Search Documents

By P. S. Roller

P. S. ROLLER,* College Park, Md.-The author introduces the idea of a grind limit, or lower size limit of the particles formed directly by grinding, and this is stipulated to be 0.70 micron. I wonder where all the particles come from that lie between the resolving power of the microscope and 0.7 micron, which are seen under a microscope in examining a sample of grind? There is also introduced the idea of excess fines; this is all right in itself but the author's quantitative definition seems to be virtually impossible to justify, and, moreover, is obscure. Just why E, the excess fines, should be defined as the distance between an F line and a [?] line is hard to follow, and just where the distance should be measured does not appear to be stated, as it should be since the F and [?] lines intersect and the distance between will depend upon where that distance is measured. The basis of the presentation is the relation of log weight against log size that was "discovered" by Schuhmann. While this relation is implied to fit, the only evidence that I can see is a slightly irregular line in Fig. I between 2 per cent and 18 per cent by weight and then a precipitous drop at 18 per cent away from any sort of straight line. Some of the conclusions are rather startling. It is stated that E, the excess fines, remains constant for all the comminution products of any particular material, except that it may decrease for some coarse crusher products. No reason is given for this assertion. One would, on the contrary, expect that excess fines, if such a concept is permissible, would generally vary with different conditions instead of being constant. Again, "as much net energy is required to reduce a ton of quartz to particles of the grind-limit diameter (namely, of course, 0.7 micron) by grinding as is required to melt the quartz," This would require many data to prove, but no proof is given; any considera¬tion would reveal the impossibility of such a result. Still further, ore samples "ground to widely different sizes, and with different circulating loads and different size distributions or product, show the same production of new surface area per revolution of the mill, when this grind limit is used." This statement is confuted by ordinary experience, which teaches that grinding efficiency varies generally with the condition of material. For example, the cushioning effect of fines is widely accepted, while the inefficiency of a ball mill when the feed is too coarse is also well known. R. SCHUHMANN, JR.,* Cambridge, Mass.¬Contrary to Mr. Bond's stand, there is a great deal of miscellaneous experimental evidence that comminution does not stop at around 0.7 micron, nor does it cease consuming energy at that point. Comminution is in fact recognized as an important method of producing colloids.5 Mechanical comminution methods were used by von Weimarn6 to produce colloidal dispersions of a variety of solids, including the minerals chalcocite, galena, hematite, stibnite. Utzino7 gives average particle sizes ranging from 0.02 to 0.09 micron for colloids produced by mechanical comminution of several crystalline solids. Andreasen and Berg8 have ground

Jan 1, 1941

By E. D. Glass, R. C. Hessing

The volume of gas in storage reservoirs may be compuled from estimates of hydrocarbon pore volume and gas density. However, both are dificult 10 estimate accurrately. Further, no adequate method has been presented for estimating reservoir performance during operation for Volumetric gas-storage reservoirs. Normally, gas-storage reservoirs exhibit small pressure gradients in the area conr(zilliitg the storage wells, even under conditions of maximum injection or withdrawal. A zone of low permeability usally surrounds, and is in communication with, tile permeable zone which causes the storage reservoir to exhibit a definite pressure lag, as evidenced by lack of pressure stabilization after extended periods of shut-in. An analysis is presented for determining pore volume and gas in place, and for predicting future reservoir behavior of such reservoirs. The mathematical development is presented and a solution is shown for an operating storage reservoir. INTRODUCTION Depleted gas reservoirs arc often used for gas storage. The reservoir usually contains a permeable area in which numerous storage wells arc located. The permeable zone permits high deliverability and injectivity with relatively small pressure loss. However, pressure equilibrium may not be attained even after long periods of shut-in because of the fact that the permeable zone is in communication with an adjacent tight section. Most of these gas-storage reservoirs are sand lenses, such as ancient beaches or offshore bars. The areal geometry of these reservoirs is normally such that flow in the tight zone will be approximately linear, and the solutions presented here should apply. DISCUSSION For purposes of this paper, a gas-storage reservoir is considered to consist of a volume of permeable rock in communication with tight rock, all of which is surrounded by an impermeable barrier. Pressure gradients across the permeable zone are small due to low flow resistance and well density. Hence, the permeable zone is considered to be at a uniform pressure which is only a function of time as gas is injected or withdrawn. The tight zone is considered to be linear with uniform permeability and porosity, and the pressure in the tight zone is a function of both position and time. l'wo solutions for flow in this system are used to define the reservoir and determine its properties. The first solution is for an infinite linear tight zone in communication with the permeable zone. The expression for constant gas-withdrawal rate is Eqs. I and 2 ore superposed to account for varying rates of production or injection and are used to fit gas-storage field reservoir history. The solution for the infinite case, Eq. 1, is first applied to field data for early times following a shut-in period. Constants BDa and c are ohtained by a least-squares fit of pressure and rate data. After the pressure wave from gas injection or withdrawal reaches the external boundary of the tight zone, the infinite solution will no longer apply and pressure behavior should be more nearly represented by Eq. 2 for a gas reservoir in communication with a finite linear tight zone. The values of the constants found from curve-fitting Eq. 1 to the early time data may then be used in Eq. 2 that only the coefficient- L containing the unknown

By J. H. Jackson, P. D. Frost, C. H. Lorig, A. C. Loonam

THIS paper describes (1) the effect of sodium on the tensile ductility of magnesium-lithium base alloys, and (2) the precautions necessary to avoid sodium contamination. Effect of Sodium on Properties The harmful effect of sodium in magnesium-lithium base alloys has been recognized by other investigators.1,2 However, quantitative data on its effect have not hitherto been published. The effect of sodium on the ductility of an 87 pct Mg-9 pct Li-4 pct Zn alloy is shown in fig. 1. The sodium was introduced by adding sodium chloride to the flux, which consisted of 50 pct LiC1-50 pct KC1. Moderately high ductility was obtained at an indicated sodium content of about 0.05 pct. Actually, the value probably was less than 0.05 pct, be-cause analyses of the lithium, which were more accurate than those of the alloy, indicated that the highest ductility was obtained when its sodium content was 0.02 pct or less. In fact, in alloys having somewhat greater strength than this particular composition, the sodium content of the lithium must be under 0.01 pct to avoid embrittlement. The effect of sodium on the ductility of several alloys having fairly low strength levels is shown in fig. 2. These alloys were melted with a flux containing 75 pct LiC1-25 pct LiF, to which various quantities of sodium chloride had been added. Sodium had no significant effect in pure magnesium but markedly decreased the ductility of the magnesium-lithium-zinc alloy, even though the alloy contained only 3 pct Li. As the lithium content of the binary alloys was increased, the ductility of the sodium-free alloys also increased. However, the effect of sodium on ductility became progressively more pronounced. Sodium raised the minimum hot-rolling temperatures, particularly those of the high-strength alloys, but its effect, in this respect, was less critical than its effect on tensile ductility. For optimum cold roll-ability, a low-sodium content was required. No deleterious effect of sodium on the stability of properties at slightly elevated temperatures was found. Control of Sodium in the Alloys Optimum ductility is obtained in the magnesium-lithium base alloys when sodium-free materials are used. Commercial lithium was formerly the chief source of contamination. Now purified lithium is usually available. When lithium of ordinary purity is used, the effects of sodium can be minimized by using a flux consisting of 75 pct LiC1-25 pct LiF. Bubbling nitrogen through the melt also helps remove sodium. The damaging influence of sodium on the ductility of the magnesium-lithium base alloys was observed very early in the alloy development program. The discovery came about as a result of chemical analyses made on the lithium used in the alloys. Ductile alloys had been made from lithium containing about 0.2 pct Na, whereas the lithium used in the brittle alloys contained about 0.9 pct Na. The first magnesium-lithium alloys were melted under commercial fluxes ordinarily used in melting magnesium alloys. These fluxes were unsuited for scavenging the contaminating sodium, and the resulting alloys were not ductile. Next, a flux containing 50 pct KC1-50 pct LiCl was tried. This was somewhat more effective but still was not satisfactory. After a number of different experimental fluxes had been investigated, one containing 75 pct LiC1-25 pct LiF was adopted. This flux made it possible to obtain good ductility in the alloys. The

Jan 1, 1951

By H. W. R. Wardlaw

This paper shows how the cost of the drilling operation per 100-ft drilled can be related directly to bit weight, rotary speed and hydraulic horsepower variations. This enables optimum drilling conditions to be determined in a rapid and concise manner. The paper also shows how data for such calculations cun be obtained quickly and accurately by the "drill-off" method suggested by Lubinski. Examples are given to illustrate the principles and methods used. INTRODUCTION The problem of drilling oil wells at lowest cost and hazard needs little introduction, since it has been with us virtually from the earliest days of our industry. With better understanding of the effect of the various factors of bit weight, rotary speed and hydraulic jet action, however, the problem has recently taken on a more scientific aspect, and engineers or groups of engineers have been assigned to study many of the facets of this problem not immediately apparent to the general observer. Moreover, the availability of better equipment and materials has enabled certain practices to be adopted which were hitherto impractical or uneconomic. This has paved the way for further efforts to reduce footage costs, and so on. The improvements to be made in drilling technology and cost reduction are by no means exhausted, and we still frequently hear of new records being set and new materials and equipment becoming available. Some of the analytical methods used by the author to assist in promoting faster and more economic methods of drilling, particularly in Venezuela, form the basis for this descriptive study. STATEMENT OF THEORY AND DEFINITIONS In a highly competitive industry, it is no longer permissible to regard penetration rate as the sole criterion for drilling efficiency. Every new technique, however spectacular, must nowadays be viewed in the cold light of whether it will eventually lead to lower cost per foot drilled. In the exposition which follows, therefore, actual drilling rates will play only a secondary role, and attention will be focused on the more important aspect of cost per foot drilled. Changes in drilling factors such as weight and rotary speed will be studied in the light of how they directly affect the cost of operations. The advantage of this method is twofold—(1) cost increments are indicated directly; and (2) the relationships are, surprisingly enough, simpler than those relating to drilling rates. To achieve these advantages, it is necessary to introduce a new unit of measurement which, although unfamiliar, is readily understood. The unit in question will be referred to as the "actual drilling cost factor" and will be designated by the symbol h. The unit is defined as the number of rotating hours taken to penetrate 100 ft of formation. The unit can be obtained immediately from the actual drilling rate (r) in feet per hour; thus, This latter transformation is so easily effected that it may be convenient to use h as an international symbol for drilling cost, and to delay conversion to dollars, bolivars, pounds, etc, per foot until actually required (e.g., for fiscal reasons, etc.). In actual operations, h was found to vary in a relatively simple manner with increases in bit weight. The variation in h with rotary speed was found to be somewhat more complex. However, this also could be enormously simplified by introducing a further factor known as L or relative bit life. The factor L is also defined very simply by dividing 1,000 by the rotary speed (N) in revolutions per minute.

By H. R. Stahi

THE ore of the Southeast Missouri lead district consists essentially of galena in dolomitic limestone. The galena usually is in a very finely disseminated condition. The experience of recent years in the concentrating plants of the St. Joseph Lead Co. has shown that dry crushing carried to a point where about 90 per cent of the ore passes a 14-mesh screen liberates the larger portion of the galena, and minimizes further necessary reduction in rod or ball mills. Concentration practice has attempted to carry size reduction to a point where economic unlocking of the galena is obtained, without causing undue sliming of the very friable galena. While 60 per cent of the tonnage of the district is treated by the flotation process, only 40 per cent of the total concentrates is produced by that means, as efforts have always been made to keep the grade and tonnage of flotation feed as low as possible. The milling practice of the district for some years has recognized the generally accepted idea of better selective grinding in rod mills as compared with ball mills, with the attendant decrease in production of fines. The function of rod and ball mills in the various concentrating plants is the grinding of table middlings and coarse tailings that carry lead values too high to be discarded. The simplified flowsheet in Fig. I illustrates the place of grinding mills in the circuit. The arrangement of classifiers, tables and mills constitutes a closed circuit within somewhat flexible limits. In 1937 experiments were conducted in replacing the ball charge in a 6 by 4-ft. ball mill with rods 2 in. in diameter and 51 in. long. This step was taken with some hesitancy, as it was easy to picture a mass of tangled and crossed rods. However, no difficulties developed, and after a short trial period, six more ball mills were converted to rod mills. These mills are of the spur gear and pinion, belt-driven type, using in most cases a 75-hp. motor and having a 9 1/2 -in. open trunnion discharge. As ball mills, they carried a charge of 16,000 to 20,000 lb. of balls, 2 in. in original size. The feed to the mills was dewatered by dewatering wheels and introduced at 70 to 75 per cent solids, with a single scoop feeder. One change in mill design was found advisable when rods were introduced. The existing end liners had an outward flare of 3 ½ in. in 12 in. These were replaced by linings having vertical interior faces, in order to prevent undue end travel of the rods. This change gave an effective working length of 5 2 ¾ in. inside the mill. No alterations were made in the shell liners, which are of the smooth type, made of manganese steel 2 in. in original thickness. The rods at present in use are 51 in. long and 1 3/4in. in diameter. The composition has been varied at times; at present they contain 0.70 per cent carbon and 0.70 per cent manganese. Some breakage and etching occurs as the rods become small, but this is considered preferable to using softer rods, which may kink and cause tangling of the charge. The mill charges are renovated occasionally, broken pieces being removed and the charge brought up

Jan 1, 1945

LIBERIAN ore has been called the richest iron ore mined in the world, and thus far, the output from the Liberian Mining Co.'s Bomi Hills mine has lived up to its reputation. Iron content has been variously given as 68.47 pct to 68.825 pct. A recent United Nations survey gave it first place, surplanting the previously top-rated Swedish product. The first shipment of Liberian ore to the United States reached Baltimore in June, 1952 and went into blast furnaces at Republic Steel's Cleveland plant, where excellent results were reported. The physical make-up of the ore was hard, dense lumps, free of fines and with ideal size distribution. The Bomi hills mines were scheduled to be in full operation by the end of 1952. At the close of 1951, 200,000 tons of ore had been delivered to Baltimore. Liberian Mining Co. states that production is about 90,000 tons of ore per month at present. Between 1200 and 1500 men are employed at the mine site, with about 25 pct of them Americans. The Bomi Hill project was financed with the aid of the Export Import bank, which loaned Liberia Mining part of the sum needed to initiate and put the development into production. No definite expansion plans are in the making but extensive diamond drilling is being pursued. Original estimates of annual production for Bomi Hills ran in the neighborhood of 1.5 million tons, but the figure is no longer applicable. Liberia's potential as an iron producing country has been known for a long time, but the physical obstacles in the way of large scale production were such as to prohibit even contemplation of the project. Until recent years, the only route from the mine site to Monrovia was a foot trail. Once to the coast, over swamps and unbridged rivers, the ore was still useless because of the lack of loading facilities. Transportation Solved One problem was solved with the completion of the artificial harbor at Monrovia. Harbor construction was well under way when the mine project began in 1948. The island location of Monrovia also made the building of a bridge across the St. Paul River necessary. The 980 ft reinforced concrete structure, accommodating railroad and vehicular traffic, was opened in 1949. Meanwhile, preliminary work on railroad construction started. Grading was completed quickly. In April, 1951, the first train made the trip from Monrovia to the Bomi Hills. About half of the cost of the entire project went into railroad construction. More than 2200 men, often forced to resort to primitive methods, were engaged in clearing a path through the heavy forests between the mine and port. Swamps were drained and filled, and bridges, including the William V. S. Tubman, named for the Liberian President, were constructed. At the port, a Robbins Conveyor Belt Co., orehandling plant was installed. Ships can be loaded from the stockpile at the rate of 3000 tons per hr and vessels need spend only one day in port. Liberian Mining is operating under an 80 year lease, giving the company exclusive mining rights to all minerals, except gold, platinum, and diamonds, within a 40 mile radius of Bomi Hills. The contract provides a fixed royalty to the Liberian Government of five cents per ton of ore extracted, and a premium -based on the New York market price for Bessemer-grade pig iron. The entire project started because an American,

Jan 1, 1952

By F. Brons, W. C. Miller

Pressure information for use in material-balance calculations is obtained, where possible, from pressure build-up surveys in shut-in wells. Using proper extrapolation methods, static pressures are obtained which, by averaging, give the field static pressure. Often, particularly in fields with a large number of wells, only spot pressure readings are available. A graphical method is presented here which enables the determination of the corrections which must be applied to the spot readings to give an estinzate of the static pressure. Although the static pressures thus obtained are not so reliable as those from build-up surveys, they are more nearly correct than the spot readings themselves for use in material-balance calculations. INTRODUCTION The average static pressure in a bounded reservoir is determined from a number of individual-well pressure measurements. Where pressure build-up curves are available, extrapolation methods such as described in Ref. 3 may be used to determine the static pressure in the drainage area of each individual well. The possibility of using pressure build-up curves presupposes that each well being surveyed is closed in during a certain period and that, during this time, a pressure recorder is left suspended in the hole. For surveys in fields with a large number of wells, this is usually not feasible for economic and other reasons. Either a large number of pressure recorders would have to be used simultaneously or, using a limited number of recorders, the time interval needed to complete a field-wide survey would be too long. The data obtained for this incomplete coverage are insufficient to permit reliable averaging of the few calculated static pressures over the entire reservoir. In many such cases another type of pressure information is available, i.e., spot pressure readings. During this type of pressure survey all wells are closed-in simultaneously and, after a reasonable shut-in time (e.g., three days), pressure readings are taken by running the recorder in and out of each successive well. These surveys, which are less expensive, are often taken at regular intervals, say once or twice a year. In fact, such surveys are required by the state regulatory bodies in many portions of the united States. These spot readings are normally lower than the static (fully built-up) pressure due to the relatively short shut-in time. No direct means of extrapolating such a reading to the static pressure is available because no pressure build-up record is obtainable. Consequently, these data are inferior to pressure build-up information as a basis for estimating the stabilized reservoir pressure in the drainage area of a specific well. For the common case where abundant spot readings but only a few pressure build-up surveys are available, our studies suggest that it may be possible to design a program which will combine the two types of information and provide an estimate of the average reservoir pressure that is more reliable than that obtainable solely from either set of data. in time) gives a point marked p. The actual static pressure at the time the well was shut-in is marked by p.. The illustration shows that, even if a build-up curve is available and can be extrapolated, a correction

By Harold Margalin, Elmars Ence

M. H. Mueller, M. V. Nevitt, and H. W. Knott (Argonne National Laboratory)—We have reported1 the crystal structure of Ti2CU as body centered tetragonal with a. = 2.944A and co = 10.786A, six atoms per unit cell, and space groupl4/mmm(MoSi2 type). This cell is larger but definitely related to the unit cell reported by the authors. It permits definite positions for the titanium and copper atoms in the cell located as follows: 2 Cu (a) 0 0 0; 1/2 1/2 1/2 4 Ti (e) 0 0 z; 1/2 1/2 1/2 + z; 0 0z; 1/2 1/2 1/2- z where z = 0.341 The face centered tetragonal cell reported by the authors can be transformed to a small body centered tetragonal cell with flo (small bet) = ao(fct)v2 Both unit cells have the same CO. Neither the face centered tetragonal cell with four atoms per unit cell nor the small body centered tetragonal cell with its two atoms per unit cell lend themselves conveniently to an orderly arrangment of the 2 to 1 ratio of titanium to copper atoms. When the co of this small body centered cell is tripled, however, a large body centered tetragonal cell results with six atoms per unit cell as described above. We had also noted that the X-ray patterns of Ti2Cu indicated that the structure could be approximately described by the face centered tetragonal cell as reported by the authors; how- ever, there was evidence for a few additional very weak lines whose positions indicated a larger unit cell. Since the X-ray scattering amplitudes of titanium and copper are nearly equal, it is evident that it is difficult to distinguish their positions especially from X-ray powder patterns. It was therefore decided to look at a powder pattern obtained by neutron diffraction since the scattering amplitudes of the two elements are considerably different in magnitude and opposite in sign. The sample used for neutron diffraction was selected from a group of three alloys prepared by arc melting using crystal bar titanium of 99.92 pet nominal purity and 99.999 pet pure Cu. The alloys designated as A, B, and C had the following respective compositions: Ti-32.8 at. pet Cu, Ti-33.3 at. pet Cu, and Ti-33.8 at. pet Cu. On the basis of the melting losses incurred, it was established that the actual concentration of either component in any of the alloys probably did not deviate from the intended value by as much as 0.1 at. pet. The alloys were annealed for 5 days at 700°C in evacuated Vycor capsules and quenched in tap water. The metallographic structures of these alloys are shown in Fig. 1. None of the alloys contained sufficient second phase to permit its positive identification. Since alloy B contained the least amount of second phase and since its composition placed it closest to the ideal stoichiometry, it was chosen for the neutron diffraction study. The oxygen content of alloy B was found by chemical analysis to be 0.06 at. pet. This particular intermetallic compound is especially suitable for investigation with neutron diffraction since if the titanium and copper atoms have a random statistical distribution, a near null-matrix would result in which there would be practically no coherent scattering of neutrons, and hence no Bragg diffraction peaks. Such a condition could come

Jan 1, 1962

By G. C. Weaver, G. T. Harley

Beds or lenses of potash and magnesium salts are found in a thick salt section (Salado) overlain by Rustler Red Beds in several members of which water is present and from one of which the rehery process water used by International is obtained. Most of the holes drilled are less than 1000 ft. deep, but some wildcat holes have exceeded 3- ft. cable or rotary tools are used to drill down to the salt. when samples of water are desired the cable rig is used, but when this is not necessary, and where caving material is present, the rotary rig has an advantage. The rotary rig has a further advantage in that only one machine is required to drill and core the entire hole, whereas the cable rig must be moved off to make way for the core drill. Casing is placed in the upper section of the hole as required to shut off water, and a 5 3/16 in. string is set in the top of the salt before coring starts. With the rotary rig drilling is often carried to the salt with no casing, as drilling mud seals off caving ground and water-bearing strata, and a string of 4-in. line pipe or 4-in. flush-joint casing is landed on top of salt. Several types of gasoline-powered core rigs have been used in the field and all recover a 2 1/2-in. core. More recently the Sullivan Model 37 rotary rig has been tried, and a Sullivan Model N has been adapted to use a rotary bit in the upper section. Core is recovered in 20 ft. sections with a double-tube core barrel, and the bit used is a bottom-discharge type set with diamonds. The drilling fluid must be a saturated brine to prevent solution and etching of the minerals present. Blowouts are apt to Occur from nitrogen gas trapped under the several polybalite beds, and some have been strong enough to blow the tools out of the hole. Core recovery is practically loo per cent. All holes are plugged after drilling, with cement through the salt section and opposite each water horizon* and mud through the remainder of the hole. Cores are carefully logged and preserved, and in sections containing potash and magnesium minerals they are quartered and analyzed. The correlation of Ore beds between drill holes is briefly described. Introduction Beds Or lenses of potash and magnesium salts are found at varying depths in a thick deposit of salt (Salado formation) in what is known as the Permian Basin, in southeast New Mexico and West Texas. The Salado formation is Composed chiefly of common salt (halite) with numerous beds of anhydrite, polyhalite, and the soluble potash minerals carnallite, sylvite and langbeinite. Many other minerals are present in small amounts but their Presence has no bearing On the coring technique. Before starting to core the salt section it is necessary to penetrate 300 to 600 ft. of unconformably overlying Permian sediments (Rustler formation) which are composed of red and gray clay, sand, anhydrite, dolomitic limestone and occasional lenses of salt. The upper part of the formation is predominantly anhydrite, and the lower part predominantly clastic. A thin bed of dolomitic limestone, known

Jan 1, 1948

By J. E. Gruzleski, O. Knacke, M. J. Meixner

The evolution of CO from small specimens of solidifying steel was investigated using a thermobalance technique. Gas evolution begins in the temperature range between 1400" to .1450°C. This behavior is independent of the initial carbon content and the cooling rate. At a constant temperature above 1220°C, the rate of gas evolution is temperature independent. Out-gassing of the specimen is thought to proceed through small channels formed by the coalescence of bubbles between dendrites. LIQUID metals have a higher solubility for gas than the solid. During freezing, these gases form bubbles which may either escape or remain trapped in the solid. Recent work1 on the evolution of oxygen from solidifying silver showed that the gas is not liberated until a large fraction of the melt has solidified, and that the gas is evolved in an almost explosive manner. Because of the commercial importance of CO evolution in solidifying steel, a similar investigation was undertaken to study the mechanism of gas liberation in steel. EXPERIMENTAL The thermogravimetric method of Knacke and eixner1 to follow gas evolution by weight changes was employed. Weight differences of 0.1 mg could be reproducibly measured on specimens weighing 2.5 g. These specimens were taken from alloys with varying carbon contents from <0.01 to 0.11 wt pct and <0.01 wt pct of each of silicon, manganese, and aluminum. Specimens contained in 8 mm ID magnesium alumi-nate crucibles were heated to 1550°C under 1 atm of very pure (99.995 pct) CO with a tungsten resistance furnace. The specimen temperature was measured with a Pt-Pt/Rh thermocouple located directly beside the crucible. It is felt that the temperature thus measured was within 5" to 10°C of the specimen temperature since on fusion the weight of the specimen increased due to uptake of CO, and this weight increase was always observed to begin within the temperature range of 1530" to 1540°C. Even after 3 hr there was a small weight increase with time indicating that the molten specimen was not completely saturated with CO. Therefore, complete saturation was not attempted, and all specimens were simply held for 3 hr at 1550°C before cooling. Two types of cooling experiments were performed. One consisted of continuous cooling at linear rates from 5°C to 20°C per min down to 1020°C. The second type of experiment involved continuous cooling at 10°C per min to a predetermined temperature, and then holding the specimen at this temperature for a certain time. During all of these experiments, the sample weight was recorded as a function of time and temperature. OBSERVATIONS Continuous Cooling Experiments. No gas evolution was detected above 1450°C. The initial weight decrease always occurred between 1400" and 1450°C with the average at 1430°C. With further cooling, gas evolution continued to about 1150°C. Between this temperature and 1020°C no further outgassing was observed. This behavior did not seem to depend on the initial carbon content nor on the cooling rate in the range investigated. Typical patterns of gas evolution are shown in Figs. 1 and 2 for different initial carbon levels and cooling rates. The total amount of gas liberated during solidification varied with the rate of cooling. The total weight change on continuous cooling at a rate of 10°C per min

Jan 1, 1970

By P. S. Roller

P. S. ROLLER,* College Park, Md.-The author introduces the idea of a grind limit, or lower size limit of the particles formed directly by grinding, and this is stipulated to be 0.70 micron. I wonder where all the particles come from that lie between the resolving power of the microscope and 0.7 micron, which are seen under a microscope in examining a sample of grind? There is also introduced the idea of excess fines; this is all right in itself but the author&apos;s quantitative definition seems to be virtually impossible to justify, and, moreover, is obscure. Just why E, the excess fines, should be defined as the distance between an F line and a y line is hard to follow, and just where the distance should be measured does not appear to be stated, as it should be since the F and y lines intersect and the distance between will depend upon where that distance is measured. The basis of the presentation is the relation of log weight against log size that was "discovered" by Schuhmann. While this relation is implied to fit, the only evidence that I can see is a slightly irregular line in Fig. I between 2 per cent and 18 per cent by weight and then a precipitous drop at 18 per cent away from any sort of straight line. Some of the conclusions are rather startling, It is stated that E, the excess fines, remains constant for all the comminution products of any particular material, except that it may decrease for some coarse crusher products. No reason is given for this assertion. One would, on the contrary, expect that excess fines, if such a concept is permissible, would generally vary with different conditions instead of being constant. Again, "as much net energy is required to reduce a ton of quartz to particles of the grind-limit diameter (namely, of course, 0.7 micron) by grinding as is required to melt the quartz." This would require many data to prove, but no proof is given; any consideration would reveal the impossibility of such a result. Still further, ore samples "ground to widely different sizes, and with different circulating loads and different size distributions or product, show the same production of new surface area per revolution of the mill, when this grind limit is used." This statement is confuted by ordinary experience, which teaches that grinding efficiency varies generally with the condition of material. For example, the cushioning effect of fines is widely accepted, while the inefficiency of a ball mill when the feed is too coarse is also well known. R. SCHUHMANN, JR.,* Cambridge, Mass.-Contrary to Mr. Bond&apos;s stand, there is a great deal of miscellaneous experimental evidence that comminution does not stop at around 0.7 micron, nor does it cease consuming energy at that point. Comminution is in fact recognized as an important method of producing colloids.5 Mechanical comminution methods were used by von Weimarn6 to produce colloidal dispersions of a variety of solids, including the minerals chalcocite, galena, hematite, stibnite. Utzino7 gives average particle sizes ranging from 0.02 to 0.09 micron for colloids produced by mechanical comminution of several crystal-line solids. Andreasen and Berg8 have ground

Jan 1, 1941

By Bernard Greifer, Walter Rose

The purpose of this note is to call attention to a source of difficulty which can be responsible for inaccurate estimates of wetting liquid pressure gradients in mixture flow experiments of the Hassler1 type. It has been our experience occasionally to observe what were indicated to be negative pressure gradients, and these were regarded as anomalous in the sense that physically they seemed to contradict the requirement that the flux be in the direction of the potential field. At first it was thought that our scheme of assembling barriers and core sample led to such asymmetry in the distribution of isobars that the manometric leads at each end of the core were responding to internal pressures (within the core sample) which could be either positive or negative with respect to each other, and which in fact might tend to reflect on different effective core lengths in an unpredictable fashion as the wetting liquid saturation was changed. However, no compelling evidence of this possibility has been established. In addition, we were cognizant of the requirements to be satisfied in selecting manometric instrumentation to minimize the delay in response of the manometer to the differential pressures being measured, and were thus assured that our observations of anomalous pressure gradients could not be explained by faulty instrumentation. What we suggest is that if a leak sink occurs in the fluid lines connecting the core to either side of the differential manometer, errors exceeding 100 per cent may result in the estimation of wetting liquid pressure gradient, even though the leak is of such small magnitude that its existence would tend to be obscured by evaporation effects. This is because in a system with leaks at the specified places, the manometer rec- ords at steady state the pressure drop across the manometer barrier elements in addition to the pressure drop across the core sample. This possibility will be illustrated below analytically and by numerical example for the special case where the leak occurs on the high presure side of the core, since we are especially desirous to explain the significance of our observations of the so-called negative pressure gradients. It will be indicated, however, that leak sinks on the low pressure side of the core can account for observations of anomalously high values of pressure gradient, and that multiple leaks on both sides of the core can provide for some compensation so that the errors in pressure estimation are reduced. However, the major point to be emphasized is the fact that the impractical-ity of detecting the small leak sinks which give rise to the difficulties makes it a complicated task for the experimenter in a particular case to know with certainty that his estimation of pressure gradient lies within the limits of error which can be tolerated. Fig. 1 shows the electrical analog of the flow system we have employed in our studies. The inflow and outflow barriers. Rbl and Rb2, the manometer barriers, Rm1, and Rm2, and the core sample, Rc— for a particular fluid saturation and distribution— are all designated as resistance elements; the flowmeter is designated by an ammeter, A; the pressure source is designated by E; and the manometric device is designated by the voltmeter, V. Included in the circuit are the variable resistances, Rx and Ry, which may be regarded as the leak sinks, and which, in general, can vary in magnitude from zero (i.e., a large hole or shunt in the system) to infinity (i.e., no

Jan 1, 1950

By Wei-Kao Lu

A general rate equation is derived for gas -solid reactions in metallurgical processes by considering the contributions of chemical reaction at inter-phase boundaries and diffusion through the solid product layer simultaneously. This equation takes care of the whole range continuously from purely diffusion control to purely chemical reaction control. The parabolic law,&apos; McKewan&apos;s equationa derived with the assumption of interface area control and the equation derived by Kawasaki et al. with the assumption of purely diffusion control are found as the limiting cases of the corresponding forms of the general rate equation. Many metallurgical processes involving gas-solid reactions have been extensively studied. The common examples of these processes include the reduction of oxides to metallic products, the roasting of sulfides to oxides or chlorides and the oxidation of metals. Gas-solid reactions consist of at least two kinds of steps; one involves the transport of reactant and product to and from the reaction zone and the other is the chemical reaction taking place at the interphase boundaries. In the limiting cases one or the other of the above mentioned steps will be much slower than the other and is then rate controlling. For example, the oxidation of metals is diffusion controlled and obeys the parabolic law.&apos; On the other hand, if the diffusion rate is much larger than the chemical reaction rate then it will be controlled by the chemical reaction rate. This latter case occurs for the reduction of iron oxides studied by McKewan. If the rate of diffusion and of the chemical reaction have the same order of magnitude, a more detailed analysis is necessary. Thus at the beginning of the reaction when the solid and gaseous reactants are in direct contact, the reaction should be chemical reaction rate controlled as no diffusion process is involved. At the other extreme as the reaction zone moves further into the solid, the diffusion distance becomes greater and the diffusion rate continuously decreases. Eventually, therefore, for a sufficiently large sample the rate must become diffusion controlled. As the transition from one controlling step to the other must be continuous, we cannot describe this transition period by either of the two limiting cases. In this paper the author attempts to derive a general rate equation for gas-solid reaction systems. This is effected by considering the rates of diffusion and chemical reaction simultaneously from the state of pure diffusion control to pure chemical reaction control. I) DERIVATION The system under consideration is made of at least three phases, namely the gaseous one, the solid reactant, and the solid product as shown in Fig. 1. The notations are listed at the end of the paper. The mass transfer process involves the diffusion of reactants and products across the solid product layer. It is reasonable to consider that no chemical reaction will take place in the solid product layer with respect to both gaseous reactant and product. For the characteristics of local equilibrium of the system, the diffusion process is really in a quasi-steady-state. It will be proved in the following paragraphs. Then we can invoke the law of conservation of mass valid for all chemical species passing through the layer. The solid reactant sphere has an initial radius, ro, after a fraction of completeness of the reaction, the unreacted core has a radius ri, where r, is in the solid product region. Let Cg be the activity of gaseous reactant in the

Jan 1, 1963

By R. Adair, W. R. Hudspeth, W. T. McDaniel

ANEW method for concentrating the flake mica either from present washing plant tailings or from new feed has been developed. In this paper, flake mica refers to that which occurs in weathered granites, alaskites, pegmatites, and schists in a form too fine to be considered sheet mica. It ranges in size from 1 in. or more down to 200-mesh or finer and occurs with kaolin, quartz, and partially kaolinized feldspar. Scrap mica, a term often used indiscriminately to describe everything other than sheet mica, will be used to denote the waste product from sheet mica mines and fabricating plants. Both flake and scrap mica are used in the mica grinding plants to produce wet-ground or dry-ground mica, although some flake mica recovered from clay mining operations is fine enough to be marketed without grinding. Until 1930, ground mica was made almost entirely from scrap. From 1930 until 1935 a gradually increasing amount of flake mica was mined, and most of this was sold without grinding. A large percentage of this mica was recovered from clay washing plants. In this type of operation the recovery of any sheet mica was .of secondary importance and was considered a byproduct. North Carolina produces more flake mica than any other state, accounting for 73 pct of the total domestic production.&apos; In the following review of current practice in jig plants, there is no intention to disparage these operations. Within the limitations under which they were built, they are examples of mechanical ingenuity. The term jig plant is a carryover from early practice since there are no jigs in operation at the present time. Current practice, as shown in Fig. 1, is based on successive stages of differential crushing with rolls followed by trommel screens to remove the more easily crushed gangue minerals. The oversize mica product then is usually dried, and passed through another set of dry rolls and a final screen to remove the remainder of the grit. This process is effective in recovering up to 50 pct of the mica in the original ore, the balance being lost in the screen undersize, usually —% in. Jigs and tables were at one time used to treat this screen undersize and were capable of recovering a part of the mica larger than about 60-mesh.V ine mica lost in these plants is normally washed away in streams and can be considered irrecoverable. There are ten or more washing plants in the North Carolina area that are losing approximately 50 pct of the mica in the feed. Since the three cooperating agencies, the Asheville Minerals Research Laboratory of North Carolina State College, the North Carolina Dept. of Conservation and Development, and the Tennessee Valley Authority, are concerned with conservation of minerals, they have investigated means for improving the efficiency of these mica plants. The mica industry has been cooperating fully with the laboratory to establish a more efficient method of recovery for flake mica. Choice of Methods The initial investigation was concentrated on finding a means of reducing the losses from plants now using the conventional flowsheet shown in Fig. 1. To accomplish this it would be necessary to collect the screen undersize products through the plant and remove from them the mica from Vs in. down to

Jan 1, 1952

By A. M. Gaudin, C. S. Chang

Adsorption was measured for barium ion and laurate radical, using radioactively marked agents, over wide range of concentrations. Laurate adsorbed in absence of barium fails to float. With barium, flotation-effective laurate coating needs not exceed 5 pct of monolayer. Barium adsorption then is several fold larger. No barium laurate coating is obtained. IN general, fatty acids or their alkali salts do not cause clean quartz to float. However, the presence of multivalent ions, such as the alkaline-earth metal ions and heavy-metal ions, can activate quartz and thus make it respond to soap flotation. The mechanism of activation and soap flotation of quartz has been studied.&apos;-&apos; The problem was attacked by making flotation tests rather than by careful adsorption measurements on suspensions of ground quartz in solutions containing the activator and the collector. Attempts have been made to determine the adsorption of barium ion and calcium ion on quartz. The results have only qualitative value as the analytical methods were not adequately precise. Artificial radioactivity and progress in radioactivity detection have introduced a new analytical tool, which can have a sensitivity several orders of magnitude greater than that of other methods."-" In the experimental work described in this paper the radioactive tracers, barium 140 and carbon 14, were employed. Lauric acid was synthesized, having a radioactive carbon atom and was used with Ba" to study the adsorption of barium and laurate ions from aqueous solution on ground quartz. Flotation tests, using the vacuum flotation technique,&apos; were run in conjunction with the adsorption experiments. The object of this investigation was to understand better the reaction mechanism of barium activation and soap flotation of quartz. Methods and Materials Quartz Sample Preparation: Lumps of hand-picked quartz, washed free of stains, were reduced in size to —20 mesh in laboratory crushers. The product was boiled in concentrated hydrochloric acid to remove the abraded iron. The cleaned product then was ground dry in a laboratory porcelain mill, using siliceous pebbles, to pass through a 200 mesh screen. The ground product was deslimed by sedimentation at 37 microns. This —200 mesh +400 mesh material was subjected again to boiling with concentrated hydrochloric acid, followed by washing with distilled water and further desliming. The cycle was repeated until no trace of iron could be detected in the wash liquor. During the final washing, the pH of the wash water was checked to insure the complete removal of the acid from the mineral. The quartz then was rinsed with conductivity water. The resulting product was dried in a drying oven and stored in mason jars. The impurities (spectrographic) were as follows: Al, 0.01 to 0.001 pct; Li, Mg, Mn, Na, each less than 0.001 pct; Fe and Cu not detected. The specific surface of the quartz sample was determined by the krypton gas adsorption method.&apos; The average of two measurements was 881 sq cm per g. The calculated specific surface for 37/74-micron quartz particles, using a surface factor of 1.96; is 900 sq cm per g. This is in good agreement with the measured value of 881 sq cm per g. Beta-Counting Method: The radioactive isotope of barium used in the experimental work was barium 140. It has a sufficiently long half-life, 12.8 days, emits high-energy beta particles (up to 1.0

Jan 1, 1953

By R. Hastings Keller, Ralph E. Esarey

Oil and gas activities in Indiana during 1943 continued to decline at about the same rate as in 1942. New development, production, and prospecting, all showed the results of Federal regulation, low price for crude oil and shortages of labor and materials. Several local stripper-well operations were shut down, but no large-scale abandonment has occurred as yet. During the year, 277 holes were drilled for oil and/or gas, of which 99 were completed as oil wells, 19 as gas wells and 159 as dry holes. This record was a decline of 21 per cent over last year in the total number of completed tests, 20 per cent in the number of oil wells and 9.5 per cent in the number of gas wells. As in the previous year, most of the drilling (82 per cent) was in the southwestern part of the state, Gibson County ranking first, with 77 completions and 49 oil wells. Posey County was second, with 59 completions and 20 oil wells. The remaining 18 per cent of the drilling was in the old Trenton oil and gas area. The total footage drilled during the year was 482,799 ft., a decline of 18 per cent from the previous year. Of this, 146,559 ft. was wildcat footage. The total initial production of the oil wells completed was 6362 bbl. and of the gas wells was 2,562,000 cu. feet. Three tests in the old Trenton area were drilled into the Cambrian and reported nothing more than tar residue. A Devonian test was to be made in the Griffin field in the Wabash River but owing to the delay caused by Federal restrictions this test did not get started in 1943. It is expected to spud in March 1944. The new discoveries in 1943 were virtually all in the southwestern part of the state, and all, so far, one-well pools, except the North Owensville pool, in Gibson County. The latter was brought in during the month of February and by the end of the year nine wells were producing in the pool, with a reported initial yield of 1260 bbl. per day. By the end of the year it had a daily average production of 401 bbl. and an accumulated production of 39,129 bbl. The average gravity of the oil is reported to be 35.6. On the Ohio-Indiana state line, early in 1943, a new pool was opened up in the old Trenton area. The discovery well was on the Ohio side of the line. An offset well was completed on the Indiana side, with an initial yield of 64 bbl., and settled down to 5 bbl. per day. This well produced 1154 bbl. up to Dec. 31, 1943. There was no further development here, because of the Federal Conservation Order M-68, restricting drilling to one well on 40 acres. The development pattern usually followed in this area is one well to 5 acres. The E. Rogers field was extended by one pool (Table 4). Most of the development in the oil-field pools during the year took place in the Kirksville pool, in Gibson County, and the Caborn pool, in Posey County. The crude-oil production in 1943 was approximately 5,273,000 bbl., a decline of 29 per cent from the previous year. The Griffin field, of Gibson and Posey Counties,

Jan 1, 1944

By E. R. Morgan, J. C. Shyne, V. F. Zackay

NITROGEN used as an alloying element in steel has received increased attention in recent years. Its merit as an austenite stabilizer and hot strengthener has long been recojinized, particularly by European investigators.&apos; Recent work&apos; indicates that nitrogen contents of 1/2 pet or more allow the addition of several percent of ferrite-forming hot strengtheners, such as molybdenum, without subsequent austenite decomposition at elevated temperatures. The room temperature properties of such alloys are characterized by high yield and tensile strengths with excellent ductility, even after cold working, as shown in Tables I and 11. The 1350°F creep-rupture properties of Cr-Mn-Mo-N steels are comparable to those of commercial alloys such as Timken&apos;s 16-25-6, containing 16 pct Cr, 25 pct Ni, and 6 pct Mo, as shown in Table 111. Cr-Mn steel containing 1/2 pet N must be made under pressure if ingot gassing is to be avoided. However, there are difficulties attendant to this practice. Pressure melting restricts the ingot size, and the equipment required is costlier and more complicated than comparable air melting facilities. An effort was made, therefore, to determine more economical methods of adding nitrogen. The pressures generated in centrifugal casting usually range from 5 to 10 atm and are sufficient to prevent the gassing of high nitrogen Cr-Mn steels. A centrifugal casting was made of an alloy containing 16 pct Cr, 18 pct Mn, 2 pct Mo, and 0.73 pct N. A risk, 8 in. diam and 11/2 in. thick, was cast at 1000 rpm. It has been calculated that a pressure of 9 atm was developed at the periphery of the disk at this speed. Metallographic inspection of specimens taken from the disk revealed no gas holes or other defects. The elevated temperature properties of radial and circumferential specimens taken from this casting are given in Table IV. The creep-rupture life of this cast nickel-free austenitic steel was equivalent to that of the wrought Cr-Mn-Mo-N steels. Although centrifugal casting appears to lend itself to the manufacture of high nitrogen austenitic steels, its applicability is limited to relatively small sizes. A process applicable to the production of large ingots of high nitrogen steels is suggested by low carbon rimming steel practice. Gas holes formed during ingot solidification are not detrimental so long as they weld shut during hot rolling. An economic advantage of gassing ingots is the increased ingot-to-slab yield, especially if the ingots are mechanically capped. Also, it is probable that higher nitrogen levels can be achieved with mechanically capped ingots than with hot-topped ingots. A series of gassy heats was made under both atmospheric and higher pressures, as indicated in Table V. Several of the air melted ingots were mechanically capped by chilling the molten ingot tops with a steel plate. The ingots were allowed to freeze in steel molds without hot-tops unless otherwise indicated. A comparison of Table 111 with Table V shows no significant difference in the creep-rupture properties of the gassy and the nongassy ingots after hot work- ing nor between the gassy ingots produced under different conditions. References 1 F. Rapatz: Application of Stainless and Heat Resisting Steels Alloyed with Nitrogen. Sinhl and Eisen ( 1941) 61, pp. 1073-1078. &apos;V. I?. Zackay, J. F. Carlson, and P. L. Jackson: High Nitrogen Austenltlc Cr-Mn Steels. Trans. ASM 119551 48, PP. 509-523.

Jan 1, 1957

By R. M. Waterstrat, B. N. Das, Paul A. Beck

The central region of the Ti-Mn system was studied by means of x-ray diffraction and metallography. The following intermediate phases were found: an MgZn2-type Laves phase having a wide range of homogeneity, the p-phase of unknozvn structure, and the + -phase with a large tetragonal unit cell, stable at 930°C and lower temperatuves. MaYKUTH, Ogden, and jaffee1 determined the soh-inh bility limits of manganese in and titanium. They also confirmed the structure of the intermediate phase TiMn2 which was previously identified as a Laves phase of the MgZn2 type.2 In the equiatomic region they found an additional intermediate phase, y, which presumably formed by a peritectic reaction at approximately 1200°C. This last observation was disputed by Rostoker, Elliott, and McPhers0n,3 who reported the formation of TiMn by a peritectoid reaction between 900" and 1000° C, but found no evidence of any intermediate phase other than the Laves phase at higher temperatures. Elliott and Rostoker4 reported furthermore that the X-ray diffraction pattern of TiMn could be identified as that of a o phase, although their data were not consistent with this inter~retation. Margolin and Ence6 presented X-ray diffraction data for the and TiMn, phases, and obtained metallographic indications for the existence of additional intermediate phases. They concluded from metallographic observations that two Laves phases are formed, separated by another phase, designated as &. A phase diagram for the Mn-Ti system has recently been published by Savitskii and Kopetskii,&apos; which shows considerable disagreement with the work of Hellawell and Hume-Rothery,&apos; in particular with regard to the terminal solid solutions based on the allotropes of manganese. The present investigation was undertaken primarily to clarify the phase relationships in the equiatomic region of the Ti-Mn system. The metals used in preparing the alloy specimens Were iodide titanium and electrolytic manganese, the latter having a nominal purity of approximately 99.9 pet. The electrolytic manganese was covered with a thick coating of a brown manganese oxide which was reduced by heating the manganese flakes for several hours at a temperature of 1100°C in an atmosphere of dry hydrogen. Three alloys, see Table I, were made using sponge titanium and manganese not treated in hydrogen. The alloys were prepared by arc-melting in an atmosphere of helium or argon. The composition of each alloy was calculated by assuming that the entire weight loss during melting was due to evaporation of manganese. Subsequent chemical analysis of certain alloys indicated that the compositions calculated in this manner were usually correct within 1.4 pct, as shown in Table I. All alloy specimens were wrapped in molybdenum foil before sealing them in silica tubes containing argon at a pressure of 1 atm at the annealing temperature. After quenching in cold water, the alloy specimens were bright and clean, with no evidence of any surface reaction. The alloys listed in Table I were examined metallographically, using an etchant consisting of 20 pet HNO3, 20 pet HF, and 60 pet Glycerine. powder for X-ray diffraction examination was prepared by crushing a portion of each annealed alloy

Jan 1, 1962

By William P. Roe, James F. Ellis, W. P. Fishel

This paper reports the distribution of sulphur between iron and titanium in Fe-Ti-S alloys and in Fe-Ti-C-S alloys. The compound TiS was found, which exists as a separate phase at temperatures below 1200°C. Titanium in steels reacts with sulphur in preference to carbon. THE general relationship between sulphides and oxides suggests that titanium should form a sulphide having properties similar to those of titanium oxide. Specifically, titanium should form a sulphide having a high melting point, indicating a stable compound of low solubility in iron. In order for titanium to function as a desulphurizer or sulphur-fixing agent, it must form a stable sulphide. In order for titanium to be useful in steels as a sulphur-fixing agent, it must react with sulphur in preference to carbon; and to prevent red shortness, the sulphide formed must have solubility characteristics such that grain-boundary precipitation of a low melting phase be prevented. The terms desulphurizer and sulphur-fixing agent when used in this paper mean that titanium effectively ties up the sulphur as a stable sulphide of titanium and prevents the formation of FeS. The partition of sulphur between iron and titanium was studied first in a series of Fe-Ti-S alloys with carbon absent, and then in similar alloys with carbon present. It has been established1 a that titanium forms in Fe-Ti-S alloys a sulphide of undetermined composition which can be identified microscopically, and that the titanium-sulphur compound does not respond to sulphur printing. It has been suggested" that titanium acts as a sulphur-fixing agent, although little proof has been given. A report% f The Advisory Committee on Metals and Minerals summarizes the existing knowledge concerning substitutes for manganese in steels. Distribution of Sulphur Between lron and Titanium with Carbon Absent The alloys for this investigation were made by -melting Armco iron, metallic titanium, and iron sulphide in alundum crucibles in an induction furnace. The titanium was added after the other ingredients had melted. In these alloys it is recognized that some TiO, is present, since the melts were made in the air and since no other deoxidizing agent was used; therefore, some of the titanium was not available for union with sulphur. Thirteen ingots were made. Ingots 2 to 8 were cast in baked-sand molds; ingots 12 and 13 were cast in metal molds; and ingot 9 was allowed to cool in the crucible in which it was melted. All ingots were annealed at 900°C for 3 hr and were cooled slowly in the furnace. Subsequent information indicates that this anneal had no effect upon the titanium sulphide particle size, since the sulphide does not go into solution at this temperature. The ingots as-cast were approximately 8 in. long and % in. diam. The compositions of the eight ingots and the results of the partition of sulphur are recorded in Table I. Total titanium and total sulphur were determined by standard gravimetric methods. Sulphur, as iron sulphide, was determined by the evolution method. Boiling HC1 (1:4) liberates the sulphur from iron sulphide, but leaves the titanium sulphide practically undecomposed. A complete separation of titanium sulphide and iron sulphide cannot be made by digesting these compounds in dilute HC1 since titanium sulphide was found to dissolve slowly. The amount of titanium sulphide decomposed will depend upon the time in contact with the acid and upon the particle size of the sulphide. The point

Jan 1, 1952

By Fritz Wald, Martin Weinstein, Herbert E. Bates

ONE of the methods for preparing a fiber-reinforced composite is that of directional solidification of a eutectic melt. Its major drawback lies in the fact that the matrix in this case is a pure constituent (or a terminal solid solution of fixed concentration) and the ratio of fiber to matrix volume is fixed. Only very few, if any, pure metal matrices possess technically useful properties, and the necessity for alloying is the inevitable consequence of this fact. Nevertheless, structures similar to binary eutectic ones can be obtained in higher-order systems by pseudoeutectic reactions of a type where crystallites and a solid-solution matrix are simultaneously precipitated from a liquid.1"3 Trials to solidify such alloys directionally—the matrix consisting of a solid solution of chromium, nickel, and silicon and the reinforcing phase being Cr3C2— will be reported elsewhere.4 This note describes constitutional studies carried out primarily to determine -the position of the pseudoeutectic and further to insure that no undesirable reactions between the reinforcing phase and components of the matrix occur in the solid state at elevated temperatures. In the investigation it was specifically desired to reinforce a matrix which is as close as possible to the 80/20 Nichrome composition with Cr3C2 to insure good oxidation resistance and improve the strength at high temperatures. No konstitutional information on the quaternary C-Cr-Ni-Si svstem in the range of interest was found in the literature. Nevertheless, the two most important ternaries c-cr-Ni5 and have been completely investigated. On C-Cr-Si some investigations have been conducted,7"9 and C-Ni-Si was studied with a few alloys in the course of the present investination 9 Eight alloys were prepared to investigate the changes in the constitution of the ternary C-Cr3Ce-yNi eutectic5 with silicon contents up to 10 wt pct. Five alloys were made with varying chromium and carbon contents around compositions containing -60 to 70 wt pct Ni and 5 to 6 wt pct Si. The latter alloys were designed to locate the y(Ni-Cr-Si-C)-Cr3C2 pseudoeutectic reaction and to determine the maximum chromium content in solid solution compatible with a two-phase equilibrium between y and Cr3C2 only. The alloys were prepared from electrolyte nickel, semiconductor-grade silicon, high-purity chromium (99.999 pct pellets), and semiconductor-grade graphite. They were melted inductively in high-purity alumina crucibles under flowing argon. Twenty-gram charges were prepared and investigated by optical microscopy (including hot-stage investigations), X-ray powder methods, and electronmicroprobe analysis. The alloys were studied in the as-cast condition and also after a homogenization annealing treatment of 48 days at 900°C in evacuated quartz ampoules. The structures of all alloys based on the ternary eutectic plus silicon in the as-cast state consisted generally of an inhomogeneous solid solution and at least two other phases. Annealing of the alloys showed that the solid solution was homogenized but that an incomplete solid-state reaction had taken place in all alloys with silicon contents over 5 wt pct. This reaction is clearly visible in Fig. 1 around the large Cr3C2 crystals (gray areas). Below this concentration, the alloys consisted of a homogeneous matrix and precipitates of Cr3C2 plus graphite, Fig. 2. The Cr3C2 was entirely unreacted here. From these observations and subsequent X-ray and microprobe investigations, it was deduced that the

Jan 1, 1969