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By T. A. Read, M. S. Wechsler, D. S. Lieberman

A theoretical analysis of the austenite-martensite transformation is presented which predicts the habit plane, orientation relationships, and macroscopic distortions from a knowledge only of the crystal structures of the initial and final phases. THIS paper presents a new theory of the formation of martensite. This theory makes possible the calculation of the austenite planes on which the martensite plates form, the orientation relationship between the austenite and martensite crystal axes, and the macroscopic distortions which are observed. The only input data needed are the crystal structures and lattice parameters of the austenite and martensite. Considerable effort has been devoted over the past thirty years to the development of an understanding of the crystallographic features of martensite reactions. Much of this work has been done on steels and iron-nickel alloys, for which a great deal of data has been accumulated concerning the shape and orientation of the martensite plates, the relative orientations of the austenite and martensite crystal axes, and the observable distortions which result from transformation. These observations are reviewed in refs. 1, 2, and 3. The first major step toward an understanding of these phenomena was made in 1924 by Bain,' who showed that the a body-centered cubic structure can be produced from the 7 face-centered cubic structure by a contraction of about 17 pct in the direction of one of the austenite cube axes and an expansion of 12 pct in all directions perpendicular to it. Since that time, most of the efforts at further interpretation have been made by investigators who have worked from the phenomenological data, incorporating some of the information from the lattice properties, and have sought an analysis into likely deformations which would produce the observed results."- "11 but the three most recent papers on the subject have already been reviewed in some detail." Machlin and Cohenl0 measured the components of the distortion matrix and verified that the habit plane is a plane of zero distortion and rotation for the (259) case. They showed that the measured distortion matrix, when applied to the parent lattice, does not yield the product lattice and hence some inhomogeneous distortion must occur. Frank,u working from the lattice properties and taking some clues from the observations, considered the correspondence of close-packed rows and planes in the austenite and martensite. He predicted substantially the observed lattice relationship and habit plane for certain steels which have a (225) habit. Geisler12 suggested that there is a natural tendency for the habit plane to be a (111) and postulated certain slip processes to account for the fact that the experimentally observed habit plane is irrational and deviates from the assumed one. The present work differs from previous treatments of martensite formation in that it permits calculation of all the major manifestations of the process. Habit plane indices, orientation relationships, and observable distortions are all calculated from a knowledge of the crystal structures of the initial and final phases alone. The calculations contain no adjustable parameters. The agreement found between calculated results and the observations reported in the literature constitutes powerful evidence in favor of the mechanism of martensite formation proposed. The theory is applicable to systems other than steel (as is discussed later in this paper) which exhibit a diffusionless phase change but because of the wide-spread interest in the austenite-martensite transformation, particular attention will be given to the iron-base alloys. For other systems which undergo a similar face-centered cubic to face-centered tetragonal transformation, the mathematical treatment is identical with that presented here. Hence the theory successfully describes the transformation in the indium-thallium alloy.'" Homogeneous Transformation to Martensite The distortion which any homogeneously transforming volume of austenite undergoes in order to become martensite is shown in Fig. 1, as was first suggested by Bain.' (This distortion will hereafter be referred to as the "Bain distortion.") This specification of a contraction along one cube axis ;ombined with an expansion in all directions perpendicular to this axis describes what is properly called the "pure" distortion associated with this transformation. The distinction between a "pure" and an "impure" distortion plays an important part in the discussion which follows. A "pure" distortion is characterized by the existence of at least one set of orthogonal axes fixed in the body which are not rotated by the distortion. (These are called the "principal axes" of the distortion.) No such set of axes exists in the case of an "impure" distortion. On the other hand, an impure distortion can always be represented as the result of a pure distortion combined with the rotation of the specimen as a rigid body. For a given impure distortion the corresponding pure distortion

Jan 1, 1954

By D. Havlena, A. S. Odeh

The material balance equation used by reservoir engineers is arranged algebraically, resulting in an equation of a straight line. The straight line method of analysis imposes an additional necessary condition that a successful solution of the material balance equatiott should meet. In addition, this algebraic arrangement attaches a dynamic ineuning to the otherwise static material balance equation. The straight line method requires the plotting of one variable group vs mother variable group. The sequence of the plotted points as well as the general shape of the resulting plot is of utmost importance. Therefore, one cannot progrm the method entirely on a digital computer ar is usually done in the routine solution of the material balance equation. If this method is applied, then plotting and anaIysis are asential. Only the appropriate equations and the method of analysis and interpremtion with comments and discussion are presented in this paper. Illustrative field examples for the various cases treated are deferred to a subsequent writing. INTRODUCTION One of the fundamental principles utilized in engineering work is the law of conservation of matter. The application of this principle to hydrocarbon reservoirs for the purpose of quantitative deductions and prediction is termed "the material balance method of reservoir analysis". While the construction of the material balance equation (MBE) and the computations that go with its application are not difficult tasks, the criteria that a successful solution of the MBE should fulfill have always been a problem facing the reservoir engineer. True and complete criteria should embody necessary and dcient conditions. The criteria which the reservoir engineer uses possess a few necessary but no sufficient conditions. Because of this, the answers obtained from the MBB are always open to question. However, the degree of their acceptability should increase with the increase in the number of the necessary conditions that they should satisfy. Generally, the necessary conditions commonly used are (1) an unspecified consistency of the results and (2) the agreement between the MBE results and those determined volumetrically. This second criterion is usually overemphasized. Actually, the volumetrically determined results are based on geological and petrophysical data of unknown accuracy. In addition, the oil-in-place obtained by the MBE is that oil which contributes to the pressure-production history,' while the volumetrically calculated oil-in-place refers to the total oil, part of which may not contribute to said history. Because of this difference, the disagreement between the two answers might be of paramount importance, and the concordance between them should not be overemphasized as the measure of correctness of either one. In this paper, a third necessary condition of mathematical as well as physical significance is discussed. It is not subject to any geological or petrophysical interpretation, and as such, it is probably the most important necessary condition. It consists essentially of rearranging the MBE to result in an equation of a straight line. This straight line method of the MBE solution has invalidated a few long time accepted concepts. For instance, it has always been advocated that if a water drive exists, but one neglects to take it into account in the MBE, the calculated oil-in-place should increase with time. The straight line method shows that in some cases, depending on the size of the neglected aquifer, the calculated oil-in-place might decrease with time. The straight line method requires the plotting of a variable group vs another variable group, with the variable group selection depending on the mechanism of production under which the reservoir is producing. The most important aspect of this method of solution is that it attaches a significance to the sequence of the plotted points, the direction in which they plot, and to the shape of the resulting plot. Thus, a dynamic meaning has been introduced into the picture in arriving at the final answer. Since the emphasis of this method is placed on the interpretation of the sequence of the points and the shape of the plot, one cannot completely automate the whole sequence to obtain "the best value" as normally done in the routine application of the MBE. If one uses the straight line method, then plotting and analysis are musts. The straight line method was first recognized by van Everdingen, et al,2 but for some reason it was never fully exploited. The advantages and the elegance of this method can be more appreciated after a few cases are carefully treated and worked out by it. SOLUTION OF THE MATERIAL BALANCE EQUATION SATURATED RESERVOIRS The MBE for saturated reservoirs written in AIME symbols is

By R. M. Waterstrat

The phases occurring in the V-Mn system were studied by means of X-yay diffraction and metallo-paphic techniques, using are-melted alloy specimens annealed in the temperature range 800° to 1150°C and quenched. The bcc solid solution extends at 1250°C all the way from vanadium to 6-manganese. Below 1050°C the a-phase is formed, and the terminal a-manganese phase is stabilized up to about 900°C by vanadium in solid solution. IN the only previous general survey of the V-Mn system Cornelius, Bungardt and Schiedtl reported the existence of three intermediate phases corresponding to the approximate compositions VMn,, VMn, and V5Mn. The phase VMn8 has recently been identified as a o phase2 but the alloy VMn was found to have a bcc structure2 corresponding apparently to the vanadium solid solution rather than to the large cubic unit cell reported by Cornelius et al. 1 Subsequent work by Rostoker and Yamamoto3 has shown that the vanadium-base bcc solid solution extends to at least 15 pct Mn at 900°C. An alloy corresponding to the composition VMn, was examined by Elliott,4 who reported that the as-cast sample as well as samples annealed at 1200o and 1300°C had bcc structures, but that annealing at 1000°, 800") and 600°C produced two phases. One of these phases was apparently the bcc solid solution and the other resembled the o phase structure. Hellawell and Hume-Rothery5 established the phase relationships in manganese-rich alloys above 1000°C, and showed that the o phase in this system is replaced by the 6 Mn (bcc) solid solution at temperatures above 1050°C. These results suggest that a continuous bcc solid solution may exist above 1050°C between vanadium and 6 Mn. The present investigation was undertaken in order to develop more complete information in regard to this system. EXPERIMENTAL METHODS The alloys used in the present work were prepared by arc-melting electrolytic manganese having a minimum purity of 99.9 pct and vanadium lumps with a purity of 99.7 pct. The major impurities present in these metals were carbon, nitrogen, and oxygen and this would account for the small percentage of nonmetallic inclusions observed metal-lographically. The arc-melting was at first performed under a helium atmosphere and it was necessary to keep the melting times as short as possible in order to minimize the loss of manganese by vaporization. It was later found that the evaporation of manganese was considerably reduced when the melting was done under argon atmosphere. The final composition of each alloy was calculated by assuming that the total weight loss during melting was due to evaporation of manganese. Compositions which were calculated in this manner agreed reasonably well with the results of chemical analysis, as shown in Table I. Spectrographic analysis revealed the presence of contamination by tungsten, but in no case was the percentage of tungsten greater then 0.4 at. pct. The specimens were in each case broken in half and the fractured section was examined visually and microscopically for evidence of inhomogeneity. Each specimen was homogenized at temperatures near l100°C, as shown in Table I. After this treatment most specimens consisted of large columnar grains of the bcc vanadium solid solution. The etchant used in most of the metallographic work consisted of 20 pct nitric acid, 20 pct hydro-flouric acid, and 60 pct glycerine. It was found that this etchant would clearly delineate the phases present in these alloys although it does not produce any striking contrast between the phases. For certain manganese-rich alloys, a 1 pct aqueous solution of nitric acid was used. This etchant gave a brown color to the a-manganese phase, whereas the o phase was virtually unattacked and appeared very light as shown in Fig. 1. The etchants used by Cornelius et a1.l were found to produce spurious effects in some of these alloys. In particular, the vanadium-rich alloys etched in hot sulfuric acid often appeared to consist of two phases when both X-ray diffraction and etching with the glycerine-acid mixture indicated the presence of single phase bcc solid solution. A few percent of what appears to be an oxide or nitride phase was found at the grain boundaries and in the interior of the grains, especially in the vanadium-rich alloys. All alloys were annealed in sealed silica tubes containing 1 atm of pure argon and these tubes were then quenched in cold water. Although some manganese loss occurred during annealing, the loss seemed to be confined to the surface of the speci-

Jan 1, 1962

By J. Campbell

Various models are discussed for the evaluation of the negative pressures which may occur in solidifying materials which exhibit various deformation modes: elastic-plastic, Bingham, viscous, or creep flow. The inadequacy of the previously proposed elastic-plastic solution for solidifying metals is revealed by comparison with the more reliable creep results which are given graphically for aluminum, copper, nickel, and iron. The maximum tensions experienced in the liquid phase of solidifying spheres ranging in size from large castings to submicron powders are in the range from —10' to —105 atm for these metals. THERE has been much recent interest in the negative pressures associated with the volume change on solidification and in the possibility of the occurrence of cavitation. Considering the freezing of a highly supercooled liquid, an attempt to evaluate the stresses in the liquid ahead of the rapidly moving solidification front has been made by Horvay1 on a microscale and by Glicksman2 on a macroscale. In a casting of a wide freezing range alloy, the pressure differential due to viscous flow of residual liquid through the pasty zone has been discussed by Piwonka and Flemings,3 In a previous publication4 the author has attempted to estimate the negative pressure occurring in the residual liquid of a spherical casting, employing an elastic-plastic model to describe the collapse of the solidified shell under the internal tension. An earlier model assuming a rigid shell was shown to be inaccurate by many orders of magnitude. The elastic-plastic model is critically reviewed here, and other models are developed which are thought to be more closely related to metals and other materials near their melting points. The spherical geometry (Fig. 1) is chosen because the highest shrinkage pressures would be developed, although the analyses are readily adaptable to cylindrical geometry. A parallel sided casting experiences little internal tension because of the relatively easy dishing inward of the sides. (This commonly observed phenomenon has previously been attributed solely to atmospheric pressure.) Furthermore, small regions of confined liquid in a large solidified volume of a casting approximate reasonably well to spherical geometry. ELASTIC-PLASTIC MODEL The author has shown4 that as solidification proceeds the internal hydrostatic tension builds up until the elastic limit of the shell is exceeded. At this point the internal pressure is closely -2Y/3. Subsequently a plastic zone spreads from the inner surface toward the outer surface of the shell. When the whole casting is deforming plastically a rather more generalized analysis taking account of the externally applied pressure PA + 2y/b gives the internal pressure as: P = Pa + 2y/a + 2ys/b - 2 Y In(b/a) [1] The 2y/a and 2ys/b terms result from the tendency of the liquid-solid and solid-vapor interfaces to shrink, reducing their energy, and thereby helping to collapse the solid phase and compress the liquid phase. The 2y/b term would be important only for powders. The last term arises because of the plastic restraint of the solid, resisting collapse and so effectively expanding the residual liquid. From Eq. [I] it is easily shown that there is a minimum in the pressure at the radius amia= y/Y [2] which is of the order of 103K for the metals aluminum, copper, and iron, and corresponds to the minimum pressure Pmin = 2 Y[l-ln(bY/y [3] The results of a fully worked out elastic-plastic solution are given in a previous reporL4 The main criticism which may be leveled at this analysis when applied to metals at their melting points is the strong dependence of the yield stress on the strain rate. The strain rate varies with both solidification conditions (e.g., whether chill-cast or slowly cooled) and during solidification, as is indicated in the following section. Thus an appropriate choice of Y is very arbitrary. Before proceeding to a discussion of models which are strain-rate-dependent, it is necessary to evaluate the strain rate as a function of the rate of solidification. SOLIDIFICATION RATE Various empirical relations have been deduced5 for the rate of thickening of the solid shell by pour- out tests on partially solidified spheres. These, however, are unsatisfactory for our purposes since they become very inaccurate when the liquid core is very small. A theoretical approach is therefore necessary, and some solutions are set out below. Making the assumptions of constant surface temperature of the casting during freezing, no superheat, and a material freezing at a single temperature, Adams8 deduces the approximate solution: which becomes when b » a: Employing a semiempirical approach vallet6 finds:

Jan 1, 1969

By H. M. Cyr, T. F. Steele, C. W. Siller

The development of a new metallurgical roasting device is described. It consists of a refractory column into which air is injected at various levels, forming several superimposed fluidized beds with no supporting grates. When pelleted zinc sulphide concentrates are charged, the roasted product needs no further sintering before reduction to metal. WHEN a gas such as air is blown upward with increasing velocities through a loose mass of solid particles, marked changes in the physical behavior of the particles are noted. At first, when the velocity of the gas is insufficient to support any of the solid, the mass constitutes a "fixed bed." As the gas velocity increases until the pressure drop through the bed approaches the effective weight of the bed per unit area, the bed expands until the solid particles are supported by the air rather than by the lower particles. Some vibration of the particles becomes apparent, but little mixing occurs. This condition is called a "quiescent fluid bed." A further increase in gas velocity imparts more separation and more motion to the individual particles until a condition of turbulence is reached. This "turbulent fluid bed" resembles a rapidly boiling liquid with the characteristic highly agitated diffuse surface and many small eruptions of the boiling mass. Different degrees of turbulence can be generated and all produce excellent mixing. The final stage occurs when the gas velocity becomes so great as to create a "dispersed suspension." Here no surface of the mass is defined and the gas carries solid particles out of their original positions. These changing conditions of fluidization have been studied carefully and pertinent nomenclature standardized by a committee of the American Institute of Chemical Engineers.' Many mathematical analyses2-3 have been made of the forces acting in a fluid bed. These analyses are invaluable, especially for the design of column sizes and selection of equipment. However, in a metallurgical process involving solids of many sizes with changing densities, varying temperatures, and changing gas compositions within the bed, calculations based on theory become approximate. Optimum operating conditions then are best determined experimentally. Many applications have been made of the principles of fluid-bed action by mechanical, chemical, and metallurgical engineers. Especially when good con- tact between reacting solids and gases is desired, very effective results are obtained from fluid beds. They permit excellent temperature control and uniformity throughout a mass of solids in fluid action. Heat transfer to walls and any coolers is high, and fast reaction rates are attained because the solid surfaces are continuously swept clean. The main disadvantages of fluid-bed operations are the danger of short-circuiting in a single bed, danger of incipient sintering which stops action, the necessity of avoiding large changes in particle size or density during roasting, and dust losses when particles of the charge are carried out with exit gases. In the metallurgical field the roasting of sulphide ores to form oxides and sulphur dioxide appears to combine several operating conditions which can be achieved to advantage in a fluid bed. Roasting involves a solid-gas reaction where a high reaction rate is necessary for high capacity, where good temperature control is important in order to prevent sintering, where good heat transfer is needed, and where the density of the solids, when changing from sulphides to oxides, is not largely changed. Short-circuiting, however, constitutes a major problem when a single fluid bed is used. Because of the turbulence of the bed, an entering particle may be in the region of the discharge before it is roasted. Hence, to attain a satisfactorily low sulphur in the calcine, a long average residence time with correspondingly low capacity is required. The solution to this difficulty is the use of multiple stages, which in the conventional fluid-bed design requires separate hearths with feed and discharge mechanisms for each stage. A further practical difficulty in fluid-bed roasting of flotation zinc concentrates is their fine particle size which makes a true fluid action without excessive carry-over of dust very difficult to attain, especially when the large air volumes necessary for high capacity are used. A New Design After considerable experimentation in the laboratory and on a semipilot-plant scale, a new method and equipment for roasting were devised which provided a unique solution to these problems. A detailed account of this development appears in the patent literature," and many of the variations of this development reported herein are the subject of

Jan 1, 1955

By R. H. Gautschi, F. C. Langenberg

Rare-earth additions were made to laboratory heats of Type 310 stainless to observe their effect on as-cast ingot structure, nitrogen and sulfur contents, and nonmetallic inclusions. Lanthanum had a grain-refining effect in 30-lb ingots, but results with 200-lb ingots were inconsistent. Cerium, lanthanum, and misch metal lowered the sulfur content when the sulfur exceeded 0.015 pct and the rare-earth addition was greater than 0.1 pct. The rare-eardh content in the metal dropped very rapidly within the first few minutes after the addition. The size, shape, and distribution of nonmetallic inclusions was not changed in 30-lb ingots, but changes were noticed in larger ingots. RARE-earth* additions have been made to austenitic Cr-Ni and Cr-Mn steels to improve their hot workability. The high alloy content of these steels often results in a considerable resistance to deformation and inherent hot shortness at rolling temperatures, particularly in larger ingots. Rare earths in the metallic, oxide, or halide form are usually added to the steel in the ladle after deoxidation although they can be added in the furnace prior to tap or in the molds during teeming. The literature- indicates that the effects of rare-earth treatments on these stainless steels are not consistent, and sometimes even contradictory. Since no mechanism has been presented which satisfactorily accounts for the claimed improvements, the effects of rare earths are a qualitative matter. The work described in this paper was initiated to expand the knowledge of the effects of rare-earth additions on melting variables such as ingot structure, chemical analysis, and nonmetallic inclusions. REVIEW OF LITERATURE Ingot Structure—Rare-earth additions to stainless steels have been reported to cause a change in primary ingot structure in that there are fewer coarse columnar grains. However, the results are inconsistent. While one investigation1 has shown a large reduction in coarse columnar crystals, another2 has been unable to observe this effect, particularly when small ingots were poured. Post and coworkers3 observed ingot structures for a number of years and found that the columnar type of structure is not definitely a cause of any particular trouble in rolling or hammering, provided the alloy is ductile. Knapp and Bolkcom4 found rare-earth additions to be quite effective in preventing grain coarsening in Type 310 stainless steel. Chemical Analysis—Many effects of rare-earth treatment on chemical analysis have been claimed in the literature. Russell5 observed that some sulfur is removed by rare-earth metals, and that a high initial sulfur content improved the efficiency of sulfur removal. Lillieqvist and Mickelson6 report that rare-earth treatment causes sulfur removal in basic open-hearth furnaces, but not in basic lined induction furnaces. Knapp and Bolkcom found no sulfur removal in acid open-hearth and acid electric furnaces, probably because the acid slag can not retain sul-fides. snellmann7 showed that sulfur could be lowered apprecfably with rare-earth additions; however, a sulfur reversion occurred with time. Langenberg and chipman8 studied the reaction CeS(s) = Ce(in Fe) + S(in Fe), and found the solubilit product [%Ce] [%S] equal to (1.5 + 0.5) X 10-3'at 1600°C. Results in 17 Cr-9 Ni stainless were about the same as those in iron. Beaver2 treated chromium-nickel steels with 0.3 pct misch metal and observed some reduction in the oxygen content. He also noted an inconsistent but beneficial effect of rare earths when tramp elements were present in amounts sufficient to cause difficulty in hot working. It is not known whether rare earths reduce the content of the tramp elements or change the form in which these elements appear in the final structure. No quantitative data are available concerning a possible effect of rare-earth treatment on hydrogen and nitrogen contents. However, Schwartzbart and sheehan9 stated that additions of rare earths had no effect on impact properties when the nitrogen content was low (0.006 pct), but served to counteract the adverse effects of high nitrogen content (0.030 pct) on these properties. Knapp and Bolkcom4 analyzed open-hearth heats in the treated and untreated conditions and found the nitrogen content (0.006 pct) to be unaffected. These two results lead to the speculation that rare-earth additions can reduce the nitrogen content to a certain level. Decker and coworkers10 have observed that small amounts of boron or zirconium, picked up from magnesia or zirconia crucibles, increased high-tem-

Jan 1, 1961

By Glenn S. Wyman

CHUQUICAMATA, located in northern Chile in the Province of Antofagasta, is on the western slope of the Andes at an elevation of 9500 ft. Because of its position on the eastern edge of the Atacama Desert, the climate is extremely arid with practically no precipitation, either rain or snow. All primary blasting in the open-pit mine at Chuquicamata is done by the churn drill, blasthole method. Since 1915, when the first tonnages of importance were removed from the open pit, there have been many changes in the blasting practice, but no clear-cut rules of method and procedure have been devised for application to the mine as a whole. One general fact stands out: both the ore and waste rock at Chuquicamata are difficult to break satisfactorily for the most efficient operation of power shovels. Numerous experiments have been made in an effort to improve the breakage and thereby increase the shovel efficiency. Holes of different diameter have been drilled, the length of toe and spacing of holes have been varied, and several types of explosives have been used. Early blasting was done by the tunnel method. The banks were high, generally 30 m, requiring the use of large charges of black powder, detonated by electric blasting caps. Large tonnages were broken at comparatively low cost, but the method left such a large proportion of oversize material for secondary blasting that satisfactory shovel operation was practically impossible. Railroad-type steam and electric shovels then in service proved unequal to the task of efficiently handling the large proportion of oversize material produced. The clean-up of high banks proved to be dangerous and expensive as large quantities of explosive were consumed in dressing these banks, and from time to time the shovels were damaged by rock slides. As early as 1923 the high benches were divided, and a standard height of 12 m was selected for the development of new benches. The recently acquired Bucyrus-Erie 550-B shovel, with its greater radius of operation compared to the Bucyrus-Erie 320-B formerly used for bench development, allowed the bench height to be increased to 16 m. Churn drill, blasthole shooting proved to be successful, and tunnel blasts were limited to certain locations where development existed or natural ground conditions made the method more attractive than the use of churn drill holes. Liquid oxygen explosive and black powder were used along with dynamite of various grades in blast-hole loading up to early 1937. Liquid oxygen and black powder were discontinued because they were more difficult to handle due to their sensitivity to fire or sparks in the extremely dry climate. At present ammonium nitrate dynamite is favored because of its superior handling qualities and its adaptability to the dry condition found in 90 pct of the mine. In wet holes, which are found only in the lowest bench of the pit and account for the remaining 10 pct of the ground to be broken, Nitramon in 8x24-in. cans, or ammonium nitrate dynamite packed in 8x24-in. paper cartridges, is being used. This latter explosive, which is protected by a special antiwetting agent that makes the cartridges resistant to water for about 24 hr, currently is considered the best available for the work and is preferred over Nitramon. Early churn drill hole shots detonated by electric blasting caps, one in each hole, gave trouble because of misfires caused by the improper balance of resistance in the electrical circuits. Primarily, it was of vital importance to effect an absolute balance of resistance in these circuits, the undertaking and completion of which invariably caused delays in the shooting schedule. Misfires resulting from the improper balance of electrical circuits, or from any other cause, were extremely hazardous, since holes had to be unloaded or fired by the insertion of another detonator. The advent of cordeau, later followed by primacord, corrected this particular difficulty and therefore reduced the possibility of missed holes. After much experimentation, the blasting practice evolved into single row, multihole shots, with the holes spaced 4.5 to 5 m center to center in a row 7.5 to 8 m back from the toe. Sucti shots were fired from either end by electric blasting caps attached to the main trunk lines of cordeau or primacord. The detonating speed of cordeau or primacord gave the practical effect of firing all holes instantaneously. Double row and multirow blasts, fired instantaneously with cordeau or primacord, proved to be unsatisfactory in the type of rock found at Chuquica-

Jan 1, 1953

By R. H. Walpole, J. M. Kane

IN all probability the mining and metallurgical industry as a whole can demonstrate a larger ecorlomic return from installation of dust-control equipment than any other major industrial group. This fact has partially accounted for the marked increase of dust-control installations made during the past decade. While the primary objectives for installation of dust-collecting systems are improved working and operating conditions for men and equipment, the fact that an economic return can be anticipated on salvageable materials is an added advantage which shows in partial or complete equipment write-off. The conditions apply to most phases of the mining, milling, and smelting industry, both non-metallic and metallic. As with any mechanical devices, selection of suitable dust collector equipment involves evaluation of available products with characteristics most nearly meeting conditions of the application at hand. When there is valuable product to be collected, and/or when there are possibilities of air pollution or public nuisance, collector selection is often guided by the maxim of "highest available collection efficiency at reasonable cost and reasonable maintenance." A brief review of dust collector designs will permit outlining of major characteristics of each group. Final selection will involve detailed data against a background of the problem under consideration. The dry centrifugal collectors, see Fig. 1, represent a group of low cost units with minimum maintenance. They are subject to abrasion under heavy abrasive dust loads and to plugging with moist materials. Efficiency drops off rapidly on particle sizes below the 10 to 20 micron group. Because of the large amounts of —10 micron particles in most mining dust problems, they will normally be used as primary collectors and will be followed by high efficiency units. This combination is cspecially popular where the bulk of material is desired in a dry state with wet collection indicated for the final cleanup portion. In remote plant locations, dry centrifugal~ can be used alone if product in dust form has no value or if dust loading is light enough to eliminate a nuisance in the plant area. Where high efficiency dust colleotion equipment must be selected, choice will normally involve fabric arresters, wet collectors, or high voltage Electro-Static precip-itators. Fabric arresters, see Fig. 2, rely on the passing of dust-laden air at low velocity through filter fabric. Velocity ranges from 1 to 3 fpm for the usual installation and may be as high as 10 to 20 fpm in arrangements where automatic frequent vibration or continuous cleaning of the filter media is employed. Fabric is normally suspended in either stocking type or in an enlvelope shape. Collection efficiency is excellent even on sub-micron particle sizes. Equipment is bulky, must be vibrated to remove the collected dust load, and is restricted in applications from temperature and moisture standpoints. Condensation of moisture on the fabric filter mcdia causes plugging of the passages with great reduction in air flow. Temperatures for the usual medias of cotton or wool are 180" and 200°F maximum, although the introduction of synthetic materials such as nylon, orlon, and glass cloth have increased the possibilities of this type of collector for higher temperature applications. The wet-type collector may employ a number of different principles so that entering dust particles in the gas stream are wetted and removed. Principles usually include impingement on collector surface or water droplets, often in combination with centrifugal forces. Variety of wet collector designs is indicated by typical collectors illustrated in Figs. 3 and 4. Collection efficiency is a function of the particular design, although the better collectors will have high collection efficiency on particles in the 1-micron range. Wet collectors have the advantage of handling hot or moist gases, take up small space, and eliminate secondary dust problems during the disposal of the material. At times collection of the material wet is a disadvantage. Wet collectors may also be subject to corrosion and freezing factors. The high voltage Electro-Static precipitator, see Fig. 5, is probably the most expensive type of high efficiency collector. It finds its applications generally in problems in which collectors previously discussed cannot be employed. Its collection efficiency is based on its design features and can be excellent on the finest of fume particles. Material is normally collected dry. Gas temperatures are of no great concern as long as condensation does not occur within the dry type of precipitator and the temperatures do not exceed the limits for materials used in its construction. As with the fabric arrester, provisions

Jan 1, 1954

By E. F. Thomas

THE late George S. Rice was active in the inves--I- tigation of bumps, particularly in the last ten years of his career as chief mining engineer of the U. S. Bureau of Mines. Since most of his investigation was carried out in Great Britain, continental Europe, and—to a lesser extent—Canada, his thinking on prevention was influenced considerably by the experience of those countries. It is not surprising, therefore, that when he was called upon a few years before his retirement to investigate bumps in the U. S. and suggest ways to prevent them, he turned to longwall mining. A longwall method had been most successful in combating the bump hazard in mining coal under deep cover, especially in Great Britain, but the prevailing method there at the time was advancing longwall mining, which he knew was uneconomical under U. S. mining conditions. For this reason he proposed a modified retreating longwall system that he believed included the best features of the advancing method. As brought out by Rice,' if the cover is 2000 ft and 50 pct of the coal is extracted, the static load on the remaining pillars will be about 4000 psi, which exceeds the ultimate crushing strength in most instances. If the pillar coal is overloaded before a pillar line is established, then the abutment zone preceding a line of extraction is no place to split pillars or extract them by any method other than an open-end system. Rice therefore advocated open-end mining, preferably by longwall, but he was willing to compromise with long-face mining if the longwall method was not acceptable. Rice's system was put into operation in a mine in Harlan County, Kentucky,3 but subsequent experience has shown that it did not take into account two important factors—avoidance of pillar-line points and maintenance of adequate development in advance of the pillar-line abutment area. For ten years after Rice's retirement the USBM did little investigation and research on bumps, chiefly because so few were occurring that there was not much cause for alarm. But in 1951 there were three occurrences involving fatal injuries, and the Bureau began a statistical survey in that year. C. T. Holland, head of the department of mines at Virginia Polytechnic Institute, was retained as a consultant. The resulting study' of 117 case histories brought out these important conclusions: 1) Almost invariably the bump occurred in a locality affected by the abutment zones of one or more pillar lines. 2) In most cases the locality of the bump was influenced by the abutment zones of more than one pillar line. The term pillar-line point has been used for many years in the Appalachian region for such a situation. Point is used in the geographical rather than the mathematical sense. 3) In pillar-line extraction the following practices are safest in preventing bumps: a. The mine layout should provide for pillars of uniform size and shape along the extraction line. b. The mine layout should be planned so that no development need be done in the abutment zone of a pillar line. c. The layout should permit open-end extraction of pillar lines from the next goaf, so that it will not be necessary to resort to pocket mining, splitting pillars, or any practice that will involve driving in the direction of the goaf within the abutment zone. d. Pillars should be large enough to support area without undue roof and floor convergence before establishment of a pillar line. These are, of course, generalities, and while they are useful in laying out areas where bumps can be expected, they are of limited help in many mines that were committed to a system of mining before it was realized that they were subject to bumps. Under such conditions it becomes necessary to choose between the following alternatives: 1) Abandon the territory, except for pillars that offer no extraction problems. 2) Through experience select the pillars that are most heavily loaded, and, by augering, induce bumps from a safe vantage point so that impinged loads are relieved. This method was first developed at the Gary, W. Va. mines of U. S. Steel Corp. and later adapted to mining thick coal beds at Kaiser Steel's Sunnyside mine in Utah. No scientific method is available to determine where to drill within a loaded pillar. Although this method of unloading has worked very successfully at Gary—with one exception—

Jan 1, 1959

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' 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.' 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' 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 Glenn S. Wyman

CHUQUICAMATA, located in northern Chile in the Province of Antofagasta, is on the western slope of the Andes at an elevation of 9500 ft. Because of its position on the eastern edge of the Atacama Desert, the climate is extremely arid with practically no precipitation, either rain or snow. All primary blasting in the open-pit mine at Chuquicamata is done by the churn drill, blasthole method. Since 1915; when the first tonnages of importance were removed from the open pit, there have been many changes in the blasting practice, but no clear-cut rules of method and procedure have been devised for application to the mine as a whole. One general fact stands out: both the ore and waste rock at Chuquicamata are difficult to break satisfactorily for the most efficient operation of power shovels. Numerous experiments have been made in an effort to improve the breakage and thereby increase the shovel efficiency. Holes of different diameter have been drilled, the length of toe and spacing of holes have been varied, and several types of explosives have been used. Early blasting was done by the tunnel method. The banks were high, generally 30 m, requiring the use of large charges of black powder, detonated by electric blasting caps: Large tonnages were broken at comparatively low cost, but the method left such a large proportion of oversize material for secondary blasting that satisfactory shovel operation was practically impossible: Railroad-type steam and electric shovels then in service proved unequal to the task of efficiently handling the large proportion of oversize material produced. The clean-up of high banks proved to be dangerous and expensive as large quantities of explosive were consumed in dressing these banks, and from time to time the shovels were damaged by rock slides. As early as 1923 the high benches were divided, and a standard height of 12 m was selected for the development of new benches. The recently acquired Bucyrus-Erie 550-B shovel, with its greater radius of operation compared to the Bucyrus-Erie 320-B formerly used for bench development, allowed the bench height to be increased to 16 m. Churn drill, blasthole shooting proved to be successful, and tunnel blasts were limited to certain locations where development existed or natural ground conditions made the method more attractive than the use of churn-drill holes. Liquid oxygen explosive and black powder were used along with dynamite of various grades in blasthole loading up to early 1937. Liquid oxygen and black powder were discontinued because they were more difficult to handle due to their sensitivity to fire or sparks in the extremely dry climate. At present ammonium nitrate dynamite is favored because of its superior handling qualities and its adaptability to the dry condition found in 90 pct of the mine. In wet holes, which are found only in the lowest bench of the pit and account for the remaining 10 pct of the ground to be broken, Nitramon in 8x24-in. cans, or ammonium nitrate dynamite packed in 8x24-in. paper cartridges, is being used. This latter explosive, which is protected by a special antiwetting agent that makes the cartridges resistant to water for about 24 hr, currently is considered the best available for the work and is preferred over Nitramon. Early churn drill hole shots detonated' by electric blasting caps, one in each hole, gave trouble because of misfires caused by the improper balance of resistance in the electrical circuits. Primarily, it was of vital importance to effect an absolute balance of resistance in these circuits, the undertaking and completion of which invariably caused delays in the shooting schedule. Misfires resulting from the improper balance of electrical circuits, or from any other cause, were extremely hazardous, since holes had to be unloaded or fired by the insertion of another detonator. The advent of cordeau, later followed by primacord, corrected this particular difficulty and therefore reduced the possibility of missed holes. After much experimentation, the blasting practice evolved into single row, multihole shots, with the holes spaced 4.5 to 5 m center to center in a row 7.5 to 8 m back from the toe. Such shots were fired from either end .by electric blasting caps attached to the main trunk lines of cordeau or primacord. The detonating speed of cordeau or primacord gave the practical effect of firing all holes instantaneously. Double row and multirow blasts, fired instantaneously with cordeau or primacord, proved to be unsatisfactory in the type of rock found at Chuquica-

Jan 1, 1952

By H. Becke, D. Stolnitz, D. Flatley, W. Kern

Extensive solid solubilities of CdTe (zincblende-type struckre) and InTe (B37 type) in each of the rock salt-type compounds, PbTe and SnTe, have been observed. Partial phase diagrams have been determined by thermal analysis and X-ray metallography. The limiting mol fraction. X,, of the solute in the rock salt-type a phase and the corresponding eutectic temperatures, T,, are: (PbTe)l-x(CdTe)x: X,- 0.2, T, =866°C; (PbTe),-,(lnTe),: X, - 0.35, T, = 646"C; (SnTe),-,(CdTe),:X,- 0.11, T, = 784 "C; (SnTe),-,(lnTe),: X,-0.53, T, = 630°C. The lattice parameters of the a phase decrease linearly with X, even in (PbTe),-,(CdTe),, where a, (CdTe) = 6.481A > %(PbTe) = 6.459A. This is taken as proof that the cadmium atom enters an octahedral interstice of the tellurium atom sublattice; i.e., the formation of the a phase entails the direct replacement of lead by cadmium. The a, us X curve extrapolates to 6.16A at X = 1, in agreement with the value predicted for an ionic crystal of CdTe; it is also consistent with the reported lattice parameter It is well-established that the compositions of intrinsic semiconducting and semimetallic compounds conform to normal valence rules.1"5 The apparent exceptions can be explained by taking account of anion-anion covalences, as in CdSb, or of multiple cation valences, as in In Te.'This empirical generalization is the basis of the chemical approach to semiconductors2 by which the properties of an intrinsic semiconductor are rationalized in terms of the ionic-covalent bonds necessary to saturate the valence-electron complement of the anion sublattice. A fundamental shortcoming of this approach is its disregard of long-range crystal interactions. It cannot, accordingly, deal with the phenomenon of charge conduction except through the use of ad hoc postulates. As an example, a crystal of CdTe would, in the valence-bond picture: be described in terms of electron pairs, each with the same discrete energy, localized between every cadmium and tellurium atom. This is, of course, contrary to the Pauli of the high-pressure rock-salt form of CdTe, corrected for decompression from 36 kbar. The free energy of formation of the a phase of pure CdTe at room temperature is calculated from the phase diagram to be +10 kcal per g-atom, in accmd with the value calculated from the transformation pressure. The standard enthalpy is +13 kcal per g-atom, and the standard entropy is +9 eu. The latter value implies the formation of extra classical particles, such as vacancies, interstitials, or nondegenerate charge carriers, but these alternatives are not consistent with the semiconducting properties and the densities of the a phase. The extrapolated values of the lattice parameters of (PbTe),-,(lnTe), and Spectively, for the rock salt-type modification of InTe. The corresponding interatomic separation is intermediate between monovalent and trivalent indium. The qualitative implications of the results are considered from the viewpoints of both valence-band theory and energy-band theory. principle, and is corrected by methods of molecular-orbital theory in which bonds are replaced by bands of molecular orbitals whose energies, E, depend upon a crystal-momentum vector, k.7 Each band is derived from a linear combination of atomic orbitals with two electrons required to saturate each molecular orbital in the band. In a crystal of N atoms containing z atoms in its primitive unit cell, there are 2 N/z states per band which lie in a region of k space bound by the Brillouin zone. In an intrinsic semiconductor, the bands can be grouped into valence bands which are normally filled and conduction bands which are normally empty. If the highest valence band and the lowest conduction band overlap anywhere within the Brillouin zone, the material is rendered semimetallic. The quantity defines the effective mass, m*, of an electron in a conduction band (or of a hole, i.e., an electron deficiency, in the valence band) and leads directly to the concept of charge mobility through the relation \x = et/m*, where e is the electron charge and t is the lattice relaxation time.

Jan 1, 1964

By G. J. Dickman, W. N. Thomas

The factors involved in selecting, a drive system for a grinding mill, including gearing arrangement, motor selection, and electric-supply system limitations are reviewed. Equipment costs are evaluated on a dollar per horsepower basis and starting performance is analyzed, the latter making use of a recently developed computer program. The present interest in large-diameter grinding mills and the rapidly rising cost of the drive in proportion to the cost of the mill establishes a need to explore the drive system in detail. The drive for a 7 to 8,000 hp mill will cost from $350 to $600,000 which represents nearly half the total cost of the mill package, and requires as much building space as the mill proper. Some of the drive components require a longer time to manufacture and the installation time and costs are equal to those for the mill. It is apparent that the drive has become a dominant factor in the overall mill design consideration. A new plant requires a tremendous investment in both time and money. With the high degree of competition in the mining industry, low capital cost is an exceedingly important factor in the development of a project. It is just as important, however, to insure that the equipment is going to perform satisfactorily for the anticipated life of the project without adding unplanned operating, maintenance, and replacement costs. This is the position in which the engineer or designer usually finds himself and is really one of his primary functionsestablishing a realistic compromise between cost and performance. Large primary grinding mills can represent a major portion of the cost of a plant and the drive system accounts for a substantial part of this cost. These facts justify a careful analysis of the drive requirements and all of the factors involved in ultimately arriving at the best system for the lowest cost. A number of excellent papers have been published on various aspects of mill drives. The authors have used information from these papers in making a study of a drive for a 7600-hp, 32 x 12-ft autogenous mill. This was a broad investigation, including mechanical and electrical design, layout, performance, and maintenance considerations, and overall economics leading to the selection of the optimum system for that mill. This paper was developed from the study and includes information applicable to large mills in general. For a successful and economical application, the drive must be considered as a system and all of the components analyzed to determine their effect on the system. This requires the participation of the mill designer, gear, and other mechanical component manufacturers, motor manufacturer, plant designer, and operator. A careful analysis of the mill torque requirements is first necessary to provide a basis for selection of the mechanical and electrical drive components. Because of manufacturing limitations on the size and capacity of gearing available for these low-speed applications, an evaluation of costs for various gear ratios and configurations in conjunction with motor speeds is required. Consultation with the power-supply company is necessary to determine limitations on motor-starting disturbances and to analyze operating costs on the basis of existing rate structures and low characteristics. Finally, consideration must be given to the drive arrangement and its affect on plant layout, space requirements, and location of auxiliary equipment for the selected process. There are presently three basic schemes for driving a mill. The first is a direct drive by a motor operating at mill speed with no gearing. The second is by a higher-speed motor through a trunnion-connected gear drive. The third is by a motor through the more conventional pinion and ring-gear with the option of an interposing gear box to permit greater motor speed than pinion shaft speed. [(Fig. 1.)] Several manufacturers have or are presently developing motors which drive the mill directly without the use of gears. One arrangement has the motor armature intimately surrounding the mill shell-discharge spout or extension of the trunnion journal. The motor stator is then supported independently and surrounds the shell or trunnion-mounted rotor. Another variation utilizes an independent mill-speed motor connected to the trunnion by a drive shaft and flexible couplings.

Jan 1, 1972

By J. S. Aronofsky

The first step in a quantitative analysis ot the mechanism of oil displacement by water in a fractured reservoir is usually conceded to be the solution of the differential equation describing the saturation distribution of two immiscible fluids flowing in a porous medium, where the capillary pressure is taken into account. In such a system the production mechanism may consist of displacement of oil both by the flow of water due to natural or artificially imposed pressure gradients and by imbibition, which implies a flow of water not due to external pressure gradients. Owing to the presence of the two oil displacement mechanisms, the mathematical model given by the differential equation intended to describe the system may not properly represent the behavior of the physical system. In fact, in the reservoir the rate of water advance may be very slow, and in the case of a fractured reservoir with a great number of large fractures, the pressure difference determining the flow of water through the matrix may be much less than 1 lb/'psi over lengths of a few feet. In such a case, imbi-bition (the exchange between oil in the matrix and water in the fractures resulting from capillary forces) may become, with time. a significant element of the production mechanism. It occurred Lo the authors, however, that without going into a physical analysis of the process of production, it might be possible by means of simple abstract reasoning to throw some light on the variation of recovery with time under conditions occurring in a highly fractured oil reservoir with rising water table. The object of this paper is to present both the reasoning and its application to a reservoir of the highly fractured type. Specifically, the analysis given here was undertaken to try to explain the increase of recovery (as defined later) with time as observed in this reservoir, without having to assume unlikely variations in the reservoir parameters with depth. This attempt has been successful as will become clear upon comparison of the computed recoveries with the actual field data. ABSTRACT MODEL Let us consider a small volume. of porous matrix saturated with oil at time, t = 0. Let the process of oil displacement by water start at time, t = t,,. At some time, t, the process will have terminated. Then a volume of oil equal to or smaller than the original oil contained in the matrix will have been produced. The first basic assumption that describes the model and guides the forthcoming reasoning is that the oil production from the small volume. dv, is a continuous monotonic function of time and that it converges to a finite limit. Such an assumption is not inconsistent with the results of laboratory waterflood tests as well as with results of imbibition tests where this is, in fact, observed. The second basic assumption is that none of the properties which determine the rate of convergence change sufficiently during the process to affect this rate or the limit. Let it be assumed that the form of the function of time relative to production from the matrix volume. dv, is given by where V,(t) is the volume of oil produced up to that time t. R is the limit toward which the recovery converges, A is a constant giving the rate of convergence, and V,(t) is the volume of oil originally in place in the volume, dv. It should be noted that recovery at time I will be understood here to he It follows that r. tends to R as t tends to infinity. CONSTRUCTION OF RESERVOIR FROM ABSTRACT MODEL Let the reservoir consist of a series of identical blocks of porous matrix stacked vertically and separated by fractures. Let each of these blocks satisfy the conditions of our abstract model. These conditions are: (1) that recovery is a continuous mono-tonic function of time converging to finite limit, and (2) none of the properties that determine the rate of convergence change sufficiently to affect the rate or the limit. Let water be rising in fractures so that oil production from any part of the block starts when the water comes in conLact with it. TOTAL RECOVERY FROM THE RESERVOIR COMPARED TO RECOVERY FROM A SINGLE MODEL ELEMENT As stipulated in Eq. 1, in the case of the abstract model,

By B. R. Queneau

Steel has been chosen as the metal whose solidification will be used to tie in the principles discussed in the previous papers. Although steel is the most important [ ] practical example that could be chosen, its solidification is complicated by the presence of many elements added either intentionally or present as impurities. The liquid steel bath in an open-hearth furnace contains carbon, manganese, phosphorus, sulphur, and many residual alloying elements such as nickel, copper, and molybdenum. The bath also contains oxygen, the concentration of which is a function of the carbon content, as shown in Fig. I. Deoxidizers such as silicon and aluminum may be added either before or after tapping depending on the grade of steel being made. There are four types of steel ingots: Killed, semikilled, capped, and rimmed, and these differ from each other in their state of oxidation. Each type of steel has advantages in the production of specific steel products, and it should be emphasized that no one type should be considered superior to the others. Fully deoxidized steels, known as killed steels, have little or no gas evolution on solidification. When the steel solidifies in the mold, shrinkage occurs which causes a large void known as "pipe." To minimize the amount of metal that has to be discarded on account of pipe, a big-end-up mold is used together with a refractory "hot top" which supplies molten steel to the main body of the ingot while solidification proceeds, Fig. 2. A section through a 32x32 in. ingot is shown in Fig. 3. The "hot top" volume is about 15 pct of the ingot, and the yield from killed steel in billet form is about 8o pct of the ingot weight. High quality machinery and tool steels are rolled from killed-steel ingots, but at the present time this represents less than 20 pct of the total steel production in the United States. In order to reduce cost and to increase

Jan 1, 1951

By John Henderson

Densities of melts 0f the lime-iron oxide-silica system in contact with solid iron have been measured by the maximum bubble pressure method in the temperature range 1250° to 1440°C and the composition range 0 to 40 mol pct lime, 15 to 100 mol pct iron oxide, and 0 to 55 mol pct silica. Densities range from 4.65 g cm 3 for wustite at 1440°C to 2.75 g cm-&apos; at 1350°C for a melt containing 30 mol pct lime, 20 mol pct iron oxide, and 50 mol pct silica. The results are interpreted in terms of a postulate that the melts can be regarded as a random array of oxygen ions in which regions of local order exist to satisfy the coordination requirements 0.f the cations. An understanding of the nature of metallurgical slags is basic to the development of a sound theoretical description of heavy metallurgical extractive and refining processes. Because these liquids are complex, direct measurements of their properties has not thrown much light on their structure. This has led to the approach of measuring the properties of simpler liquids, and building up their complexity until slag compositions are reached. In this way the density of liquid iron silicates was measured in a previous study1 and the present work represents a further stage in this synthesis. EXPERIMENTAL The technique used in the measurement of density was the maximum bubble pressure method. Details of the apparatus and procedure were similar to those previously reported,&apos; with the exception that a constant voltage transformer was used to supply the power input to the furnace and six silicon carbide resistance elements were used in place of the molybdenum winding. With these modifications melt temperature could be maintained within 1 centigrade degree during the course of a run. The silica used to prepare the melts was washed natural quartz ignited at 1000°C; wustite was prepared by air-melting A.R. grade ferric oxide in an iron crucible and lime was prepared by air ignition, at 1000°C, of weighed quantities of A.R. grade calcium carbonate, previously air-dried at 110°C. The finely ground constituents were intimately mixed in a glass ball mill prior to melting. Temperatures quoted are accurate to * 5°C and the standard deviation of the density values, calculated by the method of least squares, ranged from 0.5 to 1.8 pet. However, replicate determinations of density on different melts of the same nominal composition at the same nominal temperature did not vary by more than 1 pct, Table I, and this figure has been taken as an estimate of the accuracy of the density results. The density of carbon tetra-chloride was also measured as a check on the absolute performance of the experimental method. At 20°C a value of 1.593 * 0.002 g cm"3 was obtained; this compares with the literature value2 of 1.595 g cm"3. Results of experiments designed to measure the dependence of the density of lime-iron oxide-silica melts, in contact with solid iron, on composition and temperature are shown in Table I. Because iron sometimes precipitated in the sample during quenching, the Fe203 chemical analyses were only poorly reproducible and should be taken as a guide rather than as absolute values. Fig. 1 shows the data from various sources for the density of liquid iron silicates and Fig. 2 shows isodensity contours at 1410°C for lime-iron oxide-silica melts, calculated by graphical interpolation of smoothed curves drawn through the experimental results, together with the 1400°C results of Adachi and ogino3 and Pope1 and Esin.4 Fig. 3 shows the isothermal variation with composition of the volume of melt per gram ion of oxygen at 1410°C and Fig. 4 shows regions in which the temperature coefficient of this volume is negative, positive, or negligible (<0.005 cm3 deg-I). DISCUSSION a) Disparity Between Reported Density Results. Consider the system iron oxide-silica, the results for which are summarized in Fig. 1. Although there is some difference in the temperatures at which the various densities apply, this difference is not sufficiently large to account for the observed discrepancies. The reliability of the present results for the low-silica region has been confirmed by measurement of the density of liquid wustite by three different techniques. At 1410°C the density measured by a balanced-column method was 4.55 g cm"3, by a combination balanced-column and gas-densitometer method 4.59 g emd3, and by a pycnometer method 4.53 g cm"3. Schenck, Frohberg, and Hoffermann&apos; have also reported a value of 4.55 g cm"3 for the density of liquid wustite at 1400°C. It must be concluded, therefore, that neither Pope1

Jan 1, 1964

By J. P. Hammond, C. J. McHargue

The wire textures for cold rolled and recrystallized iodide titanium and the sheet textures for this material produced by cold and hot rolling, and recrystallization at a series of temperatures were determined. &apos;The effect of the a + ß transformation on the sheet texture was noted. UNTIL recently it was believed that all hexagonal close-packed metals deformed by slip on the basal plane, (0001), and that rolling should tend to rotate this slip plane into the plane of the rolled sheet. The pole figures of cold rolled magnesium&apos; are satisfactorily explained on this basis. There is a tendency for the <1120> directions to align parallel to the rolling direction, and the principal scatter is in the rolling direction. Zinc% as a rolling texture in which the hexagonal axis is inclined 20" to 25" toward the rolling direction. Twinning is believed to account for the moving of the basal plane away from parallelism with the rolling plane. The texture of beryllium3 places the basal plane parallel to the rolling plane with the [1010] direction parallel to the rolling direction, and the scatter from this orientation is primarily in the transverse direction. Cold rolled textures reported for zirconium&apos; and titanium5 how the [1010] directions to lie parallel to the rolling direction and the (0001) plane tilted by approximately 25" to 30" to the rolling plane in the transverse direction. Rosi has recently reported that the mechanisms for deformation in titanium are distinctly different from those commonly reported for hexagonal close-packed metals. The principal slip plane is the prismatic plane, {1010), with some slip also occurring on the pyramidal planes, (1011). However, there is no evidence for basal slip. The slip direction is reported to be the close-packed digonal axis, [1120]. In addition to the twin plane commonly reported for metals of this class, {1012), Rosi found the twin planes (1122) and {1121), with the dominant twin plane being (1121). Information regarding the recrystallization and hot rolling textures of hexagonal close-packed metals is limited. Barrett and Smigelskas report that rolling beryllium at temperatures up to 800°C and recrystallization at 700°C produce textures not differing from the cold rolled sheet texture.3 McGeary and Lustman find that hot rolling at 850°C produces the same basic texture in zirconium as rolling at room temperature.&apos; These investigators also report that the texture for sheet zirconium recrystallized at 650 °C differs from the cold rolled orientation inasmuch as the [1120] direction, instead of the [1010] direction, is parallel to the rolling direction. In the case of titanium, it is not possible to deduce which direction is preferred in the recrystallized state from the pole figures presented by Clark." The purpose of this paper is to report an extensive investigation of the preferred orientations in iodide titanium. Since the deformation mechanisms for titanium are different from those commonly given for hexagonal close-packed metals, it is not surprising to find distinct differences between the textures of titanium and other metals of this class. Materials and Methods This investigation was carried out on iodide titanium obtained from the New Jersey Zinc Co. with an analysis as follows: N2, 0.002 pct; Mn, 0.004; Fe, 0.0065; A1, 0.0065; Pb, 0.0025; Cu, 0.01; Sn, 0.002; and Ti, remainder. The crystallities of titanium were broken from the as-deposited bar and melted to form 20 g buttons on a water-cooled copper block in a vacuum arc-furnace. Hardness tests conducted on the material before and after melting differed by only two or three Vickers Pyramid Numbers, indicating no or insignificant contamination. The buttons were hot forged, ground, and etched to sizes and shapes suitable for the rolling schedule, and vacuum annealed at 1300°F. Specimens for determination of the wire textures were reduced 91 pct in diameter to 0.027 in. in 24 steps using grooved rolls. In order for the orientation of the central region to be studied, portions of these wires were electrolytically reduced to a diameter of 0.005 in. using the procedure described by Sutcliffe and Reynolds.&apos; The sheet textures were determined on titanium cold rolled 97 pct to a thickness of 0.005 in. A reduction of approximately 10 pct per pass was used, and the rolling direction was changed 180" after each pass. Specimens used for determination of the recrystallized textures were annealed in evacuated quartz tubes at 1000°, 1300°, and 1500°F. The grain size of the 1000°F specimen was sufficiently small to give satisfactory X-ray patterns with the specimen stationary. However, it was necessary to scan the surface of the other recrystallized specimens. The microstructure of each annealed specimen was that of a recrystallized material. The diffraction rings all showed the break-up into spots typical of recrystallized structures.

Jan 1, 1954

By N. A. Gokcen, J. Chipman

Aluminum and oxygen dissolved in liquid iron were brought into equilibrium with pure alumina crucibles and atmospheres of known H2O and H2 contents to study the reactions: 1—Al2O3(s) = 2 Al + 3 0; 2—Al2o3(s) + 3H2(g) = 2Al+ 3H2o(g); and 3—H2(9) +O = H2O(g). Aluminum strongly reduces the activity coefficient of oxygen and similarly oxygen reduces that of aluminum. Values of the product [% All" • [% O]3 are much smaller than those found in previous experimental studies and are of the order of magnitude of the calculated values. ALUMINUM is the strongest deoxidizer commonly A used in steelmaking, but the extent to which it removes dissolved oxygen has been debatable. The relationship between aluminum and oxygen has not been determined reliably not only on account of the usual experimental difficulties at high temperatures but also because of uncertainties in the analyses of very small concentrations of oxygen and aluminum. The earliest experimental attempt of Herty and coworkers&apos; was followed by a more systematic study of Wentrup and Hieber.&apos; These authors added aluminum to liquid iron of high oxygen content in an induction furnace and considered that 10 min was sufficient to remove the deoxidation products from the melt. Parts of the melts thus obtained were poured into a copper mold and analyzed for total aluminum and oxygen (soluble plus insoluble forms), assuming that the insoluble parts were in solution at the temperatures from which samples were taken. It is conceivable that the furnace atmosphere in their experiments, consisting of mainly air at 20 mm Hg pressure, was a serious source of continuous oxidation and therefore that their oxygen concentrations were correspondingly high. Scattering of their data was explained to be well within the maximum inaccuracy of 10°C in the temperature measurements and errors of ±0.002 pct each in the oxygen and total aluminum analyses. Maximum and minimum deoxidation values, i.e., values of the product [% All&apos; . [% O] differed by factors of 10 to 15; mean values of 9x10-11 and 7.5x10-9 ere reported at 1600" and 1700°C, respectively. Hilty and Craftsv determined the solubility of oxygen in liquid iron containing aluminum, using a rotating induction furnace. Pure alumina crucibles used in their experiments contained the liquid iron which in turn acted as a container for slags of varying compositions consisting mainly of Al2O3, Fe2O3, and FeO. The furnace was continuously flushed with argon, and additions of aluminum and Fe2O3 were made in the course of each experimental heat. The inner surfaces of their alumina crucibles were covered with a substance other than pure Al2O3, containing both iron oxide and alumina. Although frequent slag additions can change the composition of slag in the liquid iron cup formed by rotation, the inner surface of the crucible must depend upon the transfer of oxygen or aluminum through the liquid iron for any adjustment in composition. It is not clear that their metal was in equilibrium with the crucible wall, but it is clear that it was not in equilibrium with Al2O3. Their deoxidation product, [% A].]" • [% O]3, varied by a factor of more than 50; the average values of 2.8x10- and 1.0x10-7 were selected for temperatures of 1600" and 1700°C, respectively. Aside from the experimental determinations, attempts have been made to calculate the deoxidation constant for aluminum indirectly from thermody-namic data. Schenck4 combined the thermodynamic data for Al2O3 and dissolved oxygen in liquid iron by assuming an ideal solution. His calculated values are 2.0x10-15 and 3.2x10-13 at 1600" and 1700°C, respectively. Later, Chipman5 attempted to correct for the deviation from ideality and derived an expression which led to deoxidation values of 2.0x10-14 and 1.1x10-12 at 1600" and 1700°C, respectively. The errors in these treatments originate mainly from inaccuracies of thermal data and uncertainties regarding the activity coefficients of dissolved oxygen and aluminum. The purpose of this investigation was to study the equilibria represented in the following reactions in the presence of pure alumina: Al2O3(s) = 2Al + 3O K = aAl2.ao3 [1] Al2O3(s) + 3H2(g) = 2Al + 3H2O(g) H2O K2 = aAl2(H2O/H2 ) [2] H2(g) +O = H2O(g) K3 = 1/ao (H2) [3] The experimental method consisted of melting pure electrolytic iron, usually with an initial charge of aluminum, in pure dense alumina crucibles under a controlled atmosphere of H,O and H2 and holding

Jan 1, 1954

By K. D. Sheffler, R. W. Hertzberg, R. W. Kraft

The Ni-Ni,Ti lamellar eutectic alloy responds to unidirectional solidification by alig-nment of the platelets of the two phases roilghly perpendicular to the moving solid-liquid interface. X-ray diffraction and zetallographic experimenls show that the following cryslallographic relationships exist in the directionally solidified ingots: lamellar interface w [0001] Ni3Ti 11 (111 }Ni local growth direction (1120) Ni3Ti n (110) Ni These relationships are rationalized on the basis of the crystallography of the constituent phases. In recent years researchers in several laboratories have shown that an aligned or "controlled" micro-structure can be produced in many binary eutectic alloys by forcing the alloy to solidify unidirectionally in an elongated crucible.&apos;-3 It also has been found that a unique crystallographic relationship often develops between the two phases in such a structure.4, 5 Since a parallel arrangement of phase particles in an alloy should, and, in many cases, does, produce unique and highly anisotropic properties, the alloys so produced are of potential interest in a variety of applications. Studies of these alloys are also of interest at the academic level because they help to improve our basic understanding of the way in which metals solidify. This paper reports the metallographic and crystallographic structure of another of these alloys—that between nickel and Ni3Ti. In a separate paper, hot tensile and stress rupture properties as well as deformation and fracture behavior of the same alloy are discussed.19 EXPERIMENTAL DETAILS AND RESULTS General Characteristics of the Eutectic Reaction. Several small exploratory Ni-Ti melts were made to resolve a conflict in Hansen6 over the eutectic composition. The alloy containing 16.4 at. pct Ti, Fig. 1, appeared metallographically to be closest to the eutectic composition, although small dendritic patches of proeutectic nickel are present. The eutectic morphology is lamellar and appears quite sharp and angular, indicating a strong influence of crystallographic factors upon the microstructure. Some of the eutectic nickel platelets are continuous with the primary phase, whereas in hyper eutectic melts a "halo" of nickel generally surrounded the proeutectic inter-metallic platelets, indicating that nickel is the nucleating phase in this eutectic reaction. Extensive precipitation of Ni3Ti has occurred in the primary nickel phase in a Widmanstatten pattern which makes the dendrites difficult to differentiate from the eutectic mixture. This phenomenon may account for discrepancies among the eutectic compositions reported in the literature. The sloping solvus line bounding the nickel primary solid solution explains the presence of the intermetallic precipitate. Poole and Hume-Rothery7 report that the solubility of titanium in nickel drops from 13.8 at. pct at the eutectic temperature (1304°C) to approximately 9 at. pct at 700°C (Ni3Ti, on the other hand, melts con-gruently and exhibits essentially no solubility for nickel or titanium). Lever rule calculations based upon Poole and Hume-Rothery&apos;s phase diagram indicate an increase from 21 vol pct Ni3Ti in the eutectic mixture at the eutectic temperature to 61 pct at room temperature. Selective point count measurements of the volume fraction of eutectic Ni3Ti platelets in the eutectic microstructure yielded a value of 46 i 2 vol pct, which required extensive precipitation of Ni3Ti from the nickel solid solution onto the eutectic intermetallic platelets during cooling from the solidification temperature. Influence of Directional Solidification. Nickel bar and titanium sponge of 99.9 pct purity were induction-melted in zirconia crucibles under argon in the proportion of 13.8 wt pct (16.4 at. pct) Ti to produce 14-lb master heats which were cast into investment molds yielding 5/8-in.-diam pins approximately 10 in. long. These pins were remelted under argon in horizontal zirconia boats and unidirectionally solidified to produce 1/2-in.-sq ingots approximately 44 in. long. Details of the experimental techniques and apparatus are described elsewhere.&apos;

Jan 1, 1970

By L. W. Holm, A. K. Csaszar

A series of laboratory experiments was conducted in which oil was displaced from a porous medium by water-driven slugs of alcohols or similar solvents. The solvents used were soluble to some degree in both oil and water and covered the range of solubilities from complete solubility in oil to complete solubility in water. Displacement experiments were conducted on 2-and 3%-in. in diameter consolidated cores and 1-and 2-in. in diameter unconsolidated sand packs. The cores and sand packs ranged in length from 1 to 30 ft, and they were saturated with brine and crude or refined oils. The solvents used included ethyl alcohol, isopropyl alcohol (IPA), tertiary butyl alcohol (TBAI, secondary butyl alcohol (SBA), n-amyl alcohol (NAA), methyl ethyl ketone, acetone and others. It was found that all of the oil present in a porous medium could be miscibly displaced by injecting a slug of mutual solvent and driving it with water. The oil-recovery efficiency was dependent upon (1) the relative solubilities of the solvent in oil and water, and (2) the distance traversed by the flood. For complete oil recovery from cores, a smaller amount of a preferentially oil-soluble solvent was required, compared to the amount of preferentially water-soluble solvent needed. The size of the solvent slug required varied inversely with the linear flooding-path length raised to the 0.65 power. Water-driven dual solvent combinations (an oil-soluble solvent slug followed by a water-soluble solvent slug) were found to effect complete oil recovery with less total solvent than any single solvent used. In these dual-solvent displacement experiments, the slug size required varied inversely with length raised to the 0.55 power. Based upon the experimental results, a theory was developed to describe the displacement of oil and water by mutual solvents, and equations are presented to predict the production history in a linear system. These equations take into account the properties of the solvents and the porous medium. INTRODUCTION Oil-recovery processes which utilize displacing fluids that are miscible with the reservoir fluids have been studied extensively in recent years. Because of the poor contact efficiency and high pressure requirements of the LPG-gas displacement process there has been considerable interest in the alcohol-water process, and a number of studies have been made on the recovery of oil through the use of solvents which are mutually soluble in oil and water. An investigation by Sievert, Dew and Conleyl indicated that the use of mutual solvents would be limited because the presence of water in a porous medium would cause a phase break in the leading edge of the displacing solvent. Their study also showed that, in consolidated cores containing. only oil, the displacement of oil by a water-driven mutual-solvent slug of tertiary butyl alcohol (TBA) was affected by the viscosity ratios of the fluids involved. Gatlin and slobod3 concluded that an isopropyl alcohol (IPA) slug acts as a miscible piston, completely displacing both oil and water until the alcohol content of the mixing zone falls below the concentration necessary to maintain miscibility. Their study was conducted on uniform unconsolidated sand packs. They concluded further that IPA could be used effectively to recover oil from a watered-out sand. In a paper by Taber, Gamath and Reed4 relating an investigation on sandstone cores, it was stated that the displacement of oil by mutual solvents, particularly IPA, was not a miscible displacement and that no improvement in efficiency could be expected with increase in flooding-path length. However, their analytical analysis of the displacement mechanism using TBA is, in fact, one which indicates that the displacement is controlled by miscible mixing. They suggested that the lack of improvement in efficiency with flooding - path