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By M. B. Kasen

eralized strain equation appear quite different they are really identical. This identity can be shown in a simple mathematical rearrangement. Referring to Eq. [I], the substitution of ln(1 +?E) for ? (where ?E is the engineering strain) yields: a =s8- (s8?so)e?1/?c) ln(1 +s8??E) [4] or s =s8?(s8?so)(l +?E)-1 [5] Replacing (1 + cE) with its equivalent, (l/lo): s=s8?(s8?so)l/€c [6] Adding and subtracting the quantity (s8 - so) yields: s=s8?(s8?so)+ (s8-so) l-(^)l/£c [7] = s0+(s8?so)ri-(^0)1/^l [8] Letting n = 1/?c it follows that: This equation will be seen to be identical to the generalized strain expression in Eq. [3], the coefficients a and b being equal to so and (s8?so)/?c or n(s8?so), respectively. Confirmation of the relationship noted above was provided by fitting experimental stress-strain data5 for 304 stainless steel (total strain from 0.072 to 1.15) at room temperature at a true total strain rate of 4 x 10-3 sec-1 to Eqs. [I] and [3]. In this study nonlinear regression analyses were used to evaluate the equation constants. The results of this evaluation are presented in Table I. When these constants are used in Eqs. [I] and [3] the curves obtained are coincident. It will also be noted that the value of n is equal to YE,, the value of a is equal to so, and the value of b is equal to (s8 - so)/Ec. These equalities are illustrated in Table I where the Voce constants in parentheses were calculated from a, b, and n values. Excellent agreement is seen to exist. It is concluded, therefore, that the Voce equation is, indeed, identical to the linear expression involving generalized strain. In applying the generalized strain concept to stress-strain data, it has been noted3 that data within what was termed the transition region (first few points in the plastic flow region) must be excluded from the analysis in order to obtain the linearity in Eq. [3]. A similar exclusion applies in the use of the Voce equation since an expression which would describe the strain-hardening region would usually be found to be not too effective in the elastic-plastic fillet2 which connects the elastic line to the strain-hardening region. These considerations further substantiate that Eqs. [I] and [3] do indeed describe identical behavior. Note added in proof. Since this manuscript was submitted for publication, correspondence with Professor S. R. Davies, University of Edinburgh, (of Ref. 3) indicated that his associates have also confirmed the identity between the Voce equation and the generalized strain concept. It is sometimes desirable to obtain a small grain size in super-purity metals. Unfortunately, the absence of impurities, the low recrystallization temperatures, and the low activation energies for boundary motion result in rapid growth of relatively few grains when conventional recrystallization practices are used. For example, Arajs et al.1 were unable to obtain a mean grain diameter less than about 0.4 mm in iron containing 38 at. ppm of impurity. The <1 at. ppm purity range in many modern superpurity metals results in yet larger limiting grain sizes. It is the purpose of this communication to describe a technique by which mean grain diameters as small as 30 µ have been produced in aluminum having an impurity content < 0.5 at. ppm. The technique should be applicable to superpurity metals other than aluminum and has potential applications to grain refinement of commercial alloys. The necessary and sufficient requirements for production of a small grain size are as follows: a) a large number of potential nucleating sites must be created randomly within the cold worked material, b) sufficient thermal energy must be provided at a rate that will simultaneously activate a large fraction of the potential nuclei, and c) the grain growth must be stopped shortly after grain impingement.&apos; A high density of potential grain nuclei may be cre-

Jan 1, 1970

By J. W. Stewart

THE aim of this paper is to point out the various ways in which design is taught in standard four-year mining engineering curricula in American colleges and universities; to discuss the reasons apparently justifying the great diversity of procedure that exists along this line; and to indicate possible solutions to the difficulties now attending the teaching of design. A study was made of typical four-year mining engineering curricula in 15 colleges and universities of the United States (Table I). Great diversity in content of the various design courses, if taught as described, became apparent immediately. It appeared that it might be both interesting and instructive if the cause for this colorful variety could be found. Was the cause local mining conditions, departmental individuality, or actual necessity? Study of the school catalogues revealed that the various design courses could be classified into two fairly distinct types; namely, (I) those dealing with mine equipment or structures, and (2) those dealing with underground layout of mine workings. It seemed reasonable to label the first type of course "mining design" and the second type "mine design." Some such distinction must have been in the minds of mining educators, for we find both names (i.e., mine design and mining design) used to designate the design courses included in the various mining curricula. Unhappily, the name and the apparent content of such courses do not, in all instances, conform to the classification made here, although in most instances such conformation does obtain (Table I), It is possible that in the exceptional case the method of teaching a design course, and consequently the content of it, may have been changed at some time without change in the name of the course. Apparently only 3 of the I5 schools made an effort to cover both types of design in their curricula (Table I). Two other schools did not list or describe any course of study that could with certainty be taken as a design course. Of the remaining 10 schools, 9 gave a mining-design type of course and one gave a strictly mine-design type. Obviously, the preponderance of design courses afforded are of the mining-design type. Why? To answer this question, it is necessary to make a study of the prerequisites for the two types of design courses taught. In all schools the design courses were taught in the senior year; in eight, design was taught throughout both semesters; in two, it was taught during the first semester only; and in three it was taught during the second semester only. Consequently, excepting the last three cases noted, it is obvious that all prerequisites for a design course should be completed by the end of the junior year. PREREQUISITES What are the probable prerequisites for the different kinds of design courses as they appear to be taught? Let us take first the case of the mine-design type of course, having to do with layout of underground workings. Here it would seem that logical

Jan 1, 1941

Sponsored by the Utah Section of AIME, the 1960 Annual Rocky Mountain Minerals Conference brought more than 500 members of the mineral industry to the conference rooms of the Newhouse Hotel in Salt Lake City during the period October 5 to October 7. Under the overall supervision of J. M. Ehrhorn, general chairman, and H. E. Wright, assistant general chairman, the Conference offered 14 papers of technical interest in four sessions during the first two days, and it culminated in the Exploration Symposium held on the morning of October 7. The opening session of the Conference, meeting on the morning of October 5, heard R. P. Work, utility engineer of the Mountain Fuel Supply Co., discuss the growth and economic factors related to the use of natural gas in Utah. In his paper entitled Natural Gas Market In Utah, Mr. Work emphasized the very good prospects of natural gas serving almost all parts of that state. He mentioned that recently the El Paso pipeline, built to serve the Pacific Northwest with gas from Texas, has permitted natural gas to be available to Eastern Utah, and a tentatively planned pipeline from Green River, Wyo. to Southern California will bring natural gas to the central and southeast areas of the state.

Jan 11, 1960

By D. K. Deardorff, H. Kato

THE transformation temperature of hafnium from hcp to bcc is 1750° + 20 °C in contrast to previously published values by Duwez and fast2 which are believed inaccurate. The Bureau of Mines determined this value by measuring the temperatures at which a sudden decrease was observed in the electrical resistance of hafnium alloys containing up to 18 at. pct Zr, and then extrapolating the temperatures to 0 pct Zr. Duwez1 previously reported a transformation temperature of 1310°C, obtained by use of a rapid-cooling thermal-analysis technique. Later, Fast2 published a much higher value of 1950° 100°C based upon: a) his own determination of 1690°C for the beginning of the transformation range for hafnium containing 5.7 at. pct Zr, and b) his belief that lattice parameters reported3 for Duwez&apos; hafnium indicated i contained 19 at. pct Zr and that the 1310°C value should be associated with that composition. Russell4 subsequently stated that Duwez and Fast probably were reporting their lattice parameters in different units and reassigned a value of 6.0 at. pct Zr to Duwez&apos; hafnium. If this assumption were true, Duwez&apos; hafnium was of approximately the same purity as Fast&apos;s, but his transformation value was considerably lower. The reason for this divergence is not known. Fast&apos;s value for hafnium containing 5.7 at. pct Zr agrees well with our data. McGeary5 determined a transformation value of 1660°C by metallographic examination of heat-treated specimens, but the zirconium content of his hafnium was not definitely established. The specimens used in this investigation were 1/8-in.-diam rods 6 1/2 in. long, fabricated by hot-swaging bars cut from arc-melted buttons. These buttons were melted from crystal bar hafnium and crystal bar zirconium. The chief impurity in the hafnium

Jan 1, 1960

By A. Perry Thompson

I. GENERAL DISTRIBUTIONS APPEARANCE, AND RELATIONS MINING in Butte has seldom encountered covellite in commercial quantities. The notable occurrences, extending vertically and laterally perhaps several hundred feet, presented scattered, bladed lenses of this mineral in a matrix of quartz, pyrite, enargite, bornite, and chalcocite. But under the microscope covellite is seen to be widespread in the Butte ores, though in such small quantity that megascopic examination will not discover it. In early mining operations covellite was found in the Gray Rock vein, a member of the Blue vein system. As described by Sales,1 it persisted as an occasional ore mineral, associated with quartz, pyrite, tetrahedrite, bornite, chalcocite, and minor amounts of sphalerite, from the 700-ft. to the 1,200-ft. level of the East Gray Rock and Diamond mines. Crushed quartz-monzonite and later quartz completed the vein filling between

Jan 4, 1915

By W. C. Ellis, M. E. Fine

INTRODUCTION IN the iron-cobalt system there are several property-composition relationships of theoretical importance. The alloys are ferromagnetic exhibiting a maximum saturation at approximately 33 at. pct cobalt;1 the interatomic distances at room temperature show large positive deviations from Vegard&apos;s rule;2 and in the region of 50 at. pct cobalt there is an order-disorder transition which has a large effect on other properties. The thermal expansion data for the system have been reported3,4.5,6 only from room temperature to 100°C, and for this temperature range a minimum in coefficient in the region of 50 at. pct cobalt is well established. 3,4,6 This paper reports the expansion properties in the iron-cobalt system for temperatures from 30 to 850°C, shows the correlation with other physical properties, and proposes an explanation for the interrelation. EXPERIMENTAL PROCEDURE The thermal expansivities of iron-cobalt alloys in 10 pct steps in composition were determined from 30 to about 850°C. The major series of alloys was prepared by induction melting from Armco iron and cobalt rondelles with the addition of about 0.6 wt pct manganese. This series was supplemented by three compositions: high purity electrolytic iron, high purity electrolytic cobalt, and an alloy of 52.1 wt pct cobalt prepared from the high purity electrolytic iron and electrolytic cobalt by vacuum-melting. The compositions of both series of alloys are given in Table I. The specimens for measurement were prepared from the cast ingots by first hot-swaging followed in most instances by cold-swaging to 0.180-in. diam rod. Some alloys in the region of 50 pct cobalt, because of inherent cold-brittleness, were swaged hot to the final diameter. The expansion specimens were approximately 2 ½ in. long with flat-ground and polished ends. Before measurement they were annealed in hydrogen for r hr at 900°C. In the apparatus used for the thermal expansivity measurements (Fig I), a modification of that designed by Hidnert and Sweeney,7 the differential expansion between the specimen and vitreous silica was measured by a dial micrometer reading to 0.001 mm. An invar fixture supported the dial gauge, and this fixture, in turn, was connected to a split invar collar which supported the outer silica tube. A pyrex "T" was placed over the outer silica tube and was sealed to the invar collar by Apiezon Q sealing compound. A movable hollow vitreous silica rod which rested on top of the test specimen slid through a flexible rubber diaphragm cemented to the top of the pyrex "T." The specimen rested on a vitreous silica pedestal fused into the outer tube and was kept centered by a silica sleeve. A non-oxidizing atmosphere was obtained by admitting hydrogen into the

Jan 1, 1948

By I. C. Edwards, W. J. Dengler, D. F. Lillie

By 1959 the milling plant of Bicr oft Uranium Mines Ltd. had been in operation for two years. During this time many changes, both physical and chemical, had been made in an effort to improve plant efficiency. Investigational work had been facilitated greatly by the fact that the Bicroft mill consists to a large extent of two parallel circuits; however, the number of interrelated variables is great, and it is difficult to measure, let alone control, all of them simultaneously. The Bicroft mill staff feels that much progress has been made, but there is much yet to be done and many questions still to be answered. The Addendum outlines noteworthy developments to the fall of 1960. The property of Bicroft Uranium Mines is located in Cardiff Township in the Highlands of Haliburton County, southeastern Ontario, some 140 miles northeast of the city of Toronto and 70 miles due north of Lake Ontario. The ore-bearing dikes are classed as mafic. nonsegregated, syenite and granite pegmatite. Microline, microline perthite, quartz, sodic plagioclase, and pyroxene or biotite are the main constituents with uraninite, uranothorite, fluorite, titanite, cryolite, calcite, pyrite, pyrrhotite, and molybdenite as important accessories. The mill was designed for a nominal capacity of 1000 tpd with all sections except grinding, leaching, and clarification capable of 1500 tons. Feed was entered to one mill unit on Oct. 29, 1956, and to the second parallel line on November 13. The total feed rate was at 1000 tpd on November 15. The first precipitate was produced on November 12. By August 1958 the average entry had been increased to 1410 tpd without the installation of any additional major equipment. The overall recovery has been greater than 93.0 pct on several months, but a recovery of 92.5 pct appears to be most economical. Operating time for eight months in 1958 has been 97.1 pct although all grinding mills were relined during this period. A generalized flowsheet is shown in Fig. 1 and the plant layout in Fig. 2. The ore is crushed in three stages to -3/8-in. and then is wet ground in two stages to approximately 50 pct -200 mesh. The pulp is thickened to 68 pct solids and agitated by air for 33 hr at a pH of 1.9 as maintained by the addition of 93 pct sulfuric acid. Separation of the uranium-bearing pregnant liquor from the residue is accomplished by five-stage countercurrent decantation (CCD), following which the liquor is clarified for removal of fine slimes. The clarified liquor is treated in the ion exchange equipment in which both the upgrading of the solution and the removal of some impurities is accomplished. Addition of magnesia to the high-grade eluate from ion exchange precipitates the uranium as sodium-magnesium diuranate. After being filtered and dried this precipitate is packaged in drums for shipment. The residue from CCD and the barren solution from ion exchange are neutralized by the addition of lime and pumped to the tailings disposal area. CRUSHING All crushing is done on surface in a conventional three-stage unit consisting of a 36x48-in. Birdsboro Buchanan jaw crusher, a 4 1/4-ft standard Symons cone crusher, and a 4-ft Symons shorthead. Both the two primary and the two secondary screens are 4X 8-ft Symons rod decks. Moisture content of the feed ranges from 3 to 4 pct. In the initial installation the primary screens were equipped with 5/16-in. rods at 1 1/4-in. spacing and the secondaries with 5/16-in. rods at 1/2-in. spacing. With this arrangement the average crushing rate was better than 180 tph. To reduce the work required of the grinding circuit as tonnage was increased, several changes were made in the crushing circuit. By September 1957, both secondary screens had been reduced to 3/8-in. spacing, and rod size had been reduced to 3/16-in. to provide increased open area. Average throughput was reduced to 160 tph. Later, 50 pct of one primary screen was decreased to 1/2-in. opening to better balance the load between the two cone crushers. In May 1958,

Jan 1, 1961

By C. H. P. Lupis, Henri Gaye

There are m -I conditions for the stability of a solution of m components with respect to infinitesinzal flucturations. However, in most cases, only one of these conditions has to be considered to determine the domain of instability and the existence of this more restrictive condition greatly simplifies the calculations. It may be used advantageously for the prediction of miscibility gaps and the method is illustrated in details for the case of the Ag-Pb-Zn system. THE thermodynamic conditions for the formation of a miscibility gap may be viewed as a necessary consequence of the conditions for spinodal decomposition. A previous article1 has examined in detail the form of these conditions for multicomponent systems. There is only one condition for the stability of a binary system (with respect to infinitesimal fluctuations), but there are two conditions for a ternary system, and m — 1 conditions for an m-component system. The probability of violating a stability condition, and thus forming a miscibility gap, obviously increases with the number of components, a result which is rather intuitive since the atoms of the solution have now many more ways of redistributing themselves and introducing complexities in the form of the free energy hy-persurface. It is of interest to take advantage of this possibility of precipitating new phases and to examine which stability condition is the likeliest to be violated, that is, which stability condition is thermodynamically the most restrictive. The finding of such a condition would greatly simplify the application of the stability criteria since only one condition could then be considered, instead of m - 1. In Ref. 1, coherency strain energy terms were neglected, thus restricting the applications of the treatment to solutions where they are negligible, such as liquid alloys. In the following study the same assumption will be made. To generalize the treatment to systems where the strain energy terms are sizable, the reader is referred to Cahn&apos;s classical article on spinodal decomposition.2 Let us designate by Gij the second derivative of the Gibbs free energy with respect to the number of moles ni and n j. There are several equivalent sets of m — 1 stability conditions.&apos; The one considered here expresses that the successive diagonal determinants of order 1, 2, ... m — 1, associated with the symmetric Gij matrix (for 2 5 i, j 5 m) are positive.&apos; For a binary solution 1-2, the condition for stability is: O(u=G22^0 [1] For a ternary system 1-2-3, the condition [I.] is re- tained (the value of G22 will differ, of course, according to the concentration of 3) and another condition is introduced: £>(21 = G22G33 - Gl3 ^ 0 [2] In a composition diagram, these two conditions define two domains of instability. Starting at a point where the solution is stable (for instance at a point where the solution is very dilute) we gradually change the composition until the condition [I] or [2] is violated. As already noted in the literature, e.g., in the work of Prigogine and Defay,3 it is the boundary of the domain (2) which is first crossed. For if we assume that the boundary of the domain (1) is reached first, at this point G22 = 0 and the second condition is necessarily violated (D(2) = -& 5 0), in contradiction with our original assumption. An exhaustive study of the ternary regular solution case may be found in the work of Meijering.4 Moreover if the boundaries of the two domains have a common point, they also have a common tangent. For if the two lines were to cross each other as is illustrated in Fig. 1(a) any point M in the line QP would be such that £> = 0 and 0"&apos; > 0 which, as shown above, are incompatible results. Thus, the two lines must be tangent at their common point Q as illustrated in the example of Fig. l(b). The reasoning of Fig. l(a) implies that the point Q is not a "singular" point for either boundary line. This singularity may be of two types. First, the lines meet without crossing each other and without being tangent. Second, the tangent at Q for D"&apos; or 0"&apos; is not single-valued. Other types of singularity are unlikely because of the usual analytical forms of D"&apos; and 0"&apos;. The exception to the common tangent requirement due to the first type of singularity was pointed out by John Morral;5 it occurs when the common point, Q&apos; or (3" in Fig. l(b), is located at a boundary of the composition diagram, e.g., at the line X3 = 0. It may also be noted that at the common nonsingular point Q of D(1) and D(2), Fig. 1(b), G23 is necessarily equal to zero, whereas at a point such as Q&apos; or Q", this conclusion is no longer valid because the product G22G33 is now indefinite (of the form 0. a). The exception to the common tangent requirement due to the second type of singularity occurs when two branches of the same boundary line intersect, for example when D(1)or D(2) decomposes into a product of functions, at a point which belongs to the boundary of the other condition. It is possible to show by a simple analytical calculation that, in this case, if Q is a singular point of D(1), then it is necessarily a singular point of D(2), and that the reciprocal is true except if G33 = 0 at Q. For the present article, however, more elaboration on these singularities appears to be unwarranted. To generalize the previous results to an m -component system, we use the mathematical theorem stating that if the diagonal determinant D(r) = 0, then

Jan 1, 1970

By Dr. O’Neil Thomas J., Donald W. Gentry

"If you can hold a board of cross directors In happiness against their gauzy schemes; If you can dodge the wrath of the electors Till dividends will flow as in their dreams; If you can make a mine pay from the grass roots No matter what the time or place or year; Then on my soul until the final blast shoots We&apos;ll add the title "MINING" when we call you "ENGINEER. " -From Wayne Darlington, "The Mining Engineer, " The Engineering and Mining Journal, Vol. 98, Sept. 26, 1914, p. 587. REASONS FOR THIS CHAPTER Up to this point in the book the material presented has addressed the fundamental parameters associated with the evaluation of mining ventures and other investment alternatives. The procedures and techniques illustrated represent the basic approach to project evaluation, particularly when considering the normal engineering economic analysis of investment proposals. The evaluation procedure is complicated, however, when dealing specifically with mining-related investment proposals. These complicating factors are invariably directly related to the many special characteristics of the mining industry. Many of these special features such as tax policies, revenue and cost estimation, and ore reserves have been addressed in previous chapters. In addition to the many special features of mining which directly impact, in a quantitative manner, the various line items in the pro forma income statements of a cash flow analysis, other features of mining affect evaluation procedures in a more general manner. The main purpose of this chapter, then, is to provide an exposition of some of these more qualitative factors. We hope this will enable the reader to recognize these factors in mine evaluation practice and to make suitable adjustments in his or her analysis.

Jan 1, 1984

By M. Metzger, L. E. Hendrickson

Grain boundary attack in 16 pct HCl was found to be substantially the same at low penetrations in zone-refined aluminum (individual impurities 0.1 at. ppm), superior electrolytically refined aluminum (51 at. ppm), and aluminum with various impurities at much higher levels. It was concluded that impurity atom segregation affecting corrosion would have been detected and that the corrosion susceptibility did not originate in this segregation but in the structure of the boundary. It was pointed out that the significance of most previous studies in this system had been obscured by an unrecognized autocatalytic copper reaction. Although general corrosion rate was also impurity-insensitive , there was shallow pitting attributed to iron seg-regation and a hillocked surface texture associated with copper; these were interpreted as due to cathodic damage affecting- cathode distribution. THE grain boundary corrosion suffered by high-purity aluminum in hydrochloric acid has been the object of some interest (the earlier work is summarized in Ref. 1). The central metallurgical question here is whether the corrosion susceptibility of a boundary originates in its structure or in impurity atom segregation. An attempt to study this question revealed large catalytic effects associated with small quantities of the copper impurity in the aluminum, 220 ppm, or in the corrodent, and these magnified the preferential boundary attack and obscured the intrinsic susceptibility question.&apos; After the catalytic effects had been examined,"&apos; test conditions could be designed to avoid them and methods developed for studying the shallow boundary penetrations prevailing when they were absent.= It then became possible to determine whether intrinsic boundary corrosion in aluminum involves impurity segregation. The older work provided no firm information on grain boundary segregation of specific solutes influencing corrosion although several studies suggested iron segregates.4,5 perryman4 found in 10 pct HC1 that the slowly developing (microns per month) grain boundary grooves were deeper in material of higher iron content (range 10 to 550 ppm, with 5 to 80 pprn Cu) but he did not measure the depths of general corrosion, which were probably several times greater, and his reference surfaces may have varied more than did the groove depths. Metzger and Intrater&apos;s results for 20 pct HC~,&apos; which yielded higher time-average rates (mm per month) and deeper penetrations, suggested that boundary segregation of iron (range 4 to 230 ppm, with 22 pprn Cu) decreased the penetration rate. However. in the stronger acid the autocatalytic l.E. HENDRICKSON, Student Member AIME ,and M. METZGER, Member AIME, are Research Assistant and Professor of Physical Metallurgy, respectively, Department of Mining, Metallurgy and Petroleum Engineering, University of Illinois, Urbana, Ill. Manuscript submitted March 11, 1968. IMD effect of the copper impurity is greater1,2 and it is now evident that their corrosion rates had been much magnified by this effect and did not provide a proper basis for the analysis of segregation. In exploratory studies of other solutes (made under the same conditions), 1000 ppm additions of Mg, Mn, or Si or 100 ppm Ca were without effect.1 Montariol6 noted that boundary attack in 22 pct HC1 persisted after zone refining although with fewer deep fissures (ranges 0.06 to 4 ppm Cu, 4 to 23 ppm Fe). Autocatalytic effects may have influenced these results also and those of Perryman.4 The present objective was, as a first step, to see whether quantitative tests designed to exclude autocatalytic influences would indicate the existence of low-level impurity effects on intrinsic boundary corrosion. Comparison of electrolytically refined with zone-refined aluminum of lower copper and iron contents revealed no differences in boundary corrosion, but certain impurity-sensitive differences in general corrosion morphology were noted and investigated further at higher impurity levels. I) EXPERIMENTAL PROCEDURE A) Material. A selection from commercially available material was made with the cooperation of several producers. An electrolytically refined lot (III-A, 1 ppm Cu, 2.4 ppm Fe) studied previously3 provided a starting point. Since material of substantially higher copper content could not be used if the autocatalytic corrosion reaction were to be avoided,2,3 a lot (III-B) of about the same purity was added as a check, two zone-refined lots (I-A and I-B) with copper and iron an order of magnitude lower were selected for comparison, and an intermediate lot (11) was included. Analytical data are given in Table I. For copper in I-B and iron in I-A and I-B, the actual concentration is thought to be near the limit given, i.e., about 0.1 ppm. For titanium, vanadium, and chromium in III-B, the actual amounts are thought to be, like those in III-A, substantially lower than in the zone-refined lots (these elements concentrate in the solid on freezing). Data are given later for some additional lots surveved. B) Specimen Preparation and Procedure. one by 3 cm blanks with a notched stem were cut from 1-mm cold-rolled sheet, annealed 24 hr in air at 650°C, and quenched in an air stream (42°C per sec initial cooling rate, 100 sec to cool to 100°C). The high annealing temperature maximized diffusivities and approach to equilibrium impurity distributions. A water quench, in principle more efficient in preserving the distribution established, was undesirable because the boundaries were almost plane and they tended to shear and migrate during the quench and thus to be separated from any existing impurity atmosphere. Test procedures, previously described,3 involved electropolish-ing, etching 2 min in 10 pct HF at 24.0°C, and expos-

Jan 1, 1969

By G. M. Ault, G. C. Deutsch

The paper reviews the efforts made to utilize powder metallurgy to solve problems encountered when using alloys at high temperatures. The following subjects are discussed: comparison of wrought and sintered super alloys, sintered aluminum powder, porous materials for transpiration cooling, molybdenum, and cermets. TODAY, powder-metallurgy techniques are be-JL coming increasingly important, producing materials which at the present time cannot be made by any other method. Among these materials are high melting-point metals such as tantalum and tungsten, porous metals for filters and lubricant-containing bearings, electrical contacts and electrodes (i.e., tungsten-copper), cutting tools of the cemented carbides, and magnet alloys. In addition, the powder-metallurgy industry has grown because, through the application of its methods, shaped parts, such as gears, can be produced at tremendous rates with a low scrap loss in comparison with manufacture by casting or forging and machining. In the past few years, by utilizing powder metallurgy, several approaches to the problem of high temperature materials have been investigated. In this paper the authors will review some of the studies on application of materials having properties above 1350°F. Since their experience lies in the aeronautical field, possible uses of powder-metallurgy techniques in the production of aircraft en- gines will be emphasized. A large portion of the work is classified, but such a restriction has not eliminated any field from discussion, although limiting the amount of detail it is possible to present. The development of high temperature alloys from powders and porous metals for cooling will be discussed as well as research done on molybdenum and molybdenum alloys, cermets, and intermetallics. Some studies on the utilization of cermets will be considered along with special problems that have been revealed through these investigations. Special interest in certain factors of sintered aluminum powder (SAP) makes it desirable to discuss this material although its use temperature is below the 1350°F limitation set for this paper. High Temperature Alloys from Powders In general, components made from powders for room-temperature application have been investigated with the primary purpose of taking advantage of the high production rates and economies made possible by powder metallurgy. From the standpoint of properties, the ojective of the powder metallurgist has been to obtain a product comparable to the cast or wrought product that is to be replaced. Comparisons of the properties of alloys made by powder-metallurgy methods with those of similar compositions made by casting and forging are reported in the literature.12 In many cases equivalent

Jan 1, 1955

By Walter E. Hopper

The Caddo oil-field, shown in Fig. 1, is located in Caddo parish, northwestern Louisiana. The known producing territory of oil is covered by townships 19 N, 20 N, 21 N, 22 N, and ranges 15 and 16 W., shown in Fig. 2. The center of the field may be taken as Oil City, 24 miles north of Shreveport, on the Kansas City Southern railway. Shreveport is the second largest city in the State, and is the connecting point of five railroad-lines. Drilling at the present time, however, is going on all over northern Louisiana, especially in Caddo and the neighboring parishes. During the winter of 1908-09 I spent &apos;four months in the Caddo field, under the direction of the Louisiana Geological Survey and the U. S. Geological Survey. II. History and Development. Natural gas escaping at the surface is found at numerous places in northwestern Louisiana. At Shreveport, the plant of the Shreveport Ice & Refrigerating Co. has been lighted by natural gas for 20 years. The well was drilled for water, but was abandoned on account of the gas. A test well, put down in 1905 near the western limits of Shreveport, was driven to 1,650 ft., and encountered indications of gas and oil at various depths, but did not succeed in finding enough to be profitable. Attention was first attracted to the Caddo field in 1895 by indications in water-wells from 40 to 60 ft. deep, in which the pressure of natural gas was noticeable. This indication of gas in a shallow well led to the drilling of the first well in the Caddo field, the old Savage Brothers & Morricell, or the Caddo Lake Oil and Pipe Line No. 1. The rig for this well was erected in May, 1904, and drilling began in June, 1904. The well was bailed Mar. 23, 1905, with a small amount of oil. It was deepened July, 1905, and converted into a "gasser" Jan. 3, 1906. It was abandoned January, 1907. In consequence of the finding of oil in the Savage well, drilling-operations were pushed with energy; and in April, 1905,

Jan 1, 1912

By L. D. Lamont, W. C. Muehlhof

Throughout the anthracite industry&apos;s history, the problem of handling and disposing of refuse material has been one of major importance. In the early days of the industry&apos;s activities, only the large sizes of anthracite were satisfactory for market requirements and the smaller sizes were deposited as refuse. As time went by, more efficient coal-burning equipment was developed and a market demand was created for the finer sizes of coal. This demand made it possible to recover marketable coal from banks deposited as refuse in the earlier days of mining. The same condition holds good today and we can see a definite trend toward the objective of utilizing the No. 4 and No. 5 buckwheat sizes. Under present conditions, we know that a limited market exists for No. 4 and 5 buckwheat; we know that the finer sizes of anthracite can be burned as pulverized fuel; we know that these fine sizes can be processed and sold as briquettes, and we also know that the heat content of these fine sizes of coal is comparable with that of the sizes now merchandized. Therefore, it should be the objective of the industry to conserve these fine sizes of coal in such manner that if and when processing facilities are developed, and market demands exist, the material can be salvaged and made available to meet these demands. From the standpoint of economical future reclamation, the ideal method of disposing of refuse containing the finer sizes of coal is to deposit it in banks separate and apart from other breaker refuse. By this procedure, the minimum amount of material will be processed in future salvage operations to recover the No. 4 and No. 5 buckwheat sizes. Many collieries in the anthracite region are operated under leases that have a specific term. These lease agreements provide that, upon expiration of the lease by term, exhaustion or otherwise, the ownership of the refuse material reverts to the lessor. Therefore, a question was raised in the minds of these operators: Why should I assume the expense of such procedure when I have no assurance that I will benefit by the future processing and sale of these fine sizes? Under present conditions, the answer to this question is not one of choice but of necessity, as restrictions imposed by legislation make it mandatory on operators of preparation plants to take necessary steps to properly treat silt-laden waters before permitting them to enter the streams. By reason of these restrictions, the disposal of fine sizes becomes a secondary rather than a primary operating consideration, and the procedure for disposing of this refuse becomes a matter of choice governed principally by topographical and geographical conditions. Methods of Handling Refuse There are three principal methods of handling and disposing of fine-size refuse; namely:

Jan 1, 1947

By W. C. Muehlhof, L. D. Lamont

Throughout the anthracite industry&apos;s history, the problem of handling and disposing of refuse material has been one of major importance. In the early days of the industry&apos;s activities, only the large sizes of anthracite were satisfactory for market requirements and the smaller sizes were deposited as refuse. As time went by, more efficient coal-burning equipment was developed and a market demand was created for the finer sizes of coal. This demand made it possible to recover marketable coal from banks deposited as refuse in the earlier days of mining. The same condition holds good today and we can see a definite trend toward the objective of utilizing the No. 4 and No. 5 buckwheat sizes. Under present conditions, we know that a limited market exists for No. 4 and 5 buckwheat; we know that the finer sizes of anthracite can be burned as pulverized fuel; we know that these fine sizes can be processed and sold as briquettes, and we also know that the heat content of these fine sizes of coal is comparable with that of the sizes now merchandized. Therefore, it should be the objective of the industry to conserve these fine sizes of coal in such manner that if and when processing facilities are developed, and market demands exist, the material can be salvaged and made available to meet these demands. From the standpoint of economical future reclamation, the ideal method of disposing of refuse containing the finer sizes of coal is to deposit it in banks separate and apart from other breaker refuse. By this procedure, the minimum amount of material will be processed in future salvage operations to recover the No. 4 and No. 5 buckwheat sizes. Many collieries in the anthracite region are operated under leases that have a specific term. These lease agreements provide that, upon expiration of the lease by term, exhaustion or otherwise, the ownership of the refuse material reverts to the lessor. Therefore, a question was raised in the minds of these operators: Why should I assume the expense of such procedure when I have no assurance that I will benefit by the future processing and sale of these fine sizes? Under present conditions, the answer to this question is not one of choice but of necessity, as restrictions imposed by legislation make it mandatory on operators of preparation plants to take necessary steps to properly treat silt-laden waters before permitting them to enter the streams. By reason of these restrictions, the disposal of fine sizes becomes a secondary rather than a primary operating consideration, and the procedure for disposing of this refuse becomes a matter of choice governed principally by topographical and geographical conditions. Methods of Handling Refuse There are three principal methods of handling and disposing of fine-size refuse; namely:

Jan 1, 1947

By Walter Crafts, C. M. Offenhauer

In a previous study of the carbide phase of chromium steels, it was shown that chromium carbide (Cr7C3) is 2 more stable carbide than cementite (Fe3C) at tempering temperatures above about 500°C. in quenched steels containing from about I to 7.5 per cent chromium. Below this temperature the carbide phase was found to consist of cementite in all of the steels examined. In the present investigation it was found that a similar change in the carbide phase takes place in molybdenum and in chromium-molybdenum steels at approximately the same tempering temperatures. The approximate ranges of occurrence of the alloy carbides Cr7C3, Cr4C, and Mo2C have been determined in steels containing up to 5 per cent chromium, I per cent molybdenum and 0.60 per cent carbon. The experimental steels were made in small high-frequency furnaces from Armco ingot iron, low-carbon ferrochromium and calcium molybdate and were poured into 2-in. ingots, which were forged to I-in. round bars. Sections of the bars were heated to the carbide-solution temperature and cooled rapidly to induce austenite transformation at relatively low temperatures. The steels were subject to some segregation, therefore comparatively high temperatures were necessary to dissolve all of the carbides. The microstructure of the quenched specimens was examined to verify the belief that the carbides had been dissolved, and also to determine the nature of the nonmartensitic structure of incompletely hardened steels. Ferrite and pseudo-martensite were present in some of the specimens, but all specimens were quenched at such a rate that pearlite was not formed, in order to ensure the formation of the carbides at temperatures lower than about 500°C. It had been found previously1 that in chromium steels the chromium carbide could be produced by transformation of austenite in the pearlite range above 500°C., and, in order to start the tempering treatment with the carbide present as cementite, the absence of pearlite was considered to be essential. The specimens were tempered for 64 hr. to permit the formation of the stable carbide. After tempering, the specimens were submitted to microscopic and X-ray examination. The carbides were separated either electrolytically in a 5 per cent HCl electrolyte or with an acidified copper-potassium chloride solution. The crystal structure of the carbide phases was determined by the Debye-Scherrer type of X-ray patterns. The X-ray patterns obtained from specimens near the boundaries between two types of carbides were generally weak and diffuse as compared with the other patterns. For this reason, and because the more soluble components may be masked, the boundaries between the phases could not be determined precisely. However, the X-ray analysis of electrolytically separated carbides has given consistently reproducible results

Jan 1, 1943

By Walter Crafts, C. M. Offenhauer

In a previous study of the carbide phase of chromium steels, it was shown that chromium carbide (Cr7C3) is 2 more stable carbide than cementite (Fe3C) at tempering temperatures above about 500°C. in quenched steels containing from about I to 7.5 per cent chromium. Below this temperature the carbide phase was found to consist of cementite in all of the steels examined. In the present investigation it was found that a similar change in the carbide phase takes place in molybdenum and in chromium-molybdenum steels at approximately the same tempering temperatures. The approximate ranges of occurrence of the alloy carbides Cr7C3, Cr4C, and Mo2C have been determined in steels containing up to 5 per cent chromium, I per cent molybdenum and 0.60 per cent carbon. The experimental steels were made in small high-frequency furnaces from Armco ingot iron, low-carbon ferrochromium and calcium molybdate and were poured into 2-in. ingots, which were forged to I-in. round bars. Sections of the bars were heated to the carbide-solution temperature and cooled rapidly to induce austenite transformation at relatively low temperatures. The steels were subject to some segregation, therefore comparatively high temperatures were necessary to dissolve all of the carbides. The microstructure of the quenched specimens was examined to verify the belief that the carbides had been dissolved, and also to determine the nature of the nonmartensitic structure of incompletely hardened steels. Ferrite and pseudo-martensite were present in some of the specimens, but all specimens were quenched at such a rate that pearlite was not formed, in order to ensure the formation of the carbides at temperatures lower than about 500°C. It had been found previously1 that in chromium steels the chromium carbide could be produced by transformation of austenite in the pearlite range above 500°C., and, in order to start the tempering treatment with the carbide present as cementite, the absence of pearlite was considered to be essential. The specimens were tempered for 64 hr. to permit the formation of the stable carbide. After tempering, the specimens were submitted to microscopic and X-ray examination. The carbides were separated either electrolytically in a 5 per cent HCl electrolyte or with an acidified copper-potassium chloride solution. The crystal structure of the carbide phases was determined by the Debye-Scherrer type of X-ray patterns. The X-ray patterns obtained from specimens near the boundaries between two types of carbides were generally weak and diffuse as compared with the other patterns. For this reason, and because the more soluble components may be masked, the boundaries between the phases could not be determined precisely. However, the X-ray analysis of electrolytically separated carbides has given consistently reproducible results

Jan 1, 1943

By A. T. Sweet, C. E. Plummer, H. W. St. Clair, S. F. Ravitz

The ammonium sulphate process for recovering alumina from clays was proposed by Rinman, Buchner, and others many years ago, and more recently various modifications have been investigated both here arid abroad.l-Is The process, as it is being investigated by the Bureau of Mines, involves the following essential steps: (I) sulphating the alumina in the clay by baking with ammonium sulphate at approximately 4m°C., (2) leaching with water to extract aluminum sulphate and ammonium sulphate, (3) crystallization of ammonium alum, (4) conversion of the alum to aluminum hydroxide by the ammonia gases given off during the baking, and (5) calcining the aluminum hydroxide to alumina. The over-all reaction between kaolinite (the most important alumina mineral in most clays) and ammonium sulphate during the baking step may be expressed as follows: Al2O8.2SiO2.2H2O + 4(NH4)2SO4 = (NH4)2SO4.A12(SO4)3 + 6NH3(g) + SH2O(g) + 2SiO2 [i] There is definite evidence. as is shown later. that the aluminum occurs in the baked product as anhydrous ammonium alum rather than as aluminum sulphate, provided the baking temperature is not much above 4o0°C. It is probable that the reaction takes place in the following steps: (NH4)2SO4 = NH3 -(- NH4HSO4 [2] Al2O3.2SiO2.2H2O + ;}NH4HSO4 = A12(SO4)3 + 3NH3 + 5H2O + 2SiO2 [3] A12(SO4)3 + (NH4)2SO4 = (NH4)2SO4.A12(SO4)3 [4] In onc modification of the process, the ammonium sulphate is first decomposed according to Eq. z, and the clay is treated with the resultant ammonium acid sulphate. Reaction I is highly endothermic. On the basis of the best available data, the following values have been estimated for the changes in heat content and free energy in calories per mol: -AH = 331,040 - 34T -AF = 331~040 - 78.3T log T - 671.8T The heat required&apos; at 25OC is 320,900 cal. per mol, or 5650 B.t.u. per pound of alumina. The theoretical equilibcium pressure of NHa, based on the free-energy equation given above, increases very rapidly as the temperature rises above 3o0°C., and reaches one atmosphere at 43s°C. Ammonium alum solutions are almost ideal for crystallization. A saturated solution of ammonium alum contains 126. grams of AlzOa per 1000 grams of water at 90°C. and only 16.5 at 20 C. Hence most of the alum crystallizes out when the

Jan 1, 1944

By Arthur F. Taggart

Tins paper proposes two units (believed to be new) for designating, respectively, capacity and efficiency for primary and intermediate crushers. CAPACITY Operators know that the tonnage of rock that can be put through a particular crusher in a give time depends primarily upon the crusher setting (open setting for jaw and primary gyratory types; closed setting for cones and rolls), and secondarily on the lithological character of the feed, its maximum size, the amount of fines that it carries, its moisture content, and the method of feeding. The proposed capacity formula states a relationship between all of these factors in one empirical equation, as follows: TR = TR80KK&apos;K" [I] in which Ts, called "reduction tons per hour" is the new capacity unit; T = tons of feed per hour requiring crushing; R80 = "80-per cent reduction ratio"; and K, K&apos;, and K" are constants that approximate quantitatively lithological character, moisture content and method of feeding, respectively. T is the hourly weight of material in the feed that is coarser than the coarsest discharge of the crusher. (For marked departures from the normal specific gravity range - 2.5 to 2.7 - T should. be adjusted to corresponding volumes.) In the crushers under consideration, discharge from the crushing zone is limited by the minimum dimension of a particle. Passage of the particle through a square-mesh screen is limited by its intermediate dimension. Sheppard1 has shown that the ratio of maximum: intermediate: minimum average dimensions of broken rock fragments may be taken as about 1.5 : 1 : 0.6 for nonslabby rocks such as are normally crushed, while for slabby rocks, such as thin-bedded sedimentaries, foliated schists and gneisses, and the blocky metamorphics, the ratio is of the average order I.3:I:0.3, with the possibility of an even smaller relative value for the thickness. Gaudin2 finds the corresponding ratio for 150 + 200-mesh pyrite to be 1.4: 1 : 0.6. The hourly tonnage of crusher-plant feed, therefore, should be diminished by the amount thereof that would pass a square-mesh screen of which the aperture is S times the crusher setting, where S is a shape factor with the value 1.7 for rocks that break exceptionally cubic; 2 for ores and rocks that exhibit no appreciable tendency to slabbiness; and 2.5 to 3, or, in exceptional cases, even larger, when the tendency to slabbiness is marked. R80 is the quotient of the theoretical square-mesh aperture that would pass 80 per cent of the feed, divided by the aperture that would pass 80 per cent of the product. It is best determined from screen tests.

Jan 1, 1941

By Albert C. Harding

For several years barite (BaSO4) production has topped such better known minerals as feldspar and fluorspar in annual tonnage and is now well established among our foremost nonmetallic industries. Description Of widespread occurrence in the United States, barite is relatively soft (hardness, 2.5 to 3.5), inert, heavy (specific gravity, 4.5), and is also called "barytes" (sometimes pronounced "bear-i-tease"), "heavy spar" and "tiff." It occurs with granular and earthy structure but usually has coarse, orthorhombic crystallization. It is commonly white, and although occurring in various shades of red, yellow, and gray, barite always gives a white streak and can be scratched by the ordinary prospecting pick. Its softness and weight are usually sufficient to establish field identification. Another aid is the "rotten egg" odor of H2S which sometimes emanates from freshly fractured surfaces. Barite can be readily distinguished from two other minerals which have somewhat similar occurrences and characteristics. The relatively rare celestite, strontium sulphate, has a specific gravity slightly less than 4.00 and gives a distinctively crimson color in a flame test. The very rare wither-ite, barium carbonate, has a specific gravity of 4.3 but readily effervesces with dilute hydrochloric acid. Occurrence All known commercially important barite deposits in the United States occur as re- placements or cavity fillings in schists, shales, limestones, dolomites, and quart-zites, or as residual deposits from such formations. In Georgia, Tennessee, North Carolina, and Missouri, high-grade barite is recovered as fragments from residual clays covering quartzites and dolomites. In Arkansas, Nevada, and California, barite, usually of lower grade, is mined from replacement deposits. These seven states contributed the entire production reported by the U. S. Bureau of Mines for 1945 as given in Table I. Table i—Crude Berile Sold or used by Producers in 1945 State Short Tons Arkansas................... 260.660 Missouri.................... 225.467 Georgia..................... 110,393 Tennessee................... 32,812 Nevada..................... 28,919 Other States............... 37,811 Total..................... 696,062 (&apos;other states,, included only california and North Carolina. Such a large part of this undistributed production came from California that that state probably outranked both Nevada and Tennessee. For the second consecutive year Arkansas was the leading producing state, particularly worthy of note because there was no commercial production of barite there prior to 1941 and because the entire production of Arkansas has come from one large deposit which is being operated by two companies. Consumption The 1945 consumption as reported by the Bureau of Mines was: 482,442 tons as ground barite 139,288 tons in making lithopone 99,173 tons in producing barium ------- chemicals Total 720,903 tons

Jan 1, 1949

By J. W. H. Clare

An equilibrium isotherm at 550°C is given for ternary alloys rich in aluminum containing 0 to 15 wt pct Cr and 0 to 20 wt pct Mn. Phases encountered are: aluminum solid solution; stable temary compound G; and MnAl8,. In addition, a limited number of alloys were annealed at 450°C to check for differences from the 550°C isotherm. These experiments limited the composition of the G phase at 550°C and further experiments established that the G Dhase is not stable above 590 ± 1 °C. RAYNOR and Little1 established several equilibrium isotherms using aluminum-rich alloys containing up to 5 pct of the elements chromium and manganese. These isotherms showed that, in addition to an a solid solution and the two binary intermediate compounds 8 and MnAl6, a stable ternary compound ex-isted in the system. This phase was called the G phase because of its ghostly appearance in micro-sections. Little, Raynor, and Hume-Rothery2 found a similar phase in binary aluminum-manganese alloys and, as a result of their experiments, concluded that the G phase was a metastable phase existing in the aluminum-manganese system. This phase was stabilized by the addition of chromium. The constitutional work of Raynor and Little,&apos; although showing the existence of the G phase in the aluminum-chromium-manganese system, did not establish precisely the composition limits of the phase or the temperature above which the phase became unstable and disappeared from the system. The work described in this paper was undertaken to provide this information. An examination of the previous work&apos; showed that an isotherm at 550°C would provide the required data regarding composition in the shortest possible time, as the (a + G) phase field narrows considerably at higher temperatures. The same work showed that the G phase ceased to exist between 580° and 595°C. A more precise estimate of the transformation temperature would therefore be obtained by annealing selected alloys at successively higher temperatures beginning at 580°C. EXPERIMENTAL PROCEDURE 1) Materials Used—The experimental alloys were prepared from four master alloys containing 15.00 w pct Cr, 15.28 wt pct Cr, 19.80 wt pct Mn, and 19.38 wt pct Mn which were made from super-purity aluminum, electrolytic chromium, and electrolytic manganese. The aluminum was of 99.996 pct purity, the principal impurities being copper, iron, and silicon. The chromium and manganese were 99.9 pct pure, the principal impurities being antimony, calcium, and magnesium. 2) Preparation of the Experimental Alloys—The alloys were prepared in 20-g quantities by melting together the requisite amounts of master alloy and aluminum in an alumina-lined crucible and, after thorough stirring with an alumina rod, cast into a cold copper mold to give ingots 1/4 in. in diam and 3 in. in length.

Jan 1, 1960