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By T. L. MacKay

Oxidation rates of single-crystal and poly crystalline zirconium in oxygen at temperatures from 307° to 815°C obey the parabolic rate law for short ex-posure time, 4 to 6 hr. The activation energy for the oxidation of single-crystal zirconium between 420° and 790°C is 42.6 ± 0.7 kcal per mole, and in the temperature range 307" to 600°C the activation energy for oxidation of poly crystalline zirconium is approximately the same. The high-activation energy is indicative that diffusion through the bulk oxide film is the primary mode of mass transport for both types of metal. The higher oxidation rates for poly -crystalline zirconium in this temperature range were attributed to differences in the orientation of the grains in the metal with respect to the oxidizing surfaces. Above 600°C, vain growth was observed in polycrystalline zirconium, and the oxidation rates approached those of single-crystal zirconium. ThE kinetic data of previous oxidation studies1-' of zirconium in oxygen have been interpreted by both parabolic and cubic rate laws. There is some evidence that there is a transition from the parabolic to the cubic rate law at prolonged exposures, but the question is still controversial. For the parabolic rate law activation energies are reported in the range 18.6 to 35 kcal per mole, and for the cubic rate law in the range 38 to 47 kcal per mole. So far as the mechanism of zirconium oxidation is concerned, inert marker studies10,11 have indicated that the oxidation proceeds by oxygen (anion) diffusion through the oxide film toward the metal-metal oxide interface. Pemslerl2 observed that the orientation of the grains in the zirconium metal substrate affected the rate of formation of the oxide film on the surfaces of the grains and that the orientation dependence of the corrosion rate persisted beyond the initial stages of reaction. The rate of oxidation was a minimum when the c axis of the grain was parallel to the surface of the sample, and rose to a maximum when the c axis was inclined at about 20 deg to the plane of sample surface, and decreased again at higher inclinations. cox13 observed that in 300°C steam a thin oxide film was formed initially on zirconium and that this oxide film, which exhibited interference colors, became dark first along the grain boundaries and then over the whole surface in an inhomogeneous manner as the film thickened. Cox proposed a mechanism in which oxygen diffused along preferred paths created by grain boundaries in the metal and formed a much thicker film at or near the grain boundary than on the central zone of the grain. In the present study, the oxidation rates of single crystals of zirconium were measured in oxygen and compared with the oxidation rates of polycrystalline zirconium of the same bar stock. It was felt that such a comparison would elucidate the role of grain boundaries in the metal substrate. SAMPLE PREPARATION Single crystals of zirconium were prepared by following the procedure of I3apperport,14 starting with 1/4-in. rod purchased as crystal-bar zirconium. Zirconium rods 2 in. long were wrapped in tungsten foil and sealed in quartz tubes at pressures of less than 10-6 mm of mercury. Large single crystals were grown by thermal cycling above and below the a-/3 transformation temperature, 862°C. Several specimens were simultaneously subjected to the same cycling procedure, heating to 1200°C, holding for 4 hr, then cooling in the furnace and holding at a temperature of 840°C for 5 to 10 days. This cycle was repeated five or six times for each set of specimens. The grain size of the crystal-bar zirconium before thermal cycling was between 10 and 30 p. Fig. 1 shows the microstructure of an end section of as-received crystal-bar zirconium. A longitudinal section of each zirconium rod after thermal cycling was polished and examined under polarized light, see Fig. 2, and the largest single crystals were selected for this study. Zirconium rods 1/8 in. in diameter and 1/2 in. long with spherical ends were machined from the single crystals and from the as-received bar stock. An X-ray examination showed that the c axis of the single crystals made either a 34-deg or an 89-deg angle with the rod axis. The specimens were chemically etched for 2 min in solution consisting of 15 parts hydrofluoric acid (48 pct), 80 parts nitric acid, and 80 parts water. The chemical polish removed 1 to 2 mils from the surface. EXPERIMENTAL The Sartorius vacuum microbalance used in this study has a sensitivity of 0.5 pg and a capacity of

Jan 1, 1963

By P. G. Carpenter, H. B. Fisher, W. E. Bergman

The conditions under which barium carbonate can be used to remove sulfates from drilling muds are limited The amount of sulfate remaining in solution in the system after treatment with barium carbonate is shown to be a function of the concentration of the carbonate and barium ions and the concentration of other electrolytes. Barium hydroxide may advantageously replace barium carbonate when the contamination is not entirely due to anhydrite (calcium in the system is then stoichiometrically less than sulfate) or when the carbonate concentration is high. The effect of substances such as quebracho, phosphates, and chromates, which form complexes or precipitates with barium, is discussed. INTRODUCTION As the complexity of the operations in drilling for oil has increased, more attention has of necessity been directed to the problems pertaining to the maintenance of good drilling mud properties. As a result, chemical treatment of muds has become an important factor in recent years. Some of these treatments have been designed to eliminate the deleterious effects of contaminants in aqueous mud systems by precipitation or other means. The most common of substances encountered during drilling include sodium chloride, cement, and calcium sulfate while various other contaminants: usually in small amounts, may be introduced from the water, clays, and other materials used in preparation of the mud. In certain cases, for example where continued salt-water flow is encountered or massive anhydrite is drilled, special muds may be used so that the physical properties of the mud will remain satisfactory for drilling. In other cases, it is desirable to remove the contaminants so that soluble electrolytes in the system are maintained at low values. For sulfate contamination, the conlmon practice in the field is to add barium carbonate to precipitate the sulfate as barium sulfate Ordinarily such a procedure gives satisfactory results. There have been important instances, however, where addition of barium carbonate was not effective in removal of soluble sulfates from drilling muds. and it is to these cases that the present paper is directed. While it is generally known that barium carbonate is not always effective in removing soluble sulfates from drilling muds, certain inconsistencies appear in the literature as to the limitations of its use, and little explanation for the limitations are given. Varnell and Kimbrel state that "the treatment (with barium carbonate for removal of sulfate) is simple and consists in maintaining a pH of 9 with caustic soda and quebracho." They caution that concentrations of quebracho greater than 1 lb./bbl. may inhibit the reaction. In another publication', a pH of 10.5 is considered "the maximum desirable," and the indication is that as much as 2.5 lb. quebracho per barrel may be present in the particular mud under discussion. Lancaster and Mitchell5 state that appreciable amounts of phosphates in the mud will inhibit the reaction with barium carbonate and that the phosphate treatment should be discontinued at least 24 hours before addition of the carbonate. Experimental work was initiated to ascertain the factors involved in using barium carbonate for the removal of sulfate contamination in drilling muds. While the experimental data herein reported are limited, they focus attention on the pertinent factors which must be considered for successful treatment. These factors are discussed from a practical and a theoretical view, the latter being supported by equilibrium data found in the literature. Further, it will be appreciated that the factors involved in this specific study will be closely analogous to those in certain of the other chemical treatments which involve a precipitation of the soluble contaminant. A thorough comprehension of these factors should result in a more fruitful application of this type of chemical reaction to the treatment of drilling muds. EXPERIMENTAL A. Reagents Two muds were used during this investigation. For one series of tests, bentonite suspensions were prepared by dilution of a stock suspension containing 8 per cent by weight of bentonite (Aquagel). For another series, a 6.4 per cent ben-tonitic mud weighted to 9.7 lb./gal. with barium sulfate (Mag-cobar) was used. Distilled water was used in all preparations. The quebracho (72% tannin extract) was obtained from the Thompson-Hayward Co. of Tulsa and contained 11.4 per cent moisture (105 C.). All other materials were reagent grade, and concentrations were corrected for water of crystallization, if any. All concentrations are expressed in pounds per barrel (42 gallons). B. Technique The systems — either mud or water — were contaminated with either sodium or calcium sulfate after treatment with the desired amounts of sodium hydroxide and quebracbo. For treatments with barium carbonate an approximately 3-fold excess (5 lb./bbl.) was used over that computed to be required to precipitate all the sulfate as barium sulfate. Barium hydroxide was used in concentrations of 2 lb./bbl. — about

Jan 1, 1949

By A. N. Holden

A DISLOCATION mechanism has been described by Cottrell' by which metals can yield locally, I. form Liiders bands, giving rise to a characteristic stress-strain curve with a sharp yield point and appreciable strain at constant or decreasing stress. It is undoubtedly the best mechanism that has been suggested to date." In its present development, however, the dislocation mechanism provides a more satisfying explanation for the sharp yield point than for the extensive localized flow occurring at the lower yield stress. The primary objective in this paper is to extend the dislocation mechanism to account for localized cataclysmic flow by a dislocation collision process and to give experimental evidence to support such a process. Only the yielding of iron containing carbon -will be discussed, although other metal-solute systems are known to behave similarly. Cottrell Mechanism In brief, Cottrell explains the yield point in the following way: The dislocations in iron which must propagate to produce slip usually lie at the center of local concentrations of carbon atoms, since segregation about these dislocatlons relieves some of the local stress resulting from them. A dislocation surrounded by a "cloud" of carbon atoms is thus anchored, and a higher stress is required to set it in motion than to move a free dislocation. Considering all available dislocatlons to be anchored in this fashion, the iron exhibits a yield point when the first dialocations break free and move through the lattice causing slip. This first breaking away of a dislocation enables other dislocations to break loose by "interaction" and the process becomes a cataclysm producing local deformation or Luders bands. The yield point in the stress-strain diagram for iron is absent in freshly deformed material, but returns gradually with time; the phenomenon is one aspect of what is called strain aging. The rate at which the yield point returns following straining depends on the temperature of aging. According to Cottrell the rate of return of the yield point in strained iron is limited by the rate of diffusion of carbon at the aging temperature, the mechanism is onr: of reforming the solute atmospheres around carbon-free dislocations that had stopped moving coincident with the removal of stress. If the specimen is retested immediately after straining and unloading, carbon will not have had time to diffuse to, and re-anchor, dislocations and the yield point will not occur. The carbon diffusion limitation for the rate of strain aging apparently applies if the criterion for strain aging is either the change in hardness" or the change in electrical resistance" of the strained speci- men with aging time. The possibility exists, however, that the yield point actually returns to strained iron at some rate other than that deduced from hardness or electrical resistance data. Therefore, as a preliminary experiment, the rate of yield point return in a rimmed sheet steel strained 6 pct in tension was measured at 27°, 77°, and 100°C. A plot of yield-point elongation for each of these temperatures against aging time appears in Fig. 1. The aging process is described by curves which rise to a plateau value of elongation that seems independent of temperature, but at a rate that depends on temperature. Very long times lead to a further rise in the yield-point elongation above the plateau value. However, if the later increase in yield-point elongation is ignored and the log of the time to reach half the plateau value of elongation is plotted against 1/T, a straight line results for which an activation energy of about 25 kcal pel- mol may be assigned. Within the accuracy of this sort of experiment this is approximately the activation energy for the diffusion of carbon in iron (20 kcal per mol), and the carbon diffusion limitation suggested for the yield-point return on strain aging is valid. The Cottrell mechanism thus explains in a qualitative manner the occurrence of a yield point in iron and its return with strain aging. It fails, however, to explain some of the other experimental observations that have been made of the yielding behavior of iron. For example, it is known that the yield point in iron becomes less pronounced with increasing grain size. Annealed single crystals of iron have very small yield-point elongations .if indeed they have any,' compared to a polycrystalline steel. If the only requirement for a yield point is that the dislocations in the lattice of the annealed. material be anchored by carbon atoms, the difference in the behavior of single crystals and polycrystals is not explained. That a dislocation mechanism may be entirely consistent with little or no yield point in an annealed single crystal will become apparent later when dislocation interaction is discussed. Strain aging produces a definite yield point even in single crystals. This accentuation of the yield-point phenomenon in single crystals after strain

Jan 1, 1953

By Joseph R. Stephens

Deformation substructures of' polycrystalline tungsten, W-2, 9, and 24 pct Re, and W-3 pct Ta were studied by tra?zsrnission electron microscopy. The stress-strain curve for unalloyed tungsten showed gradual yielding followed by work-hardening. Electron nzicrographs indicated a gradual increase in dislocation density with increase in strain up to 5.0 pct. Dislocations, although frequently jogged, were straight over moderate distances and were in a randorn array. Stress-strain curves for alloy specimens of W-2 and 9 pct Re and W-3 pct Ta exhibited a drop in stress at yielding followed by only slight work-hardening. Electron micrographs of these specimens after strains of 0.05, 0.1, and 0.5 pct revealed no change in dislocation substructure from the unstrained specimens. After 2.0 pct strain, the three alloys exhibited dense networks. W-3 pct Ta was characterized by straight, frequently jogged dislocations comparable with the dislocation structure in unalloyed tungsten after a similar amount of strain. In contrast, W-2 pct Re exhibited dislocations that contained widely spaced jogs, while W-9 pct Re had developed a cell structure after the relatively srnall strain of 2.0 pct. The W-24 pct Re alloy contained a few dislocations after 0.1 pct strain, while after 0.5 pct strain twins were evident. Dislocation slip bands apparently preceded the twins. The stress-strain curve for the alloy indicated that twinning commenced after approximately 0.25 pet strain. These results indicate that the primary effect of low rhenium concentrations (2 and 9 pct) in tungsten is to increase dislocation multiplication after macroyielding by reducing the Peierls-Nabarro force (lattice resistance to dislocation motion). The dislocation bands that precede twins in W-24 pct Re may be caused by localized internal stresses resulting fro a metastable structure, for example, clustering of rhenium atoms. The effect of high rhenium additions (22 a 65 pet* Rproperties of tungsten. Klopp, Witzke, and Raffo5 reported bend transition temperatures as low as -100°F (200°K) for dilute electron-beam-melted W-Re alloys tested in the worked condition. Recrystalliza-tion increased the bend transition temperature, but alloys with 2 to 4 pct Re were still markedly superior to unalloyed tungsten. Fractographic examinations of tungsten and W- 3 pct Re and W- 5 pct Re alloys by Gilbert 6 revealed that these low rhenium alloys showed a greater tendency toward cleavage failure than did tungsten. Garfinkle 7 showed that rhenium additions, up to 9 pct, to (100) oriented tungsten single crystals increased the proportional limit stress and decreased the flow stress and the rate of work-hardening. In addition, while deformation in unalloyed (100) oriented crystals apparently involved both (110)100) and (112) slip, crystals with rhenium contents of 5 andpct or more deformed primarily by (112) slip. The mechanism by which high and low rhenium additions affect the mechanical properties of tungsten is still not well-established. The present investigation was undertaken to determine by transmission electron microscopy the effects of low rhenium additions, 2 and 9 pct, and a high rhenium addition, 24 pct, on dislocation substructure in the early stages of deformation of polycrystalline electron-beam-melted tungsten. Unalloyed tungsten and a W- 3 pct Ta alloy were included for comparison. EXPERIMENTAL PROCEDURES Materials. Triple electron-beam-melted tungsten, W-2, 9, and 24 pct Re, and W- pet Ta were used for this investigation. Chemical analyses of the cast ingots are given in Table I. A description of the starting metal powders and melting and fabrication pro-cedures for unalloyed tungsten and the W-Re alloys is reported.5 The W-3 pet Ta alloy was processed in a similar manner. Compression specimens measuring 0.300 in. (7.6 mm) in length by 0.130 in. (3.3 mm) in diam were machined from swaged rods. All alloy specimens were annealed in a vacuum of 8 x 10- 6 Torr (10'2iVper sq m) for 1 hr at 3600°F (2255°K). The recrystallized grain size ranged from 0.06 to 0.08 mm diam for the alloy specimens. Unalloyed tungsten was annealed at 2400° F (1589°K) for 1 hr to produce a recrystallized grain diameter of approximately 0.12 mm. Specimens were electropolished in a 2 pet NaOH solution to a diameter of 0.125 in. (3.18 mm) to remove surface notches resulting from grinding and to improve reproducibility of t data. The ends o the compression specimens were ground flat, parallel to each other, and perpendicular to the longitudinal axes with 4/0 emery paper. Compression Tests. The compressive stress-strain apparatus used for compression tests is described in detail by Stearns and Gotsky.9 Room-temperature compression tests were conducted at a crosshead speed of 0.01 in. per min (0.25 mm per min).

Jan 1, 1969

By S. K. Tarby, W. O. Philbrook

The rate of FeO and MnO reduction from silicate and aluminate slags by carbon-saturated iron is dependent on both slag composition and temperature. Owing to variable stirring rules during- the course of reaction, the reduction processes occur in two stages in stationary graphite crucibles. In the first stage the slirring caused by CO evolution results in forced convection conditions in the slag. As the boiling action subsides during reduction, the flow conditions within the system become those defined by natural convection. Analysis of the data by an unsteady-state penetration model indicates that the rate of reduction is controlled by the rate of cocur-rent flow of' cations and anions from the bulk slag to the interface during both stages of reduction. The hearth reactions of the blast furnace have been extensively investigated from the viewpoint of equilibrium by deductions from controlled laboratory experiments. However, the literature contains relatively few studies that have been directed toward the area of slag-metal reaction kinetics, and the majority of these have been concerned with the rate and mechanism of sulfur transfer between iron and slag. More recently, attention has turned to the kinetic factors in the reduction of oxides from liquid slags. One of the first investigations in this field was conducted by Dancy1 on the reduction of pure liquid FeO and pure liquid Fe3O4 by carbon-saturated iron. The integrated form of the rate equation indicated that the reduction of FeO was of the first order up to 80 pct reduction. Over the initial 30 pct of reduction, the magnetite reaction was also interpreted as a first-order process. The rates of reaction were extremely rapid as indicated by a 30 pct reduction of the liquid magnetite in a time interval of about 1 to 2 sec. Phil-brook and irkbbride' studied the reduction of FeO from a lime-alumina-silica slag by carbon-saturated iron and solid graphite in stationary crucible assemblies. With the use of the differential form of the rate equation, the rate of the reduction reaction was found to be proportional to the second power of the concentration of the reactant FeO for both the slag-metal and slag-graphite reactions. These authors presented a number of comments in an effort to explain the difference between the molecularity of the above reaction and the observed second-order relation. One such argument, recently extended by Wagner,3 was that the rate-limiting step may well be one of transport control. A complete discussion of the proposed mass-transport control mechanism of this reduction reaction will be presented in a later section of this paper. Kinetic studies of the reduction of other oxide species, namely chromous oxide, titania, and silica, have also been reported. McCoy and philbrook4 used rotating crucible assemblies to investigate chromium reduction from Ca0-SiO2-A12O3 slags. First-order kinetic law was obeyed for the slag-metal reduction reaction as determined by the integration method of data analysis. Due to scatter of the data, any dependence of the rate constant on temperature or slag composition was obscured. Concentration-time data for the reduction of titania from blast-furnace type slags under reducing conditions were obtained by Delve, Meyer, and Lander.' The data were too few, however, for a formal kinetic interpretation. Kinetic data for the very slow reaction of silica reduction have been observed by McCoy and Philbrook6 and Fulton and chipman.7 The former experimenters found the reaction-rate constant for silica reduction by carbon-saturated iron to be 20 to 60 times less than the rate constant for chromium reduction. In their work on the reduction of SiO2 in mechanically stirred systems, Fulton and Chipman also found small values for the specific reaction rate and derived a high value for the energy of activation for the reaction. In both of these investigations, the reduction process was assumed to follow first-order behavior. schuhmann8 proposed that the rate-limiting step in this reaction is the diffusion of oxygen from the interface through a boundary-layer film to the bulk metal phase, and Rawling and Elliott9 have reported experimental confirmation of this hypothesis for temperatures below 1600°C. Turkdogan et a1.10 have also concluded that the rate of reduction of silica is a slow process and controlled by the diffusion of oxygen in the metal, but only in the absence of carbon monoxide bubbles at the slag-metal interface. In the presence of bubbles, achieved either by injecting carbon monoxide at the slag-metal interface or by blowing it through the metal and slag, these investigators found a rather rapid reduction of silica which appeared to be controlled by an interfacial reaction involving the de-sorption of silicate ions from the slag-metal interface to the metal phase as silicon and oxygen atoms. In view of the unresolved nature of the kinetics of FeO reduction and the lack of kinetic data for the manganese reaction,

Jan 1, 1968

By Thomas P. Bispham

INTRODUCTION From the lender's point of view, the greatest period of risk in a project financing occurs during the construction phase. Therefore, most projects are supported during this phase by completion guarantees supplied by some credit-worthy entity; for example, the project sponsors in the case of a joint venture, such as that of W.R. Grace, Hanna Mining, and Liberty Capital; or by a strong contract with a utility or foreign importer/investor, such as the way North American Coal Corporation has structured many of its transactions. THE PROJECT LOAN REQUEST A project financing is a more complex form of term loan than one using an ongoing corporation's balance sheet as a credit base and, therefore, a potential lender will subject the loan request to a more rigorous analysis. To support this effort, the financial institution will require a more detailed information memorandum to evaluate the credit and technical risks of a project - especially since much of the necessary information is not otherwise available. (See also Gibbs and Sroka paper above in this Chapter) A well-prepared information memorandum will impress prospective lenders with the planning ability and general competence of management. Also, the process of preparing the memorandum will help management understand the true extent of the opportunities and limitations of its venture. Appendix A outlines the topics which comprise a complete information memorandum for a loan request. Of some interest to those seeking to sell a currently-operating mining venture, a similar information memorandum should be prepared i.e., covering the same topics discussed here, for the investment banker to use with potential purchasers. For that purpose, though, one might not want to prepare quite as much detail as for a lender. Appendix B outlines the elements of a reserve evaluation and feasibility study, often called the "Bankable Document." Usually, these studies are prepared by independent engineering consultants for use by potential borrowers and sellers. (Potential purchasers will often contract for a similar study for their use in evaluating an acquisition.) ELEMENTS OF A LOAN INFORMATION MEMORANDUM 1. Outline: The information memorandum should begin with a one-page outline covering such topics as: a. The borrower's name, address, and phone number. b. Principal contact at the borrower. c. Amount of the proposed financing. d. Desired pricing and structure of the proposed financing. This should include a description of the amount and source of equity in the proposed financing. e. Use of the proceeds and timing. 2. Summary: The most important information contained in the memorandum should be summarized in two pages. Note - If the borrower is a recently-formed single-purpose company, then elements 3, 4, and 5 will apply to the guarantors of completion and to the project sponsors. 3. The Borrower: This section is a brief biography of the borrower including such topics as date and state of incorporation, founder(s), lines of business, etc. 4. Capitalization: The borrower's current debt and equity position, including leases and trade debt, with amounts, terms, names, and addresses enumerated. Equity owners, and percentage ownership, should be identified. 5. Historical Financial Statements: If available, a minimum of five year's worth of complete historical financial statements

Jan 1, 1985

By C. E. Birchenall, W. Batz, H. W. Mean

AT temperatures between 1095' and 1347ºC, 13 runs have been made on the diffusion of silicon in iron. In two of the runs the couple compositions were entirely within the loop of the Fe-Si system; the base alloys contained 0.0 and 1.0 wt pct Si. The other runs in which the couples consisted of a 2.3 wt pct Si alloy welded to a 3.7 pct Si alloy were a phase throughout. The alloys were made by melting electrolytic iron and commercially pure silicon. Trace impurities were not determined. The butt welded couples were diffused in an inert atmosphere of argon or hydrogen. After cooling, cuts of two mils or greater thickness were machined parallel to the weld interface. The chips were analyzed for silicon. When the concentration and distance data were examined on the conventional probability plot,' no marked variation of the diffusion coefficient with concentration was indicated over the narrow composition ranges employed. Therefore, a single diffusion coefficient is recorded for each run in Table I. The temperatures and atmospheres of the diffusion anneal are listed along with the compositions of the base alloys. For a iron the silicon diffusion coefficient is given by the equation: D = 0.44 e-48,000/RT where R is the gas Constant and T, the absolute temperature. However, since the scatter in the data is fairly large, the activation energy may be in error by as much as 6000 cal per g-atom. The two ? iron points do not permit estimation of the activation energy with any reasonable certainty. They do indicate that the coefficients reported by Smithells2 and Jost,3 based on the work of Fry,4 are much too high. Although very dilute solutions in iron of manganese,5 nickel,8 cobalt,7 and molybdenum8 give dif- fusion coefficients quite close to those obtained for self-diffusion in iron, the silicon values are considerably higher, though of the same order of magnitude. This seems to be consistent with the fact that the silicon atom is the only one of those listed here which is appreciably smaller than iron. The authors are grateful to the Office of Naval Research for their support of this work under Contract N6ori-47, Task order IV, Project NR 031-184, to J. K. Stanley and the Westinghouse Research Laboratories for the silicon alloys used, to several steel companies for the silicon analyses, and to C. Wells, of the Metals Research Laboratory, Carnegie Institute of Technology, for numerous suggestions during the course of the work. References 'C. Wells: Atom Movements (1951) Chapter 2. A.S.M. 2C. J. Smithells: Metals Reference Book (1949) p. 405. Interscience. Pub. Co. 3W. Jost: Diffusion (1952) p. 223. Academic Press. 4 A. Fry: Stahl und Eisen (1923) 43, p. 1039. T. Wells and R. F. Mehl: Trans. AIME (1941) 145, p. 315. "C. Wells and R. F. Mehl: Trans. AIME (1941) 145, p. 329. 7M. F. Hawkes and R. F. Mehl: Trans. AIME (1947) 172, p. 467; Metals Technology (August 1947). .J. L. Ham: Trans. A.S.M. (1945) 35, p. 331.

Jan 1, 1953

Open pits and quarries are the major sources of all hard rock tonnages mined today. Normally, ore is fractured from the pit bench face by blasting and then truck-hauled to a primary crusher on the pit rim for conveying to the concentrator; overburden (or waste) is generally truck- hauled from the pit to dump areas. This system of truck haulage of blasted ore and waste accounts for some 25 to 50% of total mining costs in a typical open pit operation and commonly is the largest single cost factor. In an effort to reduce these costs mining companies are continually replacing their older and smaller trucks with larger ones and at the same time are studying conversions to conveyor haulage, since this method of transport seems to offer possibilities of lower costs per ton of material handled. In these studies it seems logical to consider installing these conveyors as close as possible to the mining face for maximum use of conveyors and minimum use of trucks, since a conveyor is much less labor intensive than a trucking system. However, direct loading onto a conveyor near the retreating bench face is impractical because of excessive lump size. Consequently, it is conventional practice to load the blasted rock into trucks at the face for haulage to the crusher. After crushing, the rock is fed onto a conveyor. The closer the crusher is to the mine, or if it is in the mine, the shorter the truck haul and the greater the potential for cost saving. At first appearance, in-pit conveying might seem to be an obvious solution, yet many mining companies have considered such a move while few have implemented it. To determine why, it is necessary to consider the factors influencing such a decision. A conveyor system tends to be capital intensive but can have a lower operating cost per ton than trucks. Trucks, on the other hand, tend to cost less to buy but more to operate; thus, economics tend to favor the conveyor if it can be operated close to capacity and if a sufficiently large tonnage can be passed over the system to mitigate the effect of its high capital cost. In a typical hard-rock, open-pit mine it is necessary for a variety of reasons to mine several faces of both ore and waste at the same time; and any mining system must have sufficient flexibility to permit loading from one face or another at short notice. This makes it impractical to place small mobile crushers and their consequent network of conveyors at each working face. To date no alternative has been found to using trucks for bringing all rock of the same class to some convenient marshalling point from which it may be crushed and conveyed from the pit. Another advantage of trucking when compared to conveying is flexibility. Many mines wish to deliver their shot-rock to different locations, depending on the mineral content. For instance, a typical copper mine may have to simultaneously mine milling ore, leach ore,

Jan 1, 1978

By Strathmore R. B. Cooke, Marcus Digre

That calcium will activate quartz for flotation with anionic collectors such as soaps has been known for a number of years,12'3 and the method has been applied to the concentration of various iron ores4 5 6 as well as to other materials. Clemmer gives the permissible pH range for the activation of quartz by calcium as between 9.5 and 12.5 with an optimum value of 11.0, and Hertzog7 has determined the calcium abstraction of iron ores and of Ottawa sand, but otherwise little quantitative data are available. The purpose of the present investigation was to determine those conditions under which dilute solutions of calcium chloride will activate quartz for anionic flotation and to determine the amount of calcium ion adsorbed by quartz. Those variables, the influence of which was primary, such as calcium, sodium, and hydrogen ion concentrations, were studied. Other factors which influenced the results are discussed in the description of the experiments. Three different experimental methods were used to determine the activation of the quartz. The first method was to treat a sample of ground, deslimed quartz with solutions of calcium chloride and sodium hydroxide, and to determine the calcium adsorbed by chemical analysis of the solution. Advantages are that a direct value for the amount of calcium adsorbed is obtained, and introduction of a collector which may affect the adsorption is avoided. The difficulty is to determine the extremely small differences in the calcium concentrations before and after the abstraction. The second method was to determine the activation of the quartz by observing the contact angle obtained on pressing an air bubble against the polished surface of a piece of quartz immersed in a solution of calcium chloride, sodium hydroxide, and sodium oleate. This is the well-known "contact angle" method. The third method was to determine those conditions under which an air bubble will pick up quartz particles immersed in a solution of the desired reagents. Disadvantages common to the sec- ond and third methods are that introduction of a collector is necessary to obtain bubble contact with the mineral; very small quantities of contaminant can give erroneous results; and no information is obtainable concerning the amount of calcium adsorbed. Advantages are that tedious analyses, which at best give approximate results, are avoided, and the conditions, especially for the third method, closely simulate those of actual flotation. Although the most consistent results were obtained with the third method, this paper describes the techniques used and the results obtained from each of the three methods. Distilled water and analytical grade reagents were used throughout except where otherwise stated. Between tests all glass articles were cleaned with a chromic-sulphuric acid mixture and sodium hydroxide. Abstraction Method The raw material for the tests was Ottawa sand sized between 14 and 28 mesh. Batches of 500 g of this sand were dry ground for 1 hr in an Abbe porcelain mill with 2700 g of quartz pebbles. The batches were split down to samples weighing approximately 30 g, and the weight of each sample was adjusted to exactly this value.

Jan 1, 1950

By T. D. Tiemann

The extraction of iron oxide from several types of siliceous iron ore by digestion in sulfuric acid solutions was investigated in the temperature range from 80 to 258°F at concentrations from 10-60% H2SO4. Degree and rate of extraction were found to be critical with respect to both temperature and acid concentration, requiring temperatures in excess of 250° F and acid concentrations of 50% to effect 80-90% extraction of iron oxide in 60 min. Rate of extraction is directly related to surface area and apparently controlled by the rate of chemical reaction. The activation energy was found to be approximately 25,000 cal per g mol. Extraction characteristics of the several ores were found to be similar. The Department of Minerals and Metals Engineering at the University of Wisconsin conducts a continuing research program to develop processes for the utilization of low-grade iron ore deposits such as the nonmagnetic taconites of the Wisconsin Gogebic range. Methods under study include high-intensity magnetic and high-tension electrical concentration, flotation, direct reduction and chemical extraction. Concentration of a number of low-grade siliceous iron ores by the extraction of silica in sodium hydroxide solutions has been reported by the writer and others in the literature.1,2,3,4 Residual concentrates containing as much as 65% iron have been obtained from taconites containing less than 30%. Currently an alternative approach is under study, namely the dissolution of iron oxides from siliceous gangue by extraction in acid, with subsequent recovery of the iron from solution. Such a process for extracting iron oxide from the low-grade oclitic deposits of Alberta with hydrochloric acid was investigated and described by Gravenor, Govett and Rigg.5 The current investigation, which is entirely devoted to extraction with sulfuric acid, is the subject of this paper. In part, the objective of this research was to determine the possibility of leaching iron ores in situ with sulfuric acid. The investigation of a recovery method is left to later research. The ores investigated contained hematite (Fe2O3) as the principal iron-bearing mineral and minor amounts of goethite (Fe2O3 .H2O), magnetite (Fe3O4), and iron-bearing silicates. The gangue was essentially quartz occurring as single crystals and microcrystalline aggregates. Detailed description of the ores may be found in the literature.1,2,4,6 A comparison of the specific ores used in this investigation is shown in the partial analyses of Table I. EXPERIMENTAL METHODS All extractions were made at atmospheric pressure using beakers with a mechanical stirrer for those tests run below the boiling point, and Erlenmeyer flasks for those at the boiling point. A thermostatically controlled hot plate was used for heating, and this maintained the mean temperature within the solution to within ±3°F during extraction. The sample of ore for extraction was weighed to + 0.02 g. For tests below the boiling point, the sample was added to the beaker containing the sulfuric acid solution at the desired temperature. For the boiling-point tests, the sample was added to the Erlenmeyer flask followed by distilled water and the

Jan 1, 1968

By E. W. Hart, W. R. Hibbard

The room temperature flow characteristics of a series of Cu-Cr alloys are found to be related to the amount and characteristics of the chromium-rich precipitate. The results are consistent with the theory of Fisher, Hart, and Pry, which considers the trapping to dislocation loops around precipitate particles. THE use of dispersed phases for hardening metals, particularly high temperature alloys, has been practiced for several years. Current theories describing the effects of dispersions on the basis of dislocation theory have been proposed by Mott and Nabarro.l-4 rowan," and Fisher, Hart, and Pry.8 These theories recently have been discussed by Hart,' who pointed out that the experimental data available for comparison with theoretical calculations for overaged alloys are confined to studies by Gensamer, Pearsall, Pellini, and Low,8 Shaw, Sher-by, Starr, and Dorn,9 and some hitherto unpublished data on Cu-Cr alloys which will be described herein. Certain difficulties were associated with the interpretation of the first two studies. Gensamer, Pearsall, Pellini, and Low were concerned principally with interparticle spacing and did not document particle sizes. Shaw, Sherby, Starr, and Dorn included all the data in their report which was necessary for the theoretical evaluation, but their study is subject to some question1" due to difference in prior history of various specimens and due to systematic differences in particle characteristics" which appeared to be associated with differences in the magnification of their microscopic examination. The current study was undertaken to obtain critical data in a manner which avoids some of these limitations. Specimens of eight Cu-Cr alloys remaining from a previous investigation'' were available in the form of 100 mil sheet of the compositions shown in Table I. At 200 mil this material had been previously an- nealed ½ hr at 1000°C, water quenched, and cold rolled 50 pct. Samples were further cold rolled to a thickness of 20 mil for a total reduction of 90 pct. Sheet tensile specimens of standard geometry with a gage length 2½ in. long and 0.20 in. wide were machined from this sheet. These specimens were then annealed in a dried N2-H2 atmosphere for the times and temperatures indicated in Table I. This anneal served two purposes: 1—to obtain a fairly uniform and nearly constant grain size of 0.015 to 0.022 mm average diameter, and 2—to precipitate a fairly uniform, nearly spherical over-aged dispersion of essentially pure chromium. A typical microstructure illustrating the grain size and the uniformity of the dispersion is shown in Fig. 1. A typical dispersion illustrating the range of sizes and shapes of the particles is shown in Fig. 2. The tensile specimens were tested in the Instron equipment described in ref. 11 at room temperature at a strain rate of 0.04 in. per in. per min. Three tests were run on each alloy with sufficient repro-ducibility that a single true stress-true plastic strain curve could be drawn through the three sets of data for each alloy. Typical curves are shown in Fig. 3.

Jan 1, 1956

By Paul J. Bennett, James E. Castle

The kyanite minerals, including kyanite, silimanite, and andalusite, are anhydrous aluminum silicates with the formula Al2O3SiO2. lumortierite and topaz are also included in this group because they are closely allied in composition and thermal behavior. However, leither are mined commercially today. They ire typical metamorphic minerals which are found in metamorphic rocks on every continent. Kyanite minerals are prized chiefly for their refractoriness and are important components in a broad range of acid refractory products, especially in mortars and castables. World production of kyanite minerals is currently about 250,000 tpy and has been increasing at the rate of 3 to 8% a year for many {ears (Anon, 1972). While kyanite minerals lave widespread occurrence, the consumption of kyanite minerals is concentrated in the relatively few highly industrialized areas where refractories are manufactured, and which in turn are typically close to the major iron and steel producing regions of the world. Thus northern Europe, England, the United States, and Japan are the principal consumers of refractories and kyanite minerals. Of these countries only the U.S. is a significant producer. The largest use for domestic kyanite, both raw and calcined, is in the manufacture of refractory mortars, cements, castables, and plastic ramming mixes. In these applications kyanite constitutes from 10 to 40% of the mixture, the balance being refractory clays and coarser grog materials. A certain proportion of raw kyanite is used to offset the shrinking of the clay binder, whereas the calcined kyanite is used in the coarser sizes for body. Certain high grade refractory shapes and kiln furniture are composed chiefly of graded sizes of calcined kyanite bonded with a little ball clay. In most applications, however, kyanite is used only to fortify the mix or as a minor ingredient. Some manufacturers of mullite brick use domestic calcined kyanite for the fine-grained portion of the brick, but seldom in quantities exceeding 10% by weight. Kyanite is used by ceramic manufacturers of wall tile and sanitary ware to offset shrinkage and cracking after firing. Indian kyanite has traditionally been used to make the coarse grog sizes needed for the manufacture of mullite brick, and for the grog portions of mortar, cement, and castable mixes. It has been largely replaced, however, by synthetic mullite, high fired flint, and diaspore clay or calcined bauxite for other applications. Indian kyanite is available in lump as well as in granules and concentrates, as are sillimanite materials from India and South Africa. Andalusite is available as 8 mesh granules and as fines with chemical analyses not unlike those of domestic kyanites. From time to time and from place to place minor amounts of other aluminum silicate minerals such as dumortier- ite, topaz, and pyrophyllite have been made available at various prices and specifications, but have not been significant factors in the refractory industry. A mixture of kyanite and sillimanite of relatively fine grain-size and of low iron content was at one time offered to the molding sand industry by the Du Pont Co. as a byproduct from their deposits of heavy

Jan 1, 1975

By J. D. Edwards, C. S. Taylor, A. S. Russell

Solubilities and densities are reported for alumina hydrates in NaOH solutions under the conditions of the Bayer process employed to purify alumina for aluminum production. The equilibrium constants for reactions to form A102- accurately correlate the effects of NaOH concentration on the solubilities; the temperature coefficients of the equilibria lead to heats of solution of 7.34, 4.76, and 5.24 kcal for the a trihydrate, a monohydrate, and ß trihydrate, respectively. Moderate calcination does not change the solubility of the a trihydrate but increases that of the a monohydrate. At solution temperatures above 100°C, the trihydrates transform rapidly to a monohydrate. Hydro-thermally produced a monohydrate is the stable phase above 20°C. CORNERSTONE of the aluminum industry, the Bayer process for alumina production is based on the difference in solubility of hydrated alumina in NaOH solutions between 170" and 40°C. Solubilities and densities in this system were determined thirty years ago by M. Tosterudl under the direction of J. D. Edwards and C. S. Taylor. The solubility data, which still appear more concordant than those in the literature, are employed to sort out the complex equilibria in this system. The nomenclature for the alumina hydrates is that presented in The Aluminum Industry' and more recently summarized in Alumina Properties.3 a trihydrate is the purified phase analogous to gibbsite, a monohydrate is similarly akin to boehmite, and ß trihydrate is sometimes known as bayerite. Literature Goudriaan4 reviewed the early literature on alumina in NaOH solutions and presented data on the solubility at 30°C of several forms. The gel precipitated from aluminum sulphate or chloride with ammonia and dried at 130" to 140°C had the highest solubility but was unstable, its solubility being cut to one-tenth on standing in NaOH. The solubility of the gel was decreased one-third on calcination at 300" to 400°C, while that of the gel transformed in NaOH at room temperature to a crystalline trihydrate was unchanged on calcination. The preparation produced by slow hydrolysis of sodium aluminate solution with CO, in the air gave solubilities in rough agreement with later values for a trihydrate. The solubility of this form was increased slightly by the calcination. Fricke and Jucaitis5 extended the work to 30" and 60°C with a trihydrate. They also gave solubilities for ß trihydrate (bayerite b), although the residue belonging to the left branch of the solubility curve corresponded to a trihydrate. The residues at high NaOH concentrations were mono and trisodium aluminates, the latter being 3Na2O . AL2O3 • 6H2O. Solubilities of a trihydrate at 95°C and of a monohydrate at 150" and 200 °C come from the work of Magarchak.6 a trihydrate transformed to a monohydrate above 95°C. A few measurements on the solubility at 25°C of an alumina trihydrate prepared by slow hydrolysis of NaAIO, solutions were presented by Sprauer and pearce.? Their analyses showed the monosodium aluminate to have the composition Na2O.Al2O3-2.5H2O.

Jan 1, 1956

By R. C. Buehl, M. B. Royer

A three ton per day blast furnace using blast temperatures up to 2200°F was operated to recover manganese from open-hearth slag and manganiferous iron ore. The spiegel product containing 12 to 24 pct Mn was blown in a basic converter to high manganese oxide slag as described in the paper which follows this one. HE steel industry, which uses over 900,000 short tons of contained manganese yearly, depends on the maintenance of supplies of manganese from foreign sources such as India, Gold Coast, and Union of South Africa. Domestic ores furnish less than 10 pct of the yearly requirements of metallurgical ores and, unfortunately, the United States has no large, or even moderate, grade deposits that could supply a major portion of the requirements for many years. It may be necessary to obtain manganese from low grade materials, even though the cost is appreciably higher than when imported ores are used. An American Iron and Steel Institute questionnaire prepared in 1949 revealed the startling fact that open-hearth slags contained about as much manganese as is imported annually. The quantity of slag produced each year by integrated steel plants (not including cold-melt shops) and the manganese contained in the slag follows: Flush slag, 4,500,000 net tons a year containing 408,000 net tons Mn; tap slag, 7,230,000 net tons a year containing 364,000 net tons Mn; for a total of 772,000 net tons Mn. Some of this slag is returned to blast furnaces but about 75 pct would be available for the recovery of manganese. In several decades, when many of the present sources of high manganese ores have been depleted, open-hearth slag and manganiferous iron ore may become the major source of the manganese required for steel production. In 1948, manganese-ore imports from the U.S.S.R. were 427,000 net tons or 34 pct of the foreign ore. Imports from this source were only 81,000 tons in 1949. It became evident that adequate stockpiles of manganese could not be accumulated for many years. The Bureau of Mines initiated an extensive research program on methods for increasing domestic production of manganese and entered into a cooperative agreement with the American Iron and Steel Institute for recovering manganese from open-hearth slags whereby the Institute provided additional research funds and trained operating personnel. Description of Process and Equipment Numerous pyrometallurgical and hydrometallur-gical processes and combinations of the two can be applied to the recovery of manganese from open-hearth slag and manganiferous iron ore. The process being developed at the Pittsburgh Station of the Bureau of Mines consists of: 1—Reduction of open-hearth slag in a blast furnace to recover the iron and manganese as spiegel-eisen with 12 to 24 pct Mn. 2—Selective oxidation of the manganese from the blast-furnace metal in a basic-lined converter to produce a high manganese slag with up to 60 pct Mn content, which is a synthetic ore for the production of ferromanganese. This process has the advantage over other processes in that the equipment which would be used, such as blast furnaces and converters, has been highly

Jan 1, 1953

By P. F. Gnirk, J. B. Cheatham

An experimental investigation has been made to study the effects of tooth angle and confining pressure on the force required for indentation of dry rock samples under confining pressure. In these experiments static single-tooth indentations were made on two standstones, a limestone, a schist and a slate under confining pressures of 5,000 to 15,000 psi. Sharp wedge-shaped teeth with included angles ranging from 30°to 105°were used. In all cases the force-displarenzent curves are approxinlately linear for rocks tested under confining pressure, as contrmted with the discontinuous curves obtained for silnilar rocks at atmospheric conditions. Photonzicrographs of thin sections through the craters indicate very little evidence of chip formation during indentation. In general the experimental data lie between calculated values for perfectly rough and perfectly smooth tooth-rock interfaces. INTRODUCTION The variables associated with ctrilling at depth are numerous and, in many instances, are not easity isolated and controlled. Also, many of these variables are interdependent; however, if the variables which are sufficiently independent of each other are held constant and are appropriately introduced, the dependent variables can be more easily studied and evaluated. It is the purpose of the present investigation to study experimentally the simplest case of bit tooth-rock interaction. In essence, the experiments have involved the static loading of sharp wedge-shaped tools on the surfaces of effectively semi-infinite dry rock samples subjected to confining pressures. During this investigation the following quantities have been maintained constant: 1. Pore pressure: atmospheric (i.e., the rock is dry). 2. permeability: although the rocks may be permeable, they are "jacketed" in such a manner that the fluid is not permitted to flow through them. 3. Temperature: 75F. 4. Rate of loading: essentially static (approximately 0.002 in./second). 5. Bit tooth: a sharp wedge-shaped tooth loaded normal to the rock surface. 6. Rock surface: smooth and flat. 7. Drilling fluid: hydraulic oil. In addition to the above conditions, the rock surface was free of a layer of cuttings. There was no indexing to previously indented surfaces, and the influence of corners of the borehole was neglected because the rock samples were effectively semi-infinite. The following quantities have been varied, and the resulting effects have been studied: 1. Rock a) Type (e.g., limestone, sandstone, etc.). b) Mechanical properties ) Necessarily dependent c) Porosity on type of rock. 2. Bit tooth (included angle of wedge-shaped tooth). 3. Depth of penetration. 4. cobfining pressure. On the basis of laboratory experiments which have shown that many rocks exhibit ductile behavior at confining pressures corresponding to depths usually reached during drilling,'.' a theory has previously been advanced which predicts the force required for penetration by a single sharp wedge tooth into a ductile* material.' The material is assumed to be isotropic, homogeneous and rigidly plastic (no elastic deformation) and to obey the Coubomb yield criterion (Figs. 1 and 2). Equations governing the force required for penelration are given for two extreme cases:** 1. A smooth, perfectly lubricated tooth-rock interface (fridionless boundary), in which the solution gives a minimum value for the penetrating force: in which F = force of the tooth, b = length of tooth edge,

By E. T. Turkdogan

In this theoretical paper, a transport-reaction mechanism is suggested for the enhancement of the rate of vaporization of metals, or other materials, brought about by the process of convection and condensation. Based on theoretical reasoning it is predicted that, when there is a temperature gradient surrounding the heated object, a slight increase in the rate of vaporization brought about by natural convection is enhanced further as a result of condensation of the vapor close to the surface of the object. In order to derive the effective vapor-concentration profile, consideration is given to the super saturation of the vapor necessary for the homogeneous nucleation of the condensed phase from the vapor. For the formulation of the rate equation based on this mechanism, adequate information should be available on the transport properties of gases and vapors at elevated temperatures and the method of calculating nucleation pressures of the vapor. The methods of deriving such data are discussed in the paper and a rate equation is calculated for the vaporization of iron in helium for various temperature gradients in the boundary layer. Considerations are also given to the effect of oxygen on rates of vaporization of metals under isother ma1 and nonisotherma1 conditions. THE study of the methods of enhancement or retardation of basically transport-controlled vaporization reactions is of theoretical interest with possible practical implications. There are several ways of bringing about such effects. For example, Turk-dogan, Grieveson, and Darken1 demonstrated that the rates of vaporization of metals under isothermal conditions were increased by several orders of magnitude, up to the maximum rate attainable in vacuo, by introducing some reacting gas into the inert atmosphere surrounding the metal. Such an effect is brought about by the interaction of the counter-diffusing reacting gas and metal vapor in close proximity of the metal-gas interface. Another case which is often experienced in practice is that of vaporization affected by natural convection. Although several theoretical and experimental studies of this problem have already been made, no detailed consideration has been given to the possible effects of condensation of the vapor in the boundary layer on vaporization rates, which constitutes the subject matter of this paper. BASIC CONCEPT OF ENHANCED DIFFUSION-LIMITED VAPORIZATION When a heated object, for example a metal sphere, is submerged in a fluid medium, heat is transferred from the surface of the object to its surroundings by four simultaneously occurring transfer processes: i) free-molecular conduction at the gas-solid interface over the distance equal to the mean free path of the gas, ii) conduction and convection through the fluid-dynamic boundary layer, iii) latent heat of mass transferred, and it!) radiation. The free-molecular conduction becomes pronounced only when the total pressure in the system is very low, e.g., below 1 mm Hg. Therefore, this term may be neglected for ordinary pressures. At low mass-transfer rates, due to vaporization of the object, the contribution of process iii) to the heat transfer can also be neglected. Finally, since the temperature profile is not affected by radiation, the latter need not be considered. The natural convection is brought about by gravitational body forces arising from a decrease in the

Jan 1, 1964

By R. V. Smith, E. J. Dewees, R. H. Williams

Before this paper was presented, the terms "f" and "v1/f in flow equations had been called "friction coefficient" and "friction factor," respectively. However, many authors have used the names for the terms interchangeably, with the result that considerable confusion has existed in the literature on flow of fluids. In an effort to establish a clear and concise terminology for these terms, the authors have designated f as "resistance coefficient" and dm as "transmission factor." Transmission factors for the flow strings of gas wells were determined by tests on wells in six fields of Oklahoma, Texas, and Louisiana. Data were collected on 14 gas wells while gas was flowing through 1 1/4, 2, 21/2, and 3-in. tubing and 51/2 and 7-in. casing. The transmission factors developed from these data were found to follow the resistance equation published by Nikuradse1 for turbulent flow in "rough" pipes with an average absolute roughness of 0.000,65 in. This average absolute roughness is believed to be representative generally of the flow strings of gas wells with the exception of those that have been exposed to corrosion or other conditions causing undue roughness. Also, use of an absolute roughness of 0.000,65 in. probably will result in transmission factors that are more representative of gas-well conditions than those in common use today, provided that the transmission factors are calculated by the methods outlined in this report. The authors believe that use of these transmission factors will permit calculation of more accurate subsurface pressures in gas wells than those in use today. Engineers use transmission factors in flow equations to compute subsurface pressures in flowing gas wells from measured wellhead pressures. These subsurface pressures are necessary for determining the productivity of gas wells from "back-pressure" tests. The methods of calculating pressures in flowing gas wells, as outlined by Rawlins and Schellhardt2 in Bureau of Mines Monograph 7, are based on the transmission factors in the well-known Weymouth formula" Natural gas engineers have realized that errors introduced by using transmission factors as given by the Weymouth formula are relatively unimportant in testing low-capacity gas wells; but they also know that such factors are important considerations in testing large-capacity wells. Accordingly, the research described in this report was undertaken to fulfill the need for more accurate means of calculating subsurface pressures in gas wells. This report is part of the Bureau's study of the flow of natural gas through pipe and commercial pipelines. Several of the conclusions regarding flow in the flow strings of gas wells were influenced by the results of the study of flow in horizontal pipelines. These instances are mentioned in the text of this report. METHODS OF COMPUTING TRANSMISSION FACTORS Methods and procedures for computing transmission factors for turbulent flow in circular pipes from test data taken on gas wells have been derived and explained in detail in a paper by Smith4. As these methods were used to calculate the transmission factors presented in this report, a summary is included. Starting with

Jan 1, 1955

By C. H. Li, R. J. Stokes, T. S. Liu

Poly crystalline silver chloride specimens having different micro structures were prepared by extruding monocrystals or precompressed powder at varying temperatures. Extrusion at high temperature (370°C) produced a Fully recrystallized equi-axed material which was brittle and fractured by simple cleavage below its transition temperature (60°C for notched specimens). Extrusion at low temperatures (down to -196°C) produced an extremely tough cold-worked material with grains elongated in the extrusion direction. This material fractured below its transition temperature (which varied from —20" to —200°C) by an intergranular mode. It is concluded that cold work increases the resistance to transgranular cleavage. Fracture is then de-flected to propagate intergranularly ouer surfaces making an angle with the tensile axis with a corresponding improvement in notch toughness. This deflection is greater for extruded powder than for extruded monocrystals because of the extremely elongated grain shape. PREVIOUS work1,' on the fracture properties of fully recrystallized silver chloride under notch impact conditions has indicated that the fracture behavior is rather invariant. Thus polycrystalline materials originating from a number of sources ranging from relatively pure Harshaw single crystals to commercial-purity powder and prepared by hot extrusion at 370°C all fracture by simple cleavage with a ductile-brittle transition at 60°C, see Fig. 1. This behavior is also typical for fully annealed single crystals and appears to be fundamental to silver chloride for the particular loading conditions and notch configuration adopted. One way in which this characteristic fracture behavior can be modified is by the dispersion of alumina particles in the silver chloride matrix.' Then, particularly when the alumina particles are in the shape of spheres, the duc tile-brittle transition is much less sharply defined and occurs over a range of temperature, the transition temperature is lowered considerably and the fracture mode changes from simple cleavage to a fibrous or shear mode. The reason for this change is believed to be partly due to the nucleation of cavities in the region of the dispersed alumina particles which influences the state of stress locally and the distribution of plastic constraint at the root of the notch. It is emphasized however that this is not the sole factor responsible for the increase in toughness. In extruded materials it is reasonable to expect that the resultant grain size, texture, and the residual cold work will be influenced by the presence of a dispersed phase. The present experiments were initially designed to determine the extent to which cold work influenced the fundamental ductile-brittle transition temperature and fracture mode of silver chloride. They were stimulated by the discovery that silver chloride could be extruded at temperatures as low as -25°C when commercial-purity powder was used and as low as -196°C when single crystals were used. This working temperature was far below that previously used (370°C) and it was found that such extruded material was exceedingly tough. The reasons for the enhanced toughness and its relation to the previous work on silver chloride-alumina compacts will be discussed. EXPERIMENTAL PROCEDURE Materials Used. Two forms of silver chloride were used in this study: i) silver chloride powder

Jan 1, 1964

By D. D. Pollock, D. I. Finch

PREVIOUS work' on a four-component system (Cu-Mn-Ni-Fe) resulted in the development of consistent relationships between the chemical composition and the electrical properties of these manganin-type alloys. The present study was undertaken to develop alloys with temperature coefficients of resistivity and thermoelectric properties more desirable than those of the manganin type, and to determine whether the relationships which had been found for the Cu-Mn-Ni-Fe alloys could be extended to five-and six-component alloy systems. PREPARATION OF ALLOYS The five- and six-component alloys were made from OFHC copper rod, ARMCO iron, electrolytic nickel, manganese cathode chip, and special high-purity aluminum and indium. A Cu-Mg master was used for deoxidation. The compositions of these materials are given in Table I. The alloys were melted and cast as indicated previously,' except that no cover was used and the melts were teemed. EXPERIMENTAL PROCEDURE The finished wire size of 0.0808 in. diam was again selected for the reason that the test specimens would be self-supporting, thus eliminating any extraneous effects which might have been introduced by winding smaller diameter wire on spools or forms. The following tests were performed on the finished wires: A) Chemical Analyses—All of the components of these alloys were determined by premium gravimetric methods. The principle residual impurity contained in these alloys is 0.008 pct Mg (by weight), maximum. Tables II and 111 present the mean analyses computed from at least four determinations of each constituent. B) Electrical Characteristics—The test procedures forthe determination of the thermal emf vs copper, resistivity, and the resistance-temperature relationship are given in detail in Ref. 1. It was also shown that the resistance-temperature relationship is of the form Pt = P,n + At -B/2, [1] where P, is the resistance at t C expressed in percent of a given standard resistance, P, is this per- 0 centage at 0°C, and A and B are constants for a given, properly treated, composition. The results of the electrical tests are also given in Tables II and 111. INTERPRETATION OF DATA The modified Linde2 formula, known to be valid for dilute binary solid solutions, was applied to the complex Cu-Mn-Ni-Fe type alloys. This relationship is of the form where Ap is the increase in resistivity. Zb the valence of the solute, Zn the valence of the solvent, and K and K are designations assigned by Robinson and Dorn3 to the constants for the periodic group to which solute element belongs. The units of K, are -cm per valence difference squared per at. pct; those of K2 are pa-cm per at. pct. For the sake of brevity, these units are not repeated in the following discussion. Calculations were made to determine if this formula was accurate for five-component systems. No experimental studies were made of the Cu-In and Cu-A1 binary systems. The effects of indium additions upon the resistivity of copper are given by Linde. These data, along with those of the effects of Mn and Fe, permitted the calculation of the effects of Al.

Jan 1, 1959

By M. E. Wadsworth, J. S. Cho

Colloidal silica dissolved in aqueous suspensions of high surface area thoria was permitted to adsorb on the thoria surface. Silica in three forms was identified by means of infrared spectroscopy and characterized by Si-O vibrational modes in the 9.0 to 10.5 region. A band at 9. 1 was attributed to free unad-sorbed silica, while adsorbed silica resulted in bands of greatest absorbency at 9.4 or 10.3p, depending upon the character of the surface bonding. Size considerations, the amount of silica adsorbed, and the character of the infrared spectrum for strongly adsorbed silica has been interpreted as evidence of formation of chain or layer silica linkages on the thoria surface. Infrared spectroscopy has been used successfully in recent years for the identification of minerals and synthetic inorganic substances. Unfortunately, infrared studies of silicate minerals have not proceeded to the point that unambiguous assignments may be made for Si-O vibrational modes as affected by lattice substitutions and structures. Several investigators have presented infrared spectra of silicate minerals, 1,7 and their work has pointed to the potential value of infrared to identify minerals as well as to note the sensitivity of various Si-O modes to structural and substitutional environments. The strong infrared absorption of various Si-O vibrational modes makes the infrared method a very useful tool for the study of silica adsorbed on the surface of solids of high surface area. Spaepen et a18 studied the adsorption of silica on thoria surfaces from aqueous solutions of silicic acid up to 300°C. A general review of the application of the infrared method to the study of surface adsorption has been presented by Pliskin and Eischens.9 EXPERIMENTAL Du Pont colloidal silica, Ludox SM (16.36% SiO2) and Ludox HS (30.25% SiO2) in aqueous suspensions were used as the sources of silica for the adsorption studies. The thoria used for this study was supplied by Union Carbide Nuclear Co.* under the preparation *Chemical Technology Div., Union Carbide Nuclear Co., Oak Ridge, Tenn. . number D-16-650. It was fired from the oxalate at 650°C and had a surface area of 32.4 ± 0.6 m2 per g measured by nitrogen adsorption. The chemical analysis of the D-16 thoria in ppm was: CO3= (5500), NO3- (< 4), so4±PO4 (100), and Si (20). The Ludox silica was placed in contact with the thoria in a two-liter autoclave10 between room temperature and 250°C. Slurries were made by placing 1 g of thoria in one liter of distilled water containing various additions of the silica sol. Samples were removed from the autoclave by opening a small stainless steel valve. Extracted samples were centrifuged for 10 min and excess solution was decanted. The sample was then dried for several hours at 80°C. Infrared samples were prepared by thoroughly mixing 10 mg of the thoria with 400 mg KBr in polystyrene vials on a "Wig-L-Bug" amalgamator. Pellets were pressed under vacuum and had final dimensions of 1.2 cm diam and 1.2 mm thick. Infrared spectra were

Jan 1, 1963