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By R. L. Ash, R. R. Rollins, C. J. Konya

An investigation was performed to determine conditions for optimizing the spacing of simultaneously initiated multiple explosive columns. This was done by using models of mortar, dolomite, and Plexiglas with 10-grain mild detonating fuse as the explosive charge. It was desired to simulate blastholes with multiple primers initiated by detonating fuse or when high-velocity explosives are used in low-velocity materials. It was found that optimum spacing between multiple charges was strongly influenced by charge length. At less than optimum charge length, the spacing at which complete shearing was possible between adjacent charges decreased exponentially with a subsequent loss of broken material volume. For charges fired simultaneously, larger burdens and spacings were possible as compared to those necessary for single-crater charges. For each material studied, there was a characteristic optimum charge length and a maximum attainable spacing at any given burden. Proper selection of the spacing distance between charges is fundamental to successful blasting. Its value directly affects the cost of drilling and explosives used per unit of broken material. In addition, the choice of a spacing that is Compatible with a given set of blasting conditions aids in the control of fragmentation sizing, ground vibrations, overbreak, and throw which in turn, influence other production costs. For example, normally loaded blastholes that are spaced too closely invariably promote overbreak and usually give coarse fragmentation. Unless care is taken, airblast and violent flyrock will occur and under certain conditions cutoffs and misfires may result. Too large a spacing, on the other hand, frequently leads to conditions that form bootlegs or toes. The choice of a particular spacicg to use, however, is largely a matter of individual experience and judgment, usually based on trial and error. Very little is known or can be found in the literature with regard to how the spacing between charges is related to field conditions and charge geometry. As a general rule, the firing time sequence of adjacent charges and properties of a material are thought to have the most significant influence on the spacing distance best suited for any given field condition. For example, delayed initiation of adjacent charges usually always requires a closer spacing than when charges are fired at the same time. This should be expected if one considers that the energy normally dissipated and lost in the surrounding ground from charges fired independently would be captured and utilized for breaking material between charges when they are initiated together. Spacing can be extended also when charges are aligned with structural planes of a material, such as jointing, along which shearing is relatively easy. It is customary to relate the spacing (S) between charges to their common burden (B) in the form of a spacing ratio, or SIB. The burden normally is considered as the optimum depth or distance from any single charge perpendicular to the nearest free or open face at which the desired fragmentation and maximum crater yield are obtained. For production blasting, value of the ratio is generally considered to vary from 1 to 2, depending on conditions.1-6 When adjacent charges are fired independent of one another, the value varies from 1 to about 1.4, the closer amount being employed to square corners or produce craters having the ideal 90" apex angle. The larger ratio is the geometric balance value for craters having an apex angle of 135". The basic ideal crater forms in the plane of the charge diameter for charges fired independently are shown in Fig. 1. In the event charges are fired simultaneously, geometric balance in the plane of their charge diameters suggests that a spacing ratio near 2 would be appropriate, as illustrated by Fig. 2. In practice, however, some compromise ratio value must be selected to conform with the specific ground conditions. An example would be where the jointing planes tend to produce 60° or 120° crater angles, the appropriate geometrically balanced charge arrangement being given by Fig. 3. In this condition, the spacing ratio is 1.15, not 1 or 1.4 as suggested for the 90° cratering of independently fired adjacent charges. In view of the foregoing, it would seem logical to assume that whenever charges all having the same burden are fired at the same time, spacing distances always can be greater than those permitted by charges fired independently. In practice this is not the case, however.

Jan 1, 1970

By G. T. Jackson

PROPERTY AND LOCATION THE Alaska Gastineau Mining Co.'s mine is located at Perseverance, about 4 mi. east of Juneau, Alaska. Its property consists of a group of claims, the lode system traversing these claims for a distance of 11,000 ft. (3352 in.). It is shown in Fig. 1. FIG. 1.-ALASKA GASTINEAU MINING CO.'S PROPERTY NEAR JUNEAU. The ore deposits occur in what is known as the Juneau gold belt. This gold belt can be traced for about 50 mi. northwest of Juneau and 40 mi. southeast. It consists of a single band, several hundred feet wide, in which stringers and veins of quartz, carrying gold, occur in a slate formation near its contact with some altered volcanic rock known as greenstone.

Jan 9, 1919

By Alexandre Mussard

How do you make a mine out of 500 000 kg (1.1 million lb) of uranium oxide contained in a low-grade deposit with high clay content? Confronted with these limited reserves and mineralogical problems at France's St. Pierre du Cantal uranium deposit, engineers adopted a rather unusual mill process to keep capital costs to a minimum. Crushing and grinding were eliminated by using a high pressure monitor to slurry ore directly from stockpiles. After mechanical separation of the high-grade fines from low-grade sand, the sand fraction is washed of clay to increase permeability and then it is heap leached. Fines are sent to agitated vessels for acid leaching, and the pulp from this operation is diluted by pregnant solution from the heap leaching.

Jan 10, 1979

By V. F. Zackay, W. A. Goering

ALLOYS of iron and aluminum up to 35 wt pct aluminum are single-phase solid solutions, and are of potentially wide applicability.1-3 In spite of early and continued interest1-4 little progress has been made until recently in the preparation and evaluation of sound alloys containing more than 6 wt pct aluminum. Vacuum-melting techniques for the production of ductile Fe-A1 alloys have been described recently.1-7 A. procedure for air melting these alloys is presented here. Low-carbon iron is induction melted without a slag in a rammed magnesia crucible. At the beginning of melt-down, aluminum pig (99.95 pct Al), charged in a clay-graphite bottom-pouring crucible is placed in a pot furnace at 1800°F. The primary deoxidation of the molten iron after melt-down is effected by the addition of 0.1 pct aluminum and 0.5 pct manganese. (Hilty and Crafts" have reported a significant increase in the deoxidation efficiency of the aluminum and manganese combination over that of the aluminum alone.) A more drastic deoxidation designed to reduce the oxyen content to the lowest possible level is accomplished by plunging metallic calcium to the bottom of the melt. This is done by wiring small cubes of the metal to a steel rod. A circular shield larger than the diameter of the crucible opening is attached to the rod so that any spa'ttering of the molten metal will not endanger the operator. Since the temperature of the molten metal is above the boiling point of calcium, the bath is vigclrously purged by calcium vapor. It is believed that the calcium-vapor treatment permits a homogeneous distribution of calcium in the melt. Owing to the vigor of the reaction the temperature of the molten metal should be kept below 2900°F prior to the calcium addition. A total of 0.05 pct calcium is added in two stages in this manner. The second calcium deoxidation is made just before charging the molten aluminum into the iron, in order that an excess of calcium be present for the remainder of the melt. The aluminum, which has been removed from the holding furnace, is then hydrogen degassed by bubbling chlorine through a quartz tube immersed in the molten aluminum. The hydrogen-chlorine reaction is an exothermic one preventing the solidification of the aluminum during the 5-min chlorination. Approximately 0.1 pct calcium, based on the amount of aluminum, is then added to the aluminum. A further excess of calcium is introduced into the melt in this manner. The oxide dross is removed, fluorspar is added to the molten iron, and the molten aluminum is poured through the fluorspar slag. The fluorspar should be dried thoroughly prior to its use, as any water present will react with the aluminum. Aluminum oxide formed during the pouring operation reacts with the fluorspar slag to form gaseous aluminum fluoride and calcium oxide. A forced-draft ventilating system is required for this operation as aluminum fluoride is toxic. As soon as the molten aluminum has been added, vigorous manual stirring of the melt is required because the slag-aluminum oxide reaction is highly exothermic and tends to take place near the top of the melt. The combination of high temperature and the slagging action of the fluorspar quickly erodes the crucible at the slag line if the aluminum is not stirred uniformly into the melt. It has been found that at least 4 min of manual stirring combined with induction stirring are necessary to ensure homogeneity. The power is shut off 1 min prior to pouring to allow metal and slag to separate. As much slag as possible is removed from the melt, which is then poured directly into cast-iron molds. A mold wash of aluminum oxide is used to prevent ingot sticking. For slab ingots which are to be rolled into sheet, a carbon-tetrachloride vapor atmosphere or a chlorinated-pitch mold wash is desirable, as the aluminum oxide formed in the pouring operation is subsequently removed by the chlorine in the presence of carbon." As in vacuum melting, a pouring temperature of about 2900°F is recommended. Adequate hot-topping is important as iron-aluminum ingots are subject to very deep piping. Ingots are removed from the molds and buried in vermiculite, where they are allowed to cool slowly to room temperature. The ingots are radiographed,

Jan 1, 1959

By R. D. Vaughn

The analysis of laminar flow of power-law non-Newtonian fluids in narrow, eccentric annuli is employed in this paper to discuss the problems of lubricant flow in journa! bearings and of errors introduced by eccentricity in experimental studies with concentric annuli on extruders and wellbore annuli. The velocity profile and pressure loss-flow rate equations are developed for the laminar flow region. In addition, the expected error in flow rate and pressure-loss measurements for concentric annuli as a result of eccentricity is determined. For example, a 10 per cent displacement of the core of an almost concentric annul us would cause a 1.8 per cent decrease in the observed pressure loss for a fluid with a power-law exponent n of 0.25. The corresponding increase in the observed volumetric flow rate would be 7. 5 per cent. INTRODUCTION Non-Newtonianism and eccentricity occur simultaneously in two engineering problems: (1) flow of lubricants in journal-bearings and pressure-reducing bushings, and (2) flow of non-Newtonian fluids in plastic extruders and wellbore annuli. The lubricants used for moving parts are often non-Newtonian in character — often they are plastic in behavior. A solution to the problem of flow of non-Newtonian fluids in narrow eccentric annuli is particularly pertinent to this problem. In all experimental studies of laminar flow of fluids in concentric annuli, such as in extruders and well casings, the error due to eccentricity must be estimated or studied. A number of publicationsl-4,6,7 have dealt with this problem for Newtonian fluids; however, I am not aware of work for non-Newtonian fluids. This work is directed to the non-Newtonian problem. Before the solution to the problem is given, the pertinent conclusions from the work on Newtonian fluids will be reviewed. Heyda 1 and Redberger and Charles 2 have published general solutions to the problem of the laminar flow of Newtonian fluids in eccentric annuli, apparently without knowing of the earlier work of Ca ld well 3 and Bairstow and Berry,4 which is reported by Dryden, et al. 5 Although several mathematical routes are encompassed by the work of these authors, the results appear to be equivalent. Redberger and Charles2 show that the error caused by eccentricity in concentric annuli is negligible for small diameter ratios (k < 0.5); however, for large diameter ratios (k + l), the error in the predicted flow rate can be as great as 100 per cent or more. Partial solutions to the problem are available from the work of Dryden,5 Tao and Donovan6 and Piercy, et a1.7 Tao and Donovan examined the case of flow in narrow, eccentric annuli (k + 1) with and without rotation of the annular core. These authors also reviewed previous work on this subject and verified their approach with experimental data. Dryden gives the solution for the limiting case of complete eccentricity or tangency. Piercy, et a l. published an early solution to the problem of narrow eccentric annular flow. The conclusions of Redberger and Charles2 and the experimental proof of Tao and Donovan6 both suggest that the region of large diameter ratios (k + 1) is of main interest and that the parallel planes approximation to the solution in this region is satisfactory. This method will now be extended to the laminar flow of non-Newtonian fluids in narrow eccentric annuli. THEORETICAL SOLUTION The geometrical aspects of the problem are illustrated in Fig. 1. To represent the non-Newtonian fluid the power-law model was selected. This model has many disadvantages which have been pointed out; nevertheless, its simplicity, its frequent and wide applicability justify its use in this work. Fredrickson and Bird8 and Savins9 have used it as a basis for a theoretical study of laminar flow of non-Newtonian fluids in concentric annuli.

Jan 1, 1966

By Paul Weir

BLEACHING clays have been used extensively in the oil-refining industries for a number of years. Their use in water purification is relatively recent and less extensive. They are frequently classified as (1) those naturally active and (2) those active after artificial activation with acid. Fuller&apos;s earths fall within the general group of bleaching clays and gradually the commercial term "fuller&apos;s earth" has been applied to natural clays originating, for the most part, in the south-eastern section of the United States. Bentonitic clays are usually classed as activable and are prevalent in the southwestern section of the United States.1 Bleaching clays vary considerably in chemical and physical properties throughout the country, or even in different parts of one deposit (Table 1). TABLE 1.-Physical and Chemical Composition of Different Types of Clays Composition, Per Cent Constituents and Properties Bentonite Fuller&apos;s Mud Clay (Wyoming) Earth (P Peak)&apos; (Alluvial) Silica (SiO2) 60.0 69.6 61.67 Alumina (Al203) 19.0 14.3 20.34 Ferric oxide (Fe203) 2.8 3.90 7.70 Titanium oxide (Ti02) 0.13 0.70 1.01 Lime (CaO) 0.5 0.80 0.00 Magnesia (MgO) 2.0 1.00 0.24 Soda (Na2O) 2.4 0.40 0.31 Ignition loss 4.7 6.77 8.51 Moisture content 8.0 10.0 20.0 Isoelectric point, pH 9.2 9.2 6.5 Average swelling volume 15.0 0.0 0.0 Weight per cubic foot, lb 55.0 40.0 65.0 Material passing 200-mesh sieve, per cent 90.0 85.0 50.0

Jan 1, 1939

By J. H. Anderson, G. A. Parks, B. K. Jindal

Both temperature and humidity are important variables in electrical separations. By independent control of temperature and water vapor pressure, it is shown that charge depends primarily upon relative humidity, not on temperature or water pressure alone. All the effects of temperature changes other than modification of humidity are apparently small. However, humidity control can be achieved through manipulation of the more tractable variable, temperature. Both temperature and humidity are important variables in electrical separations. Some of their effects have been described in general by Fraas1,2 and by Ralston.3 The influence of temperature and humidity may be considered a problem in industrial separations because humidity, especially, is not normally controlled, hence leads to unpredicted quality control problems. The problem is significant whether the separation is based on conductivity differences or on contact electrification. There is no doubt that water adsorbs from humid atmospheres onto mineral surfaces resulting in conductivity changes.&apos;-8 An empirical relationship between conductivity and humidity is1.4-6,8 where a and b are constants peculiar to the substrate, and s and h are the conductivity and relative humidity respectively. Some authors have taken advantage of differences in a and b among minerals to make or propose electrodynamic separations in which h is the key control variable. Fraasl,2 reported that the charge developed on quartz and feldspar through contact with aluminum is sensitive to humidity, as illustrated schematically in Fig. 1. Lawver9 found that the charge developed by quartz on contact with several materials was tem-

Jan 1, 1967

By Wing L. Lew

LEAD, silver, and zinc, in quantities of economic importance are found in the Aguilar ore. The lead occurs as galena and the zinc as sphalerite and marmatite. The silver occurrence is more complex. Pyrargyrite and proustite have been identified only occasionally. Galena is silver-bearing and so are the various other sulphides, though to a lesser degree. The principal nonsulphide gangue minerals include calcite, quartz, wollastonite, rhodonite, and various calcium, aluminum, and iron silicates. A typical analysis of, the ore is: Pb, 11.2; Zn, 12.3; Fe, 4.4; Mn, 2.2; Cu, 0.15; S, 11.2; CaO. 7.6; "Insoluble," 50.1 percent; and Ag 263 g per ton. Crushing is done in three stages. The primary crusher at the mine is set to deliver about a 5-in. product. The crushed ore is then transported to the secondary crushing plant by a 3-km cableway where the general practice is to hold the product of the first Symons to about 1 in., and of the second Symons to about 1/2 in. When the ore is soft and rich the product of the first Symons is by-passed directly to the ball-mill bins. Dust control in the secondary crushing plant is effected by enclosing the dust-emanating sources and exhausting the enclosures with a Clarage No. 18 fan driven at 1300 rpm by a 20-hp motor. Use is also made of spray water at one or two critical places. Every effort is made to keep the air free from visible dust. Dust counts taken at the working places show dust concentrations of less than six million particles ner cu ft. The composition

Jan 1, 1947

By H. L. Luo

SYSTEMATIC investigations of the Fe-Ga system have been reported only very recently."&apos; The maximum solid solubility of gallium in a-Fe is nearly 50 at. pet at 900°C. However, from Ref. 1, it was not evident that the a-Fe solid solution could be retained successfully beyond 17.38 at. pet Ga by solid-state quenching. The present investigation reports the retention of a-Fe solid solution at 50 at. pet Ga by quenching from the liquid. An ordered phase of CsC1-type B2) structure was observed between 34 - 50 at. pet Ga. Since it was considered very probable that the CsCl structure would occur at high temperatures, alloy filings were also quenched by the conventional quenching technique. In addition to the CsCl phase, a second ordered phase of BiF3-type (DO3) structure was found in the range 25 - 30 at. pet Ga. Alloys were prepared from elements of 99.99 pet purity or better, induction-melted in alumina crucibles under hydrogen. After reweighing, the alloy ingots were cast into wires under argon. From the weight losses and the reproducibility of the lattice parameters, compositions were estimated to be within ± 0.3 at. pet. The techniques of quenching from the liquid state were reported before.3-5 Prior to quenching, a small section of cast alloy wire (30 - 50 mg) was melted in an alumina insert under an argon flow. The quenched specimens were in the form of flakes, typically 1 sq mm in area. For solid-state quenching, alloy filings passing 200 mesh were sealed under a reduced helium atmosphere in a thin-bottom silica tube. After heating for -30 min at the desired temperature in a vertical furnace, the sealed tube was rapidly dropped into ice water so that the thin bottom of the tube broke on impact. The quenched alloy filings were always ferromagnetic and could be collected by a magnet. X-ray diffraction patterns were taken with iron-filtered cobalt radiation in a 114.6-mm-diam Debye-Scherrer camera. High-angle lines were always resolved and spacings were computed relative to ? (CoKa1) = 1.78890Å. Lattice parameters were obtained from extrapolation against the Nelson-Riley function. The variation of lattice parameter with composition is shown in Fig. 1, together with the data of Dasarathy and Hume-Rothery.1 Uncertainties in the present measurements are within the dimensions of the symbols or indicated by error bars. X-ray diffraction lines indicative of CsC1-type ordering could be clearly identified for alloys of 34 - 50 at. pet Ga quenched from the liquid and alloys of 35 - 41 at. pet Ga quenched from -975°C. Within experimental error, no difference in

Jan 1, 1968

By P. K. Foster, C. M. Payne, A. McNabb

Some measurements of the rate of loss of hydrogen from cylinders of iron and steel are analyzed in terrns of a trapping theory. The apeement is encozcraging and gives rise to estimates for the density of traps in the metal. Furthermore, the results appear to be consistent with the assumption that the diffusion constant D for the steady flux of hydrogen through iron lies on the extrapolation of the linear curve (log D us 1/T) obtained at higher temperatures. IT has been shown in a previous paper1 that the literature relating to hydrogen diffusion in steels contains sufficient evidence to invalidate the application of Fick&apos;s laws to hydrogen diffusion in steel at low temperatures. To account for the deviations from simple diffusion behavior a trapping theory was developed, and the fundamental laws modified accordingly. In the present paper, new data on the evolution of hydrogen from long cylinders are presented, and together with some published data are examined in terms of the requirements of the theory.&apos; Properties of solutions of the basic equations derived previously1 and appropriate to the present experimental conditions are outlined in the first section as they determine the form in which the experimental data must be presented for a comparison of theory and experiment. Two new solutions not previously reported have been derived as described in the Appendix. The experiments and treatment of data are then given in Sections 2 and 3 and the results are discussed in Section 4. 1) THEORETICAL CONSIDERATIONS The hydrogen-diffusion theory1 applied in this paper postulated that lattice-dissolved hydrogen at a concentration C diffuses randomly in accordance with Fick&apos;s first law, but that the lattice contains a uniform density N of traps which can capture and delay hydrogen atoms. kC atoms are trapped per second in a region containing unit concentration of vacant traps while p are released per second in a region containing unit concentration of occupied traps. On this basis it was shown (Eq. [30] of Ref. 1) that, when local changes in the lattice-dissolved concentration C are slow enough for equilibrium to be established with the traps, the system is represented in dimensionless form by the equation:

Jan 1, 1965

By Nev A. Gokcen

PUBLISHED data on the solubility of nitrogen in liquid manganese are widely discordant. Ochsenfeldl observed that a Mn-N alloy with 3.6 pct N consisted of solid and liquid phases at 1260°C but at 13 00° to 1340°C the solid phase disappeared and the liquid contained 0.1 to 0.13 pct N. In contrast Saito&apos;s measurements at approxhately 1400°C yielded an average value of 0.97 pct N. Sieverts3 volumetric method is not suitable for the determination of gas solubility in manganese because the liquid metal vaporizes readily and condenses in the cooler parts of the apparatus where it can absorb large amounts of nitrogen. The purpose of this investigation was, therefore, to determine the solubility of nitrogen in manganese accurately in the range of 1273" to 1500°C. The experimental method used in this study was as follows. Electrolytic manganese of 99.9 pct purity, and weighing 50 g, was melted in a pure alumina crucible placed in a tubular furnace. Temperature was measured with a Pt-10 pct Rh thermocouple and controlled within ± 5°C. Pure nitrogen was bubbled through the melt for 1 hr at constant temperature and then a sample was taken by sucking the metal into a 2-mm clear silica tube and immediately quenching in water. The sample was broken into small pieces and found to be sound to the naked eye. The entire sample was then dissolved and the resulting solution was analyzed for nitrogen by the well-known solution-distillation method. The experimental results are presented in Fig. 1. All the nitrogen solubility data refer to 1 atm of nitrogen pressure; hence the observed values are slightly corrected by using Sieverts&apos; equation when the atmospheric pressure differed more than 5 mm Hg from 760 mm, and further, in the case of the runs at 1500°C the correction arising from the presence of Mn vapor has been made. (Cf. Eq. [3]). Scattering in the data is within ±5°C in temperature measurement and ±3 pct in total nitrogen analysis. The straight line in Fig. 1, correlating the percentage of dissolved nitrogen, %N, with the inverse temperature in OK-1, can be expressed by:

Jan 1, 1962

By John P. Hager, Adolfo R. Zambrano

The thermodynamics properties of the liquid Pb-Ag-Au system have been determined from galvanic cell measurments five pseudobinary systems of fixed XAg/XAu ratio. The galvanic cell employed a molten PbO-SiO2 electrolyte and was operated between 775" and 1030°C. The thermodynamics properties of lead were obtained directly from the experimental data, whereas the excess integral molar properties and the activities of silver and gold were obtained by means of the Gibbs-Duhem equation. The calculated proper -ties of the Ag-Au system agree well with the published properties. The liquidus surface of the Pb-Ag-Au system has been determined at 1200°K. In Parts I&apos; and 11&apos; the galvanic cell Mo,Pb(l)|(PbO-SiO2)(l) SiO2(s) |Pb(alloy)(l), Mo [I] was used to determine the thermodynamics properties of the Pb-Ag and Pb-Au systems, respectively. The use of cell I, in which a molten oxide electrolyte is employed rather than the classical fused-salt electrolyte, was shown to result in more accurate ther-modynamic measurements in that the displacement reaction at the alloy-electrolyte interface (l/Zb)B(alloy) + (I/za)A+Za = (1/Za)A(alloy) + (1/Zb)B+zb [1] was minimized. For the Pb-Ag system, where B = Ag(l), the error in the activity of lead (aPd) re-sulting from Reaction [I] was calculated to be less than 1 pct.l For the Pb-Au system the error in aPb was calculated to be less than 0.5 pct.2 Also, electro-transport measurements on the PbO-SiO2 electro-lyte1,2 established that the conduction was entirely ionic and that lead was present only in the divalent state. In the present study the experimental technique is extended to an investigation of the thermodynamic properties of the liquid Pb-Ag-Au system. Five pseudobinary systems of fixed X Ag/XAu ratio were selected. The experimental results were then combined with the previous measurements on the component binary systems Pb-Ag1 and Pb-AU2 to calculate, by means of the Gibbs-Duhem equation, the molar and partial molar mixing properties for the liquid ternary solutions.

Jan 1, 1970

By M. Herman, G. J. London

A series of dilute Fe-Co alloy specimens were prepared from zone-melted iron. The iron used was given ten floating-zone passes in purified hydrogen. Alloying was accomplished by plating a gradient of metallic cobalt onto a 1/4-in.-diam bar of material from a pure plating bath. The deposit was radioac-tively tagged by a small addition of co80 to the plating bath. Alloy homogenization was accomplished by passing an additional floating-zone pass through the plated section of the bar. The final concentration gradient achieved was from residual to about 1000 ppm (0.1 pct) added Co. After homogenization, the bar was swaged to 0.057 in. wire and cut to tensile specimen blanks 1-1/16 in. long. Each specimen was counted in a well-type scintillation counter. The concentration of cobalt was computed by combining the normalized specimen counting rate with the known weight (measured to the nearest 0.01 mg) of cobalt added during plating. Fig. 1 shows the cobalt gradient obtained, exclusive of residual cobalt which was estimated by neutron-activation analysis to be approximately 50 ppm. Results of recrystallization for 1 hr at 750°, 700°, 650°, and 600°C with specimens containing 40, 170, 470, and 870 ppm added Co are shown in Figs. 2, 3, 4, and 5. Recrystallization was accomplished in a dynamic vacuum of less than 1 x 10 torr. The specimens were mounted and polished parallel to the swaging direction. The recrystallization re- sults show a radical coarsening of structure with increasing cobalt concentration. The coarsest structure obtained was at the 470-ppm level (at all recrystallization temperatures) with a slight grain refinement at the 870-ppm level shown. A more normal solid-solution effect is seen in Fig. 5 (60OoC, 1 hr) where the only specimen completely recrystallized is the one containing 47 ppm Co. Since the possibility existed that some atmosphere-contamination phenomenon might be responsible for the observed results the recrystallization series was repeated using sealed quartz capsules. These were filled with purified argon which was gettered during heat treatment on metallic titanium included in the capsules. The results obtained were essentially the same as shown. Another identical alloy was made using a 5-mm-diam bar of Westinghouse Puron. Specimen preparation and recrystallization treatment were also as described. The results obtained (see Table I) were somewhat erratic. Some coarsening in structure does appear at the highest cobalt concentrations. There is no apparent relationship between the coarse, columnar, directionally solidified structure of the as-melted bars and any of the recrystallized grain structures observed, all of which are much finer than the former structure. A recrystallization process can be most simply described as one of homogeneous nucleation of recrystallized grains in a cold-worked matrix followed by more or less uniform growth of the strain free grain into the cold-worked material.&apos; An impurity element may influence one or both of these processes. Aust and utter&apos; have shown that impurities strongly influence grain boundary mobility in lead. Moreover, definite recrystallization textures are produced by small concentrations of tin,3 while additions of silver or gold4 produce large boundary in-

Jan 1, 1964

By Marsh S. Watson, Robert A. Howard

An electrolytic model study has been made of the relative productivity of perforated casing. The results show that the relative productivity is roughly one-half to two-thirds of that predicted from point-sink calculations. This discrepancy can he satisfactorily explained on the basis of casing interference. The relative productivity was found to he insensitive to spatial arrangement of the perforations. INTRODUCTION The relative productivity of wells with perforated casing has been treated mathematically by Muskat.&apos; In his treatment the perforations are represented by small spheres, thus enabling him to approximate the flow by that to point sinks. By neglecting interference between the flow into the various spheres representing the perforations, one may deduce the approximate formula Qu 1/rp - 1/rw • (1) I + b/2n _________ 2n In re/rw Where Q is the flow rate into spherical sinks located in the median plane of the sand laver Qo is the open hole flow rate (uncased hole of radius rw) b is the thickness of sand layer under consideration (always one foot in this paper) n is the number of spherical sinks in the sand layer rp is the radius of a spherical sink rw is the well radius re is the drainage radius of the well. Fig. 1 is a comparison of values calculated from equation (1) with Mus-kat&apos;s values. It will be observed that the approximate values are in error by less than 21/2 per cent. The disturbing effect of the casing on the flow line; is not explicitly taken into account by Muskat, but he reasons that interference between the flow pat- terns from the various point sinks will result in a negligible flow in the region within the well radius, and that no appreciable error will result from not explicitly excluding flow from this region. This line of reasoning is open to question, however, as the flow impedance, in the immediate vicinity of the well, of a hemispherical cavity in contact with an opening in the casing is twice that of a spherical sink in a continuous sand layer. There will be a difference, moveover, between the flow into a hemispherical cavity and that into a circular hole in the casing. .This latter difference can best be appreciated by comparing the flow from a semi-infinite medium into a hemispherical cavity at the edge of the medium with the flow out of a circular opening of the same radius as the hemisphere and which is also on the edge of the medium. In this case2 ------= —2, showing that the difference in flow rates into differently shaped cavities is quite appreciable. It was the purpose of the investigation reported here to determine to what extent flow into circular perforations in casing departs from the values calculated by Muskat. For this purpose an electrolytic model was used.

Jan 1, 1950

By T. W. Watson, D. R. Flynn

Measurements have been made of the thermal conductivity and electrical resistivity of two specimens of 0.005-cm (2-mil) Be-Cufoil over the temperature range -140° to +200°C. The thermal cmductivity of Be-Cu Alloy 125 was found to be significantly higher than that of Alloy 25, which had a higher impurity content. The thermal conductivity and electrical resistivity values obtained were in concordance with the Smith and Palmer relation for copper alloys which states that ? = 0.0239 (T/p) + 0.075 where A is thermal conductivity (W cm-&apos; deg-&apos;), p is electrical resistivity (microhm-cm), and T is absolute temperature (OK) , indicating that the Smith and Palmer relation can be used to predict the thermal conductivity of this type of Be-Cu alloy over the temperature range -140" to + 200°C to within an accuracy suitable for most engineering applications. THE National Aeronautics and Space Administration, Goddard Space Flight Center, Greenbelt, Md., requested the National Bureau of Standards to measure the thermal conductivity of 0.005-cm Be-Cu foil which was to be used to fabricate antennae for use in Radio Astronomy Explorer space craft. It was felt that measurements had to be made on the actual antenna material, rather than on a thick bar of similar material, since the cold-working and age-hardening properties of the alloy would make it very difficult, if not impossible, to place a thick bar in the same metallurgical state as the 0.005-cm material. As discussed in Section 111 of this paper, the thermal conductivity of many copper alloys can be computed rather accurately from electrical resistivity data using an empirical relation developed by Smith and palmer&apos; and further confirmed by owell.&apos; However, there has been very little work done on correlating thermal and electrical conductivities on Be-Cu alloys. We were reluctant to assume the validity of the Smith-Palmer relation on Be-Cu alloys because of the very large mass difference between beryllium and copper. As will be seen, however, our experimental results indicate that the Smith-Palmer relation is valid for the two alloys on which we made measurements. I) DESCRIPTION OF SAMPLES Samples of Be-Cu Alloy 25 and Be-Cu Alloy 125 were furnished to NBS in the form of rolled strip material, 0.005 cm thick by 5.08 cm wide. The Be-Cu Alloy 25 was purchased by NBS in the form of a continuous length of flat strip. The Be-Cu Alloy 125 was furnished to NBS by Goddard Space Flight Center in the form of two lengths of strip that had been formed into "tubing" of about 1.2 cm diam, with one side of the strip overlapping (but not connected to) the other side by about 90 deg. This "tubing" had been opened up to be flat and then rolled onto spools of about 3 cm diam. When unrolled, the strip would spring back into tubular form. Samples of both of the Be-Cu alloys were chemically analyzed by the NBS Analytical Chemistry Division. The beryllium content was determined quantitatively by the phosphate-gravimetric method to be 1.83 wt pct for Alloy 25 and 1.80 wt pct for Alloy 125. The two alloys were also examined by a general qualitative spectrochemical method for metallic elements. The Alloy 25 specimen contained between 0.1 and 1.0 wt pct Al, Co, Fe, and Si and between 0.01 and 0.1 pct Ni. The Alloy 125 specimen contained between 0.1 and 1.0 pct Co and between 0.01 and 0.1 pct Al, Fe, and Si. Other metallic elements present were in amounts less than 0.01 pct. The NBS Engineering Metallurgy Section measured the tensile strength, elongation, and Knoop hardness of both the Be-Cu Alloy 25 and the Be-Cu Alloy 125. These are given in Table I. The NBS Engineering Metallurgy Section also prepared photomicrographs of the two alloys; these are shown in Figs. 1 and 2.

Jan 1, 1969

By R. J. Charles

SEARCH for a consistent theory to explain the relationship between energy input and size reduction in a comminution process has accumulated, over the years, an enormous amount of plant and laboratory data. Although some correlation of these data has been possible for purposes of engineering design and for the advancement of research in fracture, there is still great need of a means of predicting behaviour of a solid when it is reduced in size by mechanical forces. The best known hypotheses proposed to describe the energy-size reduction relationships in crushing and grinding stem from a common origin. The present article analyzes problems of comminution in the light of the precepts of this origin. Its object is to reconcile points of difference between these well known hypotheses and to present relationships more widely applicable to comminution studies. Theoretical Considerations: Most existing relationships between energy and size reduction of a brittle solid stem from a single, simple, empirical proposition.&apos; Although this proposition can be demonstrated by observation and experiment, no theoretical derivation is yet possible. Mathematically, the proposition may be stated as follows: dE = -Cdx/xn [1] where dE = infinitesimal energy change, C = a constant, dx = infinitesimal size change, x = object size, and n = a constant. Eq. 1 states that the energy required to make a small change in the size of an object is proportional to the size change and inversely proportional to the object size to some power n. No stipulations are placed on the exponent n in either magnitude or sign. In 1867 Rittinger2 postulated that the energy required for size reduction of a solid would be proportional to the new surface area created during the size reduction. As far as can be determined there is as yet no physical basis for Rittinger&apos;s hypothesis. Rittinger&apos;s hypothesis can be stated mathematically as follows: E, = K(oa-a-0 . [2] Er = energy input per unit volume, K = a constant, <ti = initial specific surface, and o2 = final specific surface. In the size reduction of particles of size x, to particles of another smaller size, x2, Eq. 2 becomes the well known relation: ET = K&apos; {l/x2-l/xx) [3] where K&apos; is a constant. Eq. 3 may be arrived at from the proposition given in Eq. 1 by integrating and by assigning a value of 2 to the exponent n. J dE = J - C dx/x2 E = Kt (1x - 1x) where K&apos; = C. In 1885 Kick3 proposed the theory that equivalent amounts of energy should result in equivalent geometrical changes in the sizes of the pieces of a solid. For example, if one unit of energy reduced a number of equal-sized particles to particles of one half the size, then the same amount of energy applied to the particles resulting from the first test should result again in a size reduction of one half or a final size one quarter the original size. The Kick concept may be expressed as follows: Ek = K" log x1/x2 [41 K" = a constant and E, = energy per unit volume. The expression for Kick&apos;s law may be arrived at by again integrating Eq. 1 and in this case assigning a value of 1 to the exponent n. dE= J - C dx/x Eh = - C In {x/x2) = K" log (x,/x,) where K" = 2.3 C. Application of Kick&apos;s and Rittinger&apos;s laws to comminution has met with varied success. Gross and Zimmerley4 and Piret5 have shown that Rittinger&apos;s equation applies under certain conditions of experimentation. Walker and Shaw6 express the belief that in metal turning and shaping and in grinding of both metals and minerals the production of very fine particles (less than lP) follows Kick&apos;s hypothesis, whereas Rittinger&apos;s concept is valid for the size reduction of coarse particles. For the practical case of crushing and grinding, however, neither of the above hypotheses has received general acceptance. Bond&apos; has lately proposed that since neither Kick&apos;s nor Rittinger&apos;s hypotheses seem correct for plant design work, an energy-size reduction relationship somewhere between the two would be more applicable. The fundamental statement of Bond&apos;s work

Jan 1, 1958

By R. W. Heckel, J. H. Smith, H. W. Paxton

The habit plane of proeutectoid cementute precipitated from austenite was studied using a two-trace surface (analysis coupled with a direct X-ray back-reflection determination of the orientatzon of the parent Matrix. The results were the same as those obtained by earlier workers1 - 2 who used Indirect methods to determine the orientation of the matrix. The data indicate that a range or habit planes in the parent phase is operative which is not in agreement with accepted existing theories. In 1933, Mehl, Barrett. and smith1 obtained a non-unique result in a study of the Widmanstätten habit plane of proeutectoid cementite. Their data are shown in Fig. 1. These results were verified by Greninger and Troiano,2 although they did not pre-sent any data. Both of these investigations relied upon indirect methods of determination of the matrix lattice orientation. Theoretical considerations of Widmanstatten precipitates have usually proved or assumed an orientation relationship between matrix and precipitate; the habit plane expressed relative to the matrix is generally unique, although it may be irrational.3 Opinion is divided as to the factors influencing the choice of habit plane. Mehl4 felt that a nucleus which was plate shaped to minimize strain energy and surface energy during its formation could maintain this shape by the point effect of diffusion; the

Jan 1, 1961

By D. G. Westlake

Polycrystalline tensile specimens of various Zr-0-H alloys have been tested at 298°, 178°, and 77°K. Solute oxygen and hydride precipitates in quenched alloys made individual contributions to the yield strength at 0.2 pct strain which combined to produce a resultant strength increment, a,., Ductility changes which were ohserved can he interpreted in terms of the various oxygen and hydrogen concentrations, testing tem -peratures, and dispositions of the hydride. ADDITIONS of oxygen in solid solution were known to increase the yield and tensile strengths of polycrystalline zirconium as early as 1951.&apos; More recently, the critical resolved shear stress (CRSS) for prism slip in zirconium single crystals was also shown to be affected by the solute oxygen impurity.&apos; This latter work also demonstrated that large increments of strength could be contributed by the finely dispersed zirconium hydride precipitates that are present in quenched Zr-H alloys.3 It was concluded that the combined strengthening due to alloying could be expressed by where to is the increase in the CRSS due to solute oxygen alone and TH is the increase due to finely dispersed hydride precipitates. Eq. [I] is analogous to one used to express the combined strengthening effects of work hardening and neutron radiation damage.4 Eq. [1] was verified only indirectly and for only small amounts of the impurities—up to 0.14 at. pct 0 and 0.63 at. pct H. The present investigation was undertaken to obtain a more direct verification of the validity of the form of Eq. [1] for this system and also to determine the combined effects of oxygen and finely dispersed hydride precipitates on the tensile strength and ductility of polycrystalline zirconium. EXPERIMENTAL PROCEDURE Tensile specimens were machined from the same rolled billet of Kroll zirconium used in the earlier study.&apos; These measured 38 by 4.7 by 0.5 mm and had 10-mm gage lengths which were 2.8 by 0.5 mm. Each specimen was ß-annealed in vacuo at 1173°K for 15.5 hr and a-annealed at 1073°K for 4 hr to D. G. WESTLAKE, Member AIME, is Associate Metal l ur-gist, Metallurgy Division, Argonne National Laboratory, Argonne, III. Manuscript submitted July 17, 1964. IMD______________ give an equiaxed structure with grain diameters averaging 0.06 mm. Oxygen was added by allowing the metal to react with a known quantity of oxygen during the 0 anneal and known quantities of hydrogen were added during the a anneal. Each alloy was encapsulated in Pyrex under vacuum, annealed at 873°K for 4 hr, quenched into ice water, and polished by immersion in a solution of 46.75 vol pct H2O, 46.75 vol pct concentrated HNO3, and 6.5 vol pct HF (49 pct) at 298°K. Special heat treatments given to a few specimens are described in the results below. Tensile tests were done on an Instron machine and were begun within 20 min after the quench, except where specified otherwise. Tests at 298°K were in air, at 178°K in acetone, and at 77°K in liquid nitrogen. All tests were at a strain rate of 8x sec-1. RESULTS AND DISCUSSION Yield Stress at 298°K. The compositions of alloys and the corresponding yield stresses (0.2 pct strain) are given in Table I. A plot of the yield stresses of the oxygen alloys, A, B, C, and D, indicates that varies linearly with CO1/2, where Co is the oxygen concentration, Fig. 1. This is in accord with Fleischer&apos;s6 theory for solution strengthening if the oxygen atoms do not cluster, or the cluster size remains constant with increasing oxygen concentration. In Fig. 1, it appears that if one could prepare some oxygen-free zirconium its yield stress would be very low. Therefore, we shall assume that for the oxygen alloys is equivalent to O0, the strength increment contributed by the presence of oxygen. The relationship between0.2and Co is expressed by 0.2 = 31.3 CO1/2, when the yield stress is in kg per sq mm and the concentration is in at. pct. Each of the hydrogen alloys, Al, A2, A3, and A4, contained 0.081 at. pct 0 as an impurity. In Fig. 1, it appears that this small amount of oxygen makes a significant contribution to the strength which cannot be ignored when we evaluate the contribution of the finely dispersed hydride. Let us assume the validity of the following equation: a0.2 = (a2o+a2R)1/2 [2] which is analogous to Eq. [I] for single crystals, and calculate values of UH for the hydrogen alloys by using the experimental values of 0.2 and o (0.081 at. pct) = 8.9 kg per sq mm. For 0.36 at. pct H, oH = 6.47; for 0.72 at. pct H, OH = 11.30; for 2.16 at. pct H, OH = 19.4; and for 3.60 at. pct H,

Jan 1, 1965

By F. W. Bremmer

THE drilling of constantly deeper oil wells has made it imperative that the manufacturer of casing be ever searching for new methods and new materials to meet the increased demands. One phase of this search constitutes the subject matter of the present paper. Starting from the original Drake well in 1859, the depth of wells has advanced by leaps and bounds, until today depths of 6000 to 8000 ft. are common. Conditions point to the necessity for drilling to greater depths, and it is not at all improbable that much of the future oil produc- FIG. 1.-TREND IN DEPTH OF OIL WELLS. tion will be from, wells around 10,000 ft. deep. The trend to deeper wells is nicely illustrated by the curve in Fig. 1.1 The points plotted are the maximum depths attained in the various pools which were at their peak for the years given. It may be of some interest to observe the change that has taken place in the tensile and collapsing stresses over the last 25 years. The pressure at the bottom of a 900-ft. hole, equivalent to the pressure caused by an equal column of water, is only 450 lb. per sq. in. The equivalent

Jan 1, 1930

By G. D. Rieck, M. M. P. Janssen

Chemical diffusion coefficients and heats of activation for diffusion in the NizAh fy), NiAl (6), and Ni3A1 (E) intermetallic phases and the solid solution of aluminum in nickel (( phase) were calculated from layer growth experiments. No finite diffusion coefficient for the NiAl3 ((3) inter metallic phase could be calculated. The values of the diffusion coefficients are dependent both on the method of calculation and the type of diffusion couple. The heat of activation for diffusion in the y phase was found to be 47 kcal per mole in the temperature range oj 428" to 610°C. Heats of activation of 41, 12, and 48 kcal per mole were found for diffusion in the 6, E, and ( phases, respectively , in the temperature range of 655" to 1000°C. Experiments with markers in the diffusion zone demonstrate a very pronounced Kirkendall effect. It appears that only aluminum atoms take an active part in the diffusion process during the formation of the 0 and y phases at temperatures of about 600°C. During the formation of the 6, E, and < phases at higher temperatures only nickel atoms are moving. It is suggested that the great stability of the intermetallic compounds in the Ni-A1 system governs the Kirkendall effect. SOME factors controlling layer growth during inter-diffusion in the Ni-A1 system (phase diagram, see Fig. 1) were studied by Castleman and Seig1e.l&apos;~ They found the NiA1, ((3) and NiAl3 (y) intermetallic compounds to appear in the diffusion zone of Ni-A1 couples at annealing temperatures of 400" to 625°C; the NiAl (6) and Ni3A1 (E) intermetallic compounds appeared in y-Ni couples at annealing temperatures of 800" to 1050°C. These authors carefully examined metallographically Ni-A1 couples after 340 hr annealing at 600°C. Besides the (3 and y phases they found very thin layers of the 6 and E phases. ~n~erman~ and Castleman and Froot4 observed a much more rapid growth of the 5 and E phases at 600°C in Ni-A1 couples in case a crack was present at the /3-A1 interface. Numerous layer thickness measurements carried out by Castleman and Seigle on the y phase prove that the layer growth of this phase obeys the parabolic law after a certain transient period. From this they concluded that the layer growth of the y phase is controlled by volume diffusion. The growth of the 13, 6, and E phases appeared to be volume-diffusion-controlled also. The authors estimated that at 600°C and at atmospheric pressure Dp was 1.8 x lo-"ll sq cm per sec, D, 9.1 x 10" ™ sq cm per sec, Qp 27 kcal per mole, and Qy 31 kcal per mole. The present work was carried out to obtain more quantitative data about the kinetics of growth of the phases of the Ni-A1 system and the reactions that occur during the formation of these phases. Because in this system the diffusion process results in the formation of several distinct intermetallic compounds, the current term reaction diffusion is used in the title of this paper. In order to obtain layers of the fl phase compound of uniform thickness, a new technique for preparing diffusion couples was developed. The kinetics of growth of the y phase in 6-Al, E-Al, and Ni-A1 diffusion couples was studied at different temperatures. The kinetics of growth of the 6, c, and ( phases in Ni-y, Ni-6, and Ni-c diffusion couples was also studied at different temperatures. The calculation of the diffusion coefficients Dp and Dy by Castleman and Seigle are critically considered in this paper; by means of a revised method of calculation more reliable val-ues of , and Dg were found. These values are in good agreement with the values of the diffusion coefficients obtained by the method of Boltzmann-Matano. From the temperature dependence of the diffusion coefficients the heats of activation for diffusion were calculated by means of an Arrhenius-type equation. The investigation of the Kirkendall effect has been used to obtain information about the ratio of the intrinsic diffusion coefficients of the separate atoms5 and the mechanism of diffusion. Moreover porosity as a result of a distinct Kirkendall effect would be of practical importance in connection with the bonding of diffusion coatings. The analyses of the diffusion couples were carried out by metallographic methods. The values of the concentrations at the phase boundaries and the concentration profile in each of the phases, which are needed for the calculation of diffusion coefficients, were obtained by electron-pro be X-ray microanalysis. EXPERIMENTAL PROCEDURE A) Materials for Diffusion Couples. The intermetallic compounds 6 (50 at. pct Ni) and E (74 at. pct Ni) were prepared from the pure metals by high-frequency induction melting in argon atmosphere. Use was made of aluminum wire (99.99 wt pct Al) and nickel sheet (99.95 wt pct Ni). The 6 and E phase melts and the nickel shiet (thickness 0.1 and 0.5 mm) used for preparing diffusion couples were annealed for 64 hr at 1200°~ for homogenization and grain coarsening (final crystal size 1 to 3 mm). composition and homogeneity of the intermetallic compounds were checked by mi-crohardness measurements and X-ray diffraction. From the 6 and E phase melts discs of 0.5 mm thickness were prepared by means of a water-cooled rotat-

Jan 1, 1968