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By J. H. Swisher

The solubility of graphite in CaO-S2O2-Al,O3 slags was measured by equilibrating slag samples with graphite crucibles and CO gas. Carbon contents as high as 2 ut pct were obtained in CaO-saturated, CaO-A1,O3 slags, and 1.3 wt pct in slags of the composition CaO.Si0,. Although the observed conditions for Sic formation were in agreement with those predicted from thermodynamic data, CaC, was found to form at a lower temperature than predicted frotn thermodynamic data. From measurements of the equilibrium carbon content as a function of CO Partial pressure, it was found that carbide ions dissolve in CaO-A12O3 melts with a valence of minus two. The carbon content increased with CaO concentration in Ca0-Al,O3 melts and increased with SiO, content along the CaO'AlO3-CaOSi0 join in the ternary system. When solid CaC2 was added to CaO-A12O3 and CaO-SiO2-A12O3 slags, it was found that one of the oxides in the slag was reduced by the carbide (Al2O3 in the forrner and SiOz in the latter). In electric furnace steelmaking, a double-slag practice is frequently used to meet alloy specifications. Initially a flush slag, which is oxidizing in nature, is used to remove phosphorus and carbon from the steel bath. Later in the refining period, the flush slag is replaced by a highly reducing carbidic slag. When calcium carbide is formed in or added to a finishing slag, the slag is effective as a desulfurizing agent and also permits alloying elements such as chromium, vanadium, and tungsten to be added to the slag in the form of oxides. The oxides are readily reduced by calcium carbide, thereby minimizing the use of expensive ferroalloys. More work has been done on the thermodynamics of silicon carbide in slags than on calcium carbide. Baird and alor' and Kay and alor' determined the free energy of formation of Sic by measuring the partial pressure of CO in equilibrium with solid silica, silicon carbide, and graphite. Using a similar technique, they determined SiOz activities in CaO-SiOz and Ca0-Si0,-A1203 slags. Rein and chipman3 also determined the free energy of formation of Sic using slag-metal equilibrium measurements. A literature survey has uncovered only one experimental study of the behavior of CaC, in slag systems. Shanahan and cooke4 report the results of some preliminary experiments on the solubility and stability of CaC, in a CaO-A1,03 and a Ca0-Si0-A1,03 slag at a temperature of about 1500". The carbon solubility as CaC, in a slag containing 50 pct CaO and 50 pct A1203 was reported to be 0.6 pct. They also review earlier work on the binary CaO-CaC, system. A eutectic exists in this system, but various investigators disagree on the eutectic temperature and composition. eal has given an explanation for carbide furnace erruptions in terms of the thermodynamic properties of CaC,; his analysis is not based on experimental data, but on compiled data for the free energies of formation of CaC, and CO.' , These data for steel-making temperatures are all extrapolated from the results of low-temperature measurements. In the experiments described in this paper, slag samples were equilibrated with graphite crucibles and with mixtures of CO and argon or with CO gas at 1 atm total pressure for measurement of the carbon solubility. Most of the work was done on Ca0-A1203 binary slags, although in some experiments CaO-SiO, and Ca0-Si0,-A1,03 slags were used. EXPERIMENTAL Slag samples of the desired composition for the solubility measurements were obtained by blending pre-fused master slags. The master slags were prepared by fusing mixtures of reagent-grade CaC03 with either A1,03 or Si0, in a graphite crucible. The master slags were crushed, then decarburized in air in a muffle furnace at 1200O C. A schematic diagram of the apparatus is shown in Fig. 1. The source of carbon for the solubility meas-

Jan 1, 1969

By M. G. Benz

The superconducting performmce of diffusion-processed Nb3Sn is influenced by its micro structure. High isotropic transverse current density may be achieved in this material by a process which forms a precipitate of ZrO, within the Nb3Sn. FOR an ideal type-I1 superconductor, little or no transport current can be carried in the mixed state; i.e., little or no transport current can be carried above the lower critical field H,,, where the field penetrates abruptly in the form of current vortices or fluxoids, even though full transition to the normal state does not occur until the upper critical field H,,.' Fortunately, nonideal type-I1 superconductors can be readily obtained and these carry large transport currents up to the upper critical field H. Both theoretical and experimental investigations have attributed this current-carrying capability for nonideal type-I1 superconductors to pinning of the fluxoid lattice by heterogeneities in the microstructure of the superconducting material. These heterogeneities may take the form of dislocations or dislocation clusters,2"5 grain boundaries: structural imperfections introduced by phase transformations; radiation damage,8"10 or precipitates.11"15 Nb3Sn formed by diffusion processing is a type-I1 superconductor. Heterogeneities are needed for high superconducting critical currents above H,,. This paper will cover: a) what the microstructure of diffusion-processed NbSn looks like; b) what changes in the microstructure take place when the system is doped with precipitates, and c) how these changes in microstructure influence the superconducting critical currents. EXPERIMENTAL Preparation of Samples. Diffusion processing was used to form the Nb3Sn. The procedure used was as follows: a) coat the surface of a niobium tape with tin; b) heat-treat this tape at a temperature above 930°C to form a layer of Nb3Sn at the Sn-Nb interface. Such a layer of NbsSn is shown in Fig. 1 The thickness of the NbsSn layer formed was controlled by the time and temperature of the heat treatment. The same general procedure was used for preparation of both undoped samples and samples doped with a precipitate. An additional step was included in the preparation of the doped samples which consisted of internal oxidation of zirconium to form ZrOn. The details of the doping process will be reported in a later paper. Sample Testing. The Nb3Sn tape samples were soldered to a copper or brass shunt. Current and voltage leads were then attached to the sample in the usual four-probe resistance measurement configuration. The sample was cooled to 42°K. In some cases it was cooled in the presence of a high magnetic field and in other cases with the field turned off. The results were the same for both cases. The samples were oriented in a configuration with field transverse to current but could be rotated such that the angle between the field vector and the wide side of the tape sample could be changed. Measurements up to 100 kG were done in a superconducting solenoid and measurements above 100 kG in a water-cooled copper magnet at the MIT National Magnet Laboratory. Once the test field was reached, the current in the sample was increased until voltage was detected across the sample. The critical current was taken as the current at which voltage was first detected in excess of background noise. In most cases this was 1 to 2 x 10~6 v for a— in.-wide sample carrying several hundred amperes with a in. separation between voltage leads and with a 10 "-ohm shunt resistance. RESULTS AND DISCUSSION Microstructure. Examination of the microstructure of the undoped Nb3Sn shows rather large-diameter (1 to 2 columnar grains growing outward from the niobium surface toward the tin surface. As the layer is made thicker by longer diffusion times, these grains grow longer. Few new grains are started. Transmission electron microscopy shows little or no second-phase material within the bulk of the Nb3Sn layer. The microstructure of a diffusion-processed NbsSn layer changes quite drastically when the system is doped so as to form a precipitate within the NbsSn layer. Instead of large-diameter columnar grains of NbaSn forming, smaller-diameter (0.5 to 1 ) equiaxed grains of Nb3Sn decorated with the precipitate form. Fig. 2 shows a transmission electron micrograph of a Nb3Sn layer doped with zirconium oxide. This layer has been etched so that one may look between the grains

Jan 1, 1969

By P. C. Johnson, B. A. Stein

This paper describes a study of the effects of combined heat treatment and explosive loading on the mechanical properties of high-strength steels. nis program investigated two distinct areas: 1) the effect of shock waves, without gross irreversible defmmution, on a 3-Cr steel at various stages of heat treatment; and 2) the effect of rapid deformation (explosive forming) on H-11 and D6-AC steels in the metastable austenitic state. The mechanical properties of these steels were improved, in some cases markedly, as a result of these treatments. ,AUSFORMWG, which requires the plastic deformation of metastable austenite, is a process which can appreciably improve the properties of selected alloy stee1s.l,2 The Ausform process significantly increases the strength of these steels without decreasing their ductility. The properties at high temperatures are also improved through a change in the response of the steels to tempering. Although the mechanism by which ausforming alters the properties of these steels is not fully understood, it appears that the dislocation arrays produced by deformation of the metastable austenite influence the structure of the martensite on subsequent transformation. This, in turn, affects the strength, ductility, and tempering response of the martensite. This research used chemical explosives to deform steels at various stages in their heat treatment in order to improve the properties of these steels. The explosive energy is used in two ways; 1) high-pressure shock waves are propagated through the steel to produce extensive microscopic shear strain without causing a large irreversible change in shape, and 2) explosive energy is used to cause extensive macroscopic plastic strain in the metastable austenitic state (explosive forming). I) AUSFORMING WITH INTENSE SHOCK WAVES The steel used in this phase of the research was an alloy having a nominal composition 0.43 pct C, 3.0 pct Cr, 1.5 pct Ni, and 1.5 pct Si. The steel was subjected to intense shock waves in three conditions: 1) in the metastable austenitic state, 2) in the tempered mar tens itic state, and 3) in the tempered martensitic state after ausforming by conventional techniques. The specimens were in the form of disks 2.75 in. in diam and 5/16 in. thick. These were incorporated into a specimen assembly consisting of two disks pressed into a 5 by 5 by 1 in. block of stainless steel, Fig. 1. Spalling (or scabbing) is confined to the front disk. The specimen is protected from oxidation and decarburization by the surrounding metal. The temperature of the assembly is monitored by a thermocouple inserted into one side of the stainless steel block. The assembly is positioned over an oil reservoir which serves both as a means of catching the disks and as quenching medium for the disks shocked under ausforming conditions. Plane shock waves are introduced into the assembly by a metal driver plate impacting the top surface of the block. The driver plate is accelerated by a chemical explosive sheet supplied by E. I. du Pont de Nemours & Co. All the specimens were subjected to plane shock waves having a peak pressure of approximately 430 kbar. The pressure is that quoted by G. E. Dieter for the plane wave generator used in this work.' The driver plate used was 1/4 in. thick, so that the initial pulse was essentially a l/2-in.-wide square wave. The attenuation of the peak pressure during the subsequent 1/4 in. is estimated to be less than 5 pct. The shock front induces a temperature rise, a portion of which is irreversible. Rough estimates (+25 pct) of this temperature rise have been made for iron shocked at room temperature.4 For a 500-kbar shock wave, the temperature rise in the shock front is about 700°F, and is held for a time of the order of microseconds. The irreversible temperature rise, which remains after the shock wave passes, is about 450°F.4 The disks are quenched to room temperature within a few seconds of the shock treatment. It should be emphasized that the temperature rises given above are estimates for pure iron at room temperature, and are not necessarily true for the tests made in this work. The disks shocked at temperatures in the metastable austenitic range were austenitized in the stainless steel assembly in a furnace protected from the firing area. The assembly was removed from the furnace and placed over the recovery reservoir. The plane wave generator was then positioned and

Jan 1, 1963

By W. J. Maraman, D. R. Harbur, J. W. Anderson

A device for rapidly quenching liquid metals into thin platelets has been developed at the Los Alamos Scientific Laboratory. This rapid quenching equipment is built around the technique of catching a molten drop of metal between a rapidly closing plate and a stationary plate. The design and operation of this unit are described. The closing speed of the smasher plate at impact is 12.6 ft per sec. The quenching rate for this device is controlled by the interface resistance between the plates and the platelet, and is dependent upon the heat content and density of the material being quenched. The initial quenching rate down to the freezing point of the platelet material is lo5º to 106ºC per sec. After an isothermal delay, which is poportional to the heat of fusion of the platelet material, the final cooling rate down to the temperature of the smaslier plates is l04ºto 105cº per sec. RAPID heating of metals by capacitor discharge and other methods has provided the metallurgist with a useful tool for probing into the kinetics of phase changes and the many nonequilibrium phenomena which occur during rapid temperature changes. Equally interesting studies can also be made on metals and alloys which are rapidly cooled from the liquid state.' Studies in this field have been limited, however, because the rates at which metals could be cooled were many orders of magnitude slower than the rates possible for heating. In recent years many new laboratory methods have been developed to rapidly cool metals from the liquid state to ambient temperature and below.2"4 All of these methods involve spreading a liquid drop of metal into a thin foil in a very short time. The methods developed have varied from ejecting a drop of molten metal at the inside surface of a rotating cylinder or stationary curved plate to catching a falling drop of molten metal between rapidly closing plates. The equipment which has been developed at the Los Alamos Scientific Laboratory for rapidly cooling molten materials uses the latter of these two approaches. The basic design, operation, and initial results of this rapid quenching device are given in this report. APPARATUS The drop smasher, which is now being used to obtain rapidly cooled metal foils, is shown in Fig. 1. Basically the device consists of a smasher plate which is driven by a solenoid into a stationary plate. The solenoid is activated by a drop passing through the photoelectric cell and is powered by discharging an adjustable 350-v capacitor bank with a 66-amp peak current into it. This power supply is designed so that the solenoid is powered for 2 m-sec after plate closure to minimize the rebound effect. There is an adjustable time-delay mechanism between the photoelectric cell and the solenoid. Both smasher plates have changeable inserts so that a variety of materials can be used to smash the molten drop. The shaft of the moving plate is guided in an adjustable housing which has ball-bearing walls. The cabinet shown to the left of the drop smasher in Fig. 1 contains the power supply and receiver for the photoelectric cell, the time delay mechanism, and the capacitor bank. The drop smasher can be placed inside a vacuum chamber, for use with radioactive materials, with the upper plate forming the lid, as shown in Fig. 2. On top of the vacuum lid is an induction coil, powered by an Ajax induction generator, which is used to melt drops from the end of the rod extending through the vacuum seal on top the quartz tube. OPERATION The drop smasher shown in Fig. 2 is operated in the following manner. The smasher plates are separated and the unit is lowered into the vacuum chamber using a pressurized cylinder. The induction coil, quartz tube, and lid with sliding vacuum seal are then assembled on top the vacuum chamber. A rod of the material for rapid quenching studies is connected to the rod extending through the sliding vacuum seal. The vacuum chamber is then evacuated and the desired atmosphere established. The photoelectric cell is turned on, and the capacitor bank is charged and armed. Power is supplied to the induction coil, and the rod of material for rapid quenching studies is lowered into the induction field. A molten drop forms on the end of the rod, drops off, falls through the light beam of the photoelectric cell, and is then caught between the smasher plates. .

Jan 1, 1970

By D. V. Edmonds, C. J. Beevers

Tensile tests were carried out on a! titanium containing 850, 1250, and 2700 ppm 0, and up to -500 ppm H. The tests were performed at -196", -78", 20°, 150°, and 300°C at a strain rate of -1.0 x 10??3 sec-1. Increasing oxygen content, increasing grain size, and decreasing test temperature resulted in enhanced embrittlement of the a titanium by the hydrogen additions. Metallographic observations showed that this can be correlated with the influence of these parameters on the introduction of cracks into the a! titanium by fracture of titanium hydride precipitates. CRAIGHEAD et al.1 reported that the hydrogen content normally found in commercial-purity a! titanium (60 to 100 ppm) was sufficient to cause a substantial lowering of the impact strength, and they attributed this embrittling effect of hydrogen to the precipitation of titanium hydride. Lenning et al.&apos; found that in commercial-purity a titanium there is an almost complete loss of impact strength at about 200 pprn H, which is approximately half the value needed to eliminate the impact strength of high-purity a titanium. They also showed that the presence of 3000 ppm hydrogen reduces the room-temperature tensile ductility of commercial-purity material to a value of the order of 10 pct; the corresponding hydrogen concentration for high-purity titanium is over 9000 ppm. It thus appears that the detrimental effect of hydrogen on the mechanical properties of commercial-purity titanium becomes evident at much lower hydrogen contents than for high-purity titanium. The main difference between the two types of a titanium might be expected to be the higher level of interstitial impurity in the commercial-purity grade. Jaffee et a1.3 studied the influence of temperature and strain rate on the hydrogen embrittlement of high-purity and commercial-purity ! titanium. In general, the behavior was the same for both materials; embrittlement was enhanced by decreasing temperature and increasing strain rate. Recent results from tests on commercial-purity a titanium containing 850 ppm O and varying amounts of hydrogen up to -500 ppm showed that the degree of embrittlement by hydrogen is intimately related to the fracture characteristics of titanium hydride precipitates.4 The present paper considers the interrelationship between the mechanical properties and micro-structural features of commercial-purity a! titanium containing 850, 1250, and 2700 ppm 0 and varying amounts of hydrogen up to -500 ppm. 1. EXPERIMENTAL PROCEDURE Three types of commercial-purity titanium supplied by IMI* were used in the investigation, and for the *Address: Witton, Birmingham 6, United Kingdom. purpose of this paper are designated Ti 115, Ti 130, and Ti 160. The principal impurity elements are given in Table I. The material was received in the form of 12.7 mm diam bars having a fully recrystallized structure. Tensile specimens with a round cross-section of 4.5 mm diam and a gage length of 15.2 mm were machined from the bars. In order to develop the same grain size (mean linear intercept of grain boundaries) in each of the three types the specimens were annealed under a dynamic vacuum of <10?5 mm Hg, Table 11. Specimen hydriding was carried out in a modified Sieverts apparatus;&apos; hydrogen was taken into solution at 450°C and after holding the specimens at this temperature for 24 hr they were furnace-cooled to room temperature at an average rate of -100 C deg per hr. By this method nominal hydrogen contents of 0, 50, 100, 250, and 500 ppm were introduced into specimens of Ti 115, Ti 130, and Ti 160 (100 ppm (wt) -0.5 at. pct). The actual hydrogen contents were calculated from the weight differences obtained by weighing the specimens before and after the hydriding treatment. Tensile tests were carried out at temperatures of -196", -78", 20°, 150°, and 300°C on a 10,000 kg In-stron machine at a nominal strain rate of -1.0 x 10-3 sec-1. Fractured specimens were sectioned in planes parallel to the tensile axis, mechanically polished to 0.25 µm grade of diamond paste, and then attack polished using a solution containing by volume 99 parts H2O, 1 part HF, and 1 part HNO3. Although the latter treatment unavoidably opened out cracks and voids visible after mechanical polishing, it did reveal the grain structure, titanium hydride morphology, and deformation twinning structure.

Jan 1, 1970

By T. S. Lovering

The use of geothermal energy is expanding very rapidly. This type of energy has proven commercially profitable for generation of electricity, for space heating, process heating, auxiliary heating of water in conventional steam power plants and for recovery of chemicals contained in natural hot water and steam. Two types of geothermal energy sources are recognized: 1) hot springs in regions of nearly normal heat flow that tap a deep reservoir through which water moves slowly to a hot springs conduit and then rapidly to the surface; 2) hyperthermal areas in which the water is heated by a relatively concentrated heat source related to volcanicity. If there is a geologic trap that provides a geologic analog to a steam boiler, as at Larderello, Italy, the hyperthermal area will have a convecting system that develops superheated water at relatively shallow depth and may provide natural steam in large quantities. If a hyperthermal area is to be productive for a long time, the underflow into the reservoir should be slow enough to allow the heat source and convective system to heat the underflow to the working temperature, and the production rate must not exceed this rate of underflow. A model based on a typical aquifer suggests that the rate of movement of water through the reservoir be such that a few years are spent in transit between isotherms that are spaced about 2°F apart. The possibility of finding blind geothermal areas is illustrated by discussion of the techniques developed in evaluating the subsurface temperatures in the East Tintic district of Utah where a map of isotherms at water level (2000 to 2000 ft below the surface) shows that a hyperthermal area may exist a short distance southeast of the mining district. Very nearly all of the energy that man currently uses comes ultimately from the sun&apos;s radiation. This includes water power, fuels such as wood, peat, coal and petroleum, the wind and all our animal power. In the paper summarizing a conference on solar energyl6 the average amount of solar energy received daily on the earth is taken at about 1 cal per m2 per min or slightly less than 2 pcal per cm2 per sec; this is almost exactly the amount of energy on the average that the earth liberates in regions of normal geothermal gradient due to its own internal heating. In many places, however, the energy released is many times the average and in some of these hyperthermal areas, geothermal steam is used for generation of electricity, and hot springs are used for heating buildings and private dwellings, process heating, auxiliary heating of water in conventional steam power plants, and chemicals may be recoverable from both hot water and steam. The use of hot springs waters for heating houses goes back hundreds of years but until recently was confined to a few dwellings close to the hot springs. In Korea, some houses had hot spring water channeled through conduits in the floor centuries ago and thus the Koreans can be credited with pioneer development of radiant heating. In Iceland at present nearly a third of the population uses natural thermal water for domestic heating." The Reykjavik system pipes hot spring water at about 94°C throughout the city and has devised insulated double pipes that allow the water to be piped for some 25 km with a drop of only 1°C for every 5 km. The actual cost to the Icelandic consumer is only one-third the cost of heating by imported coal and yet the industry is one of the most profitable in Iceland. The most profitable use of geothermal energy has been its conversion into electricity which can be transmitted economically much greater distances than hot water. The largest installation at the present time is that at Larderello, Italy, where the Count of Larderello began to experiment in the production of electricity from geothermal steam 60 years ago — in 1904. He installed his first steam turbine, with a capacity of only 250 kw, in 1912 as the result of a local quarrel with the power company which furnished the current required in the Larderello chemical industry - an industry that then dated back nearly a century. As experience was gained in drilling deep holes to tap geothermal steam and in converting it to electric power, the capacity of the installation of Larderello gradually increased, but was all destroyed by the Germans during their retreat from Italy in the closing

Jan 1, 1965

By M. D. Arnold, P. B. Crawford, P. C. Hall

An experimental study has been made to determine the optimum flooding pressures for four different oils. The oil formation volume factors ranged from 1.08 to 2.13, and solution gas-oil ratios ranged from about 200 cu ft/bbl to 2,250 cu ft/bbl. Viscosities ranged from 0.38 to 0.95 cp at the respective bubble points of the fluids and from 0.7 to 20 cp at atmospheric pressure. Water floods were conducted at various pressure levels from run to run. The recovery as a function of flooding pressure was found to be different for each fluid, with optimum gas saturations ranging from 7 up to 35 per cent. The data indicate that substantially higher recoveries may be obtained if water floods are conducted at an optimum pressure and that this optimum pressure is a function of fluid properties. The same core was used for all tests, and the reproduction of saturations for various runs indicates that wettability in the predominantly water-wet core did not change. INTRODUCTION A paper was presented by Bass and Crawford&apos; which described an experimental study of the effects of flooding pressure and rate on oil recovery by water flooding. This work was conducted using high-pressure models operated in a manner similar to that of an actual reservoir, with gas saturations being obtained by a solution-gas-drive mechanism. They found that the greatest oil recovery was obtained for the system studied by flooding in the presence of a 5 to 7 per cent gas saturation. Another experimental study simulating field conditions was presented by Richardson and Perkins.&apos; They used an unconsolidated sand pack containing kerosene-natural gas fluid and interstitial water. They flooded at various pressures and flooding rates. For their system it was found that the recovery was independent of the pressure level at which the water flood was performed. Kyte, et al," found that oil recovery by water flooding was increased as the free gas saturation at waterflood initiation was increased. However, after the initial gas saturation was increased above 15 per cent, the increase in oil recovery tended to level off. All of their runs were made at the same pressure using a light oil saturated with helium. The desired gas saturation was obtained by injecting helium into the core. Dyes&apos; made calculations which showed that an optimum gas saturation of 12 to 14 per cent may result in an increase in oil recovery of 7 to 12 per cent over that obtained by flooding at the bubble-point pressure. Others have also found that the presence of a free gas saturation may increase the waterflood oil recovery. In each case shrinkage was small and changes in fluid properties with respect to pressure were small. A careful review of the literature reveals that at the present time there is a wide difference of opinion on the factors affecting waterflood recoveries. This diversity of opinion is probably due to the fact that very little research has been done which has taken into account the many variables existing in an actual field being water flooded. Since the literature contains little information on high-pressure waterflooding studies using various types of reservoir fluids, it was believed appropriate that such a study should be made. EQUIPMENT AND PROCEDURE All tests were made using the same consolidated sandstone core. Torpedo sandstone was used to turn a cylindrical core 13.5-in. long and with a 2.92-in. average diameter. The core had a porosity of 28 per cent and a permeability to brine of 146 md. This brine was made up by adding 20,000-ppm sodium chloride and 30,000-ppm sodium nitrite to distilled water. This was used as connate water and flooding water. No fresh water was ever brought in contact with the core, as tests showed the sandstone contained argillaceous material which swelled in the presence of fresh water and plugged the stone. The core was sealed in a section of 6-in. N-80 tubing with Woods metal filling the annulus. The core was mounted horizontally; an injection well was placed in the center of one end and a production well in the center of the other. Pressure control was maintained by placing a back-pressure regulator (upstream control) on the producing well. The "live" oil was stored in a separate bottle and water was injected into this bottle to displace the oil for saturating the core using a two-cylinder standard-proportioning pump. This same pump was used for water flooding the core at a constant rate. This system was enclosed in water jackets and the temperature was automatically main-

By S. T. Wlodek

The surface and subscale oxidation reactions were followed by means of continuous weight-gain and metallographic techniques over the range 1600" to 2200°F (871° to 1204 °C) for up to 400 hr. Full identification of all scale and subscale reaction products was obtained by electron and X-ray diffraction. At or below 1800°F (982°C) a linear rate of reaction (QL = 46.0 kcal per mole) governed the oxidation process, extending for up to 100 hr at 1600°F (871 "C). During linear oxidation the surface scale consisted of an amorphous SiO2 film overgrown with Cr 2O 3 and NiCr204. This initial linear process was followed, and above 1800°F completely replaced, by two successive parabolic rate laws (Qp = 60 and 57 kcal per mole). This parabolic reaction involved the formation of a complex scale consisting of Cr2 O3 and smaller amounts of NiCr2O4. Parabolic oxidation appeared to coincide with the disruplion of the silica film present during linear oxidation and was followed by subscale (internal) oxidation of crystobalite and NiCr2O4. The balance between the subscale and surface oxidation reactions controls the oxidation of this commercial alloy. The amorphous silica film appears to result in the linear rate and diffusion through Cr2O3 is the more likely rate-limiting step during parabolic oxidation. THE oxidation of a multicomponent composition is a complex phenomenon not presently amenable to a rigorous classical interpretation. Nevertheless, even a qualitative understanding of the scaling and subscale reactions that occur in a commercial composition can illuminate the reactions that limit its high-temperature stability in an oxidizing environment. This study of the oxidation of Hastelloy Alloy X presents the first of a series of studies with the above approach in mind. Hastelloy X exhibits one of the best combinations of strength and oxidation resistance available in a wrought, solution-strengthened, nickel-base alloy. Although during long time exposure some precipitation of M6C and M23C8 carbides as well as a complex Laves phase occurs, the amounts are probably small enough to have no appreciable effect on the chemistry of the matrix. Radavich has identified the oxidation products on Hastelloy X oxidized for 5 min to 10 hr at 1115°F as NiO and the NiCr2O4 spinel. Oxidation for 5 to 15 min at 1500°F produced a scale of spinel, NiO, and a rhombohedra1 phase, probably Cr2Os. Sannier et 2. have reported continuous weight-gain data for Hastelloy X at 1650" and 2010°F and internal-oxidation measurements after 150 hr at 2010°F. In addition, much of the data on binary Ni-Cr alloys recently reviewed by Kubaschewski and okins&apos; and Ignatov and Shamgunova4 as well as studies of binary Ni-Mo alloys5 are also pertinent to the oxidation of this composition. EXPERIMENTAL Continuous weight-gain measurements and metallographic measurements of subscale reactions were the main experimental techniques used in this study. X-ray and electron diffraction backed up by a limited amount of electron-microprobe analysis served to characterize the nature of the scale- and subscale-reaction products. Two heats of commercial sheet of the composition given in Table I and identified as A and B were used in the bulk of this study. Internal-oxidation measurements were made on a third heat of material in the form of a 0.5-in.-diam bar. In order to assure homogeneity, all heats were reannealed 4 hr at 2175°F prior to sample preparation. weight-Gain Measurement. All specimens (1.5 by 0.4 by 0.03 in.) were abraded through 600 paper, electropolished, and lightly etched in an alcohol-10 pct HCl solution. An electrolyte of 150 cu cm H,O, 500 cu cm HsPO4 (85 pct conc), and 3 g CrO3 at a current density of 0.9 amp per sq cm or a solution of 10 pct HaW4 in alcohol used at 4 v and 0.3 amp per sq cm was used for electropolishing. The resultant surface exhibited a finish of 3 ± 1 p rms. Continuous weight-gain tests were made at 1600°, 1700°, 1800°, 1900°, 2000", and 2200°F on auer&apos; type balances capable of recording a total weight change of 110 mg with an accuracy of k0.1 mg. All tests were made in air dried to a dew point of -70°F and metered into the 2-in.-diam reaction

Jan 1, 1964

By Shiou-Chuan Sun

THE presence of an excessive amount of tough froth in the flotation of minerals, particularly coals, may create trouble in dewatering, filtering, and handling. Froth is also a nuisance in many chemical industries.&apos; This paper presents a study on the destruction of extremely tough froths with intense high-frequency sound. The data indicate that sound waves can be employed for continuous atandsoundwavescan instantaneous defrothing. A powerful high-frequency siren was used in obtaining the data. Also tested was an ultrasonorator of the crystal type with a frequency range of 400, 700, 1000, and 1500 kc per sec and a maximum power output from its amplifier of 198 w. The results, not presented, indicate that as now designed this machine is not suitable for defrothing. Although the sound generators of the magnetostriction type2,3 and of the electromagnetic type&apos;.&apos; were not available, it is beelectromagneticlieved they are capable of producing the required sound intensity for defrothing. The use of ultrasonics for defrothing was suggested by Ross and McBain1 in 1944. Ramsey8 reported in 1948 that E. H. Rose mentioned a supersonic device that broke down flotation froth but with low capacity. The writer has not been able to find any published literature containing practical experiments. Theoretical Considerations The mechanism of defrothing by sound is attributed to the periodically collapsing force of the propagated sound waves and the induced resonant vibration of the bubbles. The collapse of froth is further facilitated by the sonic wind and the heat of the siren. Sound waves can exert a radiation pressure&apos;," against any obstacle upon which they impinge. When a froth surface is subjected to the periodic puncturing of sound waves, the bubbles are broken. According to Rayleigh9 and Bergmann,12 the radiation pressure of sound, P, in dynes per sq cm is given as: P = 1/2 (r+1)i/v where r is the ratio of the specific heats of the medium through which sound is traveling and is equal to 1 on the basis of Boyle&apos;s law; i is the sound intensity in ergs per sec per sq cm, and v is the sound velocity in cm per sec. In this case, the accuracy of the formula is only approximate, because a perfect reflection can hardly result from a column of froth. In addition to the radiation pressure, the propagated sound waves cause the bubbles of the froth to have a resonant vibration.&apos;" he vibratory motion of the bubbles causes collision and coalescence, thereby weakening if not breaking the bubble walls. Sonic wind and heat were also generated." The sonic wind can exert pressure on the froth surface, and the heat can evaporate the moisture content of the bubble walls as well as expand the enclosed air. Apparatus The defrothing apparatus, shown in Figs. 1 and 2, consists of a powerful high-frequency siren, a glass or stainless steel beaker of 2-liter capacity with 12.4 cm diam and 17.1 cm height, and a metal reflector. The beaker was placed 2 in. above the top point of the siren. The metal reflector was adjusted to reflect and focus the generated sound waves into the central part of the beaker. Fig. 2 shows the crystal probe microphone used to measure the acoustic intensity and the mandler bacteriological filter employed to introduce compressed air into the beaker for frothing. The apparatus was enclosed in a soundproof cabinet equipped with a glass window. The siren, shown in Fig. 3, consists of a rotor that interrupts the flow of air through the orifices in a stator. The rotor, a 6-in. diam disk with 100 equally spaced slots, is driven by a 2/3 hp, Dumore W2 motor at 133 rps. The frequency of the siren can be varied from 3 to 34 kc. The maximum chamber pressure is about 2 atm, yielding acoustic outputs of approximately 2 kw at an efficiency of about 20 pct. The siren itself is relatively small and can be operated in any orientation. A detailed description of the siren has been given by Allen and Rudnick.11 Collapse of Froth To study the sequence of the collapse of froth, the glass beaker was partially filled with 920 cc water, 100 g of —150 mesh bituminous coal, 0.3 cc petroleum light oil, 0.2 cc pine oil and 1.54 cc Pyrene foam compound. This mineral pulp was agitated for 5 min and then aerated through a mandler filter until the empty space of the beaker, approximately 9 cm high, was filled completely with min-

Jan 1, 1952

By Craig S. Hartley

Expressions are obtained for the energy changes associated with the reaction of (a& (111) slip dislocations on intersecting (110)planes in anisotropic bcc metals. An energy criterion for assessing the likelihood of dissociation of the products of such reactions is also presented. It is found that the "burrier reactions" which form a(100) dislocations at the intersection of two active {110) slip planes are more energetically favorable in metals which exhibit a high value of Zener&apos;s anisotropy factor, A, than those which have a low value. The results are presented in a form which permits the stacking fault energy to be obtained from a measurement of the separation between par-tials in a dissociated configuration. However, until accurate calculations or measurements of the stacking fault energies involved are available, it is not possible to assess the physical importance of dissociated dislocations. In a recent paper,&apos; the energy changes associated with several types of reactions between two slip dislocations, (a/2)(111){110), in bcc structures were calculated.* Isotropic elasticity and the approxima- tion v = -3- were employed. The purpose of this work is to present calculations of the energy changes for many of the same reactions using anisotropic elasticity. The problem of dissociation of a(100) and a(110) dislocations is also considered, and maximum fault energies for which dissociation will be energetically favorable are calculated for several bcc metals. Two general types of reactions are considered; those for which the reactant (a/2)(111) dislocations have long-range attractive forces and those for which the reverse is true. An example of the former is: (a/2)[lll] + (a/2)[lll]-a[l00] while the latter are typified by: (a/2)[lll] + (a/2)[111] -a[011] Only reactants lying in different slip planes are considered; therefore, the products must lie along (111) or (100) directions, which are the intersection of two {llO} planes. It will be assumed that the reactants and products are infinitely long parallel dislocations, since in this case the energy change associated with the reactions is a maximum.&apos; THEORY The self-energy per unit length of a straight mixed dislocation in an anisotropic medium can be written? where b is the Burgers vector, K is an appropriate combination of the single-crystal elastic constants, and R and ro are, respectively, outer and inner cut-off radii of the elastic solution. The energy given by Eq. [I] does not account for any variation of the core energy with orientation. This could be manifested by an orientation dependence of the core radius or, equivalently, the Peierls width, of the dislocation. However, the energy contribution due to this source is expected to be small, and current models of the dislocation core are not sufficiently accurate to justify such a refinement. It has already been shown that for the isotropic case the energy contributions due to nonzero tractions across the cores of the reactants and products exactly cancel one another in the reaction.&apos; Accordingly, it will be assumed that this contribution to the total energy change in the anisotropic case is small. In the subsequent discussion it is also assumed that the core radii of the reactant and product dislocation are the same and that, where stacking faults are formed, the faulted region is bounded by the centers of the partials. Consequently only changes in elastic energy due to the reactions will be considered. When the dislocation is parallel to either the (111) or the (100) directions, K may be written:375 K = (Ke sin2 a + Ks cos2 a) [2] where K, and Ks are the combination of elastic constants corresponding to an edge and screw dislocation lying along the same direction as the mixed dislocation, and a is the angle between the direction tangent to the dislocation line and the Burgers vector. Eq. [2] should not be confused with the isotropic approximation to the variation in energy with line Orientation.6 It should be noted that the essentially isotropic expression for K is a result of the characteristic symmetry of the (111) and (100) directions and is not, in general, valid for other dislocation directions in anisotropic cubic metals. The energy* change for a reaction in which the re- actant and product dislocations are parallel perfect dislocations can be written: where Ep and E, refer to the self-energies of the products and reactants, respectively. For dislocations parallel to (100) and (111) directions, Eq. [3] becomes:

Jan 1, 1969

By F. E. Werner, B. L. Averbach, Morris Cohen

THAT bainite formed near the M, temperature bears a striking r esemblance to martensite tempered at the same temperature has been shown by the electron microscope.&apos; By means of electron diffraction,&apos; it has been established that carbide and cementite are present in bainite formed at 500°F (260°C); these carbides are also found in martensite tempered at 500°F (260°C).&apos; The investigation reported here is concerned with an X-ray study of the matrix phases in lower bainite and tempered martensite. These phases have turned out to be dissimilar in structure; the matrix of bainite is body-centered-cubic while that of tempered martensite is body-centered-tetragonal. A vacuum-melted Fe-C alloy containing 1.43 pct C was studied. Specimens of 16 in. diam were sealed in evacuated silica tubing and austenitized at 2300°F (1260°C) for 24 hr. One specimen was quenched into a salt bath at 410°+7 °F (210°+4°C), held for 16 hr, and cooled to room temperature. The structure consisted of about 90 to 95 pct bainite, the re: mainder being martensite and retained austenite. A second specimen was quenched from the austen-itizing temperature into iced brine and then into liquid nitrogen. It consisted of about 90 pct martensite and 10 pct retained austenite. The latter specimen was tempered for 10 hr at 410°+2°F (210°+1°C). The specimens were then fractured along prior austenite grain boundaries (grain size about 2 mm diam) by light tapping with a hammer. Single aus-tenite grains, mostly transformed, were etched to about 0.5 mm diam and mounted in a Unicam single crystal goniometer, which allowed both rotation and oscillation of the sample. Lattice parameters were measured by the technique of Kurdjumov and Lyssak. This method takes advantage of the fact that martensite and lower bainite are related to austenite by the Kurdjumov-sachs orientation relationships Thus, the (002) and the (200) (020) reflections can be recorded separately, permitting the c and a parameters to be determined without interference from overlapping reflections. According to these findings, the matrix phase in bainite is body-centered-cubic and, within experimental error, has the same lattice parameter as ferrite (2.866A). On the other hand, martensite, tempered as above, retains some tetragonality, with a c/a ratio of 1.005t0.002. Most workers in the past have assumed that bainite is generated from austenite as a supersaturated phase, but the nature of this product has not been established. The question arises as to whether bainite initially has a tetragonal structure and then tempers to cubic, or if it forms directly as a cubic structure. If it forms with a tetragonal lattice, it might well be expected to temper to the cubic phase at about the same rate as tetragonal martensite. The martensitic specimen used here was given approximately the same tempering exposure, 10 hr at 410°F, as suffered by the greater part of the bainite during the isothermal transformation. About 50 pct bainite was formed in 6 hr at 410°F. On tempering at this temperature, martensite reduces its tetragonality within a few minutes to a value corresponding to 0.30 pct C.&apos; Further decomposition proceeds slowly, and after 10 hr the c/a ratio is still appreciable, i.e., 1.005. Thus, even if the bainite were to form as a tetragonal phase with a tetragonality corresponding to only 0.30 pct C, which might be assumed to coexist with e carbide, it would not be expected to become cubic in this time. It seems very likely, therefore, that bainite forms irom austenite as a body-centered-cubic phase and does not pass through a tetragonal transition. The carbon content of the cubic phase has not been determined, but it could easily be as high as 0.1 pct, within the experimental uncertainty of the lattice-parameter measurements. It has been postulated that retained austenite decomposes on tempering into the same product as martensite tempered at the same temperature. There is now considerable doubt on this point. The isothermal transformation product of both primary and retained austenite at the temperature in question here is bainite," and the present findings show that bainite and tempered martensite do not have the same matrix. Acknowledgments The authors would like to acknowledge the financial support of the Instrumentation Laboratory, Massachusetts Institute of Technology, and the United States Air Force.

Jan 1, 1957

By A. B. Cook

Gas cycling is generally considered a much less efficient oil recovery mechanism than water flooding. HOWever, recoveries from some fields have been exceptionally high as a result of gas cycling. Recovery from the Pick-ton field, for example, was calculated to be 73.5 perceni of the stock-tank oil originally in place. In evaluating pressure maintenance projects, determining how much of the recovery is due to displacement by gas and determining how much is due to vaporization of the imrnohile oil in the flow path of the cycled gas is very difficrilt. Even though most of the oil is recovered by displacetr~ent, the success of a project may depend on the amount of oil vaporized. A limited number of experiments have heen performed with a rotating model oil reservoir that simulates gas cycling operations and allows a separation of the oil from, tile free gas flowing into the laboratory wellbore at reservoir conditions, thus revealing which is displaced oil and which is vaporized oil. It Iras been determined that the amount of varporizatio&apos;n is .significant if proper conditions exist These experiments show that oil vaporization depends on pressure, temperature, volatility of the oil and amount of gas cycled. Increases in each of these conditions increase the volume of oil vaporized. Data from six experiments affecting vaporization are presented to illustrate reservoir condition that range from favorable to unfavorable. 111 these eaperitnenis recovery by vaporization ranged from 73.6 to 15.3 percent of /he immobile oil (oil not produced by gas displacerrlt). INTRODUCTION Between 1930 and 1950, gas cycling was a popular. oil recovery practice. especially for the deeper reservoirs. Later, with many case history-type studies published for both gas cycling and waterflooding, it was generally believed that waterflooding was far superior to gas cycling, even when gas cycling was conducted as a primary production procedure by complete pressure maintenance. A good example illustrating the advantage of water-flooding over gas cycling is given in a paper by Matthews&apos; on the South Burbank unit where gas injection was followed by waterflooding. The author concluded in part that "Early application of water injection, without the intervening period of gas injection, would have recovered as much total oil as ultimately will be recovered by waterflooding following the gas injection, and total operating life would have been shortened". This appears to be a logical conclusion. However, it should not be applied to all fields. Pressure maintenance with gas in the Pickton field, as reported by McGraw and Lohec;&apos; will result in a much larger percentage of oil recovery than was obtained in the South Burbank unit. The great success in the Pickton field resulted partly from vaporization of the immobile oil in the flow path of the cycled gas. The amount of vaporization is related to the following conditions: volatility of the oil as reflected by the APT gravity of the stock-tank oil; reservoir temperature; reservoir pressure during gas cycling; and the amount of gas cycled. Therefore, the U. S. Bureau of Mines is investigating these effects on vaporization in a research project using a model oil reservoir. Three different stock-tank oils having 22, 35 and 45" API gravities are being used as base stock to synthesize reservoir oils. Experiments are being performcd to determine vaporization at 100, 175 and 250F and at 1,100, 2,600 and 4,100 psia. This is a progress report showing the results from six experiments. Other Bureau of Mines reports"- concerning vaporization are listed. LABORATORY EQUIPMENT AND PROCEDURES The equipment &apos; consists of an internally chromium-plated steel tube packed with finely sifted Wilcox sand. The tube is approximately 44 in. long and has an ID of 13/4 in. The sand section contains approximately 570 ml of voids, has a porosity of 32 percent, and a permeability to air of 4.3 darcies. A unique feature of the laboratory reservoir (Fig. 1) permits the tube part to rotate at 1 rpm while the outlet and inlet heads are held stationary. The outlet end contains diametrically opposed windows to permit observatlon of the flowing fluids, and two valves, one on the top and the other at the bottom. Oil and free gas. when being produced simultaneously, can be separated by manipulating the two valves to keep a gas-oil interface in view through the windows. Thus, only gas is produced through the top valve and only oil flows through the bottom valve. The laboratory equipment was designed to study vaporization. Therefore, a uniform reservoir was made using dry sifted sand as opposcd to using a consolidated sand core with interstitial water. Furthermore. the reservoir was tilted to minimize fingering of gas. This tilting also in-

By W. B. Gogarty, W. C. Tosch

A new recovery process for producing oil under both secondary and tertiary conditions utilizes the unique properties of micellar solutions (also known as microemulsions, swollen micelles, and soluble oils). These solutions, which displace 100 percent of the oil in the reservoir contacted, can be driven through the reservoir with water and are stable in the presence of reservoir water and rock. Basic components of micellar solutions are surfactant, hydrocarbon and water. They may also contain small amounts of electrolytes and co surfactants such as a1cohol.r. The specific reservoir application dictates the type and concentration of each component. A salient feature of [he process is the capability for mobility control. Micellar solution slug mobility, by way of viscosity control, is made equal to or less than the combined oil and water mobility. Mobility control continues with a mobility buffer that prevents drive water from contacting the micellar solution. Laboratory and field flooding have proven that the process is technically feasible and that surfactant losses by adsorption on porous media are small. Introduction projects are under way to recover the maximum amount of oil under the most favorable economic conditions.&apos; : New techniques are being developed to increase oil recovery,3" Polymer solutions are becoming an important means of controlling mobility in a waterflood. Thermal methods such as in-situ combustion and steam injection are being used in reservoirs containing highly viscous crudes. Surfactant flooding is receiving attention as a method of reducing interfacial tension to increase recovery.*&apos;" Exotic recovery processes have been considered primarily for &apos; perations. Economics are unfavorable in most cases for tertiary recovery. studies at the Denver Research Center of the Marathon oil CO. have led to a new oil recovery method.* Micellar solutions (sometimes called microemulsions, swollen micelles, and soluble oils) are used to recover oil by miscible-type waterflooding. Basically, these solutions contain surfactant, hydrocarbon, and water. The method can be used in either secondary or tertiary operations. First, thc concept of thc process is considered in terms of the requirements for an effective miscible waterflood ing operation. Next, micellar solution properties are described including structure, composition, and phase behavior with reservoir fluids. Fluid characteristics are then considered as related to mobility control, and, finally, laboratory and field results are presented to illustrate the efficiency of the process. Concept of the Process Unit displacement efficiency and conformance determine the effectiveness of any oil recovery mechanism. In theory, a miscible waterflood should be capable of a 100-percent unit displacement efficiency with a correspondingly high conformance. Requirements for the slug of a miscible waterflood include (1) 100-percent displacement of oil in the reservoir contacted, (2) controllable mobility, (3) the capability of being driven through the reservoir with water, (4) a low unit cost to enhance economics, and (5) the ability to remain stable in the presence of reservoir water and rock. Micellar solutions satisfy requirements for the slug of a miscible waterflood process. Our discovery that these solutions acted as though they were miscible by displacing all fluids in the reservoir and by being displaced by water solved the miscibility problem. Adequate mobility control is possible by variations in solution viscosity through compositional changes. Economic requirements are met since micellar solution costs below $6/bbl appear possible, Mi cellar solutions stabilize surfactant in the presence of reservoir rock and water, thus reducing the importance of the problem of surfactant loss by adsorption. Fig. 1 illustrates schematically how these solutions are used. Operations start with injection of a micellar solution slug that serves as the oil displacing agent. Next, a mobility buffer of either a water-external emulsion or water solution containing polymer (thickened water) is injected to protect the slug from water invasion. Finally, drive water (water used in a regular waterflood) is injected to propel the slug and mobility buffer through the reservoir. Reservoir oil and water are displaced ahead of the slug, and a stabilized oil and water bank develops as shown in Fig. 1. Stabilized bank saturations are independent of original oil and water saturations. This means that, for a particular type of reservoir, the displacement mechanism is the same under secondary and tertiary recovery conditions. Oil is produced first in a secondary operation. For tertiary conditions, water is produced first. Movement of the slug through the reservoir is stabilized by the mobility buffer. An unfavorable mobility ratio usually exists at the interface between the buffer and drive

Jan 1, 1969

By J. W. Clark, C. T. Sims

Wrought chromium-base alloys containing yttrium, cubic monocarbides of the Ti(Zr)C type, and similay alloys containing manganese and rhenium have been melted and fabricated. Strength has been studied by hot hardness and elevated-temperature tensile and rupture measurements, low-temperature ductility by tensile testing, and surface stability by oxidation testing. In additiod, studies have been conducted of the carbide stability, and of aging behavior. The carbide dispersion generates effective elevated-temperature strength, which is further enhanced hv strain-induced precipitation. The dispersion exhibits classical dissolution and aging response. The ductile-to-brittle transition temperature of these alloys is above room temperature. The alloys reported show fairly good oxidation resistance, but nitrogen contamination can cause fortnation of a hard Cr2N layer under the oxide scale. Manganese does not appear to be a promising alloying element in chromium. In the years 1945 to 1950, the metal chromium was considered as a possible base for alloy systems due to its considerably higher melting point than superalloys, its low density, its high thermal conductivity, and its apparent capacity for strengthening. However, this interest in chromium was short-lived. It was found difficult to melt and cast, to be exceptionally sensitive to the effect of minor imperfections, to have a lack of ductility at both room and elevated temperatures, and to be subject to a deleterious effect of alloying elements upon the ductile-to-brittle transition temperature.&apos; Since then, chromium, as a practical alloy base, has remained virtually unstudied. Further, purposeful ignoring of chromium has been promoted by statements that its bcc structure would not allow it to be strengthened to useful values, when compared to the "austenitic" alloys.2 Recently, a new look has been taken at chromium-base alloy systems. Study of the literature will show that chromium, providing some of its disadvantages could be eliminated or minimized, actually has a rather attractive potential as an alloy-system base. Analysis of rather scattered data suggests that chromium is quite capable of being strengthened to high levels. Also, significant strengthening of its two sister elements in Group VI-A, molybdenum and tungsten, has been demonstrated in a number of commercial and exploratory alloys. Chromium should be similar. Since chromium does not readily form a volatile oxide like tungsten or molybdenum, it offers a much higher probability of giving birth to alloy systems with useful oxidation resistance. Concerns about possible high elemental vapor pressure have been mitigated by recent data.3 In addition, the physical properties exhibited by chromium are attractive for application as a high-temperature structural material. For instance, its thermal conductivity varies from 49 to 36 Btu-ft/hr-sq ft-°F over its range of usefulness (which is two to four times higher than most superalloys), its density is about 7.2 g per cc (20 pct less than most nickel-base alloys), its coefficient of thermal expansion varies from 4 to 8 x 10-6 per OF, and it has a relatively high modulus of elasticity, approximately 42 x 10&apos; psi.4 Alloying studies on a chromium base in the past have usually encompassed rather sweeping solid-solution alloy additions for strengthening. This is not consistent with contemporary alloying practice in Group VI-A. For instance, molybdenum, also in Group VI-A, is primarily alloyed for strength improvement by use of heat-treatable carbide dispersions.5 Chromium and molybdenum are similar in their chemical activity and other properties. Thus, strengthening of chromium by carbide dispersions was studied. Chromium-base alloys are plagued with room-temperature brittleness, although high-purity unal-loyed chromium can be made ductile.4,8 Use of yttrium as a scavenger has done much to improve ductility and resistance to nitrogen embrittlement in chromium systems,7 so it was utilized in this program. It has also recently been found8 that small rhenium additions (1 to 5 pct) create improvement in the ductility of Type 218 tungsten wire. This is apparently related to the remarkable effect of rhenium additions near its terminal solid solubility in all Group VI-A metals.9&apos;10 Investigation to establish if dilute concentrations of rhenium would also be effective in chromium appeared to be logical for this program. Since rhenium is too expensive to be practical in alloys for application as structural components, ductility improvements through solid-solution alloying were also sought by substitution of manganese for rhenium; manganese, like rhenium, exists in Group VII of the periodic system. The optimum amount of carbide dispersion for chromium-base alloys was obtained by analogy with molybdenum. Strengthening in molybdenum is achieved by use of Ti-Zr carbide dispersions. A

Jan 1, 1964

By M. E. Shank, J. Wulff

At the present time, the designer of dies for metal powder pressing is handicapped by relative ignorance of stress distribution and frictional effects at the interior surface of the die. Unckell was the first to develop a method for the study of wall friction. He used three Brinell balls on which the die rested during pressing. The total frictional wall force was determined by the size of impression these balls left on a soft metal plate. Since the method does not give radial pressures, or distribution of such pressures, coefficients of friction could not be determined. Although Unckel measured density distribution, he was not able to determine radial or shear stresses. Shaler2 has proposed theoretical expressions for the stress and density distribution within cylindrical compacts during pressing, in accordance with the experimental results of Kamm, Steinberg, and Wulff.3 By application of Siebel&apos;s method,4 Kamm et a13 plotted stress trajectories for two compacts. From the stress trajectories they calculated coefficients of friction from point to point along the die wall. As pointed out by Shaler in the discussion of Ref. 3, these values are based on progressive point-to-point calculations on finite size grid squares across the compact. In the region of the die wall such calculated values may therefore have considerable cumulative error. The purpose of the present paper is to develop an experimental method by which the nonhydrostatic pressures and shears acting on the interior wall of a cylindrical die can be measured. Such measurements can then he correlated with existing data to aid in the explanation of the pressing process. The method used is based on the elastic: properties of the thick-walled tube used as the die. The principle of super-position of force systems on an elastic body is assumed to hold. Electric strain gauges were mounted in adjacent positions on the exterior die wall in order to get an exact measurement of the variation of tangential strain over the length of the die during pressing. While in this paper, measurements in terms of only tangential strains are considered, it is well to note that similar calculations may be set up for axial strains. The latter are not preferred, since they tend to be smaller than the tangential strains and therefore permit less sensitive measurements. Discussion in this work is restricted to compacts pressed from both ends, since the elastic deformation of the die is then more amenable to analysis. Before choosing the electric strain gauge method, a more direct line of attack was considered and discarded. The discarded idea was the insertion of a pressure gauge through a hole in the die wall.* The gauge would have been in the form of a small piston. If pressure were exerted against such a gauge, it would move outward along a radius of the die. One disadvantage of the scheme is its inability to measure shears along the die wall. Another more serious disadvantage is the disturbance caused by the device itself. It would serve to change the forces it was designed to measure. No matter how small the movement of the gauge, when pressure is applied a discontinuity would exist in the wall surface at that point. Due to the stress concentration caused by the hole, abnormal deflections of the die wall would occur around the gauge. During pressing, powder would be forced into the resulting depression. The depression would then become larger with increasing compacting pressure. Powder, not being a fluid, is capable of supporting shear. The ease with which it would flow into the die wall depression to further move the piston is an indication, not of the radial pressure at that point, but of the state of shear retarding the movement. Thus the "pressure" gauge is really a criterion of flowability, and of the capability of the powder to support shear. For these reasons, it was decided that the electric strain method, herein employed, was more reliable, if more indirect. The gauges and lead wires, mounted on the external die wall do not in any way affect the behavior of the metal powder or the die during pressing. Theory of the Method THE EFFECT OF RADIAL PRESSURE ON THE DIE WALL Effect of a Single Small Band of Hydrostatic Pressure Consider a die which is a thick-walled cylinder of outer radius R. and inner radius Ri. If over a small finite length L there is a normal pressure P, a tangential strain distribution at the outer wall results. This is shown schematically in Fig 1. The exact shape of the curve may he predicted by an extension of the theory of a semi-infinite beam on an elastic foundation.6 This

Jan 1, 1950

By A. Ghosh, T. B. King

The kinetics of the discharge reaction: 20&apos;- (in silicate melt) = O,(g) + 4e- at a platinum anode in lithium silicate melts have been studied al 1350°C by galvanostatic methods. Plots of the sleady-state overpotential, q, as a function of the logarithm of the current density, i, showed injlections and were linear only at high current densities. The value of the overpotential was influenced by bubbling gas through the electrolyte. The ocer potential was also studied as a function of time. The rise and decay of overpotential were very slow processes. At low current densities transport is the likely rate-controlling process but at high current densities passivation of the electrode, Presumably by an oxide film on the surface, seems to be a contributory functor. IT is well-established that molten silicates behave as electrolytes&apos;5 and, except in a few cases,6 conduction is entirely ionic. Moreover, it is supposed that a possible, and perhaps predominant, mechanism for phase boundary reactions between metals and slags is similar to that in corrosion whereby anodic and cathodic processes occur at unrelated sites, the metal serving to conduct electrons.1&apos;8 Thus electrochemical studies of some slag-metal reactions would seem to be a useful way to diagnose the rate-controlling steps in the overall reaction. The electrochemical method is, in principle, a better diagnostic tool than the direct chemical method for the following reasons: 1) The partial electrochemical reactions, which are simpler than the overall reaction, may be studied individually. 2) The rate of reaction can be controlled at will and independently of the concentrations of reactants. 3) Fast reactions can be studied by relaxation methods.&apos; Esin and his coworkers5&apos;10"12 have pioneered such studies in silicates and have deviloped some ingenious techniques. Not all of their findings, however. can be accepted without a good deal of further work. In this investigation, the kinetics of the oxygen discharge reaction: 202- (in silicate melt) = Oz(g) + 4e- [I] at a platinum electrode were studied by both steady-state and transient galvanostatic techniques. Interest in this reaction was first developed as a result of the findings of Fulton and chipman13 that the reduction of silica, in a silicate slag, by carbon, dissolved in liquid iron, is a very slow reaction. Subsequent work, for example, by Rawling and ~lliott,&apos;~ has demonstrated that the reaction under these conditions must be slow, because the rate is limited by diffusion of oxygen in the iron to the metal-crucible phase boundary at which a CO bubble may be nucleated. Further work by Tarassof,&apos;~ in which the reduction of silica by aluminum dissolved in copper was studied, has shown that under these conditions, where carbon monoxide evolution is not involved, control of the reaction rate resides in diffusion of silica in the slag phase. However, there is no practical way of inducing sufficient convection in the system to make it clear that the phase boundary reaction is indeed fast. The overall reaction of silica reduction involves the discharge of silicon ions at cathodic sites and oxygen ions at anodic sites. In the examples cited, the discharged ions are dissolved in a liquid metal. In the present study of oxygen ion discharge, gaseous oxygen may be evolved at high current densities or oxygen may simply dissolve, possibly as oxygen molecules, in the silicate at very low current densities. The discharge of an oxygen ion at an anode must, in silicates less basic than the orthosilicate composition, be preceded by a reaction in the vicinity of the electrode, such as: which makes oxygen ions available. Platinum was chosen as the working electrode since it is comparatively inert to oxygen and is, therefore, expected to come rapidly into equilibrium with the electrolyte and with gaseous oxygen. Minenko, Petrov, and Ivanova16 have measured the electromotive force at a platinum electrode in molten silicates as a function of the partial pressure of oxygen in the atmosphere, the concentration of oxide ions in the melt, and the temperature. They found platinum to behave as a reversible oxygen electrode. At two different oxygen pressures, Po2 (I) and Pq (11). the electromotive force is given by: where F is the Faraday constant, equal to 23,060 cal per v equivalent, indicating that the electrode reaction is as written in Eq. [I.]. Platinum has been similarly used in molten silicates by other inve~ti~ators. "&apos;~~ In this investigation platinum was used only as an anode, since a current deposits other elements on its surface and changes its characteristics.

Jan 1, 1970

By T. E. Watkins, M. D. Nelson

INTRODUCTION Deeper drilling for oil is becoming more and more the rule rather than the exception. With deeper drilling come additional problems, perhaps the greatest being those brought on by the higher temperatures encountered down the hole. particularly in the Gulf Coast region of Texas and Louisiana. Temperature gradients of the order of 1.8° to 2.0°F/100 ft are not unusual, and a gradient of 2.3"F.&apos;100 ft is found in some areas of Texas. With a mean surface temperature of 74oF, the following temperatures could be expected for a geothermal gradient of 2.0°F; 100 ft: at 10,000 it. 271°F. 12,000 ft, 314°F: 14,000 ft, 354,oF; and 16.000 ft. 394°F. Severe gelation of lime-base drilling fluid in wells that have high bottom hole temperatures has become perhaps the most serious difficulty enconntered in drilling under such conditions. Lime-base drilling fluids have been very succesefully and widely used in the drilling of wells in the Gulf Coast region because of their inherent stability toward contaminants. their ability to suppress the swelling dispersion of bentonitic shales, and their ease of maintainance. The gradual recognition: during the past few years, that these muds were. in themselve. the cause of many difficulties experienced in drilling has led to wide-pread efforts by the drilling industry. to determine the reasons for the failure of these mud systems and to develop mud systems capable of performing satisfactorily under high-temperature conditios. MANIFESTATIONS OF HIGH-TEMPERATURE GELATION it is generally possible to recognize the symptons of high-temperature gelation early enough that advance predictions can be made of serious difficulties. in mud control, and the useful life of the drilling fluids can be extended by proper treatment. Following i.; a list of the manifestations of high-temperature gelation: (1) The drill string &apos;takes weight&apos; while going in the hole after a trip. In early stages of high-temperature gelation it is possible to notice a slight reduction in drill string weight as the drill pipe is lowred near the bottom of the hole. (2) Excessive pump pressure is required to .tart the circulation of drilling fluid at or near the bottom of the hole when going hack to bottom after a trip. As the severity of the gelation increases it may be necessary to break circulation a number of times when going in the hole. (3) The drilling fluid from the bottom of the hole is thick and often granular or lumpy when pumped up after making a round trip. In a severely gelled drilling fluid system such a condition may be irreversible; that is, it cannot be stirred or chemically treated to produce a satisfactory drilling fluid. (4) Completion tool.. such as logging tools or perforating guns will not sink to the bottom of the hole. On some occasions completion tools will become stuck and require a fishing job to retrieve them if the wire line attached to them is broken. It is often difficult to determine whether the condition of the drilling fluid is responsible for sticking the tool or whether the wire line becomes key seated in a crooked hole and causes the allow difficulty. When there are 110 other symptoms of high-temperature gelation. then the difficulty may usually be attributed to the latter cause. (5) In extreme cases of high-temperature gelation it is necessary to "wash" and "ream" when going back to bottom after coming out of the hole. (6) In many -instance. it has been found to be extremely difficult and expensive to 1111 production packers 2nd tubing in moderately deep oil wells which had been drilled with a lime-base drilling fluid. In such instances-the original mud had apparently "set" to a consistency approaching that of a weak cement. CAUSES OF HIGH-TEMPERATURE GELATION Extensive test; have indicated that a lime-base mud does not develop a highly gelled condition at temperatures below 250°F. whereas above that temperature such condition often develops rapidly. (Fig. 1) concurrently. the following changes are evident ill the mud: (1) The alkalinity of the mud decreases to a very low value. with both caustic soda and lime being consumed. (2) The quartz content of the mud decreases sharply. (3) The bentonitic content of the mud decreases or di-appears, with concurrent decrease or loss of base exchange capacity of mud solids. (4) New compounds formed in the mud have been found to be cal-cium silicate, calcium aluminum silicate, and calcium sodium aluminum silicate. (5) The mud loses the ability to form a filter cake of low permeability. The above characteristics have been discussed, in part. by other authors

Jan 1, 1953

By V. Steve Reed, Ed L. Reed

The in situ leach mining process generates a waste stream that is high in sulfates, total dissolved solids, and radium 226. During the mining phase, the volume of the waste stream is relatively low and consists primarily of the bleed stream. During the restoration phase, larger volumes of waste water are generated. These waste streams require environrnentally sound disposal. The low net evaporation rate in the Coastal Bend area precludes pond evaporation as a feasible disposal alternative. Reverse osmosis is a practical method of reducing the volume of the waste water handled, but the concentrated waste stream from the reverse osmosis unit must be disposed properly. Deep well injection into highly saline reservoirs is considered a sound method of disposing of the liquid waste generated by in situ mining in the Gulf Coast uranium district. Thirteen injection wells have been permitted to serve the disposal needs of the leach mining industry in Texas. Of these 13, 11 have actually been drilled. Seven applications are pending. The injection zones for the permitted wells range from depths of 3050 to 6200 feet. Pressure limitations imposed on these wells range from 500 psi to 1350 psi. The following criteria are used to determine the desirability of a disposal well site: 1. A minimal number of nearby, improperly plugged borings which penetrate the disposal zone; 2. Minimal crustal disturbance; 3. Sufficient salinity of the water contained in the disposal zone; 4. Protection of oil and gas producing zones; and 5. Sand of sufficient permeability and areal extent to handle the desired volume without fracturing the reservoir. 1. Improperly plugged borings: During the early part of the century, oil wells, gas wells and test holes were drilled using cable tool equipment, often with a minimum amount of surface casing. Production casing, when it was set, was often partly removed when the holes were abandoned. Thus, wells drilled prior to 1940 frequently have less than 100 feet of surface casing and either no production casing or the upper part of the production casing removed. Additionally, these holes are often plugged only with mud. The close proximity of these holes to an injection well location are a concern in that they can provide an avenue for injection-depth fluids to migrate up the bore hole and jeopardize shallower fresh water reservoirs. Usually, where there are more than 6 or 8 poorly plugged borings in a 2 1/2 mile radius of the well site, it is preferable to examine deeper zones for disposal well potential. The deeper zones are especially attractive where the borings are not in a cluster, which renders monitoring more difficult. Often, even the deeper disposal zones are penetrated by a few improperly plugged borings. When this condition arises, the potential for leakage through the borings can be addressed in the following ways. a. Demonstration that the static head in the boring is higher than the anticipated increase in bottom hole pressure generated at the boring by the disposal well. A 100 psi differential between these two pressures is recommended. The calculated increased pressure at a boring caused by injection should be refined using annual bottom hole pressure measurements in the disposal well. Figure 1 illustrates an injection pressure map which can be overlain on the oil well map to determine the anticipated increase in pressure expected at each oil, gas or abandoned hole. b. Shallow ground water monitoring. A shallow monitor well is drilled next to the boring and both pressure and quality measurements are made periodically in the shallow well. c. Disposal zone monitoring. Recently there has been a tendency for regulators to require disposal depth monitor wells instead of shallow well monitoring. We consider disposal depth monitoring to be a less effective method of monitoring because it provides only indirect evidence of potential problems. Assumptions have to be made for the unplugged borings, such as mud weight, that are not addressed by the disposal zone monitoring program. There is little improvement with this system to that discussed in "a" above. A shallow zone monitoring program, however, yields direct evidence of a developing problem with an unplugged boring. Leakage by the boring will be detected quickly by an abnormal increase in pressure in the shallow well. Quality monitoring will detect upward migration of poor quality fluids. The pressure data provide an early warning of impending leakage; the quality monitoring will detect actual fluid migration.

Jan 1, 1980

By S. J. McCarroll

Fig. 1—The belt shown at right carries filter cake to mixing station over calciner. Crude ore conveyors appear in right background. THE Three Kids mine, some six miles east of Henderson, Nev., is in a typical southwest desert area, with high dry summer heat and cool to cold winter seasons. The manganese deposit was located during World War I.&apos; During this period 15,000 to 20,000 tons of ore assaying up to 41 pct manganese were shipped. Interest in the deposit was not revived until the middle thirties, when experiments on the ore were initiated. Test work indicated possible recovery of only 70 pct by flotation, but in 1941 additional work was done at the Boulder City pilot plant of the U. S. Bureau of Mines and also by M. A. Hanna Co. As a result, the Manganese Ore Co. was formed and a plant utilizing the SO2 process was constructed. Numerous operation difficulties ensued, and the plant. was closed when the manganese situation in the country eased. In 1949 Hewitt S. West initiated negotiations to acquire the plant. In 1951 Manganese, Inc., was formed and contract entered into with the General Services Administration to supply 27 million units of metallurgical grade manganese in the form of nodules to the national stockpile. A second contract was made to upgrade 285,000 tons of stockpile ore. Test work was undertaken by the Southwestern Engineering Co. and likewise by the Boulder City pilot plant at the U. S. Bureau of Mines. Results obtained indicated the commercial feasibility of the flotation process. Construction of the plant, which is shown in Figs. 1 and 2, was started in June 1951, and operations on a break-in basis began in September 1952. Apart from the usual starting difficulties two major disasters caused serious setbacks, one a kiln failure in February 1953, and the other a fire that destroyed the flotation building in June of the same year. The nodulizing section of the plant resumed operation in November, and the flotation section in January 1954. The ore minerals are chiefly wad,* with minor amounts of psilomelane, and occur in sedimentary beds of volcanic tuff. The ore is overlain with beds of gypsum which outcrop or may be covered with surface gravel. Intermediate beds of red and white tuff occur frequently with lenses of red and green jasper and stringers of gypsum and calcite. Small amounts of iron are present; lead content averages about 1.0 pct and minute amounts of copper and zinc are found. Barite, celestite, and bentonite are present. Since these are made up of minute asicular crystals, moisture content is very high, averaging about 18 pct. Ore reserves have been estimated at 3 million tons averaging 18 pct Mn2 and up to 5 million tons after grade is dropped to 10 pct Mn. A good part of the orebody was stripped of overburden by the previous operating company . Approximately 50 pct of the ore, representing more than 60 pct of the manganese, can be mined by open-cut methods. A system for underground min- . ing has not yet been decided on. Open-cut mining with benches of 20 ft has proved satisfactory. Although the ore is soft and appears dry and dusty it has a certain resilience, probably due to the porosity and moisture which makes drilling and fragmentation difficult. Wagon drills have been abandoned in favor of the Joy 225-A rotary drill which will put down a 43/4 -in. hole at the rate of 2 ft per min. Holes are spaced in a pattern with 8 to 9-ft centers. Forty percent powder has been used, but better breaking to 2-ft size is obtained with low velocity bag powder of 30 pct strength. Loading is done with one 21/2-yd shovel, and cleanup follows with one D-7 bulldozer. The ore is hauled with Euclid trucks about 1000 ft from the pit to a blending pile, where the daily mine production is spread in layers by bulldozing until approximately one month&apos;s mill feed is accumulated. A new pile is then started and mill feed is drawn from the first pile by one 13/4-yd shovel and Euclid trucks, with a haul of approximately 500 ft. Mining is performed by an independent contractor with engineering and supervision by the company staff. Early test work indicated that the manganese could be floated with soap, a wetting agent, and fuel oil to give a recovery of better than 75 pct with a grade of 43 pct Mn. The concentrate when nodulized with coke would upgrade to 46 pct Mn or over, and the lead volatilized to 0.6 pct residual.

Jan 1, 1955

By Lawrence H. Van Vlack, Otta K. Riegger

Periclase-type oxides were examined microscopically after being exposed to siliceous liquids. The rate of grain growth was found to be inversely proportional to the grain diameter. Grain growth proceeds more rapidly at higher temperatures, but is retarded by increasing liquid contents. aMag-nesiowiistites with higher MgO contents grow less rapidly than those with higher FeO contents. The growth rate is reduced by the presence of a second solid phase. The silica-containing liquid penetrates as a film between the individual magnesiowus tite grains. This is independent of time, temperature, amount of liquid, or the MgO/ Fe0 ratio. When present, olivine and spinel-type phases can provide a solid-to-solid &apos;&apos;bridge" between magnesioustite grains. THIS paper presents the results of a study of the microstructures of periclase type oxides in the presence of a silicate liquid. The purpose was to learn more about the effect of service factors such as 1) time, 2) temperature, and 3) liquid content upon A) grain growth, and B) liquid location among the solid grains. This study was prompted by the fact that periclase refractories are known to have very little solid-to-solid contact when the phases which are present are limited to periclase and liquid. Such a micro-structure gains industrial significance because it permits fracture during service when stresses are applied at high temperatures. The details of ceramic microstructures have not received extensive attention. This is in contrast to the extensive attention given to a) the phase relationships pertaining to refractory compositions, and b) the details of the microstructures of comparable metallic materials. A brief review will be made of the pertinent phase relationships and microstructural considerations in general, as well as of refractory compositions. a) Phase Relationships. This investigation was limited to those compositions in which (Mg, Fe)O was the solid phase. MgO and FeO form a complete series of solid solutions. Pure MgO has the name of periclase. The related FeO structure is called wustite. Both have the NaC1-type structure: however, wustite possesses a cation deficiency so that the true composition is Fe<10 even in the presence of metallic iron. The phase relationships involving solid (Mg, Fe)O and a silicate liquid are shown in Fig. 1. In this case. the liquid is saturated with (Mg, Fe)o. There-fore its SiOz content is below that encountered in orthosilicate liquids. As a consequence the liquid phase specie:; are primarily the following ions: and 0-&apos; plus occasional Fe+ ions. Two features are of importance: a) the liquid contains relatively small species and b) the liquid contains large quantities of the same species as the solid. viz., Fig. 2 shows the system, FeO-SiOz, which will be used in some of the discussions that follow. This diagram is the right side, vertical section of Fig. 1. Here, as pre\iously, the composition at the FeO end of the diagram is nonstoichiometric, varying from Feo.950 when the liquid oxide is in contact with the solid iron, to about Fe 0, when the solid oxide is in equilibrium with an atmosphere of equal proportions of CO and C02 at the solidus temperature. The Fe/O ratio will be maintained in wustite in the presence of SiO,. However, the FeM/Fe++ ratio in the liquid will be lower because of the effect OIF the SiO, on the activity of the FeO. With the addition of MgO to wustite, the over-all composition (IvZg, Fe)@, has a value of x lying between 0.9 and 1.0 when the COz/CO ratio is 1.0&apos;. b) Microstructures. In general, published attention to refractory microstructures has been directed toward the phase analyses that accompany compositional variations. This is illustrated by Harvey6 in his work on silica brick and by wells7 in his work on periclase brick. In each case, a series of altered zones is encountered which provides a sequence of phase associations. If due consideration is given to reaction kinetics, such an examination reveals phases that are compatible with equilibrium studies. Admittedly, however, it is often necessary to determine more complicated polycomponent systems to account for all the phases present.8 Relatively little attention has been given to microstructural geometry in ceramic materials. Certainly less attention has been given to this aspect of ceramic microstructures than to the size, shape, and distribution of the constituent phases in metals. Burke has pointed out that the grain size of oxides follows the same growth rules as for metals, viz.,

Jan 1, 1962