purpose—that of finding out how to mine oil shale most cheaply—entails considerable mining research. Since the drilling of oil shale probably will be one of the major items of the mining cost, fattors that influence drilling speeds were investigated. One phase of the drilling program was the development of a successful hard-surfaced rock-drill bit1 that can be used for drilling long holes in oil shale without changing bits. During that phase, drilling speeds were measured and drill cuttings were collected for assay and for screen sizing, when the cuttings from the individual tests were compared, a definite relationship was shown to exist between the drilling speed and the coarseness of the drill cuttings. Procedure Measurement of Drilling Speeds All drilling in the tests was done with a Model UMB-D-99W Gardner-Denver wagon drill. Round-lugged drill steel, 141 in. in diameter was used in 6, 12, and 18 ft lengths with 1/8-in. Timken H-type bits hard-faced with acetylene tube borium. The tests were made in the lower adit of the Bureau of Mines oil-shale mine; the grade of the oil shale drilled ranged from 18.6 to 42.7 gal per ton, During the experiments air pressure varied between 75 and go psi while water pressure remained fairly constant at 75 psi. All holes were drilled almost horizontally. Each drill hole was thoroughly cleaned of cuttings after being collared. The advance of the drill along the carriage was measured and recorded as the drilling speed in inches per minute. Method of Collecting Cuttings TO facilitate the collection of cuttings, each drill hole was collared in a smooth, vertical portion of the oil-shale breast (Fig 1). A galvanized-stee1 sheet was fastened to the breast below the area in which the hole was to be drilled. The cuttings washed out of the drill hole by the drilling water were guided by the steel sheet into a wooden box. The cuttings that remained in the drill hole after the drill was withdrawn were scraped out and placed in the same box as the cuttings washed out of the hole. After the cuttings from a I-min. drilling interval were collected, they were placed in flat metal pans. The fines were allowed to settle and the clear water was decanted. The cuttings then were dried in the sun.

Jan 1, 1949

By R. Scot Clark

The reactively sputtered silicon dioxide capacitor -was developed for use in the monolithic integrated circuit, thereby requiring the capacitor to be fabricated on an oxidized silicon slice. The silicon dioxide film is formed by reactively sputtering silicon in a 50 pct Ar-50 pct 0, atmosphere and the physical properties are indistinguishable from silica. A practical upper limit of capacitance per unit area of 0.25 pf per sq mil is reached with the capacitors fabricated from the Si0, film. Methods shown to enhance this value are double stacking the thin-film capacitors and addition of aluminum to the silicon cathode during sputtering to form an aluminosilicate dielectric. A major contribution to the field of diffused integrated circuits would be the addition of auxiliary passive devices for replacement of diffused passive devices which are currently being fabricated for many integrated circuits. This accomplishment would increase the flexibility of the diffusion process in order to concentrate more on transistor parameters rather than on fabricating diffused resistors and capacitors. To pursue this technique, a thin-film SiOz capacitor was developed which has _ proven applicable to many integrated circuits. Although the SiOl capacitor has an application, the major deterrent from extending the integrated circuit into many useful areas such as low-frequency linear application is the low dielectric constant of SiO, and the low usable capacitance in an integrated circuit. This paper will cover the sputtering techniques used for forming the SiO, film, the physical and electrical properties of sputtered SiO,, and yield information concerning the application of these capacitors to integrated circuits. REACTIVE SPUTTERING Reactive sputtering is simply the sputtering of a metal or semiconductor in glow discharge in the presence of a reactive gas. The material to be sputtered is seen in Fig. 1 as the cathode of the system and is bombarded with positively charged ions of the ambient gas. In this investigation, the ambient gas is composed of 50 pct Ar and 50 pct O2 at a pressure of approximately 2.5 X l0-' Torr. The ions are accelerated toward the cathode in the cathode dark space by the electric field and upon striking the cathode effect a momentum energy transfer to the surface atoms. The transferred energy is sufficient to dislodge a surface atom or atoms and these particles diffuse through the glow discharge to the collector or anode. Several mechanisms of oxidation in the reaction chamber can be considered during reactive sputtering. 1) The surface of the cathode is oxidized and the oxide is subsequently sputtered. 2) Oxidation and sputtering occur simultaneously. 3) Sputtering of the cathode material is followed

Jan 1, 1965

By Howard Clarke, Cyril Stanley Smith, Clarence Zener

The importance of cold-working in the fabrication of metals is in marked contrast with our ignorance as to the nature of the accompanying internal changes. The failure of the usual methods of testing to provide a consistent picture of these internal changes has given the incentive for the present program of investigating the effect of cold-working upon the internal friction of metals measured at small strain amplitudes. It has previously been shown1,2,3, that the internal friction introduced by cold-work is removed by annealing at comparatively low temperatures. In the present paper the following additional factors are investigated: (I) frequency of measurement, (2) amount of cold-work, (3) temperature of measurement. Special care was taken to minimize the internal friction due to causes other than cold-work. The present paper is purely experimental, the authors believing that any conclusions drawn from this work as to the nature of the internal changes accompanying cold-work would be premature.'l PrEpaRations oF Specimens In this study of cold-work it was desirable to eliminate as far as possible all other sources of internal friction. One disturbing source is the lack of temperature equilibrium between adjacent crystallites of a vibrating specimen.4 Temperature equilibrium is aided by small grain size and by the use of low-frequency vibrations. Unless unduly long specimens are used, low frequencies can be obtained only in transverse vibrations. But in transverse vibrations a second source of internal friction is present, that arising from transverse thermal currents.5 The effect of these may be reduced by increasing the transverse dimensions. The grain size and transverse dimensions of the specimens were thus so chosen as to render the vibrations simultaneously nearly isothermal with respect to intercrystalline thermal currents, and nearly adiabatic with respect to transverse thermal currents. The material for this investigation was a 70-30 brass, part of the same lot of material that was used for some earlier studies on the effect of strain on proportional limit.10 It had the following composition: copper, 70.59 per cent; zinc, 29.39; lead, 0.01; iron, 0.01. This material was received from the brass mill in the form of 3/4-in. rod, cold-drawn to final size by a 65 per cent reduction of area from material that had been annealed to give a grain size of 0.060 mm, Rods of appropriate length were annealed for 2 hr. at 400°C. and cooled in air. The annealed rods (grain size, 0.012 mm.) were then stretched in a tension-testing machine, using special grips1' to ensure axial application of the load. The rate

Jan 1, 1942

By E. C. Babson

In order to explore the possibility of predicting reservoir performance from laboratory data, behavior of a hypothetical low-permeability reservoir has been estimated by applying data and methods currently available in the literature. A method of calculating decline in productivity index is discussed, and recoveries by internal gas drive, external gas drive and water drive are estimated. Introduction During the life of a producing oil property an operator is faced with many perplexing problems. Any attempt to determine proper well spacing, optimum rate of production, or the desirability of pressure maintenance requires the evaluation not only of a host of economic and practical operating factors but also of the future performance of the reservoir. Although in some cases economic or operating con-siderations may be of primary importance in planning a development or production program, the anticipated effect on ultimate recovery is more likely to be the decisive factor. The soundest basis for evaluating reservoir performance is past experience with oil fields, but pertinent data are difficult to obtain or apply under conditions normally encountered in California fields. Many of these fields are characterized by thick sections of alternating sands and shales complicated by faulting and rapidly changing lithologic conditions. Further compli- cations are introduced by haphazard development and production practices resulting from competitive conditions, changing demand for oil, and insufficient knowledge of structural conditions during early development. Even in the rare cases where development has been systematic and adequate, production policy has usually been controlled by economic and competitive factors rather than a desire to obtain information for use in future operations. In other words, comparable reservoirs in which development and production practices have been systematically varied are seldom found. In the light of these conditions, conclusions based on experiences usually lack the certainty required for decisions involving large sums of money. Some other method of attacking these problems is needed to supplement and orient field experience. Progress in laboratory investigations of the flow of oil, gas, and water through sands has been so rapid in recent years that these data may furnish such a supplementary approach in the near future. In order to explore this possibility the author has attempted to predict the behavior of one type of reservoir by applying published data and methods. Basic Data and Assumptions Calculations outlined in this paper are dependent upon a detailed knowledge of the properties of oil, gas, and water present in the reservoir and the portion of the total pore space originally filled by each. It is also necessary to know how the permeability of the sand to oil, gas, and water varies with the saturations of these fluids 20

Jan 1, 1944

By John Wulff, J. H. Hollomon

THE opinion has been widely held that the corrosion of alpha brass occurs by the selective solution of zinc. As late as 1939, Fink1 and Evans2 suggested that in the initial stage of the corrosion the selective solution of zinc produces copper-rich areas, which subsequently act as cathodes that facilitate the deposition of copper. Previously Abrams3 had shown that most of the copper found on corroded brass was redeposited and not residual, although he did not consider the nature of the initiation of the corrosion process. Enough data exist to indicate that any metal surface is sufliciently heterogeneous to provide the anodes and cathodes necessary for the initiation of corrosion .4,5 Thus, for alpha brass we need postulate not an initial selective solution of zinc, but only the existence of local differences in electrochemical potential. The heterogeneity responsible for potential differences may arise from oxide and other films as well as from alloy concentration gradients. Even though selective corrosion may be possible in some alloy systems, previous data on alpha brass as well as the experimental work reported in this paper do not require this supposition. Bengough and May,6 in a discussion of Abrams' paper, first suggested the addition of small amounts of arsenic (0.03 per cent) to alpha brass to prevent the redeposition of copper (dezincification). This suggestion was later substantiated by Nixon.7 In their studies of the inhibiting influence of arsenic, Evans and Fink were led to the selective-solution hypothesis in order to explain the initial corrosion of such alloys. "I hey believed that an arsenic film covered anode areas and prevented the formation of copper-rich spots capable of acting as cathodes for the deposition of copper. They (lid not consider, however, the mechanism by which the arsenical film was formed. Bengough and May,8 in another publication, suggested that the commonly found gray or black arsenical film may act to: (1) carry oxygen to the brass surface, (2) increase the discharge potential of copper on brass, (3) precipitate cuprous chloride from colloidal suspension in chloride solutions. Masing9 assumed that the film found upon arsenical brass was metallic arsenic. His measurements of hydrogen overvoltage of bulk arsenic led hint to believe that the inhibition of dezincification occurs because the overvoltage of a brass surface is lowered by the arsenic film. He reasoned that this film would permit the more rapid reduction of hydrogen at the cathode, which would prevent the redeposition of copper. In a recent paper,10 we have pointed out that the film formed (luring the corrosion of arsenical brass and copper in stagnant solutions is essentially arsenious oxide (AS208)

Jan 1, 1942

By E. C. Babson

In order to explore the possibility of predicting reservoir performance from laboratory data, behavior of a hypothetical low-permeability reservoir has been estimated by applying data and methods currently available in the literature. A method of calculating decline in productivity index is discussed, and recoveries by internal gas drive, external gas drive and water drive are estimated. Introduction During the life of a producing oil property an operator is faced with many perplexing problems. Any attempt to determine proper well spacing, optimum rate of production, or the desirability of pressure maintenance requires the evaluation not only of a host of economic and practical operating factors but also of the future performance of the reservoir. Although in some cases economic or operating con-siderations may be of primary importance in planning a development or production program, the anticipated effect on ultimate recovery is more likely to be the decisive factor. The soundest basis for evaluating reservoir performance is past experience with oil fields, but pertinent data are difficult to obtain or apply under conditions normally encountered in California fields. Many of these fields are characterized by thick sections of alternating sands and shales complicated by faulting and rapidly changing lithologic conditions. Further compli- cations are introduced by haphazard development and production practices resulting from competitive conditions, changing demand for oil, and insufficient knowledge of structural conditions during early development. Even in the rare cases where development has been systematic and adequate, production policy has usually been controlled by economic and competitive factors rather than a desire to obtain information for use in future operations. In other words, comparable reservoirs in which development and production practices have been systematically varied are seldom found. In the light of these conditions, conclusions based on experiences usually lack the certainty required for decisions involving large sums of money. Some other method of attacking these problems is needed to supplement and orient field experience. Progress in laboratory investigations of the flow of oil, gas, and water through sands has been so rapid in recent years that these data may furnish such a supplementary approach in the near future. In order to explore this possibility the author has attempted to predict the behavior of one type of reservoir by applying published data and methods. Basic Data and Assumptions Calculations outlined in this paper are dependent upon a detailed knowledge of the properties of oil, gas, and water present in the reservoir and the portion of the total pore space originally filled by each. It is also necessary to know how the permeability of the sand to oil, gas, and water varies with the saturations of these fluids 20

Jan 1, 1944

By John Wulff, J. H. Hollomon

The opinion has been widely held that the corrosion of alpha brass occurs by the selective solution of zinc. As late as 1939, Fink1 and Evans2 suggested that in the initial stage of the corrosion the selective solution of zinc produces copper-rich areas, which subsequently act as cathodes that facilitate the deposition of copper. Previously Abrams3 had shown that most of the copper found on corroded brass was re-deposited and not residual, although he did not consider the nature of the initiation of the corrosion process. Enough data exist to indicate that any metal surface is sufficiently heterogeneous to provide the anodes and cathodes necessary for the initiation of corrosion.4,5 Thus, for alpha brass we need postulate not an initial selective solution of zinc, but only the existence of local differences in electrochemical potential. The heterogeneity responsible for potential differences may arise from oxide and other films as well as from alloy concentration gradients. Even though selective corrosion may be possible in some alloy systems, previous data on alpha brass as well as the experimental work reported in this paper do not require this supposition. Bengough and May16 in a discussion of Abrams' paper, first suggested the addition of small amounts of arsenic (0.03 per cent) to a1pha brass to prevent the redeposition of copper (dezincification). This suggestion was later substantiated by Nixon.7 In their studies of the inhibiting influence of arsenic, Evans and Fink were led to the selective-solution hypothesis in order to explain the initial corrosion of such alloys. They believed that an arsenic film covered anode areas and prevented the formation of copper-rich spots capable of acting as cathodes for the deposition of copper. They did not consider, however, the mechanism by which the arsenical film was formed. Bengough and May,a in another publication, suggested that the commonly found gray or black arsenical film may act to: (I) carry oxygen to the brass surface, (2) increase the discharge potential of copper on brass, (3) precipitate cuprous chloride from colloidal suspension in chloride solutions. Masing9 assumed that the film found upon arsenical brass was metallic arsenic. His measurements of hydrogen overvoltage of bulk arsenic led him to believe that the inhibition of dezincification occurs because the overvoltage of a brass surface is lowered by the arsenic film. He reasoned that this film would permit the more rapid reduction of hydrogen at the cathode, which would prevent the redeposition of copper. In a recent paper,10 we have pointed out that the film formed during the corrosion of arsenical brass and copper in stagnant solutions is essentially arsenious oxide (As2O3),

Jan 1, 1942

By John Wulff, J. H. Hollomon

The opinion has been widely held that the corrosion of alpha brass occurs by the selective solution of zinc. As late as 1939, Fink1 and Evans2 suggested that in the initial stage of the corrosion the selective solution of zinc produces copper-rich areas, which subsequently act as cathodes that facilitate the deposition of copper. Previously Abrams3 had shown that most of the copper found on corroded brass was re-deposited and not residual, although he did not consider the nature of the initiation of the corrosion process. Enough data exist to indicate that any metal surface is sufficiently heterogeneous to provide the anodes and cathodes necessary for the initiation of corrosion.4,5 Thus, for alpha brass we need postulate not an initial selective solution of zinc, but only the existence of local differences in electrochemical potential. The heterogeneity responsible for potential differences may arise from oxide and other films as well as from alloy concentration gradients. Even though selective corrosion may be possible in some alloy systems, previous data on alpha brass as well as the experimental work reported in this paper do not require this supposition. Bengough and May16 in a discussion of Abrams' paper, first suggested the addition of small amounts of arsenic (0.03 per cent) to a1pha brass to prevent the redeposition of copper (dezincification). This suggestion was later substantiated by Nixon.7 In their studies of the inhibiting influence of arsenic, Evans and Fink were led to the selective-solution hypothesis in order to explain the initial corrosion of such alloys. They believed that an arsenic film covered anode areas and prevented the formation of copper-rich spots capable of acting as cathodes for the deposition of copper. They did not consider, however, the mechanism by which the arsenical film was formed. Bengough and May,a in another publication, suggested that the commonly found gray or black arsenical film may act to: (I) carry oxygen to the brass surface, (2) increase the discharge potential of copper on brass, (3) precipitate cuprous chloride from colloidal suspension in chloride solutions. Masing9 assumed that the film found upon arsenical brass was metallic arsenic. His measurements of hydrogen overvoltage of bulk arsenic led him to believe that the inhibition of dezincification occurs because the overvoltage of a brass surface is lowered by the arsenic film. He reasoned that this film would permit the more rapid reduction of hydrogen at the cathode, which would prevent the redeposition of copper. In a recent paper,10 we have pointed out that the film formed during the corrosion of arsenical brass and copper in stagnant solutions is essentially arsenious oxide (As2O3),

Jan 1, 1942

By F. L. Fayers, J. W. Sheldon

It has been a matter of concern to the petroleum industry to determine what effect the capillary term has on saturation profiles, since these profiles determine ultimate economic oil recovery. In their original solution of the non-capillary two-phase flow problem, Buckley and Leverett&apos; showed that solutions of the two-phase flow equations became multiple-valued in saturation. Since it is physically unrealistic for saturation to have more than one value at a given position, Buckley and Leverett resolved this difficulty by introducing a saturation discontinuity or "shock". They evaluated the strength and position of the shock from material balance considerations. Recently Sheldon, Zondek and Cardwell&apos; have demonstrated that the method of characteristics can be used to solve this and related problems. Owing to capillarity it is physically unlikely that saturation shocks occur in an oil reservoir. Buckley and Leverett suggested that the effects of interfacial tensions should be included in the fluid flow equations, if the solutions are to give continuous single-valued saturation distributions. An analogy to this problem is encountered in the theory of wpersonic compressible flow where the introduction of the viscous term eliminates shock behavior. However, the mathematical form of the capillary term in the porous medium fluid flow equations differs from the viscosity term in the Navier-Stokes equation. This paper demons~rates that capillarity does eliminate the triple-valued behavior and indicates the magnitude of capillary and gravity effects on the saturation profile for a variety of systems. DERIVATION OF LAGRANGIAN AND EULERIAN FORMS OF THE DISPLACEMENT EQUATION The equations representing capillary, two-phase flow are given by applying Darcy&apos;s law and a material balance condition for each phase. The familiar Darcy and material balance equations are as follows. ^U*^0......<3, ^^r=°......c« S is the saturation of the water phase, and B the inclination of the porous medium. Pressures in the oil and water phases differ by the capillary pressure associated with the interfacial tensions, P.-P.=pAS)..........(5) where p, is determined experimentally. The porous medium is said to be oil-wet if P, > Po and water-wet if P, < Po. If the fluids are incompressible and the total volumetric flow rate, q, is constant, then v, f v,, = q...........(6) By eliminating P.. /»,, and v, from Eqs. 1, 2, 5 and 6 it follows that It is convenient to normalize the variables in the problem after the manner of Handy and Hadley% so that Eqs. 3 and 7 become where j,, = —, the fractional flow of water; (F, is a capillary pressure normalizing constant.) Substitution of Eq. 9 into Eq. 8 gives In this partial differential equation for an initial value

By S. W. Poole, J. A. Rosa

THE object of this investigation was to determine the distribution of chemical elements within a large, killed alloy-steel ingot, by sulphur printing and quantitative chemical analysis. With regard to segregation in steel ingots, the factors of melting practice, pouring temperatures, mold design, etc., and the mechanics of solidification and its attendant functions, have been the subject of intensive study over a period of years by many investigators. Notable among these is the Sub-Committee of the No. 5 Committee of the British Iron and Steel Institute. For an exposition 0f the many problems awaiting the researcher on the heterogeneity of steel ingots, the methods of attacking these problems, together with some of the answers, the reader is referred to the nine reports and numerous papers of that committee and its membership. STEELMAKING DATA Melt.-The heat was made in a 70-ton direct arc-type basic electric furnace. Furnace charge consisted of nickel-chromium-molybdenum steel, most of it in the form of heavy mill scrap. All of the scrap was charged cold. The standard double-slag method was employed, the heat being finished under a strong carbide slag. Aluminum was used as the deoxidizing agent. Tapping temperature, as observed with an optical pyrometer, was 3000°F. Detail of Mold and Ingot Size.-Details of the ingot mold and the ingot itself are given in Fig. I. Teeming.-Teeming temperature of the subject ingot as observed with an optical pyrometer was 2830°F. The ingot was top-poured directly through a 1 ½ in. nozzle and required 7 min., 44 sec. to fill up to the hot top. An additional 2 min., 30 sec. were required to fill the hot top. The ingot was stripped hot and charged with the remainder of the heat into an annealing furnace. There it was given the standard annealing cycle used for these ingots. Subsequent heat-treatments are described elsewhere in this paper. SECTIONING AND PREPARATION FOR SULPHUR PRINTING Because of previous favorable experience in sectioning ingots up to 21 ½ in. square with an oxyacetylene torch, it was decided to use this method to section the subject ingot. The ingot was preheated to 900°F. for 12 hr. It was then placed on a steel bed in a horizontal position so that the longitudinal axis was level and parallel to the track of the motor-driven torch carrier. Two cuts were made, one 2 in. above the longitudinal axis and the other 2 in. below. The time required to make each cut was approximately 75 min. The cutting operation is shown in Figs. z and 3. The 4-in. thick slab was mmediately charged into a furnace standing at 1200°F.

Jan 1, 1944

By H. L. Washburn, G. L. Judy

DURING recent years thermal drying of fine coal has increased at a tremendous rate, but very little discussion of the several excellent units has found its way into print. This is particularly true for operating and maintenance costs, and the operator who is faced with his first fine coal drying installation must depend chiefly on hearsay when trying to determine what his drying costs will be. This article presents the maintenance and operating costs for the seven drying units now installed in preparation plants of Pittsburgh Consolidation Coal Co. and lists other cost information that should be obtained before any particular drying unit is selected. Plant Description: The fine coal drying section of Plant A consists of four Flash driers operating on two furnaces equipped with spreader stokers. Plant B has one Flash drier operating (together with a screen-type coarse coal drier) on one furnace equipped with a spreader stoker. Plant C has two Cascade-type driers operating on a single furnace with a spreader stoker. Plant Costs: Tables I, 11, and III present the cost of operating these units as reported by the individ- ual plants. As can be seen, the labor rates vary, depending on the amount of overtime paid and the time interval the cost figures cover. There is considerable variation in thermal units used and in the value given to the thermal units at the particular mine. There is some difference in power costs at a given location and, finally, the power costs of conveyors leading to the drying plants were included in some cases. Maintenance supply costs are simply a tabulation of the cost of all supplies used on the driers, furnaces, and stokers. Depreciation is not included, and the tables omit items such as supervision, plant cleanup, oiling, general maintenance, taxes, and insurance. These tables in themselves do not offer a true comparison of costs. This information must be reduced to some common basis. But what is this common basis to be? A mine operator must calculate his overall mine and preparation plant costs in terms of tons of clean coal sold. Unit costs on pieces of equipment within a given plant are correctly calculated on the same basis. However, when unit costs are compared between one plant and another, some thought must be given to unit design. This reasoning is similar to the old argument concerning i solid bowl vs basket-type centrifuges—the units themselves are not directly comparable without their supporting equipment. So it is with thermal driers. Except in rare cases of low entering moisture i content, driers are sized on the basis of water that must be evaporated, while the wear each unit gets de-

Jan 1, 1958

By A. Lubinski, W. D. Greenfield

Bumper subs are currently used in offshore operations to permit a constant weight to be carried on the bit while drilling, regardless of the vertical motion imparted to the drill pipe by drilling vessel heave. As shown in this paper. the vertical motion of the lower end of the drill pipe (the bumper sub end) may be appreciably greater than the vessel heave. Therefore, the necessary stroke of bumper .rubs for successful operation is greater than thought in fie past. Also, there is an appreciable tendency of the drill pipe to buckle above the unbalanced type of bumper sub. Thus, more drill collars than previously used should be carried above unbalanced bumper subs to keep drill pipe straight. INTRODUCTION Drilling bumper subs are placed in the drilling string for various reasons. This paper is concerned with their use only as an expansion and contraction joint while drilling from a floating rig. In this application the bumper subs are normally located just above the drill collars and their function is to allow the driller to maintain accurate weight control on the bit regardless of up-and-down movement of the drilling vessel. This paper analyzes the effects of bumper subs on the drilling string and presents recommendations for their future use. When subjected to vertical oscillations, the drilling string behaves like a long, distributed system of mass and spring. The magnitude of vertical motion at the bumper sub is always greater than the heave of the drilling vessel due to the dynamic reponse of the drilling string. The ratio of these motions increases with the length of the drilling string, and may reach values of 1.5 or even 2 with strings 16,000 ft long. Thus, the total travel required in bumper subs can be considerably more than the motion of the drilling vessel. Lack of knowledge of this fact could have contributed to problems previously experienced with bumper subs. This fact can also lead to fatigue problems in the drilling string for very deep wells. Satisfactory operation should be obtainable whether hy-draulically balanced or unbalanced bumper subs are used in the drilling string. Theoretically, the balanced sub is preferable since its use does not require placing drill collars above the bumper sub to prevent drill-pipe buckling, an inherent characteristic of the unbalanced bumper sub. The current method of calculating weight of drill collars required to prevent helical buckling of drill pipe above unbalanced bumper subs is erroneous. Placing drill collars above the sub to prevent drill-pipe buckling has the same effect on dynamic response as increasing the length of the drilling string by an equal weight of drill pipe. Thus, total travel required in the subs is increased. Means for calculating the correct weight, which is much greater than previously thought, are given in this paper. BALANCED VS UNBALANCED BUMPER SUBS A drilling bumper sub is essentially a telescopic joint capable of transmitting torque at every position of its stroke. Thus, it allows the operator to isolate the weight of the drilling string from the weight of the drill collars above the bit. This permits the driller on a floating rig to maintain accurate control over the weight on bit — a control that is unaffected by vertical motion, due to wave and tide action of the drilling vessel. UNBALANCED BUMPER SUBS The unbalanced bumper sub is simply a splined tele~copic joint (Fig. I). Ordinarily, this arrangement will operate satisfactorily, but the presence of drilling fluid under pressure results in a pressure force that acts downward on the drill collars and bit, tending to open or extend the bumper sub. This downward force is equal to the pressure drop across the bit times the area indicated by diameter d2 in Fig. 1. Denoting this force by Fd, and the pressure drop across the bit by ?p yields Fb = (p/4)d22(?P) .........(1) There is also an upward-directed force given by Fu = (p/4) d22-d21)(?p) .......(2) which puts the drill pipe immediately above the bumper sub in compression, resulting in helical buckling. However, buckling is actually more severe than expected in that buckling occurs as if the compression were equal to Fd, rather than to Fu. This surprising phenomenon is well known as far as tubing is concerned;1-3 but, in contrast with the case of tubing, this force may shorten drill pipe only a few inches. Thus, this cannot explain the operating difficulties that sometimes have been encountered. However, having the drill pipe in compression and helically buckled is contrary to current practice; therefore, drill collars whose weight in mud is equal to the force Fd should be added above the bumper sub. Since the value of Fd depends on the pressure drop across the bit, the