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By AIME AIME

WILBER JUDSON is one of that fairly large group of mining engineers that graduated at an Eastern college, worked his way up in various jobs in the West and in the Latin-American countries, and finally came to New York as an administrative officer of a large mining company. Since coming to New York he has been a loyal friend and counselor of the A.I.M.E., and is now serving a second term as one of its Directors, his first term covering the period 1935-1938. Mr. Judson was born in the capital city of Michigan on July 26, 1880. Probably it was his roving spirit rather than professorial urging that was responsible for the fact that four colleges were responsible for his education: Northwestern, University of Michigan, Harvard (S.B., 1901), and the Michigan College of Mines (graduate study). In his first ten years of professional work he might have been found serving as mine foreman at Catamarca in the Argentine, as superintendent of the Inter Ocean mine at Sunshine, Colo.. as foreman for the Cerro Prieto, Magdalena, Sonora, as superintendent of the Hudson Iron Co. in New York, or as manager of the Dolores mine in Chihuahua for the Mines Co. of America, or in Spain, or in Panama.

Jan 1, 1942

The AusIMM came into being to foster the spread of professional knowledge within its client industry. It has evolved along with the Australian mining industry for more than a century. If the AusIMM is to become the world's premier mining professional body (whether by design, or by default), then draw up a strategic plan to get there. Define your strengths and market them Define the priority areas for professional competence which we most need, Target the groups within Australia who can produce the people, Create an industry peer group to speed technology transfer.

Jan 1, 1994

By Luther Wagoner

HAVING given in a former paper1 the results of assays of sea-water, bay-mud, dredgings from San Francisco bay, etc., and' believing it might be interesting to extend the work to include' some deep-sea dredgings, I procured from the Smithsonian Institution six samples taken by the U. S, steamer Albatross, and marked as follows

Jan 9, 1907

By Gary E. Geiger

laboratory studies were performed to assess the calcium carbonate inhibition performance of phsophonates and polymeric dispersants at high hardness and high pH conditions. Based on the results of these studies, a new calcium carbonate treatment program was developed that shows excellent Inhibition characteristics and Is not sensitive to calcium hardness.

Jan 1, 1998

By C. E. Turner, J. R. Elenbaas, R. D. Grimm, J. A. Vary, D. L. Katz

A technique is presented by which well samples and core plugs of dolomite formations are classified by microscopic examination into seven different porosity grades. Quantitative values of porosity and permeability are determined for each grade by a statistical correlation of the core plug test data with the porosity grading system. These quantitative values are applied directly to the grades exhibited in the well samples for the purpose of estimating the reservoir void space for wells that were not cored. The procedure is described for estimating the gas reserves per unit area lor the South Hugoton gas field. but a reserve estimate for the field is not given. INTRODUCTION The miscroscopic examination of well sample; and the graphic recording of their lithologic qualities and other distinguishing characteristics of various geologic formations drilled is both a science and an art of long standing and wide application. Usually the primary objective of a geologist who "sits on the well" and examines the samples are: to identify the formation being drilled, determine the total depth, casing point. and completion interval. In most cases the porosity is described. if done at all, in general terms. such as: trace, scattered, fine, poor, fair, medium. good, excellent, or in some other relative terms. In fields where various geologists have examined samples and recorded observations on many wells considerable variations in lithologic terms and porosity descriptions occur unless there is primary effort to establish uniformity of logging observations and standards of recording observable porosity. When an estimate of the pore volume of a reservoir is made a geologic concept of the processes that control the magnitudes of porosity and permeability is developed by microscopic examination of well samples. The characeristics and appearances are then mentally related to rather general quantitative units of porosity based on physical core data from the same reservoir or on such data or experience in other reservoirs that have similar qualities. The reliability of such estimates depends largely on the variations of the lithology of the formations, the geometric properties of its void system. the extent of comparisons of sample appearances with porosity data, as well as the uniform recording of all relevant characteristics. This statement is particularly significant for dolo-mitized limestone formations of substantial thicknesses and heterogeneity such as the Permian Dolomites of the Hugoton gas field. Jn this field, as well as in most of the Permian Dolomite fields, the producing formations are of relatively great thicknesses in which the porosity and permeability of the reservoir varies substantially in all directions, depending on the crystalline structure. degree and kind of impurities, kind of fossils anti cementation thereof, degree of dissolution. and fracturing. The variations of the lithologic texture of the dolomites and post deposition alterations have resulted in porosities and permeabilities of such magnitudes that only a part of the gross thickness can be counted as "pay." At the time of this study insufficient gas production had been experienced to apply the pressure decline production method in the South Hugoton Field and the electric logs are not definitive enough. The problem of estimating gas reserves in the south part of the Hugoton Field is primarily one of determining the pay thickness and porosity from well samples and core data. The area studied embraced all that part of the field lying south of an east-west line through Guvmon, Okla., anti containing approximately 1.000,000 acres. This paper describes a technique of correlation of physical core data with well samples so that quantitative values of pay thickness, porosity. Permeability, and connate water may be assigned to well samples that are representative of a given interval, and thereby permitting the estimation of gas reserves lor a given unit area. The procedure was developed by a uniform microscopic qualitative porosity grading of the dolo. mite core plugs.. and relating these grades to the respective physical core data on a statistical basis The well sample-were also graded in a Similar manner in order that the quantitative values established lor the core plugs could be applied to the well Sample for wells that were not cored. GRADING OF DOLOMITE A group of experienced geologists was given the assignment of examining the samples on all wells in South Hugoton in order that they could log their observations in a uniform and standardized manner and grade the observed porosity so that it could be related quantitatively to the core data. The group initiated the study 011 chips from cores which bad been tested for porosity and permeability. This study continued until all of the geologists developed a common knowledge of lithologic terms and of the characteristic appearances of the samples and their relations to measured porosity. The characteristic appearances of the dolomite samples under twelve-power magnification as related to their qualitative porosities afforded a classification of the dolomite into :even grades of porosity, ranging from dolomite of no-visible porosity under twelve-power magnification to dolomite of excellent porosity. The assigned grade for a specific 10-ft interval is a weighted average of all visible grades of porosity exhibited by the cuttings representing that interval. The porosity characteristics were recorded by a color graph adjacent to the lithology column in conjunction with a numerical system for further definition of relative porosity as shown in Fig. 1. The three vertical lines to the right of the lithology column each represent 33 1/3 per cent. which lines were used to record the percentage of the samples, for any particular interval. that showed porosity under the microscope. The colors were used to denote actual pore size. i.e., orange. blue and I-ed for pore diameter of one-fourth or less. one-fourth to one-Ilalf. and greater than one-half millimeter, respectively. The area colored 1)). one or more colors represents the percentages of the samples exhibiting pores of the respective size or sizes. The numerals from one to six inclusive shown on the log in

Jan 1, 1952

By J. Weertman

The paper of Gifkins and Snowden considers the interesting but difficult problem of determining the stress dependence of secondary (steady-state) creep at low stresses. These authors have concluded that at stresses below 250 psi (2 x 107 dynes per sq cm) the secondary creep rate of lead is proportional to the stress (viscous creep) and is not proportional to the stress raised to about a fifth power. The experimental data considered by them were obtained on tests conducted at room temperature and at 50°C. The lowest stress employed was 50 psi (3.5 X 106 dynes per sq cm). The authors pointed out the main difficulty in determining the stress dependence of creep at low stresses. The creep tests must be run for very long lengths of time. However they made no estimates of when an experimental creep rate determination must be rejected because it does not represent a true steady-state or minimum creep rate. In order to be certain that a creep curve is within the steady-state region, the total creep strain should be of the order of 0.1 to 0.2. For a creep test of a year's duration this requirement implies that a secondary creep rate smaller than about 10-3 per hr cannot be measured reliably. The corresponding creep rate for a 10-year test is 10-6 per hr. The creep rates of the tests that were considered by the authors to prove the existence of viscous creep were of the order of or less than 10-6 per hr. One can conclude reasonably that this data does not prove unambiguously that large strain steady-state creep rate of lead is proportional to the stress in the stress range of 50 to 250 psi (3.5 x 106 to 2 X 107 dynes per sq cm). Another technique can be used to obtain the stress dependence at low stress levels. The creep rate is a very sensitive function of temperature. The creep rate can be increased by very large amounts merely by increasing the temperature. We carried out steady-state creep tests on lead single crystals25 at temperatures up to 320°C. We were able to obtain creep rate data down to stresses as low as 35 psi (2.5 x 10' dynes per sq cm). Our smallest creep rate was 8 x 10-5 per hr. Thus we obtained large strain, steady-state creep rates to even lower stresses than were considered by Gifkins and Snowden. No evidence was seen for viscous creep. The creep rate was proportional to the stress raised to about a 4.5 power down to the lowest stresses. Since there is no reason to believe that changing the temperature should change the stress dependence of steady-state creep, we feel that large strain viscous creep does not occur in the stress range quoted by the authors for lead single crystals or large-grain polycrystalline samples of lead. This conclusion does not imply that viscous creep may nat occur in a lower stress range or in the same stress range for fine grain material or at creep strains very much smaller than 0.1. Support by the U.S. Office of Naval Research is acknowledged. Authors' Reply R. C. Gifkins and K. U. Snowden We thank Dr. Weertman for his discussion and although, as we hope to show, we do not agree with his reservations, we do concur in stressing the importance of ensuring that creep rates are reliably obtained. Dr. Weertman appears to be content to accept n = 1 for low stresses with fine-grained material but not for single crystals. We believe our results show that the former result cannot be accepted without also accepting the latter. We will also show that the probable errors in our minimum creep rates are insufficient to alter our conclusions, that the criterion proposed by Dr. Weertman is arbitrarily restrictive and his alternative experimental approach possibly invalid. 1) A principal result of our Fig. 1(a) is that n = 1 for polycrystalline specimens at room temperature and 50°C for stresses below -250 psi. There was evidence that crystal slip and grain boundary sliding contributed approximately equaily to the overall strain in this low-stress regime. This implies that either a) grain boundary sliding controls slip within the grains or b) both grain boundary sliding and crystal slip independently occur according to mechanisms which give n = 1. Alternative a does not seem acceptable, so we were forced to consider b. This led us to reexamine work on bicrystals by Strutt and Gifkins and plot curves Fig. l(b). Previously Strutt et al. (loc. cit.), had merged these points with others using the Zener-Holloman parameter and thus, we now believe, had been led to overlook the behavior where n = 1. Curves C and D in Fig. l(b) did appear to confirm the hypotheses that n = 1 for crystal slip at these low stresses and the sliding curve F was similarly of the expected form. It was comparatively easy to find a quantitative theory to account for n = 1 for sliding and the similarity of curves C and D to curve F (all obtained from the same set of specimens) led us to feel that the single-crystal curves were valid. 2) We believe the secondary creep rates for both the polycrystalline and single-crystal specimens to be in error by factors c2. In Fig. 6 creep curves for polycrystalline specimens of lead(1) and lead(II) are reproduced as curves a and b, respectively, and curve c is for a single crystal at 100 psi. It is clear that, although the attainment of secondary creep rate takes 2 years for a and 150 days for b, thereafter the curve is linear for periods of 7 and -1 year, respectively. The single crystal has a linear portion commencing after 20 and extending to 90 days. Creep extension was measured directly using a traveling microscope reading to 0.01 mm on gage lengths marked on the specimens; the gage lengths

Jan 1, 1968

By Louis A. Panek

INTRODUCTION Block caving is distinguished from other types of mining in its requirements that the mineral body to be extracted must cave after a relatively large tabular open¬ing is created underneath it, and that the caving must be predictable and controllable for both operating and safety reasons. A mineral body that caves readily may present special problems with respect to ground support in the mining access openings, where stability, rather than failure, is the desired objective. The complete ex¬traction of a caving block, which may measure 50 to 100 m on a side and 100 to 300 m in height, tends to create a large void, resulting in significant changes in the ground stress system and a migration of rock from all directions toward the void; these events begin very early in the life of the mine and lead to additional spe¬cial problems with respect to control of ground move¬ments in the mine working areas, in the rock adjacent to the caving stope, and on the subsiding surface above the mine. Thus, the mechanics of an undercut-cave operation involve a broad range of geotechnical subject matter, including the determination of the lithostatic earth stresses, the structural properties of a large jointed rock mass, the flow properties of the fragments after caving, the stability of tunnel-like openings, and the stability and displacements of the rock structure (1 to 2 km in size) that surrounds the ore body. The objectives are achieved by in-place measurement of rock properties, laboratory tests, theoretical structural analyses of the related prob¬lems, measurements of structural behavior in and above the mine, and comparison of theoretical predictions with actual performance in order to interpret and refine the predictive capability. Undercut-cave mining (Julin, et al., 1973) differs from most other methods in that the caving is an essen¬tial and inherent part of the excavation process and in¬volves the movements of millions of cubic meters of rock over distances of hundreds of meters, as compared to the gradual 2-m settling of the overburden typical of full¬caving longwall coal mining. Undercut-cave mining may involve a rock mass on the scale of a 1-km diam open¬pit mine, but mining is several times deeper than that in the open-pit mine; furthermore, a large portion of the rock mass is in a failed, rather than a stable, state. Ground-control problems are inevitable in a situa¬tion where ground movements are occurring on such a large scale, and where the rock mass is expected con¬veniently to cave at one location but to be stable at an¬other. To devise effective responses to these ground¬control problems, one needs to know first of all what is happening, not merely at the point where the symptoms are observed, but more broadly with respect to the be¬havior of the entire mine rock-mass structure over the total area of extraction, including the effects on the sur¬face as well as those underground. Monitoring structural behavior on this scale would require a sizable group of people, if they attempted to measure all the significant mining-induced changes of ground displacement and pressure, even on a 100-m grid of points. The practical approach is to conduct the mea¬surements activity by a judicious sampling, making in¬termittent, concentrated efforts in selected areas of lim¬ited extent to study problems of immediate importance; simultaneously a minimum level of surveillance is main¬tained over the total structure to delineate long-term be¬havior trends in the major structure, which provide the baseline reference data that are so essential for interpre¬tive purposes. Mine-structure behavior depends not only on the mine extraction configuration and sequence, but on the rock structural properties and how they change from one location to another in the mine. Increasing attention is being given to determinations of rock strength and de¬formation by performing mechanical properties tests in place and in the laboratory, and by drawing inferences from the mapping of geologic structure in mine openings as they are driven and from the logging of cores ob¬tained from exploration drilling that may be performed primarily for delineating ore reserves. Monitoring of mine-structure behavior and inter¬preting it in terms of the growing excavations and the structural properties of the rocks have significant im¬pacts on designing the major undercut-cave operations of undercutting and drawing the ore, supporting the ac¬cess openings, and controlling surface subsidence. The purpose of this chapter is to call attention to the com¬mon ground-control problems encountered in undercut¬cave mining, to relate them to the mining process, and to explain the geomechanical approaches to analyzing these problems with emphasis on methods that are most readily available to the mine staff. UNDERCUTTING, CAVING, AND DRAWING A distinguishing characteristic of an undercut-caving method of extraction is the objective that the rock mass fails of its own weight in a predictable and controllable manner, after it is undercut by excavating a tabular¬shaped void, horizontal or inclined. Induced or forced caving refers to procedures that may be resorted to when the ore will not self-cave over the desired span. Block caving, panel caving, and mass caving are terms used to denote the particular sequence of undercutting and ex¬tracting the total area of the ore body on a single min¬ing level. The discussion herein is not specialized to any of these variations; the word block as used does not imply anything as to sequence. CREATING THE UNDERCUT A common undercutting method is to drive a series of drifts and blast the intervening pillars. Theoretical structural analysis and practical experience show that ground-support problems in the undercutting drifts and in underlying access openings as well (the same princi¬ple applies even if undercutting is done by belling out drawpoints from the slusher drifts) are minimized by retreat-mining across the undercut slot according to a sequence such that the growing undercut area retains a

Jan 1, 1982

By Bunting S. Crocker

PEBBLE grinding at Lake Shore is not a temporary wartime substitute. The tube milling plant, with a 1000 ton per day capacity, grinds a hard siliceous ore to 90 pct - 325 mesh. The plant, prior to using pebbles, was consuming 4.3 lb of 1 1/4-in. grinding balls per ton of ore, which amounted to 785 tons of balls per year. At September 1951 prices, $132.60 per ton, this steel cost amounted to, $104,400 per year, or $0.285 per ton milled. By the Lake Shore method of substituting screened rock for this steel, all of this cost is saved. This is one of the major economies in Lake Shore mill practice. Regardless of the ultimate price of grinding balls, a change, back to steel balls is not considered. The present pebble plant is more flexible than a steel ball plant and equally efficient. For example, if it is desired to change the size of grinding media, the pebble charge in any mill can be changed completely in 4 to 5 days, as against 76 days to change completely a 1 1/4-in. steel ball charge. To change pebble size it is necessary only to change two sets-of screens and clean out the rock feed storage bin. This will take 8 to 10 days as against 3 to 4 months to clean. out the customary supply of grinding balls kept on hand. Also it has not always been ' possible to purchase all desired sizes of balls at any price. In an analysis of savings effected by the use of the pebble mills, the flexibility of the Lake Shore grinding plant should be discussed, as it has a direct bearing on these savings. The plant has always used several units to handle tonnage rather than sending all the tonnage through a single unit. This principle may result in using small diameter mills, but no objection to that is seen. At no time has any advantage been found in cost per ton ground in large diameter mills. Both the capacity and the power of any mill varies as the diameter raised to the 2.6th power. Consequently large or small mills are equally efficient, and a plant should be designed to use as many units-or combination of units as is consistent with reasonable operating practice. Mills under 5 ft diam are harder to reline, etc. To define the case at Lake Shore, .2600 tons per day formerly were milled in seven 7x6-ft ball mills and twelve 5x16-ft (and two 6x16-ft) tube mills.* This gave an excellent test plant. and an extremely efficient-one. In this plant the ratio of ball mills to tube mills was 1: 2. When the much cheaper pebble mills were substituted for the tube mills, this. ratio -was changed to one ball mill to four pebble mills to take the greatest possible advantage of the cheaper operating mill, i.e., the pebble mill. This flexibility without loss in efficiency has been an important item in the cost savings. It is interesting to note that the use of pebbles for fine grinding was. proved first in the laboratory in a 12-in. ball mill. In fact, since 1934 all testing on -.8 mesh material has been done in this 12-in. mill. Scope of the Tests A paper on fine grinding at Lake Shore Mines was published in July 1940.1 This paper covered 7 years of intensive research on fine grinding as well as sizing methods and equipment, plant scale grinding tests on 5x16-ft tube mills with and without .grate discharges-both with 1 1/4-in. and 3/4-in. balls, the use of laboratory mills to evaluate plant changes, and several reports on classifiers and classification. In the following July the addendum report' was added in which the idea of series-circuit grinding was introduced, and the results of running five stages of tube mills and bowl classifiers were shown. Since 1940, the ball milling end of the plant has been altered extensively as a result of tests on the use of 3 1/2-in. rods in 7x6-ft mills and the use of the Tyler repulping screen with from 7 to 14 mesh screens. These tests are, lengthy and may be covered in a separate 'report later. The scope of this report is confined to ore ground by rod milling and ball milling until it passed through an 8 mesh Tyler Ty-rod screen. The -8 mesh screen undersize then was pumped to a primary bowl classifier in open circuit and the sands .from the bowl sent to the primary pebble mills, see Fig: 1. In the pebble mill circuit the ore is ground to 90 pct -325 mesh (24 pct + 28 microns) In studying the flowsheet, attention should be paid to the efficiency of the classification equipment used. The Tyler repulping screen is an efficient machine on the 8 to 10 mesh separation, and the bowl classifier is equally efficient at the 325 mesh separation. Efficient classification is a necessity for series-circuit stage grinding. The ore is hard siliceous porphyry, 60 pct SiO2, 80 pct insoluble. Its grindability at different. meshes has been shown near the top of 'the list in F. C. Bond's grindability tests' Lake Shore is not shown, but an adjacent-mine with identical ore, Wright-Hargreaves, is. Reasons for the-Changeover Since 1936 grinding balls have been rising steadily in price with no sign of stopping. For a mill that used 4.5 to 5.2 lb of grinding balls for every ton of ore ground this rise represented an alarming increase in grinding costs. In many cases the quality of the grinding balls fell off as the scrap steel became more difficult to obtain. The ratio of tube mills to ball mills increased with the use of the Tyler repulping screen in the ball mill circuit. Originally only 3.0 lb of balls per ton were used in the tube mills, and this

Jan 1, 1952

By Wilbur T. Stuart

TO assist the mining industry in attacking problems of water control, the U. S. Geological Survey has begun a program of research in mining hydrology. In certain fundamental respects water control is similar to development of water supplies from wells or to the drainage of agricultural lands, as many of the tools developed in recent years for quantitative ground-water problems are applicable, with modification, to mine-water problems. In 1952 a 30-day pumping test conducted jointly by the Eureka Corp. Ltd. and the Defense Minerals Exploration Agency provided an opportunity to gain knowledge concerning water movements around a flooded mine shaft. The methods of analyzing the data may be used as a guide for the evaluation of similar problems elsewhere. The Fad shaft of the Eureka Corp. is on Ruby Hill, 1 1/2 miles west of Eureka, Nev. The shaft was completed at a depth of 2465 ft in November 1947 at a site adjacent to the downfaulted block in which the ore was found. As the drift on the 2250 level progressed toward the ore zone, a large flow of water was encountered after the Martin fault was intersected. This flow exceeded the installed pump capacity, and an unsuccessful attempt to recover the shaft and the 2250 level was made in 1948.1, 2 Geology and Hydrology: The complex structure of Ruby Hill is that of an anticline broken first by thrust faulting and later by normal faults. The present orebody comprises several mineralized zones within a block of the Eldorado limestone of middle Cambrian age which was downfaulted 1400 to 1600 ft, and it may be related to a similar body mined at a higher level south of the Ruby Hill fault. At the depth of the largest zone of ore the block is roughly rectangular in shape, about 1000 ft wide and 1500 ft long, and dips about 30" NE, see Fig. 1. It is apparently bounded on the south by the Ruby Hill fault, on the east by the Jackson fault, on the north by the Martin fault, and the west by the Bowman fault. Within the block, but between the Ruby Hill and the Martin faults, are the Office and Adams Hills faults; west of the block and the Bowman fault are the Albion and Spring Valley faults. There are many conflicting reports concerning the water-yielding characteristics of the Eldorado limestone and the condition of the fault zones, that is, whether they are open or tight. However, the diamond-drill records indicate that open spaces as much as 2 or 3 ft across were encountered, and considerable cementing and lining of holes was necessary to maintain circulation of drilling fluid. There is also evidence that the Eldorado limestone was cavernous where it was mined in the early days south of the Ruby Hill fault. At the site of the Fad shaft the formations encountered from the surface down included the Pogonip limestone, Dunderberg shale, Hamburg limestone, and Secret Canyon shale. These formations did not yield large quantities of water to the shaft. The Pogonip limestone, which appears to be permeable and might yield water elsewhere, is above the water table in the vicinity of the shaft. The Secret Canyon shale, immediately overlying the Eldorado in some places but in most places separated from it by the Geddes limestone,3 is apparently tight and does not transmit water. During the 30-day test period the shale briefly confined the water in the underlying formations so that artesian conditions were observed in drillholes E and F, which are cased into the Eldorado limestone, whereas unconfined conditions were observed in drillholes B, C, and D, which were open to the shale. The Geddes limestone, which normally lies between the Secret Canyon shale and the Eldorado limestone, was not encountered in the Fad shaft. The Geddes, a flaggy, fractured limestone, is reported capable of yielding large volumes of water. Water stored in the interstices of this thin-bedded limestone within the Ruby Hill fault zone on the 1200 level of the Locan shaft drowned the pumps in 1923 when the Richmond-Eureka Mining Co. attempted to explore the area along the fault. Eldorado limestone was not encountered in the Fad shaft. In the ore-block area the Eldorado limestone was not entirely offset from other water-yielding formations by movement along the Bowman fault; therefore it may be hydraulically connected with the other formations. Adjacent to the Ruby Hill, Jackson, and Martin faults, the Eldorado lies in contact with other possible water-yielding formations. One of these, the Prospect Mountain quartzite, is separated from the Eldorado by thin, sheared, and broken beds of the Geddes within the Ruby Hill fault zone. A limited examination by the author of the Prospect Mountain quartzite in the Richmond mine at a higher level and south of the Ruby Hill fault indicates that the quartzite is poorly permeable. The monzonite mass south of the quartzite would be a further barrier to the flow of water. The poor permeability of this area is substantiated by records of levels at which water was encountered south of the Ruby Hill fault. In view of the normally low rate of ground-water recharge, if this desert area had been permeable, water levels could not have been maintained at altitudes of many hundred feet above the present water table west and north of Ruby Hill. Thus the ore-bearing block of Eldorado limestone is in contact with possible water-yielding rocks on at least two sides, and if the fault zones are possible conduits for water circulation the geologic and hydrologic conditions are suitable for the infinite-aquifer type of analysis as used and modified here. History of Pumping: During sinking of the Fad shaft a maximum pumping rate of 1500 gpm kept the shaft dewatered sufficiently, but in March 1948, after the 2250 level drift passed through the Martin fault into the Eldorado limestone, the pumps and shaft were flooded. Subsequently additional pump-

Jan 1, 1956

By C. P. Verschuren, P. W. Kingston, W. F. Caley

"In recent years, there has been rapid growth in the use of natural stone building and construct ion products. The largest growing markets are for pre-fabricated natural stone panels for interior and exterior wall facing and flooring. This increased demand is attributed to advanced quarrying and processing technologies, renewed appreciation for the aesthetic appeal of natural stone, and to the superior durability of natural stone when compared to glass, steel, and concrete. Geologically, plutonic rocks, and to a lesser extent marble and limestone/dolostone, are used extensively for these applications. This paper details the economic structure of the dimension stone industry, describes the necessary characteristics of a stone quarry, and outlines the geological criteria which favour an economic quarry.INTRODUCTIONThe world building stone industry has remained relatively stable over the last 5 to 6 years. In contrast, the North American stone industry has experienced substantial growth during that period. The increased use of commercial granite is primarily responsible for this growth, where use as exterior wall veneer for commercial, industrial, and institutional buildings represents the largest and fastest growing market.This increased usage in the construction industry has been sparked by the application of new technologies in quarrying, processing and installation , ma king natural stone more competitive with man-made materials which are energy intensive to produce, and have risen dramatically in price. The development of safer, time-saving stone installation techniques has also appealed to the liability-conscience construction industry. Finally, there is a trend by designers and architects toward utilizing long-la sting weather-resistant building materials, and a renewed appreciation of the aesthetic appeal of natural stone. The other large user of granite is the monument industry of North America which, although it has decreased slightly in recent years, still consumes approximately 40% of all commercial granite produced. The local use of commercial marble has remained relatively stable, and North America still relies heavily on European sources and suppliers. There has been some concern over the use of marble for exterior applications in areas subject to acid rain, and in colder climates where the marble ma y be vulnerable to freeze thaw cycles. Despite these concerns, the greatest problem is not with the structural breakdown of marble but the inability of some marbles to retain their polished finishes in the se environments. As a result, there has been an increase in the use of granite at the expense of marble for exterior applications. Finally, commercial marble is used extensively for interior flooring, wall s, counter tops, furniture, sculptures, and stairway components"

Jan 1, 1989

By Jeremy Spearman, Alan Cooper, Mark Lee, Jon Taylor, Michael Turnbull, Neil Crossouard, Bramley Murton

"SUMMARYSeamounts are underwater mountains caused by volcanic activity. They are of great oceanographic interest because of their local and basin-scale influence over ocean systems, their often unique ecosystems, and because they are associated with the formation of ferromanganese (Fe-Mn) crusts. These crusts are of particular interest to scientists because they are rich in some of the rare elements increasingly required for development of renewable technologies. When the possibility of mining such seamounts is considered, an inter-disciplinary study of seamount hydrodynamics, crust formation and ecosystem sensitivity is required. MarineE-tech is one such multi-disciplined project that has been commissioned to address these different aspects in order to improve understanding of the viability and desirability of mining Fe-Mn crusts. This paper describes a key part of this study, namely the development of a model, informed by hydrographic observations, of the currents around one such seamount and the in situ validation of a sediment disturbance plume model as a pre-cursor to assessment of the potential effects of mining activity.INTRODUCTIONFerromanganese crusts are thought to form by precipitation of iron and manganese oxides from oxygen depleted mid-water known as the oxygen minimum zone (OMZ). Oxidation rich waters, typically at depths of about 800-2200 m (Hein et al, 2000). This range of depths coincides with the flanks and summits of many seamounts. The ferromanganese crusts found on seamounts have been found to have high concentrations of rare elements, like Tellurium, that are considered critical to low-carbon energy production. Seamounts are complex environments: in terms of the local hydrodynamics that tend to trap water over the seamount due to the Taylor Cap effect (e.g. Chapman and Haidvogel, 1992) and which can enhance the productivity of seamount waters through upwelling (Lavelle and Mohn, 2010), in terms of the microbiological activity that leads to the precipitation of the Fe-Mn crusts, and also in terms of the often unique ecosystems that are the product of the “island” nature of seamounts and the upwelling of nutrients. The understanding of this complexity is a necessary journey towards the realization of viable, yet environmentally responsible, mining schemes. The formation of crusts will be influenced by the local (and larger-scale) hydrodynamics. The exploration of promising crust sites could be made much more efficient by understanding where the most prospective deposits are likely to occur. Knowing how to remove these crusts without impacting significantly on the local environment will be a pre-requisite for such mining to be acceptable to the public and funders alike."

Jan 1, 2017

By W. Crafts, J. L. Lamont

Secondary hardening in tempering has long been recognized as a typical characteristic of steels containing large amounts of carbide-forming alloys. These steels, when quenched and tempered, tend to soften somewhat after tempering at low temperatures, and to reharden at intermediate tempering temperatures before finally softening to low hardness. This behavior in tempering has been studied by many investigators, chiefly with respect to thc secondary hardening of high-speed steels. The initial softening has usually been ascribed to decomposition of martensite and growth of iron carbide particles. The secondary hardening has been explained by formation of fresh martensite from residual austenite, formation of very fine alloy carbide particles and by precipitation hardening of metallic compounds in the ferritic matrix. However, factors derived for the calculation of tempered hardness in low-alloy steels1 indicated only a tendency toward retarded softening rather than a discontinuous rehardening. A study was, therefore, made to determine whether fresh martensite from residual austenite and precipitation hardening are essential to rehardening and the degree to which tempered martensite could be rehardened by precipitation of alloy carbides. The mechanism of carbide rehardening was also investigated to determine the nature of the process. The causes of secondary hardening in high-speed steel have received serious consideration for about thirty years. Among the earlier workers Bain and Jeffries2 in 1923 recognized residual austenite as a significant factor, but in addition, emphasized the effect of the formation of alloy carbide particles of "critical" size. They conceived that freshly quenched martensite forms low-alloy iron carbide at low tempering temperatures; with further tempering the iron carbide is then modified by alloying elements that can diffuse most readily at intermediate temperatures (850°F); and that the iron carbides grow in size by coalescence so that the steel tends to lose hardness. At higher tempering temperatures carbide stability rather than alloy availability becomes the controlling influence and in high-speed steel "iron-tungsten carbide is the most stable one and forms to the elimination of the earlier formed carbides. . . . The iron-tungsten carbide particles reach approximately the size for critical dispersion after a short reheat at II00 F." This concept, the formation of alloy carbide nuclei, redissolving of the iron carbide, and diffusion of carbon to the alloy carbide, was restated for vanadium steel in 1932 by Houdremont, Bennek, and Schrader. Subsequent studies have confirmed that iron carbide changes to alloy carbide at 1000 to IIOOOF, but the mechanism of the process has received little consideration.

Jan 1, 1949

By R. H. Richman, F. Borik

A 0.3 pct C-12 pct Cr-6 pct Ni steel was rolled to 93 pct reduclion in area as austenite at 510°C, and then partially transformed as desired to ~rlartensite by qnenching to - 196°C. Pole figures for the austenitic matrix and for the martensitic product were separately determined by an X-ray transmission method. The deforitration texture of' the warm-worked austenite is characlerized by (110)(225) components, and is thus closely similar to those produced in a brasses. The pole jigure of the martensite in partially transformed material agrees well with that which can be constructed by transfortnation of the {110)(225) orientations according to either the Kuvdjuniov- Sacks or the Nishi-yatuu relatiotship. Howeuer, an important result of this construction is that me-third of the predicted orientations are missing. A graphical analysis can then be used to show that in deformed austenite certain crystallographic variants of martensite (related to the most probable austenite slip systems) are suppressed, resulting in this preferred transformation. The evidence for preferred transformation is corroborated by the measured elastic anisotropy of warm-rolled and fully transformed H-11 steel. EXTENSIVE plastic deformation of a polycrystal-line aggregate in a manner that causes flow predominantly in one direction results in a preferred orientation of the constituent crystallites. The particular orientations that are produced depend upon the crystal structure and composition of the material, as well as upon the temperature, mode, and degree of deformation; in any case, the preferred crystallo-graphic orientations, or textures, are reflected in directionality of mechanical properties. Although such anisotropy may be exploited in certain specialized applications, it is more commonly diminished or eliminated by heat treatment lest it interfere undesirably in subsequent forming operations or in structural design. In the recently developed thermomechanical treatments that significantly enhance the strength of some steels,1,2 considerable deformation of the metastable austenite prior to the martensite transformation is essential to the strengthening process. If the austenite is textured by the deformation, and if the transformation to martensite proceeds according to one of the relationships established for transformation in annealed austenite, then it must be expected that the martensite will also possess a preferred orientation even though the multiplicity of martensite orientations possible in a given austen- ite crystal will tend to restore some degree of randomness. The existence of a residual anisotropy, both mechanical 3-6 and crystallographic,' has been substantiated. In the latter crystallographic investigation, preferred orientations were determined for the martensitic structure of an SAE 4340 steel rolled 72 pct as austenite at 833°C and then quenched. However, the choice of a composition that transformed almost completely to martensite during the quench to room temperature did not permit direct measurement of the prior austenitic texture. In fact, when the "ideal orientations'' associated with well-known fcc rolling textures were converted, alone or in combination, to martensite according to the Kur-djumov-Sachs (K-s)' or Nishiyama8 relations, the agreement obtained with the observed martensite texture was only fair at best. Recently a pertinent aspect of the austenite to martensite transformation was reported by Bokros and parker,10 who found that certain habit-plane variants of martensite were suppressed by tensile deformation of Fe-31.7 Ni single crystals prior to the necessary subzero cooling. It might be anticipated that the consequences of such preferred transformation are sustained during the formation of martensite in warm-worked austenite that has a well-developed deformation texture. The present investigation was undertaken first to establish more firmly the relation between preferred orientations in plastically deformed austenite and in the resulting martensite, and second to examine the textures for evidence of deformation-induced preferred transformation. EXPERIMENTAL PROCEDURES An alloy containing 0.3 pct C, 12 pct Cr, 6 pct Ni, and the balance iron, was selected because the mar-tensite-start temperature (M,) of about -100°C allowed convenient experimental manipulation of either austenite or martensite at room temperature. Furthermore, this composition can be readily deformed as metastable austenite at moderately elevated temperatures without intervention of appreciable isothermal or athermal decomposition products. The alloy was austenitized at 1150°C, aircooled to 510°C, rolled unidirectionally at this temperature to 93 pct reduction of cross-sectional area, and finally oil-quenched to room temperature. Partial transformation to martensite was accomplished by quenching to -196°C as needed. The rolled stock was reduced in thickness from 0.067 to 0.010 in. by etching in a solution of 5 pct HC1, 45 pct HNO3, and 50 pct water, and further thinned by careful mechanical polishing to maintain the two sides of the sheet parallel within 0.0003 in. After mechanical polishing to 0.005 in., electropolishing in 1:9 perchloric-acetic acid solution produced a final thickness of 0.002 in. The preferred orientations were determined from

Jan 1, 1968

By W. F. Campbell

One of the deepest coal operations today is the Springhill No. 2 mine of Cumberland Railway & Coal Co., subsidiary of Dominion Coal Co. Ltd. Mining is now conducted at a slope distance of 14,000 ft, with 4400 vertical ft of cover. The record of Springhill No. 2 can be said to contain the history of bumps in the Province of Nova Scotia. The Springhill coal field forms part of the Cumberland field of Carboniferous age. There are seven mineable fields in the area, numbered in the order they were discovered. Mining has been carried on in all but the No. 4 and 5 seams, but present operations are confined to the No. 2. Fig. 1 shows a vertical section through the seams. Opened in 1873, No. 2 mine was first worked from parallel slopes driven from the outcrop of No. 2 seam down to the 7700 level. As the mine went deeper, a two-place auxiliary slope was driven from the 6900 level to the present workings. A transfer level at the 7800 connects the main haulage with the auxiliary haulage slope. The No. 2 mine plan is shown in Figs. 2 and 3. The seam is bituminous, averaging 8.5 to 9 ft thick. There is a well defined parting 14 to 16 in. from the roof, and this roof coal is harder than the rest of the seam. Average pitch at the outcrop is 30°; at the 6500 level, 20°; at the 7900 level, 16°; and at the 13,800 level, 10°. Immediate roof and floor strata consist of beds of variable thicknesses of shales, grading to arenaceous shales to shaly sandstones to sandstones. A characteristic of the strata is the appearance and disappearance of sandstone bands, of considerable thickness, over distances of several hundred feet. MINE HISTORY Entrance to No. 2 mine is obtained by three parallel slopes, separated by 100-ft pillars except in the upper portion of the mine where the pillars are smaller. Main haulage levels were originally driven 600 ft apart for room and pillar extraction. A parallel drainage and intake airway was driven below each haulage level and a parallel return airway (counter level) above. Pillars 80 to 100 ft wide separated these two servicing entries from the center haulage level. Inclines up to 700 ft apart were driven off the levels up pitch to the upper levels, and at 40-ft centers level rooms 12 ft wide were driven off each side of the inclines, separated by crosscuts every 50 to 100 ft. The rooms were driven up to 350 ft or until they were holed into the room from the adjoining incline—a plan followed until the level reached its boundary. When the entire area between two levels had been divided into pillars, the pillars were extracted from the boundary back to the slope pillars. As depth of cover increased, mining conditions required larger pillars, and at the 3300 level, where NECESSARY PRECAUTIONS FOR MINING AT DEPTH In Room and Pillar Extraction: 1) Pillars must not be too small. 2) Pillar size must be uniform. 3) Extraction lines must be kept as straight and as uniform as possible. If they are irregular they should be brought back into line slowly. 4) Peninsulas of coal surrounded by gob must be avoided. 5) Pillars should not be disturbed after they have been formed. If it is necessary to increase the width of an opening, stripping or ribbing should be done gradually over a long distance. If increased width of openings is necessary at certain locations and can be anticipated, the openings should be widened during development. 6) Under no conditions should small pillars be left in the gob that would interfere with caving. In Long wall Operations: 1) Pillars formed during development for retreating longwalls should be uniform in size and shape. 2) Pillars should not be split after development, but if pillar splitting does become necessary it should not be carried out within the zone influenced by the working faces. 3) Width of levels should not be increased after the levels have been developed. If it does become necessary to widen a level it must be done gradually over a long distance by stripping or ribbing the coal ribs. 4) Extraction lines should remain uniform. If several walls are worked together—one behind the other—a uniform distance must be maintained between them and the wall faces must be kept as straight as possible. If the walls do get out of line, they must be brought back into line gradually to allow for a gradual change in the stress distribution over the area. 5) Stumps of coal, roof supports, etc. must not be allowed to remain in the gob to interfere with caving. 6) Midwalls should be built as rigidly as possible and properly maintained.

Jan 1, 1959

By Janet Briggs, Michael Vertes, J. F. Shirley, R. E. Cutherbertson, Alexander Sutulov, J. D. Vincent

Molybdenum, the 42nd element in the periodic table, is now a very important metal in industry of the United States and the world. Chemically, it is a very complex element having valences of 0, +2, +3, +4, +5, and +6, and can exist as a mixture of these valences in many compounds. It was discovered and separated by Scheele in 1778. At the present time, molybdenite is the only important commercial source of molybdenum. Molybdenum occurs in some 12-14 different recognized minerals, but practically all of the production is molybde¬nite. As late as 1915, it was considered a rare metal. Germany made extensive use of it during the first world war, which resulted in the establishment of a molybdenum industry in the United States at Cli¬max, Colo., in 1917. It was not until the mid-1920s when its introduc¬tion to steelmaking developed for the US automobile industry that the metal gained important acceptance. The first notable example was the Willis-Saint Claire automobile introduced in 1921 by C. Harold Willis, formerly chief metallurgist for Ford Motor Co. Molybdenum has also been produced in the US from wulfenite (PbMoO4) from Arizona and Nevada, and a small amount of molyb¬denite was produced up to the early 1970s as a coproduct from two molybdenum-bismuth mines in Quebec, Canada. The first byproduct production of molybdenite from a copper con¬centrating operation was achieved at Cananea, Sonora, Mexico, and was followed shortly by the Utah division of Kennecott Copper Co. in the mid-1930s. This section of the Handbook is divided into four main subdivi¬sions: Chapter 2, which deals with the concentration and recovery of molybdenite from primary molybdenum ores; Chapter 3, which deals with the recovery and separation of molybdenite as a byproduct from copper production; Chapter 4, which deals with the conversion practices employed for converting molybdenite to the oxide or metal form; and Chapter 5, which covers the uses of molybdenum in steel¬making, chemicals, pigments, lubrication, catalysts, etc. The Free World molybdenum production, for the past eight years from all sources, is shown in Table 1. 2. Primary Molybdenite Ores General In 1978 there were four primary molybdenite recovery plants in operation: Climax, Endako, Quests, and Henderson. For over half a century the Climax mine has been by far the largest and most important source of molybdenum in the world. During the early 1970s a few primary molybdenite mines in British Columbia were shut down due to adverse economical conditions. The largest of these was British Columbia Molybdenum, formerly owned by Kennecott Copper Corp. The Climax plant and operations are described in detail and the opera¬tions of the other plants are summarized following a general discussion of recovering molybdenite from primarily molybdenite ores. Recovery of molybdenite as a byproduct will be taken up in Chapter 3. Geology Molybdenite production from mines in which it is the sole or preponderantly important mineral is confined to ore bodies in granites or igneous-type rocks or pegmatite dikes. More than one state of mineralization is sometimes present as at Climax, York, Hardy, and Questa. In the case of Climax, important byproduct recovery of tung¬sten, tin, pyrite, and monazite is also attained. The molybdenite is commonly disseminated very finely in the siliceous granitic gangue but does occur in stringers and blebs in silicified mineral solution channels. Treatment of Primary Molybdenum Ores All primary molybdenum ores are low in molybdenite, the content ranging from less than 0.20% MoS2 to a high of about 0.40% MoS2. A high resistance to grinding is a common ore characteristic combined with a habit of very fine dissemination of molybdenite in the siliceous portion of the ore. Because of low MoS2 content, high hardness, and fine dissemination of molybdenite in primary ores, it has been necessary to develop a special mill treatment which yields an accepta¬ble recovery at a reasonably low milling cost. Essentially, the treatment consists of intensive rougher flotation at a relatively coarse grind with production of a low grade rougher froth containing a substantial percentage of the molybdenite in the form of low grade middling particles, followed with multistage regrinding to liberate the locked molybdenite and several stages of flotation cleaning to produce a finished grade concentrate. Rougher flotation on pulp ground to ap¬proximately 35 mesh is common practice, even if complete liberation of molybdenite from siliceous gangue may require regrinding to 200 mesh. The flotation of middling particles with a very low content of molybdenite in the rougher flotation step has been made possible by the application of a combination of reagents first developed at the Climax mill and used with some modification in the treatment of all primary molybdenum ores. The reagents commonly employed and their functions in flotation are:

Jan 1, 1985

By P. G. Shewmon, H. E. Collins

A study has been made to determine the sites at which sulfur adsorption occurs on copper surfaces. measurements were made of the relative torques, Ys, at the intersection of twin boundaries with surfaces near the three low-index orientations, i.e., (100), @lo), and 011), over a range of H2S/H2 ratios. HZS concerztvations j'ro~n 3 to 1500pp)n between 830" and 1050°C were used. It is concluded that sulfur adsorption occurved preferentially though not exclusively at edge sites near the (100) and (110) surfaces in the HzS range — 700 Ppm giving rise to negative torques near these orientations. Beyond this HzS range, adsorption occurred at all sites. Near the (111) surface, 7/y little with HzS concentration up to approxiwzately 75pptn. Above this range, the results indicate adsorption is occurring OH both terrace and edge sites. SCIENTIFIC interest in surfaces and their interactions with a gaseous environment dates back to the beginning of the 19th century. The scientific luminaries of that period—Faraday, Maxwell, Rayleigh, Dewar, and Gibbs—were already concerned about such processes. However, it has only been within the past several decades that adsorption on metal surfaces has been actively studied. This increased interest in adsorption has been brought about by the advent of new and improved experimental techniques and apparatus, e.g., ultrahigh vacuum, and field-emission and ion microscopes. However, most of the work done using these techniques has been carried out at low temperatures. When adsorption studies have been made at either low or high temperatures, they usually gave no indication of the particular surface orientations or type of sites on which adsorption was occurring. In the last few years, there have been a series of studies in which the surface tension, y,, and/or its derivative with respect to orientation, 7, have been studied as a function of orientation and atmosphere.'-7 Nearly all of the work on the relative torque,* ~/y, silver annealed in hydrogen and air.6 Recently Winterbottom and Gjostein" have used a modified and more accurate Mykurian method to determine the y plot of gold in hydrogen The only work in which T/~, has been measured over a range of chemical potentials for a given solute, p2, is that of Robertson and shewmon7 on the Cu-0 system. They measured T/Y, vs Po, (10"" to 10- l3 atm) at 1000°C in various mixtures of Hz0 and HZ. From this work they estimated the value of p2 at which one half of the surface sites are occupied with oxygen, pg, as being in the range 10- l6 to 10- l5 atm of oxygen. They also found that increasing Pa increased the magnitude of ~/y, near the (111) and (100) orientations. This indicates that oxygen is not adsorbed preferentially at step edges, but uniformly over all surface sites. In addition, they did one experiment on sulfur adsorption on copper surfaces, which indicated that sulfur adsorption decreases ~/y, near the (100) orientation, while not affecting ~/y, near the (111). This could be interpreted as indicating that sulfur adsorbs preferentially at step edges near the (100). In this paper the primary objective of the work has been to carry out a study of sulfur adsorption on copper surfaces over a range of temperatures and p,. In conjunction with this work, thermal grooving at grain boundaries has been examined as a method of determining the effect of sulfur adsorption on y,. METHODS Ideally, one would like to have information on the quantity of solute adsorbed on a surface and the types of sites at which it is absorbed as a function of p2. The total quantity adsorbed or the surface excess is given by the thermodynamic equation Thus data on the variation of y, with pz indicates the value of p2 at which adsorption becomes appreciable and the quantity adsorbed. The type of adsorption site is more difficult to deduce but information on this can be obtained from the variation of rz with 8, the angular deviation of the surface orientation. This is obtained from the thermodynamic equationlg Data on t and ys as functions of p2 have been obtained by the following methods. 1) Twin Boundary Grooving—By determining the effect of adsorption on the torque, 7, where T is the variation of surface energy, y,, with orientation, it is possible to obtain some indication as to the preferred sites of adsorption. Experimentally, the torque value measured is the relative torque, 7/ys The twin boundary grooving technique suggested by Mykura'' was used in this study to determine near the three low-index orientations— (loo), (110), and (111). Mykura's equation relates 7 /yS to measurements of the di-

Jan 1, 1967

By E. Amott

A test is described in which the wellubility of porous rock is measured as a function of the displacement properties of the rock-water-oil system. Four displacemet operations are carried out: (I) sponlaneous displaceti?ent of water by oil, (2) forced displacement of water by oil oil in the same system using a centrifuging procetllrre, (3) spontaneous displacement of oil by water. and (4) forced displacment of oil by water. Ratios of the spontaneous displacement volumes to the total displucenlent volumes are used as wettability indicates. Cores having clean mineral surfaces (strongly preferentially water-wet) show displacement-by-waler ratios approaching 1.00 and displacement-by-oil ratios of zero. Cores which ([re. strongly preferentinlly oil-wet give the reverse resu1ts. Neutral wellability cores show zero values for both ratios Fresh cores from different oil reservoirs have shown wettabiltties in tlris te.st covering rrlti~ost 111e conlplrtt, range of thr: te.st. Notvever, nlo.s/ of tlle fresh California cores tested were slightly prcfere111icrlly wclter-\vet. The chrrnge.~ in coro u ('liabilities, as indicated hy this te.st, r~.sril/ing from various CO~P hanrlling procedures tt,ere oh.served. In sonie ca.sc,s /Ire ~,cttahilitio.c. of fresh cores were changell by drxi:~g or 11y e.x/rclct ing with iolcreiii~ or. dioxunc~; in o/h~r cases they were 1101 changed. Co~ltrrc/ of cort,.s ~.ith filtrc~t~c. from water-base rlrilling rilrrrls crlrc.sed littlc change in we/ /ahility ivhile contnct with filtrates frorii oil-hus~ ri1rlcl.s tlecrrascrl the prefcrerlcc, of the, cores for )I.NI Usitig thi,s test to ri.crl~lute n~r~ttubili~y, N .vt~ldy was iilarle of /lie correlmtio~i of wettability with wa/erfloocl nil recovery for orttcrop Ohio sand.stone and for Al~ln-tlunl. Resul/.v indicate thml no single correlatioti between these factors applies to different porous rock syste~n. It is thought that diflerences in pore gen~netry resrrlt in diflrrerrce.~ in this correlurio~z. INTRODUCTTON Most investigators who have reported on the wettahility of porous rock have described such rock as prcferentially water-wet or preferentially oil-wet. In some cases a third classification, neutral wettability. has been used. The efficiency of water floods in each of these wettability groups has been described in numerous publications. Several methods for characterizing porous rock wet tability more precisely have been reported,' " but it appears that because of one weakness or another. none of these has been generally accepted. Early in our studies in this field, it was found that the displacement efficiency of oil by water in a particular porous rock having a strong preference for water was quite different from that in a similar rock having only a moderate preference for water. Thus, there appeared to be a need for a practical, reasonably precise wet tability test. one which could classify porous rocks into 10 to 20 different groups rather than the two or three broad groups listed above. The test developed to meet this need is described in this paper. Also, changes in wettability, as indicated hy this test, resulting from various core handling procedures are discussed. Finally, data showing the corrclation of wettability with waterflood oil recovery for two different types of cores are presented and discussed. Some confusion has resulted from the failure of certain writers to define clearly some of the wettability terms they have used. Accordingly, the following commcnts concerning definitions are offered. The wc t ta-hility of a solid surface is the relative preference of that surface to be covered by one of the fluids under consideration. It is felt that this is the generally accepted definition. The fluids being considered must bc specified (or understood) before the term wettability has any significance. In the work reported here these fluids are water (3 per cent brine) and oil (kerosene). The term preferential wettability is sometimes used, but we think that the word preferential is redundant here and should not be used. Tn line with the definitions of Jennings', a preferentially oil-wet solid surface is regarded as a surface which will show an oil advancing contact angle less than 90" (measured through the oil) in the water-oil-solid system. Oil will spontaneously displace water, if both are at the same pressure, from such a surface. A preferentially water-wet surface is analogous. This is consistent with the wettability definition above. As Jennings has said, frequently the term oil-wet is used to mean the same thing as preferentially oil-wet. However, oil-wet also has been used occasionally referring to an oil-covered surface when the availability of water was limited. To avoid confusion from this source, we do not use the terms oil-wet and water-wet. DESCRIPTION OF WETTABTLITY TEST The following points were considered desirable in a wettability test for our purpose. 1. The test should be a displacement test resembling

By I. R. Kramer, L. J. Demer

Aluminum single crystals were pulled in an electrolytic cell allowing surface removal during the deformation. The extent of Stages I and 11 of the stress-st-aitz curve was increased and the slope decreased as the rate of metal removed from the surface was increased. An increase of the strain rate caused a decrease in the effectiveness of the metal removal. The data indicate that the work-hardening coefficient in Stage I is determined primarily by the conditions which exist on the surface of the crystal. In Stages 11 and 111, both surface effects and internal barriers are important. ALTHOUGH numerous investigations have been conducted on the plastic flow characteristics of metals in an attempt to explain the mechanism of work-hardening, relatively few studies have taken into account the influence of the surface. In all current theories of work-hardening it is assumed that the impediments to the movement of dislocations are within the crystal. The barriers due to the surface and the existence of solid and liquid films have been neglected even though it has been demonstrated that the surface exerts a large effect. A number of investigators1-l8 have shown that solid films on the surface of single crystals markedly affect their mechanical behavior. In general, the presence of a solid film tends to increase the yield stress and increase the work-hardening rate. Often, on single crystals, Stage I and, at times, Stage II regions are completely suppressed. Various mechanisms have been offered for the effects of oxide and metal films as well as the influence of electrolytes. Of these, concepts concerned with the locking of surface dislocation sources and the blocking of dislocations at the surface resulting in pileups appear to be actively considered at present. Barrett,11 Takamura,6 Gilman,19 Lipsett and King,20 Shapiro and Read,10 and Weiner and Gensamer21 are among those who have interpreted their results in terms of piled-up dislocations at the surface, while Adams.22 and Chalmers and Davis23 have explained their experimental observations in terms of locking of surface dislocation sources. In general, the change in plastic flow properties due to electrolytes has been explained in terms of the unblocking or unlocking of dislocations by the removal of the oxide films. In considering the two proposed mechanisms, it appears that the locking of sources of surface dislocations by a solid film should exert a primary influence only on the critical resolved shear stress for flow and not on the slopes of Stages I and 11. However, the blocking at the surface of dislocations from internal sources may also affect the critical resolved stress and furthermore exert an influence throughout the whole plastic range. In certain cases it does not seem feasible to explain the results of experimental observations in terms of locking of surface dislocation sources. The abnormal aftereffects found by Barrettll,12 by removing the oxide by an acid treatment are excellent evidence of the blocking of dislocations at the surface. Additional evidence in favor of a blocking due to a pileup of dislocations at the surface may be found from the observations that the critical resolved shear strength continues to increase with the thickness of the oxide layer until very heavy oxide layers are formed. If the locking of surface dislocations sources were the dominant factor, the critical resolved shear stress would not be expected to increase after all of the surface sources were locked by the formation of the oxide. This may be expected to happen after a few atomic layers of the oxide are formed. In spite of the above evidence on the strong influence of the surface on the plastic flow characteristic, this has been ignored in current theories of work-hardening. Seeger24,25 suggested that most of the dislocations may slip out of the crystal only when the specimen axis is within certain areas of the orientation triangle. In other areas the resolved shear stress in other glide systems is large enough to generate dislocations which can form Lomer-Cottrell locks, thereby decreasing the average slip distance in some directions and causing a larger hardening rate. Friede126 assumed that at the beginning of Stage 11, a large number of Lomer-Cottrell dislocations are formed by a catastrophic process which used up all the Frank-Read sources on the secondary slip-planes. In this manner a fixed number of Lomer-Cottrell locks is formed which act as barriers against which the dislocations can pile up. In Stage 111, Seeger24 and Diehl, Mader, and Seeger25 proposed that Lomer-Cottrell barriers are circumvented by the cross slip of extended screw dislocations. Cottrell and Stokes,27 Friedel,26 Cottrell,26 and stroh29 suggest that the Lomer-Cottrell dislocations collapse under the stress field of the dislocation pileup. It is the purpose of this paper to report the changes in Stages I, II,and III of the deformation process in

Jan 1, 1962

By A. E. Jenkins, L. A. Baker, N. A. Warner

Rates oj decarburization of levilated Fe-C droplets conlaining 5.5 to 0 pct C have been measured at 1660°C. Gas mixtures of 1, 10, and 100 pct 0, with helium diluenl were used at velocities of 12.5 and 62.5 cm per sec. Rates were independent of carbon concentration in the mell and in good agreement with the calculated rule of oxygen diffusion through the gas boundary layer. The effects of flow rale and total pressure are as predicled and the rates are approxitnalely 2.5 times those with CO2 as oxidant. The mass-transfer correlation used incorporaled the efject of natural convection as well as forced conrection. Graphile spheres are shown to oxidize at the same rate as Fe-C droplets under the same experimental codlions. It is concluded that, for high carbon concentrations in the melt, the rate of- decarburizalion is controlled wholly by the rate of gaseous diffusion. Rate measurements with pure CO, are reported for low carbon concentrations where CO bubbles nucleate within the droplet. Under these circumstances the decarburi-zation decreased with carbon concentration and it is proposed that carbon diffusion is significant in conlrolling the decnvburization rate. In an earlier paper1 decarburization rate measurements were reported for levitated Fe-C alloys at 1660°C but with CO2 as the oxidant. The decarburization rate was found to be independent of carbon concentration in the melt but slightly affected by total pressure. The authors were unable to explain the slight pressure effect but in all other respects the results were consistent with control by diffusion in the gas boundary layer. Subsequent work has been directed at finding the reason for the slight pressure effect and whether the kinetics with oxygen as oxidant parallel those with CO2. Recently Ito and Sano2 have shown that with water vapor-argon atmospheres the decarburization rate is gaseous diffusion controlled until an oxide film appears on the surface. In this work the melts were contained in crucibles. MASS TRANSFER IN THE GAS PHASE In the earlier analysis1 only forced-convection mass transfer was considered. Subsequent recognition of the existence of some free-convection mass transfer explained the observed small effect of total pressure on the decarburization rate. Steinberger and Treybal3 and Kinard, Manning, and Manning4 have developed correlations involving the linear addition of the contribution of radial diffusion, free and forced convection. Steinberger and Treybal&apos;s correlation was chosen as the most applicable to the present work since it correlated most of the data available in the literature and handled the low Reynolds number region exceptionally well. The correlation for (Gr&apos;Sc) < 108 is where Nu&apos; is the Nusselt number for mass transfer based upon the total surface of a sphere in an infinite medium, G&apos; is the mean Grashof number for mass transfer defined by Eq. [2], Sc is the Schmidt number (µ/pDAB)f, Re is the sphere Reynolds number (dpu,pf/µf), p is the viscosity of the gas (poise), p is the density of the gas (g cm-3), Dab is the binary diffusivity for the system A-B (sq cm sec-&apos;), dp is the sphere diameter (cm), u is the approach velocity of the gas (cm sec-I), and subscript f denotes the property value is computed at the film temperature Tf defined by Tf = +1/2(To + Tr) where To is the specimen temperature and T, is the approach gas temperature (oK). Natural convection occurs when inhomogeneities exist in gas density. These may be caused by concentration gradients, temperature gradients, or both. In the present work the temperature gradient between the sphere and the bulk gas was very large and in some cases, for example the runs with pure oxygen, the concentration gradient was also appreciable. The Grashof number defined by Mathers, Madden, and piret5 was used since it took account of both temperature and concentration gradients: where Gr&apos; is the Grashof number for mass transfer (p2fgd3|-yA-yA|/µ2f), Gr is the Grashof number for heat transfer (p2f gd3p|To - T,]/µ2fTf), Pr is the Prandtl number (cpµ/k)f, g is the acceleration due to gravity (cm sec-&apos;f, a is the concentration densification coefficient (1/p)(ap/ayA)T, yA is the mole fraction of component A at the gas-metal interface, yA is the mole fraction of component A in the bulk gas stream, cp is the heat capacity of the gas per unit mass at constant pressure (cal g-I OK-&apos;), and k is the thermal conductivity of the gas (cal cm-&apos; sec-1 OK-1). Mathers et al. tested this combined Grashof number

Jan 1, 1968

By B. G. Agnew

In evaluating fracture treatments, the need to answer such questions as "What zone or zones were actually treated?" and "What was the vertical extent of the treatment" is necessary, since determining effectiveness of the fracture treatwrent depends on knowledge of the reservoir portions stimulated and the vertical extent of the fracture system or systems. Injection of hot or cold fluids during fracturing operations will transfer heat to the surrounding formations and fracture faces mainly by heat conduction. Once injection ceases, temperature anomalies will develop opposite the zone fractured because the rate of temperature decay is less than that opposite zones heated or cooled by flow inside the wellbore or along the cement sheath. Thus, temperature surveys can he used to determine where fractures were generated outside the wellbore and to reasonahly estimate the vertical extent of the fractrires. This technique has been used successfully to evaluate over 344 fracture treatments at depths ranging from 1,500 to 15,700 ft. Examples of actual temperature surveys are presented to show how this method of fracture treatment evaluation can be used effectively and beneficially. INTRODUCTION Stimulation of producing or injection wells by fracturing is commonplace in the industry today. In evaluating these treatments, two basic questions which arise are "What zone or zones were actually stimulated?" and "What was the vertical extent of the treatment?". Knowledge of the reservoir portions stimulated and vertical extent of the fracture system or systems is vital to effect efficient and economical well completions and to assure maximum recovery from reservoirs. Not all fracture treatments are successful and many times the desired or anticipated results are not achieved. When this occurs, it is in most cases the engineer&apos;s job to determine why. The answer to this problem forms the primary basis for deciding whether to spend more money for development drilling and/or completion attempts or to forego additional expenditures to improve the particular zone. Also, knowledge of the reservoir portion actually stimulated is invaluable in planning future workovers. This is especially true in wells producing from reservoirs containing multiple porosity stringers or zones. For various reasons, all zones within a reservoir may not be perforated on initial completion with the thought of developing remaining zones through future workover operations. Because particular zones are perforated and fracturing fluid pumped down the wellbore, it cannot be assumed that all perforated zones are stimulated, and in turn drained. Occasionally, unperforated zones which are planned for future development are actually stimulated on initial completion. eliminating the need for future recompletion attempts or all zones perforated initially are not stimulated, thus requiring additional completion work to insure that these zones are adequately drained. In most cases, it is costly. if not impossible, to determine with assurance why a fracture treatment is unsuccessful after the job has been completed and the results tested. Further, unpredictable upper and lower fracture barriers. as well as possible borehole communications (cement channels), do not permit it to be taken for granted that the formation opposite the perforations is the only zone receiving the fracture treatment. Thus. evaluation of fracture treatments is necessary for efficient and prudent well completion practices and should be done simultaneously with the fracture job. Several methods have been used to locate fractured zones, all of which are based on detection of radioactive tracer materials added to either the fracturing fluid or the propping agent. These methods have been helpful in increasing knowledge of fracturing operations; however, they are seriously limited because of cost and inability to detect accurately radioactive tracer material more than a few inches away from the borehole. To develop a better diagnostic tool to pinpoint which zone or zones had actually been fractured, temperature surveys were run after fracturing with heated and cooled fluids. Significant temperature anomalies were found to exist opposite zones suspected of receiving the fracture treatment. From the analysis and examples to be shown, it can be seen that temperature surveys run in a stabilized wellbore several hours after a fracture treatment using hot or cold fluid will yield temperature anomalies which show the actual reservoir portion stimulated, the vertical extent of the fracture system or systems and a qualitative indication of the portion of fracture-fluid volume entering a given depth interval. The experience and illustrations used in this paper are based on evaluation of over 344 fracture treatments during the past three years using this diagnostic tool at depths ranging from 1.500 to 15,700 ft and fracture volumes ranging from 2.000 to 110.000 gal. Based on analyses of these temperature surveys, analyses of fracture gradients and evaluation of several open-hole packer impressions in the West Texas area,&apos; vertical fracturing is the predominant mechanism and horizontal fractures occurred frequently, if at all.

Jan 1, 1967