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By Paul F. Sikora, John W. Weeton, Philip A. Clarkin

Effects of variations in carbon content and micro-structure on the stress-rupture properties of a liquid-phase-sirztered, Izeut-resistunt alloy were studied. Using the powder rnetallurgy technique, specimens could be produced having stress-rupture profierties equivalent to or better than those of the alloy in wrought form. In addition, sintered specimens were easily hot worked even when carbon contents were as high as I pct. Correlations between stress-rupture properties and micro structure were obtained in some cases. IT is well known that many alloys used in high-temperature service are quite complex both from the standpoint of chemical composition and microstruc-ture. During the development of these alloys their complexity has in some cases led to treatment schedules which, while practical, often do not fully utilize available alloying elements. For example, in some alloys, massive carbides that form during ingot practice may never be dissolved by reasonable solution treatments and therefore are unavailable for subsequent precipitation hardening, Ref. 1. This investigation was undertaken to determine whether the powder metallurgy method of consolidation might provide distinct advantages over the conventional melting method. First, segregation of alloying elements might be eliminated or reduced and thus a more efficient use of these elements obtained. Second, the structures produced in powder specimens might differ considerably from those normally seen in castings, which would perhaps result in better response to subsequent mechanical or thermal treatments. If these results could be achieved, it was felt that both forgeability and stress-rupture properties might be improved. The alloy used in this investigation was S-816 produced by powder metallurgy methods. The influence of carbon content over the range of 0.1 to 1.0 wt pct was studied for the alloy in the as-sintered condition and after three additional treatments involving heat treatment, hot swaging, and combinations thereof. EXPERIMENTAL Standard S-816 has the composition shown in Table I. It is a cobalt-base alloy having a nominal carbon content of 0.4 pct and is classified as a precipitation-hardening alloy. This alloy was chosen because it is representative of a class of heat-resistant alloys and because adequate information was available concerning its response to various treatments and its subsequent properties. For this study, -100 mesh alloy powder having essentially the same composition as shown in Table I but without the 0.4 pct C was used. Desired carbon variations, achieved by wet mixing the alloy powder with graphite, ranged from 0.1 to 1.0 wt pct. Specimens were hydrostatically pressed and then sintered in vacuum by slowly heating until incipient melting was observed. Sintering this way gave a liquid phase in the interstices of the powder particles and achieved reasonably high densities. The average density of as-sintered specimens was about 95 pct of wrought alloy density. In order to obtain various microstructures, some specimens were heat treated (solution treated for 24 hr at 2150°F, water quenched; aged 16 hr at 1400°F, air cooled); others were hot swaged at 2150 o F (50 pct reduction in area), while still others were hot swaged (50 pct reduction in area) and then given the above heat treatment. During hot swaging it was observed that forgeability was quite good even when specimens contained as much as 1.0 pct C. Specimens were evaluated by stress-rupture test-

Jan 1, 1962

The following discussion was held on August 2.2, 1927, during the meeting of the Institute at Salt Lake City. It. IIIatch, Garfield, Utah.—The object of the meeting, I belicve, is to determine the proper chemical composition to insure long wear of a ball in a mill. This might also apply to rods. I feel that there is only one way of determining accurately the wearing qualitites of balls in mills, and that is by actually wearing out 400 or 500 tolls of different kinds of balls. Two things add to the wearing qualities of a ball: (1) the tenacity, with which the particular particles of iron or steel are cemented together; (2) the individual hardness of each partiele. My cxperience is that a ball that has a Brincll of 550 to 650 will give comparatively long life. There is considerable difference of opinion as to the best steel and iron alloys neressary to produce a satisfactory grinding ball. Some favor chrome, somr manganese and some nickel. Personally, I believe cast iron alloyed with chrome makes the best ball for small sizes; for large sizes (31/2 in. arid up) steel alloyed with mangancse is best. L. O. Howard, Pullman, Wash.—How would this method apply to hning? A. Kissock, New York, N. Y.—I do not subcrribe to that exactly. As I am connected with the Climax Molybdenum Co. you would hardly expeet me to. We have been much interested in trying to work out this grinding-ball problem and have spent a, great deal of thought and effort on it. As we are associated with the American Metal Co. and have a number of Proprrties of our own, we are in poaition to try our ideas on ourselves first. We have tried different compositions, but, have not done anything very remarkable with them yet,, except that we have been able to say that the combination of chrome and molybdenum has given the best, wearing ball. Mr. Hatch feels that iron is the best. medium for grinding, hut since toughtiess as well as hardness is a factor in wearability we feel that a well made steel is better for this purpose. Another difficulty with cast, or chilled, iron is lack of uniformity front day to day or from heat to heat, and the impossibility of protlucing balls free from blow or shrinkage holes which matcrially affect their life. If cast iron is made hard it becomes brittle and lacks strctc11 or toughness. This prevents its use in the larger sizes, aS break:lgc, is a sorious hantlic:~p to efficiency. We have trirtl hard and brittle balls and softer but very tough balls and have proved that maximum hardness, with sufficient toughuess to prevent hrcakagv, is the most important factor. in resistarnce of a steel to abrasion. It is well known that. chromium is one of the best hardening agencies for cither iron or. steel; and molybdenum is recognized as an clerment capable of introducing tonghness. In other words, the proper combination of chromium and molybdenum permits unusual hadness with still sufficient touglmess to stand up under severest impact. It is yet too early to give any actual mill results of this new Combination and though we are proving our ideas at our own properties we have to depend on others to help us make the data more general and more worth while and would like to get some testa 206

Jan 1, 1928

By H. H. Rachford, W. P. Schultz

The use of computing machines to solve physical problems has made it imperative to represent physical data in a form computing machines can use. Although curve-fitting is an old and well-practiced art, there still seems to be some interest in questions regarding what methods can be used and why. Even though there are many good references on this subject, it may be of some value to discuss a particular method which has been found to be useful, and to make available the associated tables for its use in the hope that there can be saved a few of the tedious hours necessary to reduce many of the suggested methods to practice. Although a very brief discussion of the mathematical background of the tables is included immediately below, those interested only in their use can skip directly to the section on Use of the Tables. Direct Fit of a Set of Points A frequently encountered problem is: Given data which can be represented by a smooth curve +(x), to find an easy-to-evaluate function F(x), with only a few coefficients, which is a good approximation of +(x) on the interval 0 5 x 5 x,,. Consider F(x) as a linear combination of m + 1 arbitrary, single-valued, continuous, linearly independent functions, f,(x), Ff.r) = a,f,,(x) + a,f,(x) + . . . -+ aifi(x) + If the value of Q(x) is known or can be read from the curve at the points xj: +(x,), +(x,), . . . +(xj) . . . +(x.), and if n = m, then all the coefficients of F(x) can be determined, and F(x,) — +(x,) = 0 for j =. 0, 1, . . . n. This follows from the solution of the set of linear simultaneous equations. +(x<.) = aofu(x,) + aJI(x0) + . . . + a,,f,,(x,) =a, = f,,(x,) + a,f,(x,) + . . . + a,f,,(x,) +(x,,) = a,,f,(x,,) + a,f,(x,) + ... + a,,f.,(x,,) ...........(2) for the coefficients ai. Although the set a, from the solutions of the set of Equations 2 yields an fix) which gives an exact value of $(x,) for the n + 1 points, there is no restraint on ~(xj in the interval xi < x < xj+,. Indeed, if the points +(x,) are not sufficiently smooth, very serious differences may arise between F(x) in the interval and the smooth curve +(x). Least-Squares Fit It has been suggested&apos; that if n is made to exceed m, and the set a, is found to minimize the function the function F(x) is much less sensitive to small errors in +(xi), and usually provides a much better fit to the smooth function +(x). This is known as the best fit in the sense of the least-squares error. To minimize $, set £ = 0,i = 0,1,... iTi,.....(3) The m + 1 equations from Equation 3 are k—n k=n a.^p /,,(a) f,(xt) + a, X A: = 0 k=O k = n k = n k=0 k=0 +(x.) f,(x,),............(4) one equation being obtained for each i = 0, 1, . . . m. Inasmuch as the arbitrary functions f(x) may be evaluated for x,, k = 0, 1, . . . n, and the +(x,) are known values of the curve to be fitted, the simultaneous Equations 4 may be solved for the set a,. Particular Functions Given the functions fi(x) and the set +(xj), the solutions for require evaluating a number of sums of products as well as the solution of an m + 1 order set

Jan 1, 1956

By W. A. Tiller, D. H. Lane

A simple zone melting technique for investigating the effect of ultrasonic irradiation upon ingot solidification is described. The effect of i) ultrasonic power level, ii) freezing velocity, iii) constant or variable frequency, iv) direction of irradiation, and v) transducer-cozlpling bar joint upon the grain size of type 316 stainless steel have been investigated. The results of this study support the postulate that the ultrasonic vibrations increase the nucleation probability in the layer of liquid adjacent to the freezing interface. In the first paper of this series,l the possible mechanisms whereby ingot structure may be altered by irradiating a solidifying melt with ultrasonic vibrations have been discussed. A new technique for efficiently introducing the ultrasonic energy into the solid-liquid interface region was presented and applied to ingot preparation by the consumable-electrode arc-melting technique. The experimental results support the postulate that the ultrasonic vibrations may refine the ingot structure by increasing the nucleation probability in the zone of liquid adjacent to the solid-liquid interface. The present paper describes the application of the same technique to a relatively simple laboratory-scale zone-melting apparatus. This apparatus is found to be well suited to the detailed study of ultrasonic technology in this area and allows one to determine, in particular, the minimum ultrasonic power level needed to produce a certain grain size in a particular material. The results of this study support the nucleation mechanism for grain refinement, but do not provide any deeper insight into the specific nucleation mechanism responsible for the effect. EXPERIMENT The apparatus used in these experiments is shown in Fig. 1 and illustrated schematically in Fig. 2. The material to be studied, type 316 stainless steel, was in the form of a 3/4-in. sq bar 30 in. long. This bar was placed in a square open-topped mold with open ends made from high-purity graphite coated with a flame-sprayed Al2O3, skin. Surrounding the crucible was a molybdenum susceptor with 3/8-in. diam holes spaced at 1-in. intervals along its length to serve as sight ports. A tapered transducer horn was brazed to one end of the specimen. The entire assemblage was enclosed in a quartz tube, the transducer itself projecting out one end and resting in a cooling tank. A vacuum seal between the coupling bar and the end plate enclosed one end. A similar plate and seal enclosed the other end. The system is gas tight and can be operated under vacuum or an inert atmosphere. For these experiments, flowing argon at 15 in. of Hg pressure was used. The heat source for melting the stainless steel specimen was a 2-in. long 3-kc induction coil placed around the quartz tube; this source heated the susceptor, inductively melting about a 3-in. zone of the sample by radiation. To move the molten zone at a specified rate, the induction coil was moved down the tube at this rate with the aid of a simple pulling device. Thus, the sample could be irradiated at varying ultrasonic frequencies and power levels and could be solidified at varying rates. The power input to the transducer was measured using a "John Fluke Model 120 VAW-Meter." 1 The power delivered to the interface region is probably only about 50 pct of that measured on the VAW-meter. In a typical run, when the ultrasonic power was turned on, the interface of the molten zone closest

Jan 1, 1961

By M. B. Bever, A. K. Jena

Equations for the heat effects of additions of mixtures of multicomponent phases to a solution have been derived in terms of the enthalpy interaction coefficients. By these equations and the measured composition dependence of the heat effects on addition of mixtures of Ag-Cd solid solutions and elemental gold to liquid tin-rich solutions, values of the enthalpy in-teraction coefficients in tin-rich solutions at 541°K have been obtained. Analysis of these results in terms of the regular solution theory and the quasichemical theory suggests that near the melting point of the solvent association in the solution becomes important. SEVERAL recent papers have discussed the nature of dilute solutions on the basis of measured enthalpy interaction coefficients. l-3 These coefficients contribute to the understanding of the constitution of solutions and can be used for the evaluation of thermo-dynamic data of multicomponent systems. Their experimental determination, however, is laborious. This communication reports values of the enthalpy interaction coefficients of silver, cadmium, and gold in liquid tin-rich solutions obtained from the measured heat effects of additions of mixtures of Ag-Cd solid solutions and elemental gold to liquid tin-rich solutions. These measurements were made in the course of the determination of the heats of formation of Ag-Cd solid solutions reported in a concurrent publication.4 THE HEAT EFFECT OF AN ADDITION TO A SOLVENT This analysis is concerned with the addition of a mixture of multicomponent phases from the temperature To to a solvent of known composition at temperature T. If Nmixture << Nsolvent, where N designates the gram-atoms involved, the heat effect per g-atom of mixture added can be expressed as (HT - HT )i = difference in heat contents of component element i in the mixture at T and To Q = heat effect on solution of the mixture at temperature T Also1 Q = S Xi [2] where ?Hi ,T,x = relative partial enthalpy per g-atom of i in solution at temperature T and composition of the solution x x = average of the atom fractions of the added material in solution before and after the addition The relative partial enthalpy ?Hi per g-atom of an element i in a dilute solution composed of a solvent 1 and the solutes j = 2, ... m is related to the interaction coefficients by the following equation5 ?Hi - ?Hi-?Hi,x1-1 + S n1ixj+ . . . +S...S...S nrixn2... Xnji...Xnmm+... [3] vhere Xj = atom fraction of solute j in the solution nri = the rth order enthalpy interaction coefficients, which can also be represented by the more general designation5 L(i)n2,n3...nm and nj may have any positive integral value or zero. For identification of various rth order interaction coefficients, nri is written as nn2i2,...nij,...nmm. Eqs. [1], [2], and [3] combine to give the following general equation for the heat effect on addition of a small quantity of a mixture to a large quantity of a dilute solution n = _?ga ?Ha,To + ? Xi(HT - HT o)i a. i + ?xi?Hi,X,i?1 + ??xnjixj + . . . + ???. . .?. . .?xinrix2n2. . . xnji. . . xnmm+ . . . [4] The first term on the right-hand side of this equation is the heat of decomposition at the addition temperature of the phases present in the mixture, the second term is the heat required to raise the temperature of

Jan 1, 1970

THE first session, on alloys, took place on Feb. 20, Mr. Hibbard presiding in Mr. Campbell&apos;s absence. The session covered papers on manganese, chrom-ium, and silicon-iron systems. The paper by Mr. Wohr-man was presented by title. This paper showed that manganese produced many of the effects generally at-tributed to carbon and indicated that our ideas on manganese-bearing carbon steels will have to be modi-fied. Mr. Krivobok discussed this paper at length, the gist of the discussion being that Mr. Wohrman had not taken sufficient precautions in carrying out some of his work, particularly as to the effect of silicon content. Mr. O&apos;Neill&apos;s paper was also submitted by title and no discussion was presented. Mr. Kinzel showed an accu-rate determination of .the carbon-free chromium-iron loop obtained by means of a specially constructed tele-scopic dilatometer. The loop ends at 12.2 per cent chromium and the A-3 line at 900° is unaffected by chromium content. The A-4 line runs from 1400° to 900° at 12.2 per cent chromium as an oblique hyperbola. Mr. Armstrong noted that the brittleness in high chromium alloys is often attributed to the gamma trans-formation and, according to Mr. Kinzel&apos;s diagram, this was impossible, as the alloys contain a great deal more than 12 per cent chromium. Mr. Kinzel pointed out that the presence of carbon in the commercial alloys caused the elongation of the loop and made possible local austenite transformation. Mr. Styri discussed the relations of the coefficients of expansion as shown in Mr. Kinzel&apos;s diagram. Mr. Corson presented the iron-silicon diagram as he has worked it out, showing a pre-viously unknown intermetallic compound, Fe3Si, to-gether with characteristic martensitic microstructure varying with the heat treatment. This compound formation is the real cause of brittleness in certain re-gions of the iron silicon system. Dr. Hoyt believed some of the evidence presented by Mr. Corson to be in contradiction to the phase rule. Mr. Corson, however, explained that true equilibrium does not exist under the conditions at hand so that the diagram was accurate in spite of the discrepancies. The meetings of the subcommittee on the physical chemistry of steel making were interspersed through-out the day. Mr. Feild presided and a program of future activities was outlined. This includes plans for a symposium at the next annual meeting and methods for interesting physical chemists, particularly those in the universities, in the steel man&apos;s problems. This committee has been newly formed and bids fair to be especially active. The Iron and Steel Committee met at lunch at the Engineers&apos; Club at noon, with Dr. Mathews presiding. After disposing of the food the report of the Nominat-ing Committee was called for, and the following were elected as officers for 1928: R. H. Sweetser, chairman; F. C. Langenberg and W. J. Priestley, vice-chairmen; and T. T. Read, secretary. The report of the commit-tee appointed at the last meeting to consider publica-tion and general policy for the committee was then called for and freely discussed, after which it was unanimously voted to instruct the new chairman to se-cure from the Board of Directors the necessary author-ization for the Committee to recognize as a Division. This action was reported to the Board at its meeting the following night and the authority granted.

Jan 3, 1928

By F. G. Stone

A strain anneal process has been developed for the production of iron tensile test specimens, the gauge length consisting substantially of a single crystal. Since the method has been quite successful and the results of general interest, the process will be described in detail. The essential feature of the method is the continuous, slow increase of furnace temperature during the grain growth annealing treatment. Briefly, the procedure is as follows: I. Heat machined bars to 950°C in vacuum, hold I hr, furnace cool. 2. Strain to 3.5-3.8 pct elongation (axial loading). 3. Anneal, continuously raising furnace temperature from 450 to goo°C at a rate of so0 per day. 4. Hold at 900°C for 100 hr, furnace cool. Tensile test bars with a 1/4- X I-in. gauge section have been machined from 35 in. Armco iron plate and from 3/8 in. rounds of Puron, a high purity laboratory iron. The axes of the test bars were parallel to the rolling direction. During the machining operation the bars were polished in the lathe with No. 600 emery paper, and this polished surface proved satisfactory for subsequent macrographic examination. A micrographic examination of the materials showed that both were high in FeO inclusions, those of the Puron being smaller but much more numerous. An initial vacuum anneal of I hr at 950°C followed by furnace cooling resulted in a ferritic grain size of ASTM No. 3 for Armco and No. I for Puron. Special heat treatments which reversed these initial grain sizes produced no marked improvement in single crystal production after subsequent treatments. The strain for subsequent grain growth was determined by measuring the decrease in diameter of the tensile bar during elongation. The change in diameter could be measured with a micrometer to +0.0001 in., corresponding to an accuracy of 0.08 pct for the elongation. The accuracy of the diameter measurement is also dependent on grain size because of surface roughening during elongation, and for this reason, a No. 3 grain size is preferable. Optimum grain growth was obtained for strains between 3.5 and 3.8 pct elongation. Once this critical range had been cstablished, all tests were held within it, and thus only meager data were obtained on grain growth following strains outside this range. In general, grain size decreased with increasing strain to about 5.5 pct elongation. Grain growth was occasionally ohserved with less than 3.5 pct elongation although this may have been the result of error in measurement. The program control device for the grain growth anneal consisted of a I rpm motor and worm gear drive, operating the dial shaft of a pyrometer controller through a pulley. Dial speeds of 33 and 50°C per day were tried, but the results at the slower speed were not sufficiently better to warrant the increased furnace time. The furnace tube was evacuated at intervals during the run and kept sealed the remainder of the time. When the controller reached 900°C, it was stopped and the temperature held for 100 hr. This eliminated most of the numerous minute crystallites of about 0.5 mm dimension usually visible on the etched surface of the single crystal. Since recrystallization occurs at the A3 point, the top temperature of 900°C provided a 10"

Jan 1, 1949

By R. B. Basham, C. W. Macune

Very small anomalies in oil well temperatures are detected and measured by recording the difference in temperature existing between two thermally sensitive elements which are Spaced several feet apart on a small diameter carrier and lowered into the bore hole. The elements operate in a balanced electrical circuit, sending only a reference signal to the surface as long as normal gradient temperatures are encountered. When either element enters a temperature disturbance, the circuit is unbalanced and produces large recorder deflections for minute anomalies. The system is highly sensitive to small changes in thermal gradients caused by gas or fluid movements and to the boundaries between beds of different thermal conductivities. Typical logs show successful applications of the process to locating gas and water leaks in casing, gas entry. and gas-oil contact. INTRODUCTION Temperature measurements in oil wells have been made for many years in the search for additional sub-surface data. It has been generally agreed that many conditions in the bore hole cause temperature anomalies which, if located and measured, could yield useful information. But it has been recognized also that many of the irregularities, such as those produced by small water and gas leaks in casing, might be very small and therefore difficult to detect with certainty.&apos; Although tem-perature measurements have been made which show anomalies of less than one degree deviation from normal gradient temperatures, the results are usually uncertain and unconvincing. The general practice, therefore. has been to apply temperature surveys only to well conditions known to produce appreciable disturbances, such as large gas movements and cement hydration. The recent demand for additional data has focused attention on the problem of making dependable temperature measurements in the many instances when anomalies are small but none the less significant. It was for the purpose of detecting these minute irregularities in bore hole temperature that the program described in the following text was undertaken. It is emphasized that this presentation is not offered as a complete study, for the process is still partially experimental. But it is believed that the results obtained thus far will be of interest to the industry and will indicate some of the system&apos;s potential usefulness. DIFFERENTIAL MEASUREMENT The chief value of differential measurement lies in the fact that a difference between two factors can be measured with a scale of values best suited to that difference, a scale that might be cumbersome or impractical if applied to the factors themselves. When it is known that all factors will be nearly equal in value, and therefore the differences consistently small, it is possible to select a scale which will permit very accurate measurement of the differences. This advantage is in contrast to the difficulty of measuring small increments of a single variable, where the increment must be measured by a large scale which is suitable for measuring the entire factor. Gradient temperatures are usually measured under this latter condition where temperature measuring apparatus must have a scale extending over the range of expected temperatures, say 50&apos; to 25O°F, and it becomes almost impossible to detect small irregularities of less than one degree. The use of large strip chart recorders and divided multiple ranges tend to amplify the small anomaly and make it more pronounced. But the instability of instruments, both

Jan 1, 1952

In the special number of the magazine Ocl und Kohle (Oct. 22, 1940) dedicated to the petroleum industry of southeastern Europe, there appeared several articles relating to Yugos1avia.t In the first article the author calls attention to the fact that there are two fields in Yugoslavia that have been explored and exploited by more than 100 wells at variable depths during the last 50 years. These are on the Island of Mura, at the confluence of the Mura and the Drave. Additional exploration drilling in various other parts of the country has been without result, but certain ones of these explorations have revealed the existence of gas horizons that have some value, which are being used for industrial purposes. Table I shows the production of oil and gas between the years 1900 and 1939. TaBle —Production of Oil ad Gas in Yugoslavia . n,i Natural Natural Years oil Years Gas, Cu. Years Gas, Cu. Meters Meters T900-1929 10,000 1923 507,000 1934 1,388,000 1930 277 1924 567,000 1935 1,401,000 1931 202 I92S 691,000 I936 1,483,000 t932 289 19261 1.472,000 1937 1,841,000 1933 629 1927! 1,732,000 193s 2,431,000 1934 324 1928 2,515,000 1939 2,628,000 1935 260 1929 4,700,000 1936 137 I93O 5.345,200 1937 463 1931 6,376,153 1938 1,091 1932 1,661,737 1939 1,114 1933 959,574 The number of wells drilled from 1884 to the present time was in the neighborhood of zoo and the total footage drilled was 65,000 m. up to July I940 They may be divided roughly as follows: 12 wells in the region produce natural gas; 134 wells in the two old fields of the Island of Mura and about zo wells for the exploration of the zone of the Flysch, in Bosnia. The depths of 50 of these wells was roo to 200 m.; the depths of 50 wells oil the Island of Mura ran from zoo to 500 m.; of 2 ; wells, 500 to I000 m.; of ro wells from rooo to 2000 meters. The deepest well drilling in Yugoslavia has reached 2000 in. In the Valley of Morave, in the region of Selnica, I00 km. south of Nish, there have been discovered asphalt-bearing limestones, which contain about 15 per cent asphalt and are associated with bituminous clays. Later, at Paclenica, important deposits of natural gas were found. A zone rich in gas and oil seepages runs through Pracec, Miklenska, Bujavica, etc. The author studied these regions from the point of view of stratigraphy and tectonics. The anticlines and gas-bearing domes are concentrated in two zones: the region of the Drave (Island of Mura) and the region of the Save, southeast of zagreb, which contains the gas field in exploitation in the region of Bujavica. A full description of the productive region of the Island of Mura has been given by Egon Bohm, a geologist, in the second article mentioned. This includes a history of the exploration and exploitation that took place under Hungarian rule in this region. The most important exploration was effected between 1884 and 1885 by Stavenov and Stinger, of Vienna. In 1927 work was done in the same region by Anton Raky, a well-known explorer who makes work of this kind his specialty.

Jan 1, 1941

In the special number of the magazine Ocl und Kohle (Oct. 22, 1940) dedicated to the petroleum industry of southeastern Europe, there appeared several articles relating to Yugos1avia.t In the first article the author calls attention to the fact that there are two fields in Yugoslavia that have been explored and exploited by more than 100 wells at variable depths during the last 50 years. These are on the Island of Mura, at the confluence of the Mura and the Drave. Additional exploration drilling in various other parts of the country has been without result, but certain ones of these explorations have revealed the existence of gas horizons that have some value, which are being used for industrial purposes. Table I shows the production of oil and gas between the years 1900 and 1939. TaBle —Production of Oil ad Gas in Yugoslavia . n,i Natural Natural Years oil Years Gas, Cu. Years Gas, Cu. Meters Meters T900-1929 10,000 1923 507,000 1934 1,388,000 1930 277 1924 567,000 1935 1,401,000 1931 202 I92S 691,000 I936 1,483,000 t932 289 19261 1.472,000 1937 1,841,000 1933 629 1927! 1,732,000 193s 2,431,000 1934 324 1928 2,515,000 1939 2,628,000 1935 260 1929 4,700,000 1936 137 I93O 5.345,200 1937 463 1931 6,376,153 1938 1,091 1932 1,661,737 1939 1,114 1933 959,574 The number of wells drilled from 1884 to the present time was in the neighborhood of zoo and the total footage drilled was 65,000 m. up to July I940 They may be divided roughly as follows: 12 wells in the region produce natural gas; 134 wells in the two old fields of the Island of Mura and about zo wells for the exploration of the zone of the Flysch, in Bosnia. The depths of 50 of these wells was roo to 200 m.; the depths of 50 wells oil the Island of Mura ran from zoo to 500 m.; of 2 ; wells, 500 to I000 m.; of ro wells from rooo to 2000 meters. The deepest well drilling in Yugoslavia has reached 2000 in. In the Valley of Morave, in the region of Selnica, I00 km. south of Nish, there have been discovered asphalt-bearing limestones, which contain about 15 per cent asphalt and are associated with bituminous clays. Later, at Paclenica, important deposits of natural gas were found. A zone rich in gas and oil seepages runs through Pracec, Miklenska, Bujavica, etc. The author studied these regions from the point of view of stratigraphy and tectonics. The anticlines and gas-bearing domes are concentrated in two zones: the region of the Drave (Island of Mura) and the region of the Save, southeast of zagreb, which contains the gas field in exploitation in the region of Bujavica. A full description of the productive region of the Island of Mura has been given by Egon Bohm, a geologist, in the second article mentioned. This includes a history of the exploration and exploitation that took place under Hungarian rule in this region. The most important exploration was effected between 1884 and 1885 by Stavenov and Stinger, of Vienna. In 1927 work was done in the same region by Anton Raky, a well-known explorer who makes work of this kind his specialty.

Jan 1, 1941

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 Philip R. Strand, O. F. Stewart

Vermiculite is the name generally applied to the group of hydrated ferromagnesian aluminum silicates that are characterized by the ability to expand when heated. This process, called exfoliation, results in a lightweight product of commercial value. Most uses of vermiculite are for the expanded material. The chief markets are in construction, agriculture, and horticulture, with lesser uses in general industry. Composition and Properties Vermiculite, in its natural state, has the characteristic micaceous habit, a perfect basal cleavage which causes splitting into thin laminae that are soft, pliable, and inelastic. The structure of vermiculite is basically that of a talc. The prominent monoclinic crystal faces are often marked by lines at 60° and 120°. Hardness varies from 1.5 to 2 or more; specific gravities are between 2.1 and 2.8; color varies from almost clear to amber, bronze, brown, green, or black. Vermiculite feels like talc, especially when wet. Although much research has been performed on the chemical and structural composition, there is not yet complete agreement on the exact formula. This is to be expected when different workers have examined the many different varieties. Vermiculite is not considered to be a single mineral species but a family of related minerals. The structural formula for a trioctahedral vermiculite may be written: (H20)-(Mg,Ca,K)-(A12,Fe,Mg)-(Si,A1,Fe)4O10 (OH) 2 Hydrobiotite also occurs with vermiculite and is usually considered a vermiculite for commercial uses. When heated quickly to elevated temperatures, vermiculite expands by exfoliating at right angles to the cleavage into wormlike particles. The name vermiculite is derived from the Latin vermiculare, to breed worms. This characteristic of expansion is the result of the mechanical separation of the layers by the rapid conversion of contained water to steam. The decrease in bulk density of commercial grades is usually approximately 10 times, from 50 to 5 lb/ft3, but varies depending on the quality, size, and furnace efficiency. Individual flakes may expand up to 30 times. Vermiculite may also be expanded by soaking in chemicals such as hydrogen peroxide, weak acids, and other electrolytes. Color change during expansion is dependent upon the type of vermiculite and furnace conditions. Heating in an oxidizing atmosphere produces a dull gray or tan color, whereas a reducing atmosphere can produce a bronze or gold color. The expansion of the vermiculite crystal results in large pores being formed between the platelets. Thus, exfoliation makes available a large increase in void volume which is important in the application of vermiculite as a chemical carrier. Mathieson and Walker (1954) state vermiculite must be regarded as a true clay mineral. The characteristic properties of the mineral, such as high cation exchange capacity, organic complexing ability, and variable interlamellar distance are very similar to those of montmorillonite. The cation exchange capacity of vermiculite is one of the highest of all the clay minerals. The interplatelet space is accessible to penetration by some electrostatically neutral molecules. Water and glycerine are two common substances whose molecules may be so imbibed. In the natural state and under normal atmospheric conditions, water occupies the spaces between the silicate layers. The crystal d-spacing is near 14.2A. By differential thermal analysis, it has been determined that the water

Jan 1, 1983

By Henry P. Chandler

Tripoli is the general name for a number of fine-grained, lightweight, friable, minutely porous, forms of decomposed siliceous rock, probably derived from siliceous limestones or calcareous cherts. Tripoli is often confused, both in the trade and technical literature, with tripolite, a diatomaceous earth (diatomite), found in Tripoli, North Africa. A material found near Seneca, Mo., resembling tripolite was incorrectly termed "tripoli." This name has persisted and now has a definite physical and trade meaning. However, tripoli contains no diatoms, is of an entirely different origin, and has chemical and physical properties different from those of tripolite. There are two distinct types of tripoli in the United States; that from the Missouri-Oklahoma field and that from the Southern Illinois field. Commercially the term "tripoli" is applied to the material from the Missouri-Oklahoma field, while the material from the Southern Illinois area is often referred to as "amorphous" or "soft" silica. Material similar to the Missouri-Oklahoma tripoli, but finer grained, occurs in northwestern Arkansas. In its Minerals Yearbook the Federal Bureau of Mines groups statistically under "natural silica abrasives," tripoli from these producing areas and rottenstone from Pennsylvania. The justification for this grouping is the similarity of use. Composition and Physical Properties Commercial tripoli contains 97 to 99 pet SiO2. The remainder is mostly alumina, iron, lime, soda, and potash, some of which may impart desirable physical properties to the material. Variations in the chemical composition of tripoli are not as important as its physical characteristics. The color of tripoli varies from white to cream, brown, rose, and red, the variations from white being the result of iron deposited from percolating surface waters. The commercial varieties of tripoli mined in the Missouri-Oklahoma area are designated as cream and rose. Other physical properties of the tripoli from this area are: Free quartz, less than 0.5 pct; oil absorption, 50.71 to 53.14 pct; water absorption of the crude tripoli, about 38 pct, and of prepared material, about 52 pct; fracture, uneven to conchoidal; specific gravity, 2.15 to 2.62; weight per cubic foot crude, 35 lb, ground, 55 lb; fusion point, 3,225° to 3,300°F; and refractive index, 1.54. Tripoli occupies about 1 cu yd to the short ton in place and about 1.5 cu yd when broken. Packed for shipment, the powdered material occupies about 33 cu ft to the ton. Its value and usefulness depend primarily on a low content of crystalline silica, size of the natural particles, absorption, and color. Some beds in the Missouri-Oklahoma area are of sufficiently compact material that large, fairly strong blocks may be removed. These find a market as filter blocks and have been used for building material. In the Southern Illinois field the color of the material is the important physical property; pure white is preferred. The rottenstone mined in Pennsylvania is of higher density and lower in silica content than the various types of tripoli. Its composition approximates 60 pct silica, 18 pct alumina, 9 pct iron oxide, 8 pct alkalies, with small amounts of lime and magnesia. The material apparently is a ferruginous and siliceous shale, having properties rendering it suitable as an abrasive or as a filter for paint, rubber, and plastics.

Jan 1, 1960

By R. C. Hall

The magnetic anisotropy and magnetostriction of single crystals of alloys between 25 and 59 wt pct Co in Fe have been determined in the disordered and ordered states. The magrzetostriction is large and positive for all alloys in both states of order (up to 150 x 10-6 for A,, and 40 x 10-6 for 111). The magnetic anisotropy becomes zero rzear 41 pct Co for the disordered state. Ordering shifts the zero-anisotropy composition to about 50 pct Co. COBALT-iron alloys have for many years been of interest as magnetic materials because they have unusually high magnetic saturation values.&apos; This interest was increased recently by the development of Supermendur.&apos; Supermendur is an alloy of 49 pct Co-49 pct Fe-2 pct V* which was found by suitable purification and heat treatments, including a magnetic anneal, to possess excellent magnetic properties. These included a square hysteresis loop with a low coercive force and a high magnetization, and a low total loss. Two factors which are important in determining the magnetic hysteresis properties are the basic magnetic properties of magnetostriction and anisotropy. Alloys having zero anisotropy and zero magnetostriction show improved hysteresis properties.3 Since very meager data4-&apos; are available in the literature on the single-crystal anisotropy and magnetostriction of the cobalt-iron alloys, the present program was undertaken to determine these two properties in the disordered and ordered states. EXPERIMENTAL TECHNIQUES Ingots of the cobalt-iron alloys containing from 25 pct to 59 pct Co in iron were vacuum melted from electrolytic iron and cobalt of 99.9 pct purity. The ingots were cooled at a rate of about 1°C per hr through the transformation temperature (950" to 980°C dependent on the alloy content) for the growth of large grains. From these grains, single-crystal disks were fabricated with either the (110) or (100) plane parallel to the faces of the disk. The crystals were given two heat treatments to produce either the ordered state (cooling rate of 20°C per hr from 850 "C) or the disordered state (water quench from 850 °C). Since the iron-cobalt alloys order very rapidly, the water quench may not have produced complete disorder in all cases. Due to the method of growth, the crystals were not perfect. Most of the samples had low-angle boundaries and a few had some small included grains. These imperfections could and probably did affect the accuracy of both the magnetic anisotropy and magnetostriction measurements. Much of the scatter in the data may be attributed to the crystal imperfections. The methods of determining the spontaneous saturation magnetostriction and the anisotropy have been previously described.419110 MAGNETOSTRICTION The magnetostriction in the [loo] and [Ill] directions (hO0 and ?111 respectively) is illustrated in Fig. 1 for the disordered state. Both ?100 and ?111i became large positive values as cobalt was added to iron. The state of order did not have a large effect on the magnetostriction. The effect of order did not appear too consistent in these data. Tentative curves are shown in Fig. 1 for the ordered magnetostriction near the FeCo composition; here ?100 appeared to be lowered and ?111 raised as the degree of order was increased. These results agree qualitatively with the

Jan 1, 1961

THE strike and dip of bedded formations is readily determined by trigonometry if some marker horizon is intercepted by three holes. If no marker can be identified and if orientation of the cores is unknown, as is the usual case in diamond drilling, the determination still can be made readily by the use of a stereonet if the bearing, inclination, and angle the bedding makes with the core axis is known for each of the three holes. Such determination, valid only if the bedding orientation is fairly uniform within the area drilled, is considered simpler, faster, and more effective than any solution previously published. Method: The steps of this procedure are as follows: 1) Mark north and south poles on tracing paper placed over stereonet.* 2) Plot the three diamond drillholes by drawing a line from the center of the net in the direction of the bearings of the holes. Then rotate the paper until each line, in turn, lies along the equatorial diameter of the net; the end of each line is determined by the inclination of the hole measured in degrees from the periphery toward the center of the net along the equatorial diameter. Fig. 1 illustrates a vertical hole and a hole inclined at —23" due east. 3) With the projection of one diamond drillhole along the equatorial diameter, measure from the end of the hole, to the east and then to the west, a dis- tance in degrees equal to the core axis-bedding plane angle. Find two more points by measuring the same number of degrees north then south along the great circle from the end of the drillhole. Draw a circle, the center somewhere along the direction of the projection of the drillhole, to connect these four points. Draw circles for the other two diamond drillholes in the same manner. See Fig. 1 for two circles representing core axis-bedding plane angles of 35" and 23" and Fig. 2 for these plus a third circle representing a core axis-bedding angle of 42" in a hole bearing S 50 W inclined at —52". 4) Rotate the tracing paper until one great circle on the stereonet is found which is tangent to all three of the circles or parts of circles derived in the previous step. Trace this great circle on the paper and join its ends with a straight line passing through the center of the net. The bearing of this line is the strike of the bedding. In Fig. 2 the bedding strikes N 77 E. 5) Rotate the paper until the strike is along the polar diameter of the net. The dip of the bedding is determined by measuring along the equatorial diameter the distance in degrees from the periphery of the net to the great circle derived in step 4. The dip of plane illustrated in Fig. 2 is 55" to the N 13 W. Fig. 3 indicates the solution for three holes whose bearings and inclinations are N 20 W and —43o, due E and —30°, S 30 W and 23" with core axis-bedding angles of 30°, 50°, and 43" respectively. The bedding (stippled area) strikes N 44 W and dips 70" to the southwest.

Jan 1, 1956

By Leonard B. Gulbransen, John R. Lewis, W. Martin Fassell, J. Hugh Hamilton

T. E. Leontis (The Dow Chemical Co., Midland, Mich.)—This paper is of particular interest to me because of my own work with F. N. Rhines on the oxidation of magnesium and magnesium alloys a few years ago. The authors are to be complimented on their development of an accurate and reproducible technique for measuring ignition temperatures and on their comprehensive study of the many variables that affect the ignition temperature of magnesium. It is indeed gratifying to see that they have obtained a good correlation between ignition temperatures and the oxidation rates reported by us. The correlation is valid not only with composition within one alloy system but also between alloy systems; that is, alloying elements which effect the greatest increase in oxidation rate also produce the greatest decrease in ignition temperature. There are a few points upon which I would like to comment. In attempting to correlate ignition-temperature data, one must be sure that the same definition of this quantity is used by all investigators. It does not appear to me that such is the case in the authors&apos; comparison of their data with the theoretically calculated values of Eyring and Zwolinski. The equation derived by these investigators defines the ignition temperature, To, as the temperature at the gadoxide interface, whereas the present authors use the metal temperature as the criterion for ignition. The contradiction in the effect of oxide-scale thickness on ignition temperature between the predictions of the Eyring-Zwolinski equation and the observations reported in this paper indicate that some variable has not been taken into consideration. Could that be the geometry and size of the specimen? There is a marked difference in the type of specimen used in this investigation and that used in our work which formed the basis of Eyring and Zwol-inski&apos;s theoretical treatment. Another factor which plays an important role in ignition is the vapor pressure or the rate of vaporization. Combustion can safely be assumed to take place in the vapor phase by the reaction between vaporized magnesium and oxygen. Thus, a more accurate theoretical analysis may be made on the basis of the rate of vaporization which may be the controlling rate of the process. The effect of a large number of alloying elements on the ignition temperature has been reported in this paper, but beryllium was not included. Practical experience dictates that beryllium markedly decreases the burning tendency of magnesium. I was wondering if the authors plan to study the effect of beryllium in their future work. The authors predict that concentrations of sulphur dioxide in the furnace atmosphere greater than 5.8 pct would be expected to increase the ignition temperature to values still higher than those they measured. I would like to mention that large concentrations of sulphur dioxide markedly increase the rate of combustion of magnesium once ignition has started. Although it has been shown in the paper that the ignition temperature of magnesium in oxygen increases with increasing sulphur dioxide content up to about 1 to 2 pct, in practice relatively low-melting commercial cast alloys (AZ63A and AZ92A) are being continuously heat treated at temperatures just below the melting point in air containing 0.5 to 0.75 pct SO*. In regard to the change in color of the oxide scale observed on magnesium and magnesium alloys just prior to ignition, I would like to mention that in our work alloying elements were found to color the usually white magnesium oxide even though ignition did not occur. For example, the oxide formed on Mg-A1 alloys was gray, increasing in intensity with aluminum content in the alloy. Finally, I might suggest that the authors indicate their source of the value of 0.8 g per cc for the density of MgO as it is formed on magnesium upon oxidation at elevated temperatures. W. M. Fassell, Jr. (authors&apos; reply)—The comments by Dr. Leontis are very excellent ones and I will attempt to answer them in order. First, the problem of ignition of magnesium is a rather difficult one since many factors are involved. Concerning the comparison of the To in the Eyring-Zwolinski equation, eq 4, with the experimentally determined values, it will be noted that the calculated and experimental values of the ignition temperature in Table I are not self-consistent. In the case of the 1.78 pct A1-Mg alloy the calculated value is 49°C below the experimental value; for the 3.81 pct A1-Mg alloy, 122°C below the experimental value; for Mg with 5x10-&apos; cm film, 19°C above the experimental value; for Mg with 2x10-I cm film, 28 °C below the experimental value. Thus, if it were merely a matter of difference of location of temperature measurement the calculated ignition temperature would always be below the experimental value, the difference being due to the thermal gradient through the oxide film. The possibility of a thermal gradient in the magnesium metal must be considered. From Carslaw and Jaeger,&apos;Y t can be shown that the maximum temperature gradient that could exist between the oxide-metal interface and the center of the sample is of the order of O.Ol°C. The geometry and size of the specimen could certainly have some effect on the ignition temperature. The equation for ignition that has been proposed in reference 14 is of the following type containing terms to account for this and other factors: M dT AHv(T) =Cp--------------\-J(.T—TB) + ZAHl-M A dx where AH is the heat of reaction, v(T) is the velocity of the reaction at temperature, Cp is the heat capacity of sample, M is the mass of sample, A is the area of sample, t is time, J is the total coefficient of heat transfer outward from the reaction zone, TR is the temperature of the bath or furnace, and AH,, is the heat associated with any phase change involved. Prior to the instant of ignition, the vapor pressure of magnesium is of no special significance. After ignition, neither eq 4 nor the above equation is applicable. The actual combination of magnesium cannot safely be assumed to take place in the vapor phase. While experimental data is lacking to support a hypothesis that ignition does or does not occur in the vapor phase, some observation on the pressure ignition experiments may be of interest. At high oxygen pressures, once ignition has occurred, the reaction of magnesium with oxygen approaches near explosive violence, the entire sample being consumed in probably less than 1 sec. At atmospheric pressure it usually requires 15 to 20 sec. Thus it appears that the oxygen concentration becomes the rate determining factor. Further, if burning magnesium is observed through darkened glass (Lincoln Super-visibility Shade No. 12) the magnesium sample is very much hotter than the "smoke" and the outline of the sample is retained perfectly. No "flame" is visible above the metal. No work was done on Mg-Be alloys. We do, however, intend to study this problem in the near future.

Jan 1, 1952

By R. C. Gould, A. M. Skov, L. F. Elkins

Comparison of long term decline in oil production during cyclic waterflooding or pressure pulsing of part of the Driver Unit with steady injection-imbibition flooding in the Tex Harvey area led to large expansion of flood in the Driver Unit on the steady injection basis. While the flood has been successful, the major problem has been attainment of satisfactory oil production rates in most of the wells. Large volume fracture treatments of low capacity wells were unsuccessful in achieving sustained increases in production. A two-section area in the Driver Unit has already recovered 620 bbl of oil per acre by waterflood but other areas have not performed so well. Sun Andres water containing 300 to 500 ppm H,S is sweetened to 0.5 to 1 ppm H,S by extraction with oxygen-free flue gas. This prevents contamination of gas produced in the area and apparently it has reduced corrosion in minimum investment, thin-wall, cement-lined water dktribution systems. Cement-lined tubing in injection wells has mitigated corrosion as effectively as thick polyvinyl chloride films have, and at less cost. Introduction As reported in the literature the Spraberry field of West Texas has presented unusual problems for both primary production and waterflood ing. Earlier information from the Spraberry Driver Unit included conception and evaluation of cyclic waterflooding or pressure pulsing in a nine-section pilot test as an aid to extraction of oil from the tight matrix rock and as a boost to normal capillary imbibition forces An additional 5 years&apos; operation in that area, and performance of expanded steady injection water-flood, now covering a total of 68 sq miles, are reported herein. In addition, since the Driver Unit is one of the largest waterfloods in areal extent in the U. S., many operating experiences are presented for the benefit of engineers concerned with operation of other Spraberry floods or with other waterfloods where this reservoir technology and/or water handling technology may be adaptable in part. These include: (1) attempts to improve producing well capacity through large volume fracture treatments, (2) long-term performance of water treating plants utilizing oxygen-free flue gas to extract H,S from sour San Andres water, (3) performance of thin-wall cement-lined pipe in water distribution systems including comparison between those sections carrying raw San Andres water and those carrying treated water, and (4) comparison of performance of various lining materials and subsurface equipment in water supply and water injection wells. These experiences are reported without regard to whether results are good, bad or indifferent. Since the operations reported are limited to the techniques, materials, and equipment actually used in the Driver Unit, no comparison is possible with results of other approaches used in other Spraberry floods or in waterfloods generally under different conditions. However, an attempt is made to quantify these experiences as much as possible in the space available to permit other engineers to select those parts applicable to eheir problems. Background The Spraberry, discovered in Feb., 1949, is a 1,000-ft section of sandstones, shales and limestones with two main oil productive members—a 10- to 15-ft sand near the top and a 10- to 15-ft sand near the base, having permeabilities of 1 md or less and porosities of 8 to 15 percent. Extensive interconnected vertical fractures permitted recovery of oil on 160-acre spacing from this fractional-millidarcy sandstone, but they made capillary end effects dominant. Primary recovery by solution gas drive is less than 10 percent of oil in place, with most wells declining to oil production of a few barrels per day when reservoir pressures are still in the range of 400 to 1,000 psi. Partial closing of the fractures with declining reservoir pressure is believed to be the cause of such low production rates at these relatively high reservoir pressures. In 1952 Brownscombe and Dyes proposed that displacement of oil by capillary imbibition of water from the fractures into the matrix rock might significantly increase oil recovery from the Spraberry, overcoming otherwise serious channelling of water through the fractures." A pilot test conducted by the Atlantic Refining Co. during 1952 through 1955 indicated technical feasibility of the process; but low oil production rates averaging I5 to 20 bbl/well/D failed to create significant interest in large-scale waterflooding at that time." Humble Oil & Refining Co. conducted a highly successful 80-acre pilot test during 1955 through 1958 with

Jan 1, 1969

By John C. Martin

The discovery of a new gas reservoir demands that the planning of a sottnd well-spacing program be initiated early in the development stage. It is the purpose of this discussion to illustrate by actual field examples the application of basic well-spacing principles, previously developed for oil reservoirs, to the problem of well spacing in natural gas fields. These studies are presented for the field use of geologists and engineers who are concerned with the initial planning of the proper development of the newly discovered gas reservoir. INTRODUCTION The phenomenal growth of a vigorous natural gas industry emphasizes the increasing importance of natural gas as a source of energy, fuel, and raw materials to our nation&apos;s economy. Since 1945 marketed production of natural gas for the U. S. has increased 21/2 times to a record high of 10.6 trillion cu ft during 1957. As major participants in the gas industry, we share an added interest to develop and produce our natural gas reserves with constantly improved efficiency. The subject of well spacing is vitally important to the gas industry, for the well itself plays a significant role in the development of the natural gas reservoir and in control of the recovery process. Maximum utilization of wells is an integral part of sound conservation practices. The discovery of a new gas reservoir demands that careful choice of well location and well spacing be made and that the planning of a sound well spacing program be initiated early an the development stage. With the drilling of the first development well, efforts of the geologist and engineer must be directed toward acquisition of adequate technical evidence upon which a firm recommendation for a spacing program may be based. With this technical appraisal as a foundation, operators and state regulatory agencies jointly can go far in providing a framework for sound development of gas fields to achieve a program of conservation that avoids the unnecessary well. WELL-SPACING CONCEPTS Through laboratory and field investigations of the mechanism of the recovery of oil and gas, of fluid behavior, and of effective control of reservoir and well, a crystallization of ideas regarding reservoir behavior has emerged as a well-developed technology. Associated with a better understanding of the fundamental principles underlying reservoir and well behavior has been the growth of concepts concerning the role of wells and their spacing in the development and operation of an oil or gas reservoir. In addition to serving as outlets for the withdrawal of fluids from the reservoir, wells are recognized as having two other important functions: (1) providing access to the reservoir to obtain information concerning the characteristics of the reservoir and its fluids, and (2) serving as a means by which the natural or induced recovery mechanism may be effectively controlled. Beyond a minimum number of wells required to fulfill these two functions, additional wells will not increase recovery. With particular emphasis upon well spacing in oil reservoirs, many studies of the well spacing-recovery relationship have evolved the concept that the ultimate oil recovery is essentially independent of the well spacing.&apos; These fundamental concepts are no different when regarding the role of wells and their spacing in the natural gas reservoir. They are equally applicable to the consideration of well spacing in gas reservoirs. For the gas reservoir, the problem of well spacing then revolves around the question of drainage and the degree or extent to which a well may drain gas from its surrounding reservoir environment. Theoretical and Experimental Work During the past 30 years, theoretical and experimental work carried on to study the physical principles involved in the flow of fluids through porous media has shed light upon the matter of drainage. Fundamental mathematical equations have been derived to describe the mechanism of flow of oil and gas through porous rocks. With the recent advent of high-speed digital computers, attempts have been made with mounting success to develop solutions, employing numerical techniques, to mathematical expressions that describe more rigorously the physical behavior and mechanism involved in the unsteady-state flow of compressible fluids, such as a gas, through porous rock. In 1953, Bruce, Peace-man, Rachford and Ricez published a stable numerical procedure for solving the equation for production of gas at constant rate. The results of these calculations are significant with respect to this matter of drainage, for they indicated (1) that depletion of the gas reservoir resulted in a drop in pressure at the extremity of the

By Ralph B. Stevens

INTRODUCTION Underground fires continue to be one of the most serious hazards to life and property in the mining industry. Although underground mines are analogous to high-rise buildings where persons are isolated from immediate escape or rescue, application of technology to locate and control fire hazards while still in their controllable state is slow to be implemented in underground mines. Even in large surface structures such as hotels, often only fire protection systems which meet minimal laws are implemented due to the high cost of adding extensive extinguishing systems, isolation barriers, alternate ventilation, escape routes and alarm systems. Incomplete and ineffective protection occasionally is evidenced where costs would not seem to be a factor, such as the $211 million MGM Grand Hotel fire November 21, 19801. Paramount in increasing fire safety and decreasing the threat of serious fire is early warning followed by proper decision analysis to perform the correct action. However, very complex fire situations can be produced in structures such as high-rise buildings and underground mines simply because of the distances between the numerous fire-potential locations and fire safe areas. Other complexities arise when normal activities occur that emit products of combustion signaling a fire condition to a sensitive fire/smoke sensor. For example, the operation of diesel equipment or the performance of regular blasting can produce combustion products that reach the sensitive alarm points of many sensors2. Smoke detectors for surface installations provide fire warning when occupants are at a distant location or when sleeping, thus greatly reducing injuries and property damage. However, when installed in the harsh environments of underground mines, fire and smoke detection equipment soon becomes inoperative, unreliable, or requires excessive maintenance. The U.S. Bureau of Mines has performed many studies and tests to improve fire and smoke protection for underground mine workers3. This paper describes several USBM safety programs which included in-mine testing with mine fire and smoke sensors, telemetry and instrumentation to develop recommendations for improving mine fire safety. It is hoped that the technology developed during these programs can be added to other programs to provide the mining industry with the necessary fire safety facts. By recognizing fire potentials and being provided with cost-effective, proven components that will perform reliably under the poor environmental conditions of mining, mine operators can provide protection for their working life and property equal to that which they provide for themselves and their families at home. The basis of this report is two USBM programs for fire protection in metal and nonmetal mines4,5 and one coal program6. The data was collected beginning in May 1974 and continuing through the present with underground tests of a South African fire system installed at Magma Mine in Superior, Arizona, and a computer-assisted, experimental system at Peabody Coal Mine in Pawnee, Illinois. The conduct of each program was as follows: • Define the problem and its magnitude in the industry • Develop concepts to solve or diminish the problem • Review available hardware or systems approaches to fit the concepts • Install and demonstrate the performance of a prototype system through fire tests in an operating mine. MINE FIRE FACTS Whether in coal or metal and nonmetal mines, the potential severity of fire hazard is directly related to location. As shown in Figure 1, fire in intake air at zones A, B, C or D can cause contamined air to route throughout the mine quickly if not detected, isolated or rerouted. Causes and location of former metal and nonmetal fires are represented in Table 1; the cause and location of fatalities and injuries is shown in Table 2. Coal-related fires and their impact on deaths and injuries are graphed in Figure 2; their locations are described in Table 37. Significantly the table shows that the hazard to personnel was three times greater for fires occurring in shaft or slope areas, and the percentage of deaths and injuries was four times that of other areas. Number of Persons Affected A 129-mine sample indicated that from 8 to 479 employees per shift work in underground metal and nonmetal mines, and that deeper mines have larger populations, as shown in Figure 3. Coal mining relates similar employment, and a 16-state sample of 670 mines employing at least 25 persons shows the distribution in Figure 4. Drift mines accounted for 58 percent of the sample but employ only 45 percent of the underground workers.

Jan 1, 1982

By C. Atkinson

A method is described for the numerical solution of the diffusion equation with a composition-dependent diffusion coefficient and applied to the radial growth of a cylinder; the radial growth of a sphere, and the symmetric growth of an ellipsoid. Sample applications of the method are made to the growth of particles of proeutectoid ferrite into austenite. RECENTLY&apos; we described a method for numerical solution of the diffusion equation with a composition-dependent diffusion coefficient for the case of the growth of a planar interface. In this paper we extend this method to describe the radial growth of a cylinder, the radial growth of a sphere, and the symmetric growth of an ellipsoid. In the latter case, limiting values of the axial ratios of the ellipsoid reduces the problem to one of a cylinder, a sphere, or a plane depending on the axial ratio. A check on these limiting values is made in the results section. In all of these cases we consider growth from zero size. A natural consequence of this assumption as applied to the sphere, for example, is that the radius of the sphere is proportional to the square root of the time. This is consistent with the condition that the radius is zero initially, i.e., grows from zero size. It may be argued that it is more realistic to consider particles which grow from a nucleus of finite initial size; even in this case the analysis of this paper is likely to be applicable. This can be seen if a comparison is made of the work of Cable and Evans,2 who consider a sphere of initially finite size growing by diffusion in a matrix with a constant diffusion coefficient, with the results of Scriven3 for growth from zero size. This comparison shows that the rates of growth in each case differ trivially by the time the particle has grown to about five times its initial size." This investigation is a generalization of those of Zener,4 Ham,5 and Horvay and cahn6 to the situation often encountered experimentally, in which the diffusion coefficient varies with concentration. First let us consider each of the cases separately. I) GROWTH OF SPHERICAL PARTICLES FROM ZERO SIZE In this case the differential equation in the matrix depends only on R, the radius in spherical coordinates, and can be written: ? 1 <^\ ^13D . , dt U\dRz + R 3Rj + dR dR [ J where C is the composition, t is the time, and D is the diffusion coefficient which depends on c. The boundary conditions will be: c = c, at the moving interface in the matrix, c = c, at infinity in the matrix (and at t = 0, everywhere in the matrix), c = X, is the composition in the spherical particle. Each of the above compositions is assumed constant. In addition there is the flu condition at the moving interface which can be written: , dR0 ~/3c dt \dR/H =Ra where R,, which is a function of t, is the position of the moving interface. We make the substitution q = RI~ in [I] reducing this equation to: & - m - *ws) »i where we have written D = D,F(c) or simply D,F, and Do = D(c,). Thus F[c(q0)] = 1 where q, = ~,/a is the value of the dimensionless parameter q evaluated at the interface. Multiplying Eq. [2] by dq/dc and integrating, we find: where the lower limit of the integral has been chosen so that dc/dq — 0 as c — c,, thereby satisfying the boundary condition at infinity. We require, then, to solve Eq. [3] subject to the condition c = c, when q = q, (this follows from putting R = R, at the interface) together with the flux condition which can be rewritten in terms of q as: Eqs. [3] and [4] together with the condition c = c, at q = q0 enable us to find 77, and the concentration profile c = c(q). Numerical Method. We treat Eq. [3] in the same way as we did the corresponding equation for the planar interface problem&apos; i.e., by dividing the interval c, to c, into n equal steps so that: cr = ca -rbc [5] where r takes the values 0, 1, ... n and we call no,, q1, ... nn the values of n corresponding to the compositions c,, c,, ... c,.

Jan 1, 1970