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By Ian MacGregor

INTRODUCTION I retired in 1977 and have been enjoying myself ever since doing odd jobs for various people - on occasions, politicians. I commend it to you as post retirement - get involved in situations that nobody else will take on. It produces all sorts of excitement, it is a lot of fun and, courtesy of the media, you provide great raw material for their current Punch and Judy shows. The example of the adventures in the Escondida financing made me think about the history of the minerals industry in the United States. In order to see more clearly what is going to happen in the future, we have to look at the past and what we have learned, and try to extrapolate from that a scenario for the future. This paper is about global material cycles and the financial needs of the extractive industries. Since about 2,000 years B.C. we have been living in the Iron Age. Since about 4,000 B.C. we have been living in the Copper Age. These have been the two foundations of modern industrial society. Understanding the use of metals and materials helps us understand the civilization in which we live today. Let me focus more closely on the United States, particularly in this century. In the first two-thirds of this century the United States was growing and expanding and developing. It was providing for its inhabitants a society which was enjoying the benefits of a rising standard of living, improved opportunities and the ability to support a population growing at quite a dramatic rate. It is still growing. In the Department of Commerce, there is a clock that shows how many people there are in the United States. It turns all the time; the numbers are going up. On January 21,1980, when Ronald Reagan was sworn in as President, the number on that clock was a little over 220 million, and in the last week before he left, it was travelling towards 245 million. During his period in Washington, we have added , the equivalent of one total state such as California. That gives the dimensions of the U.S. appetite for raw material inputs. The inputs are of critical importance. COMPARATIVE ADVANTAGE During the first two-thirds of this century, the minerals sector enjoyed comparative advantages. There was a clear understanding of the importance of minerals enshrined in the tax legislation of the country from the early decades of this century - the depletion allowance. No other country identified as clearly that the supplies of minerals were finite. As they are used up, just as with a wearing-out machine, one must get capital back to replace the resource when the time comes. Within the last decade, that idea has been questioned and criticised. It is now seen as giving money away to a preferred part of the economy, instead of recognizing what depletion is all about. In previous times it insured that the U.S. always was able to regenerate the capital in its minerals industry, so that it could remain competitive. The U.S. had plenty of everything, including energy. We had a growing population. Equally important, we had an agriculture that led the world in productivity, even as far back as the turn of the century. It continues to do so. We still retain that comparative advantage, an ability to produce our foodstuffs in quantity and quality at excellent cost. Look at another ingredient of our comparative advantage, the energy position. We ran through the first two-thirds of the century with ever- decreasing energy costs. In the 1940s and 1950s the coal industry went into a decline. Why? Because the $3.00 - $5.00 a ton cost of mining coal in the U.S. represented relatively expensive energy. A ton of good grade bituminous coal contains 25 million BTU's. $0.20 a million BTU's represents $5.00 a ton at the pithead. That price was undermined by the increasing ability to find petroleum at even lower costs. At the apex of the growth of development of the American dominated

Jan 1, 1990

By Eugene A. Thiers, William V. Morris

INTRODUCTION Direct reduction processes reduce the various commercial forms of iron oxide (pellets, concentrate, fines, etc.) to metallic iron at temperatures lower than that of molten iron. Thus, this technology includes practically, all iron reduction processes other than blast furnaces and electric pig iron furnaces (whose output in terms of world production is negligible). The product of these processes which is known as direct reduced iron (DRI), or sponge iron, is primarily used as a source of metallic iron in steel-making operations. Interest in DRI, which has been significant since the early 1960s, increased significantly in recent years with the rapid growth of DRI installed capacity throughout the world. The importance of the subject for the Circum-Pacific region stems directly from the influence that DRI has on iron ore consumption and on future steel development for this region. Although there is widespread agreement that electric furnaces will continue to increase their share of global steel output, and especially so in the countries of the Pacific Steel community, some doubts exist about future scrap supplies being adequate to support growth at past rates. The authors believe that such doubts are soundly based. As this paper points out, the total supply of all metallics used in electric furnaces may not be adequate to support the extrapolated rapid growth in electric furnace steel production. This paper seeks to provide perspective on the global and Circum-Pacific prospects for DRI in light of recent energy price developments and the current recession. In this regard, the demand for DRI within the context of recent evolutionary patterns in steel-making, the outlook of DRI supply in terms of prevailing production costs, and the prospects of new technology are discussed. THE DEMAND FOR DRI Although several reports published in the last 10 years predict high rates of growth in DRI, the subject remains a controversial one. Significant growth has indeed occurred, but not to the extent anticipated in the studies summarized in Table 1. The substantial difference between previous expectations and present reality can be ascribed primarily to: (1) lower growth in steel production than formerly anticipated; (2) numerous cancellations of DRI facilities that were previously announced; and (3) a fundamental and probably irreversible change in the economics of DRI production. Note that DRI capacity at the end of 1980 was about 16 million tonnes, a significantly lower figure than any of the projections above. In addition, DRI production was only about half of capacity, reflecting the abnormally low rates of capacity utilization in this industry. [ ] Before examining the current outlook in steel, it is pertinent to note that the market for DRI is usually different in the industrialized countries of the West from that in developing countries. In the former, the available infrastructure and industry's diversification extends DRI's potential markets to numerous steelmaking, foundry, and other industrial applications, although competition from scrap and other forms of metallic iron is constant. Scrap is generally available in these countries and, therefore, DRI competes with it in electric furnace steelmaking, basic oxygen steelmaking (as a coolant), cupola foundry operations, or as an additive in the metallic charge for open hearth and blast furnaces. On the other hand, DRI in developing countries is often allocated exclusively to domestic electric furnace steelmaking or, when capacity exceeds domestic captive requirements, to export. Notwithstanding quality considerations, DRI is being and is likely to continue to be used predominantly as a source of metallics in iron and steel-making. Other uses of DRI, such as in copper cementation represent a marginal market in terms of overall tonnage and can be ignored at this point. Therefore, DRI demand is-determined by the overall availability of metallic scrap in its various forms--a function of steel production and its probable evolution. The Global Steel Outlook Given the present recession, an objective appraisal of the long-term outlook for steel is particularly difficult. On the one hand, historical trends and, especially, the inertial forces associated with such a basic industry as steel must be recognized; such trends suggest that the current stagnation with

Jan 1, 1982

By W. A. Sheaffer

Electric furnace installation and tough-pitch copper-casting operation at Canadian Copper Refiners Ltd. are described. General layout, power supply and control, refractories, induction pour hearth, casting equipment, metal temperature control, and oxygen-content control are discussed. E LECTRIC furnace at Canadian Copper Refiners Ltd. was put into operation in August 1949. The installation was designed primarily to melt electrolytic copper cathodes and to produce vertically cast tough-pitch shapes. To meet emergencies during reverberatory furnace shut-downs, provision was made for casting horizontal tough-pitch wire bars. Due to the ease with which metal temperatures, melting rates, oxygen content, and casting hours can be altered to suit production demands, the electric furnace has met the requirements admirably. General Layout The plant has been described by H. S. McKnight' and by J. H. Schloen and E. M. Elkin.2 Since these papers were written, the electric furnace has been installed in a 140 ft extension of the original casting building. The extension, 264 ft in width, is divided into one 24 ft and four 60 ft bays, each a continuation of similar bays in the older building. Fig. l, a partial floor plan of the extension, shows the location of major equipment. Power Supply and Control Power for the electric furnace and its auxiliary equipment is delivered by Quebec Hydro over two 12 kv three-phase, 60 cycle pole lines. One line is in use, while the second serves as an emergency standby. Connections for the electric furnace from these lines, which also feed the refinery power house, are made to a substation on Canadian Copper Refiners Ltd. property. Connections between this substation and the furnace transformer room are in underground conduit. The are-furnace transformer, 550 v auxiliary-equipment transformer, and 12 kv switchgear are in the transformer room near the furnace. Stepdown from 12 kv to are-furnace voltages is done by a 7500 kva oil-immersed water-cooled three-phase 60 cycle transformer. This unusually high kva rating is available because the arc furnace and transformer were adapted from a steel-melting unit. Secondary voltages are available in 16 steps between 251 and 95 v (across phases). A four-position tap changer, operated from the furnace switchboard is connected to taps yielding 199, 164.5, 127, and 115 v, respectively. The transformer primary feeder is equipped with a General Electric Magne-Blast circuit breaker. The pour hearth, casting wheel, bosh conveyor, and other auxiliary equipment are fed by a 12 kv to 550 v, 750 kva three-phase 60 cycle transformer. Graphite electrodes, 14 in., with tapered nipples are held in the are-furnace electrode arms with steel wedges. The arms, fed from the transformer secondary busbars by flexible cable bundles, are actuated by winch cables. Direct-current motors operating the winches are supplied by a 250 v motor-generator set. Power input to the furnace is controlled from the furnace switchboard. The board has the usual combination of remote controls for the 12 kv breaker and transformer tap changer, meters, overload relays, automatic and manual electrode motion controls, switchgear, etc. When power is on the furnace, automatic electrode feed is used; the power draw on a given voltage tap then is governed by current-limiting rheostats. The rheostats work through a Westinghouse bal-lanced-beam control circuit which, in turn, operates reversing contactors in the winch motor circuits. Arc Furnace The arc furnace is a standard-type NT Pittsburgh 'Lectromelt with inside shell diameter 12 ft 3¾ in. By hydraulic-lift cylinders, the furnace can be tilted forward to 39 ½ ° and backward to 5" from horizontal. A roof-swing cylinder is also installed but not used. Fig. 2 shows refractory-lining details. With the lining as shown, operating life between repairs is 6 to 9 months of two-shift casting operation. Charge slot, skim door, and roof refractories then are replaced completely, and any necessary repairs made to side walls. Bottom life is longer and the original under courses are still in service, while the top course was replaced in October 1951. The launder is lined with high chrome-magnesite trough brick backed by fireclay and insulating brick. The trough is covered with fireclay brick and asbestos sheet. At intervals, openings 12x6 in. are left in the permanent covering. These are covered with

Jan 1, 1956

By E. T. Hayes, A. H. Roberson, M. H. Davies

R. F. Dornagala and D. J. McPherson (Armour Research Foundation, Chicago)—I should like to compliment the authors for a workmanlike job in determining the partial phase diagram of a system comprised of two rnetals which are certainly not easy to work with. We are completing work at Armour Research Foundation on an Atomic Energy Commission-sponsored project for the determination of eight zirconium binary diagrams. Work on the Zr-Cr system has been completed and should be published within the next year. For our work, Westinghouse Grade 3 iodide crystal bar served as the zirconium melting stock. Johnson-Matthey, electrolytic chromium, specially treated for oxygen removal, was employed. The overall constitution of the system determined at Armour Research Foundation is in very good agreement with the present work. We found a eutectic at 18 pct Cr and 1280 °C, somewhat lower than the value reported. This temperature was confirmed by thermal analysis, incipient melting studies, and regular isothermal anneals. The eutectoid was located close to 1 pct Cr and 835°C by metallographic analysis of annealed specimens. Maximum solubility of chromium in /S zirconium was 4.5 pct at the eutectic temperature. Chromium solubility in a zirconium was less than 0.28 pct at all temperatures. We found the compound at 53 pct to melt around 1700°C, with an open maximum, but determined its crystal structure to be hexagonal close-packed (MgZn, type). The lattice parameters were in excellent agreement with those determined by Wallbaum in 1942. The diagrams are in substantial agreement, and .part of the differences are undoubtedly due to the use of different zirconium melting stock. M. K. McQuillan (Birmingham, England)—I read this paper with a great deal of interest, as it covered the same field as some work of my own.' There are a number of points in the present paper on which I would like to comment. First, I should say that I, too, used zirconium prepared by magnesium reduction of the tetrachloride and electrolytically prepared chromium, and melted the alloys in a Kroll-type arc furnace. The purity of my alloys should, therefore, be comparable with the purity of those of the present authors, and any differences in our observations would not be expected to be attributable to this cause. The differences between my observations and theirs concern the presence of the eutectic, the temperature of the eutectoid, and the melting point of the compound. I would be very much interested in any further evidence the authors may have for the occurrence of the eutectic at 1380°C. During the course of my work I noted that a number of my alloys containing 60 to 90 pct Zr melted at about 1400 °C, and for a time assumed that a eutectic occurred at this temperature as described in the paper. On further investigation, however, I found that the structures of the as-melted alloys could not be made to fit in with this interpretation of the system. If a eutectic exists in this region of the system it would be expected that the as-melted alloys would show the usual type of cast structure, i.e., dendrites of the compound plus eutectic. This, however, does not occur, as may be seen from Fig. 9. The compound seen there is not dendritic in form, and the remaining material is by no means certainly eutectic. It may be argued that a compound such as ZrCrl would not form dendrites but would tend to crystallize in geometric shapes. In this case, however, I have evidence to the contrary, as on the chromium side of the compound, where a eutectic occurs at about 1545"C, the compound formed from the liquid takes on a conventional dendritic form, and the eutectic is observed in the interdendritic spaces in the usual way. There is no reason to suppose that the compound would behave differently in an alloy lying on the zirconium side of the compound composition if a eutectic existed there too.

Jan 1, 1953

By John E. Willson, Ben L. Seegmiller

A portable electronic instrument was designed and constructed to detect unknown stress magnitudes in rocks. The principle used to detect stress is based on the propagation velocity method. This method allows the stress in a rock subjected to a load of unknown magnitude to be determined, provided that the wave velocity-stress relationship for similar rock is known. This is accomplished by measuring the travel time of longitudinal mechanical waves passing between two points a set distance apart in a rock. The velocity of the wave is calculated and the stress determined from the wave velocity-stress relationship. If the sending and receiving transducer spacing is constant, a time vs. stress relationship rather than the velocity-stress relationship may be used. The method is nondestructive and tests can be made without drilling or otherwise disturbing the rocks. The first studies undertaken in the United States to determine stress in rocks using propagation velocity techniques were reported by Obert1,2 in 1939 and 1940. The Soviet Union first reported using propagation methods to study rock pressures in 1951.8 Success of the method led to the development of a pulse-type ultrasonic seismoscope4 in 1953. Using this instrument, Ivanov and Betaneli5,6 in 1963 succeeded in devising and testing under field conditions a method of investigating coal pillar stresses. In 1967 Osipova7 reported results of similar studies in the Nakhichevan salt mines. In France, Tincelin8 has used the propagation velocity method to study the stress distribution in iron ore pillars. Uhlmann9 has investigated stresses in salt and potash pillars in Germany, using velocity techniques. Determination of uniaxial stress by propagation velocity methods is limited to rocks which have readily detectable wave velocity-stress variations. The present stress detection instrument is restricted in application to rocks which have a velocity change under stress of at least 300 fps. Examples of rocks which meet this requirement are sandstones, coal, and possibly some limestones. Testing of this instrument was limited to a laboratory study and the results may or may not be indicative of what would be found in a field test. A program of field experiments to study the feasibility of using this instrument to determine mine pillar and tunnel stresses is in progress. Instrument Design The instrument has two main components: A probe and a control-display unit. The probe is a hand-held device to which two identical rodlike transducers are rigidly mounted. Coaxial cable connects the probe to the control-display unit which is mounted in an enclosed carrying case measuring 7 % x 9 x 13 in. The instrument is designed so it may be carried and operated by one man. Weight of the probe and enclosed control-display unit is approximately 20.1b. The probe consists of two transducers identical in construction. One transducer is used to convert electronic pulses to mechanical waves and transmit these waves into a rock. The other detects the transmitted waves in the rock and converts them back into electronic signals. The basic element in both transducers is a piezoelectric crystal. The crystal is a disk made of lead titanate zirconate and has a natural resonant frequency of 400 kHz ±1% in the thickness mode. A schematic of one of the two identical transducing elements is shown in [Fig. 1.] A spacing of 6 in. between sending and receiving transducers has been found to be most satisfactory. The control-display unit consists of a pulse generator and a 1-in. oscilloscope. Various electronic devices are used for the power supply, amplification, and calibration. The amplitude of the square wave from the pulse generator can be continuously varied between 0 and +20 vdc. Pulse width can be set at 1, 2, 3, or 4 µsec. The repetition rate of the pulse can have the following values: 50, 75, 100, 150, 200, 250, 300, 400, 500, 750, and 1000 Hz. The oscilloscope delay system allows the travel time of the longitudinal wave passing through a rock to be measured to an accuracy ±0.1 µsec in the range from 0 to 995 µsec. Fig. 2 is a schematic of the various electronic sections in the control-display unit. Laboratory Testing The first step in using the instrument is to develop a velocity-stress or a time-stress curve for the par-

Jan 1, 1972

By Bryan J. Travis, Thomas L. Cook

INTRODUCTION Our growing concern for adequate and secure sources of energy and minerals has stimulated vigorous exploration for new sources, research toward a better understanding of geological processes, and development of new extraction technologies. The need for control, or at least prediction, of subsurface fluid flow is important for many of these technologies for example: primary, secondary and tertiary recovery of oil; ground water and waste management; in-situ fossil energy extraction (oil shale, coal, tar sands); and solution mining of uranium, copper and other minerals. These technologies, especially the last two, are characterized by highly complex systems. A partial list of the physical processes occurring would include flow in porous/fractured media, multi-phase and multi-component flow with heat transport, chemically active fluids and soils, tracers, diffusion and dispersion, fracturing and dissolution. A great deal of understanding of how these processes behave and interact can be obtained from models. MODELS Theoretical models are valuable because they: 1) provide a frame of reference for interpreting results of laboratory and field experiments; 2) once validated by experiments, allow a variety of geometries, injection/production strategies, etc., to be examined (at relatively low cost) for efficiency and stability; 3) can provide guidance to the design of experiments and field operations. Most models are based on the fundamental principles of mass, momentum and energy balance. But from this starting point, many paths can be taken. For example, there is the theory of Payatakes (1973) which concentrates on the microscale dynamics. In this approach, the rock or soil matrix is represented by a complex of characteristic channels (such as periodically constricted smooth tubes). Detail of flow within the characteristic channel is calculated very accurately. A difficulty with this model is that description of a representative channel can require several parameters which may not be easily measured. Also, it is not clear how the channels can be combined practically to model a large scale flow. Another type of model is the "global" one described by Bear (1972). Here, the continuum equations for conservation of mass, momentum and energy are averaged over a distribution of pore channels, resulting in a set of conservation equations in which the small scale structure of the medium is replaced by quantities such as porosity, permeability, dispersion coefficients and tortuosity. This approach allows computation of large scale flows. However, additional constitutive equations are needed which relate averaged quantities such as permeability to observables such as local saturation, particle size distributions, and others. An important difference between models is the way they handle the momentum equation. Payatakes' model solves the full equation. In others, it is replaced by a simpler relation such as Darcy's law (valid for slow flow rates) or by Forchheimer's equation which extends Darcy's law to higher flow rates. These simple relations can nevertheless match a great deal of experimental data (Dullien 1975). The permeability term which appears in Darcy's law and in Forchheimer's expression has been related to other quantities such as porosity and particle surface area. The Kozeny-Carmen equation is a well-known example, valid for some ranges of porosity and particle sizes and for some materials. Several other semi-theoretical, semi-empirical formulations have been devised, none of which are entirely satisfactory. One goal of researchers has been the ability to predict permeability of a porous material from basic measurable quantities such as grain size distributions without recourse to adjustable parameters. This effort has proceeded from consideration of distributions of idealized, non-intersecting channels, to intersecting channels, to consideration of both distributed pore radii and pore neck radii. This last approach has been used successfully to predict permeabilities (ranging over several orders of magnitude) in sandstones (Dullien 1975). Additionally, studies on explicit networks of channels (e.g., Fatt 1956) and percolation theory (e.g., Larson et. al. 1981) have been used to examine interconnectivity effects in porous media with the hope of eventually being able to predict permeability accurately. To be of more than theoretical interest, a model must be compared s controlled laboratory conditions where boundary and initial conditions and relevant material properties can be accurately determined. Also, extraneous forces can be eliminated so that the interactions between processes of interest can be clearly seen and flow models can be vigorously tested. In contrast to the well-defined environment and

Jan 1, 1982

By H. Pew, J. H. Field

To alleviate pollution more restrictive legislation is being enacted, either limiting emission of pollutants or the type of fuel that can be utilized. The nature and magnitude of air pollution problems affecting the mining, preparation, coking and combustion of coal are described. Methods for combatting particulate emissions by use of mechanical separators and electrostatic pre-cipitors are discussed. Proposed methods to meet the problem of gaseous emissions currently receiving considerable attention are described, with special emphasis on methods to decrease pollution by sulfur oxides. Concern about air pollution goes back several centuries, but until very recently most effort has been aimed at coal smoke and other visible pollutants. The classic example of a 'successful' campaign for smoke abatement and control is the fruitful combined effort of the city of Pittsburgh and its surrounding Allegheny County, which eventually led to the reconstruction of downtown Pittsburgh at an estimated cost of one billion dollars. Historically, the city's downtown Golden Triangle district had been afflicted by pollutants evolving from steel mills, from a variety of other industries, and from railroad locomotives. Efforts to alleviate the situation prior to 1943 were virtually ineffective. In 1945, however, a comprehensive redevelopment plan was prepared and backed by state authority. Within a few years a clean, modern metropolis has evolved where once stood America's famous 'smoky city.' But the victory in Pittsburgh, as in various other American cities, has not solved the national problem. Current estimates indicate that 133 million tons annually of air pollutants from all sources still are emitted annually into the atmosphere above the United States. About 10% of this annual effluent is particulate matter so that most of the remaining pollution problems will be solved only when other effluents are reduced. Essentially, these are sulfur oxides, nitro- gen oxides, hydrocarbons, and carbon monoxide. Over the years, both states and local communities have tended to increase the restrictions on smoke and fly ash — problems mostly of concern in the combustion of coal. Prior to the middle 1950's, ordinances sometimes permitted emissions of smoke equivalent to as much as No. 3 on the Ringlemann scale. Since 1956, no ordinance has been passed which allows smoke of greater than No. 2. Under today's conditions of improved fuels, equipment and practice, a few communities have passed laws prohibiting emission of smoke of any density darker than Ringlemann No. 1. The majority of existing laws on fly ash emission in the U.S. limit emissions equivalent to 0.85 lb of fly ash per 1000 Ib of flue gas. In recent years, however, regulations which have been adopted give cognizance to the higher level of performance now obtainable with improved equipment. A comparison of the restrictions of five codes adopted since 1960 is given in Table I. The most stringent of these is the one for New York City which provides for a maximum emission of 0.6 lb fly ash per million Btu heat release (equivalent to roughly 0.51 lb/1000 lb of flue gas). The first comprehensive effort to restrict the emission of SO2 resulted from the passage of a 1937 law in St. Louis. This regulation stipulated that coal containing in excess of 23% volatile matter and 2% sulfur must be washed, thereby presumably producing some effective reduction in the input sulfur content. This was followed in 1949 by a Los Angeles County law which prohibited the emission of SO2 in concentrations greater than 0.2%. Most SO2-restrictive legislation passed since that date has been based on this limiting 0.2% SO2 by volume, although modifications are occasionally permitted under selected conditions, sometimes based on the fact that certain limiting ground concentrations are not exceeded — such as in the rules adopted by the San Francisco Bay area. To date, no legislation has been passed in the U.S. to limit the generation of nitrogen oxides from the combustion of fossil fuels. However, such oxides are considered to be of potential importance in air pollution control because of their possible detrimental effects on health and their reported role in the formation of photochemical smog. Interest in reducing oxides of nitrogen from powerplant and auto exhausts is increasing and regulations limiting their quantity can be expected in the future.

Jan 1, 1968

By A. J. de Witte

It has been observed that upon invasion of a sand containing oil or gas and connate water by mud filtrate the hydrocarbons are more rapidly flushed by the filtrate than is the connate water.' In time, it appears that the following displacement pattern emerges: oil (or gas) is being swept ahead by connate water which, in turn, is pushed by invading fluid. Eventually the connate water may have "banked up" sufficiently to form a zone of appreciable thickness leading the invading front. In the extreme case, Fig. 1 shows how the situation will develop.' Immediately adjacent to the borehole (radius r) there is a zone, the invaded zone proper, where the connate water has been flushed out completely and which contains only residual oil or gas and mud filtrate. This zone (extending out to a radius ri) grades more or less abruptly into the next (thickness A), which contains only residual hydrocarbon and connate water. Beyond it, again more or less sharply bounded, extends the virgin formation with the original interstitial water saturation Sw. The various saturations are indicated in Fig. 1. The three zones generally will be marked by resistivity contrasts owing to their different fluid contents. As the connate water is usually more saline than the invading fluid, the second zone having a high connate water saturation forms a concentric cylindrical ring or annulus of low resistivity R, around the borehole.',' It will be referred to as the "low zone." Fig. 1 shows schematically the resistivity profile. It is clear that the phenomenon of a low zone could not occur in invaded water sands. The presence of a low zone, therefore, would be a qualitative indication of a hydrocarbon-bearing formation. Granted that the phenomenon is real, if it were pronounced enough to be detected, one might thereby have a means of locating oil or gas in the ground. This is the aim of the "displacement logging" method.' In any case, the presence of a low resistivity zone will affect the reading of electric logs. Using conventional log interpretation techniques, one must be aware of this and, if necessary, correct for it. Whether correction or detection should be the goal will depend primarily on the magnitude of the effect. The type of log which is likely to be affected most is the induction log. The current pattern of an induction log is concentric with the borehole. Any concentric ring of low resistivity, therefore, will tend to become crowded by current lines. This zone of maximum current density will obscure the relative contributions to the current conduction by other portions of a bed making, for instance, the contribution of the virgin formation less significant and thus detracting from the value or the induction log as an R, reading device. The present note is intended to ascertain the maximum possible effect of a low zone in the extreme case portrayed by Fig. 1 on three commercially available types of induction logs, sc., 5 FF 27, 5 FF 40 and 3 F 60. Type 5 FF 27 or 5 PF 40 is currently run in conjunction with a 16 in. normal and SP log. This combination is gradually replacing the regular electric survey (two normals and one lateral) as a standard log in many areas. On the basis of the scheme of Fig. 1 (sharp boundaries between successive zones), the thickness of the low zone may be computed for various amounts of infiltration measured by the ratio of invaded zone diameter to hole diameter D,/d. Referring to Fig. 1, we can set up a material balance for the water displaced from the invaded zone proper and the water present in the low zone. The volume of connate water displaced from invaded zone upon flushing by mud filtrate is: V1 = e p (r2 - r2) ? Sw . where c = bed thickness ? = porosity Su, = original connate water satu-ration. The volume of water finally accu-mulated in the low zone is: Vz = e p [(r1 -?)2 - r2 4] ? Sw . where SW1 = 1 — Sor = "water" saturation in the invaded and low zones both. The volume of water originally present in the low zone is:

Jan 1, 1958

By R. L. Mauger, P. E. Damon, B. J. Giletti

This report includes the lead isotopic dating of a suite of galenas from Arizona and an application of the K-Ar method to the dating of a Laramide porphyry copper deposit, the Silver Bell Mining District. The lead isotopic data supports prior age assignments based upon geologic inference. The Silver Bell study illustrates the necessity of correlative geologic and petrographic investigations for the interpretation of the results of potassium-argon dating. LEAD ISOTOPIC DATING OF GALENAS FROM ARIZONA ORE DEPOSITS A group of galenas from Arizona ore deposits have been analyzed for lead isotope ratios. The results were used to calculate model ages by the. method of Russell, Stanton and Farquhar,l6 in which the age is calculated directly from the Pb206/Pb207 ratio. The use of Pb206/Pb207 ratios eliminates the errors inherent in measuring the abundance of Pb204. The Holmes-Houtermanns model is the other model commonly used for calculating model ages. Both models assume that any lead sample is composed of primeval and radiogenic components and the calculated age is the time at which the lead was extracted from its source area. Using the Holmes-Houtermanns model, a lead is ordinary if its isotopic ratios lie on an isochron. The growth curve that passes through the experimental point determines the U/Pb ratio in the source area. The RSF model assumes the source area for conformable leads is the mantle, that this has a uniform U/Pb ratio, and thus all ordinary leads must lie on a single growth curve, having a mantle U/Pb ratio. The definition of an ordinary lead differs between the models, and differences in age arise mainly from the assumptions made to evaluate parameters in the model equations. These assumptions depend on the hypothesis chosen to explain ore genesis. In the RSF Model, the Pb206/Pb207 ratio is derived as a function of time. The equation contains three undetermined parameters which are evaluated by assuming three known points lie on the curve. These are the following: 1) Primeval lead from the Canyon Diablo and Henbury meteorites, 2) Modern conformable leads which lie on Patterson's zero isochron, 3) Lead from the Bathurst, New Brunswick base metal deposits. The Bathurst deposits are postulated to be examples of "conformable base metal deposits", as proposed by Stanton.l9 A conformable deposit has a particular genetic history and, as a result, the orebody conforms to stratigraphic layering in the host rock. The metals are brought to the surface in volcanic rocks which originated in the mantle. Weathering products of these rocks, including sulfur and metals, accumulate in areas undergoing sedimentation. The formation of sulfide ion in the sediments by the action of sulfate reducing bacteria causes fixation of iron as pyrite. If the pyrite becomes concentrated in favorable stratigraphic horizons, any base metal deposit eventually formed by replacement of pyrite will have a strata bound character. Compaction and expulsion of pore water from the sediments at depth result in upward mobility of solutions containing soluble base metal chlorides. The strata with high pyrite content act as chemical traps for the base metal ions and replacement occurs. An important result of this general evolutionary model is that, if complete separation of lead and parent isotopes occurred during accumulation of the sediments, any ore deposit formed solely of metals derived from those sedimentary rocks will contain "conformable" lead. This would be true even if the actual ore deposit were formed at a later date by some epigenetic process. In this case, mineralization would be controlled by local conditions, and need not conform to stratigraphic layering. Also, any ore deposit containing lead derived from a mantle or mantle-like source, even though not conformable in Stanton's sense, will fall on the curve for conformable ores and thus give a meaningful model age. Model ages for Arizona galena deposits are listed in Table 1. Fig. 1 is a location map. Jerome-Humbolt District: Galenas from the United Verde mine at Jerome and the Iron King mine at Humbolt give model ages of 1750 m.y. and 1640 m.y. respectively. Both deposits are massive sulfide bodies in a host rock of older Precambrian

Jan 1, 1965

By P. M. Robinson, M. B. Bever

THE heats of formation at 273°K of the compounds InBi, In2Bi, and TIBi2 have been determined by metal solution calorimetry with bismuth as solvent. The published information on the thermody-namic properties of these compounds is limited. The structure of the compound InBi (mp, 383°K) is tetragonal of the PbO (B10) type.1 The compounds In2Bi (mp, 362°K) and TIBi2 (mp of highest-melting composition, 486°K) are now believed to have structures of the NizIn (B8) type273 and not the A1B2 (C32) structures reported earlier.4, 5 The compounds In2Bi and TIBi2 have limited homogeneity ranges at room temperature, while the compound InBi has no detectable homogeneity range.' Samples of the compounds were prepared from 99.99 pet Bi (Mallinckrodt Chemical Works), 99.999 pet in (American Smelting and Refining Co.), and 99.99 pet Tl (Amend Drug and Chemical Co.) by melting the components in sealed, evacuated Vycor tubes. The compounds InBi and in2Bi were of the stoichiometric compositions. The compound TIBi2 contained 40 pet T1; this composition was chosen because at room temperature the alloy of stoichiometric composition does not lie within the homogeneity range.' The melts were held approximately 50°C above the liquidus for 8 hr and then quenched into iced brine. The samples were homogenized for 48 hr at temperatures 20°C below the temperature of complete solidification. Metallographic examination of sections taken from the center and the ends of the ingots did not reveal second phases or segregation . The heats of formation were determined as the difference in the heat effects on solution of a compound and of a mechanical mixture of its components added from the reference temperature of 273°K to liquid bismuth at 623°K. The details of the calori-metric procedure and method of calculation have been described elsewhere? The calorimeter was calibrated by adding bismuth at 273°K to the bismuth bath at 623°K. The calculated values of the heats of formation of the compounds, listed in Table I, are based on a value of 4.96 keal per g-atom for the difference between the heat contents of bismuth at 273" and 623oK7 No published values are available for the heats of formation of the compounds InBi and In2Bi. Hultgren et a1.7 have calculated a value of -0.50 i 0.10 keal per g-atom for the heat of formation at 423°K of the compound TIBi2 containing 40 at. pet T1 from heat-content measurements in the range from 398" to 700°K and the heats of formation of the liquid alloy. The difference between this value and the value of -0.66 * 0.01 keal per g-atom at 273°K is too large to be attributed to a difference between the heat capacities of the compound and the components over the temperature range 273o to 423oK7 However, the direct determination of the heat of formation by a calorimetric method should give a more accurate value. The heats of formation at 273°K of the compounds InBi and In2Bi lie on a straight line when plotted as a function of composition, Fig. 1. The compound In2Bi, therefore, appears to be barely stable with respect to its neighboring phases. In order for the free energy vs composition curve to be concave towards the composition axis, the entropy change on formation of In2Bi must be more positive than that on formation of InBi. As the compounds are ordered,2, 3, 8 there is no configurational entropy change on formation. The difference between the entropy changes on formation of the two compounds, therefore, is probably associated with the vibrational entropies. In view of the low heats of formation, the changes in vibrational entropy, due to changes in bond strength on formation of the two compounds, are likely to be small. Owing to the volume contractions on formation of the two compounds,5 the vibrational entropies probably decrease slightly but the decrease of the vibrational entropy of InBi is expected to be larger than that of In2Bi.

Jan 1, 1965

By H. H. Klepfer, M. E. Snyder

UNTIL very recently, the commercial availability of the rare earths as metals has been very limited. Fabrication of mill products from these metals has not been studied in most cases. This note reports the results of the development of fabrication techniques for thulium. Thulium has a melting point of about 1550°C and a hexagonal close-packed crystal structure. It oxidizes in air to give a black oxide (TmzO,). The procedures for producing thin thulium foil were developed on one ingot weighing about 220 g and were subsequently applied in processing about a pound of metal to foil. Several alternates for the various fabrication steps were investigated and will be discussed. The metal fabricated was in the form of commercial chill-cast ingots 1 in. in diam and 2 in. long weighing approximately 220 g. Impurities in the ingots were reported by the vendor to be 4000 ppni tantalum, 2000 ppm calcium, 200 ppm nickel, 100 to 200 ppm iron, 100 ppm europium, and less than 100 ppm copper, lutetium, and ytterbium. In addition to these impurities, several salt-like inclusions as large as l/8 in. in diam were revealed along the center line of the one ingot sectioned. Preliminary tests indicated that small wafers cut from the as-cast ingot would not fabricate readily by rolling. Forging of copper-jacketed wafers was therefore attempted. At 1550"F forging was satisfactory but an apparent reaction of copper with thulium demanded investigation of lower temperatures. Therefore, the remainder of the test ingot was forged at 1450°F—with only minor cracking. All ingots forged were inserted into copper tubes of 1 in. 1D and 0.125-in. wall thickness. The jackets were sealed by flattening the ends of the tubes and welding under helium. Heating time at 1450°F was 30 min. Press forgings of 1/8 in. per pass were used, followed by 10 min reheats. When the ingots had been squared and reduced to 0.250 in. in thickness, the original copper jacket was stripped off and replaced by a new jacket in preparation for hot rolling. Hot rolling at 1450" F without edge cracking was readily accomplished after forging. Excellent results were obtained with 10 pct reductions of thickness followed by 5 to 10 min reheats. After reduction of thickness from 0.250 to 0.100 in. the copper jacket was removed. It was found, in fact, that hot rolling in air was possible. A tenacious black oxide similar to that seen on zirconium was formed during 3 min reheats at 1450°F. Reduction in air to 0.010 in. foil was possible taking 10 pct reductions per pass. The oxide coat formed during hot rolling in air could best be removed by sand blasting and pickling. Common pickling solutions containing polar solvents were found to attack the metal too rapidly and a concentrated nitric-hydrofluoric acid mixture attacked neither the oxide nor the metal. The most satisfactory pickling solution was 52 vol pct concentrated nitric acid-48 vol pct glacial acetic acid. After forging to 0.250 in., vacuum annealing and cold rolling was found to be another satisfactory alternate to hot rolling in copper jackets. After forging. a hardness of Rockwell B76 was found. Annealing in vacuum (2 X l0-5 mm of Hg) for 1 hr at 1200"F did not alter this value. Annealing at 1470"F for 1 hr brought the hardness down to R;]63. With cold rolling (5 pct per pass) the hardness returned to about R,1:76 after 20 pct reduction and edge cracking became noticeable. However, cold rolling to a total of 40 pct reduction in thickness (RI,,83) was possible before edge crack propagation became serious. Good surface finish was obtained, and the metal loss due to oxidation was minimized by cold rolling and vacuum annealing. Using this procedure the yield of 1.25 in. wide by 0.010 in. thick foil from a 1-in. diam ingot was about 40 pct.

Jan 1, 1961

By Walter Rose

In a recent technical note, Owen Thornton&apos; suggests that wetting liquid relative permeability may be derived from the relationship: where Pd/Pc is the ratio of displacement pressure to capillary pressure at the wetting liquid satdration Sw, and I is the resistivity index at this saturation. Thornton shows that this expression gives values for relative permeability in good agreement with those experimentally determined by Leverett. We have recently developed another expression for krw which seems preferable to Equation 1 since it requires fewer experimental data for its verification. This equation may be derived in the following manner. Making use of the analogy between mass transfer of fluid in a porous medium and electrical conductivity through the fluid in the same medium, we can postulate a hydraulic formation resistivity factor analogous to an electrical formation resistivity factor. From Poiseuille&apos;s Law the resistivity to flow in a tube of radius R is of the form 8u/R², where is the viscosity of the fluid. Similarly7 the resistivity to flow in a porous medium of the same dimensions as the tube is given by D&apos;Arcy&apos;s Law as where k is the permeability of the medium. The hydraulic formation resistivity factor, Ff, is thus the quotient of these two resistivities, or: But permeability is defined by the KO-zeny equation as: where: r is the average pore radius of the porous medium, L. is the average tortuous length of the average pore, + is the porosity, and L is the bed length. This gives Fr as a function of the (La/L) .-.-.tortuosity ratio and porosity, or in explicit form: where: T = (La/L)&apos;, and N is the number of pores of radius, r, which will be found in any cross sectional area, x R&apos;. That is, N = R²/r². It will be clear from the foregoing considerations that the hydraulic formation fat. tor should be dependent on the value of R which in every case must be arbitrarily selected. By analogy, the hydrauiic formation factor, Fcf, characterizing the porous medium at Sw<1,will be given by: R² NT. where ke is the effective wetting liquid permeability obtaining at Sw<l, and Te is the square of the (Las/L) ratio defined by Thornton. Therefore, the hydraulic resistivity index, If, is: However, Thrnton gives I as: and it follows that: Equation 2, when checked against the data of wyckoff and Botset² and some of the Morse et al data³, gives comparisons as shown at the bottom of this page. Computations based on Leverett&apos;s data, quoted by Thorntaon, gives computed krw values somewhat lower than the experimental values. Moreover, other instances can be cited where we find that neither Equation 1 nor Equa tion 2 is checked by experimental data. The work of Botset&apos; on the Nichols Buff sandstone and Morse et a1 on oil wetted Bradford sandstone are such instances. A reason for these discrepancies may lie in the fact that if both Equations 1 and 2 are presumed always to give accurate values for krw, it follows that PC = Pd/Sw". Such an expression, it is well known, is too simple to describe accurately the capillary pressure behavior of all porous media. However, it is believed that a particular advantage resides on the use of Equation 2, since it is not dependent for its utility on a knowledge of the capillary pressures obtaining in dynamic flow svsterns. No practical technique for measuring the capillary pressures characterizing the fluid distributions in dynamic flow systems has yet been proposed. Acknowledgment is given to Dr. Paul D. Foote, executive vice-president of the Gulf Research and Development Company, for permission to publish this note. 1. Owen F. Thornton, "A Note on the Valuation of Relative Permeability," J. Pet. Tech., 1:7, Section 1, July, 1949. 2. R. D. Wyckoff arid H. G. Botset, "The Flow of Gas-Liquid Mixtures through Porous Media," Physics, 7:325-345, 1936. 3. R. A. Morse, P: L. Terwilliger, .and S. T. Yuster, "Relative Permeability Measurements of Small Core Samples," Oil and Gas J., 46:109-125, Aug. 23, 1947. 4. H. G. Botset, "Flow of Gas-Liquid Mixtures Through Consolidated Sand," Trans. AIME, 136:91, 1940.

Jan 1, 1949

By Walter Rose

In a recent technical note, Owen Thornton&apos; suggests that wetting liquid relative permeability may be derived from the relationship: where Pd/Pc is the ratio of displacement pressure to capillary pressure at the wetting liquid satdration Sw, and I is the resistivity index at this saturation. Thornton shows that this expression gives values for relative permeability in good agreement with those experimentally determined by Leverett. We have recently developed another expression for krw which seems preferable to Equation 1 since it requires fewer experimental data for its verification. This equation may be derived in the following manner. Making use of the analogy between mass transfer of fluid in a porous medium and electrical conductivity through the fluid in the same medium, we can postulate a hydraulic formation resistivity factor analogous to an electrical formation resistivity factor. From Poiseuille&apos;s Law the resistivity to flow in a tube of radius R is of the form 8u/R², where is the viscosity of the fluid. Similarly7 the resistivity to flow in a porous medium of the same dimensions as the tube is given by D&apos;Arcy&apos;s Law as where k is the permeability of the medium. The hydraulic formation resistivity factor, Ff, is thus the quotient of these two resistivities, or: But permeability is defined by the KO-zeny equation as: where: r is the average pore radius of the porous medium, L. is the average tortuous length of the average pore, + is the porosity, and L is the bed length. This gives Fr as a function of the (La/L) .-.-.tortuosity ratio and porosity, or in explicit form: where: T = (La/L)&apos;, and N is the number of pores of radius, r, which will be found in any cross sectional area, x R&apos;. That is, N = R²/r². It will be clear from the foregoing considerations that the hydraulic formation fat. tor should be dependent on the value of R which in every case must be arbitrarily selected. By analogy, the hydrauiic formation factor, Fcf, characterizing the porous medium at Sw<1,will be given by: R² NT. where ke is the effective wetting liquid permeability obtaining at Sw<l, and Te is the square of the (Las/L) ratio defined by Thornton. Therefore, the hydraulic resistivity index, If, is: However, Thrnton gives I as: and it follows that: Equation 2, when checked against the data of wyckoff and Botset² and some of the Morse et al data³, gives comparisons as shown at the bottom of this page. Computations based on Leverett&apos;s data, quoted by Thorntaon, gives computed krw values somewhat lower than the experimental values. Moreover, other instances can be cited where we find that neither Equation 1 nor Equa tion 2 is checked by experimental data. The work of Botset&apos; on the Nichols Buff sandstone and Morse et a1 on oil wetted Bradford sandstone are such instances. A reason for these discrepancies may lie in the fact that if both Equations 1 and 2 are presumed always to give accurate values for krw, it follows that PC = Pd/Sw". Such an expression, it is well known, is too simple to describe accurately the capillary pressure behavior of all porous media. However, it is believed that a particular advantage resides on the use of Equation 2, since it is not dependent for its utility on a knowledge of the capillary pressures obtaining in dynamic flow svsterns. No practical technique for measuring the capillary pressures characterizing the fluid distributions in dynamic flow systems has yet been proposed. Acknowledgment is given to Dr. Paul D. Foote, executive vice-president of the Gulf Research and Development Company, for permission to publish this note. 1. Owen F. Thornton, "A Note on the Valuation of Relative Permeability," J. Pet. Tech., 1:7, Section 1, July, 1949. 2. R. D. Wyckoff arid H. G. Botset, "The Flow of Gas-Liquid Mixtures through Porous Media," Physics, 7:325-345, 1936. 3. R. A. Morse, P: L. Terwilliger, .and S. T. Yuster, "Relative Permeability Measurements of Small Core Samples," Oil and Gas J., 46:109-125, Aug. 23, 1947. 4. H. G. Botset, "Flow of Gas-Liquid Mixtures Through Consolidated Sand," Trans. AIME, 136:91, 1940.

Jan 1, 1949

PURPOSE OF UPDATED ESTIMATE FOR MINERAL INVENTORY BLOCKS During the production stage of an open pit porphyry copper mine it was observed that the expected production grade, as determined from the block inventory estimates, often differed greatly from the head grade of ore delivered to the mill. It was determined that, for purposes of production scheduling and monthly forecasting, better in situ grade estimates for mining blocks were necessary. Because all of the bench blastholes were being sampled, and periodic holes were being drilled into the next underlying bench, much more sample data existed than was being used for estimating the grade of nearby mining blocks. It was reasoned that, by periodically updating the mineral inventory block file over the benches scheduled for mining in the next time period using all existing data, better estimates and forecasts for production grade could be made. BLASTHOLE SAMPLE DATA MANAGEMENT The collars of all blastholes in each mined bench were surveyed and assays run on the cuttings representing the full 12-m (40-ft) bench height. The blastholes range from about 5 to 6 m (16 to 19 ft) apart along the front of the bench as well as perpendicular to the front. Overbreakage often left larger spacings between successive blasts. Fig. 17-1 is a plan map of blastholes on a typical mine bench. All blasthole data were keypunched to a format resembling that of the 12-m (40-ft) composite assay data file as follows: [Collar Coordinates Compite Assoy Hole ID Northing East Elewotion Total Copper Oxide Copper] Because of the many blasthole samples available, and in order not to use excessive computer time for running kriged estimates for 30.5 X 30.5 X 12-m (100 X 100 X 40 ft) blocks adjacent to and beneath the mined area, the decision was made to average all the blastholes falling with each 15.2 X 15.2-m (50 X 50-ft) mined block, and use the mean value of the samples as a regionalized variable for purposes of assigning kriged estimates and estimation variance to adjacent unmined blocks. In other words, instead of using individual blasthole samples for making kriged estimates, the holes were grouped by blocks and assay values were averaged and assigned to the centroid of the holes within the block, which was then treated as a single regional variable for purposes of kriging. See Fig. 17-2. VARIOGRAM COMPUTATIONS AND KRlGlNG RESULTS With the many blasthole samples it was possible to compute directional and vertical experimental variograms for both sulfide and nonsulfide copper assays falling within the enriched mineral zone for the full 12-m (40-ft) sample support. Due to the close spaced drilling, excellent definition of the experimental variograms was possible, and the spherical model exhibited good fits. A three-dimensional kriging program was then run over the two or three mine benches involved in the inventory update, and estimated grades reassigned to all mining blocks falling within the range of the new blasthole assay data according to the anisotropisms of the deposit. Better confidence limits could then be assigned to scheduled mining blocks and better short- range forecasts made. An interactive kriging computer program was also applied for the purpose of determining the kriging variance or estimation of error for larger, irregular mining blocks representing the monthly production from a particular bench. The interactive program permitted the operator to enter the limits of the irregular block onto the screen of a cathode ray tube (CRT) as a series of points around the perimeter of individual gridded blocks making up the larger irregular block. The computer then was programmed to calculate the kriging variance of the larger block using all samples fall- ing within range. Thus the limits of estimated grade could be established at any confidence level. Fig. 17-3 illustrates the output from the interactive kriging program showing the sample points entering into the grade and kriging variance computations, and also the kriging coefficient assigned by the computer to each sample.

Jan 1, 1980

By E. N. Smith, J. C. Redmond

The increasing need for materials capable of withstanding higher operating temperatures for various applications such as gas turbine blading and other parts, rocket nozzles, and many industrial applications, has brought consideration of cemented carbide compositions. The well known usefulness of cemented carbides as tool materials is attributable to their ability to retain their strength and hardness at much higher temperatures than even complex alloys. However, it has been found that the temperatures encountered in cutting operations do not approach by several hundred degrees1 those involved in the applications mentioned above where the interest is in materials possessing strength and resistance to oxidation at temperatures of 1800°F and above. At these latter temperatures, the tool type compositions which are made up essentially of tungsten carbide are found to oxidize very rapidly and to produce oxidation products of a character which offer no protection to the remaining body. As a further consideration, the density of the tungsten carbide type compositions is high, from about 8.0 to 15.0. The refractory metal carbides as a class are the highest melting materials known as shown by Table 1 which summarizes the available data from the literature for the carbides of the elements which are sufficiently available for consideration for these uses. The density is also included in the table, since as mentioned above it is an important consideration in many of the applications for which the materials would be considered. It has been established that in the tool compositions the mechanism of sintering with cobalt is such as to result in a continuous carbide skeleton and that the properties of the sintered composition are thus essen- tially those of the carbide.2 On the hypothesis that this mechanism holds to a greater or less degree in cementing most of the refractory metal carbides with an auxiliary metal, it appears from Table 1 that titanium carbide compositions would offer possibilities for a high temperature material. Titanium carbide has extensive use for supplementing the properties of tungsten carbide in tool compositions. Although the literature contains several references to compositions containing only titanium carbide with an auxiliary metal,3,4,5,6 it may be inferred from the meager data that such compositions were deficient in strength and were considered to have poor oxidation resistance.7 Kieffer, for instance, reports the transverse rupture strength of a hot pressed TiC composition at 100,000 psi as compared to up to 350,000 psi for WC compositions. The work described herein was undertaken to determine the properties of compositions consisting of titanium carbide and an auxiliary metal and to improve the oxidation resistance of such compositions. It appeared possible that the inclusion of one or more other carbides with titanium carbide might improve the oxidation resistance and also that this might be more desirable than other means from the point of view of maintaining the highest possible softening point. Consideration of the available carbides in Table 1 suggests tantalum and columbium carbides because of their high melting points and general refractoriness. The work on improving oxidation resistance was concentrated on the addition of tantalum carbide or mixtures of tantalum and columbium carbide. The auxiliary metals used included cobalt, nickel and iron. It was also desired to learn the general physical properties of these compositions. Experimental Procedure The compositions used in this study were made by the usual powder metallurgy procedure applicable to cemented tungsten carbide compositions. The powdered carbide or carbides and auxiliary metal were milled together out of contact with air. In some cases cemented tungsten carbide balls and in other instances steel balls were used to eliminate any effect of tungsten carbide contamination. A temporary binder, paraffin, was then included in the mix and slugs or ingots were pressed with care to obtain as uniform pressing as possible. The ingots were presintered and the various shapes of test specimens were formed by machining, making the proper allowance for shrinkage during sintering. Thereafter the shapes were sintered in vacuum at temperatures of from 2800 to 3500°F. Final grinding to size was carried out by diamond wheels under coolant. The titanium carbide used contained a minimum of 19.50 pet total carbon and a total of 0.50 pet metallic impurities as indicated by chemical and spectrographic analysis. It was found by X ray diffraction examination with

Jan 1, 1950

By Bob J. Rawicki

TYPES OF WET DUST COLLECTORS Basically, there are two types of wettype dust collectors. One is mechanical, incorporating pumps, motors, fans, sprays, filters, or flooded beds. These come in many forms, but their operating principle is basically the same. The screen or flooded bed is wetted by a series of sprays. Polluted air is drawn in through a fan; here the dust particles impinge on the screen and are flushed off into a settlement tank. The air then passes through spray eliminators to the atmosphere. A great deal of time and money has been spent developing this type of collector. In laboratory tests where clean water at constant pressure, air at constant volume, and dust at low but constant mass to volume were applied, these collectors gave high percentage dust collection efficiencies. Theoretically and technically, they are very good. In practice, however, under rugged mining conditions where there are no constants, and dust load, particle size, water pressure, and volume of air vary with time, their efficiency varies dramatically. Some of the most common problems are: 1. clogging of screens 2. blocking of nozzles 3. replacement. of damaged pump stators Any of the above result in loss of air flow and dust collection efficiency. Generally, there are so many mechanical moving parts, something is always breaking down. This type of collector is very expensive in parts and labor to maintain. The second type of dust collector, which I personally developed, with the help of the British National Coal Board of Great Britain, has none of the above mentioned problems and it works on a completely different principle. My collector, the Mark III Precipitaire Wet-Type Dust Collector, (an improved version of previously successful models) has no internal moving parts, no flooded beds, screens, spray nozzles, etc. It is highly efficient - 99.8% efficiency on total collected dust and 97% on respirable dust. The pressure drop through the collector is constant, hence, air flow remains constant. There is no maintenance expenditure, be it labor or parts. The only maintenance that there is, is desludging. This, of course, varies with the dust load. With a low dust load, desludqinq may only be required every 4 weeks, but this is a very simple operation that can be handled by unskilled men and it takes only a very short time. My collector is powered by a ventilation fan that is placed on the clean air side for a longer fan and motor life. Dust is extracted solely by water action. The collector&apos;s operating principle is: Dust laden air is drawn in along ducting into a tapered scrubber section, shaped to create a self-induced curtain of water. This action washes the dust particles from the air and the collected dust settles into the bottom of the tank. Cleaned air is then exhausted via a series of spray eliminators where it is completely dried and on into the atmosphere. These collectors have been sold in Great Britain, to the National Coal Board, and throughout Europe. They have been working very well for many years now. GENERAL APPLICATIONS Dust is very dangerous; silicosis and black lung kill and disable many workers each year. Many of the mine explosions can be attributed to coal dust. By the use and application of dust collectors, these deaths and disasters can be avoided. There really is no excuse. However good and efficient a dust collector may be, it&apos;s positioning and installation arrangement is critical. The collector may remove 99% of dust supplied to it, but it takes the skill and experience of a dust control engineer to design and fit ducting, hoods, etc. to remove polluted air at source and deliver it to the collector. In some cases this is very simple; in the case of longwall machines, it is very difficult. In general, dust collectors should be used to collect dust at every point of dust generation. I have noted a great reliance on dilution as a means of dust control. I personally think this is a dangerous and unreliable method. As mines become deeper and the use of larger and more productive equipment becomes more common, reliance on dilution will be impossible and dust collectors will be essential.

Jan 1, 1982

By H. Sigurdson, T. V. Cherian, C. H. Moore

The phenomenon of recovery of cold-worked metals is interesting not only because of its practical importance but also because of its fundamental significance in solid state reactions. Although extensive investigations1,2 have already been made in an attempt to discover the mechanics of the recovery process, many of the observations have not yet been satisfactorily correlated to provide a completely consistent model for the process. The wide differences in the recovery rates of various properties can be cited as a typical example of one of the difficulties that are encountered. Frequently, for example, the electrical resistance will have almost completely recovered before any recovery in tensile strength can be detected. Of course, such differences in the recovery rates of different properties might be explained by assuming that each property is a unique function of the work-hardened state, and consequently each property exhibits its own unique recovery rate. The assumption that different properties are uniquely related to the work-hardened state cannot be denied. On the other hand, the properties that recover at different rates often exhibit more or less parallel changes upon work-hardening. This suggests that the microstructural changes attending recovery are not exactly the reverse of the changes attending work-hardening. Several types of imperfections must be postulated in order to account for this apparent anomaly. The different recovery rates for various properties, then, are due to the different recovery rales of the type of imperfection to which each property is most sensitive as well as the unique dependence of each property on the cold- worked state. This concept assumes that a simple model of the work-hardened state consisting only of one type of imperfection, such as Taylor&apos;s type of dislocation patterns, is inadequate to cope with the diversified phenomena attending work-hardening and recovery. Although current models for the work-hardened state are not useful for describing all aspects of the recovery process, the general trends of the recovery of each postulated type of imperfection as a function of time and temperature should be at least qualitatively deducible from the rather well developed laws of kinetics of reactions in the solid state. Consequently, recovery data might prove useful for elucidating some aspects of the complexities of the work-hardened state of metals. A preliminary attempt to study work-hardening by investigating recovery rates of cold-worked metals is outlined in the following pages of this report. Experimental Procedure Many properties recover when cold-worked metals are annealed below their recrystallization temperature. Therefore, electrical resistivity, thermal electromotive force, X ray diffraction line widths, X ray diffraction line intensities, elastic spring back, density and other physical and chemical properties have been used to study the recovery process. Major interest, however, has generally been directed toward the recovery of the mechanical properties such as hardness, yield strength, and tensile strength. But a search of the literature suggests that the effect of recovery on the true stress-true strain curve has been neglected, in spite of the current recognition of the fundamental importance of such an investigation. An investigation on the effect of recovery treatment on the true stress-true strain curves in tension, therefore, was undertaken in the present study. Commercially pure aluminum (99. + pet Al) in the form of 0.100 in. thick rolled sheet of 2S-O aluminum alloy was selected as the material for this investigation because rather extensive correlatable data are already available on the recovery of some of its properties. Tensile specimens having a 6 in. long gauge section and a uniform reduced section width of 0.500 in. were machined from the sheet in accordance with a design that has previously been reported.3 All specimens were selected with their axes aligned in the rolling direction. In order to eliminate the effects of previous work-hardening and the effects of machining, the specimens were annealed for 15 min. at 750°F before testing. During tensile testing the loads were measured by means of a proving ring (sensitive to 1/2 lb) in series with the specimen.4 Strains were determined from the extension of a rack and pinion strain gauge sensitive to a strain of + 0.0001. The stress was recorded as the true stress, namely

Jan 1, 1950

By Margaret V. Doyle, R. W. Powers

INTERNAL friction measurements, carried out as functions of temperature, have been used extensively to obtain data on the mobility of interstitial impurities in the Group V metals, vanadium, colum-bium, and tantalum. For a given frequency of vibration, the internal friction peaks at a temperature which depends principally on the particular metal and interstitial impurity. Relaxation times and corresponding activation energies are determined from a knowledge of such peak temperatures for various vibration frequencies. These experimentally determined relaxation times are related to the average jump frequencies of the interstitial impurity atoms using the model of Snoek1 and can be related to macroscopic diffusion coefficients as shown by Polder.2 In these Group V elements, the impurity whose diffusion rate has been the hardest to pin down by the internal friction technique has been carbon. The experimental difficulties associated with carbon in these metals were not evident in the earliest investigation carried out by Ke on tantalum.3 The carbon peak was thought to occur near 140°C at 1 cps on the basis of the enhancement of the height of this peak measured on specimens which had been coated with colloidal carbon and heated in vacuo at 1200°C. Presumably on the basis of similar evidence, Wert identified a peak also near 140°C at 1 cps in columbium with the diffusion of carbon in that element." Later, the difficulties of admitting carbon into these metals without simultaneously introducing oxygen were pointed out by Ang and Wert- nd by Marx, Baker, and Sivertsen.6 The latter suggested that the peaks attributed to carbon in tantalum and columbium might arise from the presence of traces of oxygen, adventitiously absorbed during the carburization. Such suspicion was confirmed, at least in tantalum, when it was found possible to prepare C-Ta specimens without contamination by oxygen. Specimens sci prepared, aitnough high in carbon content, gave no evidence of a damping peak near 140°C.7 Recently, the authors were able to Show that carbon diffuses in tantalum at a rate only slightly slower than that for nitrogen.8 At 1 cps, the carbon peak lies at 34g °C, only a few degrees above the nitrogen peak at 341 °C. However, the two could be distinguished by a different stability with respect to aging at 400 °C and by a different dependence of peak height on solute concentration. On the basis of this experience with tantalum, it appeared worthwhile to search also for a carbon peak in columbium in the vicinity of the nitrogen peak." In this report, the location of the carbon peak in columbium is made known and its position relative to the nitrogen peak is compared with the authors&apos; previous findings in tantalum. In addition, a second source of internal friction seen only in specimens of very high purity is discussed. Columbium was obtained from the Fansteel Metallurgical Corp. as either 30 or 40 mil diam stock. The manufacturer gives the following analysis as typical of this material: 99.4 pct minimum Cb, 0.10 pct maximum C, 0.5 pct Ta, 0.04 pct Ti, 0.02 pct maximum Si, and 0.01 pct maximum Fe. Wires 13 in. long were degassed near 2200°C for 4 to 16 hr in an ultimate vacuum better than 2.10" mm Hg in an apparatus described previously.8 This treatment reduced the oxygen and nitrogen damping peaks observed in the as-received material to less than 0.0001 in Q-1 above background. The latter value varied somewhat from specimen to specimen in the range 0.0004 to 0.0008, depending chiefly on the wire diameter and the frequency of vibration. The high temperature evacuation left the specimen in a rather soft condition, the Vhn being reduced to 40 from a value near 121 in the as-received conGi~iur~. The gram size was about the diameter of the specimen. After outgassing, specimens were loaded with carbon by absorption of such hydracarbons as methane or ethylene. Carbon was also added to a few speclrnens by heating them at

Jan 1, 1958

By A. H. Geisler, J. H. Keeler

Preferred orientations in rolled and annealed titanium sheets were determined by the Geiger counter spectrometer X-ray diffraction technique. Five annealing textures dependent upon the temperature range of annealing were found, and in order of increasing annealing temperature pendent upon the temperature range are: 1—a deformation like texture, 2—a rotated inorder a-recrystallization temperature texture, 3-a retained u-recrysraIlization texture, on annealing at lower temperatures of the ß-region, 4—a transformation texture based on recrystallized a and predicted by the Burgers&apos; relationship, and 5—a ,ß-cube texture. These results are examined in terms of current theories of recrystallization textures. UMEROUS investigators have described the tex- ture obtained by cold rolling the hexagonal metals, titanium, zirconium, and beryllium, which have c/a ratios less than that of ideal packing, 1.633. The basal planes are rotated out of the rolling plane, about the rolling direction, so that the basal poles are tilted toward the transverse direction as shown schematically in Fig. la. In all instances but one,&apos; it was also reported that the [1010] direction was parallel to the rolling direction (see Fig. lb). Hot rolling has been reported as causing a similar tilt of the basal poles in the transverse direction (see Fig. la) and causing the [1010] direction also to be parallel to the rolling direction as shown schematically in Fig. lb. Annealing after deformation does not appreciably change the tilt of the basal poles in the transverse direction." Beryllium2-7 continues to have the [1010] direction in the rolling direction after annealing, and similar observations for titanium and zirconium&apos; . have been reported for annealing at fairly low temperatures, again as in Fig. lb. At higher annealing temperatures, however, the recrystallized grains of titanium" and zirconium have an orientation such that the [1120] direction is approximately in the rolling direction, although the basal poles are still inclined in the transverse direction. Figs. la and lc show the resulting orientations schematically. This change in orientation has been described as a nominally ±30° rotation of the hexagonal crystallites about the basal poles of the cold rolled texture and is apparent from the results which are summarized in Table I for investigations with the X-ray diffraction technique employing film. The angles y, , and ß are indicated in Fig. 2 which represents the stereographic projection of (1070) poles for the mean orientation of a pole figure. Texture determinations for titanium using the Geiger counter spectrometer have provided similar results except that in some instances additional components of the texture were proposed, as shown by the summary of data in the upper half of Table 11. On the other hand, the spectrometer technique, when applied to zirconium,* has revealed a splitting Recently completed studies of the textures of annealed zirconium", show zirconium to possess textures very similar to those reported here for titanium. Therefore, much of this discussion will include zirconium by virtue of its close similarity to titanium in pref erred orientations. of the intense areas of the pole figure for samples annealed at 600°C. This splitting could be described by a 7" rotation of the tilt axis about the normal to the rolling plane. Such a splitting for the annealed texture relative to the cold rolled texture was not observed in other determinations for either zirconium or titanium using the less sensitive film X-ray methoe and makes the relationship between the two types of texture more complex than the simple rotation about the (0001) pole based on film work. The more precise investigations on zirconium permit the descriptions in the lower part of Table 11, which show that the texture depends quantitatively on the temperature of annealing. When zirconium is annealed at temperatures up to 400°C, the texture is similar to the cold rolled texture, while annealing in the range 500" to 900°C produces a texture which is only approximately described as [11%] in the rolling direction. More precisely described results for zirconium show that the two types of splitting ( 1—about an axis in the rolling plane through an angle given in the second column in Table II and 2—about the normal to the rolling plane through an angle given in the third column of Table 11) depend on annealing temperature. The [1120 is the rolling direction only when the annealing temperature is in the vicinity of 900°C

Jan 1, 1957

By W. L. Martin, J. N. Dew, H. B. Steves, M. L. Powers

This paper presents and discusses the results obtained during a pilot test in the Loco field in southern Okla homa. The test was conducted in a 2%-acre pattern that was part of a 20-acre conventional waterflood pilot area. The conventional flood was well past its economic limit when the hot waterflood was initiated to obtain technical information and operating experience. Temperature data from nine observation wells showed that about 75 percent of the pattern was affected by heat, and that heat losses were severe in the pilot pattern. About 60 percent of the injected heat was lost to overburden and underburden zones during hot-water injection. Wellbore heat losses were held at tolerable levels at the shallow depth of the test by providing a low-pressure air annulus between the injection tubing and well casing. Hot water provided water injectivity increases of 200 to 400 percent. The hot water channeled across the lower portion of the oil sand in three directions through zones of relatively high water saturation. There was no conclusive evidence that natural fractures or pressure partings aflected the flow of fluids in the pilot pattern; however response undoubtedly was aflected by localized pressure gradients and by injection-production rate ratios. The results showed that hot waterflooding increased oil recovery in a reservoir containing 600 cp oil. The total tertiary oil production from the pilot pattern area was 3,896 bbl or about 156 bbl/acre-ft swept by heat. The corre.sponding WOR was about 34:I. Description of the Hot Waterflood Process To hot waterflood an oil reservoir, water is injected that has been heated to a temperature substantially higher than the original reservoir temperature, but lower than the vaporization temperature of water at the prevailing pressures. In the reservoir the hot water flows continuously into cooler sand and rapidly loses heat to the sand until it has been cooled to the original reservoir temperature. Thus, a heated zone and a region or "bank" of cooled water begins to accumulate immediately after hot water injection is started. This bank of cooled water continues to grow ahead of the heated zone, which also grows, but at a slower rate. This occurs because heat transfer is almost instantaneous, and the ratio of heat capacities of water to rock is such that two or three unit PV of hot water must be injected to heat a given unit bulk volume of the reservoir. The primary displacement mechanism is the same for both hot and conventional cold waterfloods; i.e., "piston" displacement occurs at the original reservoir temperature. The incremental benefits of hot waterflooding usually occur long after the breakthrough of cold water at producing wells, and the increased oil recovery necessarily is accomplished with high WOR&apos;s (water-oil ratios). Heat decreases the viscosity and density of oil and water. These effects result in more rapid and increased recovery of secondary or tertiary oil. If the cost of the required heat is low enough, the ultimate oil recovery of a hot waterflood should be increased substantially over what would be expected at the economic limit of a conventional cold waterflood. The economic benefits of any hot-fluid injection project depend primarily upon the cost of the heat required to produce more oil at an increased rate. This cost depends in part upon the amount of heat lost to surrounding formations. Heat loss depends upon reservoir thickness, water injection rate and temperature, the depth of the formation, well spacing, and the characteristics of the reservoir rock and surrounding formations. In general, percentage heat losses decrease as injection rate and reservoir thickness increase. Although it is an old idea, hot-water injection has not received widespread field application in the oil industry as a drive process. Much of the original field work with hot water was done in Pennsylvania fields where water permeabilities and injection rates are low. In these cases, hot-water injection was used primarily as a means of increasing in-jectivities — not as a recovery process. Ramey&apos; recently has published an excellent review of the development of hot fluid injection processes. Recent publications"&apos; and our own experiences now indicate that hot water or steam injection also can effectively increase the oil recovery from reservoirs containing viscous crudes. However, the economics of these methods as displacement processes have not been established. Several theoretical predictive techniques have been published (for a review see Ramey&apos;), but simplifying assumptions make

Jan 1, 1969