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By L. S. Darken

IT has long been recognized that the driving force in an isothermal diffusion process may be regarded as the negative gradient of the chemical potential (partial molal free energy) of the diffusing substance. This equivalence is exactly analogous to the fact that a mechanical force is the negative gradient of the potential energy. Nernst1 in 1888 expressed these driving forces in terms of osmotic pressure, which is directly related to the chemical potential and free energy. The relationship between the diffusivity, mobility and activity coefficient of electrolytes in aqueous solution has been developed in great detail by Onsager and Fuoss2 with results which have recently been confirmed experimentally by Harped and Nuttall.3 This interpretation of the driving force in isothermal diffusion has been discussed earlier by the author 4.6 who pointed out4 that "for a system of more than two components it is no longer necessarily true that a given element tends to diffuse toward a region of lower concentration even within a single phase region"; departure from the behavior of an ideal solution may be so great that the concentration gradient and the chemical potential gradient, or activity gradient, may be of different sign, thus giving rise to "up hill" diffusion. Since, in the discussion of that paper,4 some skepticism as to the validity of this conclusion was expressed, it seemed desirable to adduce experimental evidence to support it. Evidence of such diffusion in nonmetallic systems has already been reported. For example, it was observed by Hartley11 in a series of experiments in which sodium chloride diffused from a saturated solution in about 5 pct acetone in water into a solution virtually free of salt but with essentially the same acetone concentration. Specifically, a wine glass containing crystalline salt and saturated solution was totally immersed in a large glass cylinder containing the acetone-water solution without salt; the difference in specific gravity was such as to prevent convection. Diffusion of salt was accompanied by a marked depletion 0f acetone in the wine glass; after 3 days the acetone concentration decreased from 5.37 pct at the top of the glass, where the salt concentration was low (0.93 pct), to 1.84 pct at the bottom where the salt concentration was high (21.64 pct). This "uphill" diffusion was found to be such that the partial pressure (activity) of acetone was essentially equalized throughout the system when a steady state gradient of sodium chloride was maintained. In order to demonstrate the existence of such diffusion in metals, a series of four weld-diffusion experiments was made. In these measurements pairs of steel of virtually the same carbon content, but differing markedly in alloy content, were welded at the end and held at 1050°C for about two weeks. Subsequent analysis showed that carbon had diffused so as to produce an inequality

Jan 1, 1948

By H. Dykstra

Calculated wellbore pressures were obtained for parameters of radilcs ratio and permeability. In all cases bur two, after-production was allowed to occur for one day. The calculated pressure build-up data were compared with actual pressure build-up data from a condensate well. The paper discusses the comparison, gives reasons for, inability to match calculated (2nd actual data, and presents conc1usions derived from the comparison. INTRODUCTION Considerable information has been published on the decline in wellbore pressure for oil and gas fields resulting from production of fluids. Considerable information also has been published on pressure build-up for oil fields, but relatively little has been published for gas and condensate fields. Perrine' summarized the equations derived for pressure build-up in oil wells and showed how the different methods could be applied. On the other hand, very little theoretical information has been published on pressure build-up in gas or gas-condensate wells.2-4 Tracy,' by pointing out the similarity between the equation describing oil wellbore pressure decline and the equation describing gas wellbore pressure decline, showed how methods of pressure build-up analysis could be applied to a gas well. In analyzing pressure build-up data for gas or gas-condensate wells, it would be very desirable to compare actual pressure build-up data with calculated pressure build-up data. Such a comparison would lead to a better feeling for the quantitative picture of gas-well build-up curves and would help to establish a degree of confidence in a given build-up curve. This paper discusses the comparison and presents conclusions made from it. Calculated wellbore pressures are given for parameters of radius ratio and permeability. In all cases but two, after-production was allowed to occur for one day. The work was done in an attempt to evaluate the reserves of a gas-condensate field. As will be discussed later, it was not possible to make an evaluation. METHOD OF CALCULATION The method used for solving the depletion and pressure build-up behavior for a gas-condensate well was that developed by West, Garvin and Sheldon.5 or the IBM 704 computer program used, it was assumed that liquid condensing out of solution would have only a minor effect on the flow behavior. Therefore, the field was treated as having single-phase flow, with the gas saturation S, remaining constant at unity minus the connate-water saturation Sw. As will be discussed later, this may not have been a good assumption. It also was assumed that oil production could be considered as equivalent gas production by using a conversion factor based on an analysis of the liquid produced. An outer boundary condition of no-flow was assumed. BASIC DATA The data required for the study were reservoir properties, fluid properties and production data. Reservoir properties included the following: thickness h, 179 ft; porosity, 0.194; connate-water saturation, 0.43; gas saturation, 0.57; permeabilities k, 1.0, 0.5, 0.25 and 0.15 md; radius ratios re/rw, 800, 1,000 and 2,000; wellbore radius rw, 5 in.; initial reservoir pressure P, 6,529 psia; and reservoir temperature, 272°F. The selection of permeabilities was based on core-analysis data showing an average air permeability of about 2 md. With a relatively high connate-water saturation, the average effective gas permeability would be about one-half, or less, of the air permeability. The selection of radius ratios was based partly on the observed rapid decline in wellbore pressure during production and partly on the fact that the well was believed to be located in a relatively small fault block. Reservoir fluid properties were obtained from a reservoir fluid study on a recombined sample of gas and condensate. Gas formation-volume factors were calculated from pressure-volume measurements made at reservoir temperature. Gas viscosities were calculated by the method of Carr, et al,G from an analysis of the re-combined sample. These data are shown as a function of pressure in Table 1. The production schedule from the gas-condensate

THE Copperton mill in reality was a sort of proving ground. It was. designed to serve three purposes: (1) to verify the accuracy of the mine sampling by actually treating substantial tonnages of ore, (2) to demonstrate on a reasonably large scale the percentage of the copper in the ore that could be recovered, and (3) to permit the testing of various kinds of machines and devices for crushing and concentrating the ore, so as to guide the engineers in designing the proposed 6000-ton milling plant at Garfield. As a matter of fact, it contained a lot of second-hand equipment sent by MacNeill and Penrose from the works of the United States Reduction &Refining Co. in Colorado; and it was by no means a model installation. Probably it would have been even less efficient, except for the skillful work of George O. Bradley and-Frank-G, Janney, both of whom had worked with Jackling at Mercur. Bradley was a mechanical engineer employed to design the plant. Later he directed the design of the concentrator at Garfield and remained as consulting mechanical engineer to the company until 1915. Janney, who afterwards became manager of mills, a position that he retained until his death in 1916, was a skilled mechanic at the Golden Gate mill at Mercur when Jackling built the plant in 1897. His flair was making machinery do what it was intended to do. He was not an ore-dressing engineer in a technical sense; but Jackling him- self was a specialist in this respect as well as in designing and construction and Janney's mechanical aptitude made him a particularly valuable man in the Utah organization. John McDonald, another of Jackling's Mercur associates, was the first mine superintendent. With practically all of the available funds earmarked for construction of the Copperton

Jan 1, 1933

By H. A. Schwartz

THE spheroidal form of the temper carbon nodules in malleable cast iron and of the graphite mottles of "mottled" cast iron suggests that in both all the graphite in a given mottle or nodule grew from a single center of crystallization or nucleus.? There is, however, a distinct metallo-graphic difference between the mottle and the nodule; the area of mother metal impoverished in combined carbon by the crystallization of graphite in a mottle is practically coextensive with the ramifications of the graphite, while the impoverished zone of metal corresponding to a given nodule is always larger, and usually much larger, than the approximately spherical mass of graphite crystals forming the nodules. Furthermore, the mottle probably is formed during freezing and the nodule almost always at a considerably lower temperature. Assuming that the nodules or mottles can be used as a measure of the existing nuclei, (capable of germination) we propose to record certain observations as to the graphitizing reaction in the light of the number and habits of growth of these nuclei. The present authors, in cooperation with others1-7, have reported upon various cognate aspects of the problem. The fact that temper carbon is a mass of flakes and not a solid granule has become well recog-nized since the introduction of an improved polishing technique8,9,10 for cast iron. The problem of deriving from the circular planar intercepts

Jan 1, 1935

By Wolfram Ruff, H. A. Schwartz

THE spheroidal form of the temper carbon nodules in malleable cast iron and of the graphite mottles of "mottled" cast iron suggests that in both all the graphite in a given mottle or nodule grew from a single center of crystallization or nucleus. ‡ There is, however, a distinct metallographic difference between the mottle and the nodule; the area of mother metal impoverished in combined carbon by the crystallization of graphite in a mottle is practically coextensive with the ramifications of the graphite, while the impoverished zone of metal corresponding to a given nodule is always larger, and usually much larger, than the approximately spherical mass of graphite crystals forming the nodules. Furthermore, the mottle probably is formed during freezing and the nodule almost always at a considerably lower temperature. Assuming that the nodules or mottles can be used as a measure of the existing nuclei (capable of germination) we propose to record certain observations as to the graphitizing reaction in the light of the number and habits of growth of these nuclei. The present authors, in cooperation with others 1-7, have reported upon various cognate aspects of the problem. The fact that temper carbon is a mass of flakes and not a solid granule has become well recognized since the introduction of an improved polishing technique8,9,10 for cast iron. The problem of deriving from the circular planar intercepts

Jan 1, 1935

By Donald L. Katz

The equilibrium constants developed by Souders, Selheimer and Brown1' have been very useful tools in predicting vapor-liquid equilibria during the past five years. Their equilibrium constants were computed by assuming ideal solutions and recognizing the deviations of vapors from ideal gas laws. The same fundamental concepts were used by Lewis and co-workers8 for predicting vaporization phenomena by means of fugacities. The equilibrium constant is defined by equation 1: y\x =K [1] in which K = equilibrium constant, y = mol fraction of a constituent in the vapor phase, x = mol fraction of the constituent in the equilibrium liquid phase. The equilibrium constants are a function of temperature and pressure and might be considered as replacing the ratio of the vapor pressure to the total pressure of Raoult's law. They may be used to compute the vapor or saturation pressure of a liquid or the dew point of a vapor for any pressure or temperature from the analysis of the mixture much as Raoult's law is used. Likewise, the quantity of a mixture that will vaporize at any equilibrium temperature and pressure and the composition of the vapor and liquid may be calculated. The purpose of this paper is to review the available equilibrium data of interest in petroleum production, to discuss the auxiliary physical properties of the hydrocarbons, and to present typical uses of the equilibrium constants. The particular applications chosen are the computation of the results of stage separation of crude oil and natural gas, the calcula: tion of a pressure-volume curve of a crude oil-natural gas mixture, and information gleaned from a consideration of reservoir equilibria.

Jan 1, 1938

By William Hamilton Merritt

A brief paper on this subject (which might readily be made to, fill a volume) is suggested at this time by several considerations, among which may he named the meeting of the Institute on the borders of Ontario, not far from some of the mining undertakings which are hereafter mentioned ; the mineral developments made in the Province quite recently ; the highly creditable collection of most of the minerals of Ontario, for the first time got together, and exhibited this year at Cincinnati, in the Centennial Exhibition of the Ohio Valley ; and, finally, the recent appointment, by the Government of Ontario, of a Commiesion to inquire into, and report upon the mineral resources of the Province, and the best means to be adopted for their development. The Province of Ontario has no Department of Mines or official geologist, to take note, and keep record, of mining and mineralogical development and discovery. Within the last few years some valuable statistics have been collected by Mr. A. Blue, the Deputy Minister of Agriculture; but recently a desire has risen on the part of the people to obtain further knowledge of these subjects, and the government has appointed a Commission, the report of which is expected to be issued in part, if not in completed form, this winter.

Jan 1, 1889

By David K. Mitchell

The mineral pyrite (or marcasite) occurs in coal beds as balls, lenses, veinlets and bands. Several million tons are w-asted annually on the refuse dumps from coal mining and coal-preparation activities. Sporadic attempts have been made in the United States to recover this pyrite for manufacture of sulphuric acid but only a few plants have succeeded in recovering it on a commercial scale. Shipments of pyrite from coal mines during 1941 were reported as follows: Illinois -The Midland Electric Coal Corp., Henry County, shipped 12,026 long tons containing 46 pct sulphur, at its Atkinson preparation plant. Indiana—The Snow Hill Coal Corp. produced pyrite from its Talleydalc preparation plant in Vigo County. Kansas—The Mineral Products Co. produced 3902 long tons of pyrite containing 44 pct of sulphur from picking table and washery refuse. Specifications for coal pyrite for acid manufacture rangc from 5 to 8 pct maximum carhon limit, and 42 to 45 pct minimum sulphur limit. Where flash roasting is practiced, the pyrite must be minus 100-mesh. Most companies give no size specifications. One company specifies 20- mesh or finer. Impurities in pyrite other than carbon may limit its use—for instance, arsenic, probably as arseno pyrite-—even though a concentrate can be made complying with specifications otherwise. Pyrite-recovery Plants Bradford breakers and picking tables may be used to make a marketable product where the pyrite occurs largely in sizes coarser than about I in. To make a grade of concentrate acceptable to sulphuric-acid plants the pyrite must break freely from the accompanying coal and slate, which is a rare condition to find. Most pyrite occurring in coal beds requires crushing and an abrasive type of grinding to free it of entwined coal and slate, followed by mechanical concentration by jigs, tables or other concentrating devices suitable for heavy ores, to make a satisfactory concentrate. At the Snow Hill Coal Corporation's plant in Indiana and at the Midland Electric Corporation's plant at Xtkinson, Ill., the final concentrate is made by a mechanical jig. The most recent and complete mill for concentrating pyritc from coal-mine refuse is that of the Mineral Products Co. at West Mineral, Kans., in which jigs and tables are used. Many regrinds and recirculations of products are necessary to get a concentrate of uniform grade. Trends On the whole, the economics of recovering pyritc from coal is not favorable. Cer-

Jan 1, 1948

By David K. Mitchell

The mineral pyrite (or marcasite) occurs in coal beds as balls, lenses, veinlets and bands. Several million tons are w-asted annually on the refuse dumps from coal mining and coal-preparation activities. Sporadic attempts have been made in the United States to recover this pyrite for manufacture of sulphuric acid but only a few plants have succeeded in recovering it on a commercial scale. Shipments of pyrite from coal mines during 1941 were reported as follows: Illinois -The Midland Electric Coal Corp., Henry County, shipped 12,026 long tons containing 46 pct sulphur, at its Atkinson preparation plant. Indiana—The Snow Hill Coal Corp. produced pyrite from its Talleydalc preparation plant in Vigo County. Kansas—The Mineral Products Co. produced 3902 long tons of pyrite containing 44 pct of sulphur from picking table and washery refuse. Specifications for coal pyrite for acid manufacture rangc from 5 to 8 pct maximum carhon limit, and 42 to 45 pct minimum sulphur limit. Where flash roasting is practiced, the pyrite must be minus 100-mesh. Most companies give no size specifications. One company specifies 20- mesh or finer. Impurities in pyrite other than carbon may limit its use—for instance, arsenic, probably as arseno pyrite-—even though a concentrate can be made complying with specifications otherwise. Pyrite-recovery Plants Bradford breakers and picking tables may be used to make a marketable product where the pyrite occurs largely in sizes coarser than about I in. To make a grade of concentrate acceptable to sulphuric-acid plants the pyrite must break freely from the accompanying coal and slate, which is a rare condition to find. Most pyrite occurring in coal beds requires crushing and an abrasive type of grinding to free it of entwined coal and slate, followed by mechanical concentration by jigs, tables or other concentrating devices suitable for heavy ores, to make a satisfactory concentrate. At the Snow Hill Coal Corporation's plant in Indiana and at the Midland Electric Corporation's plant at Xtkinson, Ill., the final concentrate is made by a mechanical jig. The most recent and complete mill for concentrating pyritc from coal-mine refuse is that of the Mineral Products Co. at West Mineral, Kans., in which jigs and tables are used. Many regrinds and recirculations of products are necessary to get a concentrate of uniform grade. Trends On the whole, the economics of recovering pyritc from coal is not favorable. Cer-

Jan 1, 1948

By L. Kaufman

An equation is derived relating AF a", the difference in free energy between austenite and martensite, to temperature and composition in the iron-chrmnium and iron-chromium -nickel systems. This equation is used to calculate at M,for low-carbon stainless steels. A method for utilizing the energetic equations to calculate M, in stainless steels is suggested. In addition, an explanation is given for the anomalous effect of chromium in forming a y loop concurrent with the lowering of M, in iron-base alloys. It has been demonstrated that the effect of alloying elements on the Ms temperature is strongly dependent upon the influence of the alloying element on the thermodynamic properties of the y and a phases. Notwithstanding the fact that alloying elements affect the Ms temperature by changing the lattice parameters, elastic constants, and inter-facial energies of the a and y phases, their primary role is a thermodynamic one. Thermodynamic analyses have been presented for the iron-carbon'-4 and iron-nickele systems in which equations are derived for Fa and FY, the free energies of the b.c.c. (or b.c.t.) and f.c.c. phases. These equations, which can be obtained with the aid of the equilibrium diagrams, can be used to help predict the Ms vs composition curves. This paper deals with a calculation of the thermodynamic properties necessary to compute AFffLY for iron-chromium alloys. In view of the lack of direct information concerning the activity of chromium in the a and y phases of iron, the regular solution approximationa,g was utilized to bridge the gap. After the necessary equations were derived for the iron-chromium system they were used to calculate AFff— Y for stainless steels by combining them with the previous results obtained for iron-nickel alloys. The results of these calculations offer a means of rationalizing the anomalous behavior of chromium in iron-base alloys. The latter depends on the regular solution approximation in contrast to the explanation offered by zener , which is based upon the "magnetic specific heat effect." THERMODYNAMICS OF THE IRON-CHROMIUM SYSTEM If y represents the atomic fraction of chromium in an iron-chromium alloy, then the free energy of the b.c.c. phase Fa, the f.c.c. phase FY, and the difference in free energy between the f.c.c. and b.c.c. phases of the same composition are given through [3] and Fmy represent the free energies of mixing of the a and y phases; while @era -' and uFea 4Y represent the difference in free energy between f.c.c. and b.c.c. chromium and iron respectively. The values of AEFe0ff4Y in the temperature range between 0" and 900K ( .

Jan 1, 1960

By Erwin Pink

It has been pointed out in a previous publication' that, for a given strain rate, serrated yielding can be observed only at certain test temperatures. Fig. 1 shows the results obtained for polycrystalline and re-crystallized tantalum, Va group in the periodic table of elements, where serrations in the course of deformation match the decrease of total elongation. However, there was no evidence of repeated yielding when testing molybdenum or tungsten, VIa group, at similar conditions, i.e., annealed and air-cooled samples (serrations have been found in quenched tungsten2). The reason for the observed serrations in the stress-strain curve are dissolved impurities on interstitial sites of the lattice. Of importance is the concentration of oxygen, carbon, and nitrogen, which are in random solution. Two methods to calculate the limiting concentrations for serrated yielding have been evaluated and applied with more or less success in estimating the role of carbon and nitrogen in mild steel.3 J. Friedel4 has given two equations, which determine the upper and lower strain rates i for the appearance of serrations at a certain test temperature: W denotes the binding energy of the interstitial atom, -1 ev for refractory metals;5 e' the portion of continuous strain in the discontinuous stress-strain curve, measured as -3x10-3; b the Burgers vector, c the concentration, see Table 11, in atom fractions; p the dislocation density, -2 X l09 per cm2 at an average stress ?, -15 kp per mm2 at medium strain. k stands for the Boltzmann constant, T for the absolute temperature, and D for the diffusion coefficient. Table I. Eq. [1], for a given strain rate, determines the minimum temperature, and [2] the maximum temperature, below or above which no "dynamic strain-aging" can occur. According to the calculation, which uses the given quantities, serrations should only be observed within the cross-hatched temperature regions in Fig. 1, and this is in good agreement with the experiment. The effect of varying impurity concentrations is indicated schematically in Fig. 2: only Eq. [I] is affected. Setting Eq. [I] equal to [2], one can calculate the limiting concentration cmin which first causes serrations. The second method uses the aging condition by Cottrell and Bilby6,7 \£i/ \ n. a / where n(t) is the number of impurity atoms, which segregate at a dislocation line of 1 cm length after the time t (t in our case is the time between "repeated yield drops" in the serrated curve, or, the time until the above mentioned strain E' is obtained, i.e., -1 sec for medium intensity or serrations). no is the number of impurity atoms in 1 cm3 and A = Qb4ep[(1 + v)/(l - v)], derived from equations given in Ref. 7. p is the shear modulus, 18,800 kp per mm2 for tantalum, 12,200 for molybdenum, AE the tetragonal distortion caused by interstitial impurities in a bcc lattice (approximately 0.4 for both tantalum and molybdenum) and v Poisson's ratio (0.35 for tantalum, 0.30 for molybdenum). From Eq. [3], combined with the condition for a pronounced yield drop as formulated by cottrel16 (at least one impurity atom segregated at the unit length of a dislocation), again a no min, and hence the concentration Cmin can be calculated.

Jan 1, 1970

By William Campbell

In the olden days the making of alloys was an art, and the secrete of the craft were jealously guarded. To-day it has become a science, though the old ideas in regard to the secrets and formulæ are dying hard. Modern progress may be attributed to scientific research and testing of materials, and the substitution of chemical analytical control for " cook-book " methods. We know to-day a great deal about the constitution of alloys, the reasons of the change in structure and physical properties by heat and mechanical treatment; and this new knowledge is due for the most part to the physical chemist and metallographist, who have shown us that alloys are simply solutions, and as such obey definite laws. The application of the theories of solutions and of the phase rule tells us what ought to happen when our alloys are in a state of equilibrium. The working-out of the temperature-composition diagram (thermal diagram) for numerous alloys has thrown a light on the structure and constitution which has dispelled most of the darkness in which we labored before the advent of metallography. The work began with the determination of the freezing-point curves, and here the research of Le Chatelier,' Gautier,2 Roberts-Austen,3 and the other pioneers broke the ground. Next the physical chemists, Tammann,4 Rooseboom,5 Bancroft,6 and their colleagues, showed us the complete thermal diagram and its meaning. Then the researches of Heycock and Neville7 on the Cu-Sn series and of Shepherd 8 on the

Jan 1, 1913

By Frank E. Woolley, John F. Elliott

Mathematical models of the transient thermal behavior of a high-temperature solution calorimeter1-3 have been developed. The thermal behavior of the calorimeter is appoxirrzated by linear lumped-parameter models, and hence is described by sets of linear ordinary differential equations with constant coefficients The response of the models to various inputs is shown to agree with the response of the real system. Application of the modeling to experimental design and analysis of data illustrates the usefulness of simple models of complex systems. The early eperiments1,2 with the high-temperature solution calorimeter indicated that the change in the temperature of the bath resulting from the addition of a solute sample to the bath involved not only the direct effect due to the solution process but also possibly a secondary effect arising from the change in coupling between the bath and the induction heating coil. Consequently, an extensive analysis of the calorimeter was carried out, and models of the transient thermal processes of the instrument were developed to aid in improving the design and interpreting the behavior of the system. This paper describes the dynamic modeling; the use of it in treating experimental results has been reported earlier.3 The high-temperature solution calorimeter was constructed to measure directly the partial molar heats of solution of solute elements in a variety of liquid metal solvents.1-3 The calorimeter consists of an induction-heated liquid metal bath into which small samples of a solute element can be dropped. The bath temperature is recorded continuously, and the change in the measured bath temperature with time, dTm = f(t), resulting from the solute addition are the raw data from which the enthalpy change caused by the addition is determined. To extract the rmodynamic results from the data, the temperature change must be compared with that resulting from calibration additions of known enthalpy change. Accordingly, it is necessary to understand the transient thermal processes arising as a result of the addition to the bath. Neither modeling nor experimentation alone could provide the required insight into the working of the calorimeter. The alternate use of both methods in conjunction greatly assisted the design of the equipment and experiments, and the interpretation of the data. THE PHYSICAL CHARACTER OF THE SYSTEM The essential parts of the calorimeter, Fig. 1, for model studies are the thermocouple, the liquid metal bath and the surrounding refractories. The system is the solvent metal bath and those refractories around it which undergo a temperature change as a result of an addition to the bath, and which determine the way the temperature of the bath responds to an input. The inputs are the combined transient thermal effects arising when an addition is made to the bath. They include the thermal effects of the addition itself and the results of changed coupling between the bath and the induction coil. The response is the variation in the measured bath temperature, dTm(t) = Tm(t) - Tm(O), from an initial steady state resulting from the inputs. It was assumed in this study that the physical properties of the various elements of the system are independent of the inputs and time, although these properties may vary as the result of changes in the composition and size of the bath during a series of additions. This separation of inputs and the system is equivalent to assuming that the system is linear, i.e., that its behavior can be described by linear differential equations with constant coefficients. Linear behavior can be expected whenever the departure of each portion of the system from its steady-state condition is small enough to cause negligible changes in the thermal properties of the materials and in the various heat-transfer coefficients. Radiative heat transfer is important in this system, so the assumption of linearity should be valid only for small temperature deviations. Several conclusions were drawn from operation of the calorimeter in earlier experimental studies: 1) Radiative heat transport from the top of the bath is a significant portion of the total heat lost from the bath. However, for small changes in the bath temperature the change in transport by this path could be assumed to be proportional to the change in the bath temperature. 2) A very small portion of the heat input is lost through the thermocouple to its water-cooled holder. The thermal resistance and thermal capacity of the thermocouple protection tube are small, so the temperature of the thermocouple should follow closely that of the bath. 3) The remainder of the total heat lost from the bath will pass by conduction through the crucible to, and through, the other refractories, eventually being absorbed by the water-cooled induction coil or by the water-cooled sides and bottom of the enclosure. 4) The thermal resistance between the bath and crucible is very small. Thus the thermal capacity of the crucible will affect the temperature of the bath very soon after an addition of heat to the bath. 5) The thermal resistance between the crucible and the silica sleeve is large, especially if a radiation shield is placed in the gap. The effect of the thermal capacity of the sleeve thus will be significant only at longer times. The thermal resistance through the packing below the crucible also is large, so the packing and the silica sleeve will have similar effects on the behavior of the system. 6) A large temperature drop exists across the gap containing the water-cooled induction coil. Thus for relatively small changes in the thermal input to the bath, the refractories beyond the sleeve

Jan 1, 1970

By T. L. Joseph

A study is made of the amount of solution loss necessary to maintain the reducing power of the gas stream in the blast furnace. Curves are presented to show the effect of solution loss, moisture in the blast, and carbon dioxide from the flux on the nitrogen in the THE blast furnace process is based upon the counter flow of gases and solids. As CO, N, and H, flow upward from the combustion zones, oxygen is acquired from the descending charge. In the reduction of iron ore, a portion of CO is converted to CO, and similarly a portion of H, is converted to water vapor. The inert N,, from the air used in combustion, passes through the furnace with little change. Its concentration is, however, decreased by reactions such as direct reduction and solution loss which increase the volume of CO in the gas stream; Similarly the evolution of C02 from the limestone also dilutes the nitrogen. he purpose of this paper is to illustrate an idealized and simplified method of studying furnace processes by following changes in the gas phase between the tuyeres and the stock line. A more specific objective is to evaluate the effect of solution loss upon the CO/CO, ratio in that region of the furnace in which FeO is reduced to metallic iron. This last stage of reduction requires relatively high CO/CO, ratios to exceed the equilibrium requirements. Solution loss or the reaction of COZ with C to produce CO has a pronounced effect upon the ratio of the carbon gases and is necessary to maintain the reducing power of the gas stream if the consumption of coke is low. It will be explained later that indirect reduction followed by solution loss is stoichiometri-cally equivalent to direct reduction. Calculations will be made to show how the composition of the gas in the region of the mantle varies with coke consumption and with different amounts of solution loss. Such calculations show the condi- tions under which gasification of C with oxygen in the charge is needed to maintain the reducing power of the gas. Equilibrium conditions for FeO and the carbon gases will be used to establish a minimum CO/CO, ratio below which reduction will not proceed. The manner in which carbon is gasified in the furnace, at the tuyeres with oxygen in the air or above the tuyeres with oxygen from the charge, will be a basic feature of all calculations. It is hoped that greater interest in the use of top gas as a means of following furnace reactions will be aroused by the calculations and generalized curves which are presented. Bosh Gas Equations Subject to some alteration by the amount of moisture in the blast and by direct reduction, the gas leaving the combustion zones of a blast furnace has a definite composition. Because a relatively small amount of oxygen is acquired from the charge in the bosh and in the crucible, the products of combustion undergo little change until they enter the shaft. Bosh gas is, therefore, essentially the products of combustion of carbon in a deep fuel bed with air. The relative weights and volumes of reacting materials are expressed by the equation: 2~ + 0? + 3.76N, = 2C0 $ 3.76NZ [I] Air Bosh Gas -air = 79% N? = 3.76) 21% 02 Starting with a known mass of C consumed largely by air, Eq 1, but partly by water vapor, Eq 2, the volume and C + H2O = CO $ H? [21 composition of the bosh gas can be calculated and subsequent changes followed as reactions take place at various elevations in the furnace. As shown in

Jan 1, 1952

By Lawrence S. Darken

STAFF: Editor, Gerhord Derge Cornegie Institute of Technology Schenley Pork Pittsburgh 13, Pa. Editorial Assistant, M. A. Redmerski Production Editor, Otto T. Johnson THE METALLURGICAL SOCIETY: J. S. Smart, Jr., President; C. C. Long, Past-President; K. L. Fetters, Vice President; T. D. Jones, Treasurer; R. W. Shearman, Secretary. PUBLICATIONS COMMITTEE OF THE METALLURGICAL SOCIETY: C. L. McCabe, Chairman; H. L. Bishop, K. L. Fetters, J. J. Harwood, W. A. Krivsky, F. C. Longenberg, R. E. Lund, W. 0. Philbrook, L. L. Seigle, R. L. Smith, and F. L. Vogal. Review and selection of manuscripts is under the supervision of the following Publications Committees: Extractive Metallurgy Division: C. L. McCabe, Chairman; R. E. Lund, Vice-chairman; D. H. Baker, Jr., W. A. Krivsky, R. A. Lewis, Paul Qusneau, A. S. Russell, R, E. Shinkoskey, W. W. Stephens, and M. E. Wadsworth. Institute of Metals Division: L. L.Seigle, Chair-man, P. A. Back, J. J. Backer, l.l. Bessen, J. 0. Betterton, Jr., J. 8. Clark, R. S. Davis, A. E. Dwight, R. W. Fountain, R. B. Gordon, C. D. Graham, Jr., R. W. Guard, J. C. Gurland, E. L. Harmon, J. J. Hager, A. A. Hendrickson, M. Harmon, J. N. Hobstatter, H. H. Johnson, Jr., H. 5. Kolish, R. R. Nash, T. N. Rhodin, R. H. Richman, W. Rostokar, P. G. Shawmon, C. W. Spencer, L. Slifkin, J. K. Stanley, C. D. Starr, R. Steinitz, and H. C. Stumpf. Iron and Steel Division: H. L. Bishop, Chairman; 2. E. Boni, F. C. Langenberg, J. B. Wagstaff, and T. V. Wainwright. Published bimonthly by the American Institute of Mining, Metallurgical, and Petroleum Engineers, Inc., 29 West 39th Street, New York 18, N. Y. Telephone: PEnnsylvania 6-9220. Subscription PetroleumEngineers,$30 per year for non-AIME members in United States and North, South, and Central America; 332, foreign; $5 for AlME members. Single copies, $4.00, single copies foreign, $5.00. The AlME is not responsible for any statement mode or opinion expressed in its publications . . . Copyright 1%1 by the American Institute of Mining, Metallurgical, and Petroleum Engineers, Inc. . . . Registered cable address; AIME New York ..Indexed in Engineering Index, Industrial Arts Index, and Chemical Abstracts.. .Second class postage paid at New York, N. Y., and Ann Arbor, Mich. Volume 221, Number 4.

Jan 1, 1962

By Edward H. Robie

EIGHT years ago in this department we had a column on the daiquiri cocktail, which, as we pointed out at that time, was invented by mining engineers. Since then we have said little or nothing about alcoholic drinks, but having set the precedent perhaps we may now pass along a little information about martinis. We have observed that as an aperitif the martini has become increasingly popular with men, so much so that we always feel a bit peculiar if we order something else. And the call is not for just a martini either; it must be a dry martini, and some of our friends are extremely particular on this point. The matter has achieved such importance that The New York Times devoted half a page to it a few Sundays back. Several recipes are given which may be useful to some of our members when entertaining visitors from this part of the country: The standard martini is said to be half gin, half Italian vermouth, two dashes orange bitters, and a pickled onion and twist of lemon peel. This progresses to the very dry martini, which calls for five parts gin and one part vermouth with an onion, olive, or lemon peel. Several alternatives are offered for the manufacture of the very, very dry martini. One is to pour a little vermouth into the mixing glass and pour it out again before mixing gin and ice. A second way is to put the ice in a sieve, hold it over the sink, and pour a little vermouth over it; then put the ice in the mixing glass. A third is to let the draft from an electric fan blow across an open vermouth bottle toward the mixing glass. Then the ultimate is the awfully dry martini, obtained by setting the vermouth bottle next to the mixing glass and turning the bottle slowly so that the label is exposed to the gin for perhaps a second. The Times notes that Mr. Martini himself personally prefers vermouth and soda.

Jan 1, 1952

By Earl S. Greiner, Eric R. Jette

The constitution of the iron-silicon alloys containing from 14 to 33.4 weight per cent silicon has been studied by a number of investigators, whose results have been reviewed in a monograph published by the Engineering Foundation1. In these previous researches, the solubility of silicon in iron was determined by thermal, microscopic or magnetic methods and the compound Fe3Si2 was reported to be stable at temperatures from 0 to approximately 1030' C. This paper describes the results of an X-ray investigation of the solubility of silicon in iron and gives data which prove that the Fe3Si2 phase is unstable below 825" C. In order to obtain the present results on the solubility of silicon in iron, the authors have utilized the lattice constants of the alpha-phase iron-silicon alloys given in a previous publication2, and also some additional data which will subsequently be described. Materials and Methods The alloys used in this investigation are listed in Table 1. They were prepared in 30 or 50-gram ingots by melting specially purified silicon and electrolytic or carbonyl iron in alundum crucibles under a pressure of less than 1 mm. of mercury. The purity of the silicon, electrolytic and carbonyl irons has been given by the authors in previous publications2, 5. The ingots were homogenized in vacuo at approximately 1000" C. for at least 96 hr. and then slowly cooled. The homogenized specimens were ground to powder in a hardened steel mortar. This powder was passed through a 150 or 200-mesh screen and sealed in evacuated quartz, Jena or Pyrex glass tubes and heat-treated as noted in the various tables. Specimens 23A, 24B, 33 and 35 and the specimens heated for one month at 700" C. were quenched by quickly removing the tubes from a horizontal tubular furnace and dropping them into water. The other specimens were quenched by means of a drop-quenching device, which has been described previously3.

Jan 1, 1937

By C. H. Li, S. H. Bendel, J. A. Sartell

Small additions of Mg to a 62 Cu-38 Ni alloy reduce the oxidation yates drastically between 500° and 850°C. Metallographic examination of the oxide scale disclosed a continuous network, identified by X-ray diffraction as (Ni, Mg)O solid solutions, located along the former alloy grain boundaries. This network acts as a barrier to diffusion. It is thought to result from the segregation of mg to the alloy grain boundaries. Oxidation occurs preferentially in these regions because of the presence of this more active element. To substantiate these ideas, homogenized alloys ulere oxidized. No oxide network was produced and oxidation rates were the same as for the parent 62 Cu-38 Ni alloy. An interesting aspect of the oxidation of the 62 Cu-38 Ni alloy previously described by the present authors, is the relationship between oxidation kinetics and changes in the microstructure of the oxide.' The oxide scale on these alloys consists of two layers: a thin layer of CuO outside a much thicker layer of Cu20 in which particles of NiO are dispersed. In addition, immediately beneath the oxide-metal interface there is an internally oxidized zone in which NiO particles are dispersed in a copper-rich alloy matrix. Experimental observations on the growth of the individual layers under isothermal conditions have led to the conclusion that the rate-determining step in the oxidation of this alloy is the diffusion of anions through the Cu20 layer. It has been found that this diffusion rate is increased by the recrystallization of the outer CuO layer. Willardson and Gorton have reported that the addition of magnesium to the 62 Cu-38 Ni alloy causes a decrease in the oxidation rate.2 The present investigation was undertaken in order to determine the mechanism by which magnesium additions improve the oxidation resistance of the 62 Cu-38 Ni alloy. EXPERIMENTAL PROCEDURE Fifty-gram ingots of Cu-Ni alloys were prepared by vacuum melting materials of 99.999+ pct purity. These ingots were then remelted in a He atmos- phere and magnesium added in the form of a previously prepared and analyzed Cu-Mg master alloy. This procedure was necessitated by the high reactivity of pure magnesium at high temperatures. The proportion of the alloying elements was adjusted so as to obtain alloys with the desired magnesium content while maintaining a Cu-Ni ratio of 62 :38 by weight. (All alloy compositions will be given in weight percent.) Since it was considered desirable to study the oxidation of only single-phase, solid-solution alloys, it was necessary first to determine the approximate limit of solid solubility of magnesium in the 62 Cu-38 Ni alloy. A series of melts with various magnesium contents was made, homogenized by annealing for 48 hr at 975OC, then cooled in the furnace. Metallographic samples were prepared, etched in a potassium dichromate solution and examined at XI500 for the presence of a second phase. The alloys were then chemically analyzed and it was found that the approximate limit of solid solubility of magnesium in the 62 Cu-38 Ni alloy was 1 pct Mg. The ingots containing less than 1 pct Mg (0.19, 0.29, 0.58, and 0.87 pct Mg) were cold-rolled down to thicknesses suitable for oxidation experiments and specimens for gravimetric studies were cut to dimensions of 3.0 by 1.5 by 0.05 cm, specimens for inert marker studies, 0.6 by 0.6 by 0.17 cm. These were heat-treated in a protective atmosphere at a temperature (975" C) well above the oxidation temperatures and furnace-cooled to insure structural stability of the alloy during oxidation. Immediately prior to oxidation, the specimens were chemically cleaned by immersion for 1 min in a solution of 2 parts nitric acid, 1 part acetic

Jan 1, 1961

By George E. Beavers

Under the title used for this paper, the converter practice of the Tennessee Copper Co. is described in the TRansactions of the Institute.l That article contains the following statement: "The limit in capacity of the Peirce-Smith converter in terms of concentrate has not been reached but approximately as high as one ton of concentrate has been handled for each ton of liquid matte charged and good working conditions have been maintained." It is not the intention of the writer to review the conditions leading up to the introduction of a Peirce-Smith converter at Copperhill, but, instead, to offer later data to show how satisfactory the installation has been. Converter Campaign, June 15, 1930 to August 23, 1931 Blowing time (converter days)............... 349.90 Number of blows........................... 427.00 Tons matte charged........................ 38,335.00 Copper assay of matte, per cent.............. 13.67 Tons blister copper made.................... 9,246.00 Time to blow one ton of copper, minutes...... 54.50 Air used per minute, cu. ft................... 12,500.00 Blast pressure, lb........................... 12.80 Air per ton copper, cu. ft.................... 681,330.00 Air per ton iron, cu. ft....................... 199,040.00 Tons flux used............................. 11,845.00 Tons flux per ton matte..................... 0.31 Tons flux per ton copper.................... 1.28 Tons flux per ton iron....................... 0.37 Tons concentrate treated.................... 28,960.00 Copper assay of concentrate, per cent......... 17.13 Tons concentrate per ton matte.............. 0.76 Tons material charged per converter day...... 235.70

Jan 1, 1934

By J. H. Lightfoot, Stuart E. Buckley

DEEP drilling has led to the development of numerous pools which yield products consisting predominantly of natural gas accompanied by high-gravity water-white or slightly straw-colored liquid. Yields of this "condensate" or "distillate," as it is commonly called, vary from traces to over a hundred barrels per million cubic feet of gas. The fact that this liquid exists in the reservoir as an integral part of the gas phase and is formed by condensation attending simultaneous reduction of pressure and temperature after its entry into the well was apparently first recognized about 1932.1 The nature of the physical phenomena involved in this condensation was clarified for the petroleum industry in 1933 by a report by Sage, Schaafsma, and Lacey2 of the results of an experimental investigation of the behavior of mixtures of methane and propane. Search of the literature has revealed that the condensation of liquid from gaseous mixtures by isothermal pressure reduction is not a new discovery. Kuenen was led by theoretical considerations to expect such a phenomenon, and in 1892 reported3 the results of experimental verification of his deductions. This was followed by a considerable amount of work on the phase relations of mixtures by Kuenen4,5,6 and by others,7,8 mostly in European universities. Since 1930 investigators in this country have reported experimental data on the properties and behavior of a number of hydrocarbon systems.9-17 It is now generally known that in the presence of natural gas the solubility in the gas phase of the heavier and normally liquid hydrocarbons increases rapidly with increasing pressure above 500 to 1000 lb. per sq. in. Thus, at pressures of 2500 lb. per sq. in. or higher a gas may contain appreciable quantities of hydrocarbons with normal boiling points as high as 600°F., which in turn may be condensed from the gas phase by the reduction of their solubility therein attending pressure reduction. The combined effects of pressure and temperature on condensation and vaporization of mixtures are most readily visualized by the study of phase diagrams. Since the construction, application, and interpretation of such diagrams have been described fully in readily available literature, 18-20 this paper will not review the theoretical aspects of this subject. While the behavior of gas-distillate systems is now clearly understood, at least in a qualitative way, most of the published work deals with the results of laboratory studies of binary systems or of synthetic mixtures of hydrocarbons, and although several systematic investigations of the effects of separator pressure and temperature on the yields of liquid product from gas-distillate wells have been made21.22 it appears that in all cases the limitations have precluded the construction of even approximately complete phase diagrams for naturally occurring systems. There has

Jan 1, 1941