Search Documents

By Nicholas J. Grant, John S. Benjamin

A series of composite alloys containing a high volume of NiAl, Ni3Ah or CoAl, bonded with 0 to 40 vol pct of a ductile metal phase, were prepared by powder blending and hot extrusion. The binder metals were of four types: pure nickel or cobalt, near saturated solid solutions of aluminum in nickel and cobalt, type 316 stainless steel, and niobium. Sound extrusions were obtained in almost all instances. Studied or measured were the following: interaction between the alunzinides and the binders, room-temperature modulus of rupture values, 1500° and 1800°F stress rupture properties, hardness, structure, and oxidation resistance. Stable structures can be produced for 1800°F exposure, with interesting high-temperature strength and good high-temperature ductility. Oxidation resistance was excellent. A large number of experimental investigations have been made of the role of structure on the properties of cermets and composite materials. Gurland,1 Kreimer et al.,2 and Gurland and Bardzil3 have indicated the preferred particle size in carbide base cermets to be about 1 µ, with a hard phase content of 60 to 80 vol pct. The optimum ductile binder thickness was noted to be 0.3 to 0.6 µ.1 Complete separation of the hard phase particles by the binder is important in reducing the severity of brittle fracture.' The purpose of the present study was to produce structures comparable to the conventional cermets, using a series of relatively close-packed intermetal-lic compounds rather than carbides as the refractory hard phase, and to study the effects of binder content and composition on both high- and low-temperature properties. The selected intermetallic compounds were particularly of interest because of the potential they offered in yielding room-temperature ductility. The highly symmetrical structures are known to possess high-temperature ductility and room-temperature toughness. Based on a ductile binder, the alloys were prepared by the powder-metallurgy route to avoid melting and subsequent alloying of the matrix, and were extruded at relatively low temperatures. It was expected that the composite alloy would retain useful ductility. In contrast, infiltration and high-temperature sintering led to alloying of the matrix and to decreased ductility. The systems Ni-A1 and Co-A1 were selected for this study. In the Ni-A1 system the compounds NiA1, having an ordered bcc B2 structure, and Ni3Al(?1), having an ordered fcc L12 structure, were chosen. In the system Co-A1 the intermetallic compound CoAl with an ordered bcc B2 structure was used. ALLOY PREPARATION The intermetallic compounds, see Table I, were prepared by using master alloys of Ni-A1 and CO-A1, with additions of either cobalt or nickel to achieve the desired compositions. The master alloy in crushed, homogenized form, was melted with pure nickel or cobalt in an inert atmosphere, cold copper crucible, nonconsumable tungsten arc furnace. The resultant intermetallic compounds were homogenized at 2192°C in argon, crushed, and dry ball-milled in a stainless mill to -100 and -325 mesh for the Ni-A1 compounds and to -325 mesh for the CoAl compound. Finer fractions were separated for some of the composite alloys. Several ductile binders were utilized. These included Inco B nickel, 5µ ; pure cobalt, 5 µ, from Sher-ritt Gordon Mines, Ltd.; fine (-325 mesh) niobium hydride powder; fine (15 µ) type 316 stainless-steel powder; and near-saturated Ni-A1 and Co-A1 solid-solution alloys, also in fine powder form. The niobium hydride was decomposed above about 700°C in processing of the compacts in vacuum to produce niobium powder. The Ni-7.1 pct A1 and the Co-5.3 pct A1 solid-solution alloys were prepared from pure nickel or cobalt and pure aluminum by nonconsumable tungsten arc melting under an inert atmosphere. The ingots were homogenized, lathe-turned to fine chips, and dry ball-milled in air to -325 mesh powder. These solid-solution alloys are designated NiSS and CoSS; see Table I. Subsequently the hard and ductile phases were dry ball-milled as a blend. Experiments clearly established the need to coat the hard particles with the ductile binder to optimize subsequent hot compaction by extrusion. Ordinary dry mixing usually resulted in nonhomogeneous alloys which were quite brittle. Conventional cermets are consolidated by liquid phase sinteiing or infiltration, which resulis in undesirable and uncontrolled alloying of the binder phase. For this study, a loose (unsintered) powder-extrusion process was emploved, minimizing reactions between binder and hard particle, thereby permitting much greater control of composition and structure. The constituent powders were first mixed in the desired

Jan 1, 1967

By T. M Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the 'active masses' of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. dC Thus----— kc [1] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of un-floatable mineral. Hence the first order equation is It has been proposed by various workers1-' that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris- as discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1953

By T. M. Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the 'active masses' of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. dC Thus----— kc [1] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of un-floatable mineral. Hence the first order equation is It has been proposed by various workers1-' that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris- as discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1953

By B. H. Caudle, B. M. Hickman, I. H. Silberberg

Secondary recovery projects often are not started in oil reservoirs until dictated by rising GOR's or declining oil production. Such circumstances require a well dispersed injection pattern to prevent the oil bank from bypassing the production wells. Many times this purpose is served by using the five-spot or line-drive patterns, in which half of the wells are made injection wells. A regular injection pattern is proposed that is based on the same well spacing as the five-spot but in which only a third of the wells need be converted to injection wells. Fluid flow model studies on this proposed pattern show that production histories will be as good as those for a five-spot pattern at mobility ratios of three and above. The model data are presented in a form that may be used to predict recoveries from both homogeneous and stratified formations. Introduction The injection of fluid into a petroleum reservoir usually is prompted by an actual, or predicted, loss in the productivity of the reservoir. Often, an oil reservoir is produced by pressure depletion until high GOR's dictate a change in policy. Fluid injection also causes increased ultimate recovery, but this usually is of less importance to the operator than is the maintenance of the production rate. It is a fact that the ultimate recovery in any particular fluid injection project is relatively independent of the number and spacing of the injection wells as long as the reservoir voidage through the production wells is balanced by the total injection rate. However, the response of a production well is directly affected by the distance it lies from an injection well. In places where the well productivities are declining, the injection wells should be spaced as evenly as possible among the producing wells ("dispersed" injection pattern). More particularly, if the production response depends upon the formation (and subsequent production) of an oil bank, each production well should be surrounded by injection wells. Otherwise, the oil productivity will increase very little as the oil bank moves past the well.'.' An evenly spaced injection pattern has the best chance of generating an oil kick in the production wells. If reservoir voidage and injection rates are to be equal, the ratio of injection wells to production wells must be equal to the ratio of the average injectivity to the average productivity (in reservoir volumes). This ratio is in strumental in choosing the injection pattern. Table 1 shows the production well to injection well ratio for the dispersed injection patterns that have been described in the literature.'-' All of these patterns will have comparable recoveries and response times (for the same basic distance between wells). The choice of pattern there fore depends upon the relative injectivity and productivity of the wells and upon the type of drilling pattern employed. Another important factor may be the ability of the proposed production wells to make up the lost productivity from those wells converted to injectors. These factors become more pronounced at the higher oil viscosities (low gravity oils). All other things being equal, the probable ratio of single well injectivity to single well productivity will be equal to the water-oil mobility ratio of the proposed flood. Thus, high mobility ratios (due to low gravity oil) should call for fewer injection wells — as long as these injection wells are "dispersed" to obtain more even pro duction response for the remaining wells. Secondary re covery projects in oil reservoirs seldom, if ever, will have mobility ratios requiring more injection wells than production wells. Thus, only the last five of the patterns listed in Table 1 will be of significant interest in oil recovery. Almost every development drilling program calls for the wells to be drilled along equally spaced rows and

Jan 1, 1969

By D. S. Rowley, F. C. Appl

Assuming that rock behavior, during cutting with a single diamond, may be approximated by that of a rigid, Coulomb, plastic material, a theory of single diamond cutting action has been developed. Using this theory, the stresses on the diamond cutting surface and the components of the cutting force have been determined. Theoretical results agree reasonably well with available experimental data. The theory of cutting of single diamonds may be used as the basis for subsequent study of the perfonnance of surface set diamond drill bits and other surface set diamond cutting tools. INTRODUCTION For more than twenty years, surface set diamond bits have been used successfully in the mining and petroleum industries to drill and core rock formations that range from medium hard to the hardest known formations. Most extremely hard rock formations cannot be economically drilled or cored by any other means. In view of the growing use and importance of surface set diamond drill bits, it is reasonable to study the way in which these tools cut, and the factors that influence wear of the diamonds. Such knowledge will contribute to better bit design and will improve operating procedures. To analyze the cutting performance of these tools it is appropriate to consider the way in which an individual diamond cuts. Once this is established, the over-all performance of the bits can be determined by considering the total effect of the individual diamonds. In recent years there has been a growing interest in rock mechanics and there now exist several publications dealing with the deformation and failure of rocks. Notable experimental work has been reported by Bredthauer,1 Handin and Hager,2, 3 and Gnirk and Cheatham.4 An extensive collection of data pertaining to rock properties at low confining pressures has been given by Wuerker.5 Cheatham and Gnirk have recently presented an excellent review of the state of knowledge regarding analytical and experimental rock deformation and failure.6 Generally rocks are extremely weak in uniaxial tension, and weak or only moderately strong in uniaxial compression. Both tensile failure and compressive failure are generally brittle in nature. This means that there is little or no deformation or "flow" of the material before shear or tensile fracture occurs. Thus, until fracture occurs, rock behaves very nearly like an elastic material in uniaxial stress. Significant changes in behavior of rock occur under conditions of triaxial compressive stress. There is generally a marked increase in compressive strength and in ductility. It has been demonstrated that the failure becomes plastic (in a macroscopic sense) at sufficiently high values of confining pressure. Sharp wedge chisel indentation studies conducted by Gnirk and Cheatham indicate that many rocks behave in a plastic manner at relatively low triaxial compressive stress states.4 If "failure" under triaxial compressive conditions is considered to be either complete rupture or macroscopic plastic flow, then the general failure characteristics of a rock can be reasonably well represented by the commonly known Mohr envelope. This certainly is not a perfect failure criterion (since the effect of the intermediate principal stress is neglected) but it has been found to be adequate for most engineering purposes. Mohr envelopes usually are constructed graphically using experimental triaxial test data. For purposes of analysis it is convenient to represent the envelope in equation form. Among the several approximate forms appearing in the literature, a parabolic form is suggested by Cheatham,7 and an exponential form is suggested by Paone and Tandanand.8 The exponential form has been adopted for this study, and based on published data, appropriate parameters have been established for 19 rock types. A typical plot of the data compared with the approximate envelope, and a summary of the rock properties and

Jan 1, 1969

By H. Herman, M. E. Fine

Kinetics of reversion and reformation of GP I were studied resistometrically in Al-1.7 at. pct Cu. The reversion process is over in roughly 1/2 minute at 205 "C irrespective of the amount of primary aging. After reversion the occurrence of an incubation time before reaging begins has been substantiated. The reaging kinetics is drastically accelerated by plastic deformation to the specimen. An incubation time is observed only if the specimen is aged before reversion. The incubation time and the effect of plastic deformation indicate that nucleation of GP I is heterogeneous. Simple clusters of copper atoms apparently do not act as nuclei for GP I. AFTER quenching aluminum-rich copper from the solid solution range, Guinier-Preston zones of the first kind, GP I, form very rapidly at room temperature. The rate, largely controlled by the concentration of vacancies, is initially most rapid1 (that is, at the time of the first measurement: 1/2 min) and decreases with time. The nucleation is thought to be athermal. If after aging for several days the GP I is dissolved by reversion at 205°C, according to measurement of Young's modulus an incubation period is observed before GP I begins to form on subsequent aging.2 Yet if the reversion is carried out immediately after quenching, both resistivity3 and modulus changes2 show no incubation effect. The zones which form after reversion have the same structure as those which form after the quench.4 Because of entropy considerations and also the short diffusion time, the solid solution after reversion, 2 min at 205°C, is probably more highly clustered than at the original solution treating temperature, 520" to 550°C. It appears most reasonable to associate with nucleation the incubation time observed on reaging after reversion. A simple cluster of copper atoms, many of which are presumed to be present after reversion, is then not a sufficient nucleus for GP I; a GP I nucleus must contain something in addition to a cluster of copper atoms. The purpose of the present note is to give some additional experimental information concerning the incubation time and rates of reaging after reversion. The existence of the incubation time has been verified by measurement of another property, the electrical resistivity. Acceleration of reaging after reversion by plastic deformation has also been investigated. In acl.dition, the kinetics of the reversion process itself was studied. EXPERIMENTAL Electrical resistance was measured by the poten-tiometric four probe method. The 1.7 at. pct Cu alloy was furnished by the Aluminum Co. of America in the form of 1/2 in. rod. The major impurities reported by Alcoa were Fe (0.002 wt pct), Si (0.003 wt pct), and Mg (0.001 wt pct). Specimens for measuremenl of resistance were prepared by rolling to 0.010 in., cutting 2 by 0.2 in. pieces, rolling the middle pclrtions of these to 0.005 in., and splitting the 0.010 in. sections up to the reduced central section to form current and potential leads. Aluminum extension wires insulated with ceramic tubes were spot welded to the leads. The specimen was supported by the lead wires. The whole specimen assembly was annealed at 500" C in an air furnace for more than 12 hr and then aged at room temperature. This was done to assure quick and complete resolution during subsequent solution treatments. These were carried out in Houghton's "430 Temper Salt" at 520'C for 1 hr. Unless otherwise specified, quenching was accomplished by quickly transferring the specimen from a salt pot to 25°C water. Aging was carried out in a 3 liter dewar of distilled water, rotated and stirred. The temperature was controlled to +0.01°C with a thermomercurial regulator in conjunction with a transistorized relay

Jan 1, 1962

By H. Bernstein

Grain coarsening tests were carried out on AI-4.5 pct Cu and AI-4.5 pct Si alloys. The effects of three variables, melt composition, pour temperature, and mold temperature, were determined. It was found that the macrostructure generally coarsened with increased pour and mold temperatures. Coarsening was extreme in the unrefined alloys but was retarded by the active grain refiners like titanium and columbium. The effect of boron was spectacular in suppressing coarsening tendencies. The results of the investigation support the carbide theory of nucleation as opposed to the peritectic theory. A REVIEW of prior work on the subject of grain refinement in the light metals indicated that, for the most part, it had been concerned with the positive effects of addition agents upon structure. This work produced various theories of refinement. Prominent among them were the peritectic theory and the transition element carbide theory. Certain inadequacies in the former theory appeared to be explained satisfactorily by the transition element carbide theory. In order to appraise these theories further, a study was undertaken of the permanence of grain refinement effects when subjected to thermal or chemical variations. It was considered that a study of this kind would be of practical value, since the light metals are exposed to thermal and chemical variations in the normal course of foundry operations. Accordingly, two alloys of commercial importance were selected for this investigation, the A1-Cu and A1-Si alloys. Experimental Procedure and Results Raw Materials: The base alloys, A1-Cu and A1-Si, were prepared from virgin aluminum and hardeners, induction melted, and pigged for remelting. The grain refiners, including titanium, columbium, zirconium, and boron, were added as master alloys. Analyses of the raw materials and base alloys are listed in Table I. Melting and Casting Practice: The metals were melted in clay-graphite crucibles in a Hevi-Duty pit-type resistance furnace, equipped with a Leeds and Northrup Micromax Recorder and Controller. After melt down of the base alloys, the additions were made. For each casting, approximately 80 grams of metal were poured into a preheated graphite mold held in a transite flask. The mold (Fig. 1) weighed 320 grams. To determine the effect of superheat upon structure, a set of melts was heated successively to four temperature levels, 705°, 775°, 845°, and 915°C, and test specimens were cast at each level. Prior to each pour, the furnace temperature was held constant for 20 min. Next, a set of melts was put through a superheat cycle and test specimens were cast at melt temperatures of 705" and 915°C successively. The crucibles then were cooled in air to 705°C as deter-mined by immersion thermocouple and additional test specimens were poured. The cooling interval was approximately 1 min. Another set of melts was furnace cooled with the cover slightly ajar and test specimens cast as before. The interval in this case was 15 to 20 min, varying with the superheat temperature. The mold temperature for all these tests was 400°C. To determine the effect of cooling rate without superheat upon structure, a set of melts was heated to 705°C, and castings were poured at mold temperatures of 205", 400°, and 540°C. An interval of 15 min elapsed between each pour. Mold temperatures were measured by a direct reading millivolt-meter, using a chromel-alumel thermocouple located at midwall thickness of the mold. To demonstrate the phenomenon of chemical coarsening, agents such as chromium and beryllium, which had acted as coarseners in a previous investigation,' were added to the A1-Cu alloy. Test speci-

Jan 1, 1955

By T. M. Morris

THE rate of flotation of solid particles determines the percentage recovery of these particles which can be obtained during a given time interval. It is an established fact that the recovery is greatest in the fine sand size range and falls off on either side of this size range. It is also a fact that the history of the development of concentrating devices has been marked by a desire to recover the small particles: tables were developed to treat material too fine to recover in jigs; flotation recovers material too fine to be recovered on tables. As our ores become more complex and require finer grinding, a greater proportion is in the slow-floating size range. It is therefore of economic importance that the effect of size of particle upon rate of flotation be studied. Many studies have been made along this line. The present investigation attempts to quantify the rate of flotation by means of a rate equation, and then compares the effect of size upon the rate of flotation. It was found that the flotation rate constant increases with increase in diameter of particle in the size range smaller than the optimum size particle. The flotation rate constant decreases with an increase in diameter of particle in the size range coarser than the optimum size particle. This suggested two different actions taking place, and these are discussed. Suggestions for further work in order to improve the flotation rate of both fine and coarse particles are made. The usual method of expressing rate is to plot cumulative percent recovery vs time. If we knew the equation of this kind of curve we could replot it as a straight line, and the slope of the line would be an index of the magnitude of the flotation rate. It is easier to compare straight lines than curves. The basis used for obtaining the equation of the recovery vs time curve is the law of mass action. This law states that the "rate of a chemical reaction is proportional to the `active masses' of the reacting substances present at any given time." This, of course, implies a constant temperature. We can modify this for our use by using the term flotation reaction instead of chemical reaction and the term floatable particle instead of active masses. [ ] where C denotes concentration of particles, t denotes time, k denotes a proportionality constant. If the exponent of C is 1, then we have a first order rate reaction. If the exponent of C is 2, then we have a second order rate reaction. For a first order rate reaction, we obtain, after integration, [ ] where k is again a proportionality constant and has the dimension of the reciprocal of time. C, is the original concentration or weight or percentage of floatable particles. C, is the concentration or weight or percentage of floatable particles remaining in the flotation cell after time t. If all particles of the mineral to be floated were capable of floating we could take C, as 100 and C, would be 100 - R, where R is the cumulative percentage recovery of mineral at time t. But it is not usual for all of the particles to be capable of floating, because of surface oxidation, etc. Therefore, we must introduce a factor x which is the percentage of unfloatable mineral. Hence the first order equation is [ ] It has been proposed by various workers1-3 that this equation describes the flotation rate. More recently N. Arbiter4 states that a second order rate equation fits the recovery vs time curve more closely than does a first order rate equation. Morris' has discussed the evidence presented for this statement. It has been concluded that the first

Jan 1, 1952

By R. S. Rickard, M. C. Fuerstenau

Anodic polarization curves for iron dissolution and cathodic polarization curves for copper deposition and ferric and hydrogen ion reduction were studied. These results were used to predict the relative rates of the reactions: The rates of the first two reactions were found to be first order with respect to the reacting ion, and the rale-controlling step was diffusion of the ions to the callzode surface. The rate of hydrogen ion reduction in an oxygen-free solution is dependent on the reciprocal of the square root of the hydrogen ion activity. In solutions containing oxygen, the rate of hydrogen reduction becomes dependent on the square root of the partial pressure of oxygen. At equal concentrations copper precipitation occurs about twice as fast as ferric ion reduction. In an air-saturated solution containing 0.01 M cupric ion at pH 2, the rate of hydrogen ion reduction zs about 10 times slower than that of copper precipitation. At pH values below 3, the rate of hydrogen reduction in oxygen-free solutions is very slozu. COPPER cementation has been utilized industrially for many years as a means of recovering copper from aqueous solutions. There are many different types of cementation plants, but all simply provide a means of contacting iron with relatively dilute copper solutions. In the past little or no attention has been paid to rates or mechanisms of the reactions involved. As a result, this and other studies'-3 have been undertaken recently to establish the important parameters in this system. The reactions involved in copper cementation by iron are: Since transfer of charge is involved in every case, these reactions should be considered from an electrochemical point of view. In any electrochemical system involving electrolytic conduction, passage of electrons necessitates ion transfer to the respective electrodes according to Faraday's law. The current in the system will be proportional to the rate of ion transfer, and, therefore, when the current in the galvanic cell is known, the rates of the reactions involved can be determined. Galvanostatic electrode measurements5" are applicable for these electrochemical studies, in that with this technique polarization curves (current density vs electrode potential) are established by impressing a constant current through an electrode and measuring the electrode potential. Once the current density, electrode potential, and reacting area are known for each reaction in this cementation system, the rates of the iron-consuming reactions can be determined. EXPERIMENTAL MATERIALS Several metal electrodes were used in this investigation; they are shown in Table I. Solutions were made with reagent-grade salts and conductivity water. Air, nitrogen, and oxygen were employed. All of these gases were saturated with water and degreased prior to use. EXPERIMENTAL PROCEDURE The experimental work was divided into two general areas. The first area involved galvanostatic measurements, which were used to determine the rates of reactions as a function of various parameters. The second area comprised actual cementation experiments to check the rates predicted from the electrode measurements. Galvanostatic Electrode Experiments. Polarization curves were obtained using a standard galvanostatic technique5'= with the apparatus shown in Fig. 1. The various electrode polarization experiments were run the same way except for changes in solution composition, polarity, and type of measured electrode. The general procedure involved the following steps: 1) a solution was made 0.5 M in sulfate ion and to the desired cation concentrations; this solution was placed in a constant-temperature bath and saturated with the desired gas; 2) the electrodes were cleaned and placed in the electrode holder; 3) the previously cleaned apparatus was assembled and immersed in a constant-temperature water bath (temperature controlled to ± 0.05"C); 4) the electrode was then positioned; the solutions were introduced, and the polarization meas-

Jan 1, 1969

By N. K. Chen, R. Maddin

THE textures of straight-rolled and of cross-rolled molybdenum were first determined with the aid of pole figures by Custers and Riemersma.' These authors have shown that for straight-rolling, the main texture was (100) [0111 with a considerable spread about the rolling direction. After cross-rolling a second texture was obtained in which there was a [111] axis perpendicular to the rolling plane, possessing a rotary symmetry. Recently Semchy-shen and Timmons2 in their study of preferred orientation of arc-cast molybdenum sheet obtained similar pole figures. In these two investigations, polycrystalline materials were used. In the main, the textures of molybdenum in these cases are qualitatively consistent with rolling textures of other body-centered cubic metals, such as iron and Fe-Si alloys." In the present investigation, single crystals of molybdenum were deformed by rolling with respect to specific crystallographic planes and directions. Since the deformation textures of polycrystalline aggregates have been related intimately to the behavior of individual grains,4 a study of possible correlations of final textures with respect to the initial orientations of the single crystals might conceivably lead to a better understanding of the textures. Annealing treatments were carried out after deformation in order to study the recrystallization characteristics as well as the recrystallization texture. Rolling of Single Crystals Single crystals of molybdenum (about 1/8 in. in diarn x 1-in. long), grown by the method previously described,5 were cold-rolled without intermediate anneals. Except for specimens MR-1 and MR-2, which were rolled at random in the preliminary test in order to find a proper rate of reduction, all other specimens were carefully documented as to their respective orientation before deformation by means of X-ray Laue back-reflection methods. A stereo-graphic projection was constructed from which it was possible to determine a particular plane and direction to be considered for rolling. Parallel flat surfaces were then polished metallographically with respect to a reference scratch on the specimen which was mounted in sealing wax. These flat surfaces contained the plane chosen for rolling. After etching, a Laue back-reflection photogram was again made of one of the flat surfaces to determine the exact plane of rolling. The initial orientations of all the specimens (specified by two elements for each crystal, direction of rolling and the pole of the rolling plane) are summarized in Fig. 1. The specimens were mounted in plasticene between two hardened Si-Fe plates. The sandwich was then rolled down, keeping the direction constant in the case of straight-rolling. In cross-rolling (MR-10, MR-11, and MR-12) the sandwich was rotated 90" after each pass so that every first and third pass contained the same rolling direction. Specimens MR-14, MR-15, and MR-16 were compression-rolled following the manner of Barrett." In these cases the direction of the sandwich through the rolls was changed by 3" to 5" for each pass until the specimens were rolled to the desired reduction in thickness. It was found that if extreme care was exercised in rolling, i. e., the amount of deformation per pass was extremely small, the specimens could be rolled to a large reduction in thickness. The rate of reduction used, with a single exception of MR-1. was generally 0.001 in. per 10 passes. Thus, it took about 800 passes to reduce a specimen to a final reduction in thickness of 87 to 98 pct. The results of rolling are shown in Table I. Cold-Rolling Textures The final orientations were determined by the conventional Laue transmission method," using molybdenum radiation, with or without a zirconium filter. Pole figures were plotted from (1101 and (200) reflections in which three degrees of intensity— heavy, medium, and light—were determined visually. Three main types of textures have been observed from the deformed single crystals. These are: 1—a

Jan 1, 1954

By A. M. Gaudin

Activation of sphalerite with silver ions results from the exchange of two Ag ions for one Zn ion in the sphalerite lattice. The exchange is proportional to the logarithm of time and proceeds until either of the reactants is consumed. Sodium cyanide prevents the activation of sphalerite by controlling the ratio [Zn++]/[Ag+]' in solution. If sphalerite has already been activated with excess silver, cyanide ions can deactivate sphalerite provided oxygen is present. SOLUBLE: silver salts are effective activators for flotation of sphalerite with potassium ethyl xan-thate as collector.* Since sphalerite takes up large quantities of silver ions rapidly,2-4 investigation of this system appeared fruitful for studying the mechanism of sphalerite activation and deactivation. Experimental data presented here comprise studies of the mechanism and rates of activation and deactivation of sphalerite with silver and cyanide ions. A 20-lb sample of sphalerite from Ottawa County, Oklahoma, was handpicked, crushed, screened, hy-drosized to produce a 65/100-mesh fraction, and purified by electrostatic separation. The sphalerite was then leached with benzene and absolute ethyl alcohol to remove possible organic contaminants and with 0.1 M NaCN to remove any copper that might have been picked up during screening. The benzene had been specially purified by grinding with fine sphalerite for 12 hr to remove trace impurities that might have had a specific affinity for sphalerite. Ii'inally, the mineral was washed, dried in flowing nitrogen, and stored in large ground-glass stoppered weighing bottles under a small nitrogen pressure. The specific surface of two samples from the 65/100-mesh sphalerite stock, measured by krypton gas adsorption," was 420 cm2 per g. Compared with spheres of the same size and specific gravity, the prepared mineral had five times as much surface. Results of a wet chemical analysis and of a spectro-graphic study are given in Table I. Manganese, nickel, tin, indium, thallium, cobalt, and magnesium were not detected spectrographically. All solutions were preparid by -adding reagent-grade chemicals to conductivitv water. which had been previously saturated with the desired gases. Nitrogen and oxygen were purified, but no attempt was made to purify the laboratory air. Radioanalysis was carried out with prepurified Ag110 tracer obtained from the Oak Ridge National Laboratory. The complete decay scheme for Ae is complex. Its primary emissions are 0.09 and 0.55 mev ß rays and 1.48 mev y raysB and its half-life is believed to be 270 days. In preparation of solutions of AgNO3, the amount of Ag110 added was negligible compared to that of the inert silver. Abstraction of silver from solution by sphalerite was measured after agitation of 1.5 g sphalerite with 20 ml of solution in 50-ml pyrex flasks. All sample preparations were carried out in a dry box so that the proper atmosphere could be maintained. After the mineral had been agitated in the silver nitrate solution, an aliquot of the solution and a sample of the wet mineral were transferred to separate counting vials and weighed. The solid sample was dried and re-weighed and the activity in both vials measured by means of a scintillation counter. With scintillation counting it is possible to count y radiation from the solid and liquid, a distinct advantage over the older ß-ray counting techdiques.4 All measurements of silver activity were made using a sodium-iodide well-type scintillation counter. The amount of zinc in solution was measured colorimet-rically by the monocolor method of Sandell.7 Activation Studies Exchange Between Zinc Ions and Silver Ions During Activation: The increase of Zn++ in solution during activation with Ag+ is shown in Table 11. These data indicate a replacement of one Zn++ by two Ag+. Analysis of a separate two-month agitation test indicated that the replacement of Zn++ by Ag+ was, as expected, in the molar ratio of 2.0.

Jan 1, 1958

By F. V. Lenel

Resistance sintering under pressure is a method of hot pressing in which a powder compact is subjected to pressure and simultaneously heated by passing a low voltage high amperage current through it. Equipment for this process is described. Its basic characteristics such as resistance requirements for powders and compacts, temperature distribution in compacts, and gas reactions during resistance sintering are discussed. Examples of the sintering process in compacts made of a single metal or an alloy powder and in compacts made of mixtures of powders are presented. Potential commercial applications of the process are evaluated. THE usual sequence of operations in commercial powder metallurgy is to compact metal powders in a die at room temperature and sinter the compact subsequently without applying pressure while the compact is in the sintering furnaces. In hot pressing, on the other hand, the compacting and the sintering are combined. The loose powder, or in many cases a cold formed compact, is inserted into the die which is held at the hot pressing temperature, the compact is left in the die until it attains its temperature, and then pressure is applied and maintained for a given length of time. The die may be heated by surrounding it with a suitable furnace, by high frequency induction heating, or by passing current through the die. In hot pressing, compacts of high density and good mechanical properties can be produced in a relatively short time. One of the principal difficulties in hot pressing is the lack of suitable die materials which will have adequate strength at the hot pressing temperature. This difficulty can be overcome at least partially by heating only the material to be hot pressed without heating the die directly. This can be done by passing a low voltage high amperage current through the material and simultaneously subjecting it to pressure. Either loose powders or compacts can be hot pressed by this method in which the desired temperature is produced by the power dissipated in the metal powder. This method of hot pressing has been termed "electrical resistance sintering under pressure." The most significant differences between this operation and conventional hot pressing are: 1—The sintering times are very short, usually a fraction of a second and at most a few seconds. 2—The powder and die are initially cold. 3—Heat is generated within the powder itself and not conducted in from the die. 4— The pressure used is high. 5—Cooling following sintering is rapid, amounting to a quench. Resistance sintering of metal powders under pressure has been suggested repeatedly. In 1933, G. F. Taylor' proposed an apparatus which consisted of an insulating tube made of glass, or a ceramic filled with the powder to be pressed, and plungers above and below the powder. The powder was to be heated by passing an electric current through it and pressure was then applied. The apparatus was intended principally for hot pressing cemented carbides. Although the principle of resistance sintering under pressure is clearly shown in this patent, few details are given. A sintering period of a second or a fraction of a second, which is terminated by the operation of an inertia switch, is mentioned but no values for current density are mentioned. Low pressures, such as atmospheric pressure or a somewhat higher pressure exerted through a lever, are suggested. W. D. Jones' discussed electrical resistance sintering under pressure and proposed using resistance welding apparatus for this purpose. In a patent issued in 1944, G. D. Cremera described a method of electrical resistance sintering under pressure using resistance welding apparatus which was to be applied mostly to nonferrous metals such as copper, brass, bronze, and aluminum. Current densities of 400,000 amp per sq in., sintering times of 1 and 2 cycles of 60 cycle current, and pressures of 5 to 10 tsi are proposed. Cremer suggested using an

Jan 1, 1956

By A. E. Jenkins, L. A. Baker, N. A. Warner

The electromagnetic levitation technique has been successfully applied to rate studies of the de-carburization of liquid Fe-C alloys from 5.5 to zero pct C at 1660°C using gas mixtures containing 1 to 100 pct CO2 with CO or helium as diluents. The rate was shown to be independent of carbon concen -tration in the melt and only slightly affected by total pressure. The results were shown to be consistent with control either wholly or predominantly by gaseous diffusion. As well as avoiding side reactions with the crucible, levitation allowed precise calculation of reaction rates since the geometry of the system was accurately known. It also enabled interpretation using the established mass-transfer correlations for a gas flowing past a single sphere, The existence of gaseous-diffusion control and the extremely high rates of reaction reported for gas flowing past small spheres are of great practical significance in relation to current investigations into the use of spray-refining techniques in continzcous steelmaking. PREVIOUS work on the kinetics of carbon removal from liquid iron has generally involved the use of oxygen as the gaseous oxidant. Investigations have been made on commercial steelmaking furnaces and in addition laboratory-scale experiments have been attempted with a slag phase present. The general lack of agreement and the obvious difficulty in interpreting the results is at least partly due to the complexity of the systems used. This work has been reviewed by carter1 and more recently by Bods-worth2 and ward.3 The gas-liquid metal reactions have also been studied with slag-free melts by the conventional technique of blowing oxygen onto the melt contained in a crucible. The reaction between the oxidant and carbon could be followed by sampling either the melt or the gas phase. Side reactions with the crucible material have precluded direct interpretation of the results. Fujii, 5-7 using O-A mixtures, attempted to correct for the crucible reaction and reported that the reaction between carban and oxygen occurred at the melt surface and was controlled by gaseous diffusion for carbon contents greater than a critical value and by carbon-diffusion control in the metal for carbon contents less than the critical. The critical value was given as approximately 0.15 pct C. However, whereas this was the case for oxygen contents in the gas stream up to 10 pct, between 10 and 20 pct the rate was found to be approximately independent of oxygen content, and Fujii attempted to explain this observation by postulating the presence of an oxide "film".7 Turkdogan et a1.' have studied iron fuming during decarburizing and suggested gaseous-diffusion control of the reaction above a critical carbon concentration, given as approximately 2 pct C. Filippov9-12 reports in a similar manner to Fujii but makes no mention of the crucible reaction with the carbon in the melt, even though this would be expected with the temperature, crucible material (MgO), and carbon contents used. Both CO2 and O2 were used as oxidant gases. In the present work the decarburization of liquid Fe-C alloys was studied using an adaptation of the electromagnetic levitation technique, and thus the complexities introduced by crucible side reactions were eliminated. In this technique a single sphere of the Fe-C alloy is freely suspended and heated by electromagnetic induction in a flowing stream containing the oxidant gas. The full range of carbon levels up to saturation was studied, the temperature of the melts in this instance being held constant at 1660°C. The oxidant was CO2 which was used either pure or in association with a diluent gas, the total gas pressure being maintained either at 1 atm or at a reduced pressure. Thus the reaction under study was:

Jan 1, 1964

By J. J. Frawley, W. J. Childs

PREVIOUS studies have shown that if a supercooled liquid metal was perturbed by vibration the probability of nucleation is is During a recent investi- gation to determine quantitatively the amount of energy needed to cause nucleation in a given supercooled liquid specimen, it was found that at low-energy values in the megacycle range, below the threshold energy value for cavitation, the amount of dynamic supercooling was greater than the amount of static supercooling. That is, if a specimen had a given amount of thermal supercooling under static conditions, its thermal supercooling could be increased by the addition of low-energy sound waves. The material used in this investigation was high-purity bismuth. This material was chosen because of its low melting point and the high degree of supercooling which could be easily obtained. The sound gener- ator employed in this work was a commercially built Sperry ultrasonic testing unit. This unit was equipped with a barium titanate crystal as a piezoelectric transducer. These sound waves were transmitted to the supercooled melt by a Pyrex rod, as shown in Fig. 1. The transducer was coupled to this Pyrex rod wave carrier by motor oil. The Pyrex wave carrier was approximately 1.5 ft long to prevent damage to the transducer by heat. The diameter of the Pyrex wave carrier appears to be critical. In the initial tests a wave carrier of 0.5 in. in diam was employed. However, the sound waves had no effect on the amount of dynamic supercooling. When the diameter of the wave carrier was increased to approximately the inside diameter of the Pyrex crucible, the sound waves were more effective. Approximately 200 g of metal was melted under a stannous chloride flux in a 30-ml beaker by induction heating. The temperature was recorded by a Pyrex-encapsulated chromel-alumel thermocoule which was placed in the beaker after the bismuth melted. The Pyrex wave carrier was immersed slightly below the surface of the liquid bismuth and the specimen was reheated to about 500°C. It was observed that the amount of supercooling depends on the amount of superheating the specimen receives; therefore, for meaningful results the amount of superheating was the same for all heating cycles. The specimen was allowed to cool in air and the change in temperature was continuously monitored with a potentiometric strip-chart recorder.

Jan 1, 1969

By W. L. Chu

It was observed that chromized coatings on low carbon steel could be embrittled by cooling slowly through the 475°C range. A higher chromium coating was more susceptible to embrittlement than a lower chromium coating. Coating structure and composition could be correlated with the chromizing process and prior carbon content of the steel. IT was recently observed that during the fabrication of rectangular tubings made with chromized thin gage 1010 steel, a high percentage of these parts were fractured readily in a brittle manner under the light impact-forming load employed. Ensuing investigation indicated the cause for such brittle failures to be the so-called 475°C embrittlement of the chromized coatings. The experimental work and results are briefly reported. Rectangular tube sections of normalized AISI C1010 steel, which had a cross-section of 0.5x1.5 in. and a wall thickness of 0.030 in., were chromized to yield coatings having several ranges of thickness, varying from 0.0005 to 0.002 in. Essentially, the chromizing process is similar to that described by Samuel and Lockington,' which consists of heating the parts in a sealed container at about 955°C for 4 hr in intimate contact with a chromizing mixture* samples in a hydraulic press, and the other employed a simple beam bending test of the flat sides of the tubing walls. Neither, however, was able to differentiate the relative brittleness of samples known to include both the brittle and tough types, since the impact-brittle samples were found to deform readily in a manner similar to the tough ones. Both the tough and brittle samples were studied by reflection microscopy. Observation showed that the microstructures of the steel cores were similar in content and distribution of inclusions as well as in the ferritic grain size. Results of Filar micrometer measurements showed that coating thicknesses+ of the brittle samples averaged 0.0007 in., whereas those of the tough samples averaged 0.0016 in. The micrographs of Fig. 2 illustrate the structural differences between coatings of both types of samples. By dissolving the steel cores in nitric acid, coatings of both the brittle and tough samples were isolated. Simple bending tests showed that coatings of the brittle samples cracked readily at about 90" bend, whereas those of the tough samples could be bent nearly 180° without cracking. The coatings as well as the steel cores were analyzed for carbon and

Jan 1, 1956

By G. Ansel, J. P. Doan

THE important literature concerning zirconium in magnesium-base alloys is predominantly contained in patent references. Sauerwald, Eisenreich, and Holub1-4 discovered the profound grain-refining influence of small amounts of zirconium on cast magnesium and the attendant increases in strength, and ductility. They extended their work to reveal that limited amounts of certain other elements, including zinc, could be added for further improvement in strength and hardness. British. Patent 5111376 relates to the parent discoveries of Sauerwald, Eisen¬reich, and Holub. All of these references deal primarily with casting characteristics and properties in the cast state. Minor mention is made of desirable tensile strength, ductility, and toughness of wrought Mg-Zr alloys* modified by one or more additional elements. No extensive reference in the literature" has been found to the particular characteristics of wrought Mg-Zn-Zr alloys, with which the present paper is concerned. - The work to be discussed represents essentially an extrusion study of the Mg-Zn-Zr alloys and indicates that this system is highly interesting in the wrought state both academically and commercially. Accordingly, an important aim of this paper is to describe the hot-workability, tension and compression properties, metal¬lography, and toughness of these alloys. Compositional variations are limited to the ranges o to 7 per cent Zn and o to I per cent Zr. EXPERIMENTAL PROCEDURE Alloying and Casting Alloying ingredients were cell mag¬nesium, commercial zinc (intermediate grade), and a chloride salt containing approximately 50 per cent ZrCl2 and 50 per cent alkali chlorides. This salt is produced by Titanium Alloy Manu¬facturing Company. Melting, alloying, and pouring procedure essentially followed the crucible process as described by Nelson.6 The magnesium was melted using Dow No. 310 flux, and the temperature raised to 1350° to 1400°F for the alloying operations. After alloying of the zinc, the zirconium containing salt was added in small portions with vigorous stirring. This is an adaptation of the method described in British Patent 533264.7 The alloying efficiency of zir¬conium was approximately 30 per cent. During the alloying of zirconium, fluorspar was added in amount required as a flux¬inspissating agent. Next the melt was refined with Dow No. 310 flux and held quietly for 15 min. Finally billets were cast in a book-type permanent mold

Jan 1, 1946

By Warren L. Howes

LITERALLY hundreds of designs of mill liners are in use in current grinding operations, varying in contour from smooth to the roughest of surfaces, and in materials from scrap rail to alloy steels. A certain amount of variation in materials can be expected, owing to differences in local market conditions and the character of ores. However, when one attempts to rationalize the designs in current operations of a parallel nature, there is little that can be called standard practice. Designs and opinions of designs are as divergent as day and night. Somewhere in this maze there must be clues to better practice. The purpose of this paper is to examine the factors for consideration in liner design, and to propose a vigorous campaign of exploration in this phase of grinding operations. The paramount considerations in the choice of mill liners for a grinding operation are: (I) durability, in view of the cost, and (2) developed grinding capacity. Because of the strategic nature of grinding circuits in controlling the production of most mining operations,1 grinding capacity usually is the greater consideration. GRINDING CAPACITY Mill liners affect grinding capacity in two ways: (I) by establishing dead space within a mill shell, which could be usefully occupied by live grinding media, and (2) by controlling the grinding action of the media. The extent to which liner thickness affects the potential volume of live grinding media, and thereby the grinding capacity, is shown in Table I. In a ball mill 7 ft. in diameter, for example, an inch of liner thickness reduces ball volume and corresponding grinding capacity by 5 per cent. With grinding capacity at a premium, this makes consideration something more than first cost of liner materials. A low-cost, voluminous material, made extra thick to secure adequate life, can be very expensive in terms of grinding capacity. On the other hand, in increased grinding capacity, a long-wearing material in a thinner pattern can save many times its value. [ ] From the standpoint of mechanics, the function of mill liners is to provide the means of driving or tumbling the load of grinding media. In this respect, an internal-spur gear-and-pinion assembly resembles a rotating mill shell with its contained load of grinding media, the mill-shell liners

Jan 1, 1945

5. The medal committee may invite nominations from the membership of the Institute, such nominations to be accompanied by an argument in favor of the award and details sufficient for a proper citation to be drawn. This medal has been awarded as follows: 1927—David William Brunton. 1929—John Hays Hammond. 1928—Herbert Hoover. 1930-Daniel C. Jackling. Miscellaneous Prizes Many of the Local Sections have established or administer prizes for the best paper presented within the year by any junior member. Annual Lectures Howe Memorial Lecture The Howe Memorial Lecture, in memory of Henry Marion Howe, Past President of the Institute, was authorized in April, 1923, as an annual address to be delivered by invitation under the auspices of the Institute by an individual of recognized and outstanding attainment in the science and practice of iron and steel metallurgy or metallography, chosen by the Board of Directors upon recommendation of the Iron and Steel Division. The following were the Howe lectures and lecturers for the years indicated: 1924 What is Steel? By Albert Sauveur. 1925 Austenite and Austenitic Steels. By John A. Mathews. 1926 Twenty-Eve Years of Metallography. By William Campbell. 1927 Alloy Steels. By Bradley Stoughton. 1928 Significance of the Simple Steel Analysis. By Harry D. Hibbard. 1929 Studies of Hadfield's Manganese Steel with the High-power Microscope. By John Howe Hall. 1930 The Future of the American Iron and Steel Industry. By Zay Jeffries. Institute of Metals Lecture An annual lectureship was established in 1921 by the Institute of Metals and it has come to be one of the important functions of the Annual Meeting. A number of distinguished men from this country and abroad have served in this lectureship. The roll is quoted below: 1922 Colloid Chemistry and Metallurgy. By Wider D. Bancroft. 1923 Solid Solution. By Walter Rosenhain. 1924 The Trend in the Science of Metals. By Zay Jeffries. 1925 Action of Hot Wall: a Factor of Fundamental Influence on the Rapid Corrosion of Water Tubes and Related to the Segregation in Hot Metals. By Carl Benedicks. 1926 The Relation between Metallurgy and Atomic Structure. By Paul D. Foote. 1927 Growth of Metallic Crystals. By Cecil H. Desch. 1928 Twining in Metals. By C. H. Mathewson. 1929 The Passivity of Metals, and Its Relation to Problems of Corrosion. By Ulick R. Evans. 1930 Hard Metal Carbides and Cemented Tungsten Carbide. By 8. L. Hopt.

Jan 1, 1930

By Willard Jillson

WITHIN the last ten years Kentucky has become celebrated for its low-sulfur bituminous coals. Prior to this time, many investigators had discovered the abundance of this coal but the fact was unknown to the general public until the last extension of Kentucky's mountain railroads was completed. The coals of Kentucky are, broadly, separated into two geographic units-the eastern and the western coal fields. On a basis of their sulfur content these coal fields fall into three units, which are indicated on the accompanying map, by numbers, 1, 2, and 3. The area marked 1, the very southeastern part of the State, shows the lowest sulfur, all coals being averaged. In this field the maximum is 1.04 per cent. and the minimum as low as 0.68 per cent.-both averages and not individual analyses, which would of course show much greater variation. Included in this group are the following counties: Lawrence, 0.87 per cent.; Martin, 0.75 per cent.; Johnson, 0.73 per cent.; Magoffin, 0.S7 per cent.; Floyd, 0.88 per cent.; Pike, 0.68 per cent.; Knott, 1.04 per cent.; Perry, 0.77 per cent.; Letcher, 0.80 per cent.; Leslie, 0.70 per cent.; Harlan, 0.79 per cent.; Knox, 0.86 per cent.; and Bell, 0.92 per cent. The total average sulfur content for these thirteen counties is 0.82 per cent. Bordering the restricted low-sulfur field of eastern Kentucky is the No. 2 belt (slightly darker), which is commonly known as the western border of the eastern coal field. Here the maximum sulfur is 2.91 per cent. and the minimum 1.05 per cent. Again these are averages and not individual analyses. The counties in this group are: Greenup, 2.60 per cent.; Boyd, 1.67 per cent.; Carter, 1.07 per cent.; Morgan, 1.40 per cent.; Wolfe, 1.83 per cent.; Lee, 2.93 per cent.; Breathitt, 1.85. per cent.; Owsley, 1.09 per cent.; Jackson, 1.06 per cent.; Rockcastle, 2.30 per cent.; Clay, 1.09 per cent.; Laurel, 1.08 per cent.; Pulaski, 2.24 per cent.; Whitley, 1.05 per cent.; McCreary, 1.50 per cent.; and Wayne. The sulfur percentage average for this area including fifteen important counties is 1.65 per cent.

Jan 9, 1919

By David Keys

APPLIED geophysics is the youngest child of that old branch of learning, that has been known from Aristotle's time as physics-the constitution and laws of nature. The mother science, with the help of mathematics, has given birth to many children, who, on becoming of age in the scientific sense, wedded more practical applications which later became known as electrical engineering, physical chemistry and acoustical engineering. Some vigorous sons still retain the family name, among which the most recent to take its place as a division of science is geophysics, which deals with the physics of the earth. Geophysics in its broad sense has many departments-meteorology, oceanography and seismology, to mention only three. Within the past 20 years, however, a special interest has arisen in the application of physical principles to, problems in geology, mining and civil engineering, with the consequent rapid growth of what has now been called "applied geophysics." Applied geophysics in this narrower sense is the application of experi-mental and theoretical physics to determine variations: and discontinuities in underground structures that cannot be ascertained from surface inspec-tion. It seems evident, then, that the physics department is the natural place in a university where the principles and methods involved should be taught and developed, for the physicist will also have had the requisite training in mathematics that the proper understanding of many methods now used will require. A knowledge of the procedure and principles, however, is not sufficient for the successful application of geophysical, methods. One may know the dynamics of flight of golf balls and of impact and yet not play a par game of golf ! The proper selection and use of the right club to suit the circumstances of the lay of the ball and an acquaintance with the possibilities and limitations of the strokes as well as the practical execution of the play is essential to a successful game. So, also, in the application of geophysical methods, it is important for the geophysicist to use the best device under the conditions presented, to know the limitations of the conclusions deduced from his physical measurements and to make the proper interpretation of the results. A knowledge of geology will be of great aid in coming to the correct con-

Jan 1, 1938