Search Documents

By R. H. Raring, F. E. Martin

A high speed gas quenching dilatometer useful in studying phase transformations in low alloy steels is described. Changes in specimen length are measured by means of an electrical micrometer tube. The maximum instantaneous cooling rate is approximately 5000°F per sec; the maximum average cooling rate from 1650' to 950°F is approximately 2900°F per sec. Data for steels of 0.11 pct and 0.67 pct C content are included. KNOWLEDGE of phase transformations as they occur in an alloy under conditions of continuous cooling affords valuable insight into its properties. The transformation characteristics during continuous cooling of steels of high hardenability, which decompose entirely to martensite at fairly slow cool-

Jan 1, 1957

By M. S. Smani, P. Blazy, J. M. Cases

Investigations on the beneficiation of sedimentary Moroccan phosphate ores are reported in four separate parts. The first discusses electrochemical phenomena at the solid-aqueous solution interface for minerals of the apatite group. In the second part, the electrochemical properties of the calcareous gangue associated with the phosphate ore are investigated. The flotation of calcite and phosphate oölites, using both anionic and cationic collectors, is discussed in the third part. It is shown that selective recovery could be obtained for precise concentrations of collectors and pH ranges. Results of work aimed at optimizing depression of phosphate oölites during inverse flotation of the carbonate gangue with sodium oleate are reported in Part 4.

Jan 1, 1976

By Carl Swartz

IN the course of a recent investigation1 to develop a more satisfactory white-metal. bearing alloy, a number of alloy systems were studied. The cadmium-nickel system showed characteristics desirable in a bearing alloy, and the system was therefore studied in some detail. EXPERIMENTAL METHOD Alloying cadmium with nickel offered no serious difficulty. After melting the cadmium, clean sheet nickel was added while the temperature was maintained at 500° to 600° C: A few crystals of ammonium chloride2 served to clean the nickel when the latter did not "wet" readily. In preparing the individual alloys for thermal and microscopic investigation, a "hardener" containing about 10 per cent nickel was used.

Jan 1, 1933

By James E. Little

Of the several extensive deposits of brown iron-ore in Cuba, including those of Mayari and Moa, that of Mayari was the first to be systematically explored, and was selected as the scene of the first operations in the development of this class of ore. Construction-work, begun in the spring of 1907, involved the building of 16 miles of stantlard-gauge railroad and two large double-track inelines, the installation of milling-machinery, a nodulizing-plant, power-plant, shops and shipping-facilities, and the dredging of an extensive basin for deep-draught vessels. Unusual weather-conditions delayed the completion of portions of this work so that the entire plant was not in operation until December, 1900.

Jan 1, 1912

By L. O. Millard

Stockpiles in the minerals industries serve a wide variety of purposes. Usually they are for surge between stages of processing, for a dependable plant feed in anticipation of delays, or to provide for single or double shift mining in conjunction with round-the-clock milling. Often surge piles are necessary to insure a uniform feed to continuous process equipment where the material is received as an intermittent or variable flow. Extensive use is made of stockpiling for the primary purpose of blending. The significant fact about all stockpiling of bulk materials is that storing and reclaiming are two separate and distinct operations and they may present widely different problems in the selection of methods and tools.

Jan 10, 1959

By Stephen Krickovic

THERE is today no proven continuous mining machine that can be used under all the varying conditions found in most bituminous coal mines. During the last five years, however, both the machines and methods of using them have been improved. Coal operators can now exercise a fair degree of selectivity, except those who must struggle with thick and hard partings and operate in seams less than 40 in. thick. It must be realized, then, that more than one type of continuous miner may be used advantageously in the same coal mine. Although this appears to violate the standardization generally sought after, the gain can more than offset added maintenance and, to a lesser degree, operating problems.

Jan 12, 1957

By Henry Tobelmann

THE location, mode of occurrence of ore, and preliminary tests leading to the development of the present leaching process and the building of the present plant on the property of the New Cornelia Copper Co. at Ajo, have been described in previous papers presented to the Institute and in the, discussions thereof.1 It is sufficient to say that these tests began early in July, 1912, and were continued to the close of January, 1916; that small tests at Douglas were followed by the construction of a 1-ton plant at the mine; and this was followed by the operation of a 40-ton plant for 14 mo. In all about 15,000 T. of ore were treated experiment-.

Jan 2, 1919

By R. J. Wasilewski

Elastic constants have been determined on single crystals of maximum-melting-temperature NiAl compound (50.6 at. pct Al) at 25°C. Temperature variations of Young&apos;s modulus in the three principal crystal directions between -150°C and +800°C exhibit no anomalies similar to those repovted for the isomor-phous compounds brass and AuCd. It is concluded that the ordered CsCl structure is stable throughout the temperuture range investigated NICKEL aluminide is a congruently melting inter-metallic compound of CsCl (B2) structure, stable over a rather wide composition range (40 to 54 pct Al). It exhibits a lattice-parameter maximum of a. = 2.883A at 50.6 pct 1,&apos; at which composition the excess aluminum atoms are randomly distributed on the vacant nickel sites;&apos; at higher nickel compositions the excess nickel atoms replace substitutionally in aluminum sites, while at higher aluminum compositions vacancies are present in the nickel sublattice.&apos; For compositions close to NiAl no significant variation in order has been observed at temperatures up to 900°C.3 At high nickel (65 pct) compositions the existence of a martensitic structural transformation has been reported.4 NiAl is isomorphous with the ordered B-brass structure and with the high-temperature structures of AuCd and TiNi alloys close to equiatomic composition. All of these exhibit anomalous elastic behavior as the temperature of their diffusionless transformation to lower-symmetry structures is approached.&apos;-&apos; This behavior has been explained as indicative of low stability of the CsCl structure in these compounds.8 The present paper reports the experimental determination of the elastic compliances of single-crystal NiA1. EXPERIMENTAL METHOD 1) Specimens. Polycrystalline NiAl was prepared by induction melting of carbonyl nickel and high-purity aluminum and casting into solid copper molds in gettered helium atmosphere. The castings thus produced were canned in mild steel and impact-extruded at 1100 to 1200°C to 38-in.diam rod. After the steel sheath was pickled off, the rods were converted to single crystals by repeated high-frequency satisfactorily produced the desired orientations in the form of crystals 38 in. diam by 8-12 in. long. These were subsequently center less-ground to 7 to 8 mm diam by 12 cm long, electropolished (in 80 pct methanol, 20 pct HzSO, at 25°C and 15 v), and their axial orientation determined to 1 deg by the Laue back-reflection technique. The rod dimensions were determined to 0.01 mm and the densities determined from geometrical dimensions and the weight of every specimen. The average density was 5.905 0.005 g per cu cm. 2) Elastic Data Measurements. Room-temperature elastic compliances Sii were obtained by determining to 0.1 cps the fundameGtal resonance frequencies in the longitudinal and torsional vibration modes. At these frequencies the following relations hold: E =p(2lfif =pvl= Sn- 2r[Sll- S12- -£„«] g -P(2iftf = pvf = (s<« - fr(sii - sw - ±s»y where vl and vt are the velocities, and f andft the resonance frequencies of the longitudinal and torsional waves, respectively; 1 is the length of the specimen, and r is the direction cosine factor—{cos2 a • cos2 0 + cos2 a . cos2 y + cos2 0 ¦ cos2 Y}. Given three crystals of orientations (loo), (110), and (111)) i.e., F = 0, 14, and 13, we can obtain six equations for the three unknown compliances, thus providing adequate internal check. Mode mixing can readily occur in the relatively long wavelength resonance utilized here. This could cause a significant error in the torsion wave velocities calculated from the apparent resonance frequenies. Therefore, throughout the present work, a method, described previosl,&apos; was used to ensure that no significant mode mixing took place at resonance. The apparatus used was originally designed for the Elastomat, Magnaflux Corp.. Chicago, 111. determination of flexural vibration resonance frequencies. At room temperature torsional vibration mode could be established by suitable positioning of the thin wires from the input and pick-up transducers, and longitudinal mode by cementing thin magnetic foils at the ends of the specimen, and inducing the vibration by means of electromagnetic transducers. Only flex-ural resonance frequencies could, however, be obtained reproducibly over the temperature range of -150 to 800°C. These data are considerably less accurate, the error in the modulus being estimated at -0.6 pct from the specimen dimensions and the room-temperature Poisson&apos;s ratio. The reproducibility of the resonance frequencies in repeat runs was, however, to better than l in 15,000, hence the relative elastic modulus error less than 0.02 pct. Therefore, the temperature variation of the principal Young&apos;s moduli obtained by flexural resonance, relative to room-temperature values, is believed to be satisfactorily accurate.

Jan 1, 1967

By A. S. Odeh

The mechanism of two-phase flow in porous media has been a subject of wide controversy. One of the properties essential for understanding the dynamic behavior of two-phase flow is relalive permeability. Relative permeability to a certain phase is defined as the ratio of the effective permeability of that phase to its permeability when it is the only fluid present and powing. In this research, a theoretical analysis was made to determine the effect of viscosity ratio between the non-wetting and the wetting phase on relative permeability. Experimental work was conducted to test the validity of the derived equations. The experiment was conducted on four natural cores. Four oils were used as the non-wetting phases with a viscosity range of 0.42 to 71.30 cp and two wetting phases with a viscosity range of 0.86 to 0.96 cp. Oil and bring were made to flow simultaneously at various ratios, and relative permeability curves were constructed. A total of eight relative pertileability cycles representing eight viscosity ratios were run oil each sample. It was found that relative permeability to the non-~cletting phase varies with viscosity ratio. The relative effect of this variation on relative permeability values was a function of the sample&apos;s single-phase permeability, decreasing with its increase. It was concluded that, for .samples of single-phase permeability over I darcy. the effect of viscosity ratio could be disregarded, and relative permeability would be, in effect, a function of satrtration only. INTRODUCTION Two-phase as well as multiphase flow occurs in many fields of science. This type of flow is of particular interest in petroleum production. The knowledge of relative permeability, which describes the dynamic behavior of two-phasc as well as multiphase flow, is essential for solution of problems arising in that field. Thc relative permeability ot a porous medium to a given phase in multiphase flow. is generally considered to be only a function of the saturation of that phase, independent of the properties of fluids involved and ranging in value from zero to unity. Work by Leverett&apos; and Leverett and Lewis&apos; apparently supports this concept. In his experiments Leverett used a clean, packed unconsolidated sand of high permeability (3.2 to 6.2 darcies) with two phases (water and oil) flowing and a viscosity ratio range of 0.057 to 90.0. His results showed that the wide range of viscosity had practically no effect on relative permeability-saturation relationship. Recently accumulated evidence from work performed by several laboratories and a paper by Nowak and Krueger,2 in which relative permeability to oil of a few core samples in the presence of interstitial water was considerably greater than single-phase permeability to water, cast some doubt on the conclusions reached by Leverett&apos; and subscribed to by a large number of individuals in the oil industry. One explanation advanced to explain this behavior states that it is caused by the variable extent of hydra-tion of clay minerals present in the sand. The greater the water saturation, the greater will be the area of contact between water and clay minerals; therefore, the greater will be the extent of swelling with corresponding reduction in permeability. Yuster4 presents another explanation for the recently accumulated evidence. Utilizing Poiseuille&apos;s law, he analyzed concentric flow in a single capillary where the non-wetting phase flows in a cylindrical portion of the capillary and concentric with it. The wetting phase flows in the annulus between the non-wetting phase and the capillary wall. The equations obtained indicate that relative permeability to the non-wetting phase is a function of saturation and viscosity ratio. Although Yuster&apos;s equations show that fractional rel-ative permeability to oil could be greater than unity, as was indicated by the data of Nowak and Krueger,1 they failed to present an explanation to the experimental data of early investigators such as Leverett.1 Due to the importance of relative permeability in understanding the flow behavior of petroleum reservoir fluids, this work—theoretical as well as experimental —was undertaken to determine whether relative permeability is a function of saturation only as was concluded by Leverett1 or a function of saturation and viscosity ratio as was theorized by Yuster.4 THEORETICAL ANALYSIS An equation will be derived for the rate of oil flow through a porous medium that is initially filled with water. Based on this equation, an analytic expression for relative permeability will be developed. The porous medium will be assumed to consist of .straight circular capillaries of different radii. It will also be assumed that there are no interconnections among the capillaries and no mass transfer across the oil-water interface. Consider a porous sample initially saturated with a wetting phase (water). As a non-wetting phase (oil) is

By R. A. Davidson, L. Silverman

RESIDENTS of many urban centers are becoming increasingly aware of the obscuring effect of fume and smoke discharge from power, metallurgical, chemical, and other industries; and they, as well as the legislatures of these affected cities, are agitating for cleaner air. Management&apos;s most pressing problem is to find an economical way to reduce process effluents in response to the growing pressure from population and legislative demands. The removal must be done, if possible, without handicap to the current operation, since the costs of relocating are often excessive or prohibitive. In fume recovery or disposal, an important item to consider is whether or not the material being discharged has any value. If it has commercial value, the cost of its recovery may offset or aid amortization. For this reason, in making a study of the specific problem in hand, a major factor was the nature of the material emanating from the stack: in particular, its particle size, size range, and its chemical and physical composition, as well as its potential value and utility when recovered (in either a wet or dry state). Should the product have no commercial value, it must be disposed of at minimum cost in a way to prevent recontamination. Initial studies were therefore made to determine stack concentrations and volumes of material evolved from the operations. The next phase of the study concerned the physical and chemical nature of the collected fume. The third portion of this paper describes the wet and dry collector studies undertaken to recover the fume. Cleaning Requirements for Ferroalloy Furnace Operation The basic need for any effluent collection equipment is the highest possible efficiency and the lowest tolerable resistance when the power consumption involved is considered. Since the electric furnace effluent is largely composed of fume of small size (less than 0.5u), it has high light obscuring properties, and even low concentrations will cause some loss of visibility and be evident to nearby residents. The permissible limit for fly ash emission in many cities is based on a weight value (viz, approximately 0.4 grains per cu ft), but the smoke density values are dependent upon a shade of color. In the case of the Los Angeles County code, emission is restricted to pounds per pound of material processed per hour basis (but not exceeding 40 lb per hr for any one given plant operation). If an average particle size of the fume from ferro-silicon alloy electric furnaces is assumed to be 0.4u (as shown later, this is the approximate mean size) and an average loading of 1 grain per cu ft (stp), each cubic foot of stack gas will contain approximately 75x10 10 particles (based on assumed, and confirmed, spherical shape and a standard deviation of unity). When it is realized that the air in metropolitan areas, which are also general industrial areas, contains approximately 5x108 particles, the tremendous light scattering effect of this concentration becomes apparent. Consequently, nearly 100 pct collection would be necessary to equal the average concentration. Fortunately, however, discharge from a high point above ground (50 to 100 ft) will result in at least a thousandfold dilution, or the stack concentration reaching the ground in the foregoing case might result in a ground concentration of &apos; particles. If the concentration at the source could be reduced by a factor of 100 (99 pct efficiency of collection), then a concentration of 75x10" particles would be diluted to 7.5x10&apos; which would be very satisfactory. An efficiency of 90 pct (factor of 10 decontamination) at the source would result in a discharge of 75x109 articles which upon dilution yields 75x10 which is still 15 times the general air value. Another approach to this consideration is to use the value of concentration of 0.005 grains per cu ft for the value of a visible effluent as cited by Kayse.1 To attain this value with an average loading of 1 grain per cu ft would require an efficiency of 99.5 pct. Since the foregoing value is not based on any reported size of fume particles, it is felt that the numbers&apos; approach given previously is more reliable. These calculations serve to indicate the desirability of thorough cleaning, preferably at the source, and with efficiencies well above 90 pct, preferably above 95 pct (dilution 1:20). One of the most important items in any control program is to reduce the concentrations as close to their sources as possible. The use of better furnace design, deeper coverage over the electrodes, and the prevention of blows or breaks in the surface all help to reduce dissemination; consequently, all of these improvements should be made, if possible, to cut down the effluent load. In addition, in order to minimize the volume of contaminated air that has to be cleaned, the furnace should be enclosed as much as possible. Test Arrangements Before fundamental studies with collectors were made, a furnace stack selected for the test program was sampled to determine the gas temperatures and

Jan 1, 1956

By K. P. Gupta, A. K. Pal, L. Chandrasekaran, U. P. Singh

The Mn-Sn binary system, investigated at the high-manganese end and between 500° and 1000° C, shows four phases at temperatures below 727"C, namely the u Mn, the p Mn, the Mn3 Sn, and the Mn, Sn phases, while at higher temperatures only the last three phases remain stable. The solubility of tin in a Mn is very small and the maximum solubility of tin in P Mn phase appears to be abollt 10 at. pct Sn. The solubility range of the Mn,Sn and the Mn, Sn phases is 23.0 to 26.0 and 37.0 to 40.0 at. pct Sn, respectively. The lattice parameter of the 13 Mn phase increases with increasing tin content. The Mn, Sn thase is hexagonal with and appears to be basically of the Ni3 Sn type structure except that for quite a few X-ray diffraction lines the calculated and observed relative intensities do not agree well. The Mn-Sn binary system has been studied by several investigators.&apos;&apos;~ Their results indicate that three intermediate phases, namely, the Mn3Sn, Mn,Sn, and MnSn, phases, exist at high concentrations of tin. However, so far the proper phase equilibrium has not been established at the manganese-rich end and very little data is available for the composition range between pure manganese and Mn3Sn. Moreover, earlier investigators differ in their opinion about the exact composition range at which the Mn3Sn and Mn,Sn phases appear and some doubt has been cast by some investigators3 regarding the structure of Mn3Sn phase which has been reported to be isotypic with Ni3Sn (Mgscd type) structure. In this investigation an attempt has been made to establish the proper phase equilibrium between pure manganese and Mn + 50 at. pct Sn composition in the temperature range of 500" to 1000°C. PROCEDURE The raw materials used were from three different sources. For exploratory work five alloys containing 5, 10, 15, 20, and 25 at. pct Sn were prepared using 99.9 pct pure manganese and tin supplied by E. Merck & Co., Germany. The rest of the alloys and one more 25 at. pct Sn alloy were prepared using 99.9 pct Mn supplied by Gallard Schlesinger Chemical Mfg. Corp., U.S.A., and 99.999 pct Sn supplied by Semi Elements Inc., U.S.A. Weighed amounts of manganese and tin were melted in recrystallized alumina crucibles in an inert gas (argon) high-frequency induction melting furnace. By careful control of temperature and time of melting the losses were reduced to below 0.2 pct in all cases. Since the losses were very small, no attempt was made here to analyze the samples chemically. Alloys were wrapped in molybdenum foil and sealed in small evacuated fused silica capsules. The alloys were annealed at different temperatures, controlled within + l°C, for sufficiently long periods to attain proper phase equilibrium, and subsequently quenched in cold tap water. The annealing periods used at different temperatures were 15 days at 500°C, 7 days at 600°C, 5 days at 700" and 750°C, 3 days at 800°, 850°, and 900°C, 2 days at 968"C, and two alloys annealed at 968°C were reannealed for 10 hr at 1000° C. From each annealed specimen, a part was utilized for metallographic study while another piece was used for X-ray diffraction study. 1.0 pct HNO, solution and oxalic acid solutions of concentrations 0.05 to 1.0 pct were used for etching Mn-Sn alloys above and below the MnsSn composition, respectively. Since all alloys were brittle, X-ray specimens were prepared using the as-crushed -325 mesh alloy powders. Only one 25 at. pct Sn alloy powder was reannealed in an evacuated silica capsule at 800°C for 5 min and water-quenched. X-ray diffraction patterns . for the Mn3Sn phase with the as-crushed and the reannealed powders did not show appreciable change. Un-filtered iron radiation at 25 kv, 15 ma was used with either Norelco 114.6-mm-diam Debye Scherrer Camera (for phase identification) or Norelco 12-cm-diam symmetrical focusing camera (for lattice parameter determination of the 6 Mn phase). The estimated accuracy of lattice parameter determination for the focusing camera was * 0.001A. RESULTS AND DISCUSSIONS The results of metallographic and X-ray diffraction study made with different alloys are shown in Fig. 1. The variations in lattice parameter with composition for the p Mn and the Mn3Sn phases are given in Tables I and 11, respectively, and the lattice parameter as a function of composition for the p Mn phase is shown in Fig. 2. The lattice parameter of the p Mn phase increases with increasing tin content while for the Mn,Sn phase the data obtained from two two-phase alloys and one single-phase alloy indicate increase in a,, and decrease in c, parameters with increasing tin content. The results, Fig. 1, indicate that the solubility of

Jan 1, 1969

By S. R. Bardsley, S. T. Algermissen

Induction, nuclear and sonic well-logging methods were employed in a Green River formation oil-shale analysis and evaluation study conducted in northeastern Uintah County, Utah. The physical and chemical properties of an oil-shale section, previously cored and assayed by conventional methods, were used to evaluate the response of the various logs. The logging program was designed to measure variable properties of oil shale which relate to oil-yield potential in order that a relationship between assay-oil yield and log-determined properties could be identified, thereby permitting a direct determination of yield potential from logs alone. All of the major oil-shale zones and section ttlarkers are recognizable on the logs used in the study. The relationship between the response of the Density and Sonic logs and the assay-oil yield in gallons per ton was suficient to perinit the derivation of equations expressing the relationships. These equations can be used to determine the potential oil yield in gallons per ton of an oil-shale zone or section. The Neutron log response distinguishes the rich oil-shale intervals from the lean intervals, but does not appear to permit establishment of a quantitative relationship. The Garnma-Ray and Induction logs indicate only a qualitative relationship to oil-shale yields. Logging oil shale by Gamma-Gamrna Density and/ or Sonic logging methods will permit a fast, economical and accurate means of evaluating the potential yield of oil-shale deposits. Considerable sums of money have been allocated to Green River formation oil-shale evaluation during the past years by both private enterprise and Federal and state governments. The result of the work accomplished to date is commendable, but the task of fully exploring and evaluating one of the world&apos;s greatest reserves of potential energy for economical exploitation is enormous and much information is still needed. Present methods of evaluation consist primarily of sampling and assaying the oil shale for potential yield. The data received are excellent, but are both costly (inasmuch as it requires funds specifically allocated for the evaluation purpose) and time-consuming (as each representative sample must be assayed in the laboratory by specially trained personnel). At the present time, the Green River formation oil-shale province is one of the most active areas in the Rocky Mountain region for exploratory oil and gas drilling. This activity could play a twofold role and serve the function of spear-heading many oil-shale evaluation programs, as it could provide both cutting samples for assaying, geophysical data and modern well logs. This paper deals with a recent study conducted on Green River formation oil shale in which the application of induction, sonic and nuclear well logs to oil-shale evaluation was tested and proved. THEORY AND DEFINITIONS Oil shale may be described as siliceous marlstone, rich in solid organic matter called "kerogen". Kerogen is only slightly soluble in organic solvents, but it will decompose and yield oil vapors and gas when heated to destructive distillation temperatures at about 800 F. The mineral constituents of Green River formation oil shale are found in essentially uniform proportions with one another. The dominant types are, in the general order of abundance, dolomite, calcite, feldspars and quartz. Kerogen is also essentially uniform in its composition, and is about 80 per cent carbon and 10 per cent hydrogen by weight. The ratio of kerogen to the mineral constituents determines the richness or potential yield of the oil shale. This ratio also determines the total physical and chemical properties of the rock and provides a basis by which oil-shale richness may be determined from well logs. Because Green River formation oil shale was deposited in a lacustrine environment, the lithology is laterally consistent over wide areas. The vertical section consists predominantly of thin bands of alternating rich and lean oil shale which have been described by Bradley&apos; as varves. The varves are occasionally broken by thin bands of volcanic ash and tuff and by zones of lean vugular oil shale. The effect that the thin non-oil shale beds and vugular zones would have on the well log analysis was a prime concern in the study. To cope with this possible problem, the logs used in the study were of two types: (1) logs believed to respond favorably to those properties of oil shale which are dependent upon the kero-gen-to-mineral ratio, these logs being the Gamma-Gamma Density log, the Sonic log and the Neutron log; and (2) logs that would respond to litho-logies and properties within the oil-shale section not associated with yield that might result in anomalous responses by the yield-measuring logs. Logs of the second type were the Gamma-Kay log, the Induction log and the Caliper log. THE PRESENT STUDY The present logging study was con-

By H. E. Mauck

The many factors that control bumping must be carefully studied for each coal seam where bumps occur, and specifications known to exclude bumping should be incorporated in the mining plans. This calls for complete knowledge of the seam&apos;s characteristics and its adjacent strata, and in many instances these characteristics are not revealed until the seam is actually mined. Pressure and shock bumps, the two general types, occur jointly and separately. In this discussion no differentiation will be made. Whether pressure or shock, they are treated as bumps, and both must be eliminated. Bumps in mines have occurred in several places throughout the coal fields of the world. A study of many of these occurrences indicates that geologic characteristics, development planning, and mining procedure have contributed. But more specifically, there are conditions usually associated with bumps: thickness of cover, strong strata directly on or above the seam, a tough floor or bottom not subject to heaving, mountainous terrain, stressed and steeply pitching beds, and the proximity of faults and other geologic structures. Mine planning should incorporate these known factors (not necessarily in order of importance): 1) Main panel entries should be limited to those absolutely necessary to ventilate and serve the mine. This reduces the span over which stresses may be set up that will later throw excessive pressures on barrier and chain pillars when they are being removed. 2) Barrier pillars should be as wide as practicable so that they will be strong enough to carry the loads thrown on them when final mining is being carried out. 3) Pillars should never be fully recovered on both sides of a main entry development if the barrier and chain pillars are to be removed later. The excessive pressures placed on the main chain and pillar barriers by arching of the gob areas can result in bumping when these barriers are being removed. 4) Full seam extraction is better accomplished by driving to the mine boundary and then retreat-drawing all pillars. If there are natural boundaries in the mine—such as faults, want areas, and valleys —retreat should be started there. 5) Pillars should be uniform in size and shape. The entire development of the mine should call for uniform blocks with entries driven parallel and perpendicular. Only angle break-throughs should be driven when necessary for haulage, etc. 6) For better distribution of rock stresses and reduction of carrying loads per unit area, both chain and barrier pillars should be developed with the maximum dimensions. 7) Pillars should be open-ended when recovered. If they are oblong, the short side should be mined first. Both sides of a block should not be mined simultaneously, but under no circumstance should the lifts be cut together. 8) Pillar sprags should not be left in mining. If they are not recoverable, they should be rendered incapable of carrying loads. 9) Pillar lines should be as short as practicable. (Three or four blocks are adequate). Experience has shown that rooms should be driven up and retreated immediately. The longer a room stands, the more unfavorable the mining conditions. This contributes to bumping. 10) Pillars should not be split in abutment zones (high stress areas lying close to mined out areas) and if slabbing is necessary, it should be open-ended. 11) Pillars should be recovered in a straight line. Irregular pillar lines will allow excessive pressures thrown on the jutting points. Experience has shown that the lead end of the pillar line can be slightly in advance. 12) Pillar lines should be extracted as rapidly as possible. This appears to lessen pressures on the line and render abutment zones less hazardous. 13) Extraction planning should call for large, continuous robbed out areas. Robbing out an area too narrow to get a major fall of the strata above the seam tends to throw excessive pressures on a pillar line. 14) Timbering in pillar areas should be adequate but not excessive. Too heavy timbering or cribbing is likely to retard roof falls and throw excessive weight on the pillar line. 15) Experience has shown that when pillar lines have retreated 800 to 1000 ft from the solid, bumps can occur. Because this distance may vary in different seams, impact stresses should be studied for each individual condition. In any event, extra precautions should be taken against bumps in this area. This list of controlling factors may or may not be complete. It probably is not, but it covers most of the problem&apos;s significant aspects. The question is whether or not bumping can be eliminated. The answer is that bumping can be minimized and possibly eliminated if these and other established factors are thoughtfully considered and incorporated in the mining and extraction plans. If a mine has already been developed or the pattern set so that little change can be made, then it will be necessary to adjust to the most nearly practicable system that can incorporate the known factors.

Jan 1, 1959

By Robert J. De Angelis, James H. Barker

An improved method for solving dynawzical dislocation problems using a digital computer is described in this paper. Interactions between two distinct types of dislocations were studied: attractive screw dislocations; and Lomer lock forming dislocations. One dislocation is positioned in the lattice and is initially at rest, while the other dislocation is moved through the lattice on an intersecting slip plane at a constant velocity in the range 0.5 to 0.999C. (C is the transverse velocity of sound.) The results obtained from these computations indicate that screw dislocations account for a small fraction of the total strain over a wide portion of the range of velocities studied. They further indicate that mixed dislocations mainly repel other dislocations in the neighborhood of the active glide plane. From this a possible explanation for cell formation is put forth. The density of Lomer locks expected to exist after a strain of 0.2 was found to be 1.4 x 106 cm-2 which is in good agreement with indirect experimental estimates. IN the past, predictions of favorable or nonfavorable dislocation reactions were based on the associated changes in elastic strain energy. Such considerations take no account of the probability of the two dislocations coming into contact to react. Venables1 was the first to approach these probabilities by considering the interactions between two moving screw dislocations on perpendicular glide planes. Because of the restrictive types of dislocations and glide plane geometry employed, his results have limited application to metallic crystals. The work to be presented here develops a general approach to solving dynamical dislocation problems; either dislocation-dislocation interactions, presented here in detail, or dislocation interactions with any other suitably defined stress field. Two types of dislocation-dislocation interactions common to face centered cubic (fee) materials are considered: those between pure screw dislocations of opposite sign on intersecting slip planes and those between mixed dislocations on intersecting slip planes, that can react to form a perfect dislocation. This latter reaction, referred to as the Lomer reaction, produces a locked product dislocation that finds it energitical favorable to disassociate into two Shockley partials and a stair-rod dislocation. This partial configuration known as a Lomer-Cottrell (L-C) lock plays a major role in work hardening of fee crystals. seeger2 names the L-C lock as the prime contributor to Stage II hardening while Kuhlmann-wilsdorf3 and Meakin and Wils- dorf4 also state that it is a significant contributor to work hardening. However, with a few notable exceptions,5-7 direct observations of the Lomer lock and the L-C lock by electron transmission microscopy are scanty, and even these are subject to other interpretations.5,6 In a study of partial dislocations present in austenitic stainless steel, whelan8 did not observe any L-C locks at the head of pile-up groups. This result contradicted existing work hardening theories and led him to postulate an alternate theory based on the stress required to break away dislocations intersecting a pile-up group, from their stacking fault nodes. Due to the importance of the Lomer reaction in producing L-C locks which are an essential feature in current work hardening theories and because there exist no data giving direct quantitative values for the density of locks, and because there has even been some doubt expressed as to whether this important reaction occurs at all, a study of the dynamic behavior of the mixed dislocations which form the Lomer lock was undertaken. Due to their ability to cross-slip with relative ease, screw dislocations play an important role in the deformation of fee crystals. For this reason, the second type of reaction considered here is between screw dislocations of opposite sign. In addition, computations in volving screw dislocation interactions are relatively simple, thus providing a convenient check on the cornputational scheme employed. DEFINITION OF PROBLEM The force exerted on a dislocation due to a generalized stress field is given by the Peach and Koehler9 equation: Here t2 and b2 are respectively the tangent and Burgers vectors of the dislocation, and T1 is the stress dyadic defining the local stress field. The stress field may be externally applied or generated internally by the presence of a lattice defect, such as a second dislocation, as is the case in this work. Frank10 has shown that an equivalent momentum, P, of a screw dislocation can be defined by: Here, EST is the total energy of a screw dislocation and ESo is its rest energy. The left side of Eq. [2] is the time derivative of momentum and the right side is the position derivative of the energy due to the dynamical nature of the dislocation. The total energy of a dislocation is the sum of the potential and kinetic energies. Weertman11 has developed the expressions which were used here; these give the potential and kinetic energies of uniformly moving edge and screw dislocations in an isotropic medium.

Jan 1, 1969

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

By N. M. Levine

WHETHER the United States and its allies can W meet the challenge of a war brought by the Communists will depend largely on who wins the battle of steel production. At the present stage of the world situation, the United States and the other members of the Western family of nations have the lead on iron curtain countries. But we have no sure way of knowing what is happening at Magnetogorsk and other Russian iron and steel producing centers. We must also face the possibility that we may have to meet the challenge alone. The fortunes of war and world politics can strip us of friends and co-fighters quickly. The destruction of Hiroshima and Nagasaki are indicative of what the world can expect if war-madness ever grasps the earth again. Our domestic supply of high grade open-pit and underground iron ore is dwindling because of the drain of three wars and higher than ever civilian consumption. The production of iron ore and its eventual use in blast furnaces are the critical problems of an armed democracy today. The world crisis has led to efforts towards beneficiation for increasing ore supplies. The huge reserves represented by the magnetic taconites at the eastern end of the Mesabi, once in production, should provide us with a substantial portion of our native ore for many years. The estimated 10 to 20 million tons of concentrates annually can be increased in an emergency. If we had a certainty of peace for the next 50 to 100 years, the situation would be a stable, hopeful one, aided by importations of high grade ore from sources such as Canada and Venezuela. The hard truth is that we have little surety of peace tomorrow morning. Let us assume &apos;the U. S. could build sufficient processing plants for increasing production of magnetic taconites under the pressure of national emergency. We must also recognize the power of atomic warfare to contaminate an area as large as the Eastern Mesabi. Thus, it becomes imperative to seek some means of protecting our ability to produce the steel we may one day need to survive. The nonmagnetic taconites, completely dwarfing the magnetic taconites areawise as well as tonnage-wise, might provide us with this insurance. Present indications are that they will be considerably more expensive to treat, but in a desperate situation we might be very grateful for ores yielding 40 to 50 pct Fe recoveries at grades of 53 to 58 pct Fe carrying low phosphorus. The University of Wisconsin, because of the difficult iron ore situation in the state, has been working on the nonmagnetic taconite problem for the past three years in the hope of making a contribution toward its eventual solution. In Wisconsin, the Western Gogebic Range has been the state&apos;s most effective iron producing area. Today however, only two mines are in operation, both underground and approaching depths of more than 3000 ft. The range, however, does have a large supply of nonmagnetic taconites and presents a promising field for study. While the Gogebic offers one large source of nonmagnetic taconites, Michigan and Minnesota have even greater supplies of such material. Alabama, the northeastern states and the West all have low grade iron ore sources which might be utilized under extreme conditions. The Gogebic Range located in northeastern Wisconsin and northwestern Michigan has a total length of about 70 miles, about 45 of which are in Wisconsin. The iron formation averages 500 to 600 ft in width, dips 70&apos; to the north and strikes at approximately N 63° E. The formation is sedimentary and consists of six distinct members characterized by alternating divisions of ferruginous chert and ferruginous slate. The footwall is generally quartzitic and the hanging wall of a sideritic slatey character. The iron minerals are mainly hematites with some magnetites, goethites, limonites and small amounts of siderite. In the area studied, very small amounts of iron silicates were observed. The magnetites occurred mostly in the Anvil-Pabst and Pence members, mixed with hematites and representing roughly about 10 to 20 pct of the total iron in the formation, thereby characterizing it as nonmagnetic. The gangue is of various forms of silica such as chert, opal and flint. Complete liberation of iron and gangue minerals is rare. There is always some iron present in the chert ranging from jasper-like solutions to fairly coarse iron oxide specks. Likewise, one always finds finely dispersed silica within the iron minerals. In late 1943 the Bureau of Mines carried out a trenching and sampling program in the two mile stretch between Iron Belt and Pence in Iron County, Wis. Preliminary work was based on samples from one of the four trenches cut by the Bureau of Mines. More detailed work following the preliminary analysis was then undertaken on samples composited from all the trenches, thereby giving a wider and more representative coverage of the area. A study of the

Jan 1, 1952

By J. L. Huitt, B. B. McGlothlin, D. K. Lowe

Published correlations for the principal aspects of hydraulic fracturing were combined into a digital computer program to facilitate the study of interrelated variables. The computer program includes individual relationships for fracture width during pumping, fracture area generated, propping agent embedment, flow capacities of propped fractures and transport of propping agents in horizontal fractures. The effects of more than 20 treatment and formation parameters on the predicted results of hydraulic fracturing treatments were studied. The effects of these parameters were determined for (I) fracture width during injection, (2) fracture width after the overburden comes to rest on the propping agents, assumed not to be crushed, (3) generated and propped fracture area, (4) location and concentration of propping agents in the fracture when injection ceases, (5) flow capacities of the various propped sections of the fracture and (6) expected increase in the well productivity. The effects of propping agent, formation and fracturing fluid parameters on well productivity are discussed. The parameters that were found to have the most pronounced effects on hydraulic fracturing treat~nents are injection rate, treatment volume, fracturing fluid coefficient, size and amount of propping agent, spearhead volume, well drainage radius and formation capacity. INTRODUCTION Many correlations have been published for predicting effects of various parameters that are considered in the design of hydraulic fracturing treatments. The Carter equation&apos; can be used to predict generated fracture radius as a function of fracture width, fracturing fluid leakoff and other parameters. Fracture width can be determined by use of the Perkins and Kern correlation&apos; in which the fracture width is related to the fracture radius, fluid injection rate and certain formation and fracturing fluid parameters. Wahl and Lowe et aL4 have reported methods of predicting the location of propping agents in fractures when pumping ceases. The former study is applicable to the case where the ratio of propping agent diameter to fracture width is less than 0.1. The latter is applicable when this ratio is greater than 0.1. These studies showed that the propping agent placement in horizontal-radial fractures depends principally on how the individual particles are transported in the fracture by the carrying fluid. Particle transport in fractures is determined by local fluid velocity in the fracture, fluid and particle properties, and the size of the particle relative to the fracture width. The distribution of propping agents, effective overburden pressure and formation rock strength control the propped fracture width6 by controlling the extent to which the propping agent particles embed into the fracture faces. From the distribution of propping agents and the propped fracture width, fracture flow capacities can be calculated or the various regions of the fracture. The flow capacities and the radial extent of these regions can be combined with reservoir information to predict the productivity increases for fractured wells. In all these studies, the effects of certain treatment and/ or reservoir parameters on one facet of fracturing can be predicted only if other facets which the parameters affect are fixed. For instance, fracture width and radius are interrelated; that is, to calculate the value of one, the value of the other must be known. Also, some parameters influence more than one aspect of fracturing. For example, prop-pant transport is a function of both fracture width and fluid viscosity. but fracture width is itself a function of fluid viscosity. Since these calculations are complex and the parameters interrelated, it is not possible to write an equation with which the over-all effects of treatment parameters can be solved explicitly. For these reasons, the correlations for determining the effects of the parameters which are most significant in hydraulic fracturing treatments have been incorporated into a digital computer program. COMPUTER PROGRAM The program, which was written for an IBM 7094 computer, can be used to predict results of most of the combinations and values for the treatment parameters that are ordinarily considered for fracturing treatments. A spearhead of fracturing fluid and a propping agent-carrying fluid with different fluid properties can be taken into account. Also, the total volumes and relative amounts of the spearhead and carrying fluids can be varied. Two different propping agents (as used in tail-in operations) and a wide range of formation properties and injection rates are considered. The computer program (Fig. 12) consists of several sets of calculations. First, the final flooded fracture radius and average fracture width at the cessation of pumping are calculated. This is done by simultaneous solution of the Perkins and Kern fracture width equation and the Carter equation for flooded fracture radius (equations used in the computer program appear in the Appendix). The next step is to determine local fluid velocity in the fracture as a function of time and radius. Since it is not possible to write this function in closed form expression, a table of velocity values is generated by the program and stored for subsequent use. The time span from the beginning of

By H. Conrad, V. Damiano, J. Hanafee, N. Inoue

The effects of hydrostatic pressure up to 400 ksi at 25" to 300°C on the mechanical properties of three forms of commercial beryllium (hot-pressed block, extruded rod and cross-rolled sheet) were investigated. Three effects of pressure were studied: mechanical beharior under pressure, the effect of pressure-cycling, and the effect of tensile prestraining under hydrostatic pressure on the subsequent tensile properties at atmospheric pressure. For all three materials the ductility increased with pressure whereas the flow stress did not appear to be significantly influenced by pressure. An increase in the subsequent atmospheric pressure yield strength generally occurred as a result of pressure-cycling or prestraining under pressure, whereas either no change or a decrease in ductility occurred. The only exception to this was sheet material, which exhibited some improvement in ductility following a pressure-cycle treatment of 304 ksi pressure. The effects of pressure-cycling and prestraining were relatively independent of the temperature at which they were conducted. Stabilized cracks of the (0001) type were found in hot-pressed specimens and {1120) type in extruded and sheet specimens following straining under pressure. Also, pyramidal slip with a vector out of the basal plane, presumably c + a, was identified by electron transmission microscopy for extruded rod and for sheet strained under pressure. Small loops similar to those previously reported were found after straining at pressures of the order of 300 ksi. THE use of beryllium in structures is limited because of its poor ductility under certain conditions. Therefore, one objective of the present research was to determine if the ductility of beryllium at atmospheric pressure could be improved by prior pressure-cycling or prestraining under hydrostatic pressure. Another objective was to study the mechanisms associated with the plastic flow and fracture of the polycrystalline form of this metal with pressure as an additional variable. Since the early work of Bridgman,1 it has been recognized that many materials which are brittle at atmospheric pressure exhibit appreciable ductility when strained under high hydrostatic pressure. This effect has been reported for beryllium by Stack and Bob-rowsky2 and by Carpentier et al.3 and has been attributed to the operation of pyramidal slip systems with slip vectors inclined to the basal plane while cleavage or fracture is suppressed.4 That such slip may occur simply by the application of pressure alone without external straining (pressure-cycling) is suggested by the results on polycrystalline zinc5 and polycrystalline beryllium,6 where nonbasal dislocations with a vector (1123) were reported. A significant improvement in the ductility of the bee metal chromium by pressure-cycling has been reported.7 On the other hand, limited studies on the pressure-cycling of the hcp metals zinc67819 and beryllium6 indicated no improvement in ductility; there only occurred an increase in the yield and ultimate strengths. The study on beryllium was limited to hot-pressed material. Consequently, additional studies on the effects of pressure-cycling on other forms of beryllium seemed desirable, especially since for chromium some authors10 have been unable to detect any improvement in ductility while others find a large improvement.7 That the ductility of polycrystalline beryllium at atmospheric pressure might be improved by prior straining under hydrostatic pressure was suggested by the known beneficial effects of cold work on the ductile-to-brittle transition temperature in the bee metals. It was reasoned that, by straining under hydrostatic pressure, fracture would be suppressed, and during the propagation of slip from one grain to its neighbor dislocations with a vector inclined to the basal plane"-&apos;4 would operate. Upon subsequent straining at atmospheric pressure, these dislocations with a nonbasal vector would continue to operate and thereby reduce the tendency for fracture to occur, by assisting in the propagation of slip across grain boundaries and by interacting with any cracks that may develop. It was recognized that maximum improvement in ductility would probably occur at some optimum amount of prestrain under hydrostatic pressure. If the pre-strain was too small, an insufficient number of dislocations with a nonbasal vector would be activated; if it was too large, internal stresses (work hardening) might increase the flow stress more than the fracture stress, or incipient cracks or other damage could develop. EXPERIMENTAL PROCEDURE 1) Materials and Specimen Preparation. The materials employed in this investigation consisted of hot-pressed block (General Astrometals, CR grade), extruded rod (General Astrometals, GB-2 grade with a reduction ratio of 8:1), and cross-rolled sheet (Brush S200, 0.065 in. thick). The analyses of these materials and mechanical properties at room temperature and atmospheric pressure are given in Table I. The grain size of the hot-pressed block was 15 to 16 µ, that of the extruded rod 10 to 11 µ, and that of the sheet 7 to 10 µ in the rolling plane and 5 to 6 µ in the thickness, all determined by the linear intercept method. Al-

Jan 1, 1969

By R. P. Abendroth

The hydrogen-water vapor ratio at which Fe-Cr alloy, chromium oxide, andiron chromite coexist in equilibrium was determined between 900" and 1200°C. A thermogravimetric method was used to determine equilibrium conditions. The results fit a straight-line relationship in the temperature region studied, and are given by Reduction experiments were also performed to confirm the results of the equilibrium investigation. ThE oxygen pressures at which Fe-Cr alloy, chromium oxide, and iron chromite coexist in equilibrium have been previously determined by Boericke and angert,&apos; Morozov and Novokharski,&apos; and Katsura and uan. Only one determination (at 1300°C) was made by Katsura and Muan, but it agrees with the results of orozov and Novokharski. The results of Boericke and Bangert, however, differ appreciably from the results of these investigators. Previous studies have assumed that the equilibrium metallic phase is pure iron, but Dahl and Van vlack have shown that the iron contains from about 1 wt pct Cr at 1000°C to over 2 wt pct above 1300°C. The chromium oxide also contains a small amount of iron in solid solution. In the present study, hydrogen-water vapor mixtures were equilibrated with the condensed phases, using a therrnogravimetric method to determine equilibrium conditions. The reaction can be written EXPERIMENTAL General Procedure. The starting material was a sintered pellet of Fe2O3-Cr2O3 solid solution with a hole in the center, and was placed on a fused silica hook. This assembly was raised into the preheated hot zone of the furnace in a helium atmosphere, hooked onto a fused silica hangdown suspended from one arm of an Ainsworth Model RV-AU-1 recording balance, and the starting weight determined. A flowing hydrogen-water vapor atmosphere was then exchanged for the helium by evacuation, and the sample reduced until the weight loss indicated the sample composition to be in the alloy-Cr2O3-chromite field. The tem- perature was adjusted incrementally until constant sample weight was achieved for several hours, to within 0.02 mg. A hydrogen-water vapor atmosphere of different composition was then admitted, and the same procedure carried out. At the end of a series of determinations, the sample was examined by X-ray diffraction to verify the presence of the desired phases. Microscopic examination of the silica hook showed no interaction with the sample, nor did it lose any weight. Several criteria were used to insure equilibrium besides constancy of weight. For a given hydrogen-water vapor composition, equilibrium was approached from both oxidizing and reducing sides by varying the furnace temperature slightly. The resulting slow weight loss or gain was observed for several hours. Constant weight could be re-established by returning to the original furnace temperature. The last criterion used was varying the relative amounts of the phases by further reduction or oxidation, and observing any changes in temperature required for constant weight for a given hydrogen-water vapor atmosphere. None were observed. This procedure was essentially the same as approaching the equilibrium from oxidizing and reducing sides, but larger weight excursions were carried out. Sample Preparation. Reagent-grade Fe2O3 and Cr83 powders were mixed in the desired proportions and heated in air at 1250°C for 2 hr. The mixture was re-ground and heated in air overnight at 1250°C. X-ray diffraction showed complete solid-solution formation as a result of this procedure. The solid solution was then pressed into l/2-in.-diam pellets using Carbowax 4000 as a binder. The hole was drilled in the center, and the pellets were sintered 24 hr at 1250°C in air on a bed of Fe2O3-Cr2O3 of the same composition, contained in an alundum boat. After cooling, the pellet surfaces were abraded with 310 paper to remove any surface compositional differences, such as loss of Cr2O3. Chemical analysis of the sintered pellets was 67.16 wt pct CrP3 and 33.02 wt pct Fe203. Atmosphere Generation and Control. The hydrogen-water vapor atmospheres were generated by passing Matheson ultrahigh-purity hydrogen, with no further purification, through two water bubblers contained in a constant-temperature water bath. Since the water-vapor dew points required in this study were below room temperature, the bath was insulated, and was cooled by thermoelectric-immersion devices. The bath temperature was controlled to 0.0l0C. Since rather high flow rates of about 900 ml per min were used through the furnace tube, an independent check of the dew point was made to insure saturation of the hydrogen by the water vapor. Although the dew point could only be determined to within 1/2"C, the determined dew points agreed with the water-bath temper-

Jan 1, 1967

By H. A. Jr. Koch, C. A. Hutchinson

Miscible phase displacement of oil from reservoirs has been emphasized in the past few years. The reason for this emphasis lies in the high oil recovery attainable by this process. Removal of capillary effects in the reservoir leads to recoveries approaching 100 per cent in the area contacted by the miscible phase. The miscible slug process is one means of obtaining a miscible displacement. Here a band or slug of I,PG is injected into the reservoir prior to gas injection. The idea is to maintain the band of LPG "wedged" between the gas and oil phases and thus achieve a miscible phase displacement. A second method lor achieving miscibility is through the injection of a gas which is not miscible with the reservoir fluid but which develops a zone of miscibility in [he reservoir through mass trans-ier with the reservoir oil.&apos; This mass transfer results in either an enrichment of the lean injected gas by intermediates from the oil or an enrichment of the oil by intermediates from a rich injection gas or one that has been enriched on the surface by LPG addition. We are interested here in discussing the process in which miscibility is developed at the displacement front by the evaporation of interrnediatcs from the oil phase into the gas phase. This process "builds up" its own slug of miscible material at the displacement front and therefore does not require the injection of LPG to obtain miscibil-ily. Each process has its own area of applicability. Generally, the high pressure gas process is applicable only with reservoir fluids which con-!ain a high concentration of inter- mediates. If the high pressure gas process is technically feasible at pressures less than 4,500 psi, it is probably more desirable economically than the slug process. The slug process has broad applicability in the shallower reservoirs and with reservoir fluids which contain a relatively low concentration of LPG and natural gasoline constituents. This paper deals with some new concepts of the high pressure gas injection process where it is proposed that flue gas can be substituted for hydrocarbon gas without sacrificing our goal of miscibility. MECHANISM Introduction Considerable effort has been devoted to study of the mechanics ot the high pressure gas injection proc-one generalization result-ing from some of these studies was that the composition of the injected gas is relatively unimportant in establishing the miscibility pressure* for a given reservoir fluid. This generalization is correct for the composition range of gases typically encountered in the field. Two such gases are a gasoline plant tail gas containing 85 per cent methane and 15 per cent ethane, and a field separator gas containing 70 per cent methane and 30 per cent heavier components. The most important factor which sets the miscibility pressure in the operation is the reservoir fluid composition, particularly the concentration of LPG-natural gasoline constituents. The injected gas is the agent by which the LPG-natural gasoline constituents are concentrated at the displacing front to create a miscible displacement. Based on these results, it appeared feasible that some inexpensive gas, such as flue gas, might be substituted for hydrocarbon gas for use in the high pressure gas process. A re-examination of the phase relations of the high pressure gas injection process should clarify the principle behind using flue gas (essentially nitrogen) as an injection gas. Three Component Diagram The phase relations of the high pressure gas injection process have been illustrated by the use of the three component diagrams.&apos;," In Fig. 1 we have arbitrarily represented the multi-component reservoir system by three components; methane, ethane through hexane, and heptanes plus. The solid curve ABC is the phase boundary curve. It represents the locus of compositions which have fixed saturation pressure at a fixed temperature; the lower branch AB shows bubble point compositions, the upper branch BC, the dew point compositions. Point B is the coniposition of the critical mixture at this temperature and pressure. The dashed lines (tie lines) connect vapor and liquid compositions which are in equilibrium. Let us consider Reservoir Fluid D which we wish to displace in 21 miscible manner by gas injection. Let us further restrict the discussion to the case where miscibility between an injection gas and the reservoir fluid at the displacement front is developed by gas enrichment in the reservoir. For this case, any gas whose composition lies between Points C and E on the right side of the three component diagram can be used to give a miscible displacement of Reservoir Fluid D. This is true because the more mobile injected gas moves faster than the displaced oil and is in continuous contact with virgin oil at the displacement front. This leads to a continuing enrichment 01&apos; the gas at the displacement front by evaporation of the C, - C,