Search Documents

By L. C. Glenn

IN the preparation of this paper the writer has drawn freely upon the writings of Orton, Munn, Shaw, Mather, Miller, Hoeing, St. Clair, Jillson, and others, as well as upon his own personal knowledge of the fields of both states. It is to be regretted that certain data gathered by him and his assistants last fall are not available for publication. OIL IN TENNESSEE A few wells drilled for brine for salt making in Tennessee between 1820 and 1840 obtained oil, but no definite search was made for oil until just after the close of the Civil War. Active drilling was then begun in Overton and counties southwest of it on the eastern half of the Highland Rim. A number of strikes were made at shallow depths in the basal part of the Mississippian but the wells were soon exhausted and abandoned. Drilling was revived, about. 1892, when the Spurrier district in Pickett County was developed and was followed by the Riverton district in the same county in 1896. A pipe line was laid from the Wayne County, Ky., fields and about 60,000 bbl. of oil were run before a very heavy slump in the production, a failure to find an extension of the field, and excessive local taxation caused the removal of the pipe line in 1906. There was then no production in Tennessee until the discovery, in 1915, of oil near Oneida, Scott Co., at about 950 ft. (289 m.) in fissures in the Newman or St. Louis limestone. This field, however, soon failed and was abandoned. In 1916, oil was found at Glen Mary, Scott Co., in the Newman limestone at a depth of 1232 ft. (375 m.) A number of wells have since been drilled there, some of which were dry while others, close by, were producers. The largest one yielded, at first, about 340 bbl. per day and produced for several months, when it suddenly went dry. Several of the first wells began at 6 or 8 bbl. per day and are still maintaining that

Jan 1, 1920

By T. W. Guy

The increased moisture due to water-spraying for coal dust control is of interest even for mines from which the coal is to be wet-washed before screening, because in many cases wet coal dust materially de- 20-mesh X 0 sizes as pct ROM average ............................... 6.3 pct Range 2.3 to 9.5 pct 20-mesh X 0 sizes as pct 3/8 X 0 sizes average........'...................... 18.9 pct Range 12.7. to 23.9 pct creases the loading efficiency of men and equipment at the faces. It also causes a serious increase in the normal wear and deterioration of all equipment with which it comes in contact. It is important to the consumer if it increases his as-received moisture. No attempt will be made in this paper to cover the entire subject of water-spraying of coal for dust control. It will be confined to discussion and data on (I) percentages of dust in bituminous coals, (2) moisture increases due to spraying for dust control in mines or plants by methods which have proved satisfactory for periods of one to five or more years, and (3) a few suggestions for research. Size and Percentage or Coal Dust The U. S. Bureau of Mines1 classifies as coal dust all bituminous coal passing a 20-mesh sieve (0.0328 in. square opening). The percentage of such dust to be found in ROM coal, of course, varies widely with the characteristics of the seam, and with the methods used in the mining and in the preparation of the coal. ROM screen analyses of 12 U. S. coals selected at random from personal files given in Table I show: A few other screen analyses examined indicating 40 pct or more of 20-mesh X 0 dust in 3/8 in. X 0 sizes were not included because of uncertainty that they were representative. The number of results in Table I is not sufficient to give a reliable estimate of the average, or of the range to be expected, but it indicates that in every case the coal dust is a relatively large percentage of the total ROM coal, and, of course, a progressively larger percentage of resultants as the larger sizes are separated for marketing.

Jan 1, 1949

By C. R. Swarts

The use of compressed air or gas in oil wells for raising crude oil to the surface has, within the last year, become a dominant factor In production engineering. Previous application of this principle was largely confined to oil wells that had been pumping for a considerable length of time. In the Seminole area, the gas-air lift has been applied to wells imniediately after natural flow has ceased and in some cases while the wells were still flowing. This early application of the gas-air lift has been watchcd with intense interest by the petroleum industry, both with respect to methods of application and mechanical installations and to the effects on immediate and ultimate production. Methods of Application The Seminole pool at its inception contained sufficient gas associated with the oil to cause natural flow. As no frce gas was found in the main productive horizon, the Wilcox sand, all the gas present was dissolved in the oil, hence the natural flowing life was short, rarely lasting over 45 days. When natural flow ceased, high fluid levels were common, and as depth and crookedness of holes were obstacles to efficient pumping, the gas-air lift was used almost exclusively in raising the oil to the surface. Because of the high gravity of the oil, 41" A. P. I., and its high gasoline content, and also because of the desire to obtain production before offset wells became productive, quantity of production was the aim of the operators and a method of flow was used whereby this aim could be fulfilled. One of the outstanding features of this method consisted in flowing the wells through the annular space between the tubing and rasing. By this means a greater quantity of fluid could be handled than by flowing through the tubing of the size in ordinary use. Wells that would produce only a few hundred barrels on the pump were by this rnethod made to produce as high as 7000 bbl. on the lift. A few days or wreks of production at that rate meant a large recovery of oil. Casing of 636 in. diameter was commonly used as a production string with a 5 3/16-in. liner set on top of the oil sand. Upset tubing for irtro-

Jan 1, 1928

By N. M. Langdon

Spokane Meeting, September, 1909.) THE following corrections to the paper of Mr. Langdon (Bulletin No. 34, October, 1909, pp. 919 to 940), received after the publication of the paper in the Bulletin, will be incorporated in the paper as finally published in the forthcoming Vol. XL. of the Transactions, but are here presented for the information of members:' Page. 921. Table I., furnace M2, item 11. For " 8801 " read "0.8801." 922. Table II., furnace C, item 2, under gas. For " 0.011 C " read " 0.011 0." Furnace G, item 13, under gas. For " 5.074 air " read " 3.074 air." Item 15, 114.269 " should be " 4.289." Furnace H, item 2, under iron. For " 0.04 " read ".0.01." 923. Table II., furnace J, item 1, under iron. For " 0.95 " read " 0.93." Furnace N2, item 6, under gas. For " 0.66 CO," read 0.066 C02." 924. Table III., furnace C, item 17. For " 379 " read " 397." 927. Table IV., furnace C, item 4. For " 0.0124 " read 0.0424." Furnace G, item 15.. For " 4,060 " read " 4,066." 928. Line 2. Items 19 and 20 of heat-supply, should be items 20 and 21. 930. Line 3. Expunge remainder of paragraph after the word " supplied," and substitute the following : "As compared with the fusion-zone, the loss by radiation n C was 25.5 per cent. of the total (2,278 units) supplied to that zone, while in D it was 22.5 per cent. of the total heat (1,824 units) supplied to that zone. Finally, as compared with the reduction-zone, C lost by radiation 36.8. per cent. of the 1,580 units supplied to that zone,

May 1, 1910

By J. W. Freeman, J. P. Rowe, R. F. Decker

HEAT-to-heat variations in properties of an alloy of constant nominal chemical composition have been a perplexing problem to the metallurgist. These heat-to-heat differences have been especially baffling in high-temperature metallurgy. In order to establish fundamental reasons for these heat-to-heat variations, the National Advisory Committee for Aeronautics has sponsored and financially assisted an extended research program on heat-resistant alloys at the University of Michigan. As part of the program to find the contribution of processing variables, Frey, Freeman, and white' studied effects of heat treating on nickel-base alloys and Ewing and reeman' investigated effects of hot-working practice on N155 alloy. To extend the program to the study of effects of melting practice on high-temperature properties, a 55 Ni-20 Cr-15 Co-4 Mo-3 Ti-3 pct A1 composition was selected for experimentation. The objective of this program was to establish the melting-practice variables which influenced the hot-working characteristics and creep-rupture properties and, at the same time, establish the metallurgical mechanisms for the effects. During the course of the investigation of practices using vacuum melting, it became increasingly evident through chemical analyses supplied by industrial concerns that a major variable was melt reaction with the crucible. This paper is devoted to the results of this phase of the melting-practice study. EXPERIMENTAL PROCEDURES Experimental 10-lb heats were induction melted in the University of Michigan vacuum-melting unit with the melting practices outlined in Table I. Aim analysis for the basic alloy, in wt pct, for all heats was as follows: Cr Co Mo Ti Al 20.0 15.0 4.0 3.1 3.1 Mn Si C Ni 0.12 0.12 0.08-0.15 Bal. Small amounts of boron and zirconium were added to certain heats as listed in Table I. These additions

Jan 1, 1959

By James M. Stapleton

It has long been recognized by blastfurnace men that correct top distribution of materials is very important in efficient and economical furnace operation. Thousands of experiments on top design, filling, sizing, and other operations have been carried out to improving furnace performance by securing uniform distribution. Many of these experiments have worked out satisfactorily and have been passed along to the industry. Many of them have been discarded and forgotten. It has been the custom to judge the results of such experimentation by the results shown in improved or lowered furnace efficiency. While the effect on furnace practice will have to continue to be the final criterion of experimental changes, it has always been a desire of blast-furnace men to ascertain the reason for failure or success. Efficiency in the blast furnace approaching the maximum can be measured in several ways: (1) by actual operating results as revealed in the coke rate and tonnage figures, and (2), in the degree to which the top gas approaches the limiting ratio of CO/CO2 for the reduction of iron oxide by CO gas. This ratio is shown in the equilibrium curve in Fig. I for the system Fe-O-C and CO2-CO-C. The closer the top gases approach this limiting ratio (set at 1.08 by Martin1 because of the effect of limestone calcination) the more efficiently we may expect a furnace to operate. The fact that low CO/CO2 ratios parallel eficient operating performance caused us to investigate the cause of good or poor distribution and the effect of filling and other changes on furnace operation. Heretofore it has been possible to take a sample of furnace gas (usually from the dust catcher) and from it determine how closely the reducing gases were to the ideal or limiting value as set by the equilibrium diagram. However, to find just where in the top area the distribution was distorted and the top-gas CO/CO2 ratios were distant from equilibrium values has been a difficult and expensive investigational procedure. In an effort to overcome this difficulty, apparatus was designed to take gas samples and temperatures across the top radius of the furnace. It was hoped that with the aid of this apparatus traverse sampling of the ascending gases could be simplified, and the cost of sampling substantially lowered. Routine experimenting could then be maintained and results checked with furnace operation. It must be emphasized that a traverse sampling is necessary for a true picture, as single samples at any one point are unreliable. This statement can be verified by an examination of Tables 3 to 6, showing results obtained at our furnaces in a large number of tests. Apparatus The design of the equipment for taking the samples was governed by the need for providing a simple, routine test that could be performed fairly quickly without adversely affecting normal furnace operation.

Jan 1, 1943

By James M. Stapleton

It has long been recognized by blastfurnace men that correct top distribution of materials is very important in efficient and economical furnace operation. Thousands of experiments on top design, filling, sizing, and other operations have been carried out to improving furnace performance by securing uniform distribution. Many of these experiments have worked out satisfactorily and have been passed along to the industry. Many of them have been discarded and forgotten. It has been the custom to judge the results of such experimentation by the results shown in improved or lowered furnace efficiency. While the effect on furnace practice will have to continue to be the final criterion of experimental changes, it has always been a desire of blast-furnace men to ascertain the reason for failure or success. Efficiency in the blast furnace approaching the maximum can be measured in several ways: (1) by actual operating results as revealed in the coke rate and tonnage figures, and (2), in the degree to which the top gas approaches the limiting ratio of CO/CO2 for the reduction of iron oxide by CO gas. This ratio is shown in the equilibrium curve in Fig. I for the system Fe-O-C and CO2-CO-C. The closer the top gases approach this limiting ratio (set at 1.08 by Martin1 because of the effect of limestone calcination) the more efficiently we may expect a furnace to operate. The fact that low CO/CO2 ratios parallel eficient operating performance caused us to investigate the cause of good or poor distribution and the effect of filling and other changes on furnace operation. Heretofore it has been possible to take a sample of furnace gas (usually from the dust catcher) and from it determine how closely the reducing gases were to the ideal or limiting value as set by the equilibrium diagram. However, to find just where in the top area the distribution was distorted and the top-gas CO/CO2 ratios were distant from equilibrium values has been a difficult and expensive investigational procedure. In an effort to overcome this difficulty, apparatus was designed to take gas samples and temperatures across the top radius of the furnace. It was hoped that with the aid of this apparatus traverse sampling of the ascending gases could be simplified, and the cost of sampling substantially lowered. Routine experimenting could then be maintained and results checked with furnace operation. It must be emphasized that a traverse sampling is necessary for a true picture, as single samples at any one point are unreliable. This statement can be verified by an examination of Tables 3 to 6, showing results obtained at our furnaces in a large number of tests. Apparatus The design of the equipment for taking the samples was governed by the need for providing a simple, routine test that could be performed fairly quickly without adversely affecting normal furnace operation.

Jan 1, 1943

By R. A. Grange, H. M. Stewart

Man.; steel parts may crack if quenched directly into a bath near room temperature, but not if quenched at a temperature just above the range where martensite forms and then allowed to cool slowly to room temperature. This latter procedure, which is the basis for such modern hardening techniques as "martempering" and "isothermal quenching," may entail some sacrifice in depth of hardening but very little, if any, in intensity of hardening. In planning such treatments, it is necessary to know the upper limit of the martensite range, and often desirable to know the proportion of martensite that would form on cooling to any lower temperature; furthermore, the tendency of a particular steel to crack when quenched is unquestionably associated with the temperature range of martensite formation, and consequently knowledge of this range aids in selecting the optimum composition for a given application. Certain puzzling phenomena associated with the properties of quenched and tempered steel may be at least partly clarified as the manner in which martensite forms is more fully understood. Information on the temperature range of martensite formation, particularly with respect to how much martensite results from quenching to a given temperature, is lacking for most commercial types of carbon and low-alloy steel. These considerations led us to study martensite formation in 14 carbon and low-alloy steels. The resultant data may be directly used in the following ways: (I) for selecting the lowest quenching temperature at which no martensite will form; (2) for selecting the highest quenching temperature at which virtually all marten . site will form, thereby avoiding quenching to an unnecessarily lower temperature with attendant danger of cracking; and (3) in producing a mixture of tempered martensite and bainite, which in high-carbon steels has been found to possess somewhat better ductility than tempered martensite, yet does not require the full transformation time necessary for a completely bainitic (austempered) structure. PART I. OBSERVATIONS OF MARTENSITE FORMATION IN FOURTEEN CARBON AND LOW-ALLOY STEELS Transformation of austenite to pearlite or bainite can be conveniently studied by observing the change that occurs over a period of time at constant temperature; such studies are the basis of the isothermal transformation diagram (S-curve). This isothermal technique is not applicable to martensite formation, which occurs within a time interval too small to measure, at least by any ordinary means. For all practical purposes, therefore, martensite may be regarded as forming only during cooling and not during a period of time at constant temperature. As austenite is cooled below

Jan 1, 1947

By H. M. Stewart, R. A. Grange

Man.; steel parts may crack if quenched directly into a bath near room temperature, but not if quenched at a temperature just above the range where martensite forms and then allowed to cool slowly to room temperature. This latter procedure, which is the basis for such modern hardening techniques as "martempering" and "isothermal quenching," may entail some sacrifice in depth of hardening but very little, if any, in intensity of hardening. In planning such treatments, it is necessary to know the upper limit of the martensite range, and often desirable to know the proportion of martensite that would form on cooling to any lower temperature; furthermore, the tendency of a particular steel to crack when quenched is unquestionably associated with the temperature range of martensite formation, and consequently knowledge of this range aids in selecting the optimum composition for a given application. Certain puzzling phenomena associated with the properties of quenched and tempered steel may be at least partly clarified as the manner in which martensite forms is more fully understood. Information on the temperature range of martensite formation, particularly with respect to how much martensite results from quenching to a given temperature, is lacking for most commercial types of carbon and low-alloy steel. These considerations led us to study martensite formation in 14 carbon and low-alloy steels. The resultant data may be directly used in the following ways: (I) for selecting the lowest quenching temperature at which no martensite will form; (2) for selecting the highest quenching temperature at which virtually all marten . site will form, thereby avoiding quenching to an unnecessarily lower temperature with attendant danger of cracking; and (3) in producing a mixture of tempered martensite and bainite, which in high-carbon steels has been found to possess somewhat better ductility than tempered martensite, yet does not require the full transformation time necessary for a completely bainitic (austempered) structure. PART I. OBSERVATIONS OF MARTENSITE FORMATION IN FOURTEEN CARBON AND LOW-ALLOY STEELS Transformation of austenite to pearlite or bainite can be conveniently studied by observing the change that occurs over a period of time at constant temperature; such studies are the basis of the isothermal transformation diagram (S-curve). This isothermal technique is not applicable to martensite formation, which occurs within a time interval too small to measure, at least by any ordinary means. For all practical purposes, therefore, martensite may be regarded as forming only during cooling and not during a period of time at constant temperature. As austenite is cooled below

Jan 1, 1947

By Edwin R. Phelps, Charles W. Porterfield

BACKGROUND OF SURFACE MINING Surface mining refers to the process of removing the material (over- burden) overlying a coal seam and exposing the coal so that it can be loaded out and conveyed by trucks or other methods to the preparation plant or other facility. Surface mining has not always been important in the coal industry. In the early 1900s, almost all of the coal mined in the United States was mined by underground methods. In 1900, about 192 Mt (212 million st) of bituminous coal mere mined and less than 1% was mined by surface methods. The development and use of larger equipment has allowed surface mining to gain prominence throughout the years. Table 1 shows the total production of coal by mining method in the United States for ten-pear periods since 1920 and indicates the growth of surface-mined coal production. Surface mining will undoubtedly continue to account for a large percentage of the coal production in the US for many years because: 1) The capacity of surface mining excavators has continued to increase. Today, mining companies are using shovels (see Fig. 1) which are able to move up to 137.6 m3 (180 cu pd) in each bite and draglines (see Fig. 2) which move up to 168.2 m3 (220 cu pd) in each bucketful. 2) Increasing tonnage from the thick western coal seams is being produced by surface mining methods. 3) Surface mining has normally had lower costs and a better safety record than underground mining. 4) Surface mining normally recovers a much higher percentage of the coal reserve than can recovered by underground mining. 5) Some coal measures call best be extracted by surface mining

Jan 1, 1981

By Thomas C. Wilder, Walter E. Galin

Measurements of the electromotive force of the cell at 995°C have shown that the cell may be used to detennine the zinc content of molten Cu-Zn alloys to the nearest 0.05 wt pct. The cell is used for brass melts to which no ZnO is added intentionally, because essentially all oxygen in the Cu-Zn-O system is present in the form of ZnO. The cell is also used to detev,nine the thernodynamic activity of zinc for Cu-Zn al1oys of the brass composition range, from which the equilibrium partial pressure of zinc for such alloy may also be calculated. The electrochemical measurement of the concentration of the most active metal in other engneering- alloy systems by a similar technique is also considered. IN recent years there have been many studies concerning the direct electrochemical measurement of oxygen content in molten engineering metals.' These investigations employ galvanic cells in which the electrolytes are various solid mixed-oxides which conduct current by movement of oxide ion vacancies under the influence of an oxygen potential gradient. Similar cells have also been used to measure the thermodynamic properties of mixing of solid alloys,' one component of which has a much greater affinity for oxygen than the other component(s). One alloy system of obvious engineering importance comprised of two metals of greatly differing affinity for oxygen is copper-zinc. At the casting temperature of molten brasses, the zinc concentration is very difficult to control not only because its vapor pressure is very high, but also because of its high affinity for oxygen to form ZnO. Inasmuch as brass castings may be required to have a zinc concentration within very narrow limits, it would be advantageous for the brass industry to have a means for quickly measuring the zinc content in the molten alloy just prior to pouring. A calculation based on earlier thermodynamic studies of oxygen and zinc in dilute solution in binary Cu-O and CU-Z' alloys indicates that the oxygen concentration of molten 70 wt pct Cu-30 wt pct Zn brass must be much less than 1 ppm for the separate phase ZnO not to exist in the ternary Cu-Zn-O system. Thus it can be assumed that all molten brasses, both in the laboratory and in the foundry, are sufficiently saturated with oxygen for the separate phase ZnO to be present. In view of the foregoing reasons, the cell investigated at 995°C for the purpose of measuring the zinc concentration in molten copper-zinc alloys, particularly those of brass compositions. In this cell the left hand electrode is a reference electrode comprised of a mixture of the pure powders of nickel and nickel oxide, to which a platinum contact is made. The electrode of interest is a molten Cu-Zn alloy of unknown zinc concentration with a tantalum wire contact. The electrolyte is the commercially available (Zircoa) calcia stabilized zirconia. The ZnO in the Cu-Zn electrode compartment is not intentionally added, but is naturally present as a result of the infinitesimal amount of oxygen required for its formation in this alloy system. In cell [A] nickel oxide is reduced at the cathode where O= represents the electrolyte. At the anode dissolved zinc of the Cu-Zn alloy is oxidized to form ZnO(s) in its standard state Thus the overall reaction is for which the molar free energy change is where 5 is the Faraday equivalent (23,063 cal per v equivalent), is the electromotive force of the cell in volts after correction for the Pt-Ta thermocouple, FOf (ZO) and AFF (NO) are the standard molar free energies of formation of the respective oxides, and a2n is the thermodynamic activity of zinc in the molten Cu-Zn alloy. The reference state of zinc in this case is taken to be the pure liquid. At constant temperature all the terms of Eq. [I] are constant except and azn Thus since all other participants in the electrochemical reaction are in their standard states, the change in the electromotive force of Cell [A] is represented only by a change of the thermodynamic activity of zinc at constant temperature. The zinc concentration is related to the activity by sampling the alloy for chemical analysis after cell measurements are taken. EXPERIMENTAL A typical cell construction is shown in Fig. 1. This and other experimental details were similar to that described earlier with the exception of some minor modifications which are detailed below.

Jan 1, 1970

By Sandip K. Mitra, John E. Dorn

The Basinki-modified Seeger equation for deformation of metals by means of thermally activated intersection of dislocations was used to ascertain the various factors responsible for the strain hardening of poly crystalline aluminum and a series of a solid solutions of magnesium in aluminum. The analyses were based principally on the rate of change of strain rate with stress and the Cottrell-Stokes ratio. From measurements of these quantities on a series of strain-hardened states, the force displacement diagram for intersection, the density of dislocations in the entanglements, and the contributions of the athermal local stress and long-range stress interactions were deduced. Poly crystalline aluminum was observed to have the same force-displacement diagram for intersection as single aluminum crystals, and the corresponding diagram for the magnesium alloys differed in a manner that suggested that small amounts of magnesium may decrease slightly the stacking fault energy of aluminum. Although poly crystalline aluminum had about the same initial density of dislocations as the single crystals, the density of dislocations increased much more rapidly with strain in the polycrystals. This factor and the higher back stresses generated in the poly crystalline aluminum were the major factors responsible for the higher strength of the poly crystalline aluminum, As the magnesium content of the alloys was increased, the initial density of the dislocations also increased. Most of the alloy strengthening arose from this factor; alloying effects due to short-range order strengthening, Suzuki locking, Cottrell pinning, and so forth, were estimated to be quite small. The gveater rate of strain hardening in the alloys was deduced to be due primarily to the greater increase in the density of the dislocations with strain as the magnesium content was increased. It is the purpose of the present investigation to shed additional light on the average plastic behavior of a grain in a polycrystalline aggregate, by analyses, based on dislocation theory, of appropriate experimental test results. These deductions when compared with those from single-crystal data will then serve to illustrate the major items of difference between the plastic behavior of single crystals and the average behavior of a grain in a

Jan 1, 1963

By W. C. Knopf, I. M. LeBaron

In the Carlsbad area potash is dry-mined and wet-concentrated. Wet concentration involves recircu-lation of saturated brines, with resultant difficulties of brine disposal and inherent losses in recovery through solution of the KCl. A dry beneficiation process for potash minerals would offer many advantages over present methods. International Minerals & Chemical Corp. has been working on development of a dry process for some time. This article is a progress report. Potash ores available in this country are largely mixtures of halite and sylvite contaminated with slimes and often with many sulfate and magnesium minerals. Electrostatic concentration of KCl when the contaminating minerals are in low percentage has been accomplished. Continuous production of better than 60 pct K2O concentrate from high potassium sulfate ores has not been demonstrated at the present time. Laboratory Procedure: Laboratory work on electrostatic beneficiation of potash minerals has been carried out using both pure crystals and mined ore. Electrostatic separation of the two main components of potash ore, sylvite (KCl) and halite (NaCl), is accomplished by first heat-treating the ore to allow for electric charging of the particles. Preferential exchange of' electrical charges between different mineral particles through surface contact is the principal mechanism involved. After proper conditioning of the particle surfaces, the oppositely charged minerals are dropped vertically into a horizontal electric field having a field gradient from about 5000 to 15,000 v per in. The particles are free-falling in the field. A suitable power supply is used to produce the proper total voltage for the electric field. Total required voltage is, of course, a function of electrode spacing. A typical electrode arrangement with power supply is shown in Fig. 1. In general, the total voltage on any system is 75,000 to 90,000 v. Power consumption is negligible, as leakage at insulators and corona discharge are the main losses in the system. Optimum preheat temperatures for pure material components generally are not the same as for mined minerals. Fig. 2 shows clearly that maximum recovery on a 50 pct pure KCl, 50 pct pure NaCl mixture occurs at a preheat temperature of about 600°F. In this test the preheat mixture was allowed to cool to about 225°F before dropping through the electric field. The optimum temperature for sep- aration is practically never the optimum temperature of preheat. Optimum separation for mined Carlsbad potash ores occurs when a preheat temperature closer to 900°F is used, as shown by Fig. 3, which plots percent sylvite (KCl) in the concentrate and tails as a function of preheat temperatures. Here again the ore was allowed to cool to about 225°F before separation. It should be noted that efficiency of separation increases continuously with increasing preheat temperatures until incipient fusion of the ore takes place at about 950°F. The increase in optimum preheat temperature for Carlsbad ores over the pure minerals can be traced to the effect of the clay slimes intimately associated with the mined potash. If the slimes associated with the ore are removed by washing with alcohol or other non-aqueous material, the necessity of preheating to the high temperature is removed. This temperature range has been proved critical within these limits during continuous pilot plant operation. Concentrate and tailings grade and optimum recovery also are functions of the temperature of the potash mineral particles during separation in the electric field. Fig. 4 presents data indicating that feed to the rougher pass should range between 150°F and 275°F for best results. It has been found by subsequent pilot plant work that these temperatures can be maintained in actual operation.

Jan 1, 1959

By Robert T. Howard, Morris Cohen

THE purpose of this paper is to direct attention to the lower part of the austenite transformation diagram, or TTT curves, where considerable uncertainty still exists as to the blending of the bainite and martensite reactions. The dissimilarities between these two low temperature trans- formations have been summarized recently by Troiano and Greninger.1 Of particular importance to the present work is the fact that martensite formation, unlike bainite formation, is not suppressed by the fastest quenching rates attainable. However, most of the investigations in this direction have been concerned primarily with the start of the martensite transformation [(M,)] during cooling, rather than with the quantitative aspects of the subsequent reaction on either further cooling or isothermal holding. Several years ago, Cohen2 suggested a form of transformation diagram which reconciled, at least schematically, the isothermal characteristics of the bainite transformation with the insuppressible nature of the martensite transfoi mation. In essence, the diagram consisted of a family of horizontal lines (representing the progress of the austenite-martensite reaction during quenching) which were interrupted in point of time by the bainite family of C-curves extending down from higher temperatures. This picture received later support by metallographic3 and dilatmetric4,5 studies, but the techniques employed were not sufficiently critical to substantiate some of the details. Elmendorf6 and Grange and Steward7 have shown that the Greninger-Troiano metallographic techniques may be used not only for [M,] determinations, but for studying the course of further austenite decomposition. In fact, the latter investigators have observed the amount of martensite formed as a function of temperature below [M,] in fourteen commercial steels. In the present paper, the Greninger- Troiano technique has been combined with the lineal method of Howard and Cohen9 to provide a unique, quantitative treatment of the austenite-martensite reaction. Furthermore, isothermal transformations have been similarly measured on holding at temperatures above and below M,, thus permitting a fairly detailed presentation of the lower part of the conventional austenite transformation diagram. EXPERIMENTAL DETAILS Materials The steels, listed in Table I, were selected for study because they exhibit a satisfactory range of M, temperatures and have not received as much attention as the lower carbon grades. They were prepared as 30 lb induction furnace heats in the laboratories of the Vanadium-Alloys

Jan 1, 1947

By H. I. Aaronson

ONLY limited studies have been made of pro-eutectoid a morphology in hypoeutectoid Ti-Cr alloys during previous investigations, 1-3 The nature of the eutectoid reaction, ß?a + TiCr2, has been considered in somewhat more detail.1,2,4,4 Inasmuch as the proeutectoid ferrite reaction in plain carbon steels has recently been extensively investigated in this laboratory, advantage has been taken of the opportunity to compare the morphologies developed in two rather different eutectoid systems in order to permit a preliminary estimate to be made of the degree of generality which may be ascribed to the various results obtained. Experimental Procedure The chemical analyses of the alloys used are given in Table I. The 7.22 pct Cr alloy was prepared by the Battelle Memorial Institute from iodide titanium and high purity chromium. The initial ingot was made by arc melting a master alloy and iodide titanium. The resultant ingot was twice remelted. by the consumable electrode technique. Upset forging at 950°C reduced the final ingot to 3/4 in. sq bars. The 10.94 pct Cr alloy was prepared by the Titanium Metals Corp. of America from sponge titanium and chromium of ordinary purity. This alloy was melted twice by the consumable electrode technique, forged at 1200° to a 3/4 in. sq bar, and rolled at 815°C to a 1/2 in. round. Specimens 1/4 in. x 1/4 in. x 1/16 in. were cut from the 7.22 pct Cr alloy; halves of disks 1/2 in. in dlam and 1/16 in. thick served as specimens of the 10.94 pct Cr alloy. These specimens were individually encapsulated in Vycor, Each capsule was evacuated three times to a pressure of 2 to 6 x lo-' mm Hg and flushed twice with a reduced pressure of argon prior to sealing. The inclusion of two small packets of zirconium chips permitted the capsules to be outgassed and internally gettered, without heating the specimen, prior to heat treatment. The specimens were solution annealed for 40 min at 1000°C, isothermally reacted in deoxidized lead pots, and quenched in iced water. Most of the isothermal reaction treatments were carried out at 25" intervals between 725° and 550°C on specimens of the 7.22 pct Cr alloy and at 650°C on specimens of the 10.94 pct Cr alloy. Results and Discussion Morphologies of Proeutectoid a*—Kinetics of the ß?a Reaction: Fig. 1 shows the TTT-curves for the beginning of the proeutectoid a and the eutectoid reactions in the temperature range investigated in the 7.22 pct Cr alloy. The M, temperature in this alloy is below 0°C. The curve for the beginning of the proeutectoid a reaction is in good agreement with that of Frost, Parris, Hirsch, Doig, and Schwartz' for a 7.54 pct Cr alloy, appears at considerably earlier reaction times than the corresponding curve presented by Rostoker4 for a 7 pct Cr alloy, and lies athwart the curve of Spachner and co-workers11 for an 8 pct Cr alloy. Differences in criteria for the beginning of the a reaction and the relatively high rate of this reaction at lower temperatures presumably account for some of these divergences and may also explain, in part, the apparent failure of an iodide-base 7.22 pct Cr alloy to react more slowly than sponge-base 7 to 8 pct Cr alloys. Grain Boundary Allotriomorphs: Crystals of this morphology nucleate at and grow preferentially and more or less smoothly along the matrix grain boundaries.5-7 Allotriomorphs are the first morphology to precipitate at most grain boundaries at reaction temperatures from 725°C, the highest temperature studied, at least through 575°C. Fig. 2 Shows typical allotriomorphs, formed during an early stage of transformation at 725°C. Grain boundary allotriomorphs are not the predc=iaant morphology at late reaction times at any temperature studied. However; from 725" through

Jan 1, 1958

By K. K. Clark, W. H. Somerton

The feasibility of reducing pressures needed to fracture formations by heating the borehole intensively has been investigated on a laboratory scale. Reductions in fracture pressures of heated Bandera, Berea and Boise sandstone cores were observed at various confining pressures, simulating reservoir stress conditions to a depth of 5,000 ft. Both permeable and impermeable borehole boundary conditions were tested. Three heating conditions were used in the tests. One group of samples was heated in a furnace to a temperature of 600C, cooled to room temperature., and then tested for fracture pressures. Two other groups of samples were heated with a borehole heater to maximum borehole wall temperatures of 625 and 792C, then cooled and tested. In all cases, heated samples showed substantial reductions in fracture pressures compared to unheated samples. The higher temperature borehole heated samples showed the greatest reduction, amounting to 70 per cent in the extreme case. In actual application to wells, it was concluded that fracture pressures may be reduced by 20 to 50 per cent by intensive borehole heating before fracturing. INTRODUCTION Hydraulic fracturing of oil and gas wells is an accepted and effective technique for increasing well productivity in many areas, but is notoriously unsuccessful in other areas. The technique becomes less interesting for production stimulation purposes when excessive fracture pressures are necessary. Formations which do fracture but require high pressures, probably form horizontal fractures. There is good evidence to indicate that some formations do not fracture at all, but yield plastically under high pressure. Thus, any technique which will make formations easier to fracture or will reduce fracture pressures is potentially interesting to petroleum producers. Horizontal fractures generally occur in planes of weakness, such as bedding planes, and at pressures equaling or exceeding the effective overburden stress. In the unusual case where planes of weakness do not exist, a slightly higher pressure must be applied to overcome the tensile strength of the rock. In general, high fluid-loss fracturing fluids are conducive to the formation of horizontal fractures.1 Vertical fractures, when they do occur, are initiated at pressures generally lower than overburden pressure and are favored by low fluid-loss fracturing fluids or impermeable boreholes. The pressure necessary to create a vertical fracture in an impermeable, elastic medium can be predicted from the so-called "thick-walled cylinder" equation: Pf = 2 or+ st.............(1) where pf = fracture pressure or = radial stress st = tensile strength of medium. This equation indicates that to fracture it is necessary to overcome a tangential compressive stress, induced by and equal to two times the radial confining stress, plus the tensile strength of the medium. If the above equation held strictly true for rocks, very little could be done to reduce fracture pressure. The tensile strength of the rock might be reduced somewhat, but any change in st would be small compared to the magnitude of the tangential stress at reasonable subsurface depths. Laboratory tests have shown that actual fracture pressures exceed values predicted from Eq. 1 by factors of 1.4 to 28.6.1,2 Eq. 1 thus does not properly describe the

Jan 1, 1966

By J. Szekely, J. W. Evans

A formulation is given for radiant heat loss from the surface of molten steel held in a ladle and numerical solutions are presented for the resultant integrodif-ferential equations. The results are shown on plots of the net rate of heat loss against time, with the geometlry, emissivities, thermal properties of the walls, and Preheat as parameters. It was found that the rate of the vadiant heat loss from the exposed metal surface was comparable to the conductive losses from the Portion of the metal in direct contact with the walls. The radiant heat loss was markedly affected by the geometry and by the thermal properties of the walls the net radiant heat loss from the steel could be drastically reduced by providing a top cover and also by the provision of preheat. The results of this work are thought to be relevant to heat loss calculatious concerned with vacuum degassing and with operatzons in which molten metals are held in ladles. WhEN molten steel is held or transferred in ladles heat losses inevitably occur. The accurate assessment of these heat losses is of importance in process design considerations, as the temperature of the metal may have to meet rigid specifications in certain stages of the processing sequence. As shown in Fig. 1, the principal mechanisms responsible for heat loss are: i) Conduction into the walls and ii) Thermal radation from the top metal surface (possibly augmented by natural convection) The conductive heat loss may be readily calculated as much information is available in the literature on both the general theory1'' and on its particular applications to ladle heat loss calculations. Evaluation of the radiant component of the heat loss is less straight-forward. In two recent papers it has been shown that the presence of a slag layer 2 to 3 in. thick would prevent any significant heat loss for periods up to an hour,5'" however, should bare metal be exposed, constituting the radiant surface, substantial heat losses could occur. This latter case is of considerable practical interest in various degassing operations and in the vacuum processing of certain specialty steels, where the top surface of the metal is not covered by a protective slag layer. Under these conditions a substantial radiant flux is emitted by the top metal surface; part of this flux leaves the system, part of it falls in the ladle walls. The radiant flux incident on the ladle walls is in part absorbed and conducted away, and in part reflected and re radiated. The purpose of the paper is to present a general formulation of the problem, the resultant solution of which should enable computation of the net radiant heat loss from the metal surface. The principal variables, the effect of which will be investigated, include initial metal temperature, geometry, thermal and radiative properties of the wall, and various modes of preheat. FORMULATION Consider a cylindrical enclosure, shown in Fig. 2, with a diameter D, height H, and wall thickness W, the cover at the top of the cylinder having a wall thickness L. If the temperature of the base (representing the top surface of the metal) is kept constant, say, Tb, the problem is to calculate the net rate of heat loss from the base as a function of time, geometry, the initial temperature distribution, and the thermal properties of the system. As customary, the formulation will consist of the mathematical statement of the conservation of thermal energy within the system.

Jan 1, 1970

The Institute assembled in the room of the Board of Trade, Mercantile Library Building, on Tuesday evening, May 20th, at 8 o'clock P.M, Hon. W. D. Kelley made an address of welcome to the Institute, which was responded to by President Raymond. The Council reported, through the President, that the Institute had during' the past year, held three meetings, in New York, Pittsburgh, and Boston. Thirty-nine papers of great professional value had been read, and sixty-nine members and twenty-four associates added to the Institute. The finances of the Institute were in good condition, the report of the Treasurer, Mr. T. D. Rand, showing a balance on hand of $1569.74. The report was accepted. Prof. L. Gruner, of Paris, France, and Prof. Peter von Tunner, of Leoben, Austria, having been proposed by ten members of the Institute, and approved by the Council, were unanimously elected honorary members. The Council reported the names of the following gentlemen, with recommendation for election as members or associates. They were unanimously elected. *Barker, Prof. G. F.,.....Philadelphia, Pa. Chalfant, John W.,.....Pittsburgh, Pa. *Crawford, John J.,.....Philadelphia, Pa. *Crusius, O. H.,.Columbus, Ohio. Dibble, Theo. Hoyt,.....Pottsville, Pa. Everson, John,. Pittsburgh, Pa. Field, Robt. P.,.Philadelphia, Pa. *Fisher, F. A.,.Bristol, Tenn. *Fleming, W. E.,.Pioche, Nev. *Harden, E. B ,..Philadelphia, Pa. *Harding, G. E.,.New York. Hartshorne, J.,..'Philadelphia, Pa. Haupt, H., Jr.,.Philadelphia, Pa.

Annual Meeting, Philadelphia May, 1873 THE Institute assembled in the room of the Board of Trade, Mercantile Library Building, on Tuesday evening, May 20th, at 8 o'clock P.M, Hon. W. D. Kelley made an address of welcome to the Institute, which was responded to by President Raymond. The Council reported, through the President, that the Institute had during the past year, held three meetings, in New York, Pittsburgh, and Boston. Thirty-nine papers of great professional value had been read, and sixty-nine members and twenty-four associates added to the Institute. The finances of the Institute were in good condition, the report of the Treasurer, Mr. T. D. Rand, showing a balance on hand of $1569.74. The report was accepted. Prof. L. Gruner, of Paris, France, and Prof. Peter von Tunner, of Leoben, Austria, having been proposed by ten members of the Institute, and approved by the Council, were unanimously elected honorary members. The Council reported the names of the following gentlemen, with recommendation for election as members or associates. They were unanimously elected. *Barker, Prof. G. F., - Philadelphia, Pa. †Chalfant, John W., - Pittsburgh, Pa. *Crawford, John J., - Philadelphia, Pa. *Crasius, O. H., - Columbus, Ohio. †Dibble, Theo. Hoyt, - Pottsville, Pa. †Everson, John, - Pittsburgh, Pa. †Field, Robt. P., - Philadelphia, Pa. *Fisher, F. A., - Bristol, Tenn. *Fleming, W. E., - Pioche, Nev. *Harden, E. B, - Philadelphia, Pa. *Harding, G. E., - New York. †Hartshorne, J., - Philadelphia, Pa. †Haupt, H., Jr., - Philadelphia, Pa. * Members. † Associates.

Jan 1, 1874

By Katherine Mather

This paper is an attempt to put together petrographic, physical, and chemical data about the large and varied group of rocks generally called limestones. Results of the properties of these rocks on their performance as concrete aggregates are discussed. CRUSHED stone classified as limestone comprises about 70 pct of the crushed concrete aggregate and roadstone produced in this country, and the product has amounted in the recent past to about 97 million tons per year, valued at more than $123,-000,000.' Therefore an evaluation of the properties of rocks classified in these statistics as limestones and an ability to predict their behavior as aggregate in portland cement concrete are matters of considerable economic consequence to producers and consumers of aggregate. This paper is an attempt to put together petrographic, physical, and chemical data about the large and varied group of rocks generally called limestones. It also points out certain results of the properties of these rocks on their performance as concrete aggregates. The problems encountered in efforts to explain the behavior of limestones, dolomites, and related rocks as concrete aggregate lie on the frontiers between the ignorance of the geologists and the ignorance of the engineers. The mineral composition and textures of the rocks have not been adequately explored and described, and the effects of mineral composition and texture on the performance of aggregates have been worked out in only a few cases. There are instances of unsatisfactory performance and some generalizations of those experiences. There are many more examples of satisfactory performance than of unsatisfactory, but there is practically no published information as to why a particular aggregate was satisfactory in a specific use. There are all too few investigations of undeteriorated concrete; more is known about the pathology of concrete than about its normal structure and aspect. As part of the investigation of aggregates, the classification of rocks is not an end in itself, but is a step toward recognition of properties expected to influence the behavior of aggregates in their intended use. Rock of any type may perform well or poorly as concrete aggregate, depending on the physical condition of the rock, its physical and chemical properties as compared to those of the matrix in which it is placed, and the circumstances under which the concrete is exposed. Some classification and descrip- tion of the limestones is necessary to distinguish subgroups whose members will probably behave in similar ways and to evaluate the group as a whole. The group name, limestone, includes and tends to conceal the identity of rocks that have a very great range in chemical and mineralogical composition, structure and texture, conditions of deposition, subsequent history, and suitability for use as concrete aggregate. All the rocks classified as limestones include one or both of the most abundant rock-form-ing carbonate minerals, calcite and dolomite, as major constituents. Carbonate rocks is a more descriptive and more accurate group name than limestone, which has generally been used to include dol-omitic or magnesian limestones and dolomites as well as calcitic limestones. Classification and Description of the Carbonate Rocks The question of names to be applied to the carbonate rocks may appear academic, but it is of practical importance as long as there are specifications for aggregate that result in acceptance or rejection of rocks by name and as long as failure to make distinctions among the carbonate rocks leads to misunderstandings of their properties and of what may be expected of them in performance. Materials included in production statistics as limestone and tested for concrete aggregate as limestone range in fact from calcareous shale and calcareous or dolomitic sandstone and argillaceous limestone and dolomite to calcitic limestone and pure dolomite and to rocks close to calcitic and dolomitic marble. To take a simple example: consider two manufactured coarse aggregates, both dense and unweathered, one composed principally of crystalline calcite and the other of crystalline dolomite, but both known as limestone. If each is used as coarse aggregate in a standardized concrete mixture with the same fine aggregate used in both cases, made into concrete specimens, and tested in accelerated freezing-and-thaw-ing, the test results will probably be quite different, because of the differences in internal bond and thermal properties between the two aggregates. As long as both aggregates are called limestone, the test results cannot be explained, or used except in a very limited way; if the aggregates are distinguished as limestone and dolomite and the consequent differences in texture and thermal properties are recognized, the results are reasonable. Classification of the carbonate rocks used as aggregate should be based on structure and on composi-

Jan 1, 1954