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The Fifth Annual National Exposition of Chemical Industries will be held in Chicago at the Coliseum and First Regiment Armory during the week of Sept. 22. It was decided to hold the exposition in Chicago, this year, for two reasons: The U. S. Army commandeered the Grand Central Palace Exposition Building in New York, immediately on the close of the last exposition and converted it into a receiving hospital, the use for which is now, happily, nearly ceased; and it was thought that the Exposition would succeed in bringing about a greater development of chemical industries in the Chicago district and the adjoining states. The Advisory Committee of the Exposition consists of: Charles H. Herty, Chairman, Editor Journal of Industrial and Engineering Chemistry; Raymond F. Bacon, Director, Mellon Institute; L. H. Baekeland, Member Naval Consulting Board; W. D. Bancroft, President, American Electrochemical Society; Henry B. Faber, Industrial Filtration Corpn.; Ellwood Hendrick, President, The Chemists' Club; Bernard C. Hesse, Chemist, General Chemical Co.; A. D. Little, President, Arthur D. Little Inc.; Wm. H. Nichols, President, American Chemical Society; R. P. Perry, Vice-president, The Barrett Co.; H. C. Parmelee, Editor Chemical &; Metallurgical Engineering, G. W. Thompson, President, American Institute of Chemical Engineers; T. B. Wagner, Vice-president, U. S. Food Products Corpn.; M. C. Whitaker, President, U. S. Industrial Chemical Co.; and Charles F. Roth and Fred W. Payne. There is also added a special Chicago committee consisting of L. V. Redman, W. D. Richardson, A. V. H. Mory, Carl S. Minor, F. W. Willard, and Wm. Hoskins; the managers, as in the past, are Charles F. Roth and Fred W. Payne.

Jan 6, 1919

By C. F. Gates, H. J. Ramey

Recent literature shows that pronounced increases in oil recovery can result from the use of miscible systems in recovery operations. This literature also points out certain problems associated with maintaining miscibility, e.g., compositional changes resulting from alterations in pressure and temperature or from zone dilution. The purpose of the work described in this paper was to study miscible zones for possible use in waterflooding operations and examine the manner in which these zones break down. The fluid system selected for study was oil—terNary-butyl alcohol—water, which, because of its narrow range of miscibility, is ideally suited for investigation in short, linear systems. The laboratory-observed effects on oil recovery and zone stability of a number of variables are reported. Among the variables investigated were miscible zone size, viscosity ratio, length of travel, and interstitial water saturation. Results obtained with this system show that interstitial water adversely affects recoveries because of premature phase break. Salinity of the water further aggravates Ais situation. This effect, although pronounced for this system, would probably occur in any fluid system containing an alcohol. As expected, viscosity ratio has a decided influence on recovery efficiency and zone stability. The results also show that use of a suficieno zone size, even though the system has poor phase characteristics, yields higher oil recoveries than are obtained from straight water floods. Length-of-travel studies showed a square root relationship between zone growth and path length for favorable viscosity ratios. No such clear-cut dependence was observed under unfavorable viscosity conditions. INTRODUCTION As new oil reserves have become increasingly more difficult and expensive to find and develop, the oil industry has devoted more and more time and money to finding more efficient methods for exploiting known reserves. The increasing number of water floods is one manifestation of this attempt to improve oil recoveries from existing fields. Since considerable quantities of oil are by-passed by the waterflood process, it is only a partial answer to the problem. The search continues. Of late, several proposed processes have received the critical attention of investigators. Among them are gas drives, in situ thermal reactions and miscible displacements. The present paper is concerned with miscible displacements in conjunction with water floods. Recent literature1- reports that, under proper conditions, oil recoveries from LPG floods, propane sweeps, condensing gas drives, etc., approach 100 per cent of the oil in place in the region contacted by the flood. This literature also points out some of the problems associated with these processes, one of the main ones being the maintaining of miscibility. Compositional changes resulting from pressure and temperature variations or from zone dilution give rise to phase breaks and loss of miscibility, the conclusion being that, after loss of miscibility, much of the beneficial influence of the zone is lost.' The results of the subject work show that, if sufficient zone material exists ahead of a phase interface, considerable improvement in oil recovery can be obtained. In order to increase our understanding of the behavior of miscible systems as they degenerate into immiscible processes, a laboratory system having rapid deterioration characteristics was needed. The fluid system oil—ter tiary-butyl alcohol (TBA)—water possesses these properties and was the system investigated. Although these fluids are different from those generally employed in miscible studies, an understanding of their behavior can be applied to other miscible slug processes. However, under conditions where mass transfer maintains a miscible system through multiple contacts, e.g., enriched gas drive or high pressure gas drive, the observed behavior of alcohol zones would not be applicable. The variables of interest were zone size, viscosity ratio, length of travel, initial water saturation, rate and core characteristics. Of these, only the effects of the following on zone breakdown and oil recovery are reported in detail: (1) zone size, (2) path length, and (3) initial water saturation. The effects of viscosity ratio and core characteristics are discussed in conjunction with other variables. Although initial water saturation is not envisioned as important to the phase behavior of hydrocarbon systems, e.g., LPG, it is important in alcohol systems. The properties of the fluids and cores used are given in Table 1. As previously stated, this fluid system was selected because of the limited range of miscibility

By William Hurst

Muskat's depletion performance equation is here derived considering the expansion behavior of the reservoir hydrocarbon system and a simple fractional-flow equation. This nietkod of derivation leads logically to the two extensions that follow. The first of these is concerned with gravity segregation in a depletion-drive reservoir. The second is concerned with including an empirically determined tern? for water influx in the performance equation. The more general equation for gravity segregation when there is a primary gas cap and empirically-determined water influx is stated for completeness. These equations have been found useful in reservoir performance calculations in Eastern Venezuela. A discussion on the methods of solving these equations follows, and considers firstly the effect of taking finite intervals in the numerical integration, and sec-ondly, methods of incorporating the time functions involved in segregation in with the expansion behavior. The paper concludes with a brief general discussion on further extensions to the depletion performance equation. INTRODUCTION The two fundamental sources of energy by which oil is produced from a reservoir result from pressure depletion inside the boundaries of the reservoir and fluid encroachment across the boundaries of the reservoir. The wells in either case form low pressure outlets through which oil and gas may be produced by the expansive force of the reservoir fluids and associated encroaching fluids. When the reservoir pressure is higher than the bubble-point pressure of the oil, so that there is no free gas in the reservoir, these expansive forces are the only ones available for the production of oil. However, when the reservoir pressure is less than the bubble-point pressure of the oil, free gas is vaporized as the pressure falls. With both oil and free-gas phases present, the additional forces of gravity and capillarity may operate on the gas-oil system, as they have previously operated on the oil-interstitial water system. Gravity tends to segregate the free gas from the oil due to their density difference. Capillarity opposes and eventually balances gravity as the more extreme free gas and oil saturations are reached, preventing the independent move- ment of free gas until it is above a certain saturation, and the independent movement of oil when it is below a certain saturation. The type of depletion performance equation chosen for predicting the future performance of a reservoir depends on the amount of past history available. When the reservoir is somewhere past the halfway mark in depletion, some form of decline curve is often used. With less past history, material balance equations which incorporate empirical factors based on the past performance are often used. When, however, the amount of past history is small, the Muskat depletion performance equation will usually be used. The distinguishing feature of this type of equation is that empirical factors based on the over-all or macroscopic reservoir behavior are almost or entirely absent. Each parameter affecting the reservoir performance is ascribed an independent set of values based on measurements made on laboratory samples; that is, incorporating microscopic empirical factors. In establishing Muskat-type depletion performance equations, it is necessary to consider the reservoir as consisting of a number of associated blocks, in each of which the saturations and pressures may be considered uniform, and in each of which all substances have uniform pressure-volume characteristics. Thus, a primary gas cap can usually be considered as one block and an aquifer as another. Gravity segregation may be negligible for practical purposes when the rock and oil properties are adverse and/or the dip or thickness of the reservoir is too small. In this case the whole oil leg may be considered as one block, except in very large reservoirs. In very large reservoirs the fluid and rock properties may vary enough, particularly in the dip direction, for it to be necessary to divide the oil leg into a number of blocks, in each of which the relevant quantities may be considered uniform. When gravity segregation of the oil and free gas is not negligible, it is necessary to consider the space occupied by the initial oil leg as divided into two blocks, a secondary gas cap and an effective oil leg, in each of which saturations may be considered to be uniform. The total volume of these two blocks is thus constant, but the secondary gas cap grows continuously at the expense of the oil leg. Muskat1 derived depletion performance equations for the basic case of an oil reservoir with closed boundaries and without segregation, and for the case of an oil

By F. H. Spedding

From thermal, metallographic, X-ray, dilatometric, and electrical resistance data a Phase diagram is proposed for the La-C system. TWO compounds are formed, body-centered-cubic Laz C, having a large range of solid solubility, and body -centered-tetrago-nal Lac2 . The former melts incongruently at 1415°C, and at 25OC it has an electrical resistivity about two and one-half times that of bure lanthanum. The latter melts congruently at 2356°C and at room temperature has an electrical resistivity about equal to that of bure lanthanum. The addition of carbon to La lowers the melting Point of La about 115' C, raises the higher (B-Y) transfomzation about 10" C and appears to have no effect on the lower ( a-0) transformation. The addition of carbon appears to lower the oxygen content in the metal and improve the machinability; however, the tendency to oxidize is increased. Two eutectics are formed in the system. The first occurs at 2.2 pct C and melts at 806" C, while the second occurs between Lac2 and C with a melting point of2271"C. PETTERSSON,~ in 1895, was the first to prepare lanthanum dicarbide by heating the oxide with carbon in an electric furnace. He showed that this yellow, crystalline but brittle carbide had a formula corresponding to LaCz and that it reacted with water with the evolution of hydrogen and hydrocarbons and the formation of the hydrated oxide. Von Stackelberg's investigation of the crystal structure of the MCz carbides showed that the lanthanum dicarbide is of the tetragonal calcium dicarbide-type structure. He found the lattice parameters were a = 3.92 and c = 6.56.* Bredig,' using X-ray methods, showed that the room-temperature form of lanthanum dicarbide is stable up to 1750 °C and that above this temperature it transformsoto a face-centered-cubic structure, a = 6.0 ± O.lA, probably of the pyrite, FeSzstruc- typ? The reported discovery of two new cerium carbides, suggested that a corresponding face-centered-cubic lanthanum monocarbide and a body-centered-cubic lanthanum sesquicarbide might exist. This paper describes work on the lanthanum-carbon system which is part of a study of the rare earth metals and their alloys now in progress in this laboratory. EXPERIMENTAL PROCEDURE Source of Materials—The lanthanum metal was prepared at the Ames Laboratory by metal1o;hermic reduction of the fluoride with calcium metal. The results of spectrographic and chemical analysis indicated the presence of 0.01 pct Y, Al, Cu, Fe, and Si; 0.02 pct Ca and Mg; 230 ppm C; and 115 ppm N. The carbon used was of two types: 1) "National Special Graphite Spectroscopic Electrodes," and

Jan 1, 1960

By D. J. McAdam, G. W. Geil, D. H. Woodard, W. D. Jenkins

INTRODUCTION IN a series of papers, the authors and their associates have shown the influence of four important factors on the technical cohesion limit.3-16 By "technical cohesion limit" is meant the technically determinable resistance to fracture. The four important factors are the stress system, plastic deformation, temperature, and the strain rate. The influence of any one of these factors may be represented by a curve of cohesion limits, although only one point on the curve may represent actual fracture. As the term "fracture stress" has come into general use with the same significance as technical cohesion limit, the authors are here using the shorter term with the understanding that it does not always refer to actual fracture. The previously-mentioned series of papers contains ample evidence that plastic extension increases the fracture stress of any polycrystalline metal. In a recent paper of the series,15 moreover, it has been shown that the effect of plastic deformation on the fracture stress cannot be attributed to reorientation of internal flaws but is a work-strengthening effect, similar to the effect on the flow stress.[$] Plastic deformation of steels, however, may not only cause work hardening, but may result in additional hardness because of strain aging. Some evidence that strain aging affects not only the flow stress but also the fracture stress, has been presented in a recent paper.14 In that paper, a study is made of the influence of plastic deformation at room temperature on the fracture stress of ferritic steels at -188°C. The study is based on the results of two-stage tests. In the first stage each specimen was extended a predetermined amount at room temperature; in the second stage the specimen was tested to fracture in liquid air. In most of these tests the specimen remained in the testing machine during the interval of about one hour between the first and second stages, it was kept under a tensile load of about 500 lb and maintained at liquid air temperature. The effects of various amounts of plastic deformation on the fracture stress were thus determined and curves were obtained to represent the influence of plastic deformation at room temperature on the stress at brittle fracture in liquid air.14 Two specimens of ingot iron, however, were removed from the testing machine at the end of the first stage and several days elapsed before the specimens were tested to fracture in liquid air. As shown in Fig 8 of that paper," the points representing fracture of these specimens are far above the curve representing the results of two-stage tests without a rest interval between the stages. The evidence thus indicated that the fracture stresses of these two specimens were increased not only by the work-strengthening, but also by strain aging. The present paper presents results of further investigation of the influence of strain aging on the fracture stress. There are two kinds of aging of steel.

Jan 1, 1948

By A. L. Labbe, J. J. Donoso

Hard-won experience in the operation of smelting plants has pointed the way to the most efficient design and economical operation of a baghouse system for recovery of metallurgical fumes. This paper treats of the principles governing bag filtration and the factors most directly affecting bag life. Methods for temperature conditioning of smoke previous to filtering are described and the evolution of bag shaking mechanisms is brought up to date. THE history of the bag filter process, as applied to the recovery of metallurgical fumes, goes back to 1876, when the Lone Elms Works of Joplin, Missouri, erected a two-section baghouse for filtering fumes from Scotch Hearth lead smelting furnaces. By modern standards this installation was crude, but the basic operating principles remain essentially the same today. Great progress, however, has been made in the mechanization of the process by the use of devices for mechanically and automatically cleaning bags of adhering fumes, in the use of instruments for recording and controlling the variables affecting baghouse operation, and in the conditioning of the smoke previous to filtering. The principles governing bag filtration of fumes were investigated in 1929 by A. J. Seitz of American Smelting and Refining Co., and he arrived at the following conclusions: 1. Pressure on Bags at Constant Volume: The rate at which the pressure builds up in a bag with constant volume is directly proportional to the time, with the same kind of smoke and dust burden (fig. 1). 2. Pressure on Bags for Change in Volume: With equivalent smoke and dust burden, the bag pressure increases with the square of the increase in volume. The rapid increase in pressure is due to (1) increase in volume, (2) increase in amount of dust on the bag, and (3) more intense packing of the dust at higher velocity (fig. 2). 3. Effect of Intensive Filtering on Pressure: Since the pressure builds up directly as the square of the volume, with the same smoke and dust burden, it is necessary to shake the bags in that increased ratio, and vigorously, to keep the pores of the bag open and pressure down. This increases the wear on the bags and the dust losses. 4. Effect of Shaking on Dust Losses: The losses are in direct proportion to the frequency of shaking because the loss takes place immediately after shaking, before the bag seals itself. Tests show that with hand shaking the last 20 pct of the removable dust requires seven times more shaking per unit weight of dust than the first 80 pct. 5. Effect of Fineness of Dust on Losses: The fineness of the dust particle increases the loss in two ways: (1) When the bag is clean a greater percentage of it will pass through before the bag seals itself; (2) fine material penetrates farther into the bag pores, building up pressure faster and necessitating more shaking, thereby increasing the loss. The physical principle of operation of the bag filter is simply the separation of solids from a gaseous carrier through a porous medium, and is analogous to the filtration of solids from liquid carriers. Basically, the mechanics of the process are as follows: Fume-laden gases drawn from the operating fur-

Jan 1, 1951

By Frank J. Tolonen

BENEFICIATION of iron-bearing formations is one of the major problems of research at the Michigan College of Mining arid Technology. Funds for this purpose hate been supplied by the State of Michigan supplemented by contributions of labor and some supplies from the Federal Government during the depression. The iron-mining companies have cooperated in securing samples and have also furnished much valuable information relative to the various formations.

Jan 1, 1937

By J. W. Peirce

Advances in instrumentation now make it possible to measure accurately flaw of such difficult-to-mea-sure liquids as mining slurries. A mass flowmeter, which introduces no restrictions in the line, will measure the mass flow of slurry equal to the volume flow times the mms/unit volume. In the mass flowmeter system a volume flow measurement is electrically multiplied by a density measurement in order to give mass flow. Fig. 1 shows a typical mass flow measuring system. The volume flow is measured by a magnetic flowmeter and the density is measured by a gamma gauge cell, neither of which have any restrictions in the line. Consequently, there is no clogging and no more loss of head than there would be in a similm length of standard pipe. None of the parts extend into the flowing stream and there are no moving parts in the measuring elements. PRINCIPLE OF OPERATION OF THE VOLUME FLOW TRANSMITTER The magnetic flowmeter, which is used for the volume measurement, operates on the principle of a generator (Fig. 2). This principle states that when a moving conductor cuts lines of force in a magnetic field, a voltage is induced in it. Moreover, if the dimensions of the conductor remain constant, then the induced voltage is directly proportional to the velocity of the conductor. Since the voltage is strictly a function of flow velocity, the transmitter output can in no way be influenced by such factors as temperature, viscosity, turbulence, or conductivity. The linear output is detected by flush mounted electrodes (as shown) and transmitted to a potentiometer flow receiver. Magnetic Meter Construction: The metering tube is constructed of nonmagnetic stainless steel. (See Fig. 3.) To prevent shortcircuiting of the generated voltage, the tubes are lined with a suitable liner, depending upon the corrosive, erosive or other properties of the liquid or slurry to be measured. A uniform magnetic field is produced by two saddle-shaped coils of insulated copper wire. A laminated iron core focuses the magnetic field at a 90' angle to the flowing liquid. The voltage generated by a flow of liquid through the magnetic field is detected by two point electrodes mounted flush with the inner wall of the metering tube, 180' apart. Electrodes are made of Type-316 stainless steel or other material suitable to the application . Volume Flowmeter Operation: The millivolt output of the metering transmitter is transmitted to a potentiometer-type receiver (Fig. 4). The self-balancing, self-standardizing receiver has a drive unit for circuit rebalancing. The measurement change imposes an unbalanced input on the twin windings of the drive unit. The resulting unequal pull on a segmental rotor provides the motion which drives a pen, while simultaneously and steplessly repositioning a rebalancing element. Changes in conductivity have no effect on the accuracy of the measurement or the calibration of the instrument. The minimum conductivity values are determined by the length of the transmission cable and the size of the transmitter. They can be as low as five micro mhos, since there is such a small load on the electrodes. Because the meter electrically averages all of the incremental velocities in an irregular flow profile, it is ideally suited for turbulent flow and suspended solids. Straight runs of pipe are therefore not required. The transmitter can be installed close to

Jan 1, 1962

By David Cornell, Yusuf K. Sukkar

The fundamental differential equation for fluid flow has been rearranged, integrated numerically for natural gases, and the results presented in tabular and graphical form suitable for the direct calculation of vertical gas flow problems. The integration is rigorous except for the use of a constant average temperature. Both static and flowing bottom-hole pressures may be calculated by (his method without resort to trial and error procedures. INTRODUCTION Bottom-hole pressures can be determined either by direct measurement with a bottom-hole pressure gauge or by calculation. Since measurement with a bottom-hole pressure gauge is costly, calculation of bottom-hole pressures is to be preferred when possible. Calculation involves knowledge of the wellhead pressure, the properties of the gas, the depth of the well, the flow rate, the temperature of the gas, and the size of the flow line. These quantities must be combined in a suitable equation for vertical flow. Several equations have been developed for the purpose of calculating gas flow through both horizontal and vertical pipes1,2,4,5,6,7,8,9 These equations all have the basic differential equation for flow as their starting point. They differ only in the assumptions which are made in order to simplify the integration step. The bottom-hole pressures calculated using any of the better known equations are in good agreement with bottom-hole pressures calculated by a more rigorous but considerably more tedious stepwise integration method. All of these methods, however. require a trial and error solution for the calculation of flowing bottom-hole pressures. The purpose of this paper is to present a method which provides a direct solution of static and flowing bottom-hole pressures without use of a trial and error solution. At the same time the method involves fewer simplifying assumptions than any of the previous methods used, except lor the lengthy stepwise method. The derivation of any practical flow equation begins with the basic differential equation which is rigorous for almost all flow problems. The five terms in Equation 1 represent the potential energy, the potential energy of position, the kinetic energy, the shaft work done, and the energy losses due to motion of the fluid respectively. Equation 1 reduces to Bernoulli's theorem for the case of horizontal flow of an incompressible fluid, no work done, and no energy losses due to friction. In the present case, the term (v) (dv)/(g,.) can be shown to be quite negligible compared to the other terms, and is therefore neglected in the derivation. The term — dws includes work done by a pump or turbine in the system and is zero in this case. The frictional losses in the pipe are included in the term — d(lw) which is expressed in terms of the Moody friction factor as in Equation 2. (f) (v2) (dL) d(lW) = (f)(V2) (dl)/(2) (g) (d) The specific volume of a gas can be expressed in terms of the temperature, pressure, molecular weight, and compressibility factor. The velocity of the gas can be expressed in terms of the weight rate of flow, the specific volume, and the dimensions of the pipe.The reduced pressure Pr is to be used because z can be expressed as a function of Pr and Tr alone for most natural gases:

Jan 1, 1956

By R. C. Gifkins

R. C. Gifkins (CSIRO)—In this paper evidence is put forward to support the idea of grain boundary shearing in aluminum at 4.2°K and the phenomenon is explained in terms of a low-temperature "equicohesive temperature". It is stated that these findings show that grain boundary shear cannot be ruled out in any situation by temperature arguments alone. We believe, from our own recent study of "apparent grain boundary sliding'' at temperatures below 0.45 Tm, that it can be ruled out on temperature arguments alone.45 Our experiments were made principally with magnesium and a Ms-A1 alloy ("Magnox A12"). Careful examination of offsets of marker lines or steps at grain boundaries produced by deformation at room temperature (-0.3 Tm) showed many of these to result from curvature over narrow zones along the grain boundaries. A significant residuum did, however, appear as a sharp offset or step. It was found that the mean value of this step rose to a small value during the first 0.5 to 1 pct strain and then remained at this level with further strain. Moreover, exactly the same values were obtained (at the same stress) at 75°C instead of 21°C. That is, in this range of temperature the "apparent sliding" was not temperature-sensitive. At 125°C and above there was sliding which increased with strain and which was temperature-dependent. When this temperature-dependent sliding was used to calculate the sliding to be expected at room temperature, this underestimated the "apparent sliding" by a factor of 104. The conclusion drawn is that the "apparent sliding" consists of shear localized in a zone of a width measured perhaps in microns, whereas it seems that true sliding occurs on the boundary surface.46 A model for such localized shear has been developed in general terms, following a similar model of Urie and Wain.47 Other strong evidence for low-temper- ature sliding has been based on the offsets of marker lines consisting of a substructure seen using the electron microscope to resolve these.48 We have repeated this work and found that the effects can be accounted for in terms of cracks in an oxide film which is produced when the substructure is etched up; these cracks form over zones of shear rather than at actual interfaces. However, in aluminum at room temperature (-0.3 Tm) there is no "apparent sliding" which is optically detectable. We suggest that, when the temperature is much lower (and possibly the rate of strain higher, too), as in the experiments of Chin et al., then aluminum behaves like magnesium at room temperature, and a narrow zone of shear develops to accommodate differences in strain across the grain boundary. Such a zone could well give rise to triple-point cracks. It is hoped to investigate these suggestions further. B. Y. Chin, W. F. Hosford, Jr., and W. A. Backofen (authors' reply)—It is interesting to have Dr. Gif-kin's comments, although it is somewhat difficult to reply in detail without the reference containing the work that he describes. Apparently though, he does find a low-temperature temperature-independent sliding that would seem to be related to the observations reported here. Perhaps some of the difficulty has semantic origins, growing out of the distinction between ('true sliding" and "apparent sliding".

Jan 1, 1965

By W. W. Krysko

In order to calculate the critical particle size of lead oxide entrapped in lead, Stoke's Formula may be applied. All data were available except figures for the density of lead oxide at elevated temperatwes. The density of lead oxide was determined by means of dilatometmetrc and volumetric methods up to 1010°C. Graphs of the density of lead and lead oxide as well as the viscosity of lead are given and discussed, with the conclusion that the most unfavorable region for the separation of lead oxide from molten lead lies just above the melting point of lead. DURING previous investigations, particularly on the dedrossing of lead with aluminum1 and the oxygen solubility in lead,' the critical particle size of lead oxide which could be entrapped in solidifying lead was estimated by calculation, and it was shown that high values found for the oxygen content of lead may be due to the method of sampling and the entrapped particles of dross. Unfortunately, due to lack of data, an exact mathematical solution was not possible. The particle size which will separate under a given condition can be estimated from Stoke's Law, which states that where: d = diameter of the critical particle size ¦q = viscosity of lead in Poise V = constant velocity in cm sec-I g = 981 cm sec-2 61= specific gravity of lead 62= specific gravity of lead oxide The depth of the lead bath and the time before solidifying define the velocity. The viscosity of molten lead is known from earlier work by Krysko and is given in Table I. The specific gravity of lead is also known, see Table II. These values are according to Hofmann. A careful survey of literature revealed no values for the density of lead oxide at elevated temperatures. The present investigation was therefore conducted to provide this information. RESEARCH INVESTIGATIONS The density of lead oxide was determined from the following experiments: 1) Determination of the density of lead oxide at room temperature. 2) Dilatometric investigations on the same samples. 3) Determination of the shrinkage at the liquid-solid trans£ormation. 4) Density determinations on liquid lead oxide. 5) Time temperature curves during cooling. 1) The samples for the determination of the density of lead oxide in the solid state were prepared in the following manner. Molten lead oxide was cast into I hot mold which could be opened along the longitudinal (vertical) axis. The cast samples were cylinders of approximately 10 mm diam and approximately 80 mm length. The cylindrical hot mold was open at both ends and was placed on a large copper chill plate shortly before casting. The solidification of the lead oxides proceeded upwards without any formation of internal cavities. The sample was carefully checked with respect to density and crystalline appearance, and only the lower portions which were crystalline and free from porosity were used. The density was determined in a specially made pyknometer-like container, where the surplus liquid was squeezed out through an orifice approximately 1/10 mm in diameter. Mercury or water was used as the liquid. The densities of the samples were calculated from the weight, volume, and mean specific gravity of eight samples was 9.62 * 0.05 g cm"3 at 22"~. 2) The samples were machined to approximately 1 1/2 in. in length and fitted into a dilatometer. A

Jan 1, 1962

By C. E. Lundin

The Ce-H system was investigated in the temperature range, 573° to 1023°K, and the pressure range, 10-3 to 630 Torr, as a function of 'composition up to 72 at. pct H. Families of isothermal arid isopleth curves were plotted from the pressure-terr~perature-composition relationships. From these curves the solubility relationships were determined for the system. The isopleths are analytically represented by equilibrium dissociation pressure equations. The relative partial molal enthalpzes and entropies of solution of hydvogen in the systerrz were calculated fronz the dissociation pressure equulions and are tabulated. The integral free energies, enthalpies, and entropies of mixing in the Ce-H system were determined from the relative partial quantities and are also tabulated. The standard free energy, enthelpy, and entvopy of reaction of the dihydride phase at kcal per kcal per mole H2, and ?S° = -34. 1 cal per deg mole H2, respectively. The equilibrium dissociation pressure equation in the two-phase region is: UNTIL recently very little was known of the detailed solubility and thermodynamic relationships of the Ce-H system. Two previous investigations1,2 are noteworthy. However, significant discrepancies and omissions exist on analyzing them. The work of Mulford and Holley1 on cerium did not clearly delineate the boundaries of the two-phase region, Cess - CeH2-x. The plateau partial pressures were not thoroughly defined and were considerably displaced in pressure compared to those from the work of Warf and Korst.2 These latter authors concentrated their studies primarily from 823° to 1023°K in the pressure range of 1 to 760 Torr. No data were determined to outline the regions of primary solid solubility and the hydride phase. Also the establishment of the plateau partial pressures was rather limited in scope. In neither work was a treatment conducted of the relative partial molal enthalpies and entropies of solution of hydrogen in the single-phase regions and the integral thermodynamic quantities of mixing throughout the system. Therefore, it was the objective of this research to determine the complete equilibrium solubility relationships and thermodynamic data for the system by pressure-temperature-composition studies. EXPERIMENTAL PROCEDURE The cerium metal for this study was donated by the Reno Metallurgy Research Center of the Bureau of Mines. Total impurity content was 0.13 pct with only 60 ppm O. The metal was checked metallographically and contained only minor amounts of second phase compared to cerium from other sources. Specimen preparation was done in a dry box flushed with argon gas. The surface of a small rectangular piece of cerium (about 0.2 g) was filed with a clean, mill file. Final weighing was done in a tared enclosed vial containing argon gas. The specimen was then loaded quickly into the reaction chamber which was purged several times with high-purity hydrogen gas and then allowed to pump to about 10-6 Torr. The furnace was heated to the reaction temperature and the run started. The equipment used to conduct the hydriding was a Sievert's-type apparatus. Basically it consisted of a source for pure hydrogen, a precision gas-measuring burette, a heated reaction chamber, a McLeod gage, and a mercury manometer. Pure hydrogen was supplied by the thermal decomposition of uranium hydride. The 100-ml precision gas burette was graduated to 0.1-ml divisions and was used to measure the quantity of gas and admit it to the chamber. The reaction chamber was a quartz tube. Prior to each run, the cerium specimen was wrapped in a tungsten foil capsule to prevent reaction of the cerium with the quartz. Control of the temperature was achieved within ±1°K. Pressures in the manometer range were measured to ±0.5 Torr and in the McLeod range (10-3 to 5 Torr) to ±3 pct. The compositions of hydrogen in cerium were calculated in terms of hydrogen to cerium atomic ratio. These compositions were estimated to be ±0.01 H/Ce ratio. The technique used to study the equilibrium pressure-temperature-composition relationships of the Ce-H system was to develop experimentally a family of isothermal curves of composition vs pressure. The range of pressure through which each isotherm was developed was from 10-9 to about 630 Torr in the temperature interval, 573° to 1023°K. RESULTS AND DISCUSSION The hydriding characteristics of cerium are iso-morphous with those of the elements of the light-rare-earth group (lanthanum, cerium, praseodymium, and neodymium) wherein the region from the dihydride to trihydride is continuously single phase.' The structure of this phase is fcc.3 The heavy rare earths form a trihydride,2 which is hcp, separated by a two-phase region from the fcc dihydride phase. The Ce-H system is represented by the family of experimental isotherms in Fig. 1. Due to the small scale required to draw the curves, the experimental points are omitted; however, a total of 240 experimental data points were taken to prepare these curves. The solubility relationships can be deduced therefrom. Three distinct regions of partial pressure and composition can be seen. The region of cerium solid solution is represented by the rapidly rising isotherms in the dilute composition range. In accordance with Gibbs Phase Rule only one solid phase, the cerium solid so-

Jan 1, 1967

By Edward W. Parker

IF one is to place any credence at all in the reports published in the daily press, the subject of conservation has been a very lively topic of conversation during the past 60 days, and it does not appear that the temperature of the summer months has been in any way moderated by the discussion. It is a subject in which we are all vitally interested and to which, so far as our mineral resources are concerned, both the Institute and the Geological Survey have liberally contributed. The report of the National Conservation Commission, appointed by President Roosevelt, contains a series of papers on the conservation of mineral resources, all of which were prepared by members of the Geological Survey and were compiled largely from information previously collected by that Federal bureau in the performance of its regular duties. It is not the purpose of this paper to reiterate in extenso any of the material already published. The contributions of the Survey officials to the Commission's report have been published in a separate document as Bulletin No. 394. This docunient is for free distribution and may be obtained upon application to the Director of the Survey. In the preparation of this paper I desire merely to make a few suggestions regarding the possible necessity of some restraint upon or control of one branch of the mining industry with which I have been somewhat closely associated for the past 20 years-that of coal. Most of the members of the Institute are cognizant of the suits brought by the government against the anthracite-operators in Pennsylvania, or the combination of interests commonly known as the ? hard-coal trust." No defense of any illegal combination in restraint of trade is intended, but there are some facts which should not be lost sight of, and unfortu-\

Nov 1, 1909

By Ernest N. Patty

WHEN an Alaskan prospector makes a new mineral discovery he stakes out his claims and then starts prospecting for a near-by landing field. This may be a convenient lake but more often it is a gravel bar along a river, where, by cutting off some brush and throwing the larger rocks to one side and then tying a flour sack to a pole, he may decoy some intrepid aviator into testing it gingerly with the wheels of his plane. It was only a few years ago that .the fence around the baseball park at Fairbanks, Alaska, was taken down so that planes could get in and out of the ball park without ripping off boards. Today the grandstand is lost among the modern hangars which line one side of a modern and busy airport. Planes come and go all during the day, and in summer, when there is no darkness, the night stillness is frequently broken by the hum of a motor overhead. Fairbanks is the hub of interior Alaska and from that hub airplanes radiate in all directions to mining camps large and small. An air industry of comparable size is based at the town of Anchorage, and in fact there are but few towns of more than 200 population in the Territory that do not have one or more planes based there.

Jan 1, 1937

By C. E. Ells, A. Sawatzky

The effect of neutron irradiation on tensile properties of the Zr-2.5 wt pet Nb-0.5 wt pet Cu alloy has been evaluated for integrated neutron fluxes up to 3 x 1020 n per sq cm (E > 1 Mev). Specimen temperature during irradiation was normally 300°C. The material was studied both in annealed and in quenched and aged conditions. When the alloy (fab-ricatedfrom sponge zirconium) is quenched from a temperature =40°C below the a + ß/ß transformation and aged 6 hr at 535oC, neutron irradiation up to the maximum studied has little effect 071 reduction in area, although there is a drop in uniform elongation and the yield strength is increased by = 25 pct. Conversely, irradiation of the alloy in an annealed condition can result in much greater changes in properties, with yield-strength increases of up to 200 pet. The dose dependence of irradiation hardening obeys the saturation equntion proposed by Makin and Minter to explain hardening in copper and nickel: Act = C[1 - exp(-Døt)/1/2 A rnodel is proposed to explain the effect of niobium content and metallurgical condition on irradiatiotz behavior. THE Zr-Nb alloys form a class of material susceptible to marked strengthening by quench and age heat treatment. With niobium concentrations as low as 2.5 wt pet, strengths can be developed which are double those of annealed zircaloy-2.1 Since the neutron-capture cross section of the Zr-2.5 wt pet Nb alloy is nearly identical to that of Zircaloy-2, significant gains in power-reactor neutron economy could be obtained by replacing stressed in-reactor components of Zircaloy-2 with the heat-treated Zr-Nb alloy. When the solution heat-treatment temperature is in the high (a + ß) phase, then the tensile properties of the quenched (and aged at 500°C) material have been shown to be relatively insensitive to neutron irradiation.1-3 Addition of small quantities of copper to the binary Zr-2.5 wt pet Nb alloy gives a useful reduction to the corrosion rate in air and carbon di- oxide,&apos; and a ternary alloy of composition Zr-2.5 wt pet Nb-0.5 wt pet Cu was chosen for one component in the Douglas Point Reactor. This application utilizes both the corrosion resistance to a moist air-carbon dioxide environment and the strength developed by a quench and age heat treatment. By suitable choice of solution heat treatment and aging temperatures, tensile properties of the ternary alloy can be made nearly identical to those considered as optimum for the binary alloy. One effect of the copper addition, however, is to give marked changes in the aging kinetics of the alloy.&apos; For this reason, it appeared that neutron irradiation might promote overaging in the ternary alloy, particularly at the elevated service temperature (=300°C). This paper describes the effect of neutron irradiation on the tensile behavior of the Zr-2.5 wt pet Nb-0.5 wt pet Cu alloy in several metallurgical conditions pertinent to its use as a reactor material. The metallurgical conditions included those for which aging occurs in the unirradiated material, and fully annealed conditions for which a higher degree of thermal stability could be expected. 1) EXPERIMENTAL Two batches of alloy were used, both made from sponge zirconium. The fabricator&apos;s analysis of principle constituents showed no significant difference between the two batches, Table I. The rod material was rolled to 1.5 in. diam, and then swaged to final sizes of 0.5 to 0.375 in. diam at nominal temperatures of 785" and 625°C for the AM and AS batches, respectively. Subsequent hydrogen analyses at CRNL confirmed that the hydrogen concentration of the as-received rod was indeed <20 ppm. However, it was found that the (a + ß)/ß transus for the AM material was approximately 25°C lower than that of the AS material, indicating rather more difference in oxygen concentration than given in Table I. Except for a small difference in yield-point behavior, the tensile prop-

Jan 1, 1965

[†ADAIR, ARTHUR C., Corp 4th Co., 1st. Training Bat., 154th Depot Brigade. ADAMS, HENRY, Major Chemical Warfare Service. ARLUCK, A. A Headquarters, 1St Corps of Artillery. ASKIN, THOMAS B. H Ensign.] [ ]

Jan 1, 1925

By R. D. Pehlke, J. V. Gluck

The limiting activity coefficient of zinc in a "binary " solvent has been described for several dilute ternary solutions by combining literature data from recent interaction studies in the terminal regions with thermodynamic data for the central regions. Results are discussed in terms of the Alcock and Richardson random solution model. The observed behavior agrees semiquantitatively with predictions using the model. RECENT efforts in the thermodynamics of dilute ternary liquid metallic solutions have been mainly concerned with the interactions between two dilute solutes, that is, the terminal region at an apex of the composition diagram.&apos;-&apos; A more general case is that of the behavior of a dilute solute over the full range of a ternary system, or, more specifically, the limiting activity of solute i as projected to the k-j binary side of the i-j-k composition triangle. An implication of this is that the interaction between i and j at the k apex should connect in some way to the interaction between i and k at the j apex. A model to characterize various types of expected behavior has been discussed by Alcock and Richardson.9,10 Such behavior can range from a simple linear relation for In yi over the 1-k binary to more complex curved relations which depend on various assumptions regarding the heat of solution in the j-k binary system and the presence or absence of clustering with solute i. Quantitative data to assess these concepts have been limited primarily to the cases of nonmetallic interstitial solutes in ferrous solutions. Alcock, Richardson, and various coworkers9-13 have discussed the behavior of oxygen, sulfur, silicon, or hydrogen in dilute solution over the range of iron + metal systems; however, the only data considered by them for substitutional nonferrous metallic solutions were the results of Elliott and Chipman concerning the activity of cadmium in the Pb + Sn, Pb + Bi, and Pb + Sb systems.9 These data were limited and did not cover the terminal regions in detail. An opportunity to extend the Alcock and Richardson considerations to the dilute solution of zinc in various "binary" solvents has been afforded by combining results of several of the recent investigations of ter-minal-region interactions with previous data from the literature for the same more concentrated solutions. RESULTS AND DISCUSSION The results of the analysis are shown in Figs. 1 to 5 where the log activity coefficient, (In yZn)xzn--o, is plotted vs the composition of the "binary" solvent. In these plots the end points are the values for In yzn in the appropriate binary solution with Bi, Pb, Sn, Cd, or In at the indicated temperature—usually 550°C (823°K). For the terminal regions the slopes were drawn using the first-order interaction parameters taken from various studies reported in the literature. The data points shown in the central regions were obtained by the short extrapolation to infinite dilution of the logarithm of the activity coefficient calculated from thermodynamic studies reported in the literature for the same systems. The sources of the data are indicated on each figure. Second-order effects were not considered. Reasonably smooth curves can be drawn to connect the central region with the terminal regions for the Zn-Pb + Bi system, Fig. 1, the Zn-Pb + Zn system, Fig. 1, and the Zn-Bi + Cd and Zn-Bi + Sn systems, Fig. 2. As shown in Fig. 2, where central region data are available from two independent investigations, the agreement is quite good. The Zn-Pb + Cd system does not appear to be as well behaved as the previous four, Fig. 3; however, this type of behavior has also been

Jan 1, 1970

By J. J. Taber, I. S. K. Kamath, R. L. Reed

Alcohol floods of consolidated sandstone cores have shown the process to be strongly dependent on the phase behavior of the particular alcobol-oil-water system used. This means that in many cases the mechanism does not conform to the idea of a piston-like displacement. Instead, it is found that by changing the alcohol it is possible to change the relative velocities of the oil and water and, in fact, the entire mechanism of the process. The effects of rate, viscosity, initial saturation, distance travelled and hysteresis of relative permeability on the alcohol flooding mechanism are discussed. INTRODUCTION Reasons for interest in the use of alcohol to miscibly displace oil and water from a porous medium appear in the existing literature.l-7 The mechanism of the displacement has been considered and the apparent implications formulated into a theory which presumably would enable one to predict the essential features of the process.6 Unfortunately, most of the reported experiments have been performed with unconsolidated or artificially consolidated sands. With these systems some of the noteworthy facets of the process are obscured and resulting data appear uncertain. It is the purpose of this paper to show how the use of consolidated sandstones has led to revision of the mechanism and, hence, the theory of alcohol flooding. The practical result is increased pessimism toward the possibilities of commercial application of the simplest form of the alcohol-slug process. However, elucidation of the mechanism has made it possible to define the essential characteristics of a system of slugs which will behave in a nearly piston-like fashion and, thus, yield the best possible result. EQUILIBRIUM PHASE BEHAVIOR Fig. 1 is a diagram of the ternary system isopropyl alcohol (IPA)-Soltrol-calcium chloride brine. Brine was used to prevent plugging of the core and calcium chloride was used because sodium-chloride brine exhibits a solid phase with Soltrol and IPA. If alcohol is added in increments to the immiscible mixture of water and oil represented by Point A, the path followed by the successively equilibrated samples will be on the Line ABC and pass from the immiscible region to the miscible region by crossing the binodal curve at B. Consider the intersection D of this path with the tie Line EF. The quantity of oleic phase is proportional to the Segment ED and the quantity of the aqueous phase is proportional to DF. Compositions of the two phases are specified by Points E and F. It is clear that in the case shown the oleic phase is diminishing and entirely disappears when miscibility is achieved. Such will

By C.F. Pascoe, H.J. Roast

The object of this investigation was to compare the arsenical antimony-lead alloy with some of the regular bearing-metal alloys. With this end in view, the following tests were made: 1. Chemical analyses to show the composition. 2. Thermal analyses. 3. Macroscopic examination of fractures. 4.Microscopic examination. 5. Tensile tests. 6. Crushing tests. 7. Scleroscope tests.&apos; 8. Brinell tests, at temperatures from 80&apos; to 400&apos; F. Tests 3 to 8 were conducted on both sand-cast and chill-cast specimens. For the purposes of this paper the alloys have been numbered from 1 to 6; their composition is as follows: Lead. &apos; Antimony, Tin, CoppeR ARSEnic, NumBeR PeR CENT. PER CENT. PeR Cent. PeR CeNt. PeR Cent. 1 82.64 12.95 4.40 0.01 none 2 27.54 9.04 59.87 2.65 none I 3 none- 0.13 ------------1.47 4 82.95 16.85 none 0.20 none 5 83.78 15.35 none 0.05 0.82 6 0.15 7.87 . 84.13 7.85 none The analysis, in the case of the antimony, was made by solution in sul-furic acid and titrated with permanganate. For tin, the ferric-chloride method was used; for copper, the alkaline separations, and the thio-cyanate methods; for arsenic, volatilization of arsenic and titration with iodine. The lead, with the exception of alloy 6, was taken by difference. With the exception of No. 5, the alloys were taken as received from the manufacturer. About 40 lb, were melted in a graphite crucible and recast into small ingots to insure uniformity of composition; samples

Jan 1, 1923

By C. F. Pascoe, H. J. Roast

The object of this investigation was to compare the arsenical antimony-lead alloy with some of the regular bearing-metal alloys. With this end in view, the following tests were made: 1. Chemical analyses to show the composition. 2. Thermal analyses. 3. Macroscopic examination of fractures. 4.Microscopic examination. 5. Tensile tests. 6. Crushing tests. 7. Scleroscope tests.&apos; 8. Brinell tests, at temperatures from 80&apos; to 400&apos; F. Tests 3 to 8 were conducted on both sand-cast and chill-cast specimens. For the purposes of this paper the alloys have been numbered from 1 to 6; their composition is as follows: Lead. &apos; Antimony, Tin, CoppeR ARSEnic, NumBeR PeR CENT. PER CENT. PeR Cent. PeR CeNt. PeR Cent. 1 82.64 12.95 4.40 0.01 none 2 27.54 9.04 59.87 2.65 none I 3 none- 0.13 ------------1.47 4 82.95 16.85 none 0.20 none 5 83.78 15.35 none 0.05 0.82 6 0.15 7.87 . 84.13 7.85 none The analysis, in the case of the antimony, was made by solution in sul-furic acid and titrated with permanganate. For tin, the ferric-chloride method was used; for copper, the alkaline separations, and the thio-cyanate methods; for arsenic, volatilization of arsenic and titration with iodine. The lead, with the exception of alloy 6, was taken by difference. With the exception of No. 5, the alloys were taken as received from the manufacturer. About 40 lb, were melted in a graphite crucible and recast into small ingots to insure uniformity of composition; samples

Jan 1, 1923