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By Thomas L. Kesler

IF all of the 14 million tons of kaolin produced in Georgia through 1949 had been mined from a single deposit 20 ft thick, it would represent a mined-out area of less than 1 sq mile. This measure of depletion is in such strong contrast with the large size of the area shown in Fig 1, from which most of the output has come, that there has developed an optimistic outlook concerning future supplies of kaolin equal in grade to that already mined. Dozens of mines are widely but not evenly distributed within this area, which consists mostly of parts of Twiggs, Wilkinson, and Washington Counties. There are scarcely half a dozen kaolin mines elsewhere in the state that have been appreciably productive, although the areas in which they occur may have good possibilities. Most of the efforts to find new deposits, therefore, will be confined to the area shown in Fig 1, for it is here that the industry has established its facilities, and prospecting is encouraged by relatively numerous outcrops and opened deposits of kaolin. The Ounulgee and Oconee Rivers drain the area. They and their main tributaries have flat, swampy flood plains covered with alluvium and for the most part thickly wooded. The divides between these streams are intricately dissected by the minor tributaries, which are spring-fed in steep-walled, cirquelike heads. Gullying and undercutting are active along the streams but are partly controlled by unusually dense vegetation. The mean altitude is about 350 ft above sea level and locally the relief ranges up to 200 ft. The area is near the inner margin of the Coastal Plain, and is underlain by Cretaceous and Tertiary formations. Its general location is shown in black in Fig 1. The division between the Coastal Plain and the Piedmont Plateau is actually a very sinuous line, the fall line, marking the extent to which erosion has stripped the Coastal Plain sediments from the southeastward- sloping surface of the crystalline rocks of the Piedmont.

Jan 1, 1952

By J. E. Lawver, R. R. Beebe, J. L. Wright

This paper describes a separation coefficient Ks useful as a "measure of goodness" to be used in factorial and response surface methodology. The use and limitations of the separation coefficient is discussed, and a computer program for calculating the coefficient is presented (in Fortran). The application of the coefficient is illustrated in terms of electro-dynamic separation of Mesabi semitaconites. Present Mesabi wash ores and gravity concentrates with an iron to silica ratio of only 5 or 6 to 1 are no longer readily salable, for the steel industry now demands concentrates with an iron to silica ratio of at least 10 to 1 and a structure comparable to pellets. THE ELECTRODYNAMIC-FLOTATION PROCESS Wash ores and gravity concentrates consist largely of hematite, goethite, and quartz. Several methods of beneficiation are being tested at the University of Minnesota Mines Experiment Station with a view to improving the salability of these materials. A process that looks particularly promising from an economic view is the combination electrodynamic-flotation process shown in Fig. 1. This process is designed to upgrade gravity concentrates to a 5% silica content and to provide fines — from the flotation circuit —to permit agglomeration. Although electrodynamic concentration of specularite is an established and well-understood process, the electrodynamic concentration of Mesabi goethite-hematite gravity concentrates presents interesting and enigmatic problems. The gravity concentrates are by no means assemblages of clean, free mineral particles. Instead, the concentrates to be retreated consist of a great variety of middlings. More or less complete liberation would require grinding to minus 325 mesh, at which size only flotation could be expected to function, and even then grades and especially recoveries would not be satisfactory. A close examination of the gravity concentrates shows that grinding to 100% minus 48 mesh will produce a product which is approximately 70 weight % plus 325 mesh (see Fig. 1). The remaining 30 weight % can be deslimed and floated by conventional methods, but the resulting iron concentrate is

Jan 1, 1964

By W. A. Boyer

MINE hoisting ropes can be loaded to capacity only when the strength of each component is exactly known. Characteristic curves provide this information. When load and rate of acceleration are specified for an individual rope, the characteristic curve assumes a predetermined shape. The factor of safety, is the value drived by dividing the breaking strength of the rope by the total stress on the rope where it goes over the sheave wheel. Total stress is made up of the weight of skip and cage, weight of ore, weight of rope between sheave and rope clevise or attachment, friction load, and acceleration force. Bending stresses are neglected in this discussion, as sheaves and drum di- ameters are assumed to be large enough to reduce these stresses to a negligible value. Fig. 1 gives a family of characteristic curves for 13/8-in. improved plow steel hoisting ropes of 6 x 19 or 6x21 construction with hemp core. The curves will hold for all rope sizes of this specified quality when they are subjected to an acceleration rate of 1.608 ft per sec2. As acceleration rates of most deep shaft hoists are close to this value, this family of curves is presented as a standard. Deviation is slight for other rope sizes. Note that for 0, which is the most important, the deviation is within the width of the line on the graph. In a comparison of values for ropes of 5/8-in. and 21/4-in. diam the deviation ranges from a maximum 1.2 pct at 0 ft to 0.6 of 1 pct at 5000 ft. The value of 5.55 is the same for all ropes at 2500 ft. The tabulated columns to the left of the curves are values of total connected load for each size of rope from 5/8-in. to 21/4-in. diam inclusive. With each of these connected loads the safety factor characteristic curve for the rope will be as indicated.

Jan 1, 1957

By J. J. Sense

This paper deals with the Asarco method for the determination of equipment availability. It is apparent that no single availability equation will provide all of the information necessary for the complete record of equipment availability. In the Asarco system, three separate availability records are carried on major equipment. The first is mechanical availability, the second is operational availability, and the third is a historical record of equipment use. The derivation of the three items and the definitions of the time elements used in the various availability equations are presented and discussed. The Mission unit of American Smelting and Refining Co. is located in Pima County, Ariz., approximately twenty miles southwest of Tucson. The unit consists of an open pit copper mine and concentrator. The concentrate is shipped by rail to company smelters at Hayden, Ariz., or El Paso, Tex. Mine stripping began at Mission in the fall of 1959, and construction of the 15,000 tpd concentrator was started in the spring of 1960. The first ore was treated in midsummer of 1961 and full ore production was reached in August 1961. At that time over 46 million tons of material had been moved from the pit. Current production (ore and waste) averages over 100,000 tpd. Total production to date (ore and waste) is over 120 million tons. Major equipment currently in operation at the Mission pit is as follows: 3 P & H 1800 Electric Shovels - 9 cu yd dippers 1 P & H 1900 Electric Shovel - 9 cu yd dipper 2 B.E. 190-B Electric Shovels - 9 cu yd dippers 35 Euclid LLD Trucks - 60-ton capacity 6 KW-Dart 55-T Trucks - 55-ton capacity 1 KW-Dart D-2550 Truck - 65-ton capacity 1 LeTourneau Westinghouse 65 Haulpak - 65-ton capacity 3 Euclid S-18 Scrapers - 274011 capacity 4 Caterpillar D-8 Dozers 4 LeTourneau Westinghouse Tournadozers 1 Michigan 280 Dozer 2 Caterpillar Model 12 Motor Graders

Jan 1, 1964

By F. A. Crossley

The additions: o, Sn, Zr, Cb, Mo, and V were studied for their effects on the a/a + Ti3A1 boundary of Ti-A1 base ternary alloys. These additions were chosen because of their importance to commercial titanium alloy practice—oxygen as the major impurity (in terms of influence on mechanical properties) and the others as major alloying additions. The composition and temperature ranges of investigation varied according to the system as shown in Table I. The principal means of investigation was light-microscopy. The basis of the ternary studies was the recently determined titanium-rich end of the Ti-A1 system.' MATERIALS, METHODS, AND PROCEDURES Ingots were prepared in 15 g size for all alloys except those of the Ti-Al-Mo and Ti-A1-V systems which were prepared in 20 g size. These alloys were prepared from high-purity, electrolytically refined titanium obtained from the Bureau of Mines, Boulder City Metallurgical Laboratory, Boulder City, Col., and Chicago Development Corp., Riverdale, Md., for the 15 g and 20 g ingots, respectively. The alloying additions were obtained from commercial sources, and were at least of 99.9 pct purity, except the V-15 wt pct A1 master alloy for vanadium additions. This alloy was of 98.2 pct purity; the major impurities were 0.37 wt pct Fe and 0.38 wt pct Si. The ingots were prepared by multiple melting in a nonconsum-able-electrode arc furnace. The ingots were covered with Markal coating "CR-T22" (Markal Co., Chicago) to protect them from scaling. Markal coating is applied as a slurry of solids in a volatile carrier liquid. The coated article is air dried. Upon heating to the hot working temperature the coating fuses to form a glassy, protective film. The ingots were heated to 1000°C and rolled to 0.16 in. thickness, reheating after each pass. The glassy film was removed by sand blasting. Ten mils were removed from all surfaces by grinding and/or pickling (3 pct HF, 30 pct HN03, balance water). Grinding was always followed by pickling to remove traces of the grindmg medium. Hydrogen removal was accomplished by vacuum annealing at 1000°C for 2 hr to a final pressure of about 0.02 µ Hg. This was followed by annealing in evacuated Vycor bulbs (800°C, 2 hr, ice brine quenched (IBQ)) in order to maximize ductility. The materials were then lightly pickled and cold rolled to 0.12 in. thickness or cracking, whichever occurred first. After rolling, 1/2 by 1/4 in. specimens were cut for the equilibration anneals given in Table 11 and carried out as previously described.1 Final polishing of the specimens for metallographic examination was done by electrolytic means. Specimens were first examined without etching under bright field and polarized light illumination, for this is the only means of detecting the "two-phase" phenomenon.1,2 The term "two-phase syndrome" refers to Ti-A1 al-loys which appear on the bases of their chemical, phys-ical, and mechanical properties, to be two-phase, but are actually single phase and inhomogeneous in composition. Since processing and heat treatment of the alloys of the first group were completed before the two-phase syndrome was discovered, a number of the specimens were found to exhibit this phenomenon—80 pct of those of the Ti-Al-Zr system, 5 pct of those of the Ti-Al-Sn system, 30 pct of those of the Ti-Al-Zr system. The affected specimens were homogenized by heating above the ß transus temperature and quenching. The heat treatment and working schedules to accomplish this are given in Table 111. After rolling, specimens were pickled (3 pctHF, 30 pct HNO3 in water) to remove 1 to 2 mils from the

Jan 1, 1970

By George A. Packard

The shaft of the Raven mine, at Butte, Mont., is an incline 1,700 ft. in length and dipping at various angles. At the top the dip is 70" from the horizontal, but this is gradually flattened until at the 300-ft. level the inclination is only 47. This angle continues to the 1,100-ft. level, below which it curves with a 125-ft. radius to 78". In addition to these variations in dip, the shaft does not lie in the same vertical plane, with the result that the hoisting rope not only rubs at intervals on both the hanging and foot walls, but presses strongly against the west dividers near the collar, while 300 ft. below it runs close to the east end plates. The early operators used no idlers, aid wall plates cut 6 in. deep by the rope resulted. Later operators first attempted to overcome the excessive friction, and the wear of rope and wall plates, by introducing solid cast-iron idlers, 3 in. in diameter. To allow for the travel of the rope from side to side some of these had to be 3 ft. long and were very heavy. Judging from the appearance of the old idlers of this type found at the mine, they often failed to turn in the bearings, which is not surprising when it is considered that they would have to make 1,000 rev. a minute under ordinary hoisting conditions. The next rolls were made of wood, 6 in. in diameter, with an iron band about each end, and a pintle of 1-in. round steel driven in at the ends to serve as a shaft. These wore rapidly, and were soon replaced by rolls made from water pipe, 5 or 6 in. in diameter, cut to the desired length and fitted with a wooden cylinder into which the pintle was driven. Where the idlers were used on the hanging wall of the shaft the original bearing was simply a piece of + by 1 3/2 in. strap iron, 10 in. long, turned up at the end in a circle 11/8 in. in diameter to receive the shaft. A small hole served for oiling, and common black oil was generally used, although filtered oil from engine bearings and compressor bearings was also used. When the rolls were to be placed on the foot wall the bearings were made from two pieces of 1 by 3 in. steel, 6 in. long. A half circle was cut on the flat side of each piece and the two half circles together formed a bearing. Oil holes were provided, and in some cases holes were bored through the two pieces so that they could be screwed or spiked to the wall plates.

Jan 1, 1915

By F. P. Bullen

Empirical relationships between fracture stress, orientation angle, and diameter of crystal have been determined at 77°K. Orientation ranges of markedly different behavior were found—a law of constant normal stress&apos; of value a (diameter)-1/2 for the fracture of ductile crystals, and a condition of shear stress (or strain) for more brittle crystals. The observations are not consistent with current theories. An interpretation is advanced which is also applicable to observations on the effect of prestrain at room temperature on the subsequent fracture stress at 77°K and to the effect of cyclic stressing on the cleavage strength.&apos;&apos; The law of constunt normal stress&apos; and the brittle ductile transition are also explained. The interpretation is more consistent with the initiation of cracks by intersecting dislocations than with theories based on stress -concentration by dislocation arrays. ZINC single crystals are particularly suited to the study of cleavage because fracture occurs on the basal plane over a wide range of crystal orientations. Analysis of the conditions of stress and strain at fracture in crystals of different orientations should indicate which parameters control the cleavage process. Unfortunately, controversy has arisen over the correct empirical relationship between tensile fracture stress and orientation. Schmid&apos;s observations1,2 favored a &apos;law of constant normal stress&apos;, as observed in other materials.2 For zinc, however, the observed values are far below the theoretical strength and cannot represent the true limit of cohesion between neighboring atomic planes. Hence, the interpretation of such a &apos;law&apos; is not straightforward. Deruyttere and Greenough3,4 found a complex variation between tensile fracture stress and orientation; this variation did not agree with a &apos;law of constant normal stress&apos;. Two theories have been advanced to account for their observations: a) the propagation of cracks from low-angle boundaries,5 and b) the release of energy from piled-up dislocations during crack-propagation.4 The present work resolves the apparent discrepancy between the observations and shows that neither of the above theories are applicable to the tensile fracture of zinc single crystals. A phenomenological explanation, along the lines suggested by Gilman,&apos; is advanced and successfully applied to previously unexplained effects. EXPERIMENTAL DETAILS &apos;Crown Special&apos; redistilled zinc was used, except for one comparison series op tests using &apos;Tadanac&apos; electrolytic zinc. Crystals of 1 mm diam, subsequently called &apos;1 mm crystals&apos;, were grown from the melt in vacuo in precision-bore Pyrex tubes internally coated with graphite. Several specimens 1 in. long were cut from each crystal and chemically polished. Jigs were used to minimize handling strains, and crystals were mounted in the Polanyi machine the day prior to testing to allow recovery from any such strains. One-mm crystals were chemically polished for long periods to obtain 0.1 mm (approx) crystals. One-mm crystals were cemented into miniature gimbals by &apos;Araldite&apos; casting resin. The Appendix gives the reasons for using gimbals and the results obtained by other methods. More complete details of all techniques are given elsewhere.7 The symbols and terminology used are as follows: X = orientation angle (angle between tensile axis and line of greatest slope in the basal plane). T = tensile stress (on true cross-section) S,N = shear and normal stress (components of T with respect to the basal plane) ? = shear strain D = crystal diameter. The subscript &apos;f&apos; will be used to denote values at fracture. PART I-ANNEALED CRYSTALS EXPERIMENTAL OBSERVATIONS One-mm crystals were used to establish the variation of fracture stress at 77°K with orientation at fracture (Xf), Fig. 1. For 18 deg = Xf = 55 deg, a &apos;law of constant normal stress&apos; was observed. For Xf > 55 deg, the fracture condition approximated to a constant shear stress. At Xf< 18 deg, twinning occurred before fracture so that the results were not typical of homogeneous single crystals,4,8— such specimens will not be considered herein. The dependences of fracture stress upon Xf were of similar type for 6 mm,* 1 mm, and 0.1 mm crystals,

Jan 1, 1963

By A. Gittens

In two recent papers stevens13,14 has compared the methods used by various investigators to calculate the contribution of grain boundary sliding to the total creep strain. In previous work Gittins15 has shown that the strain due to grain boundary sliding Egb may be estimated from the lateral, d, and vertical, v, displacements of a longitudinal marker line as it crosses a grain boundary. He obtained the formula, where nL is the number of grains per unit length measured on the specimens before creep, and \p and ?2 are, respectively, the angle between the grain boundary and the stress axis in the plane of the surface and normal to the surface (see Ref. 13, Fig. 1). The subscript L indicates that observations are made in a longitudinal direction parallel to the stress axis. It was generally found more convenient to measure the grain size of a specimen before creep; this was necessary to characterize the specimen. Stevens13,14 however prefers to define Egb in terms of NL, the number of grains per unit length after creep. Since nL - NL(l + €1) where cr is the total strain, Eq. 1 11 becomes The quantities d and ?, can be determined directly from the marker lines, while v and ?2 are found by sectioning normal to the surface. In the interior when symmetry considerations suggest that v = d and ?1 - ?2 the strain due to grain boundary sliding may be estimated from a suitable internal longitudinal marker line by the equation To determine cgb at the surface using either Eqs. |1| or [2] requires a large number of observations, particularly if many estimates of cgb are required, and also it is impossible to determine the variation of cgh with creep strain on the same specimen because the specimen must be sectioned to determine ?2. Gittins15 found that for a number of materials the sliding vectors d and v were related to each other so that provided one was measured the other could be estimated. Also during creep, grain boundaries tend to migrate to meet the surface at angles close to 90 deg,16 so that the term nL(v/tan?2)L was comparatively small when 4,- 90 deg. By determining the frequency distributions of ?2 on control specimens this term can be estimated without sectioning the creep specimen. In this way an empirical equation based on Eqs. |1| and |2| enables cgb to be determined with reasonable accuracy with fewer observations of the angles and sliding components. For B brass, the relationship between the sliding vectors indicated that ceb = 0.85nLVL . In Table I of Stevens paper this equation is stated to be incorrectly derived. However it should be pointed out that this is an empirical equation derived from Eq. |1|, based on a knowledge of the relationship between d and v and on the frequency distributions of ?1 and ?2 for a particular material under given experimental conditions. It was used only when these relationships were known, to reduce the number of observations required to estimate cb. Therefore the comment "derivation incorrect does not apply. Gittins did in fact make use of Eqs. |1| and 131. Author&apos;s Reply I must apologize to Dr. Gittins for the inadvertent misrepresentation of his work. Other slight errors in Table I have been pointed out to me. Langdon (No. 12 Table I), Gifkins (No. 7, Table I), and possibly other workers did not measure the number of boundaries per unit length after creep (A&apos;) but the corresponding quantity bejove creep (n). As Dr. Gittins points out, a completely analogous set of equations to those given in the paper can be derived by the same method: using the number of boundaries per unit length before instead of after creep. The basic equations corresponding to Eqs. [2] and 171 of the paper, from which all the others can be derived, are

Jan 1, 1968

By E. P. Sadowski, L. P. Trudeau, C. R. Cupp

The apparent high radiation resistance of two varieties of maraging steel is described and an indication is given of some phenomena that require further study. Two aspects were included in this work: a) effects of neutron irradiation on the room-ternperature tensile properties of 18 Ni (250) maraging steel; b) effects of neutron irradiation on the room-temperature tensile properties and on the impact resistance of 12 pct Ni-5pct Cr-3 pct Mo maraging steel. Subsize tensile and impact test specimens were used. The 18 pct Ni steel was irradiated to a neutron fluence of 3.3 x 10 n per sq cm (>1 mev). The 12-5-3 variety was irradiated to two neutron fluence levels: 1) 1.5 x 10 per sq CTTL (>I mev) 2) 3.1 x 10&apos;9nper sq cm (>1 mev) In all respects both varieties showed unusually high resistance to change in strength and ductility after neutron irradiation. There was no noteuorthy over-aging in fully aged specimens. The aged 18 pct Ni variety showed a 4 pct drop in yield strength upon irradiation with little or no decrease in ductility. The aged 12-5-3 variety showed a 5 pct increase in yield strength on irradiation with a very slight decrease in ductility. The impact transition of the 12-5-3 variety only increased 35 to 45 Co upon irradiation, and the shear energy fracture value for the subsized specimens decreased only half as much as expected frorn experience with conventional steels. THE development of new reactor concepts and nuclear technology now depends to a large extent on the availability of new alloys capable of operating in stringent reactor environments. One of the most important new types of metallurgical material that has been developed in recent years is the maraging steels. This type of alloy is notable for its remarkable retention of toughness at high strength levels. In the course of a research program which incorporated ambient temperature radiation effects studies to determine the mechanisms of strengthening of 18 pct nickel maraging steel,&apos; it became evident that the resistance of the alloy to neutron-induced changes in mechanical properties should be studied. This paper is to present evidence for the apparent high radiation resistance of maraging steel. The experimental program was in two parts: a) Effect of neutron irradiation on the room-temperature tensile properties of 18 Ni (250) maraging steel. b) Effect of neutron irradiation on the room-temperature tensile properties and on the impact resistance of 12 pct Ni-5 pct Cr-3 pct Mo maraging steel, a cobalt-free alloy. EXPERIMENTAL PROCEDURE 18 Ni (250) Maraging Steel. Tensile test specimens were machined to the dimensions shown in Fig. 1 from $-in.-diam hot-rolled bar of the composition shown in Table I. All specimens were austenitized for 1 hr at 1500°F (815°C) in argon and water-quenched. Half of the specimens were then mar aged for 3 hr at 900° F (482°C) in argon. The specimens were irradiated in a fast neutron transformer rod at approximately 120°F (50°C) to a neutron fluence of 3.3 x 10 10 n per sq cm (>1 mev). The irradiation was performed in the NRX reactor of Atomic Energy of Canada, Ltd., at Chalk River, Ontario, Canada. In all cases, the thermal neutron flux was three times as high as the reported (> 1 mev) fluxes. The room-temperature tensile tests were performed with an Instron tensile testing machine operating at a crosshead speed of 0.05 in. per min. The loading frame of the machine, mounted in a hot cell, was remotely controlled. The specimens were tested either

Jan 1, 1968

By J. Chipman, N. J. Grant, H. L. Bishop

In a recent review1 of the iron-oxide activity of simple open-hearth type slags containing lime, magnesia, silica, and iron oxide, it was established that activity values were lacking in the range of slag composition most interesting to open-hearth operators. That is, laboratory experimental results were not available for simple slag compositions having basicity ratios ranging from about 1.5 to 4.0 and iron-oxide concentrations ranging from about 10 to 30 pct. However, in order to estimate iron-oxide activities for the range of slag compositions of interest to open-hearth operators, activity values were calculated using ionic concepts of slags by following the treatment of Flood and Grjotheim.2 A study of these data indicated that the desired slag compositions were probably not obtained due to the limited solubility of the basic constituents lime and magnesia. Thus, it was decided to make additional heats equilibrating molten carbon-free iron with slags containing MnO, P2O5, and A12O3 in percentages commonly found in open-hearth refining slags, and to examine the liquid slag field over a greater composition range than that of the simple slags studied previously. The ranges of the constituents used in the present study are summarized as follows: Metal Pct 0 0.070-0.201 PctMn 0.005-0.142 Slag Pet CaO 24.45-41.90 Pct MgO 3.23- 9.21 Pct MnO 1.13- 9.62 Pct FeO 20.08-48.7 Pet Fe2o3 2.23-10.14 Pct SiO, 3.51-28.50 Pct P2O5 0.23- 3.65 Pct A12O3 0.43- 3.38 The experimental procedure consisted of equilibrating the slag and metal in an induction furnace provided with a carbon-arc slag heater as described previously by Winkler and chipman.3 The data obtained were used to establish both the iron-oxide activity values and the activity coefficients for FeO and MnO in the open-hearth type slags. RESULTS AND DISCUSSION I) Iron-Oxide Activity of Open-Hearth Type Slags—The iron-oxide activities of various iron-oxide contents of the open-hearth type slags, as calculated from the oxygen content of the metal by the method of Taylor and Chipman, 4 are plotted against the molar ratio of bases to acids in Fig. 1. The data represent slag-metal studies carried out in the temperature range 1530" to 1700°C. For the purpose of this study, the constituents CaO, MgO, and MnO are assumed equal on a molar basis in their effect on the iron-oxide activity; and, PO2,5 is calculated as an equivalent to SiO2 and A12O3 on a molar basis. The total iron content of the slag is calculated to FeO and is represented by the symbol (FeO) The points presented in Fig. 1 represent the results of this investigation for iron-oxide slags containing between 1 and 8 mol pct MnO and less than about 2 mol pct each of A12o3 and PO2,5 The curved lines shown in Fig. 1 represent the results of the summary1 of the iron-oxide activities of simple slags of the system (CaO + MgO) - SiOz - (FeO),. The solid portions of the curved lines represent experimental data, and the broken portions represent iron-oxide activities calculated from ionic concepts of slags. The iron-oxide values of the open-hearth type slags of this study are in good agreement with the values of the simple iron-oxide slag system. Only a slight modification is indicated. The results of this study suggest that the line for 20 mol pct (FeO), should be lowered slightly in the basicity range 1.5 to about 3.

Jan 1, 1959

By C. M. Wayman, R. Bullough

The selection rules fir twinning in fcc, bcc, and hcp lattices as established in Part I are used to predict the positions, relative to the matrix (untwinned) reciprocal lattice, of reflections due to double diffraction from the twins. These doubly diffracted reciprocal lattice points, which often coincide with ordinary matrix reciprocal lattice points, are deduced from kinematical considerations. 1. INTRODUCTION. A Bragg diffracted beam under certain conditions may be rediffracted to produce a third beam i .e., the incident beam, the first diffracted beam, and the rediffracted beam). In such a case an initial incident beam is diffracted by a set of planes; the diffracted beam may then be rediffracted by a second set of crystallographically different planes. The initially diffracted beam acts as an incident beam for the second set of planes, thus resulting in a twice-diffracted beam. This twice-diffracted beam may be regarded as a singly diffracted beam from a set of lattice planes whose structure factor is zero. In order to find the reciprocal lattice point corresponding to the doubly diffracted beam, it is customary1,2 to combine vectorially the component reciprocal lattice vectors which head to the extra reflections. For example, hcp metals may exhibit the forbidden (0001) reflections.* As Fig. 1 shows, the (0001) reciprocal lattice point results from: (1010) + (1011) = (0001) Similarly, combination of the beams (1011) and (1010) gives rise to: (1011)- (1010) = (0001) As Pashley and stowell2 have pointed out, even if in fact the double diffraction points lie on the Ewald sphere, they may not actually represent observed doubly diffracted beams because the initially diffracted primary beam, i.e., the source, must also correspond to a reciprocal lattice point which lies on the sphere of reflection; however, the reciprocal lattice point corresponding to the planes causing the rediffraction need not lie on the sphere. A little consideration of the Ewald construction in Fig. 2 makes this argument clear. In Fig. 2 the incident beam is represented by the wave vector ko. This beam is initially diffracted by the lattice planes (hkl) which are represented by the reciprocal lattice vector r(hk1) The reciprocal lattice point (hkl) of course lies on the Ewald sphere in order that the Bragg condition be satisfied. The initially diffracted beam is represented by the wave vector kl. If now kl is rediffracted by a second set of planes (h&apos;k&apos;l&apos;) resulting in k2 the lattice point (h"k"l") must lie on the Ewald sphere if the double diffracted beam k2 is to be observed. It is clear that (h"k"l") is the vector sum of (hkl) and (h&apos;k&apos;l&apos;), and that (h&apos;k&apos;l&apos;) need not in fact lie on the Ewald sphere. In the usual1,2 analysis the origin (000) of the reciprocal lattice is shifted to the point (hkl). However the translation (000)— (hkl) automatically shifts the point (h&apos;k&apos;l&apos;) to the Ewald sphere, i.e., (h&apos;k&apos;l&apos;)— (h"k"1"). It is to be seen that (hkl), the source reflection, must lie on the Ewald sphere; if it does not the translation (000) — (hkl) places the "new" origin off the sphere, not satisfying the Ewald construction and hence the Bragg condition. The point (h"k"1") of course must lie on the sphere if it corresponds to an observed doubly diffracted beam. It is to be noted that crystals having the hcp structure, as just discussed, are intrinsically capable (i.e., without lattice defects such as twins) of double diffraction, as contrasted to fcc and bcc crystals which are extrinsically capable (requiring lattice defects or reoriented lattice regions). The "allowed" bcc reciprocal lattice points (which comprise an fcc lattice) are (110), (2001, (211), (220), (310), (2221, (321), (4001, and so forth. The combination

Jan 1, 1967

By Denton R. Wieland, Harvey T. Kennedy

The frequency of bubble formation is measured for oils from the East Texas field and the Slaughter field in cores from these fields. East Texas oil was also tested in a Slaughter core. The data checked previous results in that a definite supersatrtration may exist without the formation of bubb1es in periods of observation totaling 158 hours. Measured frequencies varied from zero at supersaturations below 14 psi to 3.1 bubbles/sec/ cu ft of rock at 40 psi supersaturation INTRODUCTION The formation of bubbles is essential in a solution gas-drive reservoir in the displacement of oil. In formations where the basic pore structure consists of fine pores or fractures connected to larger pores or crevices, bubble formation in the fine pores may be the most effective means by which the oil may be displaced, even if secondary methods are applied. The most efficient oil displacement would occur where at least one bubble would form in every pore. As the pressure decreased, the bubbles would expand, due to the gas diffusing into the bubbles, and displace the oil toward the lower pressure area around the wellbore. Kennedy and Olson&apos; have shown that the total number of bubbles formed in a reservoir is primarily dependent upon the rate of pressure drop, the rate of diffusion of gas through oil, and surface area of the rock. The surface tension of the liquid and other characteristics of the rock may also exert an important effect, which would give a different frequency to different rocks and oils and various combinations of different rocks and oils. The purpose of this paper is to present frequency measurements on two rocks and two oils. A hydrocarbon liquid is at equilibrium with gas when it contains the maximum volume of gas in solution that it can have at the prevailing temperature and pressure. If the temperature is maintained constant and the pressure is reduced below the equilibrium pressure, the liquid is supersaturated to the extent of the difference in the two pressures until the gas comes out of solution as a free phase. As shown by previous work, and confirmed in the present investigation, the number of bubbles formed per second increases with an increase in supersaturation. Kennedy and Olson showed that a tenfold increase in the rate of pressure decline resulted in a tenfold increase in the number of bubbles formed per cubic foot of reservoir rock, e.g., for decline rates of 0.1, 1, and 10 psi/day, the number of bubbles formed per cubic foot would be 40, 400, and 4,000, respectively. For another hypothetical system, Stewart, Hunt, and Berry&apos; estimated that for pressure declines of 0.1, 1, and 10 psi/day, the number of bubbles formed would be approximately 10, 100, and 1,000 cu ft of reservoir rock. Their value of bubbles formed is lower than given by Kennedy and Olson; however, the order of magnitude agreed satisfactorily. The rate of diffusion will also affect the number of bubbles formed. If the diffusion coefficient is high, the gas in solution in the oil surrounding a single bubble will tend to diffuse into the bubble rather than forming other bubbles. On the other hand, the formation of more bubbles will tend to occur if the diffusion coefficient is very low. In his thesis dated Aug., 1953, Wood" reports on the frequency of bubble formation in a system consisting of oil and rock from the Rangely field. In the present work, the apparatus of Wood was redesigned to allow greater precision, especially at low supersaturations, where bubble growth is slower, and detection is more difficult.

Jan 1, 1958

By Harold Margolin

The Ti-Al system certainly merits the investigative attention it has been receiving and this latest contribution by Blackburn is therefore to be welcomed. The titanium-rich end of the phase diagram shown here, as well as the diagram by crossley,17 Clark et a1.,18 and Tsujimoto and Adachi,29 reveal no other intermediate phase than Ti3A1, although a previous work by Ence and Margolin20 had reported an additional intermediate phase. Since then, Buerger Precession pictures of single crystals of 10.5, 11.5, and 12.5 wt pct A1 were made at New York university21 and the extinctions were found to be typical of a D6h a Ti structure. Some superlat-tice spots and/or streaks suggesting the presence of an Mg3Cd structure, i.e., Ti3Al, were also found. Since these specimens had been quenched from the y region20 and had revealed a Ti, it was concludedz1 that y does not exist and the structures interpreted as m + y could not be so considered. The absence of normal two-phase behavior in these structures had been indicated in the paper. Accordingly, the two-phase region y + 6 becomes a + Ti3Al and there is some agreement between the results of Blackburn and of Ence and Margolin20 as to both shape and existence above 900°C of an a + Ti3Al field. Crossley 17 proposed that Ti3Al is a line compound which decomposes congruently at 900°C. It is difficult to discount entirely the fact that clear two-phase structures, varying in amount with position in the two-phase field, could be detected above 900°C by Ence and Margolin.20 Precipitation during quenching could not have been so regular and the twin-phase syndrome17 does not apply. The heat treatment and quenching technique was essentially the same as was used by Crossley, whereas Blackburn, who used a rapid quenching procedure, also showed a two-phase a + Ti3A1 field above 900°C. If the two-phase field exists above 900° C, it is easy to picture a two-phase field which could account for the data of Crossley below 900°C without invoking a line compound existence of Ti3A1. All that would be necessary would be an 0 + Ti3Al/Ti3Al boundary which curved sharply to higher aluminum contents below 900° C. It is clear that the controversy must be solved by techniques which obtain data at elevated temperatures without quenching. It is conceivable that the difference in width of the a + Ti3N fields reported by Blackburn and by Ence and Margolin is due to the longer annealing times used by the latter investigators. The criterion for establishment of equilibrium in our work20 was the

Jan 1, 1969

By F. S. McNicholas

IN the writer&apos;s opinion the term "block caving" has been rather loosely applied to a general principle of mining, and therefore detailed classification of caving methods and draw, together with the characteristics required for each method, will fill a desired need. Selection of caving method should be made only after a study of all governing conditions, such as size of orebody, lateral and vertical extent of ore-body, attitude, treatment demands, possible resultant pressures, and cost and profit comparison of methods. Successful caving methods can be classified under six main headings: (1) Full-protected block caving, (2) semi-protected block caving, (3) unprotected block caving, (4) full-panel caving, (5) half block-panel caving, (6) block-panel caving. The methods and factors governing choice of method are shown in Table I. Full-protected Block Caving Full-protected block caving consists of develop-ing, undercutting, and drawing to exhaustion a block of ore of predetermined size without coordination of mining or draw with adjoining areas. The block is mined and drawn as an independent unit (Fig. 1). Full-protected block caving demands that mining and draw cannot be done against unconsolidated capping in adjoining caved and exhausted blocks. Two sides of block are against solid. Caved capping must be allowed to consolidate for a suitable period, usually not less than one year. Pillars may be left in lieu of consolidated capping (Fig. 2). The advantages of full-protected block, caving are: (1) A definite area is developed, mined, and drawn as a unit so that all work may be completed, the block drawn to exhaustion in the minimum length of time, and the area then abandoned. (? Production is not restrained because of coordina1, on with adjoining areas. (3) Records of draw and caved reserve of ore are good. Upon abanwnment of the block, accountability for production is closed out with no carry-over to adjoining areas. The disadvantages of full-protected block caving are: (1) Side dilution from the draw adjacent to the consolidated capping may be serious, particularly where capping does not consolidate sufficiently in the allotted time. (2) Boundary weakening is needed on the solid sides of the block except in weak rocks. (3) Maximum cantilever and compression effect to aid caving and fragmentation may not be secured since the draw should proceed from the solid to the consolidated capping to minimize side dilution. However, undercutting from the loose (caved capping) to the solid with a slight initial draw in the same manner will develop cantilever effect. (4) A period of at least one year, more is usually desirable, must be allowed for caved capping ,to consolidate before mining the adjoining block, depending upon the characteristics of the capping. (5) Capping may not consolidate sufficiently in the time that can be allowed. Full-protected block caving is indicated where: (1) The ground is weak and heavy and workings

Jan 1, 1952

By W. F. Baldwin

The timing circuit generally employed in Carter Reservoir Behavior Analyzers consists of a synchronous motor driven cam which actuates a microswitch. A mechanically actuated timer of this type has several inherent disadvantages, among which are the following: 1. The first time interval may not be of uniform length, depending on the exact position of the cam at the conclusion of a previous run. 2. The positioning adjustments on the microswitch lever arms are critical and require periodic readjustment. 3. The mechanical system is subject to wear which will ultimately impair its accuracy. It was therefore considered desirable to devise an alternate method which would eliminate these difficulties without introducing others that would be equally troublesome. This note describes the design of an electronic timing device which has been incorporated in the analyzer constructed by the Field Research Laboratories of Magnolia Petroleum Co. The electronic timer circuit is built around a simple neon glow tube relaxation oscillator, the complete circuit diagram of which is shown in Fig. l. The operation of the circuit is as follows: Condenser C, is charged to the critical firing voltage of the neon glow tube VR-75 by battery B1 through resistors R1 and R2. When the neon glow tube fires, the condenser C, is discharged, The charging cycle then repeats itself,

Jan 1, 1950

By A. J. Goss, M. Tanenbaum, W. G. Pfann

THE purification of two metals, antimony and tin, by zone refining is described, and a reciprocating method of passing molten zones through a straight ingot, which effects a considerable economy of time and apparatus, is discussed. Antimony has been used extensively as a doping agent in germanium and silicon. More recently, it has become important in the preparation of new semiconducting compounds such as InSb, GaSb, and AlSb. High purity is especially necessary in this latter use, since the antimony may comprise as much as 80 pct by weight of the compound. Antimony is generally available in a commercial form of about 99.8 pct purity. Column 1 of Table I is a typical emission spectra analysis of commercially available 99.8 pct antimony. Zone refining has been quite successful in increasing the purity of such samples. Columns 2 to 4 of Table I are analyses of an ingot of the antimony of column 1 after the passage of seven molten zones in an atmosphere of nitrogen. The ingot length was 9 1/2 in.; the mean zone length was 2 in.; convection was the only means of stirring. The column headings are sample locations in inches from the leading end of the ingot. The results illustrate both the merit and a limitation of zone refining. It is seen that four of the five impurities—nickel, arsenic, lead,

Jan 1, 1955

By Herbert Clark Hoover, Lou Henry Hoover

GEORGIUS AGRICOLA was not only the author of works on Mining and allied subjects, usually asso ciated with his name, but he also interested himself to some extent in political and religious subjects. For convenience in discussion we may, therefore, divide his writings on the broad lines of (1) works on mining, geology, mineralogy, and allied subjects; (2) works on other subjects, medical, religious, critical, political, and historical. In respect especially to the first division, and partially with regard to the others, we find three principal cases : (a) Works which can be authenticated in European libraries today ; (b) references to editions of these in bibliographies, catalogues, etc., which we have been unable to authenticate ; and (c) references to works either unpublished or lost. The following are the short titles of all of the published works which we have been able to find on the subjects allied to mining, arranged according to their present importance :-De Re Metallica, first edition, 1556; De Natura Fossilium, first edition, 15,46 ; De Ortu et Causis Subterraneorum, first edition, 1546; Bermannus, first edition, 1530 ; Rerum Metallicarum Interpretatio, first edition, 1546; De Mensuris et Ponderibus, first edition, 1533; De Precio Metallorum et Monetis, first edition, 1550 ; De Veteribus et Novis Metallis, first edition, 1546; De Natura eorum quae Effluunt ex Terra, first edition, 1546 ; De Animantibus Subterraneis, first edition, 1549. Of the " lost " or unpublished works, on which there is some evidence, the following are the most important:-De Metallicis et Machinis, De Ortu Metallorum Defensio ad Jacobum Scheckium, De Jure et Legibus Metallicis, De Varia Temperie sive Constitutione Aeris, De Terrae Motu, and Commentariorum, Libri VI. The known published works upon other subjects are as follows:-Latin Grammar, first edition, 1520; Two Religious Tracts, first edition, 1522; Galen (Joint Revision of Greek Text), first edition, 1525 ; De Bello adversus Turcam, first edition, 1528 ; De Peste, first edition, 1554. The lost or partially completed works on subjects unrelated to mining, of which some trace has been found, are:-De Medicatis Fontibus, De Putre¬dine solidas Partes, etc., Castigationes in Hippocratem, Typographia Mysnae et Toringiae, De Traditionibus Apostolicis, Oratio de rebus gestis Ernesti et Alberti, Ducum Saxoniae.

Jan 1, 1950

By E. L. Dougherty, J. W. Sheldon

A numerical method for computing the dynamical behavior of fluid- fluid interfaces is described. Results of studies made to assess the accuracy and economy of the method on a computer are reported. It is concluded that it is practical to obtain digital computer solutions for fluid- fluid interface problems with two space dimensions. PROBLEM DEFINITION We consider two-dimensional flow in a rectangle as shown in Fig. 1, where a fluid-fluid interface divides the rectangle into two regions. The pressure is specified along the lower boundary (the injection boundary) and upper boundary (the withdrawal boundary). There is to be no flow of fluid across the side boundaries, which are walls of symmetry for the flow, and px = 0 along these boundaries. The fluids are incompressible, the absolute permeability and porosity is constant, and each fluid has its own constant mobility. Thus, the pressure satisfies Laplace&apos;s equation in each region. Across the interface, pressure is continuous and the normal component of fluid flux is 1. These conditions are shown in Fig. 1. The velocity of a point on the interface in the direction of the normal from Fluid 1 into Fluid 2 is U,/ Q, where Un is the normal component of volumetric flux at the point on the interface, i.e., For brevity, we call the problem defined above "Problem A". Problem A is a simple example of a dynamical fluid-fluid interface problem. However, the same numerical methods which we use to solve Problem A can be extended to the solution of a large number of fluid-fluid interface problems which are mathematical models for various oil recovery processes,l hence our interest in Problem A. We define the interface by prescribing the x, y coordinates of a finite set of points. These points are called "interface mesh points" and are illustrated by triangles or asterisks in the figures. Fig. 2 is a procedural (flow) chart displaying the main steps in the solution. We comment briefly on some of these steps in this section and then give further details in subsequent sections. For mobility ratio greater than one it is well known that Problem A is unstable.*2 Small amplitude,

Jan 1, 1965

By George C. Fryburg

The oxidation of platinum at high temperatures (above 800°C) is controlled by boundary-layer diffusion, except at the lowest pressures. The rate of oxidation is determined by the rate of diffusion of oxide through the boundary layer away from the platinum. This is in contrast to the situation in tungsten and molybdenum oxidation, where the rate is controlled largely by the diffusion of oxygen through the boundary layer towards the metal. The difference in mechanism is a consequence of the much lower rate of oxidation of platinum. Evidence is presented that supports this mechanism for the oxidation of platinum. SOME years ago we reported1 a study of the oxidation of platinum in ordinary oxygen in the temperature region above 800°C, where the oxide is volatile. All the data available up to that time had been obtained with platinum crucibles or plates heated in tube furnaces. The results were widely variant and highly dependent on the flow rate of gas through the furnace: the faster the gas flow rate, the higher the measured oxidation rate. These results indicated to us2 that reaction equilibrium was being approached in the furnace and that the rate was diffusion limited. To minimize these effects, we employed1 electrically heated ribbons placed in a cooled bulb. We were primarily interested in the low-pressure region, but some experiments were performed at pressures up to 1 atm. We found a peculiar pressure dependency which was explained on the basis that the rate of oxidation was proportional to the pressure at all pressures, but that with increasing pressure an increasing fraction of the volatilizing oxide was back-reflected to the ribbon by the surrounding gas molecules. The oxide molecules are not stable at the temperature of the platinum and are decomposed on impact, the platinum being re-deposited on the specimen. Because of the limitations of space, this explanation was offered summarily without detailed support. Recently, Jaffee and associates3, 4 have suggested that the pressure dependency of our data up to 0.5 torr could be explained on the basis of a Langmuir-type adsorption of molecular oxygen followed by dissociation. Since the correct explanation of the effect of pressure on the oxidation rate is a necessary requirement to the development of the basic mechanism of the oxidation reaction, we felt that the back-reflection theory should be presented in more detail. This is the object of the present paper. RESULTS AND DISCUSSION A typical set of results showing the pressure dependency of the oxidation reaction is presented in Fig. 1, where we have plotted, on logarithmic scales, the rate of oxidation, R, vs the pressure, p. For comparison some results of Krier and Jaffee4 and of Raub and plate5 are also shown. These data were taken at atmospheric pressure using furnaces. The effect of reaction equilibrium within the furnace is apparent from the much smaller values that these observers obtained: Raub and Plate In O2 found a rate of 3.1 µg cm-2 hr-1 compared to our value of

Jan 1, 1965