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By Jean-Jacques Prompsy, J. S. Rinehart

Some time ago a study was initiated at the Colorado School of Mines in an effort to arrive at a better understanding of the stress fields developed within drill bits under dynamic loading and the influence that these stress fields could have on the failure of the bits. Particular emphasis has been given to the influence that bit shape has on the establishment of highly localized, potentially destructive stress inhomogeneities within the bit. The study has been divided into three phases involving three velocity regimes: impacts at very low velocities, 0 to 20 ft/sec, a velocity range in which dynamic effects are just beginning to be found; impacts at velocities ranging from 20 to several hundred ft/sec, a range commonly encountered in practical drilling operations; and impulsive loading through the detonation of explosives, a region in which the dynamic effects are greatly exaggerated and made more identifiable. The ultimate objective is to provide basic data which will enable the mitigation through judicious design of the frequency and severity of these transient concentrations of intense stress, thereby prolonging and increasing drill efficiency. This paper presents results of the third phase of this study — the development of a better understanding of the dynamics and mechanics of failures caused by transient stress-wave interactions. The impulsive loading or transient waves in these experiments were generated by detonating small explosive charges placed on the ends of Plexiglas cylindrical rods terminating in truncated cones. Under such intense loading, dynamic effects are greatly exaggerated and, hence, made more identifiable. The explosion generates a high-intensity transient stress disturbance which moves through the rod, reflects from surfaces which lie in its path, and establishes momentarily narrow regions of high concentrations of stress where failure of fracture may occur. The peregrinations of these waves, their interactions and their effects have already been described in the literature (see, for example, Refs. 1 and 2). The locations and extent of the regions quite apparently are strongly dependent on specimen shape, making it possible through judicious shaping of specimens to relocate the highly stressed regions, minimize their extent, or eliminate them entirely. These experiments were designed, first, to identify more clearly the vulnerable regions in geometries relevant to drill-bit design and, then, to modify these regions in a predictable way by changing specimen shape. Through this latter process, it is anticipated that better bit design might evolve. The specimens used in the experiments were cylindrical rods, 1 1/2 and 1 1/8 in. in diameter, ended by a truncated cone, with the cone angles ranging from 45° to 130°. Upper cone face diameters were 1/2, 3/4, or 1 in. A small plastic-covered electric blasting cap detonated on the axis of the specimen produced the spherical shock wave. The detonators used were Olin Mathieson No. 6, which induced into the rod a transient saw-toothed stress disturbance of about 3-microseconds duration, corresponding to a wave length of 0.3 in. in the Plexiglas. Five distinct systems of fracture were observed: (1) along the axis of the rod, a lower system of fractures, the cracks focusing around one point for specimens having cone angles of 80° and 90° and, as the cone angle was increased, spreading downward; (2) just above these fractures, an upper system of fractures composed of horizontal cracks observable in some specimens but not of any appreciable extent; (3) a third system of axial cracks still above the latter, observable in the small cone-angle specimens; (4) in the upper part of the specimen, a system of radial fractures extending from the blasting point; and (5) circular spalling due to reflection of the wave on the boundary of the cylinder. Only the first three systems, the axial fractures, are examined here. LOWER TENSILE FRACTURE SYSTEM The first, lower system of tensile fractures, shown in the photograph of Fig. 1, was observed in almost all of the specimens. It was due to the reflected tensile wave coming in from the boundary of the cylinder as shown in Fig. 2. The reflected wave was cylindrically convergent and focused

By Margaret V. Doyle, R. W. Powers

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

Jan 1, 1958

By R. W. Heckel

A method is described whereby the relationship of both the "at-pressure" powder compact density and the "zero-pressure" compact density to the applied pressure may be obtained from continuous measurements of punch movement during a single compaction operation. The rate of density increase with applied pressure for the metal powders is found to be proportional to the volume fraction of pores for pressures exceeding a lower limit which varies from 15,000 to 30,000 psi, depending on the powder. 1 HE compaction of powdered materials is carried out primarily to increase the density of the material. If the ultimate goal of the over-all process is the attainment of minimum porosity, compaction is responsible for most of the densification. For example, loose powdered metals have porosities of about 65 to 75 pct. Axial loading in a die or hydrostatic pressure may effectively reduce the porosity of most powders to 20 pct or lower with pressures of 50,000 psi and greater. The relationship between the density of a powder compact and the pressure needed to achieve that density is most commonly obtained by pressing several compacts, each at a different pressure. Measurement of the densities of the individual compacts when they are removed from the die then provides the density-pressure relationship. However, this method suffers from several disadvantages. Because a large number of compacts and pressing operations are required, it is inherently slow. Furthermore, since density measurements are made after the specimens are removed from the die, data obtained refer only to pressures greater than those necessary to form a coherent compact. The third disadvantage is the inability of this method to provide information about the compact density at the applied pressure in the die. Because the determination of density-pressure relationships has been a slow, tedious process, investigators have sought to express the compaction behavior of powders by other means. The general terms "comparability" and L'compressibility" have been used to rank powders qualitatively according to their compaction characteristics. To restore a quantitative nature to the subject and to prevent confusion, Schwarzkopf1 has proposed that compac-tibility be defined as the minimum pressure needed to produce a given green strength, while compressibility should be used to indicate the extent to which the density of a powder is increased by a given pressure. Probably the most widely used of the compaction parameters is the "compression ratio,'' which is generally defined as the ratio of the compact density obtained by pressing at a given pressure to the apparent density of the loose powder.2 However, the description of the compaction behavior of a powder by these parameters provides only limited information about the process because of their applicability to just one specific condition such as a given green strength or a given pressure. The method of obtaining density-pressure relationships which will be described in the present paper was designed to eliminate the shortcomings of the previous techniques while maintaining equivalent accuracy and precision. EXPERIMENTAL TECHNIQUES The general principle employed makes use of the fact that the linear movement of the punch during a

Jan 1, 1962

By F. L. Fayers, J. W. Sheldon

It has been a matter of concern to the petroleum industry to determine what effect the capillary term has on saturation profiles, since these profiles determine ultimate economic oil recovery. In their original solution of the non-capillary two-phase flow problem, Buckley and Leverett&apos; showed that solutions of the two-phase flow equations became multiple-valued in saturation. Since it is physically unrealistic for saturation to have more than one value at a given position, Buckley and Leverett resolved this difficulty by introducing a saturation discontinuity or "shock". They evaluated the strength and position of the shock from material balance considerations. Recently Sheldon, Zondek and Cardwell&apos; have demonstrated that the method of characteristics can be used to solve this and related problems. Owing to capillarity it is physically unlikely that saturation shocks occur in an oil reservoir. Buckley and Leverett suggested that the effects of interfacial tensions should be included in the fluid flow equations, if the solutions are to give continuous single-valued saturation distributions. An analogy to this problem is encountered in the theory of wpersonic compressible flow where the introduction of the viscous term eliminates shock behavior. However, the mathematical form of the capillary term in the porous medium fluid flow equations differs from the viscosity term in the Navier-Stokes equation. This paper demons~rates that capillarity does eliminate the triple-valued behavior and indicates the magnitude of capillary and gravity effects on the saturation profile for a variety of systems. DERIVATION OF LAGRANGIAN AND EULERIAN FORMS OF THE DISPLACEMENT EQUATION The equations representing capillary, two-phase flow are given by applying Darcy&apos;s law and a material balance condition for each phase. The familiar Darcy and material balance equations are as follows. ^U*^0......<3, ^^r=°......c« S is the saturation of the water phase, and B the inclination of the porous medium. Pressures in the oil and water phases differ by the capillary pressure associated with the interfacial tensions, P.-P.=pAS)..........(5) where p, is determined experimentally. The porous medium is said to be oil-wet if P, > Po and water-wet if P, < Po. If the fluids are incompressible and the total volumetric flow rate, q, is constant, then v, f v,, = q...........(6) By eliminating P.. /»,, and v, from Eqs. 1, 2, 5 and 6 it follows that It is convenient to normalize the variables in the problem after the manner of Handy and Hadley% so that Eqs. 3 and 7 become where j,, = —, the fractional flow of water; (F, is a capillary pressure normalizing constant.) Substitution of Eq. 9 into Eq. 8 gives In this partial differential equation for an initial value

By J. H. Brophy

Experimental evidence in recent years shows that nickel coated hydrogen reduced tungsten powder can be sintered to 98 pct of theoretical density at 1100°C. New data indicate that the sintering rate is the same for nickel contents ranging from 0.125 wt pct (mon-atomic layer) to 5.0 wt pct Ni coated on 0.561. tungsten powder. Nickel tungsten made by coreduction from oxides sinters in a kinetically similar way, but the rate tends to increase with higher nickel contents. The activation of sintering can be accomplished with a minimum amount of nickel if coated powder is used. Transverse rmpture strength was found to increase in specimens containing 0.25 pct Ni as density increased during the initial stage of the carrier-phase sintering process. Upon the onset of grain growth in the final stage, strength was found to decrease. With increasing nickel contents up to 4 pct, strength increased. Maximum strengths in three-point loading were observed at 73,000 psi for 0.25 pct Ni and 93,000 psi for 5 pct Ni. These were comparable to that of massive tungsten recrystal-lized in the presence of nickel and tested in the same way. The results were sensitive to the mode of powder treatment and ductility was negligible in all cases. The study and application of sintered bodies of nickel and tungsten is not new; however, the analysis of the mechanism by which small additions of nickel increase the sintering rate of tungsten promises to Contribute new information regarding sintering theory in general. Initial experimental evidence was presented by the authors showing that nickel-coated hydrogen-reduced tungsten powder may be sintered to 98 pct of theoretical density at 1100°C in 16 hr.&apos; Simultaneously, a kinetic analysis of the process showed that it took place in two distinct stages. In the first stage, nickel evidently served as a carrier phase through which tungsten atoms could move. This stage was kinetically similar to the &apos;solution-precipitationo step postulated when the carrier phase is liquid3 although in the nickel tungsten case, there is no evidence of the existence of a liquid phase either in equilibrium phase relationships or in mi-crostructures at the temperatures under consideration. In the present work, new data for nickel-coated tungsten powder are offered in support of the proposed mechanism, and a more explicit study of the second stage of densification is presented. Concurrently, the sintering kinetics of these coated powders are compared to Ni-W powder made by coreduction of oxides. In this way the present work can be compared to earlier results in nickel activated sintering of tungsten in which the coreduction process was emplyed. The ultimate utility of such an activated sintering process will be determined to some extent by the mechanical and physical properties of the final prod-

Jan 1, 1962

Jan 1, 1951

By R. L. Orr, O. D. Sherby, J. E. Dorn

Creep data for pure metals at temperatures above those at which rapid recovery occurs (above about 0.45 the melting temperature) are correlatable by means of the equations and These correlations were applied successfully to data for aluminum, iron, nickel, copper, zinc, platinum, gold, and lead as well as for simple alloys. For a given metal, AH is a constant about equal to the activation energy for self-diffusion. THE early recognition that creep is stimulated by thermal activation prompted numerous investigators&apos;, &apos; to apply the same laws that are valid for the viscous behavior of liquids to analyses of creep data. To a good first approximation these laws are summarized by the equation were i is the creep rate; T, the absolute temperature; R, the gas constant in cal per degree per mol; T, the applied stress; AH, the activation energy in cal per mol; S&apos;, a constant dependent on the entropy of activation, the frequency of activation, and the contribution of each activation to the strain; and B&apos;, a constant dependent upon the size of the flow unit that is activated. Although Eq. 1 describes the flow of viscous fluids very accurately, its application to the creep of solids has been disappointing. Two reasons for the failure of Eq. 1 for creep are now known: First, the decelerating creep rates during primary creep demand that the internal structures of the metals are changing and that these changes might be reflected by changes in one or more of the three creep parameters S&apos;, AH, and B&apos;. Consequently when the conventional methods of evaluating these parameters are employed by comparing the secondary creep rates at a series of temperatures and at a series of stresses, the true effects of temperature and stress are masked because of simultaneous changes in one or more of the creep parameters. Second, as will be illustrated more fully later, three distinctly different types of investigations have shown that the stress term for high temperature creep enters the analysis as and not as t/T. This rather unexpected result points sharply to a significant fundamental difference between the mechanisms for the viscous behavior of fluids and the high temperature creep of metals. More recent extensive investigations",&apos;4 on the creep of high purity aluminum and several of its dilute a solid solutions have shown that where E is the total creep strain; temperature compensated time = for a constant temperature test; t, is the initial stress in a constant load test or the constant true stress in a constant stress test;" and t is the duration of test. • Eq. 2 is valid for either constant stress or constant load tests; but the total creep curve is difierent (i.e. the function, f, is differentl for each test. The fundamental origin of the validity of Eq. 2 is thought to arise from some equivalence of substructures in constant load creep tests at equal values of 0 or E. Both X-ray and metallographic examination revealed that practically identical lattice distortions and subgrain sizes are obtained for the same creep strain under the same load over wide ranges of test temperature.&apos; Furthermore the room temperature tensile properties following equal creep strains at different temperatures under constant loads were also identical, illustrating that identical substructures were indeed obtained.&apos; " A second type of correlation was obtained by differentiating Eq. 2 with respect to time, giving For the secondary creep rate, denoted by i,, But Eq. 2 reveals that 0, is a function of a, alone because the secondary creep rate in a given constant load test is always reached at a fixed value E,". Therefore where is the function that was introduced a number of years ago by Zener and Hollomon." Eq. 4 is not only valid for creep but will permit correlations between constant load creep and tensile data where e, is the rate of tensile straining and (TC is the engineering ultimate tensile strength.

Jan 1, 1955

By W. W. Mullins

The development, by the mechanism of volume diffusion, of a grain boundary groove on an interface separating a solid phase and a saturated fluid phase is calculated under the following assumptions: 1) isotropy of the interfacial free energy &apos;s, 2) applicability of the Gibbs-Thompson formula relating curvature and chemical potential, 3) a nearly planar groove surface. The groove profile is found to have a .fixed shape and linear dimensions that are proportional to (time)1/3. The constant of proportionality is evaluated and involves ?s in such a way that the latter my be evaluated from grooving kinetics. A grain boundary that intersects an interphase interface will tend to produce a groove along the line of intersection (Fig. 1). The ultimate motivation for the formation of the groove is the reduction in interfacial free energ that occurs as the grain boundary contracts. Smith has discussed the way in which the free energy of the interfaces determines their dihedral angles at the groove root and has given a detailed analysis of the effect of the dihedral angles upon the nature the microstructures developed in polyphase alloys. Herring2 has extended the considerations to the case in which the free energy of an interface depends on its orientation and has thus added torque terms to the condition for interface equilibrium. Although the grain boundary and the interphase interface may quickly achieve the proper equilibrium angles where they intersect, as determined by the balance of equivalent tensions and torques,"2 the groove will generally continue to grow (though with decreasing speed). Thermodynamically, this continued growth reflects the fact that the lowest free energy cannot be achieved until the entire grain boundary has been eliminated; mechanistically, the growth occurs in response to inhomogeneities of chemical potential that are caused by the constraint of fixed dihedral angles at the groove root. The latter point is illustrated in Fig. 1, which shows that atoms* near the groove are on the convex side of a cylindrically curved surface and therefore have a higher chemical potential than atoms on the flat interface beyond. Thus atoms tend to leave the curved groove shoulders and force the groove to deepen. When both phases are solid additional chemical potential gradient may arise from whatever strain fields are necessary to insure continuity of material across the interface. In the discussion that follows, this complication due to strain will not occur since we always assume one phase to be a fluid and the other a solid. The transport processes that may operate to permit growth of the groove are surface (interface)* diffusion, volume diffusion in the phase on either side of the interface, and evaporation-condensation when one phase is gaseous. (Also a melting-freezing process is responsible for groove enlargement on a solid-melt interface as discussed in Sec. VI). The theory of grain boundary grooving has been discussed in detail for the processes of surface diffusion and of one type of evaporation-condensation.3&apos;4 The theoretical predictions for the mechanism of surface diffusion have been confirmed in an experimental study of grooving kinetics in copper.5 In this paper, the theory of groove development by volume diffusion will be presented. Volume diffusion can occur either by atomic migration in the solid S or by diffusion of S atoms in the fluid phase, or solvent, adjacent to the groove surface. As discussed in Sec. V, volume diffusion in the solid should be less important than surface diffusion in forming a groove. Apart from verifying this statement then, the main interest of the present analysis centers on diffusion in an external solvent. The solvent may be either a liquid (e.g., water, or-

Jan 1, 1961

By H. R. Hose

The total crude fluorspar reserves in Western Europe, including the USSR and the Soviet sphere, Thetotalare estimated to exceed 17 million metric tons. The total fluorspar production in 1951 in Western Europe, excluding the USSR and Soviet sphere, amounted to 371,109 metric tons. The chief fluorspar-producing countries in Europe, excluding Russia and the Soviet sphere, are France, Western Germany, Italy, Spain, and the United Kingdom. TOTAL crude fluorspar reserves in western Europe, containing more than 35 pct CaF2, are estimated at 12 million metric tons, while reserves in the USSR and Soviet sphere may exceed 5 million metric tons. The bulk of these reserves are concentrated in the USSR, Spain, United Kingdom, Western Germany, and Italy. Since published estimates of the tonnages of individual mines and explored deposits are not available for most countries in Europe, the estimates given in Table I for individual countries are necessarily approximate, as they have been compiled from various sources and modified and combined according to the author&apos;s judgment. The majority of fluorspar deposits in Europe belong to the late Paleozoic metallogenetic epochs and have been found within or along the borders of the Hercynian massifs. Deposits in the Central and Esterel massifs of France, those of Western and Eastern Germany, Sardinia, the Catalonian coastal range and the Sierra de los Santos in Spain, and those of Devonshire and Cornwall belong to the Hercynian metallogenetic epochs. The deposits in southern Norway and Sweden belong to the middle Paleozoic or later, those of northern England to the late Paleozoic or later, those of northern Spain and Greece to the late Mesozoic, and those of northern and probably central Italy and Switzerland to the early Tertiary. Fluorspar occurs typically in veins, lenses, and stringers, but also in bedded replacement deposits and pockets, and may be colorless to white, yellow, green, pink, brown, blue, or purple to almost black. Except for the colorless variety, the only kind which can be used for the optical industry, the color of the fluorspar does not affect its ultimate use. Crystalline cubic or octahedral fluorite of CaF2 composition, with 48.9 pct fluorine and 51.1 pct calcium, has a specific gravity of 3.18. The ore in place, however, usually averages 3.0 metric tons per cubic meter owing to the quantity of other minerals asso- ciated with the fluorspar in the veins and the physical state of the fluorspar, which may vary from sugary to massive crystalline or microscopic inter-growths with barite or quartz. Barite of 4.5 sp gr, quartz of 2.66 sp gr, and chert and calcite are commonly associated with fluorspar. Other minerals usually present in lesser amounts are celestite, wad, silver, lead, zinc, copper, iron, and uranium minerals. In some cavity fillings the rhythmic deposition of the fluorspar, barite, and quartz into successive crusts may make it possible to handpick the crude ore on belts or to separate the fluorspar from other minerals while the actual mining is carried on. Where metallurgical grade lump spar is required for foundries, separation of quartz and barite may be made also by jigs and heavy media, and in some mines the lead and zinc sulphides may be separated and recovered by vibrating tables. However, except in the case of unusually high grade deposits, acid grade fluorspar is obtained by flotation. By this process fluorspar deposits grading down to 35 pct CaF2 in the crude ore can be worked at a profit when fluorspar, lead, zinc, and silver prices are high, as the sulphides of these ores can be separated from the fluorspar by flotation and concentrated.

Jan 1, 1956

By D. R. Mash

A strong yield point with attendant enhanced mechanical properties was found in commercially pure beryllium under certain conditions of heat treatment. Beryllium specimens also responded to both quench and strain aging treatments. Based on these results, a possible means of obtaining increased normal temperature mechanical properties is suggested. COMMERCIALLY pure beryllium has been found to exhibit the yield point phenomenon and several other aging manifestations under certain conditions of heat treatment. This behavior is not surprising in view of the present accumulation of experience regarding such effects in numerous metals; but in addition to extending this experience to include beryllium, the present study is of interest because aging effects are generally associated with the presence of small amounts of solute elements. The well known lack of tensile ductility in beryllium at normal temperatures is also considered to be due to impurities in some way. In the present investigation, the view was taken that beryllium of commercial purity (greater than 98 pct) should be considered as a dilute alloy and that precipitation phenomena may well exist which could lead to enhanced mechanical properties. In connection with this possibility, published data on the variation of mechanical properties with temperature are of interest. For example, Fig. 1 illustrates the variation of ductility (percentage of elongation) and ultimate strength as a function of temperature for beryllium of various histories.* Although these curves are somewhat schematic in nature, they clearly indicate the possibility that precipitation of an unstable phase and/or aging effects occur during tests in the region 400" to 700°C. The existence of such behavior should be recognized as affecting the properties of the commercial material and perhaps can be employed advantageously. Experimental Method All testing was done at room temperature using a Baldwin universal testing machine. A microformer extensometer was used in conjunction with a micro-former-type stress-strain recorder. All tests were conducted at constant loading rate, the slowest rates being used. This corresponded to 120 lb per min in the 1200 lb full scale range and 300 lb per min in the 6000 lb full scale range. After the rate of loading was set in the elastic range of the material during the early part of the test, no further adjustments were made, allowing the test to proceed at that setting. In the plastic region of the curve above the yield region, strain rates of from 0.002 to 0.006 in. per min were regularly obtained. All experimentation was carried out on specimens machined from the same lot of beryllium, lot No. Y5036BP. Table I shows the chemical analysis as furnished by the supplier, Brush Beryllium Co., as well as the standard specifications to which commercial beryllium is produced. Both 1/4 in. thick hot pressed plate and ½ in. round hot pressed and warm extruded rods were obtained. The history of the material is reported to be as follows: 1—Beryllium metal plates, produced from a standard hot pressed QMV beryllium metal block 4x15x40 in., were cut off and finished machined to 1/4x6x12 in. flat plates. The original hot pressed block was made by hot pressing —200 mesh vacuum cast attri-tioned powder at about 1050°C under vacuum. 2—Beryllium rod was hot pressed and warm extruded. Extrusion billets were machined from the regular hot pressed blocks followed by extrusion at 425°C. The extrusion was done without jacketing, and graphite was used as a lubricant. The rods were then straightened at 720 °C, after which they were annealed at 800°C for 15 min. Both flat tensile specimens and round tensile specimens were tested. The former were machined from the plate stock into standard specimens according to specifications shown in ASTM E8-52T. The nominal cross-section was 0.25x0.500 in. and the gage length was 2 in. The round specimens were

Jan 1, 1956

By Benjamin Petkof

The mineral mica, which has been known to man since ancient times, has played an impor¬tant role in the development of our modern industry. In the latter part of the 19th century sheet mica began finding use as a dielectric insulating material in the newly developing electrical industry. Later, when modern elec¬tronics began to develop, it became a vital insulating and dielectric material for use in electronic equipment. Both the electrical and electronic industries were greatly dependent on some form of natural mica. The manufacture of electric motors of varying sizes, electronic tubes, and other electronic items was dependent on an adequate supply of sheet mica. During World War II this material was so vital to national defense that special efforts were made to mine it in the United States as well as to obtain it from overseas mica-producing areas. However, advancing technology has successfully begun the progressive development of man-made mica-based and non-mica-based materials that can replace mica for many uses. In addition, developing technology has dictated changes that have made sheet mica unnecessary for some uses. As the demand for sheet mica declined, the demand for scrap and flake mica has continued to increase. In fact, today the domestic production and use of scrap and flake mica far outshadows the necessity for good¬quality sheet mica in this country. Historically the United States has always been dependent on foreign sources of material for good-quality sheet mica and imports practically its entire requirement. In the case of scrap and flake mica, the United States is the dominant world producer and consumer. End Uses Sheet mica (Chowdhury, 1941; Rajgarhia, 1951; Skow, 1962) is available in many com¬mercial forms, but these forms can be broadly classified into manufactured and unmanufac¬tured mica. Manufactured mica consists of mica that has been shaped, punched, or other¬wise processed into some form, suitable for a particular end use. Unmanufactured mica con¬sists of partially hand-trimmed or processed material that has not been prepared for any particular end use. Unmanufactured mica is broken into two commercial classes consisting of sheet mica and scrap and flake mica. These classes of mica differ greatly in their ultimate end uses and marketed forms. Sheet muscovite mica can be classified by color, degree of prep¬aration of the crude material, thickness, size, visual appearance, and electrical quality. In addition, phlogopite mica is also classified by its thermal stability. Scrap and flake mica consist of mine, trimming shop, and factory scrap that occurs as remnants from mining, processing, and manufacturing operations, and small particle size mica that is available from the beneficiation of pegmatites, clays, schists, or other mica-rich host rock. The end uses for sheet mica vary so greatly from those of scrap and flake mica that it is necessary to discuss them separately. Sheet Mica Sheet mica consists of relatively flat sheets of material which have been mined as naturally occurring books or runs of mica. These sheets are hand-trimmed to remove any imperfections and are punched or stamped into specified shapes for industrial use. Sheet mica can be further processed and is described or specified as block, film, or splittings based on thickness. Block mica is not less than 0.007 in. thick with a minimum usable cross-sectional area of 1 sq in. Film mica is split from the better quali¬ties of block mica to specified thickness groups ranging from 0.0012 to 0.004 in. Splittings are sheets of mica with a maximum thickness of 0.0012 in. and a minimum usable area of 0.75 in. Other small sized, lower quality block mica

Jan 1, 1975

By G. S. Cole, G. F. Bolling

ThE exact type of fluid flow which occurs in a solidifying ingot is important in determining subsequent grain structure. This has been shown in studies of natural Convection" and of forced stirring or other disturbances of the melt."16 We shall explicitly argue that a complement to the existing natural flow in an ingot should be more effective than other types of flow in promoting the columnar-to-equiaxed transition. The ultimate effects of fluid flow upon structure is one of the topics of solidification that we have investigated in the alloy system Pb-Sb, under conditions simulating casting. In Fig. 1, the essential features of a bidirectional growth apparatus are shown; horizontal freezing from water-cooled faces within a low-thermal-conductivity lavite boat is a reasonable simulation of an ingot section. Details such as the furnace and thermocouples are omitted in order to show the location of four permanent. magnets, each with a pole separation of 21 in. providing a central field of about 1000 Gauss. When coupled with a dc current imposed through appropriate connections made at the cooled faces, a flow of fluid is produced as shown schematically. This technique described by Pfann and ~orsi" is ideal for our purpose since the flow indicated is the same as that we observe to occur naturally in this system. Components of 99.999 pct purity were used for study and all due precautions were taken in alloy preparation to avoid contaminants. Sections of the specimens were taken to ensure results representative of the interior as well as the surface of specimens. All proportions of equiaxed and columnar regions were measured by tracing the etched grain structures onto graph paper and simply counting squares. Four alloys were examined at superheats of 40" * 5°C and 100" * 9°C. Preliminary tests had shown that no appreciable enhancement of structure was obtained with magnetic field alone, so that this condition is not reported; it should nevertheless be noted that other configurations of magnetic field will reduce convection and thus enhance columnar growth.&apos;8 For the tests reported the magnetic field was applied to a specimen by bringing the four preassembled magnets to a position 2 mm from an outer tube immediately after the ingot was poured. The dc current of 30 amps passed through the sample throughout the entire solidification cycle, resulting in a power dissipation of 9 w per sq cm of interface. This may be compared to the evolution of latent heat of approximately 18 w per sq cm. Table I contains the results of this experiment in terms of the total area ratio, equiaxed/columnar grains, as a function of the field interactions, solute concentrations, and superheats employed. The total structure of the ingot is seen to be most influenced when both the field and current are applied simultaneously; when current acts separately, little effect may be noted in spite of the heat supplied. In addition, the change in structure from the no field-no current case to the field-current condition is more pronounced at the higher superheat, and the higher the concentration of antimony. An example of the grain refinement, as well as the increased equiaxed zone, is shown in Fig. 2. It is now important to prove that there is indeed an increase in fluid flow due to the presence of the field-current interaction; this is done by temperature measurements in the liquid metal. Let us consider a bidirectional casting in which heat is applied to and removed from the system, in such a way that rate of solidification is zero.* When a sensitive thermocouple

Jan 1, 1967

By H. S. Martin

ALTHOUGH it was known some years ago at the- Utah Copper Co. mills that fine grinding improved flotation recoveries, no accurate data were available until recently as to just how far the grinding could be carried economically. In 1923 the ultimate degree of fineness for flotation feed was supposed to be about 10 per cent. on 65 mesh, with approximately 55 per cent. through 200 mesh, which meant that most of the mineral would pass 100 mesh. At this time Garfield and Wilfley tables were used ahead of flotation and a good deal of the coarse mineral removed prior to the ball mill regrind. Shortly afterwards, mill-scale tests developed the fact that on a feed containing not over 10 per cent. +65¬mesh material flotation alone would give a copper recovery equal to that obtained by the combination of tables and flotation, but that on much coarser feed the recovery with flotation alone was lowered. This was at a time when all the sulfide minerals were floated if possible, so that the only requisite was to grind fine enough to free the mineral from gangue. EARLIEST USE op MICROSCOPIC EXAMINATION OF TAILINGS As early as 1921, the writer had attempted to determine by chemical analysis the relative proportions of copper minerals in tailings from an ore that made extremely poor flotation results. Chemical analysis was not satisfactory or conclusive, and samples of these tailings were sent to the U. S. Bureau of Mines Station at the University of Utah for micro-scopic examination. The question at that time was to determine, if possible, which copper minerals had been lost and why, but R. E. Head, in his report, revealed other interesting data. He said in part: Briefly stated, the mineral particles or grains may be classified into separate groups as follows: 1. Free grains of pyrite, chalcopyrite, chalcocite and bornite, respectively, which contain no other included sulfides; 2. Grains in which all three of the copper sulfides are represented (the core may he of chalcopyrite, partially altered to bornite, these two minerals being entirely surrounded by a shell of chalcocite); 3. Grains composed of equal amounts of bornite and chalcocite or chalcopyrite and chalcocite.

Jan 1, 1929

Jan 1, 1951

By F. F. Peng, A. D. Walters, M. R. Geer, J. W. Leonard

INTRODUCTION Washability characteristics derived from float-and-sink analysis in Chapter 4 demonstrate that coal is intrinsically heterogeneous, and this heterogeneity is also evident in the proximate, ultimate, and size analyses. A range of particles from high to low specific gravity is usually seen in the run-of-mine coal. One objective of coal preparation is to recover the high quality float material (the particles of lower specific gravity) such that the recovered product meets certain quality specifications, as well as to reject or recycle low quality particles of higher specific gravity. The maximum recovery of cleaned coal at the required quality is of major concern to plant operators. However, the quantity and quality of products of coal cleaning, determined in the operation of actual commercial cleaning plants and by theoretical calculation based on run-of-mine washability studies, are known to be different with each coal. Therefore, during the last half century, coal preparation engineers have repeatedly sought methods to predict the technical and economical efficiency of separation of cleaning equipment and systems. Many formulas have been proposed by international investigators to evaluate the efficiency of various coal cleaning operations, primarily based on the yield and/or ash of the various products of coal cleaning. Performance criteria in use today are capable of defining all of the important aspects of a given concentrating operation; however each criterion has certain restrictions or limitations. It is important, therefore, to understand how such formulas can be used and the problems involved in their construction. Technical Measure of Plant Performance Conventional means for evaluating a preparation plant&apos;s performance are generally approached through the use of equipment performance mea-

Jan 1, 1979

By John Johnston

When I was honored by being invited to give the Howe Memorial Lecture, I decided to read Howe&apos;s book, "The Metallography of Steel and Cast Iron," published in I9I6—that is, about 25 years ago—in search of a text. I found the book written in fine, clear, attractive, yet precise style, as I expected, knowing that his father— Dr. Samuel Gridley Howe, his mother— Julia Ward Howe, and all of his four brothers and sisters had lived in a New England literary atmosphere and had written books. This leads me to suggest that many authors of metallurgical papers and books could well study Howe&apos;s writings with a view to the improvement of their own. I failed to find a specific text, but did find a subject, for in reading his book I was struck by the few instances in which time is considered explicitly as a factor in the phenomena examined, and by the fact that many of the hottest arguments would have been settled had time been taken properly into account. Indeed, Howe hardly mentions time except in a rather incidental way, as in a few references to lag; for instance: The lowering of temperature at which the transformation occurs by rapid cooling . . . is a phenomenon of undercooling or lag. It is in part explained by the fact that the transformation itself is a time-consuming reaction which, setting in during a rapid fall of temperature, is naturally protracted till the temperature has sunk far below that at which it is due. He does discuss, quite fully, the iron-clrbon equilibrium diagram, as known at that time, introducing it by stating: As he is a reasonable man who is not deterred from using a hand truck for moving his trunk by the consideration that he has thus to move a trunk plus a truck, so am I reasonable in putting into your hands this useful, indeed indispensable. tool for mastering the ABC of iron metallurgy. Somewhat later, in introducing a discussion of the phase rule, which "it would be hardly proper to pass ... by without an attempt to outline its meaning," for "a clear exposition of its applicability or jurisdiction is needed to restrain the half-initiated from misusing it and the uninitiated from being misled by these misuses," he writes—and I quote his statement because we cannot properly discuss the influence of time on a process until we know the state of equilibrium which would ultimately be reached by the system under consideration: It is a most remarkable and valuable generalization, its conceptions help greatly toward getting a broad outlook on metallography, but its misconception has brought out a flood of obscuring writings. it tells us about the constitution towards which alloys tend, that which they reach when equilibrium is complete, when all tendencies have corrected themselves and have been complied with completely. Its application, in absence of true equilibrium, is like determining density with a telescope or even with a tuning fork; it is like the mad tea party at which the butter put into the works of the watch was the best butter. If the reader learns no more than to beware of attempts to dccipher the conditions of inequilibrium by a law which touches only the conditions of exact

Jan 1, 1942

By W. T. Turrall, M. J. Cook

A discussion of modern developments in beneficiation of fine sizes of anthracite, this paper includes a description of the plant flowsheet, an analysis of operating results, and a summary of fundamentals of mineral separation relative to plant operation. DURING the past several years Lehigh Naviga-tion Coal Co. has installed equipment at Tamaqua and Lansford collieries for cleaning and sizing No. 4 buckwheat —3/32 + 3/64 in. and for recovering and cleaning flotation coal —20 + 200 mesh at Tamaqua colliery. Papers describing some of these installations have been presented to AIME. At Coaldale colliery, however, methods used to clean fine coal were inadequate and inefficient. Management decided upon a thorough investigation of all existing processes before building a plant to treat the combined three sizes mentioned. Investigation, carried out on company property and at other plants, entailed a study of hydrotators, hydroclassifiers, tables, spirals, flotation machines, methods of screening, and methods of dewatering. On the basis of this investigation a flowsheet was designed which necessarily excluded many processes in use, and the writers wish to emphasize that no criticism is implied by omission of processes or machines. A determining factor in selection of many machines was the policy of standardizing equipment wherever possible. Flowsheet and Operational Data The first principle to be considered in designing of any plant is efficiency of cleaning as related to laws of classification. Beneficiation of coal may be described as separation of two materials having different specific gravities. Although there is an inherent ash content even in the purest coal which slightly affects its specific gravity, it is recognized that many particles are true middlings. The percentage of these particles allowable in the final product depends upon the ultimate ash, or specifications of the consumer. Richards in his study of free settling and hindered settling classification&apos; determined that particles of equal size and different specific gravities had unequal settling velocities. He also determined a definite size relationship for equal settling velocities with free settling and hindered settling conditions. Considering the design of machines used to clean No. 4 and No. 5 buckwheat, it may be assumed that both free settling and hindered settling conditions exist and that ratio of size for practical operations is between the established values. The essential factor is that the more efficiently and closely sized the feed is maintained, the more nearly perfect the resulting separation. Flowsheet design was governed primarily by factors related to the 1200 to 1500 tons per hr feed to the main plant. It should be stressed that because of the many sources of supply from mine and stripping, the physical characteristics of this feed varied considerably. The most significant influence was varying percentage of fines, necessitating that all machines be capable of operating under maximum peak load during periods of surge. Since cleaning practice is a wet process, about 9000 gal per min containing 14 to 15 pct solids, or 350 tons per hr, must be treated. This slurry contained all the fines, or —3/32 in., but also percentages of oversize, up to 1 % in., dependent on spills or breaks in screen jackets, a not uncommon occurrence in preparation plant practice. The flowsheet developed is shown in Fig. 1. Circled numbers represent points of sampling. Table I gives operational data including size, ash by size, composite ash, rate of flow in gallons per minute, percent solids dry weight, and average tons of solids per hour. It should be noted that these results are subject to human error in sampling practice. Detailed information reported is a representative evaluation of plant operation.

Jan 1, 1954

By John Johnston

When I was honored by being invited to give the Howe Memorial Lecture, I decided to read Howe&apos;s book, "The Metallography of Steel and Cast Iron," published in I9I6—that is, about 25 years ago—in search of a text. I found the book written in fine, clear, attractive, yet precise style, as I expected, knowing that his father— Dr. Samuel Gridley Howe, his mother— Julia Ward Howe, and all of his four brothers and sisters had lived in a New England literary atmosphere and had written books. This leads me to suggest that many authors of metallurgical papers and books could well study Howe&apos;s writings with a view to the improvement of their own. I failed to find a specific text, but did find a subject, for in reading his book I was struck by the few instances in which time is considered explicitly as a factor in the phenomena examined, and by the fact that many of the hottest arguments would have been settled had time been taken properly into account. Indeed, Howe hardly mentions time except in a rather incidental way, as in a few references to lag; for instance: The lowering of temperature at which the transformation occurs by rapid cooling . . . is a phenomenon of undercooling or lag. It is in part explained by the fact that the transformation itself is a time-consuming reaction which, setting in during a rapid fall of temperature, is naturally protracted till the temperature has sunk far below that at which it is due. He does discuss, quite fully, the iron-clrbon equilibrium diagram, as known at that time, introducing it by stating: As he is a reasonable man who is not deterred from using a hand truck for moving his trunk by the consideration that he has thus to move a trunk plus a truck, so am I reasonable in putting into your hands this useful, indeed indispensable. tool for mastering the ABC of iron metallurgy. Somewhat later, in introducing a discussion of the phase rule, which "it would be hardly proper to pass ... by without an attempt to outline its meaning," for "a clear exposition of its applicability or jurisdiction is needed to restrain the half-initiated from misusing it and the uninitiated from being misled by these misuses," he writes—and I quote his statement because we cannot properly discuss the influence of time on a process until we know the state of equilibrium which would ultimately be reached by the system under consideration: It is a most remarkable and valuable generalization, its conceptions help greatly toward getting a broad outlook on metallography, but its misconception has brought out a flood of obscuring writings. it tells us about the constitution towards which alloys tend, that which they reach when equilibrium is complete, when all tendencies have corrected themselves and have been complied with completely. Its application, in absence of true equilibrium, is like determining density with a telescope or even with a tuning fork; it is like the mad tea party at which the butter put into the works of the watch was the best butter. If the reader learns no more than to beware of attempts to dccipher the conditions of inequilibrium by a law which touches only the conditions of exact

Jan 1, 1942

By Nicholas J. Grant, Rolf Nordheim, Bill C. Gissen

The system Cb-Re was examined in detail utilizing pure metals, careful melting techniques, and heat treatments. Metallographic and X-my methods were utilized for phase identification. In addition to soli-dus determinations, the composition limits and mode of formation of the intermetallic compounds 0 and X were determined. The binary constitution diagram Cb-Re has been the subject of a number of investigations.&apos;-7 Two intermetallic phases have been described and tentative diagrams have been proposed. It was the aim of this investigation to obtain a complete and more accurate diagram through the use of purer alloys and improved techniques. The diagram worked out by Knapton7 was published after our experiments were concluded. EXPERIMENTAL METHODS The starting materials were Re powder of 99.95 pct purity, supplied by the Chase Brass and Copper Co., and Cb rondelles of 99.6 pct purity, supplied by the Electro Metallurgical Co. Typical analyses are given in Table I. The alloys were prepared from compacted Re powder and Cb rondelles; for critical compositions master alloys were used. The melts varied from 5 to 50 g and were melted nonconsumably in an Heraeus arc furnace, under ti-tanium-gettered laboratory grade argon. It was necessary to invert the buttons and remelt, followed by crushing and further remeltings, the number of which depended on the nature of the fracture of the alloy. The weight of each alloy was carefully checked after each melting operation. It was found that significant weight losses occurred only when loose Re powder was used. These losses amounted to a maximum of 3 wt pct in such instances, but did not occur after the charge was once molten, regardless of the composition. Typical loss of weight after 4 remelts, including crushing, amounted to 0.2 wt pct. Oxygen pickup during successive melting cycles did not appear to be significant, since the weight gain values never exceeded 0.01 wt pct. Control of the composition of the alloys was not too difficult; however, the problem of homogenization was a far greater one. Optimum alloying conditions were finally achieved by a combination of master alloy production and repeated crushing and remelting cycles. The compositions of the alloys utilized to determine the constitution diagram are listed in Table 11. The weight balance method was used ultimately as the sole analytical method, since repeated checks indicated that it was accurate to 0.2 pct, if the weight losses were small. The results were double-checked by a metallographic study of the homogeneity of the samples, since this was presumed to be the more serious factor. Heat treatments at 1200°C or lower were accomplished by sealing the specimens in Vycor tubes, heating in a Globar furnace, followed by air cooling. This was considered to be an adequate quench because of the very low temperatures involved. For temperatures of 1200" to 1900°, alloys were homogenized and heat treated in a tungsten-filament resistance vacuum furnace. From 1900" to 2750" a tantalum-tube resistance furnace was utilized, operating under a vacuum of 105 mm Hg. The temperature measurements up to 1200" were made by means of platinum-platinum 10 pct Rh thermocouples, with an accuracy of 5°C. For temperature measurements greater than 1200°C, optical pyrometers were used. They were calibrated against a standardized instru-

Jan 1, 1962

By J. N. Anderson

Developments in reverberatory furnace practice at Noranda over the period 1928 to 1953 are described. Features of interest are increasing furnace tonnage from 700 to 2000 tons per furnace day, the use of the suspended basic roof, and improvements in furnace fuel ratio. Points of comparison and contrast between Canadian and American furnace practice are given. OPERATIONS at the Noranda smelter have been described in a series of papers published between 1930 and 1945.1-4 In the present paper the development of reverberatory furnace practice over the 25 year period from the beginning of operations up to the present time will be reviewed and current furnace practice will be described in some detail. Particular attention will be given to those features which are characteristic of Canadian practice: first, the use of the suspended basic roof; second, the deliberate driving of furnaces to give tonnages in the range of 1400 to 2000 tons of solid charge per furnace day; and third, the continuous operation of furnaces at high tonnage for many years without any extended shutdown for repairs or rebuilding. Finally, points of comparison and contrast between reverberatory furnace practice in Canada and the United States will be discussed. The Noranda smelter, which started in December 1927, was built for the reduction of ore from the Horne mine where early development had shown a comparatively small tonnage of rich cop-per-gold ore of a grade and character which would require direct smelting. The smelter was designed for a capacity of 1000 tons of ore a day and originally consisted of eight roasters, two reverberatory furnaces, and two converters, together with necessary auxiliaries. The discovery of a large new body of ore in 1928 made it necessary to increase both the size and the scope of the reduction plant. A mill which was ultimately brought to a capacity of 3000 tons a day was built for concentrating low grade ore and in addition the capacity of the smelting plant had to be greatly expanded, so that for the next fifteen years every effort in the smelter was directed toward achieving maximum tonnage. Because the layout and design of the smelter made it impracticable to add a third furnace, it was necessary to increase the tonnage on the two original furnaces by improvements in furnace design and operation, roasting and converting equipment being added as required to meet the demands of increased furnace capacity. The smelter tonnage, which reached a peak in 1941, was curtailed in 1944 when it became necessary to operate on a one-furnace basis because of a reduction in the output of the Horne mine due to a shortage of labor. Two-furnace operation at a moderate tonnage was resumed in 1949. In the early years the smelter tonnage was made up almost entirely of Noranda ores and concentrates. However, since 1937 an increasing tonnage of ores and concentrates from other mines has been treated in the smelter. In recent years, with decreased production from the Horne mine, copper concentrates from several mines in the district, treated on a custom basis, have formed an important part of the smelter tonnage. A review of developments in reverberatory furnace practice at Noranda falls naturally into the following four periods: 1—1928 to 1935, the early years; 2—1935 to 1944, the period of greatest change and development; 3—1944 to 1949, one-furnace op-

Jan 1, 1955