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By P. K. Currie, J. V. Walters

Solution mining of magnesium salts (Carnallite and Bischofite) is taking place near Veendam, in the northeast Netherlands. The magnesium salts are used for the production of magnesia sinter, used as a base material for the manufacture of refractory bricks for furnaces. This paper gives, first, an introduction to the operational procedure followed in Veendam, with the emphasis on some of the unique aspects, and then discusses rock-mechanical aspects of the cavity behaviour. The creep behaviour of magnesium salts is described, and Finite Element calculations on cavity stability and potential surface subsidence are discussed.

Jan 1, 1985

[A AALSETH, Earl P (M 51) Consult Geol Engr 2019 Eldorado Dr, Billings, Monl ABADIE, Henry G (M 43) Asst to Supt of Oper Long Beach Oil Dev Co 925 Harbor Plaza, Long Beach 2, Calif ABBE, T Waldo (A 56) Asst Dist Oper Supt Sun Oil Co PO Box 66, Stowell, Tax ABBOTT, Fred E Jr (M 59) Abqaiq Dist Mgr Arabian American Oil Co Dhahran, Saudi, Arabia ABBOTT, Clifton T (A 60) Apartado 677, Maracaibo, Venezuela ABBOTT, J L (M 56) Partner Pritchard & Abbott 1010 Ft Worth Natl Bk Bldg, Fort Worth 2, Tex ABBOTT, Richard H (J 51) Partner Abbott & Howard 422 W Hermosa San Antonio 12, Tex .ABBOTT, William G (M 50) Div Mgr Rice Engrg & Operating Inc Box 1142, Hobbs, N. M. ABDERHALDEN, Ronald R (J 59) Prod Engr Continental Oil Co Box 100, Carmi, III ABEL, Malcolm D (J 50) Gen Partner Abel & Bancroft PO Box 865, Midland, Tex ABEL, Ramond L (M 28) Prof of Chem Eng Univ of Pittsburgh, Pittsburgh 13, Pa ABEL, Wayne K (M 57) 1513 Patton Drive, Odessa, Tex ABELL, Joseph M (J 55) Petr Engr Tennessee Gas & Oil Co Box 1772, Casper, Wyo ABELS, Thomas A (J 60) Exploit Engr Shell Oil Co Box 193, New Orleans 3, La ABENDROTH, G F (M 46) Dring Prod Supt Lyons & Logan 1500 Beck Bldg, Shreveport, La ABERCROMBIE, Bolling A (M 60) Mgr Product Planning Axelson-Garrett PO Box 2427, Longview, Tex ABERCROMBIE, C T (160) Prod Engr Cities Ser Petr Co Box 472, Brownfield, Tex ABERCROMBIE, H L Jr (J 55) Prod Engr Mobil Oil Co Room 1800 Gr PI Life Bldg, Lubbock, Tex ABERCROMBIE, Robert (J 57) Prod Engr Mobil Oil Co PO Box 1770 Shreveport, La ABERG, Richard S (M 57) Area Prod Geol Shell Oil Co of Canada, Box 100 Calgary, Alta, Canada]

Jan 1, 1961

By J. R. Hauber, N. R. Borch

A study uras made on the creep behavior of cast and extruded SR grade beryllium. It is shown that, for stresses below about 1000 psi in the temperature range 760" to 85o° c, the creep behatior is nearly exactly described by the Nabarro-Herring mechanism. The activation energy is apparently that for self-diffusion, 42 2 kcal per mole, below 1000 psi, but it rises to 88.6' kcal per mole at higher stresses. The slress exponent is 4.5 at the higher stresses. Tertiary creep in this material is clearly related to void formation apparently caused by grain boundary sliding. BERYLLIUM is becoming an increasingly important technological material. A search of the literature has revealed a lack of fundamental studies on the creep of beryllium. This investigation was therefore undertaken to gain insight into the creep behavior of this material. EXPERIMENTAL PROCEDURE Tensile creep samples were prepared from cast and extruded SR grade beryllium. The vendor's chemical analysis of the beryllium tested is shown in Table I. The test specimens were machined as shown in Fig. 1 from i -in.-diam rods. The machined creep specimens were annealed to remove surface damage and to establish a variation in grain size. The heat treatments and their resulting grain diameters are shown in Table 11. The average grain diameters were determined by the intercept method. True average grain diameters were estimated by multiplying the measured average grain diameters by 1.5.' The creep tests were performed at temperatures from 700" to 850°C in a vacuum of 105 Torr or better in a tantalum-element resistance furnace. Loads were applied by an Instron testing machine such that the stress was held constant to within 20 psi. The test temperature was measured by thermocouples on the sample. With one exception the deviation from the desired test temperature was 2"C, and in that case the deviation was 4"C. Elongations were measured by an extensometer within the vacuum chamber which was fastened to the creep specimens at the grooves in the grip ends, as shown in Fig. 1. The precision of measurement of elongation was 1 X loe4 in. The creep tests were begun only after thermal equilibrium had been established in the sample and the extensometer to insure that there was no strain rate component due to transient thermal effects. Two kinds of creep experiments were performed. First, at constant grain size, the effects of varying stress and temperature were studied. In the second set of experiments, the temperature was held constant and the grain size was varied. The activation energy for creep was determined by varying the temperature at constant stress, as described by orn, during the first set of experiments. The variable grain size experiments revealed that creep at low stress apparently occurred by the abarro-errin' mechanism. In all of the creep tests, the stress was held constant only until sufficient strain had occurred to accurately determine the creep rate. The stress was then raised to a new value and the new creep rate determined. This process was reported until tertiary creep occurred. In this manner several data points were obtained from each specimen. The assumption was made that the structure was not changing appreciably between the incremental stress levels. This was verified by reducing stresses to previously tested levels and comparing the creep rates with the creep rates obtained at the lower strains. The re-producibility was within experimental error over strains on the order of 1 pct. EXPERIMENTAL RESULTS A typical series of creep tests is shown in Fig. 2. Primary creep was not observed in any test. Transient creep, which we define as creep rates that are nonlinear with time, was not observed for any change in stress except where tertiary creep occurred. The absence of both primary and transient creep has been previously reported for hot-pressed powder beryllium.= The data for the creep behavior at constant grain size are summarized in Fig. 3. The data show two distinctly different stress dependencies. At stresses below about 1000 psi, in the equation:

Jan 1, 1969

By J. W. Snavely

A practical mathematical treatment is presented for the determination and control of conveyor belt acceleration, particularly for conditions of starting where vertical curves are involved. A typical sample problem is analyzed, with required calculations, to clarify the procedure. THE part that acceleration plays in starting a belt conveyor and its effect on belt conveyor design are well understood in a general way. Its practical importance is easily overlooked, however, and under some conditions, it is absolutely necessary to give the problem of acceleration detailed study. Most handbooks on conveyor belting design adequately present basic data for the determination of acceleration values. This paper will only attempt to present practical thinking and a convenient method of treatment of acceleration in belt conveyor design. Mathematical Analysis In working out the various problems of conveyor belt acceleration, the starting point, as presented by the handbooks, is the familiar formula of "force of acceleration is equal to the mass times acceleration." By expressing these fundamental quantities in terms of belt conveyor design, it is possible to arrive at the unsuspected conclusion that the acceleration time for horizontal belt conveyors is independent of the load, and instead, dependent upon the belt speed, the type of drive arrangement and drive pulley, and the idler coefficient of friction. The mathematics leading to this conclusion are shown in Table I, which has been prepared to show this derivation. While at first the conclusion just given may not seem to be reasonable, further reflection indicates that obviously the type of drive pulley and the type of drive do affect materially the tension in the conveyor belt, and thus, as clearly shown, the time of acceleration is dependent upon the factors mentioned. Inasmuch as all of the factors except time are predetermined by the belt conveyor design, it becomes relatively easy to establish the accelerating time and to reduce further this time determination to a simple graph from which the time in seconds can be read directly. Such a graph is given in Fig. 1. The table appearing on Fig. 1 should be explained further. For a given belt speed, the time of acceleration can be expressed as a percentage of the belt speed. The time of acceleration is also dependent on the drive arrangement, and changes in the drive arrangement consequently change the time of acceleration. It further follows that for a given belt speed, the time expressed as a percentage of that belt speed also changes with the type of drive. Obviously then, it becomes possible to graph the percentage of speed for .each type of drive against the belt speed and accelerating time, after which, for a given belt speed and type of drive, the time can be read directly in seconds. Two constants were established for Fig. 1, the first one being the limiting of the maximum acceleration tension to 35 pct of the full load operating tension in the belt. The purpose of this is to limit the total tension imposed upon the belt during the acceleration period to 135 pct of the full load operating tension, 'which is the amount required to start or breakaway the fully loaded belt conveyor from rest. The other constant is the friction factor used for the idler equipment, which has been established as 0.022. For installations where it is necessary to establish the values of acceleration, invariably high grade idler equipment is used, and it has been established from field experience that 0.022 for the idler friction factor is amply conservative. The use of this friction factor for idlers must be tempered with judgment, of course, for occasions will arise where more power than indicated is required to start, even with the very best of equipment, such as low temperature operations that tend to congeal the grease in the bearings and thus produce additional friction drag. An inspection of the table in Fig. 1 affords a convenient rule of thumb method for determining the acceleration time, which conveniently can be 5 pct of the belt speed in seconds. The 5 pct of belt speed figure is close to the average for most types of drives. In using Fig. 1 it must be emphasized that it applies accurately to horizontal belt conveyors only.

Jan 1, 1953

By J. W. Snavely

A practical mathematical treatment is presented for the determination and control of conveyor belt acceleration, particularly for conditions of starting where vertical curves are involved. A typical sample problem is analyzed, with required calculations, to clarify the procedure. THE part that acceleration plays in starting a belt conveyor and its effect on belt conveyor design are well understood in a general way. Its practical importance is easily overlooked, however, and under some conditions, it is absolutely necessary to give the problem of acceleration detailed study. Most handbooks on conveyor belting design adequately present basic data for the determination of acceleration values. This paper will only attempt to present practical thinking and a convenient method of treatment of acceleration in belt conveyor design. Mathematical Analysis In working out the various problems of conveyor belt acceleration, the starting point, as presented by the handbooks, is the familiar formula of "force of acceleration is equal to the mass times acceleration." By expressing these fundamental quantities in terms of belt conveyor design, it is possible to arrive at the unsuspected conclusion that the acceleration time for horizontal belt conveyors is independent of the load, and instead, dependent upon the belt speed, the type of drive arrangement and drive pulley, and the idler coefficient of friction. The mathematics leading to this conclusion are shown in Table I, which has been prepared to show this derivation. While at first the conclusion just given may not seem to be reasonable, further reflection indicates that obviously the type of drive pulley and the type of drive do affect materially the tension in the conveyor belt, and thus, as clearly shown, the time of acceleration is dependent upon the factors mentioned. Inasmuch as all of the factors except time are predetermined by the belt conveyor design, it becomes relatively easy to establish the accelerating time and to reduce further this time determination to a simple graph from which the time in seconds can be read directly. Such a graph is given in Fig. 1. The table appearing on Fig. 1 should be explained further. For a given belt speed, the time of acceleration can be expressed as a percentage of the belt speed. The time of acceleration is also dependent on the drive arrangement, and changes in the drive arrangement consequently change the time of acceleration. It further follows that for a given belt speed, the time expressed as a percentage of that belt speed also changes with the type of drive. Obviously then, it becomes possible to graph the percentage of speed for .each type of drive against the belt speed and accelerating time, after which, for a given belt speed and type of drive, the time can be read directly in seconds. Two constants were established for Fig. 1, the first one being the limiting of the maximum acceleration tension to 35 pct of the full load operating tension in the belt. The purpose of this is to limit the total tension imposed upon the belt during the acceleration period to 135 pct of the full load operating tension, 'which is the amount required to start or breakaway the fully loaded belt conveyor from rest. The other constant is the friction factor used for the idler equipment, which has been established as 0.022. For installations where it is necessary to establish the values of acceleration, invariably high grade idler equipment is used, and it has been established from field experience that 0.022 for the idler friction factor is amply conservative. The use of this friction factor for idlers must be tempered with judgment, of course, for occasions will arise where more power than indicated is required to start, even with the very best of equipment, such as low temperature operations that tend to congeal the grease in the bearings and thus produce additional friction drag. An inspection of the table in Fig. 1 affords a convenient rule of thumb method for determining the acceleration time, which conveniently can be 5 pct of the belt speed in seconds. The 5 pct of belt speed figure is close to the average for most types of drives. In using Fig. 1 it must be emphasized that it applies accurately to horizontal belt conveyors only.

Jan 1, 1953

By R. G. Ward, A. E. El-Mehairy

A technique was developed to calculate the density of liquid metals from the profile of a weighed levitated drop obtained by emitted light photography and calibration. The density of liquid copper was determined over the temperature range 1370" to 2100°K and expressed with the equation: The molar volumes and thermal coefficients of expansion are calculated. Copper, upon melting, was found to expand 3.34 pct. A knowledge of the density of liquid metals is essential as it is a parameter appearing in most theories of the liquid state. The liquid density is also necessary for the calculation of the commercially important contraction occurring on solidification. Unfortunately good density values are rarely available and the present work was undertaken to establish a new method of density determination free from the limitations of the existing methods. For liquid metals with high melting points, the variation and the limited number of density values reported in the literature have been caused primarily by difficulties in technique, a matter which has been adequately discussed by Urbain.' In all the methods used for measuring the density of the high-melting liquid metals there are sources of error, and corrections have to be made. For example, in the method of balanced columns, the formation of gas bubbles in one or both arms was a great source of error. In the pyknometer3 and immersed-sinker methods, corrections of certain volumes for thermal expansion to the operating temperature exceed 4 pct. In the maximum-bubble method5 a correction for the thermal expansion of the bubble-tube is used (about 1.6 pct); a second correction takes into account the change in the volume of liquid displaced by the bubble-tube (about 3.5 pct). Another method used is based on the shape of a drop of liquid, at rest on an inert refractory plaque, obtained photographically or radiographi-cally. The volume is calculated from the shape, which is sensitive to the surface tension and supporting material, by some intricate empirical relation. The drop contour changes with 90 deg rotation of the support.7 The application of all methods is limited by reaction between liquid metal and the apparatus. The successful stabilization of a mass of liquid metal during levitation melting8 afforded a technique for measuring the density of liquid metals out of contact with solid materials. The technique was developed by using copper as the density of liquid copper has already been studied extensively, thus allowing comparison with the new method. It consists of calculating the volume from the dimensioned profile of the levitated drop obtained by emitted-light photography and careful calibration. The molten egg-shaped drop spins around the magnetic axis of the coil (lined vertically) at a speed which seems to increase with increasing temperature. The method does not involve corrections and the experimental setup allows the use of vacuum or special atmospheres. The maximum operating temperature is limited by excessive evaporation of metal from the droplet. EXPERIMENTAL Levitating System. A stabilizing ring was used in conjunction with a 3-turn levitating coil energized by a 450-kc per sec, 10-kw Toccotron with a high current-low voltage output stabilized to f 1 pet.' APyrex glass tube with a Pyrex Optical flat window inserted through the coil and ring allowed the use of special atmospheres during levitation melting, as shown in Fig. 1. Optical System. Fig. 1 shows the optical arrangement used to photograph the profiles of the drop, parallel and perpendicular to the vertical spinning axis, simultaneously on the same frame of a high-speed 35-mm film. The optical system was adjusted to avoid any distortion of the profiles and to ensure that both profiles are in focus. The length of path from object to objective was about 30 cm, and a magnification factor close to 1 was obtained with the single-lens reflex-camera system. Temperature Measurement and Control. In operation, the specimen temperature was constant to 5 C for a fixed power input to the levitating coil and fixed rate of gas flow through the tube. The temperature was measured by an automatic two-color pyrometer (Coloratio) checked at the melting

Jan 1, 1963

By Ta M. Li

Expanded mine-mill capacity and greater operating flexibility are part of the dramatic changes taking place in the Tennessee mining operations of Asarco, Inc. Climaxed by the recent start-up of the $6.3 million Young mill in September 1975, Asarco has pushed Tennessee back into number one position as the leading zinc mining state. Furthermore, start-up of the Young mill is expected to insure a much-needed and continuing domestic zinc supply, at a time when US mines can only supply one-half of 1975 domestic consumption.

Jan 5, 1976

By W. T. Stuart

THE increased demand for iron ore has necessitated a re-examination of ore-bearing lands on which the presence of water previously has indicated hazardous and expensive operating conditions. In view of the importance of iron ore production to the national economy and defense, the Ground Water Branch of the U. S. Geological Survey, in cooperation with the state of Michigan, began a study of mine drainage in the Iron River district in 1945, and later extended the work to the Marquette district. The purpose of these studies was to define the principles involved in the movement of surface and ground water toward the mined areas, with the hope that the information obtained in the research would lead to the development of methods of water control and to a reduction in the total mining costs. Not only are the direct costs of drainage increasing, but also the indirect operational costs of working a wet mine are becoming a larger proportion of the total mining costs. For some wet mines the direct costs of pumping and drainage may range from 35 to 40Ø per ton of ore produced, but the indirect costs due to handling wet ore and controlling the water may be five to ten times this amount. The methods of study were formulated as the work progressed, and inasmuch as they were the first large-scale studies of their kind, they should serve as a guide for the solution of similar problems in other mining areas. The Iron River district was chosen for a pilot study because in this district the longest records of mine pumpage and water-level observations were available, including, as they do, the records for the Homer mine of the M. A. Hanna Co., where pumping from surface wells began in 1930. The results of the first investigation by the Michigan Department of Conservation have already appeared." The first section of the report on a similar study of the Marquette district, which was started in 1948, will be published this year. Methods of Study In an effort to reduce the flow of water into the mine workings in the Iron River district, about 4500 gpm was pumped from the bedrock being mined and about 9000 gpm from the glacial overburden. In the Marquette district in the vicinity of Ishpeming and Negaunee, about 5000 gprn was pumped from the bedrock and about 4000 gpm was pumped from the glacial overburden. Where the water was pumped only from the bedrock, the rate of pumping ranged from a few hundred gallons per minute in the dry mines to many hundred gallons per minute in the wet mines. In each district, pumping from the overburden was localized on a few properties where costly pumping installations had been made and the expenditures for power had been large. In each district a comprehensive ground-water investigation was made of the whole area, involving the collection and interpretation of all the available data bearing on the source and quantity of water to be controlled. Although it is not the purpose of this paper to discuss the methods of making a ground-water investigation, it should be pointed out that a drainage study follows a pattern of engineering analysis that determines the occurrence, source, movement, disposal, and quantities of water involved. The investigation in the iron-mining districts of Michigan began with the construction of a map of the buried bedrock topography. Because the ground-water reservoirs occupy the low points in the bedrock basins, this map gives information concerning their areal extent, depth, shape, and degree of interconnection. The depth to water in the drillholes and wells indicates the altitude to which the ground-water reservoirs are filled, and the logs of the material penetrated in the drillholes and wells indicate the general character of the materials filling the reservoir. The slope of the ground-water surface indicates the direction of flow through the reservoir, the water movement being from the points of higher altitude to points of lower altitude. By analysis of the rise and fall of the ground-water levels in response to additions of water through recharge and to changes in the rate of discharge through pumped wells, estimates of the total quantity of water in storage and of the rate of flow through the reservoir

Jan 1, 1952

By J. P. Zannaras

Referring to the article by R. J. Charles and P. L. de Bruyn, let us assume that W = weight of glass bar; P = weight of hammer; e = total deformation; K = unit of deformation; K = potential stress energy; E = modulus of elasticity; L = length of the bar; 7 = coefficient of inertia; h = height, ft; V = velocity, fps; and a = cross section area. The portion of kinetic energy which is effective in producing stress energy in a fixed bar struck horizontally is given by the formula" 1 + 1/3 W/P K =1+1/3w/p/(1+1/2w/p)2 P.V2/2g = Ph where 1 + 1/3 W/P ? (1 + 3W/P/(1=1/2w/p)2 [8] Putting e e = W/P =------------ From the above equation it can be seen that the maximum transfer of kinetic energy to stress energy is when e = 0 or W/P = 0 which indicates that the weight of the hammer must be very large as compared with the weight of the impacted rod. Eq. 8 diametrically opposes the conclusions reached by the authors of this article. In fact, if their suggestions were followed to the extreme when e = co when P = 0, there would be no transfer of kinetic energy to stress energy at all, as 7, becomes zero. Eq. 8 presumes that the velocity with which the stress is propagated through the bar is infinite, whereas the authors claim that the compression waves reflected are reaching the struck end of the bar prior to the complete transfer of the kinetic energy to cause such modification of the conditions there as to make them reach the reverse conclusions demonstrated by the above formula. That such interference exists is unquestionably demonstrated by the authors and others. However, if my observations are correct, such interference for this specific experiment and also for practical comminu- tion is insignificant, and the conclusions of the authors are in error and must be reversed to comply with Eq. 8. Eq. 5, w. = AE 2/2, given by the authors on page 51, is derived from the following equation (Eq. 9): K = 1/2Pe, where P = Sa, S = ?E, e = EL, and L = 1. The above formula, Eq. 9, cannot be applied in this case. This formula is applicable for static loads where the load increases from zero up to its final value, P, in such a way that the deformation at different instants is proportional to the loads acting at those instants and actually represents the area of a right triangle in the strain load diagram of base e and height P. The typical photographs shown in Figs. 3 and 4 represent the familiar strain load diagrams, and since the line of the wave marks the existence and intensity of the strain with the unquestionable conclusion that such strain has been caused by the action of a load acting continuously all along the wave until it reached the horizontal axis, the work stored at this point is represented by the area under the wave line and the horizontal axis and not by the area of the fictitious triangle given by the authors. Then if this is correct, even visual estimation of these areas at gage stations given in the typical photographs of Figs. 3 and 4 suffice to contradict the authors' calculation given in Figs. 6a and 6b and Figs. 7a and 7b. The typical photographs presented by Charles and de Bruyn show a considerable variation of the intensity of the strain at different stations but very small variation of areas which actually represent the stress energy at the corresponding stations. And, apparently, by squaring the small quantities, the authors magnified their error tenfold. J. M. Frankland's paperV iscusses the relative strain intensity and not the total energy for different types of impact loading. He states in his paper, "The reader is explicitly warned not to confuse the results in this report with those obtained when the load is applied by a blow as from a hammer. In this case the peak load rises to very large but mostly unknown values. The accompanying large deflections and stresses are the result of high values of P, not of the dynamic load factor n. According to Frankland "the dynamic load factor" is the numerical maximum of the response factor. It therefore appears that the authors followed the same procedure in obtaining the relative strain energy ab-

Jan 1, 1957

By Marcus A. Grossman

THE institute has established this lectureship to honor the memory of a great American metallurgist, one whose fame has continued long after his passing. As one scientist recently stated it," All metallurgical texts gradually become outdated and outmoded, except those of the great Henry Marion Howe." It seems fitting, then, that, once every year, we should review, in his honor, some aspects of the science in which he was such a distinguished leader, and also discuss any new data that may help to clarify our concepts. It is my privilege, for which indeed I express the deepest appreciation to this Institute, to offer for consideration a few thoughts on the toughness and fracture of hardened steels. Before proceeding to the data, I wish to express my great indebtedness to many associates whose stimulating discussions, generous contributions of data and cooperation in experiments are gratefully acknowledged. (But I wish to absolve them of any blame that may arise from the way in which their discussions and data have been used!) First, Dr. Edgar C. Bain, Vice President, Carnegie-Illinois Steel Corporation, whose amiable but merciless criticisms and always constructive and thoughtful suggestions help unfailingly to guard against pitfalls. Then to associates at Gary Steel Works, especially Mr. H. B. Wishart, Mr. W. K. Smith and Mr. V. Elliott, for great assistance in securing data. Also to Col. C. H. Greenall, Director of Laboratories, Frankford Arsenal, and his metallurgical staff, including Messrs. J. K. Desmond, H. Markus, D. F. Armiento, H. Rosenthal and Henry George, for numerous data and illustrations. Further to Dr. J. R. Low and Dr. M. Asimow, for helpful discussions of effects of states of stress during deformation and rupture, to Mr. J. M. Hodge for discussions and data, to the staff of the U. S. Steel Corporation Research Laboratory for advice and data, and to the Metallurgical Staff at Duquesne Steel Works. The individual acknowledgments will also be found under illustrations. Acknowledgment is also due to Mr. S. C. Snyder for incredibly meticulous scrutiny of manuscript. To all of these, grateful thanks.

Jan 1, 1946

By V. N. Krivobor

The study and use of low-carbon manganese steels have been curiously neglected in the general history of developments in alloy steels. Hadfield1 made an extensive study of manganese-iron-carbon alloys about 1886, which culminated in the wide application and use of Hadfield austenitic manganese steel. His alloys were nearly all high in carbon, as compared with the steels discussed in this paper, and he found a marked tendency toward brittleness in the alloys of medium and low manganese content, that is, below 7 per cent. manganese. As a result, there has been a widespread prejudice against the use of manganese steels containing between 1.0 to 3.5 per cent. manganese with comparatively low carbon. Steels of this composition have never been studied adequately in comparison with other alloy steels, and have had litttle use in practice. In a very recent paper by Sir Robert Hadfield2 the following statement appears: 6L... manganese-iron alloys with little or no carbon do not seem to offer any practical employment; they only become valuable in the presence of quite high percentages of carbon, 0.90 to 1.50 percent." The results of the present investigation are so contrary to the findings of Hadfield that the writers feel fully justified in presenting the preliminary report incorporated in this paper. Previous Investigations As early as 1886 the Terre Noire CO.3 in France produced alloys containing up to 2.5 per cent. of manganese, using ordinary low-grade

Jan 1, 1927

By K. G. Kreider, J. H. Brophy, J. Wulff

The technology of nickel-activated sintering of tungsten powder has been successfully applied to the densification of plasma-sprayed tungsten. Nickel was added by infiltration in a zinc solution followed by evaporation of the solvent. After sintering one hour at 1300°C density 95 pct of theoretical and transverse rupture strength of 74,000 psi were obtained. Shrinkage was found to be anisotropic and the mechanism of densification was comparable to that found in the nickel-activated sintering of tungsten powder. 1 HE use of a plasma spray gun for the fabrication of massive tungsten parts has become increasingly interesting. Applications now exist where a deposit in the as-sprayed condition is satisfactory. However, these deposits are generally characterized by a lamellar anisotropic microstructure containing 15 pct porosity of which, typically, two-thirds is open to the surface. Mechanically, the as-sprayed deposits fail at relatively low stress levels with a biscuit-like fracture. As a result of these problems the possibility of improving structure and strength by sintering treatments subsequent to spraying is particularly attractive. Preferably this sintering treatment should be adaptable to large bodies of sprayed metal. The similarity between the as-sprayed tungsten structure and that of a powder compact suggests that the relatively low-temperature activated sintering technique1 might be profitably employed in the densification of plasma-sprayed tungsten. It was the purpose of the present investigation to develop a technique for introducing the nickel-activating agent into the sprayed structure, to evaluate the amount and mechanism of densification obtained as a function of time and temperature, and to obtain an indication of the relative strength before and after sintering. EXPERIMENTAL PROCEDURE Powder used for spraying was purchased from the Wah Chang Corp. in several size fractions ranging from an average size of 4 to 150 . These powders were sized further for an explicit study of the influence of average feed size on densification. All powders were dried at 200°C before use. Spraying was accomplished with a Plasma Flame unit manufactured by Thermal Dynamics Corp. Several modifications of the unit were helpful in conducting the investigation. A variable speed auger feed mechanism coupled with the carrier gas mecha nism facilitated the use of fine particle sizes. A coil of ten turns of copper tubing in series with the arc power and concentrically would around the nozzle improved nozzle life and extended the range of operating currents available. The function of the auxiliary coil was to cause the arc to spin and to prevent impingement at only one point in the nozzle. Normally air sprayed deposits were made with an arc maintained at 400 amp at 50 to 70 v. The arc was blown by a gas mixture containing from 5 pct H, 95 pct N for the finest powder feed sizes ranging to 20 pct H, 80 pct N for the coarsest size. The flow rate was maintained at 100 cu ft per hr NTP through a nozzle of 0.25 in. ID. When apraying in air, the powder stream was directed toward an aluminum substrate for ease of mechanical removal of the deposit. The substrate was cooled by diverting the plasma flame with an air jet, and a second jet was directed on the deposit surface. In this configuration a gun-to-work distance of 2 to 3 in. was found to be satisfactory. Fig. 1 represents a typical as-sprayed deposit micro-structure. Laboratory studies of protective atmosphere spraying were carried out in cylindrical chamber 8 in. in diam by 18 in. in length. In operating the nozzle attached to such a chamber, particular care was required to avoid nozzle burn out due to reduced gas flow. The structure and density of the chamber sprayed deposits varied over wide ranges depending on substrate temperature. For the purposes of this investigation, flat deposits were made approximately 2 in. sq by 3/8 in. thick. From these deposits individual samples were cut an ground to a rectangular shape typically 1 1/2 in. by 1/8 in. sq such that the long dimension was perpendicular to the spraying direction. For the study of shrinkage anisotropy deposits up to one inch thick were produced. From these, rectangular samples were cut having a longer dimension parallel to the spraying axis. Prior to the addition of activating agent, the samples were deoxidized in hydrogen at 800°C for 20 min. No detectable dimensional or microstructural change was observed after this treatment. The addition of nickel was accomplished by infil-

Jan 1, 1963

By R. L. Wiegel

The two questions considered in this study are: (1) what similarity is there in the liberation of magnetite from various iron formations, and (2) how is this liberation affected by the size reduction technique employed. Twelve sample suites were obtained from commercial concentrators processing magnetite iron formation in the United States, Canada, and Australia. These samples have grades in the range 15 to 45% magnetite iron and approximate mineral grain sizes o f 50 to 2000 microns. Efforts were directed at measuring the liberation characteristics of the samples obtained from the concentrators primary grinding mills. The Davis tube, a laboratory magnetic separator, was used to concentrate individual size fractions of the size reduction product to determine liberation data. These liberation data were compared with relationships calculated from a highly idealized mathematical model of liberation in a binary mineral system. The results of these comparisons have been encouraging for nine or ten of the twelve samples tested. This leads to the conclusion that for most commercial iron formations, at least with conventional rod mill size reduction, it is possible to quantitatively characterize liberation with two parameters: an effective mineral grain size and ore grade. Limited laboratory size reduction tests as well as one of two sets of liberation data from commercial primary autogenous grinding lend some support to the contention that autogenous grinding can affect mineral liberation.

Jan 1, 1976

By E. A. Gulbransen, K. F. Andrew

In a recent communication14 we have reported on the kinetics of the reactions of zirconium with O2, N2 and H2 as a function of the time, temperature and pressure variables. A systematic study was made and the results correlated with fundamental theories of gas-metal reactions. This paper will present a similar study for titanium. Titanium and zirconium are members of the IV group of the periodic table and possess many similar physical and chemical properties as a result of their similar electronic configuration for the outer electrons. The two metals are relatively inert to both gas and liquid phase corrosion at room temperature. However, at moderate temperatures the metals become active and react readily with the common gases including O2, N2 and H2 which are of interest in this study. A study of the kinetics of these gas-metal reactions is of interest for three reasons: (1) to understand the rate of reaction of titanium and its role in the behavior of high temperature alloys; (2) to understand the practical difficulties of the reduction, refining and working of titanium; and (3) to correlate the data with fundamental theories of gas-metal reactions and crystal structure predictions. Literature Survey Several review papers8 and books4243 exist on the preparation and properties of titanium and its alloys. THE METAL Titanium has, at room temperature, a hexagonal lattice of the zinc type. Hagg19 gives a value of 2.953A for the (a) axis, a value of 4.729 for the (c) axis and a density of 4.427 at 20°C. Burgers and Jacobs6 have observed the transformation of the hexagonal to the body-centered cubic structure at 880°C and have established a value of 3.31 for the cube edge and a density of 4.31. TITANIUM-OXYGEN Carpenter and Reavell6 using a pressure change method have studied the reaction at temperatures of 742° and 1000°C and for a pressure of one-fifth of an atmosphere. The probabilities for reaction are calculated from kinetic theory and they report a value of 10-5 for O2 at 1000°C and 10-6 at 740°C. The titanium-oxygen system has been investigated by Ehrlich.10,11 Five phases are observed. Between (TiO2 and TiO1.90) an alpha-phase, consisting ofarutile lattice, is found. A beta-phase is observed between (TiO1.80 and TiO1.70). A gamma-phase is homogeneous between (TiO1.56 and TiO1.46) and has a structure of the corundum type. The delta-phase exists between TiO1.25 and TiO0.6 and has a sodium chloride structure. From TiO0.42 to Ti the metal structure is observed. The surface oxide films have been studied by Hickman and Gulbransen.20 The rutile structure is observed in the temperature range studied, 300 to 700°C. Three crystalline modifications of TiO2 exist: rutile and anatase which are tetragonal and brookite which has an orthorhombic structure. Anatase is reported36 to exist in two forms: I and II. Anatase II changes to anatase I at 642°C. Anatase I is stable up to 915°C where rutile becomes the stable modification. At 1300°C rutile transforms to brookite which melts at 1900°C. The monoxide, TiO, may be prepared from the dioxide by high temperature reduction with carbon or magnesium. Its melting point is 1750°C. TITANIUM-NITROGEN Carpenter and Reavell6 report that at 1000°C a linear rate law is observed. The probability of reaction is given as 10-8 at 1000°C. Fast12 has studied the solubility of nitrogen and its effect on the mechanical properties of the metal. The crystal structure of TiN has been shown by several workers2'21'44 to follow the sodium chloride structure. However, the calculated density is found to differ from the pycnometric value. This is studied by Brager3,4 in detail. He has suggested that the titanium sites in the lattice are only partially filled at low temperatures. As the temperature of preparation is raised the vacant sites become occupied which expands the lattice and increases the hardness and density. An (a) value of 4.22Å is given for room temperature. TITANIUM-HYDROGEN The solubility and the crystal structures observed in this system have been reviewed in a recent book by Smith.40

Jan 1, 1950

By W. R. Johnson, M. Hansen, John M. Parks

AS part of a research project sponsored by the Signal Corps Engineering Laboratories, which had the objective of obtaining a magnet wire of good conductivity and low temperature coefficient of resistance, a comprehensive literature survey, supplemented by experimental work, revealed that a linear relationship exists for solid-solution alloys between the electrical conductivity, K, and the temperature coefficient, a, as expressed by the equation: k = a/ß where ß is constant for the solvent metal and a - per°C. tR0 It has long been known that the electrical conductivity at a given temperature and the temperature coefficient of resistance of binary alloys are related to their constitution by the LeChatelier-Guertler rules, as illustrated by Figs. 1 and 2: Rule 1—In alloy systems consisting of a heterogeneous mixture of two phases, the electrical conductivity and temperature coefficient vary as a straight line, if the composition of the alloys is given in volume percentage, Fig. 1. Rule 2—In solid solutions, the electrical conductivity and temperature coefficient are always lower than those of the solvent metal, and usually a considerable decrease occurs with the first small percentage additions of the solute metal, Fig. 2b. If the two components form a continuous series of solid solutions, the electrical conductivity and temperature coefficient curves take the form of a U-curve, Fig. 2a. That some kind of proportionality exists between electrical conductivity and temperature coefficient of resistance is implied by the LeChatelier-Guertler rules and by study of data such as those given by Smith and Palmer,' Fig. 3. Dellinger2 in 1910 first recognized this relationship in a number of commercial copper wires, ranging in conductivity from 94 to 100.7 pct I.A.C.S. He gave the ratio of temperature coefficient to the percentage of conductivity as a constant, C = 0.003939. A year later, Lindeck observed that the linear relationship was valid for a wider range of conductivity in "various sorts of copper," and concluded that the product of resistivity and temperature coefficient was 6.78x10-5 1.5 pct. Since then, this relationship seems to have been entirely ignored. Conductivity-Temperature Coefficient Relationship If the electrical conductivity of copper solid-solution alloys with phosphorus, silicon, arsenic, antimony, aluminum, nickel, manganese, and iron is plotted against the temperature coefficient of these alloys, the diagrams in Figs. 4 and 5 are obtained. Similar data were found with copper-rich solid solutions containing magnesium, cobalt, and titanium,'" as shown in Fig. 6. In all cases, a linear relationship exists. The point representing the highest

Jan 1, 1952

By V. S. Gupt, W. H. Somerton

Rock may undergo great changes in physical properties when heated to high temperatures and then cooled, The temperature and intensity of reactions causing rock alterntiorl.s can he controlled by introducing certain chemicals during heat treament. Three typical outcrop sandsone samples were saturuted with common salt solutions, then heated to several maximum temperatures. After cooling, it was found that per-rrreuhilities had increased much more for salt-saturated samples than for samples not saturated with salt but heated to the same temperatures. This was only true, however. for samples heated above the melting points of the particrrlar salts. Potassium chloritle was particularly effective with Bandera sandstone. Samples saturated with potassium chlorirlr. solution and heated to 900C showed an 11-fold increa.se in perrrieahility. Samples without potasrirdrn chloritle but heated to the .same temperature. showed only a 2.7-jold incresre in permecahility. In application, it seems possible that injecting chemicalc into the formations from a wellbore followed by applying intensive borehole heating might promote reactions which would greatly improve permeability of the formalions. INTRODUCTION Great changes in the physical properties of rocks which have been heated to temperatures in the range of 600 to 800C have been reported earlier. Permeability increases of 50 per cent, and equivalent decreases in sonic velocity and breaking strength have been observed. Although there might be some suppression of certain reactions responsiblc for these changes when rock samples are heated under simulated reservoir pressure conditions. recent work has shown that these are more than offset by the increased importance of other reactions at high pressures. The reactions considered responsible for alteration of rock properties by heating include differential thermal expansion. dehydration, phase changes and dissociation of mineral constituents. Ceramists have long known that temperatures at which reactions occur can be raised or lowered by the presence of certain impurities in the sys. tcnl. Thus the possibility exists that by saturating rocks with appropriate solutions. desired reactions might be accelerated and unwanted reactions might be suppressec! when the rock is heated. Purpose of the present work was to investigate the effects of several common salts on the thermal alteration of sandstones. Types of reactions which might occur are reviewed and the findings of a number of thernlochemical alteration tests are presented. THEKMOCHEMICAL ALTERATION OF SANDSTONE MINERALS The most abundant constituent of most sandstones is quartz. Quartz can exist in at least eight different forms, but in the temperature range of immediate interest (to 900C) only two forms are of importance—alpha quartz below 573C and beta quartz above this temperature. Although the alpha-beta inversion of quartz is very important in thermal alteration of rocks. it is not particularly important here because the inversion temperature cannot be changed significantly by the presence of impurities or by the application of pressure. The transition temperature of quartz to tridynlite is 867C. but the transition is very sluggish. Tridymite is considered to be a metastable transition phase between two stable phases: quartz?cristobalite. However. the transition temperature for tridymite to cristobalite is 1,470C.4 The quartz-tridymite-cristobalite transition can be accelerated substantially by the presence of fluxing agents. Finely divided calcium, magnesium and titanium oxides accelerate the conversion while A12O may retard it. A Mixture of NaCl and CsCl reduces the temperature for cristobalite formation by as much as 420°C. Feldspars constitute the second most important mineral in sandstones. The melting points of feldspars range between 1,000 and 1,700°C depending upon the variety. No information available indicated any important role of fluxing agents in the thermal alteration of this group of minerals. Carbonate minerals are subject to dissociation at temperatures within the range of the present investigation. The dissociation of magnesite can start at a temperature of 373°C, but the reaction is sluggish and might not occur untll a temperature of 500°C or ,higher is reached. Dolomite dissociates in two stages at 500°C and 890°C, whereas calcite dissociation temperature is about 885C. Because CO2 is released in the dissociation of carbonates, ail such reactions are somewhat dependent upon CO2 partial pressure. In the absence of CO: in the surrounding atmosphere, the dissociation of calcite starts at 500°C.However. when 1 atm of CO surrounds the sample, the dissociation does not start until 900C. The other carbonates are apparently much less sensitive to a change in partial pressure of CO2. The aragonite calcite transformation can occur an!. where within the temperature range of 357 to 488°C. depending upon the presence of impurities such as barium. strontium. lead and perhaps zinc. The differential thermal analysis (DTA) curves of both magnesitc and dolomite vary with the presence of 'impuritics. The presence of iron in particular seems to affect the magnesitc DTA curve

Jan 1, 1966

By H. H. Stoek, G. W. Harries

The term " Anthracite Coal-Field " is generally used to refer to a comparatively small territory lying in the eastern-central part of Pennsylvania. This territory includes about 3,300 sq. miles of area, located chiefly in five counties in which anthracite-mining is the dominant industry. Although an area of 3,300 sq. miles is embraced in the region, less than one-fifth of this area, or only 484 sq. miles, is underlain by workable coal-seams. Moreover, this prorluctive portion is not a continuous area, but consists of a number of detached basins. These are known as the Northern, containing 176 sq. miles; Eastern-Middle, containing 133 sq. miles; Southern, containing 181 sq. miles; and western-Middle,containing 194 sq. miles. From the standpoint of business and trade, the region is divided into three districts, known as (1) the Wyoming, which includes the Northern field and the small Bernice basin; (2) the Lehigh region, which includes all of the Eastern Middle field and that part of the Southern field east of Tamaqua; and (3) the Schuylkill region, which includes the WesL ern-Middle field and that part of the Southern field west of Tamaqua. The relative commercial importance of these districts at the present time is shown by the coal-shipments as given in the following table :

Jan 1, 1904

By P. W. Bridgman

FLOW and fracture are admittedly complicated phenomena of which we are yet only partially masters. There is not even universal agreement as to the details of the language best adapted merely to describe the phenomena, to say nothing of any theory. Certain points of view, however, are widely held, and there are certain general expectations as to what sort of procedures will prove permissible. A complete description of any situation involving fracture would demand a complete history of all the variables in the control of the experimenter. The experimenter endeavors so to choose his independent variables that they are adequate to determine completely the situation in the sense that when an experiment is repeated on the same initial material and with the independent variables running through the same sequence, all the other measurable properties of the system will also run through the same course. It is, I think, the universal conviction that when we are dealing with phenomena of flow and fracture, the most important of the necessary variables are the external forces acting on the various parts of the body; when the force history repeats, with certain obvious restrictions such as that temperature must be constant and non- mechanical forces, such as magnetic field, must be small enough to be ineffective, then the history of all measurable properties will repeat; in particular the alteration of geometrical shape due to flow will run through the same course, and when fracture comes, the character of the fracture will be reproduced. Complete mastery of merely the descriptive aspects of flow and fracture would demand the ability to predict the entire course of flow and complete characterization of the final fracture for every possible combination of forces, varying in time in all possible ways, when applied to bodies of all possible shapes. This full situation is obviously of prohibitive complexity. The first stage in reduction to manage- able simplicity is to resolve the total situation into the sum of contributions made by small elements. Our thesis is that if each of the elements is adequately characterized, the behavior of the whole system is fixed. The element is usually taken of a simple geometrical shape, suggested by the mathematics best adapted to the particular situation. The forces acting across the faces of the elements are ultimately determined by the forces acting across the free faces of the whole system. The detailed process of finding the precise forces on the faces of all the elements in terms of the external forces involves the use of a special mathematical invention, the stress, a constructional quantity consisting of six components. The stress at any point in the body is determined by both the external force system and by the physical properties of the body at all points; quite different stresses prevail at the interior of bodies of the same shape under the same external forces in the elastic and plastic ranges. The forces on the faces of any element are, however, uniquely determined by the stress system at the element. Our fundamental thesis

Jan 1, 1944

By Mary Barrett, Walter Rose

INTRODUCTION The conventional way to program reservoir problems* for solution on the electric analyzer1 is one which concentrates the block units towards the inner radius of the prototype system being represented by the analog model. This is achieved by the device of considering block units of equal value radial resistance, leading to the consequences: (1) the ratio of the external to the internal radius of each block unit is a constant (i.e. r1/ri-l = constant a); (2) the external radius of the nth block unit is equal to the innermost radius of the system times the constant a raised to the nth power: (3) the capacitance of the block units increase by a factor of a2 going out radially from the innermost to the outermost boundaries of the system; (4) given n block units in the radial direction, the capacitance ratio of the outermost to the innermost block unit is a (2n-2); and (5) the vertical resistance of the block units decrease by a factor of a going out radially from the innermost to the outermost boundaries of the system, such that the ratio of the maximum to minimum vertical resistance is also given by the factor a(2n-2) PRESENT PROGRAMMING METHOD While this conventional way of programming problems is quite adequate for routine uses of the electric analyzer, it leads to difficulties in those cases where the inner radius of the prototype system (e.g., the well-bore) is small compared to the external radius (e.g.. the external reservoir boundary). To handle such a case by the conventional method would require either that a large value be selected for the constant a or that a large number n, of radial block units be chosen. In consequence, the capacitance and vertical resistance ratios become impossibly large, as will be evident from the relations given above. In any event, the size of the smallest and largest capacitors will always be limited by the size of commercially available components, and the analyzer itself will determine how many mesh points are available for radial block units. For example, it can be shown that if the maximum capacitance ratio which can be electrically handled is a factor of l04, the maximum external radius of the system which can be considered (assuming an internal wellbore radius of 0.25 ft is: It is thus seen that in direct consequence of having the inner wellbore radius small compared to the external radius, models extending beyond 30 to 40 ft are not possible when the capacitance ratio is limited to l04. (NB. A value of 10-9 farads can perhaps be taken as the minimum electrical capacitance possible if corrections for the capacitance of the lead wires are to be avoided. Values greater than 10-5 farads would require the use of unstable and inaccurate electrolytic capacitors). On the other hand, models extending out to 500 to 5,000 ft may be desired in single-well problems, and hence the need for a revised method of problem programming. * REVISED PROGRAMMING METHOD The revised method of programming is one which attempts to retain the desirable features of the conventional method (viz. concentration of block units towards the inner boundary, simplicity in computation, etc.) without having as a consequence that impossible capacitance and vertical resistance ratios will be required. Other advantages will also be achieved by the revised

Jan 1, 1957

By A. W. Demmler

RECENTLY, Chaudron,&apos; Montariol,2 and their co-workers published the results of their studies on zone-refined aluminum. They found their material to be capable of recrystallization at —50°C, but gave no data as to the percentage reduction or time required. Their material was found to be practically non-work hardening at room temperature, but after 96 pct cold work at a low temperature and storage in liquid nitrogen it was found to recrystallize to coarse grains in less than 3 hr at 20°C. Recently, a study was made of the room temperature recrystallization characteristics of zone-melted aluminum. An ingot 14 in. in length, with a 3hx3 in. cross section, was made in an extremely high purity graphite boat under an atmosphere of dried commercial argon. It was made by first passing the molten zone along the ingot of 99.99+ pct A1 15 times. At this point, the most impure 3 in. (the end toward which the eutectic forming impurities migrated) were replaced with more of the original starting material, and 15 more traverses were made. Each traverse of the zone along the ingot required approximately 3 hr. Three ingots were prepared in this identical fashion and the chemical analyses were made spec-trographically on two of these ingots, rather than on the one used in this investigation. The analyses were made at 1 in. intervals along the ingot, in duplicate. Over the purest 5 in. of the ingot, which were most important in this investigation, there was found an average composition of <0.00004 pct Cu, <0.0003 pct Fe, <0.0001 pct Si, and <0.0001 pct Mg. In addition, the following elements were not detected spectrographically even in the more impure end of the bar: manganese, zinc, nickel, chromium, titanium, vanadium, lead, tin, boron, beryllium, bismuth, gallium, zirconium, cadmium, antimony, cobalt, and molybdenum. The original ingot was given several cold reductions, with intermediate anneals in order to refine the grains. Owing to the extreme length resulting and the ease of deforming the material plastically, it was cut into eight equal lengths and identified as shown in Fig. 1, with the highest numbered section at the purer end (that enriched with peritectic formers). Section 6 was used in the exploratory work and was found to be of exceptionally high purity, giving more rapid rates of room temperature re crystallization than any previous aluminum studied at the Alcoa Research Laboratories. This fact suggested that the entire midsection of the ingot might be of very high purity. Accordingly, section 4 was chosen for the next study. A portion of it was rolled to cold reductions of 10, 25, 40, 60, 75, and 90 pct at room temperature, and the re-crystallization of each specimen was followed by X-ray pinhole transmission photographs. The other portion received the same treatment, except that both the specimens and the rolls were maintained at 160°F in the process of rolling. The latter ortion was retarded in its recrystallization as compared to the former, probably due to the stress relief effect of the warm rolling. Both of these portions of the fourth section were found to recrystallize at a considerably slower rate than did the sixth section.

Jan 1, 1957