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By L. J. Kashar

By means of a continuum-mechanics approach, the following equation for calculuting the pressure for axisymmetric cold forging has been derived: where P is the extrusion pressure; K and n are the strength coefficient and the strain-hardening exponent of the steel, respectively; a and b are strain factors, primarily dependent on the die design; and R is the extrusion ratio. Data from compression and backward-extrusion tests performed at a crosshead speed of 0.5 in. per min on a nonwork-hardening material pure lead) and on a work-hardening material (1018 steel) indicate that the predicted extrusion pressures obtained from the derived equation agree very closely with the actual extrusion pressures. Therefore, the equation can be used to predict backward-extrusion pressures from standard mechanical-test data. The decision to cold-forge a part is based on several considerations, most of them related to economics. However, the production feasibility is usually determined by the pressure needed to extrude the part and the availability of tooling that can repeatedly sustain this pressure for over several hundred thousand cycles. A major contribution to the cold forging of steel would result from the development of steels that could be forged at reduced pressures and from the development of techniques that permit the accurate prediction of the pressure necessary to cold-forge the part. The ability to predict extrusion pressures is necessary to properly design the tooling required for the cold-forging process. Attempts to theoretically predict the cold-forging pressure, especially for work-hardening materials, have been unsuccessful in previous investigations of the cold-forging process.1-3 Many empirical relations have, therefore, been proposed relating the extrusion pressure to the chemical composition,' the extrusion ratio,2 the tensile strength,273 and hard- ness, 2 and the yield strength3 of the material being extruded. However, these empirical relations are valid only for the specific conditions used in their development, and, in some cases, the empirical relations had a scatter band greater than * 10 pct. A study of the extrusion process was undertaken with the objective of developing predictor equations for extrusion pressure by using a continuum-mechanics approach to the problem of metal flow. The derivation of such an equation is summarized in this paper, and data are presented that confirm the validity of the derived equation. DERIVATION OF THE EXTRUSION EQUATION If the extrusion of a material through a die is considered in terms of a work balance, it has been shown that it is possible to derive an equation relating the work necessary to move the material through the die to the deformation work undergone by the material. The work done, W, in moving the material is given by where F is the extrusion force, I, is the distance the material was moved, A, is the original cross-sectional area of the material, P is the extrusion pressure acting on A,, Ap is the cross-sectional area of the extrusion punch, and Pp is the punch pressure acting on Ap. The deformation work, WD, is given by where V is the volume of metal deformed, ay is the flow stress of the metal, and E is the total strain undergone in the extrusion. If there are no frictional losses, the work done, W, is equal to the deformation work, WD; therefore, Since the volume of a metal does not change during deformation,

Jan 1, 1968

By L. A. Wilson, N. H. Harrison, R. R. Clay, R. L. Reed. D. W. Reed

Experiments have been performed in a linear near-adiabatic system for the purpose of extending data on reverse and forward combustion from atmospheric pressure to 1,000 psig. Results obtained from reverse combustion appear to conform qualitatively with the existing description of the process when reaction kinetics are suitably modified to account for the increase in pressure, i.e., increasing tbe pressure decreases the peak temperature and increases the combustion zone velocity. Forward combustion appears to be a fuel dominated process wherein peak temperature and combustion zone velocity are not very sensitive to changes in pressure. The moderate effects of pressure that do exist at low flux virtually disappear at high flux providing all oxygen is consumed. With this provision. increasing the pressure decreases the frontal velocity and increases the peak temperature. Results me shown graphicaIIy which demonstrate the effects of presswe on peak temperature, rate of advance, oil recovery, air/oil ratio, carbon oxides produced and temperature distributions. INTRODUCTION Numerous field tests of oil recovery using the technique of underground combustion are now in progress. Operating pressures used are always substantially in excess of atmospheric pressure. Nevertheless, the literature contains no laboratory data pertaining to the effects of pressure except for those of Martin, Alexander and Dew.1 Unfortunately, these data appear to reflect heat losses which may have obscured some of the effects of increased pressure. The underground combustion processes are exceedingly complex, and general concepts relating to them have necessarily been described in relatively simple terms. Particularly with regard to forward combustion, the interaction between multiphase fluid flow, heat transport and various rate mechanisms is so involved that to date a rigorous treatment is not available. In this paper a linear near-adiabatic system is described which was found capable of operating at pressures up to 1,000 psig and temperatures of at least 1,100 F. Using this equipment, experiments on both forward and reverse combustion were accomplished at elevated pressures and the data presented as a function of air flux or temperature. It will be assumed that the reader is familiar with the papers of Wilson, Wygal, Reed, Gergins and Henderson on forward combustion and that of Reed, Reed and Tracht3 on reverse combustion, so that it is not necessary to repeat the details of the two processes. EQUIPMENT AND PROCEDURE It was decided to study the effects of pressure over the interval 0-1,000 psig using a near-adiabatic system. Since satisfactory operation of such a system requires very thin tubing walls, it proved necessary to insert the entire combustion tube and heater assembly into a high pressure jacket in such a way that the internal tube pressure could be counterbalanced by pressurizing the annulus. This meant that the numerous heater and thermocouple wires had to pass through gas-tight seals that required considerable effort to maintain effective. An unexpected difficulty arose in connection with the nature of the annulus pressuring gas. It turned out that the composition of this gas critically affected the functioning of the adiabatic control system. Operation of the equipment at elevated pressures was not particularly difficult. The problems that did arise were a direct consequence of the large numbers of connections and seals involved. EQUIPMENT The equipment consisted of the combustion tube which fitted within the pressure jacket. This jacket was in turn mounted on a wheeled carriage which facilitated movement and provided for rotation of the

By E. J. Roberts

F. C. Bond (Allis-Chalmers Mfg. Corp., Milwaukee) —This paper considers comminution as a first order process, with the reduction rate depending directly upon the amount of oversize material present. The data show that other factors should be taken into account, and it is possible that in time these may be evaluated as simultaneous or consecutive reactions: Development of the theory of comminution has been retarded for many years by the assumption that surface area measurements constitute the sine qua non of the work done in crushing and grinding, and it is encouraging to note the belated growth of other ideas. In the Abstract the term "net power" should be changed to "net energy." Throughout the paper the term "hp per ton" should be changed to "hp hrs per ton", or "hp hr t." The term "Probability Theory" in the title does not seem appropriate, since it is not clear how the probability theory is used in developing the ideas in the paper. There seems to be a contradiction between the large calculated advantages of closed circuit operation and the statement following that the closed circuit test results showed no significant change in grinding behavior, when compared with the batch grind curves. Tables I and II show that between 75 pct and 50 pct solids the energy input required decreases with increasing moisture content and may indicate the advisability of grinding at higher dilutions in certain cases. The calculation of the hp-hr per ton factor indicates an input in the laboratory mill of only 7.32 gross hp per ton of balls; this casts some doubt upon the accuracy of the factor used, since the power input in commercial mills at 80 pct critical speed is customarily much higher. The tests show that within fairly wide limits the amount of ore in the laboratory mill may be varied and a product of constant fineness obtained, provided that the grinding time is varied in the same proportion. This has often been assumed, and confirmation by actual testing is of value. The Cavg corrections for differences between the plant and laboratory size distributions do not seem very satisfactory, since in many cases the plant/laboratory ratio is farther from unity after correction than before. The following equation has been derived from the data in Table VI: Relative Energy (log new ball diam in in. + 0.410) Input = --------------—--------------- from which the relative energy inputs for balls of different sizes can be calculated and compared. The relative energy input is unity for balls of 2.715 in. diam. The equation indicates that the work accomplished by a ton of grinding balls per unit of energy input is roughly proportional to the square root of the total ball surface area; provided, of course, that the balls are sufficiently large to break the material. The data in support of this statement are admittedly meager, but are fairly consistent when plotted. The relative grindability values listed in Table VI for 200 mesh multiplied by 4/5 apparently correspond approximately to the A-C grindability at 200 mesh.' It would seem that for open circuit tests comparable accuracy could be obtained much more simply by the old method' of plotting the test grind, extending the mesh grinds to the left of zero time if necessary, and determining from the plot the equivalent time required to grind from the plant feed size to the plant product size, using the average of several mesh sizes. The en- ergy input value of one time interval could be determined by tests on materials of known grinding resistance, and this multiplied by the interval required should give the desired energy input value. The relative grindabilities would be the relative time intervals required for a specified feed and product size. When the plotted mesh size lines of a homogeneous material are extended to the left beyond zero time they meet at one point at zero pct passing. The horizontal distance of this point from zero time indicates the equivalent energy input required to prepare the mill feed. The author's results show that the closed circuit grinding tests give about the same K values as open circuit tests, from which he concludes that open circuit tests are satisfactory in many cases. The value of the closed circuit test is its ability accurately to predict energy requirements in closed circuit grinding for both homogeneous and heterogeneous materials. If the material is homogeneous, the open circuit test gives satisfactory results; but if the material contains appreciable fractions of hard and soft grinding ore, the open circuit tests will not be accurate because of the accumulation of hard grinding material in the circulating load. Since in most cases it is not possible to determine a priori whether the material contains hard and soft fractions, the closed circuit tests are preferable and more reliable. B. S. Crocker (Lake Shore Mines, Ontario)—Dr. Roberts probability theory of grinding is very similar to our log pct reduced vs. log tonnage method of plotting and evaluating grinding tests at Lake Shore. However, although we both seem to start at the same point we finish with different end results. Shortly after publishing our grinding paper (referred to by Dr. Roberts) in 1939, we did pursue the subject of the "constant pct reduction in the pct +28 micron material for each constant interval of time. We ran innumerable tonnage tests on the plant ball mills, rod mills, tube mills with 11/4 and 3/4 balls, and lastly pebble mills, with tonnage variations from 180 tons per day to 950 tons per day. We found that when we plotted the log of the tonnage against the log of the pct reduced of any reliable mesh, we had a straight line up until 90 pct of the mesh is reduced. We have also tested this in our 12-in. laboratory mill with the same results. We have used this method of evaluating grinds for the past 8 years and developed the recent four stage pebble plant on this basis. By pct reduced we mean the percentage of any given mesh that is reduced in one pass through a mill at a given tonnage (or time). For example, if the feed to a rod mill is 90 pct +35 mesh and the discharge at 500 tons per day is 54 pct +35, the pct reduced is 90 — 54/90 = 40 pct. If the feed had been 80 pct +35 the discharge would have been 48 pct +35 or pct re- duced 80-48/80 = 40 pct as long as the tonnage re- mained constant at 500 tons per day. Thus we can easily correct for normal variation of mill feeds. This log — log relationship derived from the tonnage tests of all our operating mills has proved of tremendous help in checking laboratory work and in designing alternate layouts or new plants. The difference between the log — log and the semi-log plot is only shown up when the extremes in tonnages are plotted. When the relationship between the pct reduced and the tonnage was first investigated, we used semilog

Jan 1, 1952

By L. H. Robinson

A proposed method of drilling utilizes sequential detonation of two types of explosive charges delivered to the hole by a conventional drilling fluid through pipe. A shaped charge first produces a long thin hole. A second charge, called a gauging charge, is pumped into the thin hole, explodes, and enlarges it to full gauge. In one suggested configuration, the explosives are contained in cylindrical capsules 4- to 5-ft long. A drilling head is positioned 2- to 3-ft off bottom. When a capsule is pumped to the bottom of the hole and stops, pressure inside the drill pipe increases and differential pressure detonates the explosive. Primacord along the length of the capsule disintegrates the capsule. The debris circulates out of the hole with the drilling fluid. The gauging-charge action was examined using 2-ft cubes of Berea sandstone. The sandstone (18 per cent porosity) was vacuum-saturated with water and surrounded by concrete. To prevent reflected tensile pulses, the acoustical impedance of the concrete matched that of the sandstone. A 750-psig pressure was applied to a drilling fluid within a shaped-charge-formed hole before detonating the gauging charge. In these tests, 190 gm of rock were removed for each gram of explosive detonated. A similar series of experiments on low-porosity (0.40 per cent) metamorphic limestone provided a more rigorous test. Extensive fracturing from the gauging-charge action prevented an estimate of the volume of the expanded hole. The diameter of the hole was irregular, ranging from about 2 to 4 in, or about fivefold larger than the hole created by the shaped charge. INTRODUCTION In the past, explosives have found a varied but somewhat limited use in the oil field. An early well stimulation technique, still in use today, in vo lves shooting ni troglycerin to en large the borehole and to increase matrix permeability near the wellbore. Bullets and jets are now the conventional method of perforating casing and tubing. Detonations are frequently used to "back off" or "shoot off" stuck pipe and, of course, seismo-graphic work has long depended on explosive-generated pulses to find attractive earth structures. Efforts to use explosives in drilling operations have received limited examinations1-3 which, in one respect, is surprising because explosives offer one way to transmit large amounts of energy directly to the bottom of the hole. Recently, Ostrovskii l described a Russian method of drilling with explosives that utilizes small capsules containing explosives which are pumped into the mud stream to the bottom of a borehole. As the capsules leave the end of the drill pipe, a detonation mechanism is armed. Impact detonates the explosive. In another configuration, two liquids (an oxidizing agent and fuel) in the capsule are separated by a frangible membrane. Separately, the liquids are nonexplosive. Near the bottom of the drill pipe, the capsule passes through a narrow constriction which ruptures the membrane. The mixing of the liquids produces an explosive. Impact is again used to detonate the explosive. About the time Ostrovskii was developing his method, Humble Oil & Refining Co. was independently doing research on another explosive drilling concept based on the sequential detonation of two types -of explosive charges. The proposed technique consisted of alternately pumping a shaped charge and then a "gauging charge" down a well. The shaped charge should drill a long, tapered hole, and the gauging charge should expand it to full gauge. This paper presents the visualized drilling concept and gives results of laboratory work which showed that sequential firing of charges can form hole at an attractive charge-to-rock removal ratio. An important finding was that increased environmental pressure, which is to be expected in wells, increases rock removal efficiency of the gauging charge.

Jan 1, 1966

By L. T. Sanchez, W. T. Rogers

This paper presents the results of a production experiment evaluating the effect of the use of oxygen in the bessemer converter with respect to its relation to blowing time, amount of steel scrap or cold pig iron which can be melted and teeming temperatures of the molten steel. THE simplest and probably the most efficient and economical means for producing steel from molten pig iron is the bessemer converter. This process is essentially a mechanism for bringing oxygen into contact with the molten metal thereby oxidizing the elements silicon, manganese, and carbon. The source of oxygen is that contained in the air and is limited by that amount. In theory, there is an intermediate chemical reaction due to the mass action of iron and air which is represented by the equation: 2Fe + O2- = 2Fe0 The FeO then becomes the oxidizing agent and reacts with the silicon, manganese, and carbon to produce silicon dioxide (SiO,), manganese oxide (MnO), and carbon monoxide (CO), in the order mentioned. The balancing of the various chemical reactions shows that in the ordinary bessemer process more than sufficient oxygen is delivered to satisfy these reactions, which indicates that a further increase in the oxygen content of the air would serve no useful purpose. In spite of this reasoning, however, it will be demonstrated that when melting stelel scrap, commercially pure industrial oxygen introduced into the normal air blast will permit the melting of a larger proportion of steel scrap than is possible with air alone and also will increase the production rate by a substantial reduction in blowing time. It will also be shown that when melting cold pig iron, it is possible to melt up to twice as much pig iron with the use of oxygen as can be melted without added oxygen with a nominal increase in blowing time. At the plant where this study was made, oxygen is purchased from a nearby supplier and piped directly into the plant. It is supplied to the bessemer converter by being introduced directly into the blast line through an automatic control system which may be adjusted by the bessemer operator or "blower." Normally, iron is supplied by five blast furnaces to an open hearth shop containing twelve 165-ton capacity stationary furnaces and a bessemer converting shop having three 30-ton converters. The molten metal from the blast furnaces is allocated to one shop or the other in a manner to best provide for maximum production. At various times, it is desirable to direct the maximum amount of available blast furnace production to the open hearth shop. This has the effect of reducing the production of the bessemer shop. When this condition exists, it is desirable to supplement the reduced iron charge at the bessemer with steel scrap or cold pig iron to the maximum possible extent in order to maintain high total ingot production. This condition becomes aggravated when one of the blast furnaces is taken out of production for any repair job and particularly for extended periods such as when relining is necessary. The purpose of the investigation and experimentation with oxygen was to determine whether or not it would contribute to increased flexibility in the use of the raw materials, molten iron, steel scrap, and cold pig iron. In addition to the effect of oxygen on flexibility of operation, it was also necessary to determine the relation of its use to production rate or "blowing time," final finishing temperature of the steel or "teeming temperature," and steel quality.

Jan 1, 1953

By Fred H. Poettmann

An equation has been derived for use in calculating the sandface pressure of flowing gas wells in which the variation of the compressibility factor of the gas with pressure is taken into consideration. This variation due to compressibility has been put into both graphical and tabular form. Comparison of calculated results with field measured results were made on 20 dry gas wells from a given field and 11 distillate wells from different fields. The agreement between calculated and observed results is good. In addition to their use in the calculation of sandface pressures of flowing wells, the factors can be used in the direct calculation of the static bottom hole pressure of gas wells, the capacity of gas transmission lines, and in the calculation of the theoretical isothermal horsepower necessary to compress a natural gas. Examples demonstrating the use of various equations are given. INTRODUCTION Many of the equations used to calculate the sandface pressures of flowing gas wells from well head data do not take into consideration rigorously the deviation of the natural gas from ideal gas behavior. For low pressure wells this error is not serious. For high pressure wells flowing at high rates this error call be serious. Sandface pressures are used to determine the producing capacity of gas wells as shown in the U. S. Bureau of Mines' Monograph 7.' An equation has been derived for use in calculating the sandface pressure of flowing gas wells in which the variation of compressibility of the gas with pressure is taken into consideration. This variation due to compressibility has been put into both graphical and tabular form. In addition to their use in the calculation of sandface pressures, these "factors" can also be used to give a direct solution in the calculation of the static bottom hole pressure of gas wells, the calculation of the capacity of gas transmission lines, and the calculation of the isothermal horsepower necessary to compress a gas. CALCULATION OF SANDFACE PRESSURE OF FLOWING GAS WELLS The starting point in the derivation of any specific flow equation is an energy balance on the fluid flowing between any two points in the system. This energy balance based on one pound of flowing fluid is expressed by the well known general flow equation: jp~- vdP+ Ah + -— + wt + r. = o . . (i) P, 2 g. where V = specific volume of flowing fluid in cu ft/lb P = pressure in Ib per sq ft ah = difference in height above datum plane in, ft ------ = kinetic energy change of one lb of flowing fluid, v equals velocity in ft/sec and g. = 32.174 Ws = Work done by fluid while in flow (similar to shaft work in driving a turbine) W, = Znergy losses due to the irreversibilities of the flowing fluid This equation contains no limiting assumptions and can be made the basis of any fluid flow relationship by limitations or substitutions. The energy units are usually expressed in term* of foot-pounds of energy per pound of mass. Considering the vertical flow of gas through tubing or annu-lus, the external work W, done by the gas is zero. Let us also assume that the value of the kinetic energy function is small and can be neglected. Equation (1) is reduced to: /P: VdP + Ah+Wt = 0 .......(2) Pi The energy loszes PI can be expressed in terms of the well known Fanning equatinn: 4/Lf: rf = —— ............(3) 2g,D where L = the integrated average velocity in ft/sec of vertical flow in a uniform flow string -~h = L v = the integrated average velocity in ft/sec / = dimensionless correlating function

Jan 1, 1951

By Fred H. Poettmann

An equation has been derived for use in calculating the sandface pressure of flowing gas wells in which the variation of the compressibility factor of the gas with pressure is taken into consideration. This variation due to compressibility has been put into both graphical and tabular form. Comparison of calculated results with field measured results were made on 20 dry gas wells from a given field and 11 distillate wells from different fields. The agreement between calculated and observed results is good. In addition to their use in the calculation of sandface pressures of flowing wells, the factors can be used in the direct calculation of the static bottom hole pressure of gas wells, the capacity of gas transmission lines, and in the calculation of the theoretical isothermal horsepower necessary to compress a natural gas. Examples demonstrating the use of various equations are given. INTRODUCTION Many of the equations used to calculate the sandface pressures of flowing gas wells from well head data do not take into consideration rigorously the deviation of the natural gas from ideal gas behavior. For low pressure wells this error is not serious. For high pressure wells flowing at high rates this error call be serious. Sandface pressures are used to determine the producing capacity of gas wells as shown in the U. S. Bureau of Mines' Monograph 7.' An equation has been derived for use in calculating the sandface pressure of flowing gas wells in which the variation of compressibility of the gas with pressure is taken into consideration. This variation due to compressibility has been put into both graphical and tabular form. In addition to their use in the calculation of sandface pressures, these "factors" can also be used to give a direct solution in the calculation of the static bottom hole pressure of gas wells, the calculation of the capacity of gas transmission lines, and the calculation of the isothermal horsepower necessary to compress a gas. CALCULATION OF SANDFACE PRESSURE OF FLOWING GAS WELLS The starting point in the derivation of any specific flow equation is an energy balance on the fluid flowing between any two points in the system. This energy balance based on one pound of flowing fluid is expressed by the well known general flow equation: jp~- vdP+ Ah + -— + wt + r. = o . . (i) P, 2 g. where V = specific volume of flowing fluid in cu ft/lb P = pressure in Ib per sq ft ah = difference in height above datum plane in, ft ------ = kinetic energy change of one lb of flowing fluid, v equals velocity in ft/sec and g. = 32.174 Ws = Work done by fluid while in flow (similar to shaft work in driving a turbine) W, = Znergy losses due to the irreversibilities of the flowing fluid This equation contains no limiting assumptions and can be made the basis of any fluid flow relationship by limitations or substitutions. The energy units are usually expressed in term* of foot-pounds of energy per pound of mass. Considering the vertical flow of gas through tubing or annu-lus, the external work W, done by the gas is zero. Let us also assume that the value of the kinetic energy function is small and can be neglected. Equation (1) is reduced to: /P: VdP + Ah+Wt = 0 .......(2) Pi The energy loszes PI can be expressed in terms of the well known Fanning equatinn: 4/Lf: rf = —— ............(3) 2g,D where L = the integrated average velocity in ft/sec of vertical flow in a uniform flow string -~h = L v = the integrated average velocity in ft/sec / = dimensionless correlating function

Jan 1, 1951

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

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

Jan 1, 1951

By J. F. Held

In a study of the influence of cold rolling on mechani cal properties of steel, it was found that the specific rolling procedure influenced crystallographic orientation. Properties developed by several rolling modes were determined and related to crystallographic orientation. Sanzples were cold-rolled various amounts on a laboratory mill. When palm oil was used as a lubricant, the crystallographic orientation was homogeneous throughout the thickness of the samples. Other samples were rolled without lubrication. Variations in intensity of the recrystallization texture components persisted throughout more than 50 pct of the volume of these samples. When the average plastic strain ratio, which is dependent upon crystallogvaphic orientation, was plotted as a function of percent reduction, a maximum was found at 40 pct in the samples which contained the observed texture variation. No maximum was found for the samples with the homogeneous texture distribution. Other mechanical properties, such as tensile strength and yield strength, were not influenced by the variation in crystallopaphic orientation. MECHANICAL properties of steel sheets depend, in general, upon the structural properties of the steel. For example, the Petch equation describes the relationship between yield strength and grain size. Other properties, such as hardness, vary with chemical composition. The plastic strain ratio, or R value, has been found to be dependent upon crystallographic orientation . In the course of investigating the effect of cold rolling on mechanical properties, it was found that a previously reported relationship between percent cold reduction and R value was not always observed. Published data had shown that the optimum cold reduction for best R values is 70 pct for aluminum killed steel.' However, by altering the rolling conditions, the optimum reduction can be shifted to much lower values, or it can be moved to very high values. This paper describes the rolling conditions that lead to shifting of the optimum cold reduction. The effect of rolling conditions on the various structure parameters and the resulting mechanical properties are evaluated. EXPERIMENTAL PROCEDURE Samples of three low-carbon aluminum killed steels were cold-rolled various amounts on a 9-in.-diam 2-high laboratory mill. Hot-mill thermal practice and steel chemistries are listed in Table I. Steel A was rolled with palm oil lubrication and heavy reductions on each pass; i.e., a total reduction of 75 pct was ac- complished with a maximum of five passes. Steel B was rolled without lubrication, and light reductions were given on each pass. After approximately 75 pct reduction, it was necessary to use a small amount of light oil to achieve any additional reduction in thickness. Following cold reduction, the samples were box-annealed at 1300°F for 10 hr. The heating rate between 800" and 1300°F was 50°F per hr, and the cooling rate following the soak period was 50°F per hr. Tensile specimens were machined from the annealed material and these were used to measure yield strength, tensile strength, and the plastic strain ratio (R value). Recrystallization textures were determined by an inverse pole figure technique.*'3 Integrated intensities of the (110), (002), (112), and (222) reflections were recorded for those planes lying parallel to the sheet surface. Examinations were made on both surfaces of each sample, and also at various levels beneath each surface to approximately 0.010 in. depth by etching techniques. RESULTS R Values. The R values obtained in steels A and B are shown as a function of percent cold reduction in Fig. 1. No optimum cold reduction was observed in steel A, rolled with palm oil lubrication. It is likely, however, that an optimum reduction would have occurred had higher reductions than 88 pct been employed. The optimum reduction in steel B occurred at 40 pct. The behavior in both of these steels is quite

Jan 1, 1968

By R. E. Cook

This paper presents results of a study to determine to what extent errors in estimated free gas saturation affect the results of static pressure calculations from build-up curves in two-phase systems. Use is made of the method of pressure build-up curve analysis developed by Miller, Dyes and Hutchinson1 , with modifications for two-phase flow advanced by Perrine2. Cornhined fluid compressibility of gas-oil mixtures is included, with special attention given to the discontinuity that occurs in fluid compressibility at the saturation pressure and its effect on static pressure calculations from build-up curves. The sensitivity of static pressure calculations to errors in gas saturation is considered over wide ranges of formation and fluid properties and reservoir pressure conditions. It is concluded from this study that for analyses of two-phase pressure build-up curves (I) the use of under-saturated crude compressibility data, even when producing GOR's are at or near solution values, will result in very significant negative errors in calculated static pressrrre, and (2) except for large values of buildup curve .slope and/or very low crude gravities, rather large errors in calculated gas saturation can he tolerated without .significant effect on calculated static pressures. The conc1usions arc qualified as referring only to the effects of gas saturation and excludc possible errors in the nuumerous other factors that enter into the ancllysis of pressure build-up curves. INTRODUCTION As part of the analysis of basic data for several major Eocene reservoirs in the Bolivar Coastal field (Western Venezuela), it was necessary to analyze a large number of pressure build-up curves for oil permeability and static reservoir pressure. From performance data it became apparent that gravity segregation is very active in these reservoirs. As a result of this performance it was recognized that the free gas saturation within the oil leg of the reservoirs would be small but would be subject to certain errors in calculation due to the inaccuracy of relative permeability data in the low gas saturation region. Since all of the various methods of build-up curve analysis were originally derived for sin- gle-phase, constant compressibility systems, and since the presence of a small but rather indeterminate gas saturation would greatly affect the reservoir fluid compressibility, a study was undertaken to investigate the effect of gas saturation on the results of two-phase pressure build-up analyses. TWO-PHASE BUILD-UP CURVE ANALYSIS The method of analysis used in this investigation is basically that originally developed by Miller, Dyes and Hutchinson1, with modifications for two-phase flow advanced by Perrine2. Miller, et al, showed that the original equation for single-phase permeability away from the well bore, held for two-phase flow except that the calculated permeability is the effective permeability to oil away from the wellbore. This finding was acknowledged by Per-rine, who also presented the equation for effective gas permeability* as Since Eqs. 1 and 2 contain no terms influenced by fluid compressibility, no distinct problem should develop in calculating permeability from two-phase build-up curves. In regard to the calculation of static reservoir pressure from build-up curves in single-phase systems, the modified Miller, et al, method as presented by Perrine is briefly as follows. 1. From build-up curve data, PVT data for the reservoir fluid and the well spacing and sand thickness, calculate the climensionless shut-in time corresponding to the highest value of actual shut-in time (t, hrs) on the straight-line portion of a semi-log build-up curve. Dimenrsionless shut-in time is related to actual shut-in time by the relation, t- 0.000264 kt 2. From a graph of dimension less shut-in time vs dimensionless pressure difference (Fig. 1 1 ) find the value of dimensionless pressure difference corresponding to the value of 7 calculated under (1) above. Dimen-

By J. Szekely, G. P. Martins

A formulation is given and a general analytical solution is presented for the problem of spherical phase growth in multicomponent systems. The results are valid for describing the rate of bubble growth in the absence of translational motion and, therefore, should be relevant to heterogeneous nucleation processes. Typical worked examples are given, which include bubble growth due to the simultaneous diffusion of hydrogen and nitrogen in molten steel, and the growth of CO bubbles in steel, due to the reaction: C + 0 = CO. The calculations predict that the presence of additional components would enhance the growth rate and also increase the absolute rate at which the individual components are transferred into the bubble. These findings should be relevant to degassing processes. THE problem of spherical bubble growth in the absence of translational motion is appropriate to the description of the heterogeneous nucleation and subsequent growth of gas bubbles while attached to a solid surface. Such problems are of considerable practical relevance to various vacuum degassing and steelmaking processes where bubble nucleation and subsequent growth play an important part in the overall reaction mechanism. A definitive formulation of spherical phase growth has been given in the now classical paper by Scriven1 and the analogous problem for solid phase growth has been discussed by Horvay,2 furthermore a thorough review of the relevant literature has also become available.3 All this previous work was concerned with one component or binary systems. The purpose of this paper is to show that the technique developed by Scriven may be generalized to alow the consideration of multicomponent systems. The results of this generalization would be particularly relevant, as gas bubble nucleation in the metallurgical systems previously mentioned frequently involves the simultaneous diffusion of several components. In the following a general formulation is given and typical worked examples are presented. FORMULATION Consider the growth of a spherical gas bubble due to diffusion in a quiescent liquid of infinite extent, as sketched in Fig. 1. The formulation of such problems requires the statement of the following conservation equations relating to the liquid phase: Equation of continuity Equation of motion Conservation of the diffusing species—Diffusion Equation. The following assumptions are made: a) The fluid is incompressible, quiescent, and of infinite extent b) The bubble is stationary, i.e., it undergoes no translational motion c) Spherical symmetry is observed throughout d) The mass diffusivity is constant e) Chemical equilibrium exists at the surface of the bubble, and any reactions occurring there are fast compared to the diffusion process f) Inertia and surface tension effects are neglected. Equations derived on the basis of these assumptions by previous investigators were found to provide a reasonable representation of bubble growth while attached to a surface in organic and aqueous systems4,5 For gas bubble growth in molten metal systems assumptions a) to d) chould hold reasonably well. The exclusion of chemical kinetics as a possible rate limiting step may not be universally valid, but should hold for many systems operating at high temperatures. The effect of hydrodynamic inertia may be neglected for moderate growth rates; finally, the assumption neglecting the effect of interfacial tension on the bubble growth was made for the sake of mathematical convenience; its inclusion would have required the use of quite cumbersome numerical techniques. While surface tension effects could be appreciable on the initial values of the overall growth rate, especially at low absolute pressures and in case of small bubbles, they are not thought to modify the relative effects present in multicomponent systems, the study of which is the objective of this paper. Within the framework of the above assumptions

Jan 1, 1970

By N. K. Chen, R. Maddin

In the extension of molybdenum single crystals at room temperature, the slip planes were found to be of the type (1101; the slip direction <111>. Theories of plasticity of body-centered cubic metals have been examined, and an explanation based upon planes of highest atomic density seems plausible to explain the plastic behavior of molybdenum single crystals. IN body-centered cubic metals, slip has been reported to occur on several planes; these are of the type {110), {112), and (123). Andrade has correlated&apos; the operative slip planes in body-centered cubic metals with the testing temperature relative to the melting point. In this correlation, if T is the absolute temperature at which slip takes place and Tm the absolute melting point, then the acting slip planes in body-centered cubic metals may be arranged as shown in Table I. However, deviations from this correlation have been reported in iron and iron-silicon alloys2 and also in molybdenum deformed at a temperature around 2400°C." In the case of iron and other body-centered cubic metals or alloys in which all three slip planes have been reported to be active at any one temperature, the crystallographic mechanism of slip is not well understood. Numerous investigators in this field have assumed that the resolved shear stress is a deciding factor. Such considerations were given by Gough in his experiments with repeated torsion of iron single crystals4 and recently by Opinsky and Smoluchowski in their study of iron-silicon alloys," and will be dealt with later in this report. A third possibility is that slip lines in body-centered cubic metals may be formed by alternate slip on nonparallel (110) planes and not by slip on any plane of higher indices. This was advanced by Elam in the case of ß brass." Since (1101 planes are planes of highest atomic density in the body-centered cubic structure, this hypothesis would be in accordance with the plastic behavior of metals in other systems, namely, face-centered cubic and hexagonal close-packed. Experimental evidence in support of this hypothesis is lacking. Nevertheless the work of Barrett, Ansel, and Meh12 has definitely demonstrated that the slip resistance on {110} planes in iron is very much smaller compared with that on (112) and (123) planes. In the present investigation, molybdenum single crystals were carefully documented as to the operative slip planes and lattice reorientations after plastic extension at room temperature. In three crystals studied, (110) planes were found to be active invariably rather than the (112) planes, which were previously reported for molybdenum at this temperature&apos; and predicted by Andrade&apos;s theory.&apos; In two other cases (112) and (123) planes seemed to operate together with (110) planes as active slip planes. However, it is believed, after careful analysis of crystallite and lattice rotations, that slip on (112) and (123) planes may be spurious. Particular reference will be made to the formation of asterism, lattice rotation, and crystallite rotation in an attempt to show that slip apparently observed on {1121} and (123) planes is actually slip on nonparallel {110} planes. Materials and Treatment Sintered molybdenum rods ? in. in diam, obtained from the Fansteel Metallurgical Co., were reported to have a purity of 99.9 pct.* One specimen, Mo-13, ¼ in. in diam, was supplied by Westinghouse Electric Corp. Single crystals were prepared, using the technique devised by Chen, Maddin, and Pond." Mo-13 was ¼ in. in diam x 7 in. long with a reduced gage section of 0.200 in. in diam x 3 in. long. The center of

Jan 1, 1952

By B. L. Averbach, Morris Cohen, S. A. Kulin

The martensitic transformation can be initiated by elastic stresses at temperatures above M. in a steel containing 20 pct Ni and 0.5 pct C. Shear strains and normal tensile strains acting on a potential habit plane promote the transformation but compressive strains oppose it. THE strain sensitivity of the martensitic transformation has long been recognized. It is now well known that martensite formation can be induced by plastic deformation at temperatures below, and not too far above, the M, point.&apos; However, the exact role of elastic vs. plastic strains and of the concomitant stresses has proved to be a very elusive matter. Many investigators2-&apos; of the steel hardening reaction have made use of the stresses resulting from volume changes to account for such observations as the existence of retained austenite and its variation across the section of a bar, the change in the amount of retained austenite with quenching rate, and the self-stopping nature of the transformation when the cooling is stopped. The effect of applied strain as an independent variable has been studied in 70 pct Fe-30 pct Ni alloys by Scheil7 and McReynolds.8 Some data are also available for lithium and Li-Mg alloys,9 Cu-Zn,10 Cu-Sn and Cu-A1 alloys," and aus-tenitic stainless steels."-" It has been found that martensite formation can be induced isothermally, even at temperatures above M,, by mechanical deformation of the parent phase. The competitive nature of plastic yielding by slip and by the martensitic reaction has been depicted by Scheil.7 Basically, Scheil postulated that: 1—at temperatures not too far above M, (where austenite is less stable than martensite), a critical resolved shear stress (within the elastic range) is required to promote the transformation to martensite; 2—at and below M,, the austenite lattice becomes mechanically, as well as thermodynamically, unstable and shears over spontaneously into martensite without the application of external stress; and 3—at temperatures sufficiently far above M8, the critical resolved shear stress for martensite formation increases to a level above that required for slip, and plastic flow then supersedes the transformation when external stress is applied. Scheil&apos;s concepts were tested by McReyno1ds8 ho found no change in the elastic moduli (measured both statically and dynamically) of 71 pct Fe-29 pct Ni alloys in the vicinity of the M, temperature. Since the moduli did not approach zero or become negative on cooling to M,, it was concluded that the austenite lattice does not become mechanically unstable at M2, as postulated by Scheil. McReynolds also reported that M, is not raised by elastic stresses. Accordingly, plastic deformation was considered to stimulate the transformation by generating martensite nuclei in the distorted regions of the parent phase. The above issues with regard to the role of applied stress require clarification, if a basic understanding of the martensitic transformation is to be obtained. For example, in the reaction-path theory,&apos;"," it is postulated that strain embryos exist in the austenite that provide part of the activation energy for the nucleation process. These embryos are visualized to be regions of localized strain in which the atoms in the austenite lattice are displaced part way along the path to their ultimate positions in the martensite. Consequently, it would

Jan 1, 1953

By M. A. Goudie

The deposits occur in a large salt basin of Middle Devonian age. The potash, the final deposit in the salt basin, results from several interrupted cycles of evaporation and dessication. The deposits are extensive, and, at first glance, relatively undisturbed. With more and more wells being drilled, it has now become evident that salt solution has played a large part in changing the original deposits, resulting in some cases in partial to complete removal of the potash and the underlying halite. The most dominant factor in the removal of salt by solution appears to have been tectonic movement and consequent faulting, probably of relatively minor dimensions but of major importance. Evidence which indicates the tilting of the evaporite basin to the north and northwest is shown by the changing pattern of the basin during succeeding eras of potash deposition. The potash minerals of most importance economically are sylvite and carnallite. Reserve calculations indicate that 6.4 billion tons of recoverable high grade potash in K2O equivalent exist in the basin. The Devonian salt basin, which contains the Saskatchewan potash deposits, extends from just east of the foothills in Alberta, north as far as the Peace River area, across Saskatchewan and into Manitoba as far east as Range 10 west of the First Meridian and south into Montana and North Dakota (Fig. 1). The basin is closed everywhere except to the northwest. The known potash deposits are confined almost entirely to the Province of Saskatchewan, with the exception of a small area in western Manitoba bordering the Saskatchewan boundary. The following discussion will concern only the Saskatchewan part of the basin. The evaporite series in the basin is defined as the Prairie Evaporite Formation of the Elk Point Group, of Middle Devonian age. Recent work done by potassium-argon dating methods has indicated an Upper Middle Devonian (Givetian) age of from 285 to 347 million years for the potash. The Elk Point Group consists in ascending order of the Ashern, Winnipegosis, and Prairie Evaporite Formations. The Ashern formation, with an average thickness of 30 ft, sometimes called the Third Red Bed, consists of dolomitic shales and shaly dolomites. The Winnipegosis, is a reef-type dolomite, usually with good porosity, and in many cases oil-staining, although to date no production has been obtained. The thickness varies from 50 to 250 ft. The Prairie Evaporite formation, varying from 0 to 600 ft in thickness, consists of halite with interbedded anhydrite and shale, with considerable amounts of potassium salts in the upper part of the formation. The potassium salts are chiefly chlorides, although very minor occurrences of sulfates have been re- ported. The anhydrite beds do not appear to be continuous, although generally one or two bands of anhydrite underlie the lowest potash zone and are used as marker horizons. The shale occurs as seams interbedded with the salts, as large irregular inclusions in the salts and as very fine particles in intimate mixture with the salts. The Prairie Evaporite Formation is overlain by the Second Red Bed member, the Dawson Bay Formation and the First Red Bed Member of the Manitoba Group, listed in ascending order. The Red Beds are shales which vary in color from red to green, maroon, grey, grey-black, and reddish purples. They serve as marker horizons for coring the potash. The Second Red Bed averages 14 ft in thickness, the First Red Bed 35 ft. The Dawson Bay Formation, which everywhere overlies the First Red Bed and the Prairie Evaporite Formation in the area under discussion, is a reef type of carbonate, in some places limestone, in others limestone and dolomite, with vugular to pinpoint porosity averaging 130 ft in thickness. In some parts of the area, it has a salt section near the top of the formation, usually with interbedded shales and limestones. In other parts of the area, it is waterbearing and the salt is absent. Detailed mapping has indicated that the areas in which the Dawson Bay is water-bearing are areas which have been disturbed by faulting. Where the Dawson Bay is salt-bearing, the porosity has been plugged by salt. The total thickness of the salt varies from between 600 to 700 ft in the center of the basin to zero at the northern edge of the basin (Fig. 2).* The salt-free area in the center of the Province is believed to have resulted from removal of salt by solution. Evidence from several wells suggests that salt removal has been a continuing process from the time of deposition to the present day. One well drilled between the Quill Lakes for potash information encountered

Jan 1, 1961

By C. W. Mueller, P. H. Robinson

A technique was developed for depositing single -crystal films of silicon on single-crystal sapphire substrates via the pyrolytic decomposition of SiH4/H2 mixtures. Electron diffraction and X-ray Laue reflection examinations of these films revealed single-crystal patterns. These films were characterized by measuring conductivity type, thickness, resistivity, and Hall mobility. Chemical etching, dislocation staining, and electron-microscopy examination have indicated the presence of low-angle grain boundaries and microtwinning in the silicon films. The role of the sapphire substrate with respect to its quality, orientation, surface finish, and thermal and chemical treatment was investigated. Hall mobility as a function of sapphire orientation was measured. A mobility equal to 88 pct of bulk silicon mobility was attained. Insulated-gate field-effect transistors which have a transconductance of 4000 pmho at 6 ma were fabricated using these films. THE technology of the fabrication of microelectronic circuits with active and passive components directly on a single wafer of semiconductor material has attained a high degree of perfection during the last few years. One inherent difficulty with the single-wafer approach has been undesirable coupling between components arising from the common base material. Several techniques have been developed for minimizing this coupling and obtaining good electrical isolation between components; however, these result in extra capacitance being added to the circuit. The extraneous capacitance of the connecting leads, resistors, and capacitors degrade circuit performance and limit the high-frequency operation of these circuits. Recently there has been considerable interestl-3 in the deposition of silicon films upon insulating substrates. The use of silicon as the semiconductor allows the subsequent utilization of the large amount of technology which has been developed for silicon for the fabrication of good active components. A sapphire substrate supplies an excellent insulating material which at the same time gives a thermal-conductivity improvement of an order of magnitude over glass-type substrates. This paper will describe the results obtained for the deposition of single-crystal silicon layers upon single-crystal sapphire substrates as well as the characterization of these films. I) APPARATUS The &apos;system used for the growth of silicon films on sapphire substrates by means of the pyrolysis of SiH4/H2 mixtures is shown in Fig. 1. The all-quartz reaction chamber is water-cooled and includes a quartz optical-flat window for use in conjunction with a 5 kw, 2.5 Mc per sec constant-temperature controlled rf generator. The temperature sensor for the rf generator is a thermopile which is focused through the quartz optical flat onto the susceptor surface. With this system the rf generator will hold the temperature of the susceptor to about +0.3°C at 1200°C. The susceptor is supported on a quartz sled inclined to the gas flow to increase the uniformity of the deposition. The water jacket surrounding the reaction chamber is used to keep the quartz wall temperature low enough to limit the decomposition of the silane to the rf-heated susceptor which contains the sapphire substrates. This eliminates the reaction of silicon with the hot quartz to form SiO which deposits out on the apparatus walls, thus obscuring vision, and interferes with the optical measurements of the temperature as well as with the growth process itself. Provision is also made for "n"- and &apos;p"-type doping by using mixtures of either phosphine or diborane in hydrogen which are added to the silane during the growth process. The hydrogen used is the purest commercially available tank hydrogen further purified by use of a palladium diffuser. The susceptors which were used include 1) spectro-scopically pure graphite, 2) high-purity molybdenum, 3) high-resistivity silicon using a low-resistivity insert to facilitate coupling to the rf generator, 4) silicon carbide-coated graphite. The latter two susceptor materials produced silicon films on sapphire substrates with more reproducible resistivities. Prior to silicon deposition all connecting lines as well as the reaction chamber are thoroughly flushed with argon and then hydrogen. After the hydrogen flow is set, the growth temperature is adjusted to 11009 to 1150°C, and when thermal equilibrium is reached the silane flow is started. The flow rate of the hydrogen must be fast enough so that turbulent flow occurs in the size of reaction tube used, otherwise a streaked nonuniform silicon deposit occurs. After the deposition

Jan 1, 1967

By R. T. Hukki

The purpose of this paper is to present an analysis of the unit operation of grinding and the circulating load, of the unit operation of classification and the circulating load, and of the two superimposed into one operation to clarify the overall trend of events relating the circulating load to the capacity of the closed circuit grinding system, to the specific surface area of the final fine product, and to the energy consumption in the closed circuit grinding process. The capacity of the closed circuit is frequently, if not always, determined by the classifier, not by the mill." With these words a startling idea was expressed in 1934 by Dorr and Anable.1 Today, some 30 years later, grinding in closed circuit is universally accepted. The basic curve presenting the capacity factor versus circulating load by Dorr and Anable has obviously been so convincing that not a single authoritative investigation regarding its validity seems to have been conducted. The only major subject of discussion appears to have been the relative merits of the &apos;high&apos; and &apos;low&apos; circulating loads. It seems to be universally agreed that much still remains to be hoped for regarding sharpness of classification. BASIC CURVES FOR CONSTANT MILL LOADING Fig. 1 presents a series of basic curves for constant mill loadings in function of new feed rate and circulating load. The mill loading indicated by each curve is the sum of hourly tonnages of new feed and circulating load. For example, at point 0 where the new feed rate is 100 t/h and the circulating load 300%, or 300 t/h, the mill loading is 400 t/h. In closed circuit grinding, the new feed rate corresponds to the rate of finished classified material produced. Although at this stage the effect of classification is excluded in construction of the basic curves, the trend indicated by these curves is strongly in favor of closed circuit operation under relatively small circulating loads. A study of the curves reveals that, e.g., a mill operating under a load of 200 t/h at point P where the new feed rate is 50 t/h and the circulating load 300% will also operate under the same total load at point Q where the new feed rate is 100 t/h and the circulating load 100%. In other words, a reduction of circulating load from 300% to 100% would lead to doubling of the new feed rate and doubling of the rate of production of finished material, without a change in the total mill loading. The obvious question is, is it also possible by a reduction of the circulating load in a closed industrial grinding circuit to increase the production of finished classified material and, if so, how and to what extent? GRINDING TESTS BY DORR AND ANABLE In a series of open circuit grinding tests reported by Dorr and Anable1 limestone was ground at various feed rates in a 3-ft ball mill. The original data are reproduced in Table I. It can be seen that with the increased feed rate production of -65-mesh material increased and the energy consumption per unit weight of the -65-mesh product decreased in the reverse ratio. The data shown in Table I have been reproduced for our purposes in Table II for a scale corresponding to grinding in a large industrial mill. This reproduction

Jan 1, 1968

By A. M. Gaudin, D. W. Fuerstenau

IN concentration of certain ores by soap flotation, prevention of quartz flotation is desired; the contrary is true in treatment of some oxide iron ores. Experimental study of the flotation of quartz goes back at least a quarter of a century.&apos; Clean quartz does not float with fatty acids or soaps but does float when activated by one of many metallic salts. Control of pH is important. Broadly speaking the problem of quartz flotation has been approached from the empirical side, as was the related topic of flocculation or dispersion of quartz suspensions? Empirical studies indicated the existence of critical relationships between the additions of collector and activator. These additions were not stoichiometric for the formation of soaps of the activator cation. For example, Gaudin and Rizo-Patrona found optimum quartz flotation at pH 10.6 using barium activation and oleate collection when the molar ratio of oleate to barium was 1 to 1 instead of 2 to 1, as would be called for formation of Ba(OL),. Cooke and Digre" &apos; found that if calcium is the activator, a minimum quantity of that agent is required for effective flotation when the pH is 11.5. In other words, if the concentration of Na&apos; exceeds 10 -2.5 mols per liter or if the concentration of H exceeds 10-11 mols per liter, the concentration of Ca++ in solution required for quartz activation is increased. This suggests that H+ is much more effective than Na+ on 10-2.5 quartz, perhaps in the ratio ofp or 108. Schuh- mann and Prakash6 studied anew the activation of quartz by barium, calcium, aluminum, and ferric iron. Salts of these metals are activators provided they are in stoichiometric excess over the collector used with them. Alkaline-earth salts are effective only in alkaline circuits. Important as they are in giving data on the effect of various agents on quartz, all these studies fail to give information on the detailed structure of the quartz-solution interface. A beginning in this direction was made by Gaudin and Chang,? who measured the adsorption of radioactively marked barium and laurate ions at quartz surfaces, and by Gaudin. Spedden, and Laxen,8 ho measured the adsorption of radioactively marked sodium at quartz surfaces. Work has also been done by Bell0 n measurement of the adsorption of sodium ion and chloride ion at quartz-solution interfaces. In all these experimental measurements of adsorption, the solid was centrifuged or filtered from the bulk of the liquor with which it had been associated; and the composition of the solid surface was ascertained by analysis, after allowance was made for the ionic composition of the liquor included with the centrifuged solids on the assumption that it is identical to the bulk liquid. Adsorption measurements give only the total amount of collector adsorbed at the quartz-solution interface; they give no indication of the distribution of these adsorbed ions at the interfacial layer. It is the ionic structure of the solid-solution interface and its importance to flotation that is the subject of this paper. The structure of this interfacial layer can be investigated by means of electrokinetic methods so that by combining electrokinetic studies with adsorption measurements, a more complete picture of the quartz surface may be obtained. The following studies were made to ascertain how electrolytes used in the soap flotation

Jan 1, 1956

By Charles W. Chesterman

Perlite is a glassy volcanic rock which will, upon rapid controlled heating, expand or "pop" into a frothy material of low bulk density, valued as a lightweight aggregate. The term perlite also is applied to the expanded product. The chemical composition of perlite ranges from that of rhyolite to that of andesite; the water content ranges from 2 to 5%. Most perlite is pale gray, but some is black, brown, or even green. Inclusions in perlite are common and include phenocrysts of quartz, feldspar, biotite, and hornblende, and miscellaneous rock fragments. Occurrences of perlite are restricted to areas of Tertiary and Quaternary volcanic rocks. Perlite occurs as flows associated with thick accumulations of pyroclastic debris and other flow rocks, as dikes, sills, domes, and as selvages of intrusive bodies of volcanic rock. Individual flows of perlite range in thickness from a few feet to as much as 100 ft and commonly cover several square miles. Dikes and sills of perlite generally are small, but domes generally are substantial in size and may be as much as several miles in diameter. Most perlite deposits are large enough and topographically situated so as to lend themselves to low-cost open pit mining methods. Conventional equipment is employed in the mining of perlite. Although milling procedures may appear to be simple, care is exercised in order to produce a carefully sized product and to minimize the production of excessive quantities of fine materials, which have limited use and are considered as waste. Perlite is processed, or "popped," in one of two basic types of furnaces: horizontal rotary or vertical stationary. Both types of furnaces include product collection and dust control systems. Many commercial deposits of perlite are known to occur throughout the world, and although no exact estimate of reserves is available, there appears to be an adequate supply to meet current and future worldwide requirements. End Uses The uses of expanded perlite are many and varied and are based primarily upon its physical and chemical properties. Because of its (1) low bulk density, which generally ranges between 2 and 20 lb per cu ft; (2) low thermal conductivity "K" factor (Btu per hr per sq ft per OF per in, of thickness), which ranges between 0.267 for loosely packed perlite (bulk density of 4.48 lb per cu ft) to 1.25 for perlite gypsum plaster; (3) high resistance to fire; and (4) low sound transmission, expanded perlite is used to a great advantage as aggregate and insulating material in the construction industry. Early uses of perlite were related exclusively to the construction industry, which now consumes about 60% of the domestic production. The primary use of perlite now is as aggregate in insulation boards. Other uses of perlite in 1971 were as aggregate for plaster, horticultural aggregate, low-temperature insulation (cryogenics), and concrete. Because of its low thermal conductivity ("K" factor), high adsorption of sound, low bulk density, and fire resistance, perlite aggregate plasters hold many advantages over conventional plaster. Perlite aggregate plaster lends itself to machine applications, dries quickly, and has relatively high elasticity which reduces cracking after drying. These same features make perlite useful as aggregate in portland cement concrete and asphalt concrete for construction wherever weight reduction, insulation, and acoustics are of prime importance. Perlite-portland cement concrete offers up to 20 times more thermal insulation than conventional concrete, and because of this it finds wide application in roof decks and floor fill in multistory buildings, in

Jan 1, 1975

By D. de Fontaine

A convolution procedure is described for obtaining an analytical expression for the composition profiles of solute concentration:; in binary solid solutions. The free energy of solutions containing such compositional inhomogeneities can then be evaluated and the stability of spherically symmetric clusters (G.P. zones, for short) us a function of their radii and central compositions can be studied accordingly. Two characteristic zone radii are defined: the minimum radius (RM) corresponding to the smallest zone for which growth is energetically favored over dissolution, and the critical radius RK) corresponding to zones associated with the minimum work of formation. The values of RM and RK coincide at the two classical limits of nucleation and growth and spinodal decomposition. As the spinodal is approached, however, RM and RK diverge increasingly and it is proposed that the initial] decomposition morphology is governed more closely by the minimum, rather than by the critical radius since the former possesses no singularity within the m eta stable phase boundary. Reversion is briefly discussed in the light of the foregoing and it is shown that a set of reversion temperatures located between the coherent miscibility gap and the coherent spinodal must correspond to a given distribution of low temperature zones. SINCE the early works of Guinier1 and preston,2 the kinetics and morphology of GP zones have been studied extensively, see Ref. 3, for example, for a recent review. Nevertheless, a completely satisfying theoretical treatment of the decomposition of meta-stable or unstable binary solid solutions is still lacking due to the almost unsurmountable mathematical difficulties that one encounters in attempting to solve the pertinent diffusion equation which, in principle, describes the whole process. This equation, first derived by Cahn for the initial stages of spinodal decomposition,4 is a nonlinear partial differential equation of the fourth order in anisotropic space, a general analytical solution of which is completely out of the question. From this equation, some prediction can be made, however, concerning the later stages of spinodal decomposition5 and a recent numerical solution6 of the equation in the case of one- and two-dimensional composition modulations of a solid solution inside the spinodal has led to the following considerations: the sinusoidal morphology7 is to be expected only for solutions of average composition c, located near the center of the coherent miscibility gap.8 For solutions with co closer to the spinodal decomposition, depleted shell zones9 should constitute the first recognizable decomposition product, later evolving to iso- lated, enriched preprecipitates in a uniformly depleted matrix as this must, in all cases, constitute the final metastable morphology, barring, of course, loss of coherency and change in crystal structure. In agreement with the qualitative description of Bonfiglioli and Guinier,l0 it thus appears that the various models discussed by previous authors are, in fact, idealized successive stages of the general decomposition process. Unfortunately, corresponding calculations for a solid solution outside the spinodal are not yet available because of two additional difficulties: it is essential that the diffusion equation be solved in three-dimensional space, and furthermore, one must include in the equation an appropriate "source term" describing the statistics of the creation of composition fluctuations required to trigger the nucleation events. At present, a general treatment of this problem does not appear to be tractable. Nevertheless, by using a modelistic approach, i.e., by selecting a priori certain composition variation profiles and by evaluating their associated free energies, one can draw certain conclusions relative to the stability of these isolated composition modulations or "zones", both outside and inside the spinodal. This is the object of the present investigation. THEORY We shall now describe a standard four-parameter spherically symmetric GP zone, evaluate its free energy according to the Cahn and Hilliard11 formula (Eq. [5], below), and plot the corresponding energy surfaces as a function of zone amplitude and radius as

Jan 1, 1970

By Irving B. Cadoff, Alvin A. Machonis

The resistivity, Hall coefficient, Seebeck coefficient, and thermal conductivity were measured as a function of temperature for cation-rich alloy single crystals covering the composition range across the PbTe-SnTe system. Alloying of PbTe with up to 20 pct SnTe was found to have little effect on the energy gap. Above 20 pct SnTe the alloys were "p" type but below this range the sign could be varied by heat treatment. The lattice thermal resistivity of the compounds SnTe and PbTe is raised by alloying one with the other. Z values in the order the interesting values obtained. THE PbTe-SnTe system has several interesting features. For one, PbTe is a useful thermoelectric material and the possibility of improving its figure of merit by alloying with SnTe, an isomorphous compound, has been suggested since these pseudo-binary solid solutions generally have a more favorable ratio of electrical conductivity to thermal conductivity than either of the components.&apos; Other interesting features relate to the conductivity mechanism, band structure, and stoichiometry of the compounds and their alloys. PbTe is a semiconductor with an energy gap of about 0.29 ev2 at room temperature whose conductivity sign and magnitude can be varied from "n" to "p" by controlling the proportion of lead and tellurium with respect to the stoichiometric ratio.3 Excess lead results in "n"-type conduction. SnTe is found to exist only as a "p"-type material of relatively high conductivity. This behavior is attributed to stoichiometric deviation by Brebrick4 but Sagar and Miller proposed that the behavior of SnTe must be due in part to the presence of an overlapped band. An investigation of alloys of this system, therefore, might give additional information which would permit one to evaluate which of the two proposals is the more appropriate one. Abrikosov et al.&apos; studied the room-temperature electrical properties of these alloys and reported data for Seebeck coefficient and resistivity on poly-crystalline alloys. The present work is a more exhaustive survey of the PbTe-SnTe system. Re- sistivity, Hall coefficient, Seebeck coefficient, and thermal conductivity were measured over a wide temperature range for single crystals at 10-pct intervals of lead/tin ratio across the pseudobinary system. The relative concentration of tellurium was controlled so as to obtain metal-ion excesses in all cases. SAMPLE PREPARATION The crystals were prepared by melting elemental lead, tin, and tellurium in weighed proportions in evacuated Vycor capsules. The lead and tellurium were high-purity grades obtained from American Smelting & Refining Co. The tin was supplied by Comico. The proper calculated proportions of lead, tin, and tellurium were weighed and charged into prepared Vycor capsules prior to evacuation. The capsules were prepared from 15-mm Vycor tubing. A sharp point was worked on one end of the tube. A pyrolytic graphite coating was deposited on the Vycor walls by heating the tip to 800°C in an atmosphere of acetone-saturated argon. An additional coating of graphite was deposited on the pyrolytic coating from an Aquadag suspension. Above the coated tip the tube was reduced in diameter to form a constrictive neck. To avoid scratching the graphite coatings the charge was placed in the tube above the constriction. After a low-temperature bake, the evacuated capsule was sealed. On subsequent heating the charge melted down into the lower portion of the capsule. The crystals were grown by lowering the capsule through a Bridgman-Stockbarger furnace. The lowering rate was 1 in. per 8 hr. The upper portion of the furnace was set for 950°C and the lower portion for 800°C. In general the yield of single crystals was about 25 pct. The mixed compositions were, as expected, the most difficult to grow. The finished crystals were sectioned into 5/8-in. slices. The tip, end, and middle slices from each crystal were analyzed by X-ray fluorescence to determine the lead-to-tin ratio. The resulting values were used to plot a composition vs distance plot for each crystal. Slices were selected from each crystal, with the aid of the composition plots, to cover the complete range of compositions at 10-pct intervals. In general, the slices selected were taken from the seed end of the crystal where the longitudinal segregation (as determined from the X-ray fluorescence analysis) was a minimum. Laue single-crystal analysis and metallographic analysis was used to verify if a slice was single or polycrystal. Any grain boundaries were clearly visible in the as-cut and polished condition. In ad-

Jan 1, 1964