6th international forgemasters meeting, cherry hill 1972

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Papers Presented at the

SIXTH INTERNATIONAL FORGEMASTERS MEETING

Cherry Hill, New Jersey, USAOctober 1-6, 1972

CONTE NTS

"FLAME CUTTING PROCESSES IN THE HEAVY FORGING INDUSTRY"Reginald J. H. HuntBritish Steel Corporation, England

"AUTOMATIC FLAME CUTTING IN THE FORGING INDUSTRY"Dr. Hans HirschbergMesser Griesheim GmbH, West Germany

"ON-FORGE HEATING ON THE OPEN DIE PRESS"Thomas W. JohnsonBritish Steel Corporation, England

"NEW FORGING EQUIPMENT AND SOME STUDIES OF DIE FORGING METHOD OFLARGE FORGINGS"Dr, Shoichi ShikanoThe Japan Steel Works Ltd., Japan

"NEW SPECIAL ALLOY FORGING PLANT COMPRISING 2600T OIL HYDRAULICOPEN DIE FORGING PRESS INTEGRATED WITH 40T-M MANIPULATOR"Dr. Hideo OhsawaDaido Steel Co. Ltd., Japan

"THE HETEROGENEITY IN HEAVY FORGING INGOTS, STUDY OF THE INFLUENCEOF IMPURITIES AND ALLOY ELEMENTS ON SEGREGATION"Jacques ComonCreusot-Loire, France

"INFLUENCE OF 'A' SEGREGATIONS IN THE MECHANICAL PROPERTIES OFFORGINGS OBTAINED FROM VACUUM POURED INGOTS" .Santiago FernandezAstilleros Espanoles S. A., Spain

"THE SOURCE OF INCLUSIONS IN FORGING INGOTS"Roland B. SnowUnited States Steel Corporation, USA

"REMARKABLE CASES OF DEFECTS IN THE MANUFACTURE OF LARGE STEELFORGINGS"Tertulliano SalinettiTerni Steel Works, Italy

"PROPOSED ULTRASONIC CLASSIFICATION OF FLAWS IN LARGE FORGINGS"Dr. Giuliano CanellaCentro Sperimentale Metallurgico, Italy

"METALLURGICAL EVALUATION OF NUCLEAR HIGH PRESSURE ROTOR FORGINGS"Bohdan HasiukWestinghouse Electric Corporation, USA

"MANUFACTURE OF LARGE HEAD FORGINGS FOR NUCLEAR REACTORS"H. C. SmithBethlehem Steel Corporation, USA

"FORGINGS FROM GIGANTIC INGOT WITH 3,550 MM (140 INCH) DIAMETER AND400 METRIC TON (881,000 LBS.) WEIGHT"

PART 1 - "PRODUCTION AND METALLURGICAL DEVELOPMENTS"Dr. Saburo KawaguchiThe Japan Steel Works Ltd., Japan

PART 2 - "OPERATIONAL STRESSES AND ACCEPTANCE CRITERIA"Dr. Rudolf SchinnKraftwerk Union AG, West Germany

"IMPROVEMENT OF PROPERTIES OF BIG, HIGH STRENGTH TURBINE DISCS"Dr. Max KroneisGebriider BOhler, Austria

"METALLURGICAL CHARACTERISTICS OF FORGINGS FOR THE HYDRAULIC SYSTEMDJERDAP ON DANUBE"

Abduselam SarajligMetalurki Fakultet, Yugoslavia

"SOME METALLURGICAL QUALITY AND MECHANICAL PROPERTIES OF3% CHROMIUM-MOLYBDENUM STEEL FORGING MADE FROM ESR INGOT"Kou TakeuchiKobe Steel Ltd., Japan

"ECONOMICAL ASPECTS OF FORGING MANUFACTURE BY CONVENTIONAL PRODUCTIONAND BY ELECTROSLAG REMELTING"Dr. Manfred WahlsterLeybold-Heraeus GmbH, West Germany

"FUTURE POTENTIALS, PROBLEMS AND METHODS OF SOLUTION OF PRODUCINGLARGE FORGING INGOTS BY ELECTROSLAG REMELTING"P. J. WoodingConsarc Corporation, USA

"PRODUCTION OF LARGE FORGINGS FROM LD-CONVERTER STEEL"Hironobu Sakuda

.41Kawasaki Steel Corporation, Japan

"FATIGUE STRENGTH OF FLAW-CONTAINING LARGE MARINE STEEL FORGINGS"Keiichi ShiraishiThe Japan Forged Steel Society, Japan

"A COMPUTER-CONTROLLED AND FULLY INTEGRATED 1.800 Mp FORGING PRESS"Ola ForslundSandvikens Jernverks AB, Sweden

"NEW INTEGRATED 4000-TON PRESS PLANT"Dr. Hans HojasGebriider BOhler & Co., Austria

Abstract

This paper will describe the forging press installationcompleted by Maschinenfabrik SACK GmbH, DUsseldorf, Germany,at SANDVIK AB, Sandviken, Sweden.

The importance of bringing science into the forging shopsis reviewed and with this as a background the design of thepress and it's auxiliary equipment is discussed. The installa-tion consists of four furnaces, the 1.800 Mp forging press andtwo railmounted forging manipulators.

As practically all materials to be forged are high alloyedstainless steels the control of the forging performance hasbeen seen upon as most important in order to bring out soundproducts. The control systems for the press, comprising amongstall a processcomputer, are therefore described in a way thatthe advantages and the possibilities of such an equipment shallbe understood.

Beside controlling the press and manipulator movements thecomputer also acts as a main part of a malfunction detectingsystem. Such a system reduces the time to spot a great numberof various disturbances, hence reducing the break down timeconsiderably.

It is concluded that the forging press automatic controlalways is faster and more accurate than the manual control. Themanipulator automatic control tends to be slower than themanual control when there are few strokes per pass. However,the influence of a correct bite ratio on the soundness of theforging pieces makes the automatic control of the manipulatorsadvantageous in most cases.

Introduction

Greater efficiency and higher productivity is continuallydemanded in modern industry. This is particularly important forthe forging industry. Hampered by traditions and lack of suit-able equipment the art of press forging has varied very littlefrom the late 19th century to the middle of the 20th century.Forging techniques established within this period are still prac-tised and results, achieved by a great number of investigators,making it possible to change the forging from an "art" to ascience have been sparely used.

The new 1.800 Mp forging press installation with it's auxi-liary equipment, completed by Maschinenfabrik SACK at SANDVIK AB,offers a good example of the new thinking on the subject of opendie press forging.

General Re uirements

We know that producing sound and uniform products in coggingdown ingots to billets or forging bars and rounds can be a diffi-cult task, especially with high alloyed materials.

By using the knowledge of the deformation patterns, stressand strain relations and spread behaviour in the forging process,several advantages can be achieved.

The most important factors which determine the performanceof the material in these respects are the total reduction, reduc-tion per pass, bite, and the height-to-width ratio. Essential forthe inner soundness is that the ratio height/bite is kept underclose control. The analysis of how these factors affect the forg-ing is not the aim of this paper and therefore I only concludethat the form of the forging piece after a completed pass has tobe known and in respect of this form the bite and the reductionfor the next pass have to be chosen in a way that ensures the re-sult wanted.

In addition to these metallurgical aspects, we also have tochoose the way in which a typical workpiece should be forged inorder to obtain a high production. The possibilities to choosethe right way are limited but there are numerous ways to do itwrong. It must be accepted that it is outside the competency ofthe normal blacksmith to choose the right way. The forging cycletherefore has to be precalculated and fed into the forging equip-ment in order to provide the best result in all respects.

SANDVIK's Re uirements

SANDVIK produces high alloy steels and special alloys inform of tubes, strips and wires. The main tasks for a forgingpress equipment was to cog down ingots to billets suitable for

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further strip or wire rolling and to produce round bar materialsuitable for the extrusion presses.

The rounds, being the biggest part, about 30 000 tons perannum, are mainly forged from prerolled square or rectangularbillets to fixed round dimensions, 192 mm, 246 mm, 276 mm and345 mm. The qualities are mainly stainless steels with varyinganalysis and a minor part nickel-base alloys and some titaniumor zirconium alloys. The throughput of ordinary "easy to work"materials had to be high in order to economically compete withrolling in the blooming mill. On the "difficult to work" materialswhich generally have high ingot costs, the soundness of theforged product was of great importance, too.

We therefore made great efforts in the planning of the newforging plant to find the best equipment to fulfil our wishes.

The press force is a most important factor when designingthe press. The press has to be so strong that it can forge thematerial with the highest hot strength and with a bite ratiobig enough to ensure the inner soundness. Moreover no decreasein the speed of the press is wanted hereby. Experiences won byhot draw tests made calculations of the press force possibleby using the formulas given by Wistreich and Shutt. In additionforce measurements were made on our existing 2.000 Mp water-hydraulic press in order to check the calculated figures. Theshortest bite possible which still gave an acceptable productwas derived from studies of the macro and micro structures ofthe materials. Combining these factors we were able to estab-lish a press force of 1.200 - 1.500 Mp as a minimum.

We also put in a great deal of work to find a lay-out whichwould ensure a good material flow through the forging plant.(Fig. 1) This was essential as, when forging 6 - 12 tons roundbars per hour and every bar weighing 1,5 - 2,4 tons, a lot ofbars have to be handled.

The Furnaces

The furnaces, designed and built by Granges Engineering,Vasterås, Sweden, are all gas-fired with a propane-butane mix-ture. They consist of three car-bottom furnaces and one walking-beam furnace. The furnaces can heat the material up to a maximumtemperature of 1,3000 C and average throughputs are for thewalking-beam furnace 10 tons/hour and for the car-bottom fur-naces 3 tons/hour and furnace.

The heated material from the car-bottom furnaces is broughtto the forging equipment by a GLAMA charging manipulator (Fig. 2)and put on an up-and-down moveable turntable in the forging line(Fig. 3) where it can be taken by the forging manipulator.

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The material to be heated in the walking-beam furnace isput on a charging table (Fig. 4) by a crane and is automaticallyfed into the furnace at the end of a cycle starting with thetaking out a piece of heated material (Fig. 5). The heated ma-terial is brought out in front of the forging manipulator byfour hydraulically operated arms, the number of arms being auto-matically selected by means of an electronic device for length-measuring (Fig. 6).

The For in Press

With two manipulators, maximum access and visibility aroundthe press are necessary features. These requirements,have beenmet by a two-column, single-piece frame, pull-down press. Theupper tool is attached to the press frame in a way that thepress operator, with a single push-button, can turn it 900 ordisclamp it, leaving it standing on the lower tool. A tool slidetable operating crosswise to the forging direction can hold fivedifferent tools. The slide table can be locked in 11 positionsby push-button selection (Fig. 7).

The main features of the press are:

Press power:Press stroke:Press speed, max.:Lifting speed, max.Number of strokes, max.:

1,800/1,200 Mp1,000 mm180 mm/sec.350 mm/sec.

(2,000/1,330 tons)

(40 in.)(7 in./sec.)(14 in./sec.)

120 per min, by planishing with10 mm (3/8 in.) penetration.

The press has one press cylinder and two lifting cylinders.The load of the main plunger is transferred to the frame by apush rod supported in spherical bronze elements, thus avoidinglateral forces on the packings and bushings. Allowable full loadoff centre is within a circle of 200 mm.

The press is driven by an oil-hydraulic system comprising 6WEPUKO radial plunger pumps. Each pump has a maximum flow of 715dm3/min. (190 gal./min.) and the maximum work pressure deliveredis 315 kp/cm2 (4,500 PSI) (Fig. 8).

Most astonishing for those accustomed to hydraulic pressesis the way the new press cycles. The movements resemble more toa crankshaft drive than the typical hard up and down change-oversof a hydraulic press and intermittences at the change-over pointscan hardly be seen. This is due to an unusual way in operatingthe stroke of the big directional piston valves by means of anelectro-hydraulic feed-back system comparing a reference valuewith the actual positionsof the valves. The normal pressureshocks in the pipe system, depending on the decompression ofthe oil, are kept to a minimum, too, by means of this system.This should be of advantage for the piping.

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The For in Mani ulators

When the available speed and capacity of the forging pressare to be transferred into economical use, an efficient methodof handling the forging piece is necessary. In this installationtwo rail-bound manipulators having the following characteristicsare used:

Max. loading capacity:Max. moment of load:Max, gripping force:Max, speed of travelling:Max, speed of rotation:Full peel parallel lift:Downward tilt of peel:Lateral shift each side of centre line:Full travel:

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6.312.6

35.095018560

5200

12,000

MpMpmMpmm/sec.rev./min.mm

mmmm

(7 tons)(45 ft. tons)(40 tons)(3 ft./sec.)

(22 in.)

(8 in.)

(40 ft.)

The two manipulators make it possible to forge the pieceright through from end to end, hereby reducing the time for com-pletion of the work considerably and also offering a better uni-formity of the material forged (Fig. 9). By help of the two turn-tables, one on each side of the press, we can also continue toforge even if an eventual break-down occurs on one of the manipu-lators.

The travelling motion of the manipulators is very important.It has to be fast in order to make the available press speedefficient. It also has to be carefully controlled in order togive the bite ratios wanted. The importance of this has been men-tioned above. The drive necessary to stop and start the manipu-lators for every stroke at a stroke rate of more than 100 strokesper min, has to be powerful. Therefore in the case of this in-stallation, the manipulators travel continuously. The intermittencesin the forging cycle are taken care of by letting the peel carriertravel and by a hydraulic cushioning system stopping and accelerat-ing the peel only. This enables a high forging speed but requiresa closed loop infinitely variable speed control system which ismore expensive than operation the manipulators with a stop-startsystem. The control system for this will be described later.

Auxiliaries

In order to increase the efficiency of the forging equip-ment, a highly mechanized auxiliary equipment has been installed.I already mentioned the turntables and the charging manipulatorfor the car-bottom furnaces. For the walking-beam furnace, thetransfer arms are to be mentioned. It remains only the equipmentfor the ready-forged material to be reviewed.

The forged material is mainly put on a cooling bank of thewalking-beam type (Fig. 10). The transfer from the manipulatorto the cooling bank is carried out by four hydraulically operatedarms (Fig. 11). The number of arms is automatically chosen byinterlockings detecting the manipulator positions.

Automatic Control E ui ment

The prime part in the control equipment is taken by a pro-cess computer with a core memory storage capacity of 12 k. Thecomputer is a PDP8/I from Digital Equipment Corporation, Massa-chusetts, USA. It has been installed by ASEA, Vasterås, Sweden,who in close co-operation with SACK and SANDVIK have designedthe control systems (Fig. 12).

The computer works through a main program in real-time de-sign. This program consists mainly of interrupting routines. Bymeans of a device selector system and interrupting facilities,jump-outs of the main program occur and hence different tasksas controlling the press movements, the manipulator movementsetc, can be performed. A big interfacing system with solid statecircuits takes care of the information from and delivers the con-trol signals to the outer equipment.

Controllin the Press

By pushing a button "Forging Start" after having chosen thewanted way to bring in the forging data selected, the controlprogram for the press movement is called upon. Forging data canbe fed into the computer by three different ways.

1) Preselected values dialled in at the operator's pulpit.

2) Taking over actual forging values, achieved by the operatorwith manual control, as reference values.

3) From in core memory stored information read in by numericaltape.

Forging data necessary to start the press movement arethickness, stroke length, stroke speed and forging force wanted.The program calculates upper and lower turnpoints and starts thepress movement. The travelling of the press ram is registered bya pulse generator delivering four pulses per mm.

The program is scanning the value from the pulse generatorevery tenth millisecond. Near the turnpoints the scanning speedincreases thus enabling a higher accuracy for the signal revers-ing the press movement. The thickness is in this way held withtolerances less than I 1 mm. Corrections are made at every strokeand the actual value is displayed on figure tubes on the panelof the control pulpit.

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When the lower turnpoint is reached, the program for themanipulator control is called upon. The press program haltsfor ten milliseconds in order to give the manipulator programpriority, before it starts again.

Controllin the Mani ulators

In the case of the manipulators the control is a bit morecomplicated than for the press. By manual control a joy-stickis used for travelling, rotating and gripping movements. Whena manipulator has gripped, that is a grip pressure is built up,and the press is started, the integrated travelling is allowed.

The impulse to start is given by a selector switch withthe direction of the travelling given by the joy-stick usedby manual control.

Data for the program can be read in from devices presetat the control panel or from the core memory of the computerwhere they have been stored from numerical tapes. Data usedare bite per stroke, rotation or no rotation of the peel and,if rotating the peel, the rotation speed.

The computer calculates the available time between thepress strokes and gives out a speed reference to the hydraulicsystem. The speed reference gives a corresponding flow fromthe variable-flow radial plunger pump. The control of the pumpflow is in a feed-back regulation mastered by the actual valueof bite per ratio. This value is calculated with the help of apulse generator measuring the travelling length.

When the press starts to forge on the material, the con-trol situation becomes more complicated, and still more whenboth manipulators are to be working together which mostly isthe case.

The manipulator on the outgoing side is then operated asmaster and it's peel system is hydraulically locked to thecarrier part. When the press catches the material, it triesto stop the manipulator. In the next moment, when the presspenetrates, the flow of the material is faster than the mani-pulator and hence should try to push which would result in bend-ing of the forging piece. Therefore the master manipulator re-ference is given an additional correction that ensures that italways draws, the correction being calculated by the computer.

The slave manipulator on the feeding side has to be con-trolled in a way that it does neither push the master when thepress is open nor push the forging when the press catches thepiece and forges. The peel is in this case unlocked from thecarrier and moves hydraulically cushioned. The angle deviationsfrom the vertical position of the peel supporting system aremeasured and give additional corrections to the precalculatedspeed reference.

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Semi-Automatic Control

It is quite obvious that a highly sophisticated installa-tion of this size cannot only depend on cumputer control butrequires additional conventional control elements, too.

The manipulators have manual control elements with variablespeed control and the press is equipped with a mechanical con-trol gear. This control gear enables semi-automatic press cycl-ing at any point of the press stroke after the press has beenlowered manually to the required forging dimension.

ThT obtainable size tolerances with this equipment arewithin - 1 mm. This additional equipment was installed to per-mit operation of the forging plant in case of serious or ex-tended trouble in the computer system. It was, however, up tonow only used for a short period during the start-up of theinstallation.

Malfunction Detectin S stem

Though it is normal to end papers of this kind with themaintenance part, it is for this installation logically com-bined with the control system.

In the computer interface, a detection system mainly con-sisting of 627 digital entrances is built in. With a separateprogram all control relays and their status are constantlychecked.

When an error occurs, for instance a too high pressure,a low or high temperature etc., the system gives alarm and themalfunction is registered on a teletypewriter making it possiblefor the press operator to do the appropriate steps. By a commu-nication program it is also possible to call upon the error de-tection system from the teletypewriter and ask for the statusof the connected devices. In this way we can check up every in-terlocking, every error sensing device and a little more thanone hundred magnetic valves necessary for the performance ofthe forging. A good example how the system works, demonstratesthe valve supervision system.

The action of a magnetic valve is released by a signalgiven from somewhere and going through a chain of electric de-vices ending at a relay closing the circuit in which it actu-ates the solenoid of the valve. (In this moment a current goesthrough the coil and the valve acts.) A malfunction of the valvecan either be electric if, for instance that final relay is notfunctioning or the outgoing circuit from this relay is damaged,hence preventing the current to flow, or it can be mechanic orhydraulic in case the valve has stuck.

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In case the final relay has not responded, we ask the errordetecting system for the status of this relay and then know thatthe circuit before this relay must have a disturbance. When thefinal relay is operating but the outer circuit is damaged withno current flow, the malfunction is automatically typed out onthe teletypewriter hence locating the disturbance. With the finalrelay in the right status and a current flowing through the outercircuit, a pure mechanical fault has been spotted.

As the error detection time in the today's complicated sys-tems covers about 80 % of the repair time, a system like thisgives an enormous help in bringing down the break-down time. Ofcourse the man operating this error detection system has to havea complete knowledge of the logic of the coordination betweenelectric and hydraulic systems if the full benefits are to bewon.

The Control Desk and the 0 erator

The operator controlling the forging facility has a greatresponsibility even when the forging program is precalculatedand fed to the computer. In our installation he controls fromthe desk the walking-beam furnace discharging, the transfer armsand the turntables, the manipulators, the press and finally thecooling bank equipment. Much effort has therefore been spent onmaking the control desk functional and also selecting and edu-cating the right people.

The control desk (Fig. 13) has a front panel where allessential values are displayed. The controls for the press andpress auxiliaries are right under this panel. On either sidethe controls for the other equipment are situated, to the leftthe walking-beam furnace equipment and one manipulator and tothe right the other manipulator and the cooling bank. All start-ing-up controls are located on a special panel on the wall be-hind the operator.

The operators were selected by special tests composed toreveal the qualities necessary for the man on this job. A yearbefore t ' work sho t the w

sj.thj matnematics,_physicsl_hydraulics„materialknewledge and defnrmatiDDheory. Thid-was made in order togive them the right understanding-for their more or less pro-grammed work. The education ended with the special subject ofthe press installation comprising, among other things, a com-plete layout model of the control desk. Shortly before thestart of the installation, study trips to other forge shopsin Sweden and then also in Germany were made. This was in orderto make the operators aquainted with press forging as no-oneof the selected men turned out to have any experience in forg-ing. Having come back from the study trips, they took full partin the installation work, although the equipment was to be takenover key-ready and the installation should be made by the con-tractor.

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The For in Performance

Except for the first tries in order to test the equipment,no manually controlled forging has been made in the press. This,too, includes practically the manual setting of forging data.

The manual setting are too time-wasting on material withhard restrictions in the forging temperature intervals. In lessthan one year more than one hundred numerical tapes have there-fore been prepared and new ones are made frequently.

The tapes are prepared by the engineers responsible forthe press. The ta es are b the theor of deformation inorder to brin ou acadmaiRPoducts. The o-m.nsary_am_f2rmatioriofjhe_cross-sebtIans-af-ter-s,very_pass is calculated by, helpOf tjle-datas-givan-by_TDmijnqgli—and_Stringbr. ifowever, EcTi.1.-6c-tiSnsmust be made for materials with other spread behaviourthan those investigated by Tomlinson and Stringer. The selec-tions of reductions per pass are important and are made by shopexperiences combined with knowledge derived from hot draw testswhere deformation force and contraction are registered.

The tapes are fed into the core memory of the computerwhich is able to keep ten programs at a time. With the knowledgeof the planned production the feeding of the computer can bemade for rather long periods of work restricting the press ope-rators' responsibility to selecting the right program number onthe program dial.

After having gripped the forging piece and put it underthe press tools the operator starts the forging by pushing twobuttons "Data" and "Start". The press movement then starts andby turning the joy-stick in the wanted direction of travellingand putting the selector switch in "Integration" the manipulatorstarts. When the free end of the forging piece travels out ofthe press it is taken by the other manipulator by manual control.Having gripped the integrated travelling of both manipulatorsstarts by giving the direction also with the joy-stick of thismanipulator (Fig. 14).

The press thickness control holds the tolerances within- 1 mm when the bite is held within a tolerance of i 5 mm. Byplanishing or by the forming in rounding tools these valuesare easily kept.

Typical throughput values for the prerolled material forged-nto extrusion billets are:

270 mm square, 1 500 kg forged to 192 mm round - 6 000 kg/h

270 mm " , 2 400 kg -"- 246 mm - 10 000 kg/h 943)

330 x 290 mm , 2 400 kg -"- 276 mm " - 12 000 kg/h\v?a- ko,(1°-\

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As a comparison the throughput values for the same forg-ing, but made in our older 2 000 Mp press with one mobile forg-ing manipulator can be given:

270 mm square, 1 500 kg forged to 192 mm round - 1 700 kg/h270 mm " , 2 400 kg -"- 246 mm " - 3 300 kg/h330 x 290 mm , 2 400 kg -"- 276 mm " - 3 600 kg/h

Men occupied by the forging are for the new press threeand for the old press five. Hence the production per man andhour is increased with a factor 5-6. Our moveable costs forthe forging operations are hereby reduced with about55 %.Thishas been possible not only by the greater efficiency of themachines but also to a great extent by the precalculated forg-ing schemas. These schemas make it possible to reduce the num-ber of passes with at least two on every product to be forged.

Conclusion

The question of the degree of automation brought into forg-ing presses has been a great topic during the past few years. Ithas been said that automatic control should be restricted to theplanishing or rounding stads of the forging. The backgroundbeing that it would not be'4conomical for the cogging down state.Cogging and preforging operations were regarded as operationswhich could be carried out faster with manual controls.

With the possibilities of precalculating and presettingthe optimum forging data for the press and then starting withone or two push-buttons only, there is no doubt in my mind thatthe automatic operation is much faster than any operator. Iwould even take a pure upsetting into this judgement.

In the case of the integration between press and manipula-tors it is not this clear. For large penetrations with only afew strokes, for example cogging down the conical shape of aningot, it is doubtful whether the automatic operation is fas-ter than the manual control of the manipulator. Howevar- assooluas-tha_number of strokes to be made increase to let_us'say more ' -a-tion should be used,rt will then increase the forging speed and, most important,ensures that the bites are kept under optimum conditions hencegiving a sounder and more uniform product.

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ICURE 1

1',..-• ;74.V: •

ILURE 2

.1 4 1144f 8,1•11)- 0

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4%w

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FIGURE 9

FIGURE 10

FATIGUE STRENGTH OF FLAW CONTAINING

LARGE MARINE STEEL FORGINGS

BY HIRONOBU SAKUTA

KEIICHI SHIRAISHI

KATSURO SATO

THE JAPAN FORGED STEEL SOCIETY

ABSTRACT

For correct evaluation of flaws on steel forgings, it is

essential, the working conditions of the parts concerned and

strength of flaw containing materials should be first ascertained.

Very complicated as they are, the stresses on marine engine parts

are mainly such fluctuating ones as bending, torsion and/or

tension-compression, and their stress levels, which may differ by

kinds of parts, sometimes exceed +10kg/mm2 in case of bending

stress for example. Therefore, as material strength of marine

engine parts, fatigue strength should be principally considered.

Although much study has been made on relationship between

flaws and fatigue strength on steel forgings, the majority was

based on small test-specimens, and practically no tests were made

on large test-specimens having segregation-cracks/sand-marks at

right angles to fiber flow which are thought affecting fatigue

strength a great deal.

For clarification, therefore, of the above, large test-

specimens of 50-60mm diameter were taken from such flaw-containing

large carbon steel forgings at right angles to fiber flow. Of

these test-specimens, bending and torsional fatigue test were

made. Tests were also given them thereafter to check the effects

of removing surface flaws. These have revealed : (i) Where there-1-

is segregation-crack on surface of less than 2 mm in lengthwhich is visible, fatigue limits are 13kg/mm2 in bending and10kg/mm2 in torsion. ( As crack becomes larger, smallerbecomes fatigue limit ) (ii) Where there is sand-mark of0.1-0.3mm in diameter and 1.1-5.5mm in length, fatigue limitis 18kg/mm2 in bending. ( However, if it is of spindle shape,fatigue limit tends to become smaller as angle between sand-mark axis and test-specimen surface nears zero value) (iii) Allin all, subsurface flaws affect fatigue limits similarly assurface flaws. (iv) Removing surface flaws alone does notalways recover fatigue limits in relationship with inner flawdistribution patterns.

So, fatigue strength of large marine steel forgings wasreviewed from the above test results as well as the followingpoints

(i) Influence due to flaw direction.(ii) Size effect of defective material.(iii) Relationship between tensile strength/fatigue strength,

etc.On such overall consideration we made deduction and decision

of fatigue strength against each of stress on flaws differing inkind, direction and size.

This study was done by the Defect and Stress Level Committee,J.F.S.S., as part of collection of elementary data for compilationpurposes of the Standard for Surface Inspection of Marine SteelForgings.

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CONTENT

Page

I. Introduction....... ..... ....... ........... 1

Fatigue test with large test specimen ....... 2

1. Fatigue tests using large test specimenwith flaw contained .......... ...... owe**.

(1) Test material 2

(2) Test specimen 0040 .................. MOO 2

(3) Testing machine . ...... ........0000000000 2

(4)Test results OOOOOOOOOOOOOOOOOOOOOOOOOOOOO

(5) Study on results OOOOO . ........ 3

2. Bending fatigue tests using test specimen withgrooves 0 0•000•000.4000000•0•0 0•00•4 00•00.0.••• 5

(1) Test material and test specimen ...... 5

(2) Test methods ............ ....... oneeteme 5

(3) Dimpling methods of surface flaws 5

(4) Test results and study thereon 5

3. donclusions

III. Estimating the fatigue strength of large marinesteel forgings 6

IV. Acknowledgement 8

FATIGUE STRENGTH OF FLAW CONTAININGLARGE MARINE STEEL FORGINGS

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by HIRONOBU SAKUTAKEIICHI SHIRAISHIKATSURO SATO

THE JAPAN FORGED STEEL SOCIETY

This paper is an attempt to estimate the fatigue strengthof flaw containing large marine steel forgings by elucidatingthe effect of flaws appeared on large marine steel forgings.

I. IntroductionSteel forgings have been conventionally used as an im-

portant material in the construction of mechanical structures,and it is natural, they have had to meet the requirement expectedof them; namely, a sterner level of the intrinsic soundness ofmaterial itself than in the case of other materials. This demandfor the soundness of steel forgings has been further spotlightedin the recent years by a trend for mammoth ships and high speedships, in turn requiring necessarily ever larger units of thismaterial or having higher levels of stress.

However, such steel forgings are being turned out on anindustrial production basis, and flaws often appear in theirimportant parts: further, the utility frequency of non-destructive tests has immensely increased these days, permittingvery minor flaws which have hitherto left undetected to bebrouth to our attention as their high detecting sensitivity. Thefact has brought us home the necessity of correctly evaluatingall detected flaws whether they are injurious or not to the partcontaining such.

For connect evaluation of flaws various factors contributingto them are to be considered of course, but here in this writingwe have decided to take up the strength alone of defectivematerial; especially fatigue strength as the marine engine partsare generally subject to variable stress.

With regards the effect of the flaws contained in steelforgings on fatigue strength, there have been many test-resultreports made on small test specimens of around 10mm 0, but onlya few numbers of such reports on large test specimens areavailable. It is to be noted above all that the fatigue test ofmaterials containing such vital flaws as segregation crack andsand mark at right angle to the fiber flow of materials (i.e.large test specimens smapled from the materials with flawcontained at right angle to their fiber flow) has never beencarried out before.

In view of the foregoing, the writers have given fatiguetests to large test specimens with a view to exploring andelucidating the above point. And these tests have enable us to

give the fatigue limit of such large steel forgings with flawcontained, and at the same time we could have the fatiguestrength estimated for large marine steel forgings with flawcontained from further study of the above test results as wellas the test results in •the past. Aside from this, with regardsdimpling, a conventional method of remedy for cases where flawsappear on surface, experiments have been carried out to checkthe effect such dimpling brings on bending fatigue strength.

II. Fatigue test with large test specimen

1. Fatigue tests using large test specimens with flaw contained

(1) Test material:The test materials used contained segregation cracks

and sand marks.For use in bending fatigue tests, were Test Material

A. which contains a large segregation crack, TestMaterial B. a minor segregation crack and sand mark, andTest Material C. sand mark only; while for torsionalfatigue tests; part of Test Material A was used.

As regards their production history, chemical com-position, mechanical property and microstructure, Tables1, 2, and 3, and Photo 1 in the Annex give them res-pectively.

(2) Test specimen:As shown in Fig. 1, test specimens were taken from

the test material at right angle to its fiber flow,finishing them up to the shape and size as shown inFig.2. Further, as to the surface of such test specimenvisual inspection as well as both Dye Penetrant andMagnetic Particle Inspections were carried out to checkthe flaw conditions prior to the fatigue tests.

Table 4 shows the size and number of flaws detectedin each of the test specimens; Photo 2, typical exampleof flaw indication; and Photos3 and 4, detected segre-gation cracks and silicate-sand marks. Six of 20 testspecimens (Nos.1 - 6) of Test Material A were used forbending fatigue tests; another Six (Nos. 7, 12, 14, 16,17 and 18), for torsional fatigue tests; and the re-maining eight, for fatigue tests in 2 hereunder.

(3) Testing machine:The testing machine used for the tests was of the

type and capacity shown in Table 5.

(4) Test results:Tables 6 and 7 show the results of the bending and

torsional fatigue tests and size of the flaw in whichfatigue fracture originated. Figures 3 and 4 are S-Ndiagrams drawn on the basis of Tables 6 and 7.

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(5) Study on results:(i) Test specimen A series (with segregation crack

contained) used for bending test had, as shown inTable 4, too many, varied size flaws to permit anyadjusted S-N diagram to be obtained; but it wasdeducible that in case of segregation cracks less than2mm in length, the fatigue limit would be the upperlimit value shown in Fig.3 (13Kg/mm2), and in caseof those of more than 2mm but less than 10mm inlength, the lower limit value (11Kg/mm2). It isadded for information that this upper limit value of13Kg/mm‘ agrees to the bending fatigue limit ofmaterial containing a fresh water corrosion crack(depth less than 0.2mm and length 0.2 - 0.5mm)reported in the past.'"

Photo 5 shows typical examples of the test speci-men surface after fracture and of relationship betweenflaw and crack initiation: in both it is observedthat segregation crack in the surface was made theinitiating point and that a comparatively long crackflaw, whose faint indication (shallow flaw) wasdetected by Dye Penetrant Inspection, was not madethe originating point of fracture.

(ii) Test specimen B series (with both minor segregationcrack and sand mark contained) for bending fatiguetests.

Compared with test specimen A series, test speci-men B series contained a smaller number of flaws, someof which even could not be discerned by the naked eyes.

Photo 6 shows a typical example of fracturesurface, where:

Test specimen B3 has a crack propagated origina-ting in a segregation crack of 3mm in length at thedepth of 1.5mm from the surface, and test specimen B4also has a crack originating in a spherical sand markat the depth of 0.5mm. The S-N diagram obtainedthrough connocting these two shows the fatigue limitof 17.7Kg/mm‘.

(iii) Test specimen C series (with sand mark contained)for bending fatigue test.

Test specimen C series contains minor sandmarks as shown in Table 4, with its fatigue limit at18Kg/mm2.

Photo 7 shows an example of fatigue fracture:all fractures there originate in sand marks.However, as apparent from Table 6 all the sandmarks involved are not necessarily those detectedby the pre-surface inspection—they include someof those oversight. Sizes of sand marks, theircontacting conditions with the surrounding areas,etc. probably play respective part in fatiguefracture. Therefore, a very complex relationshipbetween the sand marks detectable by inspection and

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those possibility of developing into initiatingpoints of fatigue fractures has been brough to lightby these informative media.

We paid attention initially to the points wherefatigue fracture is thought to have originated,selecting such test specimens as of approximateapplied bending stress levels.

Table 8 shows, on this assumption, differenceof the stress cycles to fracture due to the lengthof sand marks and the angle formed between the majoraxis of sand marks and the tangential line of testspecimen surfaces at the position of sand marks.It is observed in the Table that in cases where thelength of sand marks is less than 4mm the smalleris G, the smaller is, too, the number of the stresscycles to fracture.

From the above it is deducible that at 9.900the inclusion presents itself as a near-round shapeat surface, with fiber flow at that position directedtoward center; while when 0 becomes smaller, thesection of sand marks appearing at surface becomesan oval with the major axis on the circumferentialdirection, with the result that stress concernt-ration against bending stress is intensified,leading to shortening of the fracture lifew Ofcourse, we are to admit that the structural weaknessis simultaneously playing its role, but we may assertthat sand mark with a small angle (formed againstthe surface tangential line of the test specimen)does exert more influence to deterioration offatigue strength.

(iv) Test specimen A series for torsional fatigue test.Some dispersion is observed in fatigue strength,

but generally the fatigue limit of segregation cracksup to about 2mm in length is 10Kg/mm2, which, it iscensidered, should be reduced to 9Kg/mm in the caseof large flaws. There is an example of No. 14 testspecimen having been fractured at 9.5Kg/mm2 due tolarge flaw.

Photo 8 shows the fractured surfaces of testspecimens. Cracks A and B (Photo 8-a) havedeveloped into these length but stopped furthergrowth. The crack on the right hand side isfractured.

Photo 8-b is the fractured surface in this case,with the arrow pointing to the large segregation crackwhich started the fracture. It is added that forasymmetry of this fractured surface, the spot oforigin of the crack was responsible which is in thefillet part of the test specimen. Further, theprojections located sporadically over the fracturedsurface are the showing-ups of the segregation cracks

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in the interior; however, in cases where minutesegregation cracks are contained in the interior,appearances of comparatively smooth fractured surfaceswill be presented as shown in Photo 8-c.

2. Bending fatigue tests using test specimen with grooves

(1) Test material and test specimenThe test material used are same as described in

II-1-(1). The test specimens are same, too, as in 1.for bending fatigue tests, which however having beendimpled for surface flaws, have various grooves madein the parallel and fillet parts due to the number andsize of the flaws involved.

(2) Test methodPre-surface inspection was carried out similarly

as described in 1. Photo 9 shows an example of the indi-cation detected by Dye Penetrant Inspection. The fatiguetesting machine used here is the uniform bend type shownin Table 5.

(3) Dimpling method for surface flawsOf the flaws detected by Dye Penetrant Inspection,

the crack-like indications and round-shape ones of above1mm 0 were taken up for treatment, which were completelyremoved by a small grinder, and the grooves of whichwere finished off with sand-paper. Dye penetrantInspection was given to confirm the conditions afterremoval of flaws. Photo 10 shows an example of dimpling.

(4) Test results and study thereonTable 9 shows the number of grooves, size of the

largest one among them, and fatigue test results.As noted in the Table, all fracture originated in

the remaining flaws (which we failed to detect due todeterioration of sensitivity through repeated use ofDye Penetrant Inspection) or from the flaw in thesubsurface of groove bottoms. Therefore, the resultsof complete dimpling case have not been available.

Fig. 6 represents S-N diagram based on Table 9where the results of bending fatigue tests for TestMaterial A in 1 above are concurrently recorded. Thefigure permits us to conclude that the material, whosesurface flaws have been completely dimpled but which isfractured due to the flaw in the subsurface of goovebottom has a fatigue limit of 13Kg/mm2 (this is inci-dentally a similar level envisaged for material with asmall surface flaw of about 2mm), and that, in case ofthe remaining segregation crack at the bottom of groovepart, the material affected has 11Kg/mm2 fatigue limit

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3. Conclusions

(equal to a test specimen having a large segregationcrack) or less.

(1) Summing up the test results in the foregoing, thefatigue limits of large test specimen of 60 mm $ are asin Table 10.

(2) From the results of (1) above and the fatigue testresults for material of 60 mm $ level published here-tofore(2V3) , relationship between the size of flaw andfatigue limit can be obtained as in Fig. 7. Based onthis Figure fatigue limit against flaw is deducible bythe following formula:

awd=0"Wo-Ad (1+log102)

where,avid: Fatigue limit (Kg/mm2) of material with flaw

of long. axis;k (mm)awo: Fatigue limit (Kg/mm2) of sound materialAd: Value (Kg/mm2) given by kind and direction

of flaw and kind of stress.Note: In case of torsion, wd shall be altered

to readirwd, and awo to readtwo.Ad value in the above formula may be given from Fig. 7and Table 11.

(3) Effect of dimpling the surface flaws depends on theflaw distribution in the interior of test specimens:namely, in the event distribution density of surfaceflaws is high, mere dimpling of surface flaws can nevercompletely achieve the end envisaged as there is alwaysan extremely big probability of flaws existing in thesurfacial strata, and on top of that, such dimplingwould work to intensify stress concerntration, adverselyaffecting the material concerned.

(4) The relationship between the angle (G) folmed by sandmark axis and tangential line of test specimen, andfatigue limit shall require further study based on muchmore experiment data.

III. Estimating the Fatigue-Strength ofLarge Marine Steel Forgings

As large marine steel forgings, those of sizes, 500mm -1,000mm$ are of general demand. However, these steel forgingscontain without exception a stress concentralation part, fromwhich fatigue fracture is usually initiated. Therefore, forprevention of accident of this nature it is devolved on us toensure the soundness of material for this part; not only thatit is indispensable for us to grasp the fatigue strength there.

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For this purpose, test specimens of a size approximate to thestress gradient of that part shall be used.

The test results in II were obtained by using test specimensof around 60mm$, but strictly speaking, for evaluation of steelforgings as large as 1,000m levels we should be equipped withfatigue strength data obtained through test specimens of about100mm 0 at least,(4) otherwise our present information wouldprove insufficient for designing or inspection purposes for suchlarge steel forgings.

Generally, the size effect on fatigue limit of 100mmagainst 10mm 0 in sound material is said to be 0.85 - 0.9 in caseof bending and 0.8 in case of torsion or 0.6 of large rotarybending fatigue limit,(5) but in the case of flaw containingmaterial the size effect is said to be correspondingly larger.(i) The size effect of bending fatigue of material with flaw

containedThe bending fatigue limit of segregation crack of less

than 2mm is 13Kg/mm2, agreeing to the bending fatigue limitof material with fresh water corrosion crack as stated in II.According to this information») the size effect of 100mm 0against 50mm 0 is about 0.8, from which the size effect ofbending fatigue of 100mm 0 against 60mm 0 can be deduced tobe 0.8.

(ii) The size effect of torsional fatigue of material with flawcontained

Regarding the size effect on torsional fatigue limit,no information whatever is available: furthermore, torsionalfatigue tests cannot be carried out with a test specimen of100mm 0 under the presnet circumstances. These make anyguess work about size effect very difficult. However, inview of the fact that the effect of material flaw ontorsional fatigue limit is generally pretty smaller than inthe case of bending fatigue, it is unthinkable that the sizeeffect alone is especially large, and it is logical thereforeto regard the size effect being same with, or less than thatof, bending: namely, 0.9 against 60mm 0.

From these size effects and test results in II, thefatigue strength of large marine steel forgings has beendeduced as in Table 12. The star-marks (*) in the Tableindicate our deduction on the basis of the experiments inthe past's) plus consideration for the size effects in theabove.

Table 12, which has primarily been compiled as thebasic information for "Standard for Surface Inspection ofMarine Steel Forgings", has made no attempt for too fineclassification of flaws for fear of inviting uselesscomplication. So in the Table, flaws are graded simply asA, B and C, and fatigue limit is given out there belongingto each of the three grades respectively.

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IV. Acknowledgement

This study has been carried out under the full and wholehearted support and cooperation of Nippon Kaiji Kyokai (ShipClassification Society of Japan, NK), and its results are nowvery effectively utilized in the establishment work of ourmember's own standard for surface inspection of marine steelforgings, and further, as the basic information for crank-shaft design standard and inspection standard") by NK.

The writers wish to express on this occasion their sincereappreciation and gratitude to each of the members engaged in thisstudy.

References(1) J. Hoshino - No.55 Pre-print of 39th Meeting of Japan

Society of Mechanical Engineers (Nov. 1961)(2) The Japan Cast Steel Society - No.2 Report on

Relationship between Flaws and FlawPatterns by Non-destructive Inspection(March, 1965)

(3) Y. Ueda - Internal Combustion Engine of Japan vol. 7No.21 (Nov. 1963)

(4) T. Iki - Internal Combustion Engine of Japan vol. 3No.25 (July, 1964)

(5) Design data book on fatigue strength for metal (KinzokuZairyo, Tsukare Tsuyosa no Sekkei Siryo, inJapanese), Japan Society of MechanicalEngineers.

(6) Nippon Kaiji Kyokai - Rules and detailed Rules for DieselEngine Crankshafts and these Explanations1972.

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Material

A

Table 1 History of material employed

Ingot weight Forged shape Forging ratio

Kind of flaw contained

(ton)

(mm)

15

5004 x 1150

3.7

Co-existense of large

segregation crack and

sand mark

15

500x 100

x 610

41 0 x

410

x620

6.0

Co-existense of minor

segregation crack and

sand mark

4.5

Sand mark only

Table 2 Chemical composition of material employed

Material

,- (%)

A0.35

B0.23

C0.31

Si (%)

0.25

0.29

0.22

Mn (%)

0.60

0.72

0.67

'P (%)

0.018

0.019

0.009

S (%)

0.010

0.014

0.012

Table3

Mechanical properties of material employed Remarks

Segregation crack in

fracture surface

No material flaw in

fracture surface

No material flaw in

fracture surface

Notes: 1. Tensile test specimen for materials A and B has been taken at right

angle to their fiber flow.

2. Tensile test specimen for material C has been taken in parallel to

fiber flow.

T.P. No.

Table 4 Distribution of surface flawsin the test specimens

Por Material A

Flaws detected by naked eye

Total number of flaws, Max. size of which

Al 9 6 mm crack

A2 20 25 mm crack

A3 18 18 mm crack

A4 9 5 mm crack

AS 16 • 5 mm crack

A6 16 4 mm crack

A7 many 0.5--2 mm crack

Al2 many 0.5--2 mm crack1 6 mm crack

A14 many 0.5--2 mm crack1 5mm crack at fillet

part

Al6 many 0.5--2 mm crack

Al7 Same as above

Al8 Same as above

For Material B

T.P. No.Flaws detected by dye penetrant inspection

Total number of flaw indications, Max. size of which

Bl 10 above 0.2 mm6 1.5 mm6

B2 5 above 0.2 mm6 3.5 mm4

B3 3 above 0.2 mm6 0.5 mm$

B4 3 above 0.2 mm/ 0.5 mm4

B5 2 above 0.4 mm6 0.5 mm6

B6 3 above 0.2 mm6 0.5 mm4

-A3- to be continued

Table 4

T.P. No.Potal number and size of flaws (and flaw indication)

2 x 0.14, 1 x 0.24(4 x 0.14,3 x 0.34)

02 8 x 0.14, 1 x 0.34(1 x 0.14,5 x 0.34)

C1

2 x 0.14, 1 x 0.1"T x03 (1 x 0.34, 1 x 0.3x 2.0-4)

04 5 x 0.14, 3 x 0.34, 1 x 0.1wrx 1.0(2 x 0.14, 2 x 0.36, 1 x 1.04, 1 x 0.314) x 2.0)

4 x 0.14(1 x 0.14, 9 x 0.36)

2 x 0.14,3 x 0.34, 1 x 0.1tcx 0.3 , 1 x 0.1arx 0.61 x 0.1w-x(9 x 0.34, 6 x 0.54, 1 x 1.04, 1 x 0.51-0-x 1.541 x 0. 3urx 2.51)

05

06

For Material C

Flaws detected by naked eye and dyepenetrant inspection

10 x 0.71, 1 x 0.1urx 0.51, 1 x 0.1/0-x 1.21C7 (1 x 0.1 , 6 x 0.34,2 x 0.54, 1 x 1.04)

08 26 x ur4 x 0.34, 1 x 0.1x 2.01(1 x 0.1 , 13 x 0.34, 3 x 0.54, 2 x 1.04)

8 x 0.14, 6 x 0.34, 1 x 0.1ur x 0.81; 2 x 0.1urx 1.0109 (9 x 0.34, 3 x 0/54, 1 x 0.74, 1 x 1.04, 2 x 1.54,

1 x

6 x 0.14, 6 x 0.36, 1 x 0.1,0" x 1.01, 1 x 0.5urx 1.1iC10 (2 x 0.14, 22 x 0.34, 4 x 1.04, 2 x 2.04, 1 x 2.56)

Notes: 1 The figures in the parenthesis show the flawindication detected by dye penetrant inspection.

2 ar: Width of flaws (or flaw indication) : Length of flaws (or flaw indecation)

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Table5 Fatigue testing machine used for the experiments

Kind of testing Type of machinemachine

Bending

Tortion

1 Material T.P. StressNo. (kg/mm2

A

B6

Table6 Bending fatigue test results

Al 13.7 3.78 x106

1 A2 12.0 2.64 x106

A3 11.0 8.41 x106

A4 10.5

A5 11.5

A5 13.0

uniformbend-type

cantilever-type 400

resonance-type 1000

No. of cycle

B4 18.3

>10.25 x106

>10.08 x10

>10.15 x10bm

B1 22,0

B2 16.0

I B3 21.0

B5 17.0

17.7 )10.53 x10

-A5-

Load Revolution Materials used(kg-m) (rpm) for experiment

2000 1500 material A and

1200

1800c m

Condition of fracture

Fractured from segregationcrack of 2mm on the surfaceat fillet part of test spe-cimen

Fractured from segregationcrack of 7mm on the surface

Fractured from segregationcrack of 9.5mm on the sur-face

Not fractured

Not fractured

Not fractured

Fractured, but no flaws in thfracture surface

Not fractured

Fractured from segregationcrack at 1.5mm distance fromthe surface

Fractured from sand mark at0.5mm distance from the sur-face

Not fractured

Not fractured

material C

material A

to be continued

Table 6

0.5 x 1.1= (2=4 and 2.5=4)on the surface

Note: The figure in the parenthesis show the size of indicationby dye penetrant inspection.

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T.P. No.

Stress

(kg/mm2)

Table 7 Torsional fatigue test results

No. of cycle


Not fractured. But two fatigue cracks initiated

A7

10.0

>16.0 x 106

diagonally from the tip of circumferential

segrega-

tion crack and propagated up to the length of 2mm.

1106

Fractured from diagonal segregation crack.

--.]

Al2

11.5

3.05 x

No initiation seen in axial segregation crack.

iFractured from segregation crack (25mm depth) in

A14

10.5

3.11 x 106

fillet part of specimen where there is 9.5kg/mm2

stress.

A16

9.0

>25.0 x 106

Not fractured.

Al7

12.5

1.81 x 106


Al8

14.0

0.62 x 10


Table8 Difference of the stress cyclestofailure due to the existing stateof silicate inclusion at the sur-face

Reversed bending Small 4-- (Stress cycle tostress fracture) -+ Large(kg/mm2)

18.8--19.2

18.7

18.3-18.4

CD

No. 10 -- No. 8 -- No. 5.2=5.5mm Ag=2mm 2=3.8mm0 =42° 0 =22° g=61!

No. 7 --- No. 1I1/2-3mm -e=1.5mm9 =15° 0 =80°

No. 6 - No. 2 — No. 9= 5. 3mm 1=1. 5mm bz l. 6mm=76° 9 =16° 9 =61°

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Table 9 B'ending fatigue test results of surface treated specimens

T.P. Total

Max. size of which Stress No. of cycle

No.

number

(length x width

(1g/mm2)

of grooves x depth)

in mm


A8

19

30 x 9 x 2.0

16.0

1.67 x

106

Fractured from segregation

crack in the interior

A9

924 x 7 x 2.0

14.0

1.97 x 106

Same as above

A10

920 x 7 x 1.5

13.0

>10.49 x 106

Not fractured

All

17

25 x 13 x 1.5

13.5

1.68 x 106

Fractured from segregation

crack in subsurface

6Fractured from remaining crack

aA13

14

24 x 13 x 1.5

13.0

1.20 x 10

in the bottom of grooves

\o 1

(1mm in length)

A15

20

40 x 12 x 2.0

13.0

1.50 x 106

Same as above (but 1.5mm in

length)

A19

13

30 x 17 x

3.5

12.5

1.51 x 106

Same as above (but 3.5mm in

length)

A20

25

30 x 13 x 3.0

15.0

1.30 x 106

Fractured from crack in the

interior

Table 10 Fatigue test result of large test specimen with flaws contained

Kind of

Test specimen

Flaws

Material

testinDia

Sampling direction

kind

Location

size

g (mm)

con )

A

RB

61.5

A

RB

61.5

AD

RB

61.5

A

AT

61.0

A

AT

61.0

RB

6o.o

RB

50.0

Right angle to

fiber flow

Notes

AD:Dimpling Material A

RB: Rotary Bend

AT: Alternate Torsion

Segregation Surface

crack

Fatigue

limit

(kg/mm2)

0.5 --2

13.0

SegregationSurface

2--10

11.0

crack

SegregationSub-surface 0.5--2

13.0

crack

SegregationSurface

0.5--2

10.0

crack

SegregationSurface

2--10

9.0

crack

Segregation

crack and

Sub-surface max. 2

17.8

sand mark

Sand mark

Surface

1.1 -

5.5

18.0

(0.1— 0.34)

Table ll Ad value

Kind of flaws Sampling direction of test specimen Kind of stress Ad (kg/mm2)

Sand mark

Rotary bend

5.5

Segregation

Right angle to fiber flow

Rotary bend

7.5

crack

Segregation

Alternate

crack

torsion

3.5

Noflaws

Grade A

Grade B

Grade C

Table 12 Relation between flaws and fatigue strengthin large forged steel material

Fatigue strength(kg/mm2)Kind of DirectionKind and size of flawflaw of flaw Bend Torsion

Those that cannot be L* 20fm 12fmdetected by non-des-tructive inspection* T* 19fm 12fm

Those that can be de- L* 19fm 12fmtected by non-destruc-tive inspection but T* 18fm 12fmof microscopic size*

Non metalic inclusionsdetected by visual in- L* 18fm llfmspection but not be-longing to the below T 15 *11fmcolumn

Spindle-shaped sand L* 17 9marks of 2mr0 and over T 11 9

Segregation cracks and L* 17 9porosities of 0.5--2mm T 11 9

Segregation cracks, L* 15 8porosities and pipes,9 8etc. of 2mm and over

Notes: 1 Direction of flawL: Where the major axis (direction of fiber flow

of forged material) of flaw intersects with themaximum principal stress at an angle less than 45°.

T: Where the major axis of flaw intersects with themaximum principal stress at angle of 45°--90°.

2 fm: Material Factor=11-2/3(TiK/45-1)

3 Tensile strength of material(kg/mm2)

-Al2-

ietara -14•,--e • r fl UP,

,skr -11*

254d { -4 Pi (L. 1.4 _rf 1•All•

•

A t‘ • ' —1^. 4a., .144<0-•

idip4 A

•33

gt int.),,.L.:.i ....hole.:•...„ A.-1,3t,. •# 4•1 "Pli .kc'b ah A

rtzio. 11». • ...- • "%Ai ,0.0t 7 # ,-.4x,7 444., A.?dieh sh N.') ,-ash s, ?, 7. til ;es_, - 4-crtr: at i.4s„, 4-44

x 1:0Ma LeriaL A

Material_ C

* ,

x 1.00

Le itat fir ":"toritobr-4,

Photo 1 Microstruc ture of material employed

Photo 2 .typical example of flaw indication in thesurface of the test specimen from Material A(Detected by magnetic particle inspection)

x 50Photo 3 Typical example of segregation

crack detected in the test spe-cimen

Photo 4 Typical example of sand markdetected in the test specimen

-A14-

(1) Distribution of surface flaws

(2) Location of fracture

(3) Fracture surface

Photo 5 Relation between flaw and crackoriginating point of A2 specimen

-A15-

(1) 93 specimen(Arrow shows the segregation crack atthe origin of fatigue fracture)

(2) BL specimen(Arrow shows the sand mark at theorigin of fatigue fracture)

Photo 6 Fatigue fracture surface of specimenfrom Material B

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J111111111111111111111

specimen. C7 specimen

Photo 7 Fatiue fracture surfaces and enlarged ohoto-.5r:I0h.5 of origin of fatigue fracture(Arrow shows the inplusiOn at the or.ta.in offatigue fracture)

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09 specimen

t • •„.•:1;;;StityftrtAgitt404

(I) Locationof fracture of A14 spec imen

(2) Fati4ue fracture sur face of Alit specimen

( Example of lare sej,reation crack )

• • •„:„J,t,;* • , •

•...)%akt- • •

• ,••

(":3) Patlue fracture surface of AIS specimen

( Example of small segreation crack )

Photo B •ypical example of torsional fatiue fracture

Photo 9 Flaw indication of A13 specimenbefore flaw elimination

Photo 10 Surface condition of Al3 specimenafter flaw elimination

-A19-

(1) Location of fracture

4/()!m",:, 7

(2) Fatigue fracture surface(Arrow shows the flaw at the origin of fatigue fracture)

Photo 11 Fati,,ue fracture surface ofA9 specimen with grooves

-A20-

(1 ) Location of fracture

(2) Pa tigue fracture surface

(AT ow shows the flaw at the oritai.n of fatigue fracture)

Photc 12 Fatiu e frac tor() surface ofAlp specimen with groove

-A21-

0 1 5 4 5 I I 12 /3 14 /

2 6 7 g g /0 /6

Forgeddireaion

(1) Material A

No.1 No2 No.3 Na4 No.5 No.6

50 100 100 100 100 100 50

600

8

-A22-

620

Faged direction

(2) Material B

I \ N\ \\ \ \\

\ \I \ N1

N \ \\ • \ \ \ \

\ \N \ N•

N • \ NN N \

N

\ 2 4 6 8 10

I 3 5 7 9 o9

Ccl

180 900

—100

Forged direct ion

(3) Material C

Fig.1 Method of sampling test specimens

OZZ,

Taper 1 /20

cz

500

s.

HO 40 50 100 50 40 HOd----61.50 for material A

dr--60.00 for material B

(1) Bending fatigue test specimen from Material A&B

32.61WR 1WR

Taper 1 /40 Taper 1 /40

50 IZO

280

(2) Bending fatigue test specimen from Material C

120R -- 120 12OR

Taper 1 /50 Taper I /50

Taper 1 /20

fir ,aI I

CNJ

:3180

/25 200 /25 450

(3) Torsional fatigue test specimen from Material A

Fig.2 Fatigue test specimen

—A23—

26A C 1 ' C 4 ' 0 Material A

C 5 ' 9 '24 A • Uppe r limi t

\C 3 -- -0-- — Lower Emit

22\\N N

B 1 0 Material B

6 —` • `23, ,L Material Ce B 3 -...5 20..„. •-•,.. C 5

no C8g, --, C 4

222

B 4 C 2

a>

-:- - B 6

C 1

222 B 5

;11 16 0- B 2CO

to..g. 14

A 1

-0a22)

A 6

----- A 2Ca 12 •222..,,

10

C10 C 9C 9 C 6 _

101 5 10°Number of cycles to failure

15

14

74:10a0—202-,

9

8

72

Fig.3 S--N diagram ( Bending )

Al8

Al7

tao

2n

1015 lot

Number of cycles to failure

Fig. 4 S--N diagram ( Torsion

-A24-

A 5

A-93A 4

A 7

10'

Al6

24

106Number of cycles to failure

-A25-

Fig.5 Shape of the bottom of thegrooves by dimpling

• Surface treated material81 0 Material A

22 N 0 Upper limit0 - — Lower limit

B 3.11/4 &Material Bc; 20 ........

-...., B 4---,to

18

a

% -A. _ _ 8 6

in B 5cn A 8a)Li 16-. 82(I) A2060.0a 14 A 9 Alo

^8 A11• A10c -°..A 15a) 12A13 °'---- A 2 A 6CO Al9

A 5

Z310 A 4

Fig.6 S--N diagram of surface treatedspecimen (Material A)

flaw

Aroove

22

20

8

18

16bD

14

S 12tiptti

LT-1 1 0

0 Bending, Artificial round holes Bending, Sand mark0 Bending, Segregation crackX Torsion, Artificial round holea. Torsion, Segregation crack

—A26—

•

0.2 0.5 1 2 5

Size of flaws (mm)

Fig.7 Relation between the size offlaws and the fatigue limit

PRODUCTION OF LARGE FORGINGS FROM LD— CONVERTER STEEL

KAWASAKI STEEL CORPORATION

by Hironobu SakudaDr. Hiroshi Ooi

At the Mizushima Works of KAWASAKI STEEL CORPORATION, all kinds of

large forgings are produced from ingots of vacumrdegassed LD—convertersteel.

Investigations were carried out on (1) characteristics of solidi—fied structure of a 100—ton ingot, on (2) properties of a forging madefrom a 100—ton ingot and on (3) production and properties of a commercialforging made from a 200—ton ingot.

The method and the results are summarized as follows:(1) A 100—ton ingot of LD—steel degassed by the ASEA—SKF process

was investigated in detail, The ingot shows a sound macro—structure without any marked segregation. Decreasing oxygencontent by the ASEA—SKF treatment contributed to so small anaccumulation of oxide inclusions in the bottom region of theingot.

(2) A forging of 1,140 mm in maximum diameter and 12,300 mm inlengthwasmade from a 100—ton ingot producedby the sameprocess. Macro—structure and distribution of elements andinclusions, as well as mechanical properties, were investi—gated throughout the forging, The forging shows satisfac—tory characteristics.

(3) A slab—cooler shaft of 1,200 mm in maximum diameter and 17,250 mmin length was made from a 200—ton ingot of LD—steel degassed bythe RH process. After the shaft was normalized and tempered,test blocks were cut off from its top and bottom ends. Then,macro—structure, distribution of elements and mechanical pro—perties were investigated. No defects were detected byultraSenic flaw inspection of the shaft.

From the results thus obtained, it was confirmed that forgingsfrom LD—converter steel are in no way inferior to those from electric

furnace steel or open hearth steel.

I. INTRODUCTION

PRODUCTION OF LARGE FORGING FROM LD-CONVERTER STEEL

Recent development in the building of bulky ships, huge steel mills,and big machines for heavy industries and power plants has called formuch higher qualities of considerably larger forgings.

Forgings are mainly manufactured from vacuum-degassed steel ingotsproduced either by electric furnaces or by open hearth furnaces.

Main features of the present paper cover:

1) The characteristics of solidified structure of a 100-ton ingotof LD-converter steel vacuum-degassed by the ASEA-SKF process.

2) The properties of a forging made from a 100-ton ingot producedby the same process.

3) Production and properties of a commercial forging made from a200-ton ingot of RH-degassed LD-converter steel.

2. METHOD

The dimensions of the carbon steel ingots and those of forgings areshown, respectively, in Fig.1 andFigs.2and 3.

Investigation was carried out as follows:

1) The characteristicsof solidified structure of the 100-ton ingot

Macro-structure, distribution of elements and inclusionsand formation of inverted-V segregation were studied in detailby longitudinal and transversal cutting of the 100-ton ingot.

DI

Fig. 1 Dimensions of 100-tonand 200-ton ingots.

100-toningot

200-toningot

H (mm) 3,429 4,155

D1 (mm) 2,318 2,898

D2 (mm) 1,844 2,470

C.4

Top

.czt

12,300

"A.

Fig. 2 Dimensions (mm) of the forging

made from a 100-ton ingot

r-I

Bottom

Top

1

Test Block

17,250

Fig. 3 Dimensions (mm) of the slab-cooler

shaft made from a 200-ton ingot

'Q.

-fa

0 'la

0 -E

k0

0 0

0 0

NI

r-i

00

r-4

' M

r.-.

1

Test Block

Bottom

2) The properties of the forging made from the 100-ton ingot

The forging of 1,140 mm in diameter and 12,300 mm in lengthwas cut both longitudinally and transversally, then macro-structure and distribution of elements were investigated. Also,the mechanical properties of specimens taken at several positionof the forging were investigated.

3) Production and properties of the commercial forging made fromthe 200-ton ingot

A slab-cooler shaft of about 1,200 mm diameter and 17,250 mmlong was made for a slabbing mill plant from the 200-ton ingot.After the shaft was normalized and tempered, test blocks were cutoff from the top and bottom of the shaft. Then, investigationwas carried out on macro-structure, distribution of elements andmechanical properties.

3. RESULTS

3.1 Characteristics of Solidified Structure of 100-ton Ingot

3.1.1 Process

The ladle analysis of the 100-ton ingot is shown below:

C % : 0.31

Si % : 0.28

Mn % : 0.76

P % : 0.009

S % : 0.007

Vacuum-degassing time of the melt for the ingot was 35 min. and castingtemperature was 1,575°C.

A "bar-test" method was employed to determine the rate of verticalsolidification of the ingot. The result is shown in Fig. 4. Time re-quired for complete solidification of the ingot excluding the feeder partwas about 17 hrs.

3.1.2 Macro-structure

Examination was made on the longitudinal section including the axisof the ingot as well as on the cross sections at 1/6, 1/2 and 5/6 of theingot height as illustrated in Fig. 5.

Macro-structures and sulphur prints are shown in Figs. 6 and 7 forthe longitudinal section, and in Figs. 8 and 9 for the cross sections.

5000

4000

3000

2000

1000

0

0

.111 ••••

Top of the ingot

500 1000

- 6 -

time (nin.)

Fig. 4 The vetical solidifcation rate of100-ton ingot measured by "bar-test"method.

:investigated section

Fig. 5 Location of investigatedsections of 100-ton ingot

- 7

5/6 H Top

1/2 H Middle

1/6 H Bottom

*-0

og

Top(5/6)

Middle(1/2)

Bottom(1/6)

fr6S3C4A

Fig. 8 Macro-structure on the transversalsections of 100-ton ingot

- 10 -

44-6 6

b.0

rri

In the longitudinal section, a segregation zone of inverted-V shape extendsfrom 1/10 of ingot height towards the top. The segregation zone facescolumnar crystals towards the outside and equiaxed crystals towards theinside of the ingot. The V-shaped pattern appearing at the core of theingot does not represent ordinary V-segregation. Instead, it shows dif-ferent sized grains piled up layer by layer. The ordinary V-segregationis very faint.

The inverted-V segregation zone is circular on the bottom and middlecross sections, but becomes filled circular shape on the top cross section.Segregation lines in the zone at the top cross section appear as dots whichget larger towards the center of the cross section, while the lines at thebottom cross section show up as short fibers. The appearance of the linesat the middle cross section is similar to that at the top cross section inthe outer part of the zone, and to that at the bottom cross section in theinner part.

Solidified structure of the segregation zone itself is mostly equiaxedat the bottom, columnar at the top (extensively towards inside) and inbetween of the two at the middle.

3.1.3 Distribution of Elements

The distribution of elements is shown in Fig. 10, and the highest andlowest values of each element in the ingot and their ratio to the ladleanalysis are given in Table 1. The samples were taken by a 10 mm diameterdrill.

This ingot does not exhibit any marked segregation as shown in Table1. Even carbon, phosphorus and sulphur which are usually responsible toa heavy segregation, show a maximum segregation ratio of less than 0.9.

Roque et al.1) showed the relationship between the maximum segrega-tion ratio of carbon and ingot diameters. Comparison of the present in-vestigation and our previous studies2) on medium- and small-sized ingotswith Roque's data is shown in Fig. 11.

The data3) reported in 1926 by the British Committee on Heterogeneityof Steel Ingots (C.H.S.I.) are also included in Fig. 11. The maximumcarbon segregation ratio found by us and C.H.S.I. falls within the rangeobtained by Roque. Further, the diameters of ingots studied by C.H.S.I.were 1,450 mm for a 64-ton ingot and 1,950 mm for a 110-ton ingot. Al-though C.H.S.I.'s ingots were more slender in shape than ours, it shouldbe noted that our data agree well with those of C.H.S.I, as to the rela-tion between the maximum segregation ratio and ingot diameters. Thisreveals that the maximum segregation ratio is more considerably influencedby ingot diameter than by ingot weight.

3.1.4 Inverted-V Segregation

The location and the width of the inverted-V segregation zone in the100-ton ingot are given in Table 2. The location of the zone in the 100-ton ingot as expressed by the ratio of the distance-from-surface to radius

- 12 -

0.25 - 0.37

0.55(max)

0.0050.20 (min)(min)

Fig. 10-1 Distribution of carbon andsulphur in the 100-ton ingot

- 13 -

0006 - 0.010

0013(max)

Si% Mn%

0.35 0.90(max) (max)

0.27 - 0.33 0.76 - 0.85

Fig. 10-2 Distribution of silicon andmanganese in the 100-ton ingot

- 14 -

x 0.72(nin)

P%

0.008 - 0.012

0.01ax)

Fig. 10-3 Distribution of phosphorus in the100-ton ingot

0.0(min)

Table 1 Segregation ratio of a 100 ton ingot

- 16 -

Analysis % Segregation Ratio(

Ladle analysis 0.31

Max. value 0.7780.55

Min. value 0.20 -0.355

Ladle analysis 0.28

Si Max. value 0.35 0.250

Min. value 0.27 -0.036

Ladle analysis 0.76

Mn Max. value 0.90 0.184

Min. value 0.72 -0.053

Ladle analysis 0.009

Max. value 0.017 0.888

Min. value 0.007 -0.222

Ladle analysis 0.007

Max. value 0.013 0.857

Min. value 0.005 -0.286

0.75

4-•

4.1 0.50

0bO0 0.250

0

0

:C.H.S.I.'s Data

o :Our Data

Range byRoque et al.

Fig. 11

0.5 1.0

- 17 -

0

•

ingot diameter (m)

The relation between maximumsegregation ratio of carbonand ingot diameters

100-ton ingot ofpresent investigation

1.5 2.0

•

•

Table 2 Location of the inverted - V segregation zone

in a 100 ton ingot.

Distance from the surface (mm) Zone width (mm)

Top Middle Bottom Top Middle Bottom

584 397 239 178 192 131

(0.54) 1 (0.40) (0.27) (0.16) (0.19) (0.15)

Note: The numerical values in parentheses show the ratio of

distance-from-surface to radius.

- 18 -

is almost the same as in the 30-ton and 50-ton ingots previously examinedat our laboratories.

Macro-structure of the inverted-V segregation zone is shown in Fig.12. A segregation line appearing in Fig. 12 is magnified in Fig. 13. Onecan observe sulphides and small holes existing inside the line. It isevident that the segregation line solidified, because of high solute con-centrations and resulting low melting point, after complete solidificationof its surroundings.

Tashiro and his co1leagues4) suggested that inverted-V segregation

would appear if the solidifying rate in the horizontal direction decreasesto less than 1 mm/min. Whereas Narita et al.5) indicated that it appearsat a position where the derivative of solidifying rate changes its signfrom minus to plus. Experimentally obtained location of inverted-V segre-gation zone, calculated solidifying rate in the horizontal direction andthe position where the rate is at its minimum, are shown in Fig. 14. Fromthe figure, one sees that the location of inverted-V segregation zone cor-responds to the position of the minimum solidifying rate.

Thus, in order for the inverted-V segregation to appear, the solidi-fying rate should be sufficiently small. The travelling of interdendriticsolute-enriched liquid is necessary for the formation of segregation line.As solidification gets slower, the dendrites grow coarse and the co-existence of solid and liquid phases lasts for a longer time, making thetravelling of enriched liquid easier.

As a dendrite is of a plate-likeshape, the enriched liquid is formedin a shape of plane at first. The enriched liquid tends to go afloatowing to its smaller density than that of bulk liquid, therefore it pro-trudes up through a point where the flow resistance caused by above-locat-ing dendrites is the weakest. As the difference in pressure between thepoint and its circumference gets larger, the protruding goes further up-wards. The liquid is considered to continue to travel to be a line re-sulting in the formation of a long segregation streak.

Many of the segregation lines in the top region extend very long andtheir cross sections are point-like. In contrast, the segregation linesin the bottom region stay in an intermittent state and their cross sectionslinear. In the bottom region, as sedimentation of crystals occurs in frontof the dendrites and blocks the floating of the enriched liquid, the growthof segregation streak would be immediately stopped. On the other hand, asthere exists no such obstacle in the top region, the segregation streakwould grow sufficiently.

3.1.5 Distribution of Oxygen and Inclusions

The distribution of oxygen in the ingot is shown in Fig. 15. It isa general characteristic of this ingot that oxygen content is remarkablylow because the molten metal was fully refined by the ASEA-SKF process.Namely, the oxygen content does not exceed 18 ppm on the surface and 6 ppmat the top. The bottom region is found to accompany positive oxygen se-gregation which is peculiar to the sedimental zone. Maximum oxygen con-

- 19 -

ga

position ofminimumsolidificationrate

solidificationrate

< 1 mm/min

1 - 2 rum/min

2 - 3 mm/min

> 3 mm/min

Fig. 14 The location of inverted-Vsegregation and thecalculated solidifciationrate of 100-ton ingot

- 21 -

inverted-Vsegregationzone

•

0 ppm

- 22 -

centration, however, is 73 ppm and the part of such oxygen segregation isto be discarded at the time of forging. Low oxygen content of the melt asindicated by that of surface contributed to so slight a segregation in thebottom part.

Typical types of non-metallic inclusions in this ingot are shown inFig. 16. Sulphides and alumina are found to be the only non-metallic in-clusions, no other oxides being detected. The surface and top are veryclean, with only a dash of sulphides included. Alumina clusters are whatmainly occur in the sedimental zone, with few sulphides being spotted.The distribution of alumina in the ingot is shown in Table 3,

3.2 Pro erties of Forging Made from 100-ton Ingot

Solid forging operation was done by a Push Down type 6,000-ton hydrau-lic press. In order to take test samples from every part of the ingot, nodiscardingwasdone of the top and bottom part unlike the case of ordinaycommercial forgings. Curves for subsequent normalizing and tempering areshown in Fig. 17.

toping.


Macro-structure of the longitudinal section of the forging is shownin Fig. 18. The inverted-V segregation lines run in the forging directionin parallel with the metal flow which corresponds to the shape of theforging. The location of the lines in the forging corresponds to that inthe ingot, i.e., top segregation lines are found to be located more insidefrom the surface of the forging than middle and bottom segregation lines.


Samples for analysis were taken by a 10 mm-diameter drill from the(700 mm 0), middle (950 mm 0) and bottom (700 mm 0) part of the forg-For oxygen analysis, however, block samples were taken.

Carbon, phosphorus and sulphurtop getting also higher towards thecontent is the lowest at the centercenter region of the bottom,

contents are a little higher in thecenter, as tabulated below. Oxygenof the top, and the highest at the

23 -

Fig. 16 Typical non—metallicinclusions in the100—ton ingot

Table 3 Distribution of Alumina Inclusions

(According to JIS G 0555; Microscopic Testing

Method for the Non-Metallic Inclusions in

Steel,x 400, 60 Fields)

Half ofCenter Radius

- 25 -

Surface

Top 0.014 0.011 0.011

Middle 0.004 0.004 0.011

Bottom 0.106 0.045 o

870°C

640°C

Fig. 17 Normalizing and tempering curves of theforging made from a 100-ton ingot

- 26 -

Air cool

Furnace cool

Top

A

2 oo 1114

'

In tf)

A

12,300

fl

Fig. 18 Macro-structure on the longitudinalsections of the forging made from a100-ton ingot.

- 27 -

`€3..

Bottom

3.2.4 Mechanical Properties

Mechanical properties were examined at the following part of the forg-ing, i.e., 700 mm-diameter and 1,140 mm-diameter part each at the top andbottom, and 950 mm-diameter part at the middle. The location of specimensis shown in Fig. 19 and the test results are listed in Table 4.

This forging was made to obtain the required tensile properties spec-ified for SF-50 grade carbon steel forgings of Japanese Industrial Stand-ards (JIS G3201, 1971). Under this specification, the tensile strengthis prescribed to be in the range 50 to 60 kg/mm2, the yield point over25 kg/mm2, and the elongation over 25 % in the longitudinal direction andover 20 % in the transversal direction.

The results for every specimen examined fully satisfy the requirementsfor SF-50. This is shown in Table 4 which indicates every part of theforging exhibits better tensile properties in three direction than thespecification.

Also, rotating bending fatigue tests were performed with specimenstaken in the longitudinal direction at a distance of a quarter of the dia-meter from surface from each 950 mm-diameter part at the top, middle andbottom. The tests were made on 10 mm-diameter standard specimens at 2,800r.p.m. with an Ono-type rotating bending fatigue testing machine. A uni-form value of about 24 kg/mm2 is obtained for each part at the revolutionof 107:

Endurance limit Endurance limit x100U.T.S.

kg/mm2

- 28 -

2 R3

L1 1-L2 L3

T2%my

—71

T1 --a

Fig. 19 Location of specimens for mechanicaltests of the forging made from a 100-ton ingot

- 29 -

Table 4 Results of mechanical test of the forging made

from a 100-ton ingot

- 30 -

Part

Top

Mark of

Test piece

L SurfaceIL2

L3 Center

Tensile Test

ensi e ElongationStren th

Kg/mm2

52.9

52.5

59.6

Yield

Kg/mm2

30.8

30.9

32.3

ReductionArea

%

37

36

31

o

63

61

52

Side RI SurfaceII -

R2700mmO 26.5 52.1 30 55

R3 Center 29.7 53.3 30 50

1 Surface 1 29.8 52.6 36 57

T2 1/4 Dia. 29.5 52.9 31 50

LI Surface 28.3 51.9 36 62

30.0 52.8 36 60

Top L3 Center 29.0 54.8 30 52

Side RI Surface 28.5 52.4 35 61

R21140mm0 27.9 53.1 33 57

R3 Center 27.2 54.8 28 40

T1 Surface 29.8 52.4 35 61

T2 1/4 Dia. 30.8 54.4 27 40

L1 Surface 30.0 53.5 38 64

31.2 55.2 35 61

3 Center 28.5 53.5 36 59

R1 Surface 30.2 52,9 33 50

Middle 30.2 54.5 26 51

950mmf6R3 Center 28.7 53.0 28 40T1 Surface 25.9 53.3 36 60

T2 1/4 Dia. 29.3 54.1 34 49

- 31 -

3.3 Production and Pro erties of Slab-Cooler Shaft Made from a 200-toningot

Press: Push down type 6,000-ton hydraulic press

Forming ratio: Solid forging only 9.7s at minimum dia.

4.7s at maximum dia.

No upset forging was required considering the use of the shaft. Changesof shape in forging process are shown in Fig. 20. The normalizing and temper-ing curves are given in Fig. 21.

3.3.2 Ultrasonic Flaw Inspection

Ultrasonic flaw inspection after normalizing and tempering revealeda

good result, with no flaw wave detected, as shown in Fig. 22.


Test blocks were taken from the marginal portion of the product (cor-responding to the top and bottom of ingot) for macro-etching sulphur printtest on cross section. The results are shown in Figs. 23 and 24. Thougha round segregation zone is observed on the top side, it is much less con-spicuous as compared with the segregation found in the top of 100-ton ingot.This is due partly to the fact that the top and bottom were already dis-carded, and partly to the effects of forging and heat treatment. The bottomside shows no significant segregation zone,


Samples were taken by a 10 mm-diameter drill from the blocks used formacro-structure investigation, and the distribution of elements was ex-amined. The results were as follows:

- 32 -

1 6

2

3

7530

4 20 00 ----- 00

7280

2000

200-ton ingot

6750

19,0

1600

1 5007 130

8

(Bottom cut)

9

10

17 00

12501 00 --

cut off

WOOS

1250

9500

14500

15370

8000 17250

Notes;

1. Forging ratio 2. DimensionsUnit

part A: 6.7Spart B: 4.7SPart C: 8.3S

Fig. 20 Forging diagram of the slab-coolershaft made from a 200-ton ingot.

- 33 -

18300

12000 12400

12000

12000 cut off

890°C

650°C

- 34 -

Air cool

Fig. 21 Normalizing and tempering curvesof the slab-cooler shaft made froma 200-ton ingot.

Furnace cool

1 2

3

4

5

6

7

8

9

10

11

12

13

14 15 16

17 18

12

3

4

5

6

7

kiki

likili

dia

8

9

10

11

12

13

14

15

&la

ud Li

Li_

Li

LI

LIL

LNote: Detector; USK-5M

Frequency;

1 MHz

Sensitivity; B1=100%

Probe dia.; 24mm

Couplant; Water glass

Fig. 22 Results of ultrasonic

flaw inspection

of the

slab-cooler

shaft made from a 200-ton ingot

Top Bottom

Fig, 23 Macro—structure on the transversalsections of the slab—cooler shaftmade from a 200—ton ingot.

Top Bottom

Fig. 24 Sulphur prints on the transversalsections of the slab—cooler shaftmade from a 200—ton ingot.

-36-

The above results indicate that carbon, manganese, phosphorus andsulphur are higher at the top than at the bottom, and also that they tendto be higher towards the center at the top but lower towards the centerat the bottom.

3.3.5 Mechanical Properties

Specimens were taken, from the same blocks, for mechanical testswhich were conducted in both longitudinal and transversal directions ateach position from surface to the center, as shown in Fig. 25.

The test results were listed in Table 5. Although there is some dif-ference in the elongation and reduction of area of the bottom side speci-mens between the longitudinal and transversal direction, the over-allresults are almost uniform with little variations.

- 37 -

R1

..... -... 4.1

2 .....- .-..

Fig. 25 Location of specimens formechanical tests of the slab-cooler shaft made from a200-ton ingot

- 38 -

R2R3 center

CA

Table 5 Results of mechanical

test of the slab-cooler

shaft made from a 200 ton ingot.

Tensile Test

Charpy Impact TesHardness Test

U-notch. 2mm. 20°

Part

Mark of

Test piece

see fig. 25

Yield

Point 2

Tensile

Strength

2

ElongationReduction

of Area

Brinell

Hardness

Kg/mm

Kg/mm

Kgm/cm2

Surface

24

44

41

67

13.3

116

L I24

45

36

63

13.7

116

L 223

45

37

61

11.4

118

Top

L 3 Center

24

48

30

58

10.2

126

RI Surface

24

44

36

63

12.4

118 .

R2

22

44

30

61

11.2

123

R 3 Center

22

46

36

63

11.3

126

L, Surface

26

24

45

45

40

39

62

63

12.2

13.4

123

118

23

43

40

62

13.0

116

BottomL 3 Center

24

44

40

65

11.8

118

R1 Surface

24

44

36

56

11.5

116

R 223

43

38

54

10.2

114

R 3 Center

23

44

37

53

9.0

116

4. CONCLUSION

Investigations were made on the characteristics of solidified struc-ture of a 100-ton ingot of LD-converter steel vacuum-degassed by the ASEA-SKF process, on the properties of the forging made from a 100-ton ingotproduced by the same process and on the production and properties of acommercial forging made from a 200-ton ingot of RH-degassed steel. Theresults are summarized as follows:

(1) The steel ingot showed a sound macro-structure with very slightV-segregation except for as much inverted-V segregation as normally ob-served in steel ingots. Also the solute segregation on the axis was notso marked. It has been clarified that the inverted-V segregation appearswhere the solidifying rate is at its minimum. Reducing oxygen content ofmolten steel by the ASEA-SKF treatment made it possibleto obtain a cleansteel ingot with little accumulation of inclusions in sedimental zone,that is, oxygen content of the ingot being 18 ppm at the surface and itsmaximum value being 73 ppm at the bottom region.

(2) The forging from a 100-ton ingot was found to have a sound macro-structure with slight segregation and to possess sufficient mechanical pro-perties in every part of the forging.

(3) The ultrasonic flaw inspection indicated no defects in the slab-cooler shaft for a slabbing mill plant made from a 200-ton ingot. Examina-tions made on test blocks from its top and bottom ends showed a soundmacro-structure with little segregation and uniform distribution of mechani-cal properties.

(4) The Mizushima Works manufactures various kinds of forgings fromelectric furnace steel ingots as well as from LD-converter steel ingots.This paper revealed that ingots from degassed LD-steel satisfactorily com-pete with those from electric furnace steel in the qualities of largeforgings.

At present, high quality carbon steel forgings and such alloy steelforgings as Ni-Cr-Mo steel, Cr-Mo steel and high carbon chromium steel aremanufactured from LD-steel ingots at the Mizushima Works of KAWASAKI STEELCORPORATION.

- 40 -

References

1) Roque et al.: Rev. Met., 57 (1960), P.1091

2) Ooi: unpublished

3) Iron and Steel Inst.: First Report on the Heterogeneity ofSteel Ingots (1926), P.97

4) Tashiro et al.: Tetsu to Hagane, 51 (1965), P.1893

5) Narita et al.: Tetsu to Hagane, 56 (1970), P.1323

- 41 -

ABSTRACT

AN APPROACH TO THE PRODUCTION OF LARGE FORGING INGOTSBY ELECTROSLAG REFINING

P.J. Wooding and J.M. Mowat

It is recognized that many ingot problems associated with theconventional production of large tonnage forgings would be elim-inated if electroslag refining and directional solidificationpractices presently used with smaller sized specialty steel ingotscould be applied with equal success to the manufacture of largeingots.

In spite of the important innovations introduced in steelmaking within the last twenty years, large ingots manufactured bypresent conventional practices are technically deficient withrespect to compositional uniformity, segregation, ingot sound-ness, cleanliness, and yield.

In response to the mentioned problems and the continuingrequirement for hydrogen control, and in recognition of theprojected needs of the heavy forging industry, a controlledatmosphere electroslag refining system concept has evolved whichwould provide the steel maker with a means to exercise maximumcontrol over the steel refining aspect of the ESR process, and,in addition, provide the means for allowing the ingot solidifi-cation process to proceed in a smooth, uninterrupted manner undercontrolled thermal conditions. Both technical requirements mustbe met to ensure large forged product reliability, uniformity,and reproducibility at high quality levels.

Projected operational costs of large ESR furnaces have beenestimated. These costs must be weighed against potential savingsexpected to Le realized in the areas of product yield gain, forgeprocess simplification, and higher forging design confidencelevels derived from improvements in the level and uniformity ofmechanical properties of ESR material.

While the economics of the more intangible areas of productvalue improvement for large forgings remain to be quantified,based on the success of smaller ESR ingot applications, itappears that the time has arrived for the incorporation ofelectroslag remelting into the manufacturing operations of thesteel making and heavy forging industry.

-

INTRODUCTION

Within the last few years, we have witnessed the acceleratedgrowth of the ESR process to the point where it has becomeaccepted as a means of producing high quality steels in ingotsizes ranging up to 40 inches in diameter and slabs up to30 inches by 60 inches in section. Few people question theimprovements in cleanliness, ingot structure, mechanicalproperties, uniformity, workability, and product yield obtainedon production ingots in this size range.

The question asked by those steel makers skilled in the artof producing large forgings and those responsible for the designand construction of the end product, such as power generationsystems, is whether the additional advantages of ESR can beobtained economically and reliably in the much larger ingot sizeswhich are required for current and projected applications.

Enthusiasm for the potential of electroslag refining, then,has been tempered by reservations which may be summarized asfollows:

1. What sort of ingot solidification pattern may oneexpect in an ingot 3 meters in diameter and up to6 meters high? Can reasonable melt rates beobtained consistent with an acceptable ingot structure?

2. Can the hydrogen content in the ESR ingot be reliablycontrolled within tolerable limits? Can hydrogenpick up be prevented?

3. Can thermal cracking of the ingot during ESR pro-cessing be prevented? What ingot temperatures maybe expected?

4. What metallurgical and operational considerationsmust be given to the design and operation of theESR furnace system and to the manufacturer ofelectrodes?

5. In a steel shop producing ingots up to 300 tons orlarger, refining equipment must be reliable andcomparatively simple to operate. What are theessential features of a large ESR furnace necessaryto meet those requirements?

6. Economics; what does it cost to operate a largeESR furnace? Can those costs be offset withsavings resulting from improved yield and productvalue?

-2-

An ESR furnace system designed to meet the steel makersrequirements for large ingot production is proposed.

OTHER LARGE ESR FURNACE APPROACHES

According to the literature, at the present time, there aretwo other ESR furnace system approaches either in use or in pilotoperation aimed at producing large ingots. The system placed inoperation at Roechling for producing ingots 1700 mm diameterto a projected 2500 mm uses four separate electrodes each suspendedand driven independently with provision for in-process electrodechange.(1) The electrodes are relatively short in length andtherefore a series of electrode changes are required during thecourse of manufacturing a large ESR ingot. In order to minimizeelectrode change time and to have the crucible system remain ata convenient fixed level, an ingot withdrawal device is used. Anattractive feature of such a system is that short electrodes ofmodest diameter (up to 36 inches) can be routinely produced inconventionalfashion with minimum top to bottom segregation.

Upon examining such an approach to making large ingots, wefind several technical and operational deficiencies. Electrodechanging is risky both metallurgically and operationally. Ifthe change sequence is conducted with precision and reproducibilityit may be a technically acceptable procedure, although this hasyet to be proven. Electrode change times are critical,(2)

Electrode changing four or more times per heat complicatesthe melting procedure and can only detract from the overalloperational reliability of the total furnace system. The use offour or more conventional round electrodes results in a poor fillratio which affects melting efficiency.(3) In addition, non-uniform heating of the periphery of the molten metal pool willalso occur and may be reflected in the ingot structure and ingotsurface behavior.

The large ESR system proposed by the USSR for rotor forgingsspecifies the use of a cluster of six or seven electrodes withno provision for electrode change. It, too, requires relativemovement between ingot and mold and in their case the mold movesupward as the ingot grows.(4) The USSR is operating a pilotscale version of such a furnace. A novel aspect of the USSRapproach is that there is no movement of either the ingot orelectrode assembly. The group of electrodes are equivalent intotal cross-sectional area to the forming ingot. This isaccomplished with use of a "tulip shaped" mold having an expandedupper section. The upper section contains the molten slag whilethe ingot forms in the reduced lower section of the mold with theslag-metal interface maintained in the reduced section.

-3-

Each of the two basic approaches discussed utilize relativemovementof the mold and forming ingot. Each furnace type operatesunder atmospheric conditions.

300-TON ESR FURNACE, GENERAL DESCRIPTION

A schematic of the proposed ESR furnace appears in Figure 1.The melting chamber is assembled from a series of jacketed, water-cooled cylindrical sections. These sections are arranged topermit them to easily separate from each other for furnace loadingand unloading. Also the upper section below the electrodesuspension system is designed to be completely removed duringthe melting of the relatively short ingots. The chamber assemblyis made using standard vacuum sealing techniques to preventatmospheric contamination during melting. Ports are providedfor the continuous flow of air from the dry air supply system,and for the exhaust duct to the fume collection system.

The electrical connections between the power supply systemand furnace are made so that the melting chamber itself acts asthe return current conductor. A completely coaxial conductorarrangement is thus achieved, eliminating unbalanced magneticfields in the area of the melt zone.

The electrode stub clamp connected to the electrode supportcylinder rod serves as a current conductor as well as supportingthe weight of the electrode. Each electrode assembly is weldedto a reuseable stub adapter designed for simple connection to theclamp attached to the lower end of the electrode suspensionsystem. Current to the electrode clamp is transmitted fromsymmetrically arranged water-cooled power cables thereby ensuringmaximum furnace power factor and magnetic field symmetry.

The furnace is shown in Figure 1 as having a fixed base.The base assembly could also be fitted to an ingot car shouldit be necessary for material handling reasons, such as craneavailability, to separate the melting station and the loading-unloading station.

The basepiate cooling system circuits are designed to permittemperature regulation in the region of 3000 centigrade. Thisfeature is aimed at reducing the potential hazard of ingotcracking by maintaining the ingot base above the transitiontemperature.

The furnace will also be equipped with two make-up slag andalloy feeding systems.

A multiple stage air drying system will be provided to en-sure continuous supply of dry air through the furnace chamber.This is to prevent the atmosphere from serving as a major sourceof hydrogen contamination during the electroslag refining process.

-4-

The melting controls for such a furnace will be fullyautomated in order to ensure process consistency and to guardagainst operator error. In all probability the melt would beprogrammed for a constant pool volume up to the hot-toppingcycle. An appropriate kilowatt profile would be fed into theweight based graphic controller in charge of the refining process.

Particular attention has been given to the crucible designbecause of its enormous size. A fixed crucible design is plannedin contrast with the shorter collar molds in use or proposedbythose using the ingot withdrawal or moving mold technique. Wehave developed a concept of fabricating electroslag cruciblesfrom individual hollow water-cooled, copper modules arranged ina vertical palisade fashion with a supporting frame. The modularcrucible is illustrated in Figure 2. A crucible incorporatingthese features but designed for use with smaller ingots has seenmonths of satisfactory production service. It required nomaintenance and no noticeable deformation was experienced.

The primary purpose of developing a modular crucible is toensure its reliable operation over extended periods of time andto simplify the construction and transport of such a largeassembly. Also, in the event a section of the crucible shouldbe damaged, it can be easily removed for repair or replacement.

In order to guard against the possibility of hydrogen pick-up due to moisture in the slag and to ensure the ease andconsistency of starting the use of molten slag starts is advo-cated. Molten slag will be provided to the ESR furnace from a3-phase slag melting furnace having a capacity of approximately5,000 Kg.

The furnace will be powered by several single phase ACwater-cooled, saturable controlled power supplies with balancedsupply having a total output in excess of 100,000 amps at 40electrode volts.

COMPOSITE ELECTRODES

The manufacture of electrodes for small ESR furnaces israther straightforward but becomes a more pressing issue forESR furnaces having ingot capacities in excess of 50 tons be-cause of electric furnace size limitations relative to therequired final ESR ingot size and concern over electrodecompositional segregation.

Since electrodes are merely an intermediate product, araw material for the ESR furnace, they do not have to be manu-factured to the stringent requirements of forging ingots,particularly in terms of soundness and cleanliness. However,compositional segregation in the vertical dimension of the

-5-

electrode is of concern because those differences will carry overto the final ingot. On the other hand, transverse segregation,should it exist, would be eliminated during remelting. A solutionis proposed to the electrode manufacturing problem which appearsto satisfy the complex requirements of large ingot ESR processing.

As shown in Figure 3, it is proposed that the electrode beproduced as a composite structure consisting of a group ofindividually cast sectors, the number to depend upon the requiredfinal ingot diameter, electric furnace capacity, etc. Theelectrode sectors would be joined to a common holder and supportedby a hydraulic ram assembly.

One of the major advantages of the single composite electrodeapproach is the elimination of the need for electrode changeduring processing with its attendant operational complexity andmetallurgical risks. It also permits the furnace structure to begreatly simplified and therefore ifs cost reduced.

Since each sector electrode tends to be long relative to itscross-sectional area conventional vertical casting would bedifficult. It is therefore proposed that the electrode sectorsbe cast horizontally in the manner shown in Figure 4.

There are numerous advantages to the horizontal castingmethod. There would be no compositional differences throughoutthe length of the electrode, thus assuring compositional uniformitythroughout the length of the remelted ingot. Any chemistry varia-tion which may exist in the transverse dimension of the electrodesector, or among individual sectors made from different electricfurnace heats, is eliminated upon electroslag remelting. Otheradvantages include the possibility of casting electrode sectorsto lengths exceeding that which is practical for vertically castshapes and the ability to shorten electrode lengths by the simpleinsertion of a dummy block into one end of the mold cavity or re-moving one or more sectional mold pieces.

Horizontally cast electrodes up to 170 inches in lengthhaving a cross-section of 8 inches by 22 inches have been pro-duced successfully in sectional steel molds of similar construc-tion to that shown in Figure 4 and no serious problems areenvisioned in the production of longer pieces. No mold erosionwas experienced and the relatively thin electrode did not vary instraightness more than 1/2" over its entire length.

It is recommended that the electrodes be produced fromdegassed steel in order to begin with the lowest possible hydrogencontent. In the event one feels it necessary to isolate theelectrode from the atmosphere during the teaming operation, thiscan be accomplished with a simple mold enclosure.

-6-

Hot-top losses with conventional vertically cast electrodesrange up to 15%. There are no such losses associated withhorizontal casting resulting in substantial product yield im-provements.

INGOT SOLIDIFICATION

The prime technical concern of those responsible for producinglarge forging ingots can be summed up in two words; "ingotsolidification". Can reasonable melt rates be obtained consistentwith an acceptable ingot structure? Can ingot cracking beavoided?

Since 3 meter ingots have not yet been producedby the ESRprocess we have projected answers to such important questionsbased upon extrapolation of smaller ingot experience. The finitedifference computer technique, which has proven to be accurateand reliable when applied to studies involving ingots up to 40"diameter, was usedAb)

As can be seen in Figure 5, if the melt rate, in this case8,000 lbs/hr., is maintained throughout the length of the ingotthe molten pool will continue to increase in depth reachingapproximately 1 meter after 6 meters of ingot has formed. Tosimplify our initial computer analysis, using the finite differ-ence technique, we have chosen to allow the melt rate to remainconstant. In actual practice one would select what had previouslybeen determined as the optimum pool depth and profile for agiven alloy grade and size and adjust the melt rate on an auto-matic program basis to maintain this profile.

As the melt cycle nears completion (the hot-top cycle) themelt rate is further reduced to allow the molten metal poolgradually to reduce in volume while still continuing to freezeprogressively, thus assuring minimum ingot structural disturbanceand hot-top loss.

Figure 7 details the computed liquidus and solidus isothermsof a Ni-Mo-V rotor steel ingot at the 6,000 mm ingot growthlevel while melting at a rate of 8,000 lbs/hr. The depth of the"mushy zone" at the ingot center is shown to be approximately125 mm. The depth of the liquidus-solidus zone increases asthe ingot grows, assuming a fixed melt rate.

No one really knows at this time what would be consideredthe ideal pool profile and liquidus and solidus isotherms for a3,000 mm diameter ingot. It is a function of the alloy grade,its segregation tendencies and the segregation tolerance levelestablished for that particular grade for a given application.

-7-

While quality standards, based upon ingot-billet structureand mechanical properties, have been highly refined for materialsmelted by the vacuum arc and electroslag process for use in jetengine applications, new standards would have to be developed forvery large ESR ingot products.

Other factors constant, the pool shape, its depth, and thedistance between the liquidus and solidus isotherms is influencedby the ratio of electrode to ingot cross-sectional area (fillratio), a factor which strengthens the argument for the proposedhigh fill ratio composite electrode. The pool tends to becomeshallow as the fill ratio increases, explained by the fact thataverage temperature gradient across the top surface of themolten metal pool is reduced.

We have also learned, with small ingot work(3) that internalingot quality can be adversely affected by excessive stirring ofthe metal bath due to the presence of furnace magnetic fieldinbalance. The furnace system as proposed eliminates that possi-bility because the current is returned from the ingot to thepower supply in a completely symmetrical fashion.

The computed ingot surface heat distribution at varioustimes during the growth of the ingot is shown in Figure 6. Whenthe base is maintained at 300°C no portion of the 6 meter tallingot will be less than this temperature thus protecting the ingotfrom thermal cracking.

It may be argued that concern over details which influenceseemingly minor variables of ingot structure is inappropriateconsidering that a large ESR ingot produced to rather crudestandards will be far superior to currently produced static ingots.It is the view of the authors that "minor" process improvementsinvariably follow quantum jumps as far as ingot making technologyis concerned and that the design and operation of a large furnace,which would be expected to set new standards for large forgingingots and function for a number of years, should incorporate thelatest consumable melting technology available.

One thing appears clear; based on work done with ESR ingotsup to 40" diameter round and 30" x 60" slabs, the improvement instructure of large ESR ingots over the conventional ones incurrent use should be dramatic.

HYDROGEN AND CONTROLLED ATMOSPHERE PROCESSING

It is generally accepted that the control of hydrogen in anESR ingot is dependent on several factors, including:(4)

-8-

ECONOMICS

1. The partial pressure of water vapor in the atmos-phere above the melt and the atmospheric circulationrate.

2. Quantity of hydrogen introduced into the system bythe electrode.

3. Starting technique, flux type and bulk.

There is reference in the literature to a technique understudy in the USSR in which some hydrogen is removed during ESRmelting by the injection of "gas mixtures" into the molten slagand metal pool.(6) To date, no specific information isavailable either regarding working details or the effectivenessand reliability of such an approach.

The authors feel that the most effective and reliable meansof controlling hydrogen at acceptable levels is to conduct themelting operation in a closed chamber allowing only dry air witha controlled moisture content to pass through the system. Al-ternatively, such a system can also be operated with an inert gasatmosphere should that be desired.

With the closed furnace system, one has the option of select-ing the atmosphere and slag composition which will best fit theneeds of the alloy type being processed, and has available themeans to exercise positive control over processing conditionsduring the melt.

The quantity of hydrogen introduced to the system by theelectrodes is best controlled by the use of vacuum degassed steel.

The use of molten slag starts is advocated because of itsproven effectiveness with smaller ingot sizes. While there re-mains the question as to whether or not slag melting must alsobe conducted under dry atmosphere conditions, the slag pre-melting furnace is, in fact, enclosed for the purpose of fumecollection and could, if necessary, be operated under controlledatmosphere conditions using the same air drying system thatserves the ESR furnace.

While we are not in the position to indicate in specificterms the total cost improvement potential which we feel probablyexists with large ESR ingots, based upon considerable experiencewith smaller furnaces, we can state qualitatively that signifi-cantly improved product reliability, improved product yield, andthe potential for producing large forgings from a given ingotsize, can only help to moderate process costs to the forgingmanufacturer and give the user industry a product of significantlyimproved value.

-9°

However, the direct cost of production melting of largeESR ingots by electroslag refining can be estimated.

Shown in Table I are the manufacturing costs of a largeESR furnace operated at a conservative output of 12,000 ton/year.The product mix consists of ingots ranging in size from 1.5 to3.0 meters diameter, having an overall average melt rate ofapproximately 5,000 lbs/hour. A five day schedule with an 80%furnace melting factor was assumed. Smaller ESR furnacesoperate at 90% to 95% efficiency levels because of shorterturn-around times.

Considering the net economic impact of producing largeforging ingots by ESR, one would expect that the improved valveof the final forged product plus the more visible areas ofpotential economics would offset the added cost of ESR process-ing. However, it remains in the hands of the steel maker andend-user to make that final determination.

The authors feel that in the face of increasing demands onquality and product improvement,(7) that the use of large steelingots produced by the ESR process for critical applicationssuch as power generating equipment, will become justifiableeconomically.

SUMMARY

While steel making practices have continued to improve, theinherent structural limitations of conventional ingots imposedby mass solidification remains a formidable obstacle to furthersignificant progress in large ingot making technology.

Electroslag refining has proven itsvalue as a refiningand solidification process with smaller sizes, but many steelmakers have expressed concern about the adaptability of thisprocess to large forging ingots.

The authors have presented an ESR furnace concept designedfor operational simplicity and reliability and to provide thesteel maker with the means to product large diameter ingots underclosely controlled conditions, offering solutions to the keyproblems of: control of hydrogen; ingot solidification; ingotcracking and compositional control.

-10-

REFERENCES

1. H. Hinze, H. Schedig, A. Choudhury and R. Jauch; Productionof Heavy Ingots - A New Era of ESR. Proc. 3rd Int. Sym.on ESR and Other Special Melting Technology, Mellon Inst.,June 1971.

2. R.O. Jackson, J. Luchok, A. Mitchell; An Examination ofElectrode-Change Practice in Electroslag Melting. 1972Vacuum Metallurgy Conf., Pittsburgh. To be published.

3. R.J. Roberts; Techniques for Maximum Melt Rate and MinimumPower Consumption in Electroslag Melting, Proc., 2nd Int.Conf. on ESR, Mellon Inst., Sept. 1969.

4. T. Bagshaw; The Behaviour of Hydrogen in ElectroslagRemelting. Proc., 3rd Int. Sym. on ESR and Other SpecialMelting Technology, Mellon Inst., June 1971.

5. Ingot Thermal Profile Computer Analysis by the FiniteDifference Technique, Conducted for Consarc by Prof. A.Mitchell of the University of British Columbia. Unpublished.

6. B.E. Paton, B.I. Medovar, G.A. Bojko, V.M. Baglaj,L.V. Chekotilo, O.P. Bondarenko; Electroslag Remelting ofLarge-Tonnage Ingots and Shaped Castings. Proc., 3rd Int.Conf. on ESR and Other Special Melting Technology, MellonInst., June 1971.

7. D.L. Newhouse and D.R. Forest; Meeting Requirements forLarger Generator Rotors---A Metallurgical Challenge. Int.Forgemasters, Terni, 1970.

TABLE I: Estimated Cost of Operating a 300-Ton Capacity ESR Furnaceat an Output of 12,000 Tons/Year; 5-Day Schedule.

Materials $330,000.00

Direct Labor 95,000.00

Manufacturing Overhead 390,000.00

Total Cost of Sales $795,000.00

Manufacturing Cost Per Ton $66.00/Ton

General and Administrative $ 44,000.00

Total with G&A $839,000.00

Total Costs Per TonIncluding G&A $70.00/Ton

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ABSTRACT

Economical aspects of forging manufacture by conventional production

and by electroslag remelting

By

Dr. -Ing. habil. M. Wahlster, Leybold-Heraeus GmbH, Hanau,

Dr. mont. E. Zimmermann, Rheinstahl Htittenwerke AG, Hattingen

In this paper an attempt is made to compile suitable data for

a comparison of manufacturing costs of large forgings originat-

ing from ESR and conventionally cast ingots, respectively.

According to capacity and utilization of one ESR unit, remelting

costs can vary between 260 and 600 DM/ton. Despite a mostly

much better metallurgical output with ESR ingots, costs of their

generation may increase up to twice those of conventionally

cast steel. Only a reduction of several cost elements during

the manufacture of a large forging may justify the application

of the ESR process economically.

By means of mathematical-graphical methods (nomograms) an

attempt is made to determine the influence of remelting costs,

crude steel and scrap costs, ingot yield, forging, heat treatment

and machining costs as well as rejection rate on total costs.

to quality and size, it is impossible to

one to consider, where an application

large forgings is justified.

of particular forging products relative

standardize

of these nomograms

our calcul-

of ESR material for

enables

Economical aspects of forging manufacture by conventional production

and by electroslag remelting

By

Dr. -Ing. habil. M. Wahlster, Leybold-Heraeus GmbH, Hanau,

Dr. mont, E. Zimmermann, Rheinstahl Htittenwerke AG, Hattingen

SYNOPSIS

Technical advantage through Electro Slag Remelting. Steps to

investigate the remelting costs of ESR units of different sizes and

capacity. Possibilities to influence the production costs of ESR

steel as compared to conventionally cast ingots. Separation into

cost elements, Numerical evaluation of most important cost

factors like ingot yield, raw ingot prices and scrap prices, forging

costs, heat treatment and machining costs as well as rejection

rate. Development of straight-line charts (nomograms) in order

to determine the net manufacturing cost. Conclusions.

INTRODUCTION

Steel quality is not only influenced by metallurgical reactions, e. g.

attaining low phosphorus and sulphur values or low hydrogen cont-

ents by vacuum treatment, but is also very strongly determined

by physical reactions during ingot solidification. In conventional

steel production possibilities to influence the crystallization process

in order to prevent undesired reactions like ingot segregation are

limited. The well-known large-scale remelting processes introduced

in the past 20 years (like vacuum arc, electron beam and electroslag

remelting) on the other hand permit a solidification of liquid steel

by means of a controlled crystallization and are thus reducing the

extent of typical internal defects to a permissible degree.

Because of its flexibility with regard to remelting various steel

alloys into various ingot shapes and sizes and also because of its

simple construction, the ESR process does not only continue to•

supersede the dominating vacuum arc furnace. It also enters the

field of large ingot cross-sections which are of great importance for

the production of large forgings. In the near future it should be

possible to produce all kinds of ingot sizes by this method. Evidence

for this development is the erection of a 4-electrode ESR unit at

ROchling-Burbach Steelworks, Volklingen, Western Germany, in

1971 (Fig. 1).

Typical changes of steel remelted by ESR in comparison with

conventionally cast steel are observed in density and porosity,

improvements of the oxide and sulphur cleanliness, improved

toughnes s and uniformity of all measurable prope r tie s over the

ingot cros s -section 1) (Fig. 2). But as there are numerous

possible applications for ESR material, it is obvious that not

all improvements caused by remelting are economical, i. e.

they cannot be utilized in order to obtain a higher market price

of the final product. Only in those case s, where the consumer

specifies this production proce ss or where certain properties

cannot be realized by conventional production methods, the applic -

ation of ESR is self-evident. For all othe r products it is necessary

Ingot conditionsurfarrepraramty and densityyield(output)

Chemical compositionbasic m(1talshydrogenOxygensulphurtracers

Cleanliness•a:rose/4)m

macroscopic

Ingot structuretertot segregattonscrystal segregations

Mechanical propertiestensile strengthyteld pointiMpad valueisotropy

remelting defectorag.treckles

worsened imprewal mom

Ufwhonged :reproved

—WY4---+4-4li

4 4

4

1972

Influence of

Electro Slag

Remelting

on Properties

(afterM. Wahlster)

Fig, 1

Fig.

to calculate the manufacturing net costs with the aim of finding

at least a cost balance between remelted steel and conventionallyproduced steel. As Figure 3 shows for the manufacture of large

forgings, one has to keep in mind that only certain cost elements

can be influenced or reduced by remelting. This is valid with

regard to material costs, forging costs including the necessaryheating, as well as the rejection level and all expenses to prevent

rejects. It is evident that particular costs can strongly vary accord-ing to type, size and grade of a large forging.

Mechanical treatment

1972

Material costs

Heating up andslow cooling

Reduction byforging

Relative portion of production costs

Production costs of forging pieces(after M. Wahlster)

[-Costs affected by ESRCosts not affected

work's kost

3

Production risk(rejection rate)

age-hardening

Fig. 3

With this paper we make an attempt, to bring more transparency

into the very complex cost calculations when comparing convent-

ionally cast ingots with ESR ingots. This should help to minimizethe economical risk of the installation and operation of an ESR unit.

Com arison between lar e for in s roduced b conventionall cast

in ots and ESR in ots, res ectivel .

Several factors are important when considering the economics of

an application of ESR ingots as compared with conventionally castingots:

In principle remelting leads to higher material costs. Figures 4 to 6

show specific data and cost factors which influence remelting costs,

they also show variable as well as total remelting costs as they depend

upon the capacity of a unit (1 to 300 tons). All data take into consid-eration the newest technical level and are valid for continuous oper-

ation with an 80 % time utilization of the unit (6, 820 hours/year).

4000t—g I

2000 -w

aa 0FP, le

b)ca xa

-p0

0,7c I

ts.. 2 0 5

0.71

z

0

fixed mould

Omen

11 men

1972 Specific Data of ESR Units

Movable mould

5 Ill 20 50 700 200 400Size of unit I in tons)

A

16 men Fig. 4

Ingot diameters increase from 350 mm for 1 ton ingots to 3, 000 mmfor 300 ton ingots (s. Fig. 4 a). While the investment costs ofESR units (Fig. 4 b) are relatively accurately determinable, thereare other considerable costs in connection with supplementaryinvestments for construction of buildings, materials handling devices,high-voltage connections etc. , which depend strongly upon geograph-ical location and thus cannot be standardized. This description,therefore, contains 50 % extra cost above the normal costs fora unit. Investment costs determined this way are below 1 million1DM for small units up to 10 tons. They increase in case of large -scale units up to appr. 5 million DM for a capacity of 50 tonsand attain nearly 14 million 1DM for 300 ton units.

Labor costs (Fig. 4 c) are based upon the minimum requiredpersonnel for units integrated into a steel plant. Small unitsrequire 2 men per shift, large-scale units require 4 men per shift.The total costs of 40, 000 DM per man and year result from a4 shift crew, the necessary supervision included. As the operationof an ESR unit requires well trained personnel, it is not possibleto vary the number of people according to changes in the util-ization of the unit. We, therefore, added personnel costs tofixed costs.

The normal unit costs (Fig. 4 d) are determined according tousual accounting methods; they amount to 36 % of total costs perunit and year.

As Figure 5 shows, power consumption increases with risingcapacity from nearly 600 kW1ilt in case of small units to nearly

1, 800 kWh/t for large-scale units. Data contained in this description

mostly results from operating units with single and multiple electrodes.

"1800

.F?1600 4/5r•

C..' We 1e -„-; • 3/60 • 4.-------- - In1200 . • *---r* 3/50•wog . 1/50 :1/50 ......---C----7/50 .3750 81/50E -"---tz, ±800 v. frequency in HzQ. --600 1/50

3/50' • X- Number ol electrodesr1

400

20

16

12

8

at 80% utilization 6,52014ot 50 %utilization /4,250h1

80% 20t/h

•450%

1.0t/h-

- — — 0.5 t/h

01 5 10 20 50 100 200 GOO

Size of unit lin tons/

1972 Power consumption and annual production of AESR units

Fig. 5

The annual production of ESR units increases with rising unit size

(Fig. 5 b), but depends considerably upon utilization, of which the

highest level very seldom surpasses 80 %. Taking into consideration

a utilization of 80 % and normal remelt rates, the annual production

of a 1 ton unit attains a maximum of 2, 000 tons. For 10 ton units

it rises to a maximum of 5, 800 tons, for 100 ton units to 12, 000

tons and for 300 ton units it reaches 18, 500 tons.

The variable costs (Fig. 6 a) show a linear increase from nearly

130 DM per ton for small units up to 180 DM per ton for large-

scale units. Variations in this curve result from assumptions of

prices for electricity of O. 07 DM/kWh for small units and O. 05 DM/kWh

for large-scale units. Hence, total remelting costs (Fig. 6 b) mostly

depend upon unit utilization and average remelt rate. Based on

assumed remelt rates of 250 kg up to 300 kg per hour for small

units and 2, 500 to 3, 000 kg per hour for large-scale units, we

calculate total remelting costs as shown in table 1.

Table 1: Total remelting costs of ESR units

Ingot weight Ingot diameter Total remelting costs in DM/tutilization

111m 80 % 50 %

up to 2 t up to 500 appr. 300 400

5 to 40 t 600 to 1,200 I I 275 340

50 to 100 t 1,300 to 1,800 I I 350 470

200 to 300 t 2, 000 to 3, 000 I I 400 570

7— 6 4>Price of electribity lin Opl/kWhl

5

I 80% utilization

I 50 % utilization600 i

I /%•.,

\ IIf,: 7,..i. 500 \ !

tFz"--\ ! 0250t,, 2

\ I

ct = ---- ZOO05t/h

-44O •-• c-6LS 300

200 —

1972 Proportional and total remelting costs of AESR units

Fig. 7 shows the relationship between useful ingot weight andnominal ingot weight. In case of conventionally cast ingots theuseful ingot weight first increases in the range of small ingotweights. Then it remains rather constant in the range of 5 to30 tons. But with rising nominal ingot weight the output decreasesconsiderably and attains 59 to 69 % in case of 300 ton ingots.The increasing scatter band is caused by the type of finishedproducts, e. g. solid or hollow body.

I00

(SRMg° s95

90 4 Conventionally cast ingots

85

804-4

75

70

1972

1 2 5 10 20 50 100 200 400Size of unit lin tonsI

601 2 3 5 10 20 30 50

Nominal raw ingot weight in tons

65

'Iunder the premises that maximum inrt length is remelted

/

. / .......i----/

// .....---- _

/ 1,0t/h ".1.5t/h 1" ----/ ..//

/ 2,0t/h ----

100

Output of conventionally cast ingotsand ESR i

700 300

big

small

A

Fig. 6

Fig. 7

On the other hand ESR ingots of up to 1, 200 mm diameter yield

a usable ingot weight of nearly 95 %. This usable weight decreases

with rising ingot diameter to a value of 80 - 90 % with 3, 000 mm

ingots (weight 300 tons).

The increasing scatter in the yield from ESR ingots does not only

depend upon the type of finished product, but also on the ratio

"length to diameter" (LID) of the ingot. It is evident from this

description that an economical application of ESR units for the

manufacture of large forging ingots mainly depends upon the increase

in yield over that from conventionally cast ingots.

As Fig. 3 showed, total production costs are composed of a variety

of cost elements, which influence one another and can thus hardly

be determined. Consequently an exact cost comparison between

conventionally cast and ESR ingots is only possible by solving

complicated mathematical functions, which represents unjustifiably

high expenditure. A much simpler solution is attainable through

the preparation of nomograms which lead to similar and sufficient

exact results. The following cost factors are considered in their

influence on total costs.

Rernelting costs,Total yield (= metallurgical yield x yield during

production)

Crude steel prices and scrap prices,

Forging costs,Costs for machining and heat treatment,

Rejection rate (AT-rate)

Note: Definitions and abbreviations applied in the nomograms are

explained and listed in a separate table.

A nomogram is shown in Fig. 8. Here, we can determine the increase

or decrease in costs per ton of usable material, if ESR steel is

applied in contrary to conventionally cast steel. Factors

Output X ol a conventional ingot Scrap value K5 14nOM/t1

05 80 75 15 65 60 55 1000 800 600 400 2:1 0

i WOO

71X 'WOO €I ,/

800

',so"600 1

, - 400 IIk '8ft

7a1 51 I

1/EX

'iomrt,34

800 400 '0 ran -800lower costs MK/ higher costs

I I2C6 AK - !Kw-Xs)

‘1

60 65 70 75 80 85 95

Output Y ol an (SR ingot

Aix = ,r)Ku

1972

Nomogram for

the

determination

of higher or

lower costs (MK)

per ton of

useful ingot

weight

Fig. 8

(See Page 16 on howto use Nomogram.)

that are taken into consideration are: cost of crude steel and scrap,

remelting cost, and usable weight of ingots from conventionally

cast and from ESR steel. With increasing crude steel costs Kw

or the difference crude steel minus scrap cost A K, with increas-

ing ESR yield Y, with decreasing yield of conventionally cast

steel X and declining remelting costs Ku, an ESR application

becomes more and more economical.

The large variation of crude steel and scrap costs can be seen

in Figure 9. Starting with unalloyed carbon steel up to stainless

steel, the crude steel costs range from 350 DM per ton to

1,800 DM per ton. While scrap costs only amount to one-fourth

in the case of carbon steels, they reach one-half the crude steel

costs in the case of chromium-nickel-grades.

WSW(Steel grade)

high alloyed austenitic

high alloyed ferriticNi-steel

- -steel 2 — 4.5%Ni)high alloyedCr-Mo-V-st I

- anCr-Mo-Vsteel

0 -

weldable specialstructural steel

C-Mn-steel(up to 1.5°0 Mn)

C-steel(0.25 100% C)

1972

-\

Scrap value Crude steel value

C-steel(0.05 — 0.24°0C)

0 200 400 600 800 1000 1200 1400 1600

DM/t

Crude steel value and scrap value of severalsteel grades

A

Fig. 9

A better insight into the relationship between technical factors

and their influence on basic costs is given in Figure 10. Here,

an attempt is made to balance the cost of ESR steel and convent-

ionally cast steel. Variables are different ingot yield, different

remelting costs (250, 350 and 450 DMA) and the respective crude

steel and scrap costs.

This figure reveals that crude steel cost is less important than

the difference between crude steel cost and scrap cost. With

remelting costs of 250 DM per ton it is visible that equality

of costs is obtained with considerable differences in yield, e. g.

Y : X = 1.8 with A K values of 300 DM per ton. If the yield

advantage in favour of ESR is lower, e. g. Y : X = 1. 4, we need

A K values of 650 DM per ton, even if remelting costs amount

to only 250 DM pe r ton. In general we may as sume that equalmaterial costs may only be obtained in case of alloyed steel withmore than 3 % Ni.

1300

1200

eoo450

350

•

TRemelting costs ifu10M/t of row ingot weightl

13 14 15 76 1.7 70 79 10y/x

75 70 65 60 55 50x/y in °I.

1972

Relationship

between output,

costs ofremelting and

material

(X = conven-

tional ingots,V = ESR ingots)

Fig. 10

The advantage of a better ingot yield in case of ESR steel cannotalways be fully utilized, as this value depends on the shape ofthe forging. Very often it is not possible to cut off exactly therelatively short end crops of 50 to 100 min. During preliminaryforging the ingot edge s advance mainly with higher diameteringots. This cause s higher end crops, which would normally notbe required. Therefore, an inve stigation is neces sary, whether theapplication of ESR steel results in saving s in the followingmanufacturing stage s.

The general manufacturing schedule of large forging s is shown inFig. 11. With conventionally cast ingots we may have an intermed-iate cooling after preliminary forging of the raw ingot. An ultra-sonic inspection follows in order to determine the be st suitedsection for the respective forging. In case of ESR material sucha pre -testing is not necessary.

Since ESR ingots generally do not require a larger reductionratio between 2 : 1, smaller ingot diameter s can be applied inthe case of a plain drawing operation. It is thus pos sible toreduce the number of heating cycle s as well as the duration offorging time.

Figure 12 shows the relationship between a lower reduction ratiofor ESR steels

( =Reduction ratio ESRReduction ratio cony.

9

Convettionat melting

Conventional melting by electricarc or basic oxygen process

vacuum treatment

casting of raw ingot

preliminary forging,intermediate cooling,

testing

final forging slow cookingtesting

rough working testing

heat treatment testing

final machining testing

shipment

Elecfroslag remeMng

Conventional melting by electricarc or basic oxygen process

vacuum treatment

casting of ingots for electrodeswithout or with

prelimina im in

Remelting byelectroslag process Schedule for the

manufacturing

of forging pieces

shipment of rough stampings

and the respective savings as a percentage of the forging costs

of conventionally cast ingots. With at-values of 0.5 savings of

20 % in forging costs are obtained. If forging of conventionnaly

cast steel requires an additional upsetting savings increase nearly

proportionally with Ot» and attain nearly 70 % in case of a-values

of 0.3.

-; 80

.0

-

70a

E 60

-6

50

40 -I=

I 30 •

t-910

112 0.3 04

orging during conventional manulacturkg: I

! al with additional upset ragingb1 without additional upset largIng

0,5 0,6 0,7 0,8 09 1.0

Oelorrnabbn degree n lalt/conventionall

- 10 -

Cost saving

during forging

A

1972

1972

Fig. 11

Fig. 12

Normally, no changes occur in the following manufacturing stepslike slow cooling, non-destructive testing, rough machining, heattreatment, finishing and inspection in comparison with forgingsfrom conventionally cast ingots.

Reference from literature and our own production experienceshow that by an application of ESR ingots the production riskscan very often be considerably reduced. This refers mostly torejection, additional expenses in order to prevent rejection, andreclamations from customers in case of deviations from specif-ications. All this ranges under the name "AT-rate" (rejection rate).

An analysis of particular defects leading to rejection reveals(Fig. 13) that it is rather futile to try to determine an improvementof the rejection rate (AT-rate), since causes for defects mainlydepend upon steel grades and ingot weight. The same is valid

for the absolute level of the rejection rate (AT), which may varyby one decimal factor and is usually not revealed by steel plantsor is related to differing reference bases. All further calcul-

ations bear in mind a reduction of the rejection rate by 50 % incase of ESR material as compared with normally cast steel.

Nonrnetalhcinciusions

Defects caused duringforging and machining

Flakes, porosityand separation

Defects affected by ESR

strongly partially not

Cracks andfractures

Share of single defects on total rejection

Mechanical propertiesnot attained

A

1972

Explanation of

single defects

affecting the

rejektion

rate (A1)

Fig. 13

Besides an improved yield with ESR steels and a reduction of therejection risk, some further savings - mainly during forging - canbe realized.

The calculation of manufacturing costs of large forgings as theydepend upon those factors, which mainly influence costs, is shownin Fig. 14. Additionally, two calculation examples are includedin this chart. The distance between cost curves for ESR

and conventionally cast ingots shows clearly how the higher costs

for ESR steel decrease during production and how they are nearly

balanced after an extra charge for the rejection rate is levied.

8 C A 30

•is if2400

t;

—14004.1-!! 1300

.`j 1200

2 1100

rri9,1000

900

3-; 800o.Io /

/

O 600/

500 50 70

400 59 0,

tr 300 '‘.97

2000

(rit; 700

A-50

Relectio: r5 iskP frO% of ESZet cost;

&al output X in%

Total output Vin %

413 50 1970X 100

1600 7 MOO

1200 • I 12002

_

400 WO "

0 0 e30 40 5 607080100 gm 99oe

mo 0

C0.5(5 1(0 fin OMM, M0011

5 52.:n550054§, Fig. 14

--4000 3000ttil-

.1 (See Page 17 on howto use Nomogram.)

2000 icr;

1000 5,5'

0 WV MX 3000 4000 5000 MOO WWMonulacturMg costs inclusive rejection rote *3

tin OMWI

Remelting

1972 Nomogram for the determination ofmanufacturing costs

The last nomogram contains all factors and is exact but difficult

to read. A specialist does not really need such an exact description,

but can operate with a simpler one to find a solution.

450

400 4‘61:\

\ "̀350 ,‘ ‘1‘ -tV

i555300 4

- 12 -

Strap valueKs rmawk,SW

/ •

A

1972

Influence of

manufacturing

conditions

necessary to

achieve cost

equality

Fig. 15

Figure 15 tries to show in perspective the importance of essential

cost factors, in order to obtain cost equality when producing

machined forging pieces from either conventional steel or ESR steel.

Main factors are:

Remelting costs Ku , differences in yield and rejection risk and

the difference between crude steel and scrap costs. As all further

factors like forging and machining costs could not be plotted in

a 3-dimensional graph, the following simplifications were introduced:

Scrap cost Ks is assumed at a constant level of one-fourth of

crude steel cost; Kw. According to figure 9 this signifies the mean

ratio between average scrap and crude steel costs of all alloyed

and unalloyed forging steels.

Furthermore, the following assumptions were taken into consid-

eration:

Influencing factors

Forging costs Kvper ton of finishedproduct

Costs of machiningand heat treatmentKB per ton offinished roductRejection risk "p"in % of totalmanufactur in costs

Conventionally caststeel

100 % of normal

crude steel costKw per ton charge

200 % of normal

crude steel costsper ton charged

twice

The combination of forging and machining costs with crude steel

costs should be permissible, since expenses for forging and mach-

ining increase with increasing material value.

Figure 15 contains the following factors:

The rejection risk "p" for ESR steels is plotted over the x-axis,

remelting costs Ku over the y-axis and the crude steel costs Kw

over the z-axis. The difference in yield X - Y (X = yield from

conventionally cast ingots, Y = yield from ESR ingots) are introd-

uced as parameters in steps of 10, 20 and 30 %. The parameterssymbols consist of parameter surfaces with constant differences of

X minus Y. However, parameter surfaces of 20 % may correspond

with different X- and Y -values like "X minus Y" as: 60 - 80 %,50 - 70 %, 40 - 60 %.

- 13 -

ESR steel

80 % of normal

crude steel costKw per ton charge

200 % of normal

crude steel costsper ton charged

once

As expected, the crude steel cost Kw, , necessary to achieve

cost equality, increases with higher remelting costs. The same

is true with decreasing yield differences "X minus Y", and

decreasing rejection risk "p".

This latter statement requires an explanation: As the rejection

risk of conventionally cast steel is calculated with twice the value

of that for ESR steel savings in favour of ESR steel increase with

rising rejection risk.

From this one can see: Because of the strong influence of a

reduction in rejection risk and forging costs, useful improvements

of 10 to 20 % in overall yield when applying ESR ingots lead to

cost equality at relatively low crude steel costs. This would not

be expected when comparing yields only.

In order to decide, whether ESR steel or conventionally cast steel

is more economical, one needs a complete cost comparison.

The results of such a comparison for some typical forging pieces

are compiled in table 2. Tabulated according to the end use of

the forging pieces, e. g. alloyed hot rolled bar, rings, rolls,

rotor shafts, and according to rising finished weight, are:

- Finished weight (shipped weight) in tons

- Costs of machined pieces manufactured from conventionally

cast ingots in the as shipped condition in DM per ton.

-

Manufacturing costs of forgings originating from ESR material

as compared with appropriate forgings originating from

conventionally cast steel (conventional = 100 %) computed with

remelting costs of 450 DM per ton for material input, for

forging, machining and heat treatment and for rejection risk.

Furthermore, total manufacturing costs in DM per ton are

included based upon the following remeltings costs: 450 DMA,

350 DMA and 250 DMA.

Despite a considerable scatter one can conclude that even at remelt-

ing costs of 450 DM per ton, cost equality can be attained,

especially in the case of increasing total manufacturing costs, as

a result of more expensive materials, increasing finished weights,

i. e. yield advantage in favour of ESR ingots, and increasing

manufacturing depth. With lower remelting costs of 250 resp.

350 DMA almost all forgings, bearing manufacturing costs of more

than 3, 400 DM/ton in the finished condition and originating from

ESR steel, are as expensive as or 10 % cheaper than those from

conventionally cast steel. (Fig. 16)

SUMMARY

There are no technical reasons but economical ones which hinder

a further propagation of the ESR process. They are mostly based

on a lack of suitable data for a comparison of manufacturing costs

of products made from ESR steel as compared with those from

conventionally cast ingots.

- 14 -

Based on our own production experience this paper attempts to

close this gap for the case of large forgings. Remelting costs may

differ between 260 and 600 DM/t for units with capacities from

1 to 300 tons under the premises of a unit utilization of 50 to 80 %.

The increased material costs, nearly doubled by application of

ESR ingots, can be balanced by a reduction of several cost elements

incurred in the production of forging pieces.

By means of nomograms the influence of remelting costs, crude

steel and scrap costs, ingot yield, forging, heat treatment and

machining costs and the rejection rate on total costs is determined.

The results obtained do not allow plotting a simple general diagram

or offer a pat solution for determining the economics of an applic-

ation of ESR.

In brief, it can be concluded that an application of ESR steels is

economically possible,

a) as long as a high degree of utilization of the unit is obtained,

in order to reduce the disproportionately high fixed costs of

the unit,

b) as long as only higher alloyed steels are remelted in case of

small ingot weights with a lower advantage in yield (in the area

of heavier ingot weights also low alloyed grades can be applied,

because of the better yield),

c) as long as products are manufactured with high rejection risk,

which can be reduced drasticly by an application of ESR steel,

particularly in the case of greater manufacturing depths (i. e.

large value differences between finished condition and forged

condition),

d) as long as forging pieces are produced with high forging costs,

e. g. conventionally cast ingots have to undergo a multiple

upsetting operation, in order to attain the necessary deform-

ation ratio.

- 15 -

The nomogram can be used as follows: (Figure 8)

Find at point (1) of the upper right table the appropriate material

costs and at point (2) the appropriate scrap costs, join both in

the point of intersection (3), draw a horizontal line to (4) and

find there the appropriate beam for å K. Then find the appropriate

yield value, either X for the conventional ingot at point (5), or

Y for the ESR ingot on point (6), join both and draw from the

auxiliary line Z at point (7) a vertical line upwardly to the point

of intersection of the chosen beam (4) and thus get the point of

inte re section (8). From this point of intersection draw a horizontal

line to the right. As the mirror image one finds point (9).

Starting with Y at point (6) and remelting costs Ku at point (10)

find the point of intersection (11) and transmit this by means of

mirror image to point (12). Then points (9) and (12) are joined and

you find the higher or lower costs of ESR steel compared witH

normal cast steel at the point of intersection with auxiliary line Z.

- 16 -

Use the nomogram as follows: (Figure 14)

The example included here for the manufacture of conventionally

cast ingots is marked with uninterrupied auxiliary lines and numerals

and for ESR ingots with interrupted auxiliary lines and letters. Find

in the upper right field of addition the mean scrap values (KB) for

conventionally cast steel (1) and crude steel costs (K ) at point (2)

and bring both to the point of intersection (3). Then by joining

horizontally find point (4) and the appropriate beam of multiplication

in the left field. From total yield (conventionally) X % at point (5)

draw a vertical line joining beam (4) at the point of intersection (6).

The point of intersection (6) is extended horizontally to the right

and intersects K (beam (1) at point (7) and from this point of

intersection (8) with the marginal line draw a horizontal line to

the left toward the auxiliary line "a" at point (9). In the upper left

partial field the forging costs K (in DM per ton of finished weight)V

are plotted (10) and connecting line is drawn across point (11) to

point (9). At the point of intersection (12) with the auxiliary line "c"

find by means of mirror image (13) downwardly to point of inter-

section (15) the costs of machining (KB) at (14), which is reflected

horizontally to the right. At the point of intersection (16) with

the rejection rate "p" draw a vertical line downwardly and find at

point (17) the total costs per ton of finished weight.

Cost determination of ESR steel (interrupted auxiliary lines) is done

in the same sense, but remelting costs must be added and the lower

forging costs must be kept in mind. From the mean scrap value

K (a) and the crude steel value K (b) draw auxiliary beams, which

cross at point (c). Then draw a horizontal line across (c). Starting

at the point of remelting costs K (d) draw a vertical line upwardly.

- 17 -

At the point of intersection (e) with the horizontal line (c) you bring an

auxiliary 45o

-line to intersect the marginal line (f). From there

draw a horizontal to (g). From the total yield of ESR "Y" at point (h)

draw a vertical line upwardly which crosses the auxiliary beam (i)

starting at point (h), draw a vertical line upwardly which crosses the

auxiliary beam (i) starting at point (g). Then draw a horizontal line

from this point of intersection to the right which crosses K at

point (j) and is continued under an angle of 45 ° to the marginal

line (k) and again horizontally to the left to auxiliary line ''a" (1).

In the partial figure at the lower left hand, and starting with the

deformation ratio ESR vs. conventional (m) vertically cross the

respective function (n) which is reflected upwardly. At the point

of intersection with the cost parameter (o) of the forging operation

of the conventional ingot (p) draw a horizontal auxiliary line to (q)

with (1) you find point (r) at the auxiliary line (c). Transmit (r)

by reflection at (s) to the cost parameter (t) for the machining

costs thus finding the point of intersection (u). Transmit (u) horiz-

ontally to the right partial figure and extend this at the point of

intersection for the rejection rate (AT) (v) of ESR material vertic-

ally downwardly. At point (w) you find the total costs of the forging

produced by ESR.

- 18 -

120

115

110 -

105

90

1972

‘Sgo

250

Changes of manufacturing costs(ESR vs. conventional)

Remelting costs lie M/tonsl

—450

350

5 10 20 50 100 200Finished weight lie tons/

.A

Calculations of typical forging pieces for1972 comparison purposes (rejection rate AT

R eel = 50° vent

Fig. 16

Forged product Matenal

Finishedprece corghtion

to besth

forgrag RelaWe manufacturing costs ESR vs conventionalin ao(oony. =100%)

&outlay to manufaduring Total at rernettingstagesat 450DMA remetting costs costs of IDM/t)

ValueaPPE

Weight DMA Material Forgirtg Madm Rrtoenc-250 350 450

Bar 4000 X2CCr13

meg

98 100 661-5 1650 162 112 119 126

Ring 100Cr6 6,7 3750 177 99 100 57 108 113 118

Ring X5CrNi18.9 20,2 10450 106 96 100 65 90 93 96

Cold roll 8007 4,6 5400 163 46 100 68 97 100 103 Table 2

Definition

Crude steel costs

Scrap valueRemelting costs

Forging costs

Costs for heat treatment andmachining

Deformation degree

Metallurgical output,i.e. useful ingot weightvs. raw ingot weight

Total output,i.e. finished weightvs. raw ingot weight

Amortization rate for rejection,extra charges and reclamation

Table of formulae

symbol dimension

DM/t raw ingot weight

KS DM/t

KU

VKB

ESRconventional

ESR Ycony. = X

ESR = Ycony. — X

ESR = pcony. q

DMA raw ingot weight

DMA finished weight

DM/t finished weight

1972

Table

of

formulae

SOME METALLURGICAL QUALITY AND MECHANICAL PROPERTIES OF3% CHROMIUM-MOLYBDENUM STEEL FORGING MADE FROM ESR INGOT

M. Inoue, K. Kubo, K. Takeuchi and Y. Hayashi

Kobe Steel Ltd., Takasago, Japan

ABSTRACT

In order to get a sound engineering steel forgings with reliable and highquality, there have been many improvements in ingot making and forgingprocedure.

Electro Slag Remelting process has been interested to be one of themethods for having the metallurgical un formity in addition to be high qualityof toughness.

Kobe steel ltd. has constructed a larger ESR furnace having a capacity of50 tons in weight of ingot for the purpose of producing high quality largesteel forgings such turbine as rotor shaft.

This report was the results of the investigation on some metallurgicalfeatures of steel forging made from ESR ingot.

The material was 3% chromium-molybdenum steel forging made from 10 tonsESR ingot for marine low pressure turbine rotor shaft having approximately900mm, diameter and 8.5 tons at final forging condition.

And these results have been compared with that of the conventionalforging made from vacuum degassed ingot.

1. Introduction

SOME METALLURGICAL QUALITY AND MECHANICAL PROPERTIES OF3% CHROMIUM MOLYBDENUM STEEL ROTOR FORGING FROM ESR INGOT

M. Inoue, K. Kubo, K. Takeuchi and Y. Hayashi

Kobe Steel Ltd., Takasago, Japan

Recently, there have been several reports on the metallurgical andmechanical properties of forging produced from Electro Slag Remelting (ESR)ingot and it has become of great interest in its excellent characteristics.

Kobe Steel Ltd. has studies on ESR process since 1962 and has alreadyconducted many investigations on various kinds of forgings such as crank shaft,turbine blade and hardened work roll for cold strip mill and aluminum foil millby using ESR ingot.

As the satisfactory results were obtained from these qualitative testingsand performances in actual service of products, ESR ingot might also be muchsuitable one for the production of larger forgings required high toughness andreliability such as turbine rotor shaft.

In 1970, a ]arger sized ESR furnace having a capacity of melting 50 tonsin weight were installed as the first step to the production of high qualitylarger forgings.

This paper is concerned with the results of the examination on 3% chromium-molybdenum steel forging made from 10 tons ESR ingot for marine low pressureturbine rotor shaft having about 900 mm diameter in the body and 3500 mm intotal length at finished machined condition.

And the results have been compared with those of the conventional forgingmade from vacuum degassed ingot having similar composition and size.

2, Test Material

Test material was 3% Chromium-molybdenum steel forging for marine lowpressure turbine rotor shaft made by the manufactur ng procedure as shown in

fig.2 and fig.3.

Photo.1 shows 10 tons ESR ingot and photo.2 shows the test rotor forgingproduced from this ingot.

As shown in fig,4 and fig.5, in addition to the essential differences ofingot making process, there were remarkable differences in ingot size, forging

procedure and annealing method between ESR and conventional forgings.

The yield of the final forging from the ESR ingot was approximately 85%and far higher than that of the conventional forging having about 50% yield.

And the forging procedure and annealing treatment were much simpler in

operation and shorter in operating time compared with those of the conventional.

Table 1 shows the specification of chemical composition and mechanical

properties for the rotor shaft tested.

The rotor forging for this testing was confirmed that there were not

observed any kinds of indication and defect by non-destructive testings suchas magnetic particle and ultra-sonic testings and also material inspection of

tensile and impact properties taken from the extra portion of the ends of

journal and body were within the specification of the rotor shaft.

3. Results of testing.

3.1. metallurgical quality.

A. Sulphur Segregation.

Sulphur segregation in the forging has been thought to be intrinsicundesirable phenomena and larger the ingot, the sulphur segregation becomesmore and serious.

In the ESR rotor forging, it could not be found any visible segregation

pattern and showed uniform dispersion on all over the section from top to

bottom of forging as shown in sulphur prints of photo.3, photo,4 and photo.5.

B. Segregation of chemical composition

Fig.6 shows the results of the analysis of chemical compositions at various

positions from top to bottom of the rotor forging.

Distribution of the chemical compositions were very uniform throughoutthe sections except a little lower content of carbon at the end of the bottomside journal.

C. Cleanliness

Table 2 and fig. 7 shows the result of the analysis of oxygen, oxideinclusion by chemical analysis and of cleanliness measurement by microscopicalexamination.

The contents of oxygen and oxide inclusions were not only lower than twothirds of conventional forging but also distributed much more homogeniouslythroughout the forging.

And it was observed that inclusions of oxide and sulfide were finelydispersed in the forging than those in the conventional one.

D, Macro structure

Before macroetching, dye-penetrant and magnetic particle testings werecarried out on the various sections of the ESR forging, but there were noindications of defect.

Macro structures on various sections showed also very clear and soundpattern that was free from local attack, but some dendritic patterns wereobserved in the inner portion of the body probable due to the small reductionof sectional area (forging ratio) on forging.

3.2. Mechanical properties.

A. Room temperature tensile and impact properties.

Table 3 and table 4 show the tensile and impact properties at variouspos tions and directions of ESR forging.

The tensile and impact values of the tangential direction at the outerand the midway of the body showed little differences from those of the radialones.

The values shown in the tables were the average values of two measurementsfor the tensile and three for impact.

The tensile and the impact properties of all the specimens were highenough for the specification values.

Compared these results with those of the conventional forging, there weresome superior characteristics of ESR forging as follows,

(1) The transverse ductility and toughness of the body in ESR forging werenot so different from those of the longitudinal ones. But in the conventionalforging, the transverse ductility and toughness were inferior to those of thelongitudinal direction as shown in fig.8.

(2) Variation of the strength, ductility and toughness in all portions withinthe body were much less in the ESR forging than the conventional.

And all the values of the ductility and toughness in the transversespecimens, in spite of higher tensile strength, were high enough and near orbeyond the upper limit of the values obtained by the testing of the conventionalforging as shown in fig.9.

(3) Variation of the ductility and toughness at the center core of the ESRforging seemed to be related mostly to the diameter of each sections of theforging, that is, the smaller the diameter, the better ductility and toughnesswere obtained as shown in fig.10.

But in the conventional forging, there was remarkable dependence on theposition relative to the ingot.

This characteristics of the ESR forging have indicated that the ductilityand toughness at the center core depend on only the forging ratio and quenchingvelocity, and there is not influenced by the position relative to the ingot.

As mentioned above, in spite of simplification of forging and annealingprocedure, some excellent characteristics on tensile and impact propertieswere obtained in the ESR forging.

Especially, the transverse ductility and toughness at the center core ofbody which were thought to be most critical region for rotor shaft have beenfound to be far superior to those of the conventional.

To clarify the ductility and toughness at the center of [SR forging, itwas carried out the experiment to compare the relationship between the strength,the ductility and toughness with the specimens taken from each positions andheat treated under the same condition.

Fig.11 shows the experimental results of the tensile and impact testingsof the specimens which were austenitized at 880°C, and quenched at the coolingrate of 1500C./h, followed by tempering at various strength levels.

From this experimental results, it was found that the transverse ductilityand toughness in the center core of the ESR forging were almost the same withthose of the longitudinal direction of the journal which was most superior,and far superior to those of the midway and the center of the conventional.

Then another experiment was conducted to know the effect of reduction ofsectional area during forging on the ductility and toughness at the center ofthe forging.

The samples taken from center core in various section were heat treatedas the method described before.

Fig.12 shows the relationships between the ductility and toughness andthe forging ratio which was calculated by the ratio of cross sectional areaon upsetting to that on final forging.

These relationships show that the tangential ductility and toughness atthe center core were almost the same and not close relation to the forgingrat o within the investigated range.

From these experimental results on the forging effect, it was found thatthe ductility and toughness, even at the center core of forging having adiameter of 900mm, could be expected to be high enough with the forging ratioas low as 1.34.

B. Fatigue strength

Table 5 shows the results of rotating bending fatigue tests on ESRforging and conventional.

The test pieces of this testing were 10 mm, of diameter and 50mm. inguage length, and rotated at 1800 cycles/min.

Fatigue endurance limits at 107 cycles were 41 kg/mm

2. at the outer and

37 kg/mm2. at the center of the body of ESR forging.

Compared these values with those of the conventional forging, remarkableincreases were obtained in the fatigue strength and also ratio to the tensilestrength at the inner portions of the body.

C. Low cycle fatigue strength.

Fig.13 shows the results of low cycle fatigue test of the specimens taken

from the outer and the center portions of the body in ESR forging.

The testing was conducted with the method of cyclic loading of tensilestress at the given stress on notched specimen as shown on the top of thefig.13.

This result shows that there are little scattering and are littledifference between the outer and the center of the body.

D. Creep rupture properties.

Fig.14 is the result of creep rupture test of the rotor forgings.

The rupture strength and ductility of ESR forging has not only higher butalso less scattering compared with those of the conventional.

4. Summary

The investigation of 3% chromium-molybdenum steel forging produced fromElectra Slag Remelting ingot has shown some superior characteristics ofhomogenity and mechanical properties.

The most distinguised was its high reliability and toughness of transversedirection at the inner portion of the forging.

This testing is concerned with rather small size of forging for turbinerotor shaft, but the investigation on the larger one produced from 50 tons iningot are now getting on.

3

Top discard

Diameter (D)

Forging ratiofrom upsetting

(10702/D2)

Consumable electrode

Molten slag

Molten metal

Solidified ingot

Fig.l ESR ingot making process.

820mm.dia.

380mm, 925 wm.

7.93. 1.34

Finishing forging

t.,

700 mm,

2.34 4.97

Cooling water

Ingot mould

2400 mm. 2300 mm.

820mm, dia.

ESR ingot Ready for upsetting

t 230mmm- 640mm. 750mm. ,NH 650rnme 510mm,

480mm.

Fig,2 Forging procedure of the ESR rotor forging.

1300mm

1070 mm.

After upsetting

00mm.

Bottom discard

Ingot10 tons

Iot20 tons

Final forging

A. ESR forging

Aealing after final forging

880°C

650°CFurnace cooling

60h.

Oil quenching640'0

20h. 30h.

QUenching and tempering

Fig.3 Heat treatment of ESR rotor forging

Upsetting

(0-*Upsetting

Furnace cooling

Finalforging

UPsetting

Discarding Final forging

' B. Conventional forging

Fig.4 Comparison of forging procedure between ESR and conventional forging

650°C 650°C

4 °

920°C 900°C

4 °C

670°C

4-100h, 4 250h.

For LSR forging For conventional-forging

Fig.5 Comparison of annealing cycle between ESR and convetional

forging

Carbon

Silicon

phosphorus

Suluphur

Chromium

Molybdenum

0.28

0.008

X

0.29 x0.29 X

0.013---7e0.016

0.015 x 0.016

0.015 x 0.015 X 0.014 x 0.016 x0.016 x 0.016

x 0.007 x 0.007

x 0.008 x0.007

x 0.007 X 0.008 X 0.007 x0.007 x 0.007

x 2.67 x2.67

x 2.67 x 2.672.67„2.69, , 2.67 2.68, 2.67,0.47) xl0.47) 'l0.A3) X(0.47) X 0.A7)

x(0.47) x(0.47)

x(0.48) x(0.47)1 rDistribution of chemical comcosition ( wt. 2 )

*1

150

100

50

- 75

0S.

1w-......,.\\\4\\

\ ESR forging 6-0 Conventional

. 1

\

0

forging

\\

Outer _ \

123 4 51si e

1

1 CenterESR forging

10-0 Conventional 1

forging 1

•

1 2 3 .Size

Snlfide

0

Oxide inclusion

Center

011

1 23 45Size

1 23Size

0

Fig.? Comparisons of-size distribution of'sulfide and oxide at the

. middle ofbody between l;SR and conventional foi.gings.

Size 1 c2x165m2 cros6 section<Siz,,,.2 <-4x10"mm<Size8x101.rm <1.Size. 4.<16x15<Size 5

7 <

cd

at)

• 70

4-I0

0 •

•

600

$4

k 5050 60

+40

Longitudiaal reduction of area

0

70

+20 +40

O : ESE forging

O :Conventionalforging -

Longitudinal FAT C.

Fig,8 Comparison of relationship of reduction of area and FATT

between.transversa and longitudinal directions of the body.k

70

60

-

•-Top of body0- 1didd1eof bodya-Bottom of body

_

[FLit—

o 077'

+10„.4

ee

/////' -4+20 P.

1 1 IOuter Liftay Centerhadia, Radial Tang.

Top4 Mi1dle 1Bottom

-40

-30

-20

-10

25

A

4-)

0

25

A4-)

C 1

155

oH IW W

M

104'

cd

5

•C)

S.fl

+30 Lc\

Position and direction of test Pieces

Fig.9 Variation of transverse tensile properties and impact

propertieS from outer to centerr of body

Hatched area shows the range of data obtained from the test,of

conventional rotor forging of body

--4200 Conventional rotor_

+10 Tangential

0

A— B— C B — F - G H

E3, BSR rotorV. Tangential

0-0

ESR rotor Longitudinal

•

A B C B EFGHPosition of test pieces

Pig.10 Variation of tensile srtength, reduction of area andPATT at the center from th,,:s top to -2;.317a bottom

+400.0 +20

50

0-

-4--- ar —IA

P—)0

0 -----

-20

-40

-60

-15

-

Taken from 2ESA, Conventional forging

sa

1:1:"-":"1:0

X--- -2(

70, 80 90 100Tensile strength kg/mm2

Fig.11 Relationship between tensile strength and tensile

ductilty and FATT of various kindSof test pieces after

heat treated to various strength level

Top journal Outer LongitudinalMiddle body Midway Radial

11

( From sulphur segragated area)Middle body Center. Tangential

al

(Ft0

Fig.12

•

70

•

609-10

•

50`14

1 GI 2

0-- o Longitudinal test piece

CF0 Tangential test niece

Forging ratio4 68 1 2 4

Forging.ratio

0

20

1068

20

Cd

40VEI o°rri

60•1#?..0

Effect of forging ratio on the ductility and toughness

at the center of forging.

50

60 mm

12mm •

Test specimen

k=2.52

Radial direction from outer ojf body

0 Tangential direction from center of body

1 103 104

Cycles to fracture

?ig.13 Tie result of low cycle fatigue test on ESR forging,

10

so

50

40*PA

ESH, Conventional forg-ing..0-0 —0 Outer.of body, Radial

Midway of body, Radial- - Center of body, Tangential

Ed •-•-

•0 c+ 0 0 H

.c+

E d L .+ 0

Phot

o.3

Sulp

hur

prin

t on

cr

oss

sect

ion

(x1/

10)

Phot

o.4

Sulp

hur

prin

t on

ve

rtic

al

sect

ion-

of

body

(x

1.

540)

Bottom of joy

Photo.5 Sulphur prints on the vertical sectionbf body (x1.5/a

Ph

Body

Centeraxis

o.6 Macro structures on the verti al sectIons of body

-al (x

Journal

value(2mm U notch) 5.krs-m/cr?for trans.

.*1 : Longitudinal *2 : transverse

Table 2 Distribution of Oxygen, oxide inclustion and cleanliness

Position of sample Oxygen Oxide Cleanli-ppm ppm ness

Top of journalmarked TP.1

Top of bodymarked TP.2

marked TP.4

in ESR rotor forging.

Bottom ofjournal marked

,TP.5

Middle of body

marked TP.3

Bottom of body

TP.1 ---1-----""I I x S. xST I j 1xs. TP.5

xA.132_1 (----__Ix C. )( C. x0. . x C.

I ri 1JPosition of sample.

0. 20 46 0.062C. 22 46 0.058

S. 20 43 0.051

L. 20 35 0.048

C. 21 30 0.050

S. 22 48 0.054

M. 24 50 0.061

C. 24 42 0.059

S. 19 42 0.047

17 45 0.052

C. 17 36 0.045

S. 19 37 0.050

C. 19 38 0.046

Table 3 Tensile properties of 6512 rotor forging.

TP.1 TP.2 TP.3, T4.4 TP.5

kilT2

TP.1 'T2.2 132.3 TP.4 T .5I , I

Position of sa7.ole

Table 5 Rotary bending fatigue strength

A) ESR rotor forging

Position and directionof test piece

Middle of bodyOuter Radial 41.0 80.9 0.51

Midway Radial 39.0

Center .Tang. 37.0

B) Conventional rotor forging

Middle of body

Outer Radial 39.0

Midway Radial 35.0

Center Tang. 30.0

Fatigue limit, Tensile Fatigue limit at 107 cycles strength Tensile stren-

kg/mm2 kg/mm2 gth

80.4 0.49

76.9 0.48

79.5

78.7

70.6

0.49

0.45

0.42

Abstract

METALLURGICAL PROPERTIES OF FORCINGS FOR HYDROELECTRIC POWER PLANTS- DJERDAP ON ThE DANUBE

A. Sarajlic, Zenica, Yugoslavia

Recently a hydroelectric power plant was completed in the Yugoslaw-RumanianDanube Area, having an annual capacity of about 12 billion Wh. For thispower plant and the corresponding ship locks more than 7000 tons of forgedblanks were required. The type of steel most used, namely 65% of the totalquantity, was the weldable Mn steel of type St 52, approximately 0.2% C and1.3% Mn, while the other types were low alloy and carbon steels. The individualweights of the forging blanks were in the main, namely 35%lbetween 5 and10 tons.

The notch impact strength in transverse direction at low temperatureand the elastic limit at larger dimensions proved to be sensitive propertiesof Mn steel with regard to operational stress and receiving conditions.That is why the conduct of these properties was subjected to a closerinvestigation, by considering the forging parameters as functions of theabove.

The following forging parameters were selected: a) the degree offorgeability in limits of 2.5 to 10; b) heat treatment under variousvariants, whereby the chilling speed following austenitization was withinlimits of 600 - 2°C/minute, and after the tmpering it was 100 - 0.1°C/minute.

The small samples were taken from a 3 ton ingot forged stepwise,and in a simulating manner the actual operating programs were tested andheat treated in various manners. The relationships between impact notchstrength at various temperatures and corresponding heat treatments wererecorded in diagram form. The small degrees of forging and the rapidchilling thereby produced the best impact notch values, while slow chillingand high degrees of forging at low temperatures led to a pronounced brittlenessfracture in radial direction. The tendency of Mn steel to temper-insbrittleness only showed following rapid chilling after austeniizatlon.

The fine separations were determined at the former austenite granulelimits at the total brittle impact notch samples, following a microscopicexamination, in increased intensity.

The examination of the yield point values resulted in its directdependency on the chilling speed following austenitization.

INTRODUCTION

Approximately 230 kilometers to the east of Belgrade, a hydro—electricpower plant was installed recently and called "Djerdap = Iron Gate" in viewof the proximity of the narrow pass by this same name. The construction ofthe plant was started by virtue of a decree based on a treaty betweenRumania and Jugoslavia, and it took severi years to construct it. Thispower plant which at the same time contains the locks to regulate watertraffic was built in accordance with the large hydroelectric planti'inthe Soviet Union.

Based on the rated capacity and average quantities of water, thispower plant is supposed to generate approximately 12 billion KWh annually.

Large quantities of forgings were neeoed for the mechanical equipmentof the three most important departments of this power plant, namely theturbine halls, the ship locks and the weir installations. This paperwill discuss these forgings.

One section combines the statistical data on quantity, weights offorgings and types of steel. At the same time problems of the propertiesof the most important steels are discussed and confronted with calculatedoperating stresses.

Another section reviews the properties in function of the degreeof forgeability and of the tendency toward tempering brittleness ofMn steel of shade 0.2% C, 1.3% Mn and discusses these factors. The shareof this type of steel amounted to approximately 65% of the entire supplyof forgings.

GENERAL DATA

The area with a favorable downward gradient and the large quantitiesof water from the Danube averaging 5,520 cubic meters/second formed areliable basis for building such a hydroelectric power plant in thevicinity of Djerdap.

The blocking of the Danube was effected over a width of 1278 meters.Picture 1 shows a scale model of this power system.

The weir dams /barrages/ for controlling the level of the Danube arein the center, the turbine halls are adjacent on both sides, and near theshores on both sides ship locks are built for the passage of ships.

Picture 1 - Scale Model of the Hydroelectric "Djerdap" Power Plant

at the Danube

6930NT

34.00 45.70

Picture 2 - Section through the Turbine Hall

Turbine Halls

One turbine hall each was builtfor the generation of , newerJugoslaw and on the Rumanian side.Each hall houses 6 generator uEach unit has an installed capacity of 180 NW.

The position of these unitsin the turbine hall can be seen from

picture 2. The directed water jetsoperate the Kaplan turbine which inturn is coupled directly to the rotor,The generator rotor, with adiameter of 14 meters rotates at 71 rpm,generating a voltage of15 KV which is transformed to 400 KV.

The coupling of the Kaplan turbineto the rotor is accomph shedwith a hollow forged shaft. In its finishedcondition this weighsapproximately 84 tons. Tbe shaft, whose tersdnal flangeshave a d itmeterof more than 2.3 meters was welded from 3 forgingsaccording toelectroslag method and air-tempered. Mn steelof shade St-52 was thekind of steel used.

Picture 3 shows this turbineshaft at the moment of thepreparation of the mounting.

Many other types of forghigswere likewise used in theturbine hall, for examplethe rotor shafts weighing17 tons, the connectingparts between 100 ton casthousing and blades weighing21 tons, various parts forhydro cylinders and for the ori-entation of the watercurrents installations wereused which are installed

- spirally.

Picture 3. Turbine shaftof Mn steel prepared andreadied for mounting.

Ship Locks

As a result of the dam barrage of the Danube,the upperwater levelis raised to a level of about 69 meters, while the lower level remains atabout 35 meters. This difference in height of about 34 meters is bridgedfor the necessary passage of ships by the lock chamber built for thispurpose in two stages and having an overall lengthof 2(

The upper and lowerchamber with correspondinglocking gates are emptied andfilled on the basis of thegravitation principle. Theopening and closing of thegates is accomplished bythe special hydraulic system,called servo motors, whosecylinders and piston rodswere made from forged steel.Many of these servo motorshave lengths exceeding 20meters. The inner diameterof the cylinders is as highas 600 mm and the diameterof the piston rods as highas 260 mm.

Picture 4 shows suchservo motors. The one on theleft side reaches a length of23 meters. A total of 32 motorsis required for the lock chamberlocated on both sides.

Weir Systems

7

Figure 4 - Servo jotorsLock Chamber

14 overflow attenuatmrs are locatedin the center cortion betmeenthe turbine halls, each of which having theupper and "owe'construction. These constructionsweighing 200 and/or 280 tons arepushed upward or downward each by two servomotors each, whereby thelevel of the Danube is maintained at a certainaltitilie. As mdi catedbefore, these servo motors are assembled primarily from forged cyLinders ,rings and piston rods. A total of 56 servo motors is instthis purpose.

Statistical Data

The indications available in the projects as to the characteristicsof the forgings make it possible to get insight information about thequantity, unit weights, types of steel, operating stresses and conditionsof receiving.

Grouping the Forged Blanks as to Weights

Table 1 represents the quantity of forged blanks required for allinstallations on both sides. This quantity was copied from the originalprojects after converting the finish dimensions into forging dimensions.In table 1 the weights are classified into certain intervals withinwhich the unit weights of the forged blanks also are located.

As can be seen from table 1, the total quantity of forged blanksis 7,131 tons. Of this total 64% accounts for the forged blanks in theturbine halls, 16% for the lock chamber and 20% for the weir systems.

Regarding grouping of the forged blanks according to unit weights,the lion's share = 34.2% corresponds to unit weights froM 5 to 10 tons.

The quantities listed do not include the forgings built intosideline installations, such as cranes, pumping stations, ventilation,etc. It is estimated that some more % of forgings must be added, sothat a total of about 7500 tons of forged blanks can be figured.

Grouping of the ForgedDlanksas to Types of Steel

Table 2 shows the quantity of the forged blanks according to thetypes of steel. As this table shows, this refers primarily to low-alloyedsteels. About 65% of the steel used was Mn steel of the weldable shade0.2% C, 1.3% Nn.

The other types of steel are low-alloyed steel on Cr or NiCrMobasis, while about 10% were pure carbon steel.

Quantity of Forged planks according to Types of Steel

Section

Types

of Steel

Table 2

C Steels

Mn Steels CrV Steels CrMo Steels NiCrMo Steels

total

tons

Quantity of Forged Blanks according

to Unit Weight

Table 1

Section

Grou

s of Unit

Wei

hts

total

u to 1

1 to 2 tons2to 5 tons5 to 10 tons10 to 20 tons20 to 40

over 40

tons

Turbine halls

105

41

588

1300

255

1620

636

4545

Ship Locks

110

117

192

286

415

1120

Weir Systems

169

190

252

855

1466

Totals

384

348

1032

2441

670

1620

636

7131

Share in %

5.4

4.9

14.5

34.2

9.4

22.7

8.9

100.0

Turbine Halls

462

2950

540

328

265

4545

Ship Locks

78

796

96

150

1120

Weir S stems

196

884

386

1466

Totals

736

4630

540

424

801

7131

Share in %

10.3

65.0

7.6

5.9

11.2

100.0

Stress and Receiving Conditions

The considerable differences between the temperatures in Summer and Winterin this Danube area call for higher demands for exterior installations thanfor interior installations, where, like in the turbine hall, no considerabledifferences in temperature exist, insofar as the quality of the steels isconcerned.

In this respect, the demands for the impact notch strength of the moreimportant forgings are shifted toward the lower temperatures. Certain minimulfvalues, for example -40°C, are required for the principal parts of all servomotors, such as cylinders, piston rods, flange parts and connecting parts,and even at this temperature a certain plastic share shall be determined atfracture areas.

Because, viewed from another angle, this stress relates primarily tostatic loads, the major importance is concentrated in the receiving offorgings to the yield point which was placed in relation to the calculatedstress. This relation, also called safety factor, fluctuated between 2.5and 3.

A stress in the body of up to 9.3 kp/mm2 is expected according to thecomputation for the large turbine shafts which were hollow-forged fromweldable Mn steel. The maximum stress occurring at the flanges with acci-denfal relatively unfavorable and short-lived starting or braking iscontemplated as mounting to up tg 20 kp/mm2. For the large forgings ayield point of at least 26 kp/mm4 and an impact notch strength at + 20°is demanded at a minimum of 3.5 kpm/cm2 at the Mesnager test for largeforgings at receiving and at transverse samples.

The same conditions also apply to the rotor shafts.

The cylinders of the servo motors welded together into larger lengthsare likewise forged from weldable Mn steel. The internal operating pressureof the oil rises up to 170 kplcm2 and the calculated admissible tension inthe cylinders is 14 kp/mm2. These forged blanks are required to show at thetransverse sample a yield point of at least 35 kp/mm2 and at least 4 kpm/cm2at -40°C with the DVM test.

Piston rods for servo motors are forged from steel 36 NiCr 6 accordingto original literature.

This type of steel which also contains some molybdenum contains approxi-mately 1.5% Ni and 0.5% Cr. The piston rods whose lengths also exceed 20meters are forged from one piece. The maximm operational lo24 reachesaccording to mathematical computations as high as 15.6 kp/mm . In thisconnection weights of up to 200 tons must be lifted and resistances mustbe overcome which can occur particularly in Winter due to ice formationat the construction.

The minimum yield point required is 45 kp/mm2 and the impact notch strengthmust be at least 4 kpm/cm2 at -40°C at the DVM test.

PROPERTIES OF WELDABLE Mn STEEL

The official standards still furnish only incomplete data about theproperties of forgings from weldable Mn steel, particularly in the caseof impact notch strength values at low temperatures and of yield pointvalues at larger dimensions. That is why in such cases the receivingconditions frequently are discussed and agreed upon with the client.

During the testing of the different forgings from Mn steel which weredelivered by our forging plant for the "Djerdap" power plant, thesecritical properties proved to be particularly sensitive, so that a corres-ponding examination was conducted in order to check on these conditions.

It is a matter of general knowledge that nowadays various proceduresare used on a metallurgical basis in order to lower the transfer temperaturesof steels. On the one hand, this involves different melting processes andon the other hand the refining ofthe granulation based on micro alloying.These methods are being examined very much in general and still will offera wide area of research with regard to the possible combinations.

In our case the examination of these problems is limited to the forgingparameters which are applicable as a result of the degree of forgeabilityand the different kinds of heat treatment. Special attention was given tothe impact notch strength at low temperatures in function of the degreeof forgeability and the tendency toward brittleness under tempering,expressed by the chilling speed following tempering.

Execution of the Experiment

For the testing a 3 ton forged ingot was used which was melted in thearc furnace. The melting analysis had the following composition in %:

Mn Si I' S Cr Ni Cu Al No.19 1.28 0.37 0.022 0.020 0.04 0.04 0.16 0.15 0.0104 (sic)

The 3 ton ingot was rounded or trued after heating to about 550 mm in diameterand upset to a diameter of 620mm. The center pert of the ingot then wasforged further and gradually and then stretched to the diameters listed below.

Diameter of the forged blank, in mm 390 310 250 195

Corresponding degree of forgeability: 2.5 4 6 10

Without chilling, this forged blank was subjected directly to an anti—flake treatment, in order to avoid possible brittling by hydrides /1/. Thiswas followed by a normal annealing.

Because the examination is based on the system of comparison ofvalues, and since a major number of samples was necessary, generallya distance of r/3 from the surface was selected as testing position.The longitudinal sampleshad their axis at thispoint,and the radialsamples had their notchedincision and/or their center.Figure 5 shows theposition of the samples.

2,

;

Figure 5 Diagram of Sampling.

The programmed heat treatment was developed primarily from twodifferent points of view:

a.— In the first place the chilling speed after austenitigatiod isrepresented to a certain extent as a factor

The dimension of the small samples cut out was 15 x 15 x 70 mm. Thesesamples then were subjected to different heat treatments which were soscheduled that they represented a certain imitation of the actual heattreatment for various extreme dimensions of the forgings.

which regulates the yield point values and also is a parameter whichinfluences the tenacity of 1in steel.

In order to place the practical values of the chilling speed after auste14-dzation into the program, the chilling of a bar of 100 mm in diameter inwater was taken as the fastest, and the chilling of a cylinder of 700 mmin diameter in air was considered as the slowest speed. According toBandel-haumer /2/ the average cooling speed . to 500 °C is about 600°C/minute for 100 mm diameter and about 2°C/minute for 700 mm diameter.

These cooling speeds are shown in the continuous diagram of the steeltested as curves I and 2, in figure 6.

Chemische ZusammensetzunSi t.-4n P S Al N Cr Ni GU

019 7 8 020 0 5 016

1000 Austenitisierungstempmatur,ANT

900 Halledaum =15min

1800 .

Act--> 700 - 8 •

'E 600 - • '5 2

'Z`," ,LpL400 rfrt— -

a moam

4)`,4444rmo

1 105Zeit in s

- 10 -

20/1

1

Picture 6 - Chilling Limites in the Konti

Diagram of Mn Steel

Thus, the field between curves 1 and 2 is covered by these two extreme

chilling speeds in a manner of speaking. The micro structure visible in thediagram corresponds with these chilling speeds. The structure subsequent tothe chilling speed of 600°C/minute contains also perlite and bainite besidesferrite. After the slower chilling the structure is slightly coarser andpurely ferritic-perlitic.

An intermediate speed of about 30°C/minute /curve 3/ also wasselected for many tests.

The heating spePd of the small Samples and the duration of the

austenization at 920°C were adjusted to these extreme dimensions /table 3/.

b.- The other problem to be clarified in some manner by the heat treatmentis the tendency of Mn steel toward brittleness under tempering.

The problem of tempering brittleness is interesting with heavy forgingsalready because of the strong dimensions and stress-relieving-annealing afterwhich generally slow chilling must be anticipated.

The tendency toward tempering brittleness can be expected with Mn steelalready due to the increased Mn content. It is generally known thatmanganese is one of the elements which directly promote tempering brittleness.Thus Opel and his assistants show [3] that an increase of the Mn content from0.25 to 0.66 considerably reduces the impact notch strength of a rotor steelboth at ambient and at low temperatures. W. Steven [4] also shows thatmanganese directly promotes the tempering brittleness. In many cases thereistalk, however, of a slight.effect of this element or of its indirect effectvia the P-diffusion [5].

The tendency toward tempering brittleness is on the other handdependent to a high extent of contaminating elements P, Sb, As, Sn [6,7].As for our steels we calculate on an average due to the composition of theore with a slightly higher content in Sb and As. For the test melting treated,chemical analysis established the following content in these elements [in %]:

As Sn0.022 0.035

In order to express and indicate the tendency toward tempering brittlenessthe cooling speed was varied in the program after the tempering. 100°C/minutewas selected as rapid cooling speed; it is obtained when chilling acylinder of 100 mm in diameter of 600°C in oil. 0.1°C/minute was taken asslow chilling. This corresponds approximately to the slowest coolingencountered in forging literature [3,7,8] when treating the temperingbrittleness.

Within this broad interval from 0.1 to 100°C/minute a more moderatecooling speed of 5°C/minutewasalso chosen to fill possible gaps.

Table 3 represents the types of heat treatments which were programmedfor the small samples.

0.034 0.007

Program of Heat Treatment

Impact Notch Test

- 12 -

Table 3

Type of Heat Treatment [WB]

Prior to the programmed testing of the impact notch strength, thestructural properties and the mechanical properties via the tearing teststheBaumannand deep etchtests were performed on the forged blanks. Theremoved disks showed no flaws.

Figure 7 shows in diagrams the values of the impact notch strengthat radial samples in functinn of the temperature for different degreesof forgeability [VG] and the first group of the heat treatment underA, B, C. The impact notch strength was tested on the Charpy-V sample.The testing temperature ranged from + 60°C to - 70°C (140 to -94°F).

The diagtams show that the impact notch strength highly depends on thedegree of forgeability. The tendency is the same in all diagrams. Thedifference is particularly high at plus temperatures, but decreaseswith lower temperatures.

The influence of the chilling speed after the tempering

14 drmebehan lun

13 VG 25

VG6

4

VG10

piabie_h

—VG25

Tempercaur,°C Tempercaur,°C

-13 -

drmebehandlungC_

VOL,

VG6

VG10 VGIO

—VG4

VG 5

VG6

-60 -20 +20 +60 -60 -20 +20 +60 -60 -20 +20 +60Tempercaur,°C

(VG = Degree of Forgeability)

Picture 7. lm=pact notch strength as a function of the type of heat treatment

A,B,C and of the degree of forgeability (VG). Radial samples, Charpy V.

from 100°C to 0.1°C/minute, by which the types of heat treatment A,B and

C differ is not so highly pronounced that this can be evaluated easily from

the diagrams of picture 7. Consequently a diagram is offered in picture 8whowing the degrees of forgeability 2, 5 and 10, and heat treatment A and

C - 100°C/minute and 0.1°C/minute cooling speed afger tempering.

A direct comparison of the curves for the degrees of forgeability 2, 5

and 10 shows that following the slower chilling [heat treatment CI theimpact notch strength values are lower by about 1 kpm/cm2 and that thereis a significant difference at all temperatures. However, with the rising

temperature the influence of the higher degree of forgeability on thebrittling becomes greater than the tendency toward tempering brittleness.

The hardness also was measured at the impact notch samples, in order to

evaluate the possible truth of the comparative results.

14

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Tem

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°CVG =

degree of forgeability;

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Picture 8 Comparative

Curte

of impact notch strength for

heat treatments

(WB) A and C

War

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Wei

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.

VG

25

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Picture 9 - Impact notch strength asafunction of the heat

treatment

E,F,G and of the degree of forgeability

(VG)

Radial samples, Charpy V

After heat treatment A,B,C all samples showed a Brinell hardness of about170 Hp at +20°C. The hardness measured for information increased at -60°Ccharge time 15 seconds, to 187 HB. These figures showthat there was nodifference in the hardness of the samples and that a comparison of theimpact notch values may be evaluated fully. The hardness test shows alsothat at low temperature a certain consolidation of this steel takes place.

Picture 9 shows the curves of the impact notch strength at radialsamples for the heat treatment systems E,F,G in the same manner as inpicture 7.

The tendencies in the position of the impact notch strength aresimilar to those in picture 7, only a stronger embrittlement can beseen at forgeability degree 10.The total embrittlement fractureis reached already at -40°C, regardless of whether the chilling aftertempering takes place rapidly or very slowly.

Similar to picture 8, picture 10 showsa direct influence of the chilling speedafter tempering.

It is shown that in this case thereis practically no influence of thechilling speed after the tempering, asidefrom the degree of forgeability. Thepreceding slow cooking after theaustenization had a strong influenceon the impact notch strength so thatunder the later tempering no moreeffect was achieved in this respect.

The hardness, measured at the corres-ponding impact notch samples after heattreatment E,F,G revolved around 155 + 5HB units. This is softer by one shade-than with the heat treatment samplesA,B,C which were chilled faster afteraustenization.

In order to examine the influenceof the tempering itself upon theimpact notch values, the samples withthe forgeability degree

- 15 -

VG2 5

WEG

-60 -20 +20 +60Temperatur,°C

Temperature °C

VG= degree of forgeabilityWB= heat treatment

•.)w,t4.)

01

4=

.0

54

3

10

2, 5 and 10 were only austenitized and chilled, once at 600°C/minuteand then at 2°C/minute - Heat Treatment D and H. The resultsof this test are shown in table 4.

Hardness at

- 16 -

Table 4

Testing Heat Treatment heat Treatment HTemperature Forgeability Degrees (VG) Forgeability Degrees

°C 2.5 10 2.5 10

+ 60 7.9 6.1 11.8 7.5+ 20 3.9 4.1 11.0 5.0-20 1.6 1.6 5.5 1.6- 60 0.7 0.9 1.8 1.0

+ 20°C 240 230 160 155 HB

Compared with the tempered samples, heat treatment D shows a definiteembrittlement due to the rapid chilling after austenitization and to thecorrespondingly higher hardness as shown. Under heat treatment H with slowcooling after austenitization the values of the impact notch strength arepractically equal to the values shown in the diagrams of figure 9.

The impact notch test was tested at various temperatures, also forthe case of a medium spedd of chilling after austenitization and tempering,heat treatment I. The chilling speed selected were 30°C/minute afteraustenitization and 5°C/minute after tempering. The attained values of theimpact notch strength were in the main between the values of the extremechilling speeds and showed the same tendency as already shown in figures7 and 9.

Because the official testing of the impact notch strength of the Mn steelswas carried out mostly at - 40°C, figure 11 shows for better clarify a diagramindicating the direct dependency of these values on the degree of foregeabilityat - 40°C.

The curves shown correspond to the different heat treatments. In group oneof the heat treatment, A and C, quick cooling after austenitization, there isa clear afference in the values of the impact notch strength. After a slowcooling these values are lower in all foregeability degrees after cooling[0.1°C/minute].

WB-G

WB -A

WB-E

2,5 4 6 10Verschr6edungsgrad

Forgeability Degree

Figure 11: Influence of kind ofheat treatment (WE) and ofdegree of foregeability onimpact notch values at -40°C.Radial samples, Charpy B.

- 17 -

C.4

E2220

VGIO

VG 25B-At

•

rd.u I, 16

•

-61w. 4•x :p12.ca4j 0 / 0

•

n 8w

•

,cti E 6

" 4u w2

CJCU

Temperatur,°C

WB-G VG

VG 10

-60 -20 +20 +60

Temperature °C

Figure 12: Impact notch valuesof longitudinal samples for thedegrees of foregeability (VC)2,5 and 10 and heat treatment(WE) A and G, Charpy V.

Even less favorable values ace obtained after slow cooling [2°C/minute]after austenitization, heat treatments E,G. The cooling speed aftertempering has practically no influence on these values of the impactnotch strength, these values steadily drop with the higher degree offorgeability.

The diagram in figure 12 shows the values of the impact notch testof the samples taken in longitudinal direction for forgeability degrees2,5 and 10. Regarding the heat treatment, the extreme cooling speedswere selected, that is the fastest ones after austenitization and temperingand the slowest ones.

The high values of the impact notch strength can be seen up to -60°Cat heat treatment A, particularly with the foregeability degree 10. Atheat treatment G where the slow cooling speed causes a certain embrittlementthe higher values are maintained only at the upper temperatures anddroprapidly at lower temperatures.

Examination of the Structure

In consirlering the structural properties a better fine granulation was noticedin the samples of the first group of heat treatments A to 0. The shorter holdingtine at 920°C and the quicker cooling after austeniti.zation make possihle thebuildup of the fine structure. The granule size of this group was about 9-10ASTM units, whilethegranule size after the slot.; cooling of the samples,[heat treatment E to Il] was slightly coarser and In the neighborhood of 7ASTM units. Figure 13 shows the structures after heat treatment A and E

Heat Treatment A c Troatme

Figure 13. The mic rostructure after heat treatment A :ilid L.ENO3. [ 500 : l

The ciferences in the values ol the radiai samples discovered in thetest of the impact notch strength arid caused from the side of the forgea-bility degrees also may be explained by the line structure present. Figure14 illustrates the line structure hy segregation strips.

The examination of the structure under the electron microscope waslimited to observation of the fracture areas and of the secondarystructure. Figure 15 showsfractogrophicpictures of a plastic and of abrittle fracture [irnpact notch strength 12 and 1 kpm/cm21.

Plastic Fracture

Figure 15: Fracture pictures of impact notch tests [8500:11

- 19 -

brittle Fracture

The fine segregations were discovered with the electron-microscopicexamination of the secondary structure in btittle samples of the heattreatment groups E, F, G at the forner grain borders of the austenite.Figure 16 shows such segregations which also cross the ferrite borders.

8500:1

- 20 -

2_001:1

Figure 16.- Fine segregations at the former grain limites of the austenite.

These segregations are revealed even more in the brittle impact notchsamples tested at low temperatures of -40°C and -60°C. The intensity ofthe sane samples tested at room temperatures was notably smaller. Thisphenonmenon requires additional oetailed investigations.

At an enlargement of 22,0PG times the segregations have dot-likeforms, figure 16, right, and remind one of the aluminum nitrides shownin the literature [9]. The good possibility for their segregationis obtained by slow cooling after austenitization, particularly inthe rnnge around 700°C [10,11j, providing part thereof already wasdissolved in the austenite. This also might be a reason of themore brittle behavior of the samples of the heat treatment groupE, F, G with relation to group A, 8, C.

Tearing Tests

A number of tear samples was tested to examine the influence of heattreatment and degree of forgeability upon the yield point values. The tear sampleswere taken in longitudinal and radial direction, actually after heat treatmentA and G, that is after the highest and the lowest cooling speed after theaustenitization and tempering. heat treatment I still was selected as a meancooling speed. The values of the mechanical properties are compiled in table 5.

The results of the tear test show that the faster cooling after theaustenization [heat treatment A] produces the higher strength values whichare close to the upper limit for this Mn steel type. On the contrary, theslow cooling shifts the solidity values toward the lower limit.

The constriction values are slightly higher for heat treatment Athan for heat treatment G. With the increasing degree of forgeability theconstriction values in radial direction become steadily lower than in the

- 21 -

Cooling Speed, °C/minute

Figure 17: Yield point values as afunction ofthe cooling speed

mechanical Properties

Mechasische Eigenschaftea Table

TabelleStrength Elongation Constriction

- es ig-kett- 2

591658,659,o6o,459,358,359,35918

52,45215

1 52,51 52,352,451,852,752,5

53,954,253,653,1

L = LbagsLongitudinalR = Radial

-22-

SUMMARY

Out of more than 7000 tons of forgings which were needed for thehydroelectric power plant "Djerdap" on the Danube, about 65% consistedof weldable Mn steel whose properties were examined at lower temperatureswithin the scope of this paper.

The results of the investigation showed that the forging parameters,degree of forgeability, exert a strong influence on the impact notch valuesin radial direction.

The lowest degree of forgeability [2,5] in combination with the highestcooling speed after austenitization [600°C/minute] and after tempering [100°C/min.] produced the highest impact notch strength at room and at low temperature.The contrary was true with regard to the highest degree of forgeability [10]and the lowest cooling speed after austenitization [2°C/minute] and aftertempering [0.1°C/minute].

The eensitivity of Mn steel against tempering embrittlement isdetermined on a moderate scale in samples quickly chilled after theaustenitization. Following slow chilling the tempering embrittlementpractically did not manifest itself.

With electron microscopy fine segregations were discovered at sampleswith low cooling after austenitization at the former austenite grainlimits. By appearance these segregations resemble Al nitrides. Theintensity of these segregations is particularly high for brittle samples withlow impact notch strength.

The results of the examination of the tear samples showed a directdependency of the yield point values on the cooling speed after theaustenitization. This dependency was established in diagrammaticalpresentation.

- 23 -

LITERATURVERZEICHNIS= B IBL IOGRAPHY

1/ V.I.Miheva .Gidridi perehodnih metalov.Akad.Nauk,Moskva,1960.

2/ G.Bandel1H.Haumer. Stahl und Eisen 1964,H.15,S.952/47

3/ P.Opel1C.Florin,F.HochsteinIK.Fischer. Stahl und Eisen19701H.91S.465/75.

4/ W.Steven, Journal Iron Steel Institute,193/1959,S.141

5/ P.A.Restaino10.J.Mc Mahon.Transaction of the ASM,vol.60119671S.699/706

6/ Temper EMbrittlement in Steel. ASTM Spec.Publ.nr.407

7/ I.Comon,P.Bastien.Le traitement thermique de qualitédes pieces de forge. Convegno Internazionale dellaFucinatura,19701Terni

8/ R.Liston1G.P.Smedley.Record of proceedings and discus-sion1S.11/20.International Forging Conference,Shef-field11967.

9/ J.WyszkowskilA.TereszkowskalJ.Konarski. Stahl und Eisen1969111.14,3.768/772

10/ P.König1W.Scholz1H.Ulmer.Archiv fUr das Eisenhttten-wesen.1961,H.81S. 541/56

11/ A.N.Morozov.Vodorod i azot v stali.Moskva 1968

-24-

CONVERSION TABLE FROM THE METRIC TO THE BRITISH SYSTEM

UMWANDLUNG DER METRI6CHbli IN- DIE ENGLI3OHE NJUSSE

-25--

ABSTRACT

IMPROVEMENT OF PROPERTIES OF BIG HIGH STRENGTH TURBINE DISCS

M. Kroneis and associatesBohler Bros. & Co. Ltd.Kapfenberg, Austria

By means of the destructive testing of a 16 ton upset forged disc 2000 x600 mm, manufactured from an ESR-ingot of steel 25 NiCrMo 1452, informationregarding the properties over the complete cross section and at various direc-tions of the disc is obtained from samples and the results are compared withthose obtained from a disc manufactured from a conventional ingot. Theseinvestigations were necessary in order to obtain exact data regarding theproperties over the complete cross section of the disc and in different direc-tions of loading, since from the exeptance tests conclusions regarding theproperties over the cross section can be drawn only if fundamental knowledgeregarding the trend of the properties over the cross section is available. Onaccount of the missing major segregation in the ESR-ingot, the disc shows norecognizable grain flow and segregation. The chemical composition varies overthe cross section only within close limits. Thus, for instance, the variationsof the carbon content remained within 0,24: 0,01%. The mechanical valuesdetermined by the tensile strength were generally higher and more uniform thanthose of the disc manufactured from a conventional ingot and proved almostindependent of the orientation of the sample. The values regarding tensilestrength varied between 102-105 kp/mm2 i.e. 145-150 psi x 103.

The differences between surface and centre of the disc in regard toyield point and reduction of area and somewhat more pronounced in regard to thenotched bar impact strength, are exclusively the result of the heat treatingeffect, which diminishes from the surface to the centre of the disc, or with theincrease of the percentage of bainite which occurs in the same direction. Thedifferences, however, are always smaller than those encountered in conventional-ly manufactured discs. From practical experience, as well as from the thoroughinvestigation it can be taken that bigger forged discs, manufactured fromESR-ingots offer considerable improvements not only in respect to the cleanli-ness of the steel but also in respect to the higher mechanical values and theiruniformity in comparison to ingots made by conventional methods.

Finally, an investigation regarding the release of residual stresses in big

discs has been conducted by means of relaxation tests. The results have shownthat, in addition to the Mo content, the structure excerts a profound effect onthe release of the residual stresses. In the area of greater percentages ofupper bainite, the release of the stresses takes place at a slower rate.

CONTENTS

1. Major segregation, grain flow and effect of heat treatment in forged discs.

2. Means to improve service properties.

3. Advantages of electroslag remelted ingots.

4. Investigation carried out on 16 ton disc manufactured from electroslagremelted ingot.

5. Results and improvement of overall properties.

6. Comparison between electroslag remelted ingot and conventional ingots.

7 . Residual stresses, effects of molybdenum and heat treatment.

IMPROVEMENT OF PROPERTIES OF SIG HIGH STRENGTH TURBINE DISCS

M. Kroneis - E. Krainer - H. Hojas - A. LoidlBohler Bros. & Co. Ltd.Kapfenberg, Austria

The main subjects of the Forgemasters Meeting are the developments andexperiences gained at the manufacture of big forgings. This report is con-cerned with the improvement of the properties of big, forged turbine discs,manufactured from electroslag remelted ingots.

The firm of Messrs. Bohler has carried out extensive development work inregard to forging processes and electroslag remelting as well as metallurgicalmeasures to be employed for the manufacture of highly stressed discs. The workand the results obtained have been reported on several occasions 1) 2) 3). Themechanical properties and their uniformity over the entire cross section or incertain selected areas are of paramount importance in big turbine discs sub-jected to high stresses in service. Since acceptance tests utilize onlysurface samples for the evaluation of the quality of the discs, the propertieswithin the discs can be judged only by indirect means. It was for this reasonthat systematic investigations became necessary to give a clear picture regard-ing the properties over the cross section of discs in the various directions ofloading 2). For the definition of the properties of a disc the following con-siderations should be listed:

1. The flow of the fibres is determined by the upset forging process andfollows a non-uniform but systematic pattern as shown in Fig. la.Zone "a" near the face has been deformed very slightly or not at all.In this zone conditions exist as were relevant in the starting mater-ial before the upsetting process. In zone "b", the upsetting resultsin a pronounced radial grain flow towards the outer circumferencewhereby considerable diversity prevails regarding the degree of defor-mation. The outermost zone "c" receives a relatively uniform degreeof deformation in the direction of the indicated fibre. In zone "a",which is subjected only to slight deformation, generally the defectsof the starting material persist and voids will be closed only withdifficulty. Zone "b", containing the central segregation of the forg-ing ingot, generally shows defects of a more extented type, causedeitherby segregation or non-metallic inclusions. It is in this zonethat ultrasonic inspections will reveal the greatest number ofdefects.

2. The heat treatment effects are of decisive importance for the proper-ties of a disc. A diagrammatic representation of the heat treatmenteffects on a disc with parallel faces is given in Fig. lb. Due to thefaster rate of cooling, the outer zone "1" will contain a higher per-centage of martensite than zones "2" and "3" in the interior of thedisc. The latter contain high percentages of bainite whereby, depend-ing on the type of steel, zone "3" will contain upper bainite. In

case of faulty heat treatment or if the alloy content of the steel istoo low, zone "3" may contain also certain amounts of ferrite.

3. The properties of the discs are attained by the joined effect of grainflow and heat treatment. The grain flow, modified by the forgingprocess and major segregation, is superimposed by the heat treatingeffect. The latter diminishes from the surface to the center of thedisc. The resulting typical properties in various places and direc-tions of loading of discs obtained by conventional methods have beendiscussed on several occasions. Within zone "a" (Fig. la) the heattreating effect creates favourable conditions but a low degree ofdeformation may have an adverse influence on the properties. In zone"b", inferior properties will generally be encountered. This is dueto major segregation and the slighter heat treating effect. Theductility and toughness of axial samples are very low but also radialand tangental samples have inferior properties. Zone "c" suffers noadverse effectsby major segregation, and, in addition, the good heattreating effect permits highest mechanical values to be obtained inall directions of loading. The results of the mechanical testing ofsamples taken from only one definite zone are not representative forthe properties encountered over the entire cross section of the disc.

For the improvement of the mechanical and service properties of suchdiscs, the previous considerations indicate several possibilities.

A. The chemical composition of the steel must be such that for a givendisc, quenching in oil or water results in a ferrite free structureover the entire cross section. The structure should contain a minimumof upper bainite which is usually of a coarser nature, and similar toferrite, may have an adverse influence on the properties. Beneficialmeasures comprise machining of the discs to within close machiningallowances prior to heat treatment and quenching in water. The molyb-denum content of the steel should be balanced to assure avoidance oftemper brittleness and prevention of adverse residual stresses afterhardining and tempering. Higher contents of molybdenum may berequired for the attainment of greater high temperature strengths.

B. The forging process must be carried out in such a manner that theeffects of unfavourable deforming conditions in the zones "a" and "c"are reduced as far as possible.

C. Since major segregation of the ingot is directly responsible for themechanical properties and their aniaotropy, it must be brought down toits lowest limit.

D. On account of the pronounced effect of non-metallic inclusions on theservice properties of turbine discs, the cleanliness of the steelbecomes a major consideration.

At the conventional manufacture of discs from big forging ingots, theconditions listed under C and D cannot be met satisfactory. Difficulties dur-ing the manufacture and differences of opinion regarding internal defects andinferior mechanical properties, therefore, cannot be avoided. However, in thecase of big turbine discs with high yield point, greatest demands are placed inregard to the absence of non-metallic inclusions as well as in respect tosuperior and uniform properties over the entire cross section.

In this respect, the special advantages offered by the electroslag remelt-ing process permit the manufacture of discs with improved service propertiesresulting from enhanced mechanical properties and greater homogeinity.

According to the experiences gained by Bohler in the manufacture of biggerforged discs of electroslag remelted steel, the mechanical properties demandedby specifications can be maintained with great safety. In addition, the greatcleanliness of the steel, established by ultrasonic testing, shows a distinctimprovement resulting in greater safety during production and improved homo-geinity of the material. For complete evaluation of the quality of a disc, theproperties in several directions of loading must be known for the entire crosssection. Investigations, carried out on a big disc of steel 25 Ni Cr Mo 1452,manufactured from a conventional ingot, have been reported during previousForgemasters Meetings 2). By means of a disc with a diameter of 1600 mm and athickness of 570 mm, heat treated to 105 kp/mm2, it was shown that the valuesregarding ductility and toughness over the cross section of the disc depend onthe heat treating effect, the characteristic segregation and grain flow, andmay vary considerably. Investigations, which were carried out on shafts anddiscs of electroslag remelted material of lower as-heattreated strengths haveestablished the greater uniformity of the properties over the cross sectioninfluenced by the heat treatment 3).

The present report covers the investigations which were carried out on aforged disc of electroslag remelted steel of 2000 mm diameter and 600 mm thick-ness, having a weight of 16.3 metric tons. The mechanical properties andcharacteristics of this disc will be compared with those of the disc mentionedin the previous report, the latter disc having been manufactured from a conven-tional ingot.

Manufacture and Testin of Disc

The chemical composition of the steel as well as the production method andtesting procedures of the disc are given in Table I. The steel was melted in abasic electric arc furnace and cast into electrodes after vacuum treatment.The remelting of the electrodes was carried out in a plant featuring a liftingmould. Six electrodes were required for the manufacture of the ingot having adiameter of about 1000 mm and a length of 2700 mm, resulting in a weight of theingot of about 18,000 kilograms, The shape of the disc was obtained by upsetforging of the ingot. No drawing out of the ingot preceded the upset forging.Thediscwas then subjected to isothermal annealing and cooled to room tempera-ture. Annealing was followed by ultrasonic testing. Thereafter, the disc,after austenitization, was quenched in water and tempered to a tensile strength

of 105 kp%mm2. The heat treatment provoked no cracks or surface defects. Asecond ultrasonic inspection carried out after partial machining, also revealedno defects. This was in line with the results obtained in routine productionof forgings from electroslag remelted ingots, and proved, that even with over-sensitive adjustment of the ultrasonic equipment, no defects could be detected.Next, the disc was parted in axial direction and hot etch samples were pre-pared. One-half of the disc was subjected to another tempering operation andits strength reduced to about 90 kp/mm2. The determination of the mechanicalproperties was carried out on sets of tensile and notched bar impact samples.The samples were taken in tangential and axial direction from over the entirecross section of the disc as indicated in the various diagrams.

Steel 25 Ni Cr Mo 1452

This steel, containing about 3.5% Nickel in addition to chromium and moly-bdenum, is employed for forgings where greater susceptibility to hardening andtempering as well as high strength at room temperature and optimum toughness isrequired. As can be seen from the isothermal transformation diagram (Fig. 2),an as-heat treated structure free of ferrite can be obtained also in biggersections and most certainly within the disc under investigation. Regions inthe vicinity of the surface will be martensitic with small amounts of upperbainite. At greater distances from the surface, higher percentages of upperbainite must be expected. The range of the upper bainite is greatly influencedby the molybdenum content of the steel. Regarding molybdenum, a degree ofmicro-segregation of 1.5 was found to exist in the disc. This may explain thefibrous nature and heterogeinity of the properties as far as they can still bedetermined in electroslag remelted steel.

Fig. 3 shows the tempering curve of the steel determined from tangentialsamples. It shows the optimum properties which can be obtained by quenchingand tempering. The deviation from these values indicate the effect of theslower rate of cooling, the influence of segregation and position of thesamples. The highest admissible tempering temperature, according to the dia-gram, has been established to be 660° C.

Results of the Investi ation

Homo einit of Disc

The ultrasonic tests mentioned previously as well as the examination oflongitudinal macro and micro samples confirm the absence of adverse inclusionsand heterogeneities and show the superior cleanliness of the steel. The controlof the chemical composition over the cross section confirmed the absence ofmajor segregation. Within the possible limits of deteomination, the carboncontent varied only between 0.24 and 0.25% while the molybdenum content showedvariations only between 0.43 and 0.45%.

As-Heat Treated Structure

Microstructures encountered in different positions in the disc are shownin Fig. 4. The as-heat treated structure is entirely free from ferrite andcontains percentages of bainite. In the center of the disc, the transformationresulted in a purely bainitic structure. It can be seen that the bainiticstructure in the center is of a coarser type. An influence of the segregationcould not be established by metallographical means.

Mechanical Pro erties

Table 2 shows the results of the mechanical tests for the higher tensilerange while Table 3 has been compiled for the lower strength level. The posi-tion of the samples taken can be seen from the accompanying sketches. Allsurface samples have been taken from places 40 mm below the surface of the disc.

The figures give a complete picture regarding the properties encounteredin the cross section of the disc and in respect to all relevant directions ofloading. A comparison of the properties and characteristics of the discobtained from the electroslag remelted ingot with the results obtained pre-viously from a similar disc, manufactured from a conventional ingot 2), give aclear picture of the improvements which were attained.

Fig. 5 permits a comparison to be made of the mechanical properties oftangential samples of the disc made from electroslag remelted stock and a discmade by conventional methods. Both discs were heat treated to the higherstrength level. The values of the 0.2% offset yield point and ultimate strengthof the electroslag treated disc as compared to the conventional disc indicatethe high degree of uniformity of the first. The same applies for the valuesregarding elongation and reduction of area. In the disc manufactured fromelectroslag remelted steel, all values are superior to those obtained in normaldiscs and even in the most unfavourable conditions never drop to the low limitswhich must be expected in conventionally produced discs. Thus, samples takenfrom the non-deformed central region 31, possess extremely low values regardingreduction of area in the case of conventional discs. At the disc, manufacturedby the electroslag process, samples taken from the same localities, possess amaximum reduction of area. Fig. 6 represents the same data for axial samplesat a higher strength level. Here again the values regarding the 0.2% offsetyield point and tensile strength are more uniform. A very considerable improve-ment in respect to elongation and reduction of area, however, can be noted inelectroslag remelted material. Samples taken from the positions 11 and 13 inthe center of the discs, made from conventional ingots, generally possess verylow ductility. In discs, made from electroslag remelted stock, this criticalposition is still discernible but the respective values never drop to extremelylow values.

The conditions existing at the lower strength level of about 90 kp/mm2 aregiven in Fig. 7. Here, the superior advantage in respect to the uniformity ofthe properties of the electroslag treated material is the most conspicuous.Elongation and reduction of area remain at their top levels in the samples

taken from all positions. This applies also to position 31 at the face of thedisc, which, in the case of conventional discs, possesses inferior ductility.The diagram also shows that over the entire cross section, elongation and reduc-tion of area reach the maximum values, which, according to the tempering diagram,can be obtained in the steel under discussion.

While the typical data regarding ductility obtained by the static tensiletest are considerably improved by the electroslag remelting process, the im-provements regarding notched bar impact strength and fracture appearancetransition temperature do not quite reach the same extent. To give a clearpicture of the existing conditions, Fig. 8 relates reduction of area and transi-tion temperature to the as-heat treated strength of samples taken from twotypical positions. In this manner the variations brought about by the electro-slag treatment can be recognized. As has been mentioned before, for bothstrength levels and positions of the samples, the electroslag remelted materialpossesses considerably higher values regarding reduction of area. However, onlya small reduction of the transition temperature can be observed in the mostcritical zone, i. e. in the center of the disc produced from electroslag re-melted ingots. In the outer zones, corresponding to position 35, which are themost influenced by the heat treatment, somewhat less favourable, i.e. highertransition temperatures have been determined. These higher transition tempera-tures can partly be explained by the somewhat altered position of the sampleswhich were taken from a greater distance from the surface.

By the investigations and subsequent comparisons between conventional andelectroslag remelted steels, a number of typical properties have been ascer-tained for discs manufactured from electroslag remelted ingots (Fig. 9).

The disc made from electroslag remelted steel excels by a uniform carboncontent over the entire cross section (Fig. 9A), whereas a disc made from conven-tional steel contains considerable segregation both in regard to carbon andalloying elements. For this reason, in electroslag remelted steels, the as-heattreated strength of a disc depends primarily on the rate of cooling duringquenching and thus on the dimensions of the disc.

In the case of high strength heat treatable steel, the disc made fromelectroslag remelted ingots shows tensile strength values between 102 and 105kilmm2 in tangential samples while the disc made from a conventional ingotshows tensile strength values between 105 and 115 kp/mm2 (Fig. 9B).

The ductility of electroslag remelted steel, represented by the reductionof area never drops to the low values experienced in the zones of segregation ofdiscs manufactured from conventional ingots.

In discs made by the electroslag remelting process, the scatter of thevalues regarding reduction of area remains between 64 and 52% while convention-ally produced discs show a scatter of 53 to 33%.

For the transition temperature of the notched bar impact strength, somewhatlower temperatures have been determined for the center, while in the outermostzones of electroslag remelted discs the low transition temperature of discsmade from conventional ingots are not quite reached. The results do not permitto decide whether this represents a typical behaviour of the ESR-steel under im-pact loading. It might be assumed that the influence of the structure super-sedes the influence of segregation and grain flow.

Reduction of Residual Stresses

Residual stresses in highly stressed forgings such as turbine discs may beadvantageous if the residual stresses occur in a controlled and desirablemanner 6). However, they may be the cause of breakage if they are not control-led and surpass certain limits.

The residual stresses, which exist in a forging after the heat treatment,are determined primarily by the temperature cycle of the heat treatment wherebytransformation stresses prevail. By the tempering operation, the stressescaused by the hardening operation, are partly released, and, if the rate ofcooling after tempering does not exceed certain values, no additional stressesoccur 5). The chemical composition of the steel exerts a double influence onthe residual stresses existing after the heat treatment. It influences thetransformation behaviour of the steel and the high temperature yield point.By selecting the most suitable steel and controlling the rate of cooling afterhardening and tempering it should be possible to influence the residual stresseswithin certain limits 6). The release of stresses during the tempering opera-tion is closely connected with the creep strength of the steel and the rate ofcreep at tempering temperature. In this manner, the release of stresses can beinfluenced within close limits directly by the duration and temperature of thetempering operations if the rate of stress relief in dependance of the tempera-ture is known.

For the respective investigations, we have employed relaxation tests.Relaxation tests at constant creep rates do not correspond entirely to the cond-itions existing in a forging during tempering since distortions are possible ina forging during tempering. However, relaxation tests are well suited to re-produce the conditions during the short periods corresponding to temperingcycles.

Relaxation tests have been carried out at temperatures of 500, 550 and6000 C with a constant initial creep of 0.5%. Two laboratory heats of steel25 NiCrMo 1452 with various Mo contents of 0.3 and 0.7%, respectively, as wellas samples taken from the disc have been investigated. The samples from thelaboratory heat and a sample taken from the disc have been heat treated individ-ually to a martensitic structure of about 105 kp/mm2 tensile strength.

Fig. 10 indicates the ratio between the residual stressffR to the initialstress 6'A after ten hours of loading versus the test temperature.

From a comparison of the three individually heat treated satples, theknown influence of the molybdenum content is clearly discernible. In the samemanner, the creep rate is influenced by the carbon content whereas variationsof the chromium and nickel content within some points have scarcely any influ-ence. The release of stresses however, is considerably influenced by the typeof structure as can be taken from Fig. 10 where the results obtained from asample, taken near the surface (35) and a sample taken from the center (11) ofthe disc confirm the data given elsewhere in literature (8). As long as theferrite content is negligible the upper bainite possesses the highest creepstrength. The creep strength diminishes steadily via the lower bainite to themartensitic structure. The influence of the structure may completely supersedethe effects of variations of composition and segregation.

From these results it can be deduced that in bigger discs of steels,generally used for the purpose, the release of stresses over the cross sectionduring tempering takes place at various degrees. In discs of high tensilesteels which contain upper bainite and no ferrite in the center and which aretempered at low temperatures, the relaxation of stresses in the center is re-tarded. The adjustment of the equilibrium stresses in the disc will thus beconnected with a variation of the stress distribution and deformation.

These considerations become more important if, with increasing yieldpoint of the disc, the tempering temperature approaches 600° C or some temper-ature below.

Summary

By the destructive testing of a 16 ton upset forged disc, manufacturedfrom an electroslag remelted ingot of steel 25 NiCrMo 1452 informations wereobtained regarding the properties over the complete cross section of the discfor various orientations of the samples. The results were compared with thoseobtained from a disc manufactured from a conventional ingot. Due to theabsence of major segregation in the ESR-ingot, the disc has no recognizablegrain flow and segregation. The mechanical values determined by the tensiletest, are without exeption higher and more uniform than those of the discmanufactured from a conventional ingot and show almost no dependance on theorientation of the sample. In respect to yield point and elongation, thedifferences between surface and center of the disc are somewhat •more pro-nounced. This applies to a larger degree to the notched bar impact strength.The more pronounced differences are exclusively the result of the heat treat-ment effect which diminishes from the surface towards the center of the disc andresults in increasing percentages of bainite in the direction towards the center.The differences in the mechanical properties of the electroslag remelted steel,however, remain always smaller than those encountered in discs manufactured byconventional methods.

Finally, an investigation was carried out by means of relaxation tests re-garding the release of residual stresses in big discs. The investigations showthat, in addition to the molybdenum content, the condition of the structure de-termines to a large degree the reduction of residual stresses. In the areascontaining higher percentages of upper bainite release of stresses takes placeat a slower rate.

References

1. M. Kroneis, ETUDE DU FORGEAGE DE DISQUES PAR REFOULEMENT, Revue deMetallurgie 1962, p. 953

2. M. Kraneis, E. Krainer, Th. Skamletz, Riegler, Betrachtungen aerHerstellung und Eigenschaften groBer Turbinenscheiben fUr sehr hohemechanische Beanspruchungen, Internationale Schmiedetagung 1965, Berlin

3. M. Kroneis, E. Krainer, H. Hojas, Th. Skamletz, Eigenschaften groBerSchmiedestUcke aus ESU-BlOcken, Internationale Schmiedetagung Terni 1970

4. B. Cima und P. Jubb; Factors affecting the transition temperatur offorgings used in powergenerating equipment. Journal of the Iron andSteel Inst. Dec. 1959, pp. 329/43

5. H. Bailer, A. Rose, Archiv 50 (1969) pp. 411/23

6. H. Wolf, W. Ofihm, Archiv 42 (1971)pp.195/200

7. F. Clemens, P. Hammerstein, H. Imgrund, Archiv f. d. EisenhUttenwesen 41(1970) pp. 231/33, Heat Resisting Steels and Alloys, C. G. Conway,George Newaes, Ltd., London 1953, Sect. 3

B. F. B. Pickering, Iron and Steel May 1968, pp. 206/09 Norton J. F.,A-Strang, Journ. Iron and Steel Inst. February 1969, pp 183/203.

Table 1: Disc of 25 NiCrMo 14 5 2; chemical composition, manufacture andtesting.

1. Chemical Com osition: C 0.26%, Si 0.0O%, Mn 0.39%, P 0.008%,S 0.006%, Cr 1.30%, Mo 0.45%, Ni 3.55%, V 0.12%, W less than0.02%, Cu 0.14%, Sn 0.018%

2. Manufacture of Steel: basic electric arc furnace, vacuum treatedduring tapping

3. Electrosla Remeltina) cast electrodes 500 mm V, 2000 mm long - 6 piecesb) ESR-ingot, polygonal 1000 mm x 2700 mmc) weight 17 tons

4. Forging of Disc: Upset forging without drawing out

5. Heat Treatment After For in : Air cooling to 400°C, soaking for20 hours; heating to 680°C, soaking for 20 hours, cooling infurnace

6. Heat Treatment: a) Quenching from 860°C in waterb) Tempering at 610°C, soaking for 20 hours, air cooling,

stress relief annealing at 580°C, slow cooling in furnace(higher strength level)

c) Parting of disc in longitudinal direction, taking of hotetch disc extending over complete longitudinal cross section

d) Second tempering of one half of disc at 6400C, soaking for20 hours, air cooling, stress relief annealing at 580°C,slow furnace cooling (lower strength level)

7. Testing: a) Etching in 50% hot HC1 as well as Oberhoffer-etching over

the complete longitudinal cross section of the discb) Ultrasonic inspectionc) Tensile- and notched impacted bar tests on radial, tangential

and axial samples. Determination of fracture appearancetransition temperature (FATT)

Table 2: Disc 25 NiOrMo 14 5 2: Mechanical Properties, Higher Tensile StrengthLevel

Position of test pieces

16.0 50 2.5 I 70

16.0 52 2.8 60

' 16.8 56 3.5 20

15.0 51 2.9 60

17.2

15.4

18.4

16.0

16.0

17.6

18.2

18.4

. 31 33 35r 3 21 25

11 0 13 15

2000

12.0

15.6--r

16.0 I

14.0

15.4

F 16,0

17.0 62

16.2 59

17.4

15.6 48

58

50

50

58

48

51

56

62

62

3.5 20

2.9 I 20

3.2 0

4.6 , -10

2.9 60

2.7 60

4.2 10

2.7 60

3.8 20

3.8 20

4.8 -10

5.0 -10

36 1.9 +80r-

2.0 : 70

3.1 :I20

51 2.5 20

50 2.4 20

54 4.0 ; -10

Table 3: Disc 25 NiCrMo 14 5 2: Mechanical PropertiesLower Tensile Strength Level

CO

fl


Position Test 0.2%- Tensileof test piece Yield 1 Strengthpieces Nr. Strength

kilmm2 kp/mm2

11 R

31 R

35 R

11 T

31 T

35 T

81.2

77.1

78.7

78.7

79.6

84.1

Elong.

31

61120000 I

35

Reduction Charpy-V FATTof Area

mkp °C

93.0 17.8 60 4.2 0

89.8 20.8 64 7.8 -20

90.4 19.6 67 8.8 -30

91.1 I 18.2 60 4.0 0

91.1 ' 20.2 66 5.6 -20

93.6 1 21.0 67 11.3 -50

Fig. ls Forged Disc,Grain Flow and Effect of Heat Treatment

3

Fig. 2: Steel 25 NiCrMo 14 5 2,TTT-diagramTemperature of Austenitization 840°C

BeginnStart

10 102 /05

1

Per litPearlite

10 v2

Zeit Time

BeginnStart

104

EndeEnd

103

105SPC

min

Fig. 3: Steel 25 NiCrMo 14 5 2, Tempering diagram,Hardening Temperature 840°C, Tempering Time: 20 h,Furnace Cooling

1000 psi kp/ min? mkp ft lbs.

180 Zug fes tigkeitTensile Strength

14 100

170 120 A../• A.....--,...,... 9060...7.. —_-- . 12

160//"%

Einschnikung

15GArea

I ‘110 / 80

Redurtion o fc I01

.... C50 i 10

-6,Q) eiL 0,2 Dehngrenz e I 70

::-..th 100 0,2 % —Yield Strength I

140 (1/ VI49

'.... C S.

tz 41, I 1 I 6040N , I

1\III

1 8 Et

130 -IL th l' I 0N c 90 I cc Q, 50o, L.- I I I u

icyl „.„

120 ,2 t,--, 30 ICharpy- V /

--.1.--.140

1 I 801: 1 , 1

110 c' IN e\cS' 20 Dehnung .., \... 4 30

Elcngation •••• ••N

•••••°.\

100 70 —----- 20\--- ammo. oiroasw ayes..

10 290

1060

SOf 0500 600 700

Antalltempera tar

Tempe ri ng Temperature°C

1.3

Fig, 4: Forged Disc, Structures


Fos

Fig.5: Forged Disc 25 NiCrMo 14 5 2, Higher Tensile Strength Level,Mechanical Properties of E5R and Conventional Ingots(Tangential position of test pieces)

11 13 15 31Position

Lage der ProbenPosition of test pieces

35

Zug festigkeitTensile Strength

31 35

F131 13 15

11 13 15 31 35Position

EinschniirungReduction of Area

11 13 15 31

Position

35

Fig. 6: Forged Disc 25 NiCrMo 14 5 2, Higher Tensile Strength Level,mechanical Properties of ESR and Conventional Ingots

(Axial position of test pieces)

cony. ESU

n u1000psi kp/mm2

160

150

140

130

120

110

110

100

90

80

if') 70

20

10

0

0,2 - Dehngrenze

0,2% - Yield Strength

11 13 15 31 35

Position

Dehnung

30 Elongation

11 13 15

Position

Cage der ProbenPosition of test pieces

60

40

20

0

ZugfestigkeitTensile Strength

31 35

P11 13 15

11 13 15 31 35

Position


31 35 11 13 15

Position

31 35

Fig. 7: Forged Disc 25 NiCrMo 14 5 2, Lower Tensile Strength Level,Mechanical Properties of ESR and Conventional Ingots(Tangential position of test pieces)

conv. ESU

oIWOpsi ko /mm2

0,2 — Dehngrenze0,2% - Yield Strength

I.

30

20

10

0

11 31 35 11 31

Position

DehnungElongation

11 31Position

Cage der Proben

Position of test pieces 0 11

35

60

ZugfestigkeitTensile Strength

11

Posifion


31Position

31 35

35

35

Fig. 8: Forged Disc 25 NiCrMo 14 5 2,

Reduction of Area and FATT,

Comparison of ESR- and

Conventional Ingots (Tangential

position of test pieces)

°C

Pos.31S 3/

4.4%

.%**

****

% -

- Fbs.35

Loge der Proben


«80

4,40

31

35

Pos.11

It

Pos.35

ie

-40

ie

ieew

e.•

••••

•

4.40

4maE

SU

A

----

co

mp:

80

80

90

100

110

120

80

90

100

110

120

Zug

fest

igke

itin

kp

/m

m2

Ten

sile

S

tren

gth

Fig. 9: Forged Disc 25 NiCrMo 14 5 2, Comparison of Lines of EqualCarbon Content, Tensile Strength, Reduction of Area and FATTfor ESR- and Conventional Ingots (Tangential position of testpieces).

Cony,

C in

1 II

9 •e"n_ . „ 0 ....... ot• ,....• / /07 OP, .../ _ ......" ,...• - , / 9 i AI / or" 0,13.1......C..... .."' Yeola I

0/ t,"/ 4.)7IL / / t...1_,

109

.79s

0 ° 80°C

a24 I 0,01

Zugfestigkeit, Tensile Strength in ip/htm2

FATT in 0C

!02-105

Einschnurung, Reduction of Area inA

SO o

20

NA‘, k

,

f";

/ • 7 /

.

7

VI0,

I

/ \

rcj

CO C1/41

• ( oi-5"ji) y co ) 511noti 0! jai; .011(_, 0

unwie yo JO r7 q 01 epcu /Hp

T-

\ f;.;

0 0

01

4.1 1.•1 ;,`' 1.0

;('' la_

ri tC3 LS ca

-1/

ABSTRACT

Steadily increasing energy requirements are demanding higher stressed andlarger sized low pressure turbine and generator rotors. This implies thenecessity of forgings having greater strength and fracture toughness, controlledresidual stresses and improved homogeneity.

Comparative microfractographic examinations were performed in the zone ofmaximum inhomogeneity of rotor forgings which occurs approximately at the one-quarter diameter. These examinations and tests included samples obtained fromboth large and small forgings produced in accordance with past and present dayforging practices. These results were further compared with samples obtainedfrom forgings produced from very large "A. P. Processed" ingots.

Representative internal samples were taken from selected large crosssection forgings and examined for fracture toughness properties, nil-ductilitytemperature and crack growth characteristics and for the influence of inhomo-geneities on these mechanical properties.

These test results were then compared with the estimated stress concentra-tions caused by inhomogeneities by the application of fracture mechanics andsubsequently included in a suitable form into the purchasing specifications.

Finally, the test results are reported for eleven very large capacityrotor forgings produced from 400 to 500 ton ingots within the framework of adevelopment program. Modern, very large forgings, especially those producedfrom A. P. processed ingots exhibit less inhomogeneity and as a result a greaterresistance to fracture than smaller forgings produced by more conventionalproduction techniques. There is future hope that the electro slag remelt pro-cess will yield similar favorable results.

CONTENTS

INTRODUCTION 1

DIMENSIONS OF LARGE CAPACITY ROTORS

2 and 4 pole generators

Large LP-turbine shafts

OPERATIONAL STRESSES IN ROTORS 2

RESIDUAL STRESSES IN FORGINGS 5

FRACTURE TOUGHNESS AND FATIGUE CRACK GROWTH 7

Fracture toughness

Fatigue crack growth

Correlation with acceptance test

INHOMOGENEITY OF ROTOR FORGINGS

Radial trepans oriented according to ultrasonic indications

Radial trepans not oriented according to ultrasonic indications

"A.P. Process"

INHOMOGENEITY AND MECHANICAL PROPERTIES

Porosity, fracture toughness and NDTT

Porosity and crack growth behaviour

Segregation and crack growth behaviour

ACCEPTANCE CRITERIA

Properties for use

Conventional test values

TEST RESULTS OF LARGE CAPACITY ROTORS

SUMMARY AND FUTURE PROSPECTS

REFERENCES

CONVERSION TABLES

15

21

25

25

29

INTRODUCTION

Rotor forgings for steam turbines and turbine-generators have always beenof special interest to metallurgists and engineers. Whenever the metallurgyand forging technology did not follow the demands for steadily increasing energyrequirements, severe failures have occasionally occurred. The same applies tothe development and application of a satisfactory testing technology and materialstress analysis appropriate for these special pieces as a whole.

Following a short review, the present status will be described, and anattempt will be made to indicate the trend of development for the coming years.Rotor forgings made of steels intended for high temperature application willnot be discussed. (Figures, Tables, Literature References, and ConversionTables will be found in pages following the text.)

DIMENSIONS OF LARGE CAPACITY ROTORS

A survey of the development of one p ece generator rotor forgings for3,000 or 1,500 rpm operation is presented in Figures 1 and 2, respectively.The 80 MVA generator in Figure 1 was the first of its kind, and for aoroxi-mately 20 years, the largest in the world for this output and speed.1) It wasnot possible at that time to solid-forge the rotor; therefore, it was assembledfrom three pieces. As soon as the forging producers were capable of solid-forging rotors of the required size and quality, and on schedule, with few ex-ceptions the multi-piece rotor was abandoned again. The development of largerone piece rotor forgings continued. "It is rather indisputable, however, thatin the future all two-pole rotors (3,000/3,600 rmp) will be solid-forged witha good degree of safety". 2)

The limiting factor for two-pole generators of high output is not basedon the size limitations of the rotor forging; however, the stresses in thezone of the segregations are considerably higher than in the past (Table 2).The limiting factor for greater than 50 Hz/1500 MVA generators is presentlybased on the required high 0.2% yield strength of the nonmagnetic retainingrings, which cannot be achieved at the present.

Rotors of four-pole turb ne-generators are considerably more massive thanthe two-pole type. Generator rotors with 1500 rpm (50 Hz) require a 20%larger body volume than those with 1800 rpm (60 Hz), however at considerablylower stresses than two-pole generators. Here too, the development startedout with assembled rotors (Figure 2). Forgings, which are presently requiredfor large capacity generators such as in nuclear power plants, can only beproduced by very few forging producers in the world (Table 1).

For the first time in 1969, the Japan Steel Works (JSW) produced 400 ton

*SUPERSCRIPTS REFER TO LITERATURE REFERENCES FOLLOWING TEXT.

ingots to manufacture rotors of 220 ton maximum weight and 2000mm maximum dia-meter. An extensive development program has been undertaken with JSW to satis-fy present orders for 1500 rpm generators with outputs of 1000 to 1500 MVA,thereby assuring delivery schedules as well as creating a basis for possiblystill larger or higher stressed forgings in the future.

Figure 3 represents the production time schedule. In consideration ofthe acceptance target date, each production piece was followed with a backupforging. To date, difficulties have not been experienced; in fact, the ini-tially planned back-up forgings could be utilized to fill orders. As will beshown later, a uniform rotor quality has been achieved, especially with regardto homogeneity, which justifies the step-up to 500 ton forging ingots. Basedon the uniform homogeneity, it is permissible to produce a forging such thateither 1000 or 1500 MVA rotors may be machined from the same forging configura-tion. At the same time, a forging could be produced for either 2000 MVA(50 Hz) or 2300 MVA (60 Hz) rotors (Figure 4). The applications of the 500 tonforging ingot are therefore not yet completely exhausted.

In addition to generator rotors, it became necessary to manufacture lowpressure turbine rotors with a weight of 165 ton with a maximum diameter of2700mm (106"); to be utilized for two railroad turbo-sets with 1000 rpm(16-2/3Hz), (Figure 5). The operational stress of these rotors is so smallthat the specified minimum 0.2% offset yield strength of 13.5 kp/mm2 at 300°Cis adequate. These initial rotors were forged from 400 ton ingots and werealso included as part of the development program.

OPERATIONAL STRESSES IN ROTORS

Mechanical stresses in rotating shafts may be determined by proven exper-imental and analytical methods; which can state with high accuracy the stresseson axial and radial bores, on rotating outer notches such as the blade groovesof turbine shafts or on generator winding slots which are running parallel tothe rotor axis--however, always based on the assumption of homogeneous andisotropic shaft material, which in reality is never achieved.

Figure 7a represents the cross sectional stress distribution character-istic of rotating shafts with and without an axial bore. It should be notedthat the combined tangential-and radial-stresses are of equal magnitude forboth types in the zone of greatest segregation. The shaft with an axial bore,however, has a stress peak at the bore surface, which acts normal to thelongitudinal shaft axis and which has a magnitude approximately twice that ofthe shaft without a bore. Likewise this stress peak acts normal to thedirection of the segregation stringers, over their total length in full height.Figure 7b shows the stress characteristics of a radial bore which is locatedeccentrically with respect to the shaft axis. Here too, the stress peak actslocally, normal to the segregation stringers; however, it decreases veryrapidly in the longitudinal direction of the stringers.

Therefore, crack propogation will not occur by the same manner at bothaxial and radial bores. Cracks may originate at the zone of the stress peaksduring long-term operational stressing, or at material-inhomogeneities whichhave been in existence from the beginning in the vicinity of the stress peaks.In the case of axial bores, the cracks may extend throughout the total shaftlength in a uniformly high stress zone, whereas cracks at a radial bore willpropogate into a zone having a much lower stress level and consequently beretarded, thereby not leading to shaft failures.

Figure 7a represents for generator rotors only the stress distributionfrom the rotor centerline to the bottom of the winding slots. The stresspeaks at the slot bottom of two-pole generator rotors are of secondary im-portance, since the maximum stress is always located within the core of therotor because of the large centrifugal stresses. In contrast, four-polegenerator rotors are most highly stressed in the zone of the winding slots(refer to Table 2). However, this is a zone in which the forging does notdisplay inhomogeneities because of segregations.

The crucial task in determining the usability of a rotor consists ofjudging the effect of the operational stress peaks combined with the realisticconsideration of superimposed stresses, such as

a) stress concentrations caused by material-inhomogeneities

and

b) residual stresses resulting from the forging heat-treatment process.

The effect of superimposed stresses caused by material-inhomogeneitiesmay be taken into account with the aid of the linear elastic fracture-mechanics.A portion of the inhomogeneities is indicated by the ultra-sonic test of theforging. As long as detailed information is not available about the actualcause of such indications, the first evaluation would point towards materialseparations, which had frequently occurred before in the shape of hydrogenflaking-cracks, prior to the introduction of the vacuum-degassing process.Since experience has shown that the indications are often related to elongatedsmall segregation stringers within the forging, the first evaluation is furtherexpanded on the possibility of existing elongated, elliptical cracks with theaxial ratio; crack-depth a/crack-length 2c <0.1. The indication therefore, isbased on a "worst case flaw geometry" and the size of the worst case flawderived from the equivalent flaw-size of the ultrasonic indication.

For geometrically simple crack-formations such as elliptical cracks, thestress-intensity K at the crack tip may be calculated with the aid of linear,elastic fracture-mechanics 3,4); this stress intensity results from operationalstresses and the shape of the crack. In actual operation, brittle fracture ofthe rotor is not permitted to occur; therefore, the only question of practicalinterest pertains to the critical flaw size for brittle fracture. The criticalflaw size acr depends on the operational stresses of the shaft, the shape ofthe crack and the fracture toughness Kic of the shaft material.

Figure 8 represents the formula for the critical flaw size acr, as derivedfrom fracture-mechanics; three different flaw positions are directly comparedin rotating shafts with and without bores. The crack-shape parameter 0) isassumed to be constant for this analysis. Case 1 in Figure 8 applies to afully embedded flaw in a shaft without a bore, while case 2 describes thecondition far a shaft with axial bore. As will be shown later, segregation isnot concentrated at the axis of the finished forging but rather in a cylindricalzone intersecting the approximate mid radius of the body diameter. In general,

only comparatively small axial bores can be permitted in order to maintain theshaft strength; and are therefore insufficient to remove the segregations. Onthe contrary, it must be assumed that the segregation stringer, together withthe anticipated crack have been intersected by the axial bore such that twostress peaks are now superimposed on each other, as viewed from a mechanical

stress standpoint. Consequently, the resultant critical flaw size for brittlefracture amounts to approximately 1/5 that of the permissible size for a shaftwithout a bore. Considering the uniform orientation of the axial bore, shaftstress and flaw, as described before in Figure 7, it becomes clear that thisis the worst case for practical applications ("The most critical defect forthe application is a long-shallow surface crack intersecting the axial bore ofthe rotor with the major plane of the crack normal to the tangential borestresses,"5))

For many years we have been using radial trepans located eccentriallywith respect to the shaft axis, in order to examine the material propertiesrelative to shaft uniformity (case 3 in Figures 7 and 8). Especially, whentrepans are pin-pointed by ultrasonic evaluation it happens that the borepenetrates or intersects a flaw. Here too, two stress peaks are superimposedon each other, and the criticalflawsize for brittle fracture is locally muchsmaller than that of the shaft without a bore. The following facts, howeverare of importance for comparison with the shaft with an axial bore:

Inherent material flaws in todays forgings are always much smaller than

the critical flaw sizefor brittle fracture, as based upon single loadingat maximum shaft stress, for example during an overspeed test, Under thecompound repeated alternating stresses of long-time machine operation, theflaws may increase to critical size by way of crack growth. Herein lies thefundamental difference between the shaft with an axial bore and the one witha radial bore: The tip of the growing crack in an axial bored shaft is alwaysexposed to an equally high stress-intensity, and therefore the crack continuesto steadily propogate. In contrast, the crack frowing away from the radialbore automaticallyenterszones of lower stresses and crack growth is retarded

and eventually stops, thereby confined to a sub-critical flaw size.

The interrelationship between operational stresses, flaw size and fracture

toughness of shaft materials will be discussed for two low-pressure turbineshafts of differing sizes and stresses, Table 3. The maximum tangential stress

of the shaft type 750, without a bore, is assumed to be T= 1.0. This stress

doubles, when the shaft receives an axial bore. The critical flaw size of theshaft with an axial bore, type 750 made of 34 Ni Cr Mo 74 steel, was assumedto be = 1D in accordance with the stress present in this shaft during overspeed

testing (D = Diameter of disc shaped equivalent flaw from ultrasonic indication,flaw position according to Case 2 on Figure 8). Subsequently, an estimate wasmade as to how large a flaw may be, under equal conditions, in a shaft withouta bore of the type 750 and in a higher stressed shaft of the type 875. For thispurpose, the fracture toughness Kic of three different rotor steels was takeninto account at the overspeed test temperature. The comparative figures inTable 3 indicate very clearly the damaging effect of an axial bore and thepositive effect of good material fracture-toughness. The considerations de-rived from the fracture mechanics are well suited to accomplish a realisticsteel selection and quality determination of rotor forgings.

In reality, numerous rotors of the type 750, utilizing 34 Ni Cr Mo 74steel, have been in operation for many years, and in most cases without anaxial bore. For Some time, these rotors have been produced only from 26 Ni CrMo V 85 steel which also contains 2% Ni but optimized in its fracture tough-ness. Rotors of the type 875 are manufactured from 26 Ni Cr Mo V 145 steel,which contains 3.5% Ni.

The preceding viewpoints about the effect of additional stresses caused

by material inhomogeneities still remain very conservative, since each in-homogeneity indicated by the ultrasonic test is initially still considered asan actual crack. This assumption, however, holds true only until sufficientinformation becomes available about the effect of inhomogeneities on materialcharacteristics, which are essential for fracture-mechanics considerations;more information is also needed about the fracture-toughness Kic and the crack-growth behavior under operational stresses. This will be covered later.

RESIDUAL STRESSES IN FORGINGS

The discussion about residual stresses is as old as the manufacture andusage of shaft forgings. Numerous publications refer to the origin and quali-tative effect of these stresses as well as to measurements of distortions whichoccur during machining operations. It must be concluded that the presentstresses and sizes of shafts require data about residual stresses of a givenpiece, which are backed by accurate measurements. It does not suffice tospecify general data, which are difficult to verify, about temperature-and time-cycles during heat-treat or to prescribe such patterns. Residual stress measure-ments are necessary not only for the safety of the shaft, but also for economi-cal utilization of heat-treat facilities.

Literature information is available concerning various measuring methods;some of which are being utilized for specific purposes. All methods, however,have limitations; therefore, their application is not suitable for the case inquestion. The "Ring-Core-Method"6) permits measuring of residual stresses atany time and at any point on the shaft surface, both in the axial and tangentialdirection; provided that a relaxation groove can be milled into the shaft withthe dimensions as shown in Figure 9. Subsequent to finish-machining, it ispossible also to measure, in thecese of turbine shafts, the reduction of axialstresses by the blade grooves and in the case of generator rotors the tangential

stresses caused by the longitudinal slots. Based upon the known stress patternthrough the cylindrical section and the measured residual surface stresses(which take into account the machining stresses) it is possible to evaluatethe internal residual stresses.

A long holding time at a sufficiently high temperature during the temperingor stress relieving process means that residual stress depend essentially uponthe temperature gradient occurring within the shaft during cooling, ie uponthe pooling rate within the plastic range of the steel. However, the fracturetoughness of the steel is also dependent on this process, since the selectedsteels display a trend more or less towards temper embrittlement, The residualstress and temper embrittlement dependence on cooling rate are however indirect opposition. Therefore, it is necessary to optimize the cooling ratewith due consideration being given to the steel properties, shaft geometry andfurnace characteristics. All these must be based on measured values, in orderto prevent a high toughness being obtained at the expense of excessive residualstress,

The results of residual stress determinations on 28 turbine and generator-rotors, are presented on Figure 10. Plotted values only include the measuredaxial stress of generator rotors and the tangential stress of turbine rotorsat the mid-length of body surface. It is apparent that here residual stressis independent of tempering temperature or body diameter. We do not believethat this is already the result of optimum treatment selection, Therefore,these variables alone surely do not lead to an optimum treatment selection.In fact any temperature >610°C is capable of substantially reducing the resi-dual stresses resulting from water quenching unalloyed and Ni Cr Mo V steelswith the presently specified 0.2% yield strengths. However, this may no longerbe the case for other levels of yield strength. In addition, the cooling ratesand tempering times may not have always been previously selected at optimumvalues,

Prose s ng may be performed with a preliminary specification value ofdo<-6401m . In Figure 10, this limit has been exceeded by a rotor identifiedas CI. This rotor as well as a second identified in the same manner were pro-duced from heat-resistant Cr Mo V-steel, which require tempering conditionsthat are different from the remainder of the evaluated steels.

In Figure 10, the rotors identified as have also exceeded the residualstress limit. The lowest tempering temperatures combined with considerablylonger tempering times of 98 or 95 hours, were applied, as apposed to 15 to 54hour holding times for comparable rotors of other producers. Since the coolingrate has not yet been considered on this plot, no statements can be made aboutthe admissib.lity of interchanging temperature and time.

The range of all axial measurement data has been plotted for a generatorrotor (1 each of both shaft extensions, 3 of the body). The values scatter toa much greater extent than normally experienced. The purpose is to point outthat the "Ring-Core-Method" is also well suited to provide evidence about theuniformity of residual stresses.

The residual stress measurements on the turbine shaft shown in Figure 5were particularly instructive, since previous experience regarding temperingof a forging having a body Oiameter of 2780mm was not available. A maximumresidual stress of -6 Kp/mm4 and a maximum cooling rate of 8°C/hour had beenspecified. In accordance with Figure 6, the actually applied cooling rate of6°C/hour with a theoretical temperatur gradient of 83°C resulted in undesirableresidual stresses of -13 to -19 Kp/mm . Only repeated tempering with an asso-ciated cooling rate of only 2.5°C/hour in the applicable temperature rangebrought about the expected success for this large diameter forging. The factthat the tempering temperature was increased by 30°C, had no effect on theresidual stresses of this steel with only 0.25% Wo and prior "hardening" in air.

This initial evaluation clearly indicated the necessity for accurateresidual stress measurements. This is in the interest of the producer forselection of qualitatively and economically optimized heat treatments. It isin the interest of the consumer to specify measurable and characteristic stressvalues. In this manner the responsibility can then remain under the properjurisdiction.

FRACTURE TOUGHNESS AND FATIGUE CRACK GROWTH BEHAVIOR

The total stresses of the rotors consist of the static and dynamic loadstresses, which are subject to the machine operation, and of the residualstresses prevailing within the forging. In order to evaluate the effect ofthe total stresses on the behavior of rotors, thereby applying the rules ofthe linear, elastic fracture mechanics, the following information is required:

1. The temperature relationship of the material-fracture-toughnessKIc, which already has been mentioned, together with the formulaein Figure 8, and

2. The crack growth behavior of rotor material under simulatedoperational dynamic stressing.

FRACTURE TOUGHNESS

The determination of the fracture toughness Kic is accomplished in accord-ance with the regulations described in the testing standard ASTM-E 399. Incontrast to many other structural pieces, plane-strain conditions always pre-vail within the internal structure of large rotor-forgings; the validity oflinear, elastic fracture-mechanics is premised on these conditions.

Of interest is the temperature dependence of the fracture toughness KIc atthe lowest operating temperature ranges of rotors. This is the environmentaltemperature at the location of the overspeed test and at the location of cold-starts after extended periods of shutdown. In order to maintain plane-strainconditions in this temperature range during experimental KIc measurements, verylarge samples of presently utilized ductile rotor materials are required, that

is at least up to size 015 at room temperature, sample dimensions of

125x300x315mm. Figure 11 shows fracture-toughness-samples (Compact Tension

Specimen) of sizes CT1 through 015 per ASTM-E 399, which have been used for

the eXaminations described later. Figure 12 shows the experimental setup for

Kic measurement of a CT5 sample at room temperature.

These large samples make it impossible to perform a Kic measurement as

part of conventional acceptance testing of forgings. The major difficulty is

in the procurement of representative sample material, even for basic examina-

tions. Of course, the sample material must have the same characteristics as

the zone of maximum rotor stress, which is to be calculated by using its KIc

value. The sample material must therefore agree metallurgically with the

rotor, that is inclusive of segregations, inclusions, grain flow end structures

In the case of welded rotors, this also applies to the weld material and es-

pecially to the various zones of weld to base metal transitions.

As a result, the following two tasks arise:

1. Basic examinations of representative samples from rotor

materials which will be utilized and

2. Correlation of material characteristics, which can be

tested during acceptance.

Sample material was available from various forgings and materials for the

measurement of fracture toughness and crack growth behavior. Figures 13 and 14

show two examples.

Generator rotor 7776 in Figure 13 had a forging weight of 58 tons and was

manufactured from a 100 ton ingot. An experimental piece of 990mm diameter

x 1250mm was forged onto the bottom end of the rotor and left there until

completion of the first tempering operation of the forging at 600°C. After-

wards, a disk-shaped piece was cut off, while the remainder of the experimental

piece, together with the rotor, were subjected to a final heat treatment at

67000. In this manner, uniform sample material was obtained from a single

forging, but under two different heat treat conditions.

The experimental forging 1020mm diameter x 1160mm, which is represented in

Figure 14, had a forging weight of 14 tons and was forged separately from sur-

plus material taken from the top section of the 400 ton ingot for generator

rotor 2.7901 (see Table 6, Figures 2 and 49 through 56). This experimental

forging was subjected to a simulated heat treatment, which corresponded to

that of the 1800mm diameter rotor 2.7901. The heat treatment was monitored

with thermocouples, which are drawn into Figure 140

In addition to the sample material obtained from the production forgings

described in Figures 13 and 14, several axial trepans were available which

were removed from the bodies of the large capacity rotors; refer to Figures 46,55, 56. The axial bores on some of these rotors (7642, Figure 46, 2.7901 and3.8153, Figure 55) were produced specifically for the purpose of removingporosities from the rotor-center, all of which had been detected during ultra-sonic inspection. These inhomogeneities will be discussed later. The result-ant material examinations have established the fact that all of these porositiescould have been left in the rotors without endangering the machine operationand that the axial bores would not have been necessary. On the contrary, theaxial bores of rotors 5.8154 and 6.8189 (see Table 6 and Figure 56) were per-formed for the sole purpose of obtaining representative sample material fromthese forgings which were manufactured in accordance with the "A.P. Process"of the Japan Steel Works.

The fracture toughness experiments performed on samples from the interiorof heavy forgings were supplemented with measurements on external samples.This category also includes samples from the hub bores of several large LPturbine discs. Some additional information about the materials, covered herewith regard to their fracture toughness characteristics, has been compiled inTables 4 and 5. The symbols used in these tables refer to the fracture tough-ness-temperature-diagrams, Figures 16 and 17.

During removal of the samples for KIc and crack growth measurements, thecrack plane was positioned, as far as possible, normal to the main stressdirection prevailing in the rotors: crack plane axial-radial; direction ofcrack propagation radial. This is indicated on the sketches in Figures 13 and14. These sketches also illustrate that the "radial" direction of crack prop-agation could not always be strictly adhered to because of the geometry ofsample and experimental material. In all cases, however, the crack plane waslocated parallel to the grain flow of the forging. Attention was given to thedistribution of samples over the cross-sectional area of the respective forgingsuch that the crack tips of the samples were located, if possible, in zones ofuniform mechanical characteristics, in order to obtain comparative results withminimum scatter. In order to evaluate the influence of heat treatment uponsample location, tensile, notched impact and drop weight specimens were testedover the total cross sectional area. The brittle fracture transition tempera-tures FAT7 (Fracture Appearance Transition Temperature 50% crystalline fracturesurface on Charpy-V sample) and NDTT (Nil Ductility Transition Temperature,ASTM-E 208) served as aids for comparative judgment of the KIc samples. Thediagrams in the lower right corner of Figures 13 and 14 illustrate this process.

The results of the fracture toughness measurements are compiled as Kicdiagrams in Figures 15 through 17. Figure 15 summarizes and compares valuesfor various rotor steels through the end of 1970, For comparison, measurementswere also utilized which have been,reinorted repeatedly, as for instance byGreenberg, Wessel and colleagues 5)0 (). The minimum curves on Figure 15 andtheir relationship to the respective products are explained in the legend.Today, the following items of this summary still appear to be noteworthy:

1. For NiCrNoV steel with 3.5% Ni: The American and German measure-

ments correlate quite well, This means that forgings have equallygood characteristics when they are manufactured by the same processfrom the same alloyed steels, regardless of the location of manufacture,

2. For Cr-free NiMoV steel with 3 to 3.5%; All forgings, which have beenexamined to date, have been manufactured in the U.S.A. The valuesreported by Greenberg and colleagues, as well as our own measurementson trepans of generator rotor 7642 (figure 46), agree very well witheach other. This means that the regulations established in the teststandard ASTM-E 399 yield reproducible test results, regardless ofthe location of the laboratory performing the tests.

3. For NiCrMoV steel with 2% Ni: The development effort expended onthis steel grade during the past 10 years, in conjunction with astep by step change of the chemical composition and treatment, hasresulted in a considerable improvement of the fracture toughnesscharacteristics.

Figure 16 gives a summary of fracture toughness measurements, which arepresently available for the 2% NiCrMoV steel 26 NiCrMoV 8 5, which includesthe trepans from the 197 ton 1800 mm diameter rotors 2.7901 and 3.8153; theserotors were manufactured from conventionally poured 400 ton ingots (see Table 6,and Figure 55). The measured values were grouped according to the NDT tempera-tures which prevailed at the location of the sample removal. The scatter bandis valid for all pieces having an NDTT In addition to the fracturetoughnesses plotted in Figure 15 for 2 NiCrMoV forgings up to 1000mm maximumdiameter, and the external samples from 1800mm diameter rotors, we have in-cluded the results of center samples bottom end of rotor 3.8153 and mid-bodyof rotor 2.7901. The lower boundary curve of this spread, which agrees withthe curve "1970" in Figure 15, is possibly somewhat optimistic in the rangebelow approximately -60°C; this, however, is of no significance to practicalapplication of turbine and generator rotors.

The fracture toughnesses of trepans AXK originating from the top ends ofrotors 2.7901 and 3.8153 are located below the spread range in Figure 16; theheavily lined minimum curve and NDTT 4=+15°C applies in this case. It isessential to note that this minimum curve for the 2% NiCrMoV steel lies higherthroughout the total temperature range than the corresponding curve for theCr-free 3-3.5% NiMoV steel ASTM-A 469 Class 4, although the rotors manufacturedfrom 2% NiCrMoV steel had much higher ingot and forging weights as well aslarger body diameters.

The difference in fracture toughnesses of the 1800mm diameter trepansfrom the ingot bottom and center as compared to those from the ingot tops maybe considered to be the result of the C-segregation (higher C-contents andtensile strength in the ingot-top section, see Figures 53 and 55). Furtherimprovements are therefore possible by changing the pouring and/or deoxidationprocesses; refer to the results of segregation examinations on ingots whichwere poured by way of the "A.P." Process.

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The C-content of the 1020mm diameter test block (Figure 14) which was manu-factured from the surplus material of the top of the rotor forging 2.7901, in-creases from 0.25% at the outer surface to 0.33% at the core; here too, onlythe C-content at the core agrees with that of the top-side trepans AXK/AXKSof the rotor (see Figure 55). The trepans of the experimental forging, however,display a more favorable fracture-toughness and NDT temperature than thetrepans AXK/AXKS of the rotor, regardless of the higher 0.2 yield strength(compare data in the lower right graph of Figure 14 with those of Figure 55).The separately executed, simulated heat treatment of the experimental forginghas not yielded exactly the same characteristics as the production heat treat-ment of the rotor.

Figure 17 contains the Kic values of the 1800mm diameter rotors, whichwere manufactured from 3.5% NiCrMoV steel 26 NiCrMoV 14 5. For the purpose ofcomparison, the lower boarder lines of the Kic temperature scatter bands havebeen plotted for

1. Rotors up to 1500mm diameter (trepan location 1/4 D below surface)and turbine-discs made of the same steel, transposed from Figure 15,

2, Rotors up to 1700mm diameter made from ASTM-A469 Class 4 steel,transposed from Figure 15,

3. Rotors up to 1800mm diameter made from 26 NiCrMoV 85 steel,transposed from Figure 16.

The Kio values for the 1800mm diameter 3.5% NiCrMoV rotors were combinedwithin a common scatter band in Figure 17, which is, however, not completelycorrect. It is true that both rotors 5.8154 and 6.8189 originate from the400 ton ingots, which were produced by the "after-pouring" process; however,rotor 5.8154 was produced without vacuum-deoxidation and tempered at 6350Cafter water-spray tempering, whereas rotor 6.8189 was manufactured withvacuum-deoxidation and tempered at only 6100C. In reality, these two rotorswere produced differently and consequently, acquired different mechanicalcharacteristics, see Table 6 and Figures 49, 51, 53 and 56. This was the in-tention of the development plan, but it naturally effects the comparison. Themanufacturing and treatment variations must be taken into account during dis-cussion of the measured fracture toughnesses;

The KIc values of the outer samples of both rotors are equally good andincrease above 450 Kp mm -1.5 even at -1000C. The difference between theouter and inner samples of rotor 5.8154, which was not vacuum deoxidized, isof the same order of magnitude as the previously discussed 2% NiCrMoV rotorsfrom conventionally poured ingots, Figure 16. The trepans AXM taken from thecenter of the body (trepan location, see Figure 56) determine the lower boarderline of the Kic temperature scatter band in Figure 17.

As compared to rotor 5.8154, the fracture-toughness characteristics ofthe vacuum deoxidized rotor 6.8189 is more uniform and also better regardlessof the higher tensile strength. The fracture-toughness at -2000 for center

top end samples AXK (see Figure 56) lays on the minimum curve for much smaller

forgings and surface tests transposed from Figure 15. This excellent result showsthat a high 0.2 yield strength with a simultaneous high fracture-toughness can beachieved even for large rotors, provided that proper processing is utilized.

The summarized results of Figures 15, 16, 49 to 56 draw a clear path tothe further development of rotors and their determining operating character-istics. The first steps have already been taken, especially towards the op-timization of the chemical composition, see T-ble 6, rotors 7.8309 to 9.8307(Figure 79).

The minimum values of fracture toughness Kic listed in Figures 14 through16 and the formulae in Figure 8 can be utilized to calculate the criticalbrittle-fracture-flaw-size acr for that case when rotors are loaded once withthe greatest stress at the lowest temperature ever occurring during operation.The following results correspond to an overspeed test of a 1800mm diameter -2000 MVA - generator rotor (see Table 2) at a rotor temperature of +15°C, where-by the assumption is made that a semi-elliptical crack of the shape: cracklength 2C=10 x crack depth a (worst case flaw geometry and - position) islocated at the axial bore surface:

Diameter of CircularCritical Crack Substitution Flaw ofDepth acr Length 2C Equal Area

Material (mm) (mm) mm

Greenberg and colleagues5) have obtained comparative figures of similar magni-tude. They have explained the limitations in detail, under which these compu-tations must be evaluated; these limitations apply also to our example. It isindisputable that flaws of such a large size can be already detected clearlyduring the first ultrasonic inspection at the forge shop and that such flawswill not occur in delivered forgings.

CRACK-GROWTH

Turbine and generator rotors are not only facing one-time heavy peakloading events, but they are subjected to more or less large load fluctuationsof varying frequency during many years of machine operation. Large turbine-generators at some power plants are either started up and shut down daily orthey fluctuate several times daily between full and reduced machine output8)°There are no uniform output time tables.

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Under the assumption that a crack exists in the rotor forging, the start-up and shut-down as well as the output fluctuations of the machine create avarying stress-intensity AK at the crack-tip. Depending on the magnitude andfrequency of this AK, the crack may grow in time more less rapidly to criticalflaw-size. For reasons of fracture-mechanical safety considerations of rotors,not only must the maximum tolerable stress intensity and fracture toughnessKIc of the material be known, but also the crack-growth behavior under simu-lated operational conditions of varying stress intensities.

The relationship between crack-growth rate da/d0 (a=crack depth, N=numberof stress cycles) and varying stress intensity AK acting upon the crack tiphas been described by Paris9) with the general crack-growth equation: da/dN=C04Km.A linear relationship, therefore, exists on the log-log coordinate scaleda/dN=f (log AK), whereby the constants m and c establish the slope of the lineand its point of origin. The crack growth exponent m is essential for thecomparison of various material.

Fracture toughness measurements were performed on the CT-samples (fracturetoughness test samples) ASTM-E 399, which are shown on Figure 11); the crack-growth curves were plotted automatically as a function of the load cycle numberby means of the ultrasonics or the electrical potential difference methods.

Figure 18 summarizes the results of our own crack-growth measurements, aswell as comparable data from the literature5)7)10),for the three rotor steels26 NiCrMoU 85 (2% Ni), 24 NiMoV 14 5 (3-3.5% Ni, ASTM-A 469 Class 4) and 26NiCrMoV 14 5 (3.5% Ni). For practical application, the upper border lines ofthe scatter band and their crack-growth exponent m are of interest. These arein practical agreement for the 2% NiCrMoV steel and the 3% NiMoU steel.

In addition to other samples from the 3.5% NiCrMoU steel, trepans AXF ofthe 1800mm diameter rotor 6.8189 were also examined (sample location, seeFigure 56); and their test values are denoted on the right partial graph ofFigure 18. They fit well along a line having a crack-growth exponent of m=3.The same exponent applies also to the range, which encloses in total our crack-growth values for 3.5% NiCrMoV. This is contrary to the measurements, whichGreenberg and colleagues5) 7) have reported. They found considerably lowercrack-growth exponents m=1.4 to 1.5 for the two forgings 3178 and ZV 3037;which is also indicated on Figure 18. For the time being, the reasons of thesedifferences have not been determined. Worth mentioning are the differences intest techniques, whereby Greenberg and colleagues performed their experimentsat a lower test frequency of 10 HZ (in comparison to 120-160 HZ for our experi-ments) and with higher varying stress intensities AK. We do not believe thatthis explanation is adequate. Additional testing will be required.

All crack-growth measurements presently available have been performedunder continued cyclic vibration stress and at much higher load frequencies,than turbine-generator rotors will experience during output fluctuations. Itis not yet sufficiently known how the crack-growth behavior of steels is af-fected by low rate, operational-type load-fluctuation, by the center load which

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is subject to centrifugal stresses,by environmental conditions (steam andgenerator coolants) and other operational influx-factors.

Until final clarification, all three discussed steel grades may be con-sidered equivalent to each other on the basis of previous measurements andoperating experiences; the calculation of rotors may be based on the generalcrack-growth equation:

da/dN (2,10-10) AK3 mm/load cycle.

The fundamental investigations, which have been discussed in this chapter,yielded the following summarized results:

1. The fracture toughness K/ of rotor materials depends to a highdegree on the chemical coMposition of the steel, on the manufactur ngprocess and heat treatment. Specific information is available forfurther development of large capacity forgings.

2. The crack-growth behavior of current rotor materials of various steelgrades is largely independent of the just mentioned influx-factorsand are equivalent to each other. Sufficient information is availablefor reliably dimensioning and classification of rotors. In order toachieve higher material efficiency, further basic research is re-quired (long-term experiments under operational conditions).

The processes, which were applied during the basic research, are unsuitablefor routine acceptance testing, as previously explained. This is why reliablerelationships are needed between the fracture-toughness KIc and other materialcharacteristics, which can be tested during acceptance.

The question is: Now does rotor material behave, when a crack is locatedat a highly stressed point of the forging? The most reliable answer is to beexpected from a specimen with a natural crack. It is common practice to judgethe brittle-fracture characteristics of the material by notch impact acceptancetests. On Figure 19 a natural crack is contrasted to the notch-shapes ofseveral notch impact specimens, all are drawn to the same scale. All notches,which were produced mechanically, are much less severe than a crack, and itbecomes immediately clear why a simple and indisputable interrelation does notexist between the results of notch-impact tests and the brittle-fracturebehavior of cracked component parts.

The simplest laboratory specimen with a natural crack is the drop weightspecimen developed by Pellini and colleagues.") for determination of the NDT-temperature, Test Standard ASTM-E 208. The applied stress is simulated inthis experiment such that the material is stressed at the crack location witha sudden impact beyond its elastic limit. Naturally, the attempt is made toavoid this impact type of stress during machine operation; however, it mayoccasionally occur in turbine-generators because of false synchronization orhy short-circuit stresses.

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The NOT-temperatures of the examined forgings, respectively, are plotted onthe lower border lines of the fracture-toughness scatter bands. It is noticeablethat the fracture-toughnesses definitely improve with decreasing NOT-temperature.The fracture-toughnesses also increase rapidly to very high levels at tempera-

tures beyond NDTT; this is to be expected according to the Fracture-Analysis-Diagram (FAD), which was established by Pellini11) 12). In this case, definiterelationships exist between Kic and the results of a simple laboratory experi-ment; this is also the reason that we have introduced the NDTT test into ourdelivery specifications.

Figure 23 schematizes the specimen removal for acceptance tests of gener-

ator rotors. The exciter bore is extended to 600mm below the body, in orderto obtain representative sample material for the NOTT-test. Naturally, thestatements made in the section "Operational Stresses in Rotors" regarding theaxial bore apply to this section of the generator-rotor. In looking ahead,this method of specimen removal and testing is unsatisfactory; we strive towardscomplete elimination of axial bores below the rotor body. For this purpose,extensive research is presently being performed to define other test methods,

and, especially with still smaller specimens, to establish definite relation-ships to material fracture-toughnesses which are known from basic research.

Further examinations address themselves to the question of how fracture-toughness KIc and crack-growth behavior are effected by material-inhomogeneities

which occur in large forgings. Before initial results of these efforts arereported, the types of inhomogeneities will first be discussed.

INHOMOGENEITY OF ROTOR FORGINGS

It is well known that homogeneous ingots do not result from oonventionalsolidification. This has been reported repeatedly in detail13 -16). Figure 20ashows the segregation pattern of a large ingot, as published by Kawaguchi andKudo16). Since there is only a remote possibility to change the ingot segre-gation resulting from the solidification process by subsequent forging andheat treatment, it must be recognized that corresponding inhomogeneities mustbe present within the rotor forgings. Figure 20b reproduces the results ofC- and P- analyses of two radial trepan samples, which were removed from agenerator-rotor and corresponding to the top and bottom position of a 120 toningot. Statements regarding the location of maximum segregation in largecross section forging ingots have been verified14,16). The segregation is notconcentrated along the ingot axis but rather in a cylindrical zone intersectingthe approximate mid-radius of the ingot body. Figure 22 confirms very des-

criptively the additional unfavorable effect of high C- content in the melt;with its large commensurable segregation, the alloying elements also segre-gated to a larger degree14).

In any event, rotors manufactured from these iogots have met the necessaryquality requirements up to this time. Ammareller16) has examined a generator-rotor containing extensive segregations, which has remained in operation withoutdamage, despite excessive overspeed. Based on available publications, no case

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is known whereby failures were directly associated with normal inhomogeneities.It is uncertain, however, whether the presence of these typical customary in-homogeneities will also be acceptable for rotors of the coming 80's. On theone side, operational stresses combined with resultant stress concentrationsassociated with inhomogeneities will increase. On the other side, some obscuritystill exists as to the extent and to the factual influence of these inhomogen-eities upon the operational characteristics of the forging. The latter may bebased very well on the fact that the sample material is selected frequentlyfor routine acceptance tests in such a manner that the effect of inhomogeneitiesis not detected at all during testing. The numerical results of tensile or im-pact test specimens which are removed from the rotors either radially or tan-gentially out beyond the mid radius of the body diameter or axially from thecore, are not affected by customary inhomogeneities at all. The statisticalevaluation of these numerical values is used as a basis, even today for mostacceptance test specifications. The significance of the following discussionsare to establish the status of inhomogeneities of present forgings, and usingthese as a basis for the development of the homogeneity requirements for forg-ings of the 80's.

When trepans 1 and 2 with a length of 0.6X body diameter are removed fromthe zone of maximum segregation Figures 23, or as located by ultra-sonic in-spection, and when these trepans are tensile-tested in their total length(refer to Figure 24), the first fracture will automatically occur at the weak-est point. Subsequently, the specimen is reclamped and tensile-tested again.Of course, the mentioned statistics cannot be applied in any way to judgingthese numerical results. The numerical result of the ductility values (6 /4')of such a small test sample is purely accidental and is dependent upon theinfluence of segregation stringer volume normal to specimen axis. The detectionof inhomogeneities is definitely accidental during notch-impact testing withpredetermined fracture location. A considerably better insight into the type .and extent of inhomogeneities is obtained by comparative microfractography ofindividual fractures.

TREPANS ORIENTED BY ULTRA-SONIC INDICATIONS

Prior to the introduction of the vacuum-degassing process, it was ofspecial importance to evaluate ultrasonic indications to determine whetherthey were caused by flakes or only by typical inclusions. Figure 25 showsfracture I of a radial trepan, which occurred without deformation far belowthe 0.2 yield strength level. The rotor in question was rejected in 1951because of flakes.

The radial trepan shown in Figure 24 was removed from a generator rotorwhich was produced from a 100 ton ingot. By means of ultrasonics, a singleindication corresponding to an equivalent flaw-size of 3 to 4 mm diameter wasdetected at a depth of 265 mm. Fracture I occurred without deformation at adepth of 257 mm; the inclusion within the tensile specimen had a projecteddiameter slightly greater than 5.5 mm. The fractographic and metallographicexaminations (Figure 26 a/b) resulted in the detection of MnS-inclusionswithin a highly ductile matrix.

During ultra-sonic inspection of the generator-rotor 4.7933,(which had an

1800 mm diameter, a weight of 172t and was produced from a 400 ton ingot),

several indications were obtained near the axis. Their frequency and distance

from the axis increased from the bottom to the top end of the rotor. The

average equivalent flaw size was determined to vary from 1.5 mm diameter at the

bottom end to 2.5 -4mm diameter at the top end, with a single isolated flaw of

2.5 mm diameter at the bottom end. It was decided to utilize the rotor without

an axial bore but however to evaluate by means of 6 radial trepans the type and

extent of the indications. In accordance with Figure 50 the trepans K14/18/19

were removed from the zone of the strongest indications, while trepan F22 was

aimed at the single flaw at the bottom end. The trepans were then bisected.

Figure 52 contains the test results of fracture I of the outer half. Up to a

depth of 540 mm, no effect of the segregation was detected, magnetic particle

inspection of the remaining 540 mm length of inner specimen halves did not

reveal any indications. Figure 54 contains the test results of fracture I.

Figure 27 shows all multiple fractures of the inner radial trepans of this

rotor. The zone of the ultrasonic indications and the sequence of the fracture

tests are also indicated. The correlation of tensile strength and carbon con-

tents of the radial trepans F22, F3, F2 and K19 leads one already to expect

Fracture I to occur near the forging axis as a result of negative segregation.

During grinding of the 17 mm diameter trepan F22, a spot-shaped flaw was

readily detected which corresponded to the ultra-sonic indication. The results

of the examination of fracture I is shown in Figure 28. The flaw in question

was identified as a cylindrical silicate inclusion with rounded ends proportional

to a circular area of 5.5 mm diameter, which was imbedded in a highly ductile

matrix. Similarly, all consecutive fractures also occurred after large deforma-

tions.

Fractures I and II of trepans F3/F2 and K19 were associated with zones of

large deformation, that is without the influence of inhomogeneities. Fracture

I of trepan K14 occurred outside the zone of extensive ultra sonic indications

at a distance of 275 mm from the axis, Fracture II inside this zone at a

distance of 25 mm. The metallographic and microfractographic examinations,

Figure 29, 30, revealed the presence of mnS-inclusions. The reproduced photo-

graphs a/b have all been taken at 400X magnification and show good ductility.

The 1000X enlargement shows a very ductile fracture between inclusions. The

trepan K18 fracture at a distance of 10 mm, and the results of this examination

were similar to those of trepan K14.

During ultra sonic testing, especially of rotors with large diameters,

indications are occassionally obtained from the vicinity of the axis, which

are often the cause for making an axial bore. In accordance with the state-

ments for Figures 7 and 8, the stress is thereby doubled and large stress con-

centrations develop at possible flaw locations. An axial bore, therefore,

should be large enough to remove all possible inhomogeneties. This is, however,

not possible according to the statements made for Figure 20.

During the testing of rotors 2.7901 and 3.8153, ultra sonic indications

were obtained with an average equivalent flaw size of 2 to 5.5 mm diameter

within the main body as well as within the bottom end shaft extension. In

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order to examine the cause of these indications and to determine influence onthe operating characteristics, a full length axial trepan was removed which wasof sufficient cross section such that extensive investigations could be con-ducted in both axial and radial directions (see Figure 55). Figures 31/32show that the indications within the main body were caused by porosity (seealso rotor 7642, Figure 46); those in the shaft extension however, were causedby inclusions. The radial tensile test samples from the axial trepans havinglow-deformation fractures (see Figure 55) were subsequently examined metallo-graphically. Only MnS-inclusions were confirmed, mostly of type 2 (seeFigure 33).

TREPANS NOT PIN POINTED BY ULTRASONIC INDICATIONS

Since 1936 the results of more than 2000 radial trepans are available(first removal of radial trepans in 1913). Prior to the introduction ofvacuum-degassing, when low-deformation fractures were determined with fractureelongations of approximately the most frequent cause was the effect ofhydrogen in the form of flakes or segregation cracks. Subsequent to the intro-duction of degassing, fractures were judged via stereomicroscope as "fibrousfractures", Many metallographic examinations performed jointly by Sealise 19)and Opel20) resulted in the conclusion that low ductility is caused by type 2MnS-inclusions, Figure 33. Additional knowledge could be obtained only afterthe introduction of microfractorgraphy as a tool for these investigations.

Figure 34 shows the results of the examination of a small two pole rotorproduced from a 100 ton ingot, which contrary to the rotor forging 7557EFigure 26, was not vacuum deoxidized. The elongation of the solid radialtrepan is plotted at lOmm intervals for the entire length; which resulted in auniform elongation of approximately 10%. The fractures I to III occurred withnormal fracture ductility. Fracture IV was the first to occur after reachinguniform elongation but with negligible reduction of area. Fracture III oftrepan K occurred at approximately the same distance from the axis, again withreduced deformation. The fracture displayed two crystalline stringers oneither side of a shear fracture. At the stringer location, the cross-sectionalmicrograph revealed coarse grains having increased hardness and molybdenumsegregation of 1.19%. Positive segregation was also present for Cr, Mn and Ni.The microfractography revealed an intercrystalline fracture surface within thesegregation stringers and high ductility within the zone of the shear fracture.

A similar fracture appearance was observed on the radial trepan of a verysmall rotor with a delivery weight of only 15 ton and produced from an 85 toningot, which again was not vacuum-deoxidized, Figure 35. The microfractographyrevealed intercrystalline fracture in the immediate vicinity of the segrega-tion, but elsewhere it was ductile. A metallographic examination of the totaltrepan section between the fractures II and III revealed an additional zone ofcoarse grain size. A Baumann-Print of this spot shows a sulphide line on bothsides of the coarse grain. Regardless of the 0.34% to 0.55% Mo-segregationwith an associated hardness increase of approximately 20 points and the expec-ted intercrystalline fracture behavior, no cracks could be detected.

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Subsequently radial trepans were taken from the sample stockroom of ourplant. These trepans had yielded low deformation fractures and were repre-sentative of production prior to the introduction of vacuum-degassing. Figure36 contains an example of the examination results of a rotor forged from a 96ton ingot. The deformationless fracture occurred in an inter-crystallinemanner. At a distance of 15mm from the fracture, an additional segregationspot was detected by means of a Baumann-Print; this time, eutectic inclusionsof type 2 were present with a Mo-segregation of 0.40% to 0.92% and a hardness

increase of 20 units.

An examination of low-deformation, radial trepan fractures (Figure 37) ofthe unalloyed rotor 7838 again yielded MnS inclusions of type 2. At the frac-ture, however, the microfractography did not indicate any intercrystalline

fractures.

The rotor 5.8154 was subjected to four tensile fracture tests and sub-sequently, Baumann-Prints and micrographic examinations were made throughoutthe total length of each of the trepan-inner-and-outer-halves. Coarse grainzones were found with increased hardness and Mo-segregation although the carbonsegregation was for all practical purposes eliminated completely by the "A.P.-Process" as will be shown later.

The results of the examination of the largest coarse grain zone areplotted in Figure 38. This zone was located in the top trepan approximately750 mm below the body surface. Additional coarse-grain zones of smaller sizewere found throughout the total trepan length, but always in conjunction withsulphide-segregations. Cracks could not be detected at any of these spotsdespite repeated micro polishing to various depths.

A miniature notched tensile test sample was removed from the coarse grainzone in Figure 39a in such a manner that the fracture occurred within thesegregation zone. The microfractorgraphy showed an intercrystalline fracturesurface, Figure 39b, which is embedded directly within a highly ductile matrix,Figure 39c/d. A similar coarse grain zone was retempered for 15 hours at6300C. The hardness of the coarse grains and the matrix were reduced, howeverthe hardness difference prevailed and the matrix structure did not change.A notched tensile test no longer produced an intercrystalline fracture at thisZone. The hydrogen content was locally measured at two additional coarse grainzones and amounted to 0.9/0.9 ppm; values between 0.2 and 0.3 ppm were foundin the vicinity of these zones and at other locations in the rotor, howeveroutside of the segregations.

Reduced fracture deformation was discovered in long radial trepans ofrotors, which were produced from small as well as large ingots of approximatelythe last 20 years. The fact that in most cases MnS is found in the eutecticform, indicates that the fracture occurs in the zone of the residual melt andlatest solidification. Other alloys in steels also exhibit a positive segrega-tion, however molybdenum was only used as an indicator in this case, since ithas the highest segregation coefficient. Pouillard 13) and Bastien 14) have

pointed out this fact already after their investigations of ingots. Duringheat treatment, which is adapted to the melt analysis, structure variations13, 17, 19) and variable hardness levels will naturally develop in these areas.

It is also understandable that the residual hydrogen is not uniformly distribu-ted throughout the forging; instead, it is concentrated in the segregationzones and will cause intercrystalline fractures under room temperature stresses.The size of the segregation zones appears to noteworthy, which already appearsin the centimeter range in a rotor produced from an 85 ton ingot; these zonesalso spread over the total cross sectional area.

At this time, we must point out again that the inhomogeneties describedabove have never resulted in a failure. The rotor 4318 Figure 36, has been inoperation since 1956 and has accumulated 60,000 hours. It must also be pointedout that this rotor was produced without an axial through bore or an axial testbore. However the rotor was under a stress of only 21 kp/mm2 with an internallymeasured 0.2 yield strength of 58 kp/mm2. In the future however, rotors ofequivalent body diameter will operate at double the stress, and will requireinternal 0.2 yield strength, greater than 80 kp/mm2, and must meet necessaryacceptance requirements to provide adequate safety (Table 2).

THE "A.P. - PROCESS"

At Terni in 1970 and at Philadelphia in 1972, Kawagnchi and colleagues15, 22) have referred to a process, by which the otherwise conventional ingot

displays a considerable reduction in segregations. In conjunction withFigure 21, which was taken from their report, they explained that after acertain solidification time, an after pouring is performed with reduced c-content and a high degree of purity such that the residual molten mass, whichhad accumulated by segregation is again "diluted."

Starting in January 1971, this process has been applied per Table 6 forthe forgings now under discussion. At this time, test results are availablefor the rotors 3.8154 and 6.8189 Figure 56, A Baumann-Print and a metallo-graphic examination were made throughout the total length of 1080 mm radialtrepan. In addition, the carbon content of the radial trepans was determinedat 25 mm intervals, while that of the axial trepan was selectively determinedover the total length.

Figure 40 shows a AC - 0.03% (155 individual determinations) over thetotal section of rotor 6.8189. The same analyses were performed for rotor5.8154, that is trepan K by JSW and trepan F by KWLI. The result was a mutualAC - 0.03% (76 individual determinations). Additional data about segregationanalyses are contained in the report by Kawaguchi 22). When these data arecompared with theAC - 0.07% in Figure 20b for the rotor, which was producedfrom a 120 ton ingot (see also Figures 46/54), it becomes apparent that thecarbon segregation in the 400 ton ingot has practically been suppressed com-pletely by this process. Local coarse-grain formation with positive Me-segre-gation, however, was detected in the rotor 5.8154. In rotor 6.8189, which hadbeen vacuum-deoxidized, even this segregation has reached such a low level thatit becomes insignificant. There was no evidence of any lines of demarcationresulting from the after pouring process.

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INHOMOGENEITIES AND MECHANICAL PROPERTIES

Now, what is the effect of previously discussed inhomogeneities on theservice properties, which are essential for rotor operation, namely materialfracture toughness and crack-growth behavior? The major difficulty in theattempt to answer this question again consists of obtaining representativesample material. This material must correspond in all details to the mosthighly stressed point of the forging, but in addition it must display as manylarge flaws as possible and inhomogeneities in such abundance and distributionthat precise samples can be extracted for mechanical testing. It is easilyunderstandable that such sample material can only be obtained from the forgingproducers as time goes on and with continuous, close and open-minded coopera-tion.

POROSITY FRACTURE-TOUGHNESS AND NDTT

In order to reduce these difficulties, simple model experiments were per-formed initially on specimens with artificial flaws to evaluate orientationinfluences. Drop weight specimens of size P2, ASTM-E 208, were removed from a26 NiCrMoV 85 forging and prepared as shown at the top of Figure 41. Specimentype "1" received 5mm diameter holes, oriented normal to the direction of crackpropagation and served as artificial flaws within the crack plane. The holesof specimen type "2" are oriented normal to the crack plane just below thetension-zone, which experiences maximum stress during the drop weight experi-ment. Specimen type "3" corresponds basically to type "1"; however, the flawsare simulated this time by 0.5x5mm spark erosion grooves.

As a supplement to Test Standard ASTM-E 208, we measure the crack lengthoriginating on the tension side of the specimen during each drop weight experi-ment; in this manner, we are able to represent the experiments, which lead tothe determination of the brittle-fracture-transition temperature NDT, in anillustrative form at crack length-temperature diagrams. The smallest possiblecrack length corresponds to the width of the crack-starter-weld bead and thelargest possible crack length corresponds to the width of the specimen(50.8mm = 2 inches for the P2 - specimen). The test results of the NDT-specimen provided with artificial flaws, are summarized on the crack length-temperature diagram at the bottom of Figure 41. The normal P2-sample (withoutflaws) yielded an NDT-temperature of -25°C for this material. The same resultwas obtained from specimen type "2". In contrast to this, the test tempera-ture for specimen Types "1" and "3" had to be lowered another 110°C and 250C,respectively, in order to obtain complete fracture of the specimen. The run-ning crack apparently has been "drilled out" and truly intercepted by theartificial flaws.

In another experimental series, the grooves of specimen type "3",Figure 41, were converted to fatigue cracks by means of repeated bendingstresses, before the conventional crack starter weld bead was placed onto thespecimen. Again, the drop weight test resulted in an NDT-temperature of -25°C,the same as that of the P2-specimen of flawless material.

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Porous material was available from section AXM of the axial trepan of alarge generator rotor. Figure 42, upper left, reproduces an x-ray of the crosssection of a 20mm thick disk taken from this trepan. A P2 drop-weight specimenwas prepared from this disk such that the crack plane and the porosities asdetected by x-ray were coincident. The specimen was drop-weight tested at+10°C and was not completely fractured. In order to expose the crack plane,the specimen was again cooled to +5°C and then stressed for a second time.The two pictures on Figure 42, upper right, show the fracture surface and theadjacent photomicrograph AXM 3/1, each with the respective porosities. Thetest results of the porous specimen are compared with flawless AXM-trepansobtained from equivalent locations on the crack length-temperature diagram atthe bottom of Figure 42. The behavior of the porous specimen is even morefavorable during the first stress at +100C, than that of the flawless compari-son-specimen. Here too, as during the model experiments, no ill effect of theporosity can be detected with respect to the NDT-temperature. In the meantime,these results have been verified repeatedly by examinations of porous speci-mens, which were extracted and tested by the same method. Hochstein 21) hasobtained the same results during his examinations of flaked sample material.

During fracture-toughness testing of rotor-materials, CT-samples werefrequently extracted in such a manner that the crack tip was located adjacentto flawed zones, which had been previously detected during nondestructivetesting of sample material, or such that the crack had to intersect flawedmaterial zones during fracture of the sample. The flaws were actually detectedon the fracture surfaces after the Kic-test; however, the measured fracture-toughnesses were always located in the same scatter band as the test-values offlawless samples. This is understandable, since the notch effect of pores justcannot be greater than that of natural cracks in CT- and NDT- samples. Thedecision whether a porous forging is useable can be based on a conservative andsufficiently safe assumption that the effect of porosity is identical to thatof cracks. The presumption of further impairment of the material characteris-tics is unfounded.

All too often the fact is being overlooked that modern calculation methods(that is the Pellini-FAD-, as well as the Kic concept) are based on the pres-ence of real cracks in the component part from the beginning. The conditionsin forgings with distinct segregation of the alloying elements must be judgeddifferently. It must be hereby realized that the assumed crack may be locatedin a material of considerably different chemical and physical characteristicsthan those upon which the calculation was based. The type and extent of alloysegregation in forgings of present production were discussed in the previoussection. This type of inhomogeneity cannot, or at least not adequately, bedetected with present nondestructive test techniques. Therefore, the majormetallurgical task pertains to the improvement of alloy-homogeneity in largeingots. The effect on the operational mechanical characteristics caused byunavoidable inhomogeneities because of the chemical composition requires fur-ther research effort. Our fracture-toughness experiments, which were startedin this area, have not yet been completed.

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POROSITY AND CRACK-GROWTH BEHAVIOR

Investigations to evaluate the effect of inhomogeneities on the crack-growth behavior were also initiated with simple model-experiments on artifici-ally flawed specimens. Figure 43 shows an example. The cross-sectional testarea of the standard crack-growth specimen (specimen type CT2, ASTM-E 399)received holes in a systematic pattern normal to the crack-plane. Figure 43shows a fractured CT-2 specimen resulting from this experiment. The fatiguestress was interrupted several times during the crack-growth experiment, suchthat crack arrest lines could form on the crack plane. The right side ofFigure 43 illustrates an enlargement of the crack surface. The crack, formedas a result of the fatigue stress, runs from the bottom of the picture upwardsand stops just short of the fourth row of holes; the remainder of the fracturesurface represents final static failure of the specimen. The crack arrestlines reveal that the fatigue crack has advanced faster in the web materialbetween holes and has circumvented the bore in a sickle-shaped manner. Thebore surface, which may be equated to the cavity surface of correspondinglyoriented porosities, have arrested the crack and slowed down the crack-growth.

In an additional series of model experiments, the holes were oriented inthe crack plane transverse to the direction of crack-growth. In this case,the cracks always progressed only up to the first bore, where they were com-pletely stopped ("drilled out"), since the bore surface has a much smallernotch effect than the natural fatigue crack. In another experimental series,the specimens were first drilled transversely, as in the previous example;and then forged until total compression of the bores had occurred. The speci-mens were then reheat treated and finish machined. In this manner, crack-growth specimens were prepared with several crack-like flaws, which werelocated sequentially in series within the crack-plane. Comparative experimentswith similarly forged and heat treated specimens but without crack-like flawsyielded the same crack-growth rates when referenced to the respective cross-section of the specimen. The nominal cross-section of the specimen was natur-ally traversed at a correspondingly increased rate of crack growth, as thepercentage of the flawed area increased with respect to the total cross-sectional test area. For a component part containing porosity, the conclusioncan be drawn that the crack-growth rate will increase only, when the pore orpore concentration becomes so large that the nominal cross-section of the partis considerably reduced at this point and therefore the stress-intensity hasincreased. The material characteristics "crack-growth behavior" will not bealtered by porosity.

SEGREGATIONS AND CRACK-GROWTH BEHAVIOR

In addition to the model experiments on artificially flawed specimens,crack-growth experiments were performed on specimens with natural segregationsand inhomogeneities. As discussed in the section "Inhomogeneities of Rotor-Forgings", the manganese-sulfide segregation in forgings, which is easilydetectable with the Baumann-Print, very often has a close, spatial interrela-tion with the segregation of other alloying elements. For this reason, theBaumann-Print and the results of nondestructive tests were introduced as sup-plementary means to evaluate sample material removal of the crack-growth

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specimens. Figure 44 shows the fracture-surfaces of two creck-growth specimenstogether with the respective Baumann-Prints, utilized to select and orient thespecimens, In addition to the MnS stringers, which are recognizable on theprints, the sample material contained a large number of small ultrasonic indi-cations, the cause of which is presently still under investigation. Thestringer-shaped markings on the Baumann-Prints can also be detected in similarform on the fatigue crack surfaces of the specimen, although the prints couldonly be taken at a certain distance from the crack surfaces. It is certain,however, that the sample material has inhomogeneities. The correlation ofsegregation and crack-growth directions on Figure 44 indicates that the crack-growth specimens were extracted longitudinally and transversely to the axis ofthe forging.

A large axial trepan was available from the 2010mm diameter LP turbineshaft No. 49424 of a 1000 rpm railway-turbine-generator. The shaft was manu-factured from a 216 ton ingot of modified C-steel with 0.22% C/ 0.46% Cr/0.42% Ni and haat treated as follows; Normalizing 880 to 900°C 36 hours/aircooling, tempered 630 to 650°C 60 hours; furnace cooled. Prior to performingthe axial-bore, which was structurally necessary for manufacturing this torsionshaft, numerous ultrasonic indications were detected locally at the center lineof the shaft. Sample material with and without indications was obtained fromthe core for crack growth investigations.

The results are compared in Figure 45, which represents two crack-growthdiagrams log da/dN=f (log (K). The left diagram applies to the indication-free CT2 specimen. The measured values fall within the range having a crack-growth exponent of m=3. The range limits were transposed to the right diagramtogether with the measured values of the longitudinal and transverse specimensfrom the flawed core zones. Both specimens shown on Figure 44 are alsoincluded. For practical purposes, complete agreement exists in the behaviorof specimens with and without indications. The markings, which can be recog-nized on the fracture surfaces of the specimens shown in Figure 44, have noeffect on the crack-growth behavior of this material.

The same results were obtained from experiments with flawed specimens froma 620mm diameter journal of a generator rotor forging which corresponded to thebottom end of a 100 ton ingot, 26 NiCrMoV 14 5 steel (melt 17704). In additionto flaw free zones, this journal had displayed zones containing ultrasonic indi-cations having an equivalent flaw size of 2 to 8mm diameter, Crack-growthspecimens were extracted from both zones. The results are included on the righthand diagram of Figure 18, The resultant values obtained from the flawed speci-mens having fracture surfaces containing markings similar to the specimens shownon Figure 44, do not differ from those of the ultrasonically clean specimens;the fatigue crack surfaces of the latter were completely smooth in all cases.

Investigations are still underway to determine the nature of the actualsegregations and flaws of these specimens. In addition, there are certainlymore unfavorable pieces. From presently available results, which are still in-complete, it can be expected that the unavoidable inhomogeneities in large forg-ings occurring because of alloy segregation will not materially effect thecrack-growth behavior of rotor steels.

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Based on existing research of a fundamental nature and present largecapacity forgings, distinct guidelines can be established for a justifiable

delivery specification and acceptance test of materials and productioncomponents.

ACCEPTANCE CRITERIA

The delivery specification requirements are based on the plan which isrepresented in Figure 23. It becomes apparent after the preceeding statements

that the determining factors are:

the 0.2 yield strength andthe NOT-temperature at the axial trepan,the residual stresses,the ultra sonic test results,the homogeneity test by means of radial trepans 1 and 2,with fractography at locally reduced deformation.

The evaluation of the results is to be accomplished in its entirety; that

is the results may be to a certain extent compensated against each other.

Since LP-turbine-and generator rotors are treated exclusively in this

paper, the delivery specifications should contain, in addition to the steel

type, only guideline standard data about

the melting,the chemical composition,the forging practice,the heat treatment andthe surface samples

since the true material characteristics will be measured on the delivered item.

This naturally does not apply to the same extent to rotors which are produced

from heat-resistant steels.

TEST RESULTS OF LARGE CAPACITY ROTORS

Presently available test results which pertain to the previously mentioned

large capacity rotor forgings by Japan Steel Works, will now be discussed.These forgings reach a weight of up to 200 tons and have never been manufactur-

ed before. Required for all generator-rotors, and therefore specified ascritical quality factors were:

0.2 yield strength of trepan ?Hi kp/mm2, at slot-bottom ?.46 kp/mm2,

NDT - temperature 5.4-15°C at axial trepan,

residual stress, axial <-5 kp/mm2

equivalent flaw size relating to ultrasonic single indications<5mm diameter,

- 25 -

DEVELOPMENT PROGRAM

bore not required, but permissible up to 400mm diameter.

Material separations revealed by homogeneity testing are notpermissible.

As previously discussed simultaneous answers should be obtained if possi-ble, even if only partially, to the following questions:

Now does the homogeneity compare for forgings produced from 400/500 ton ingots

to present smaller ingots up to 200/300 tons?

What improvements are possible with the "JSW-A.P. Process"?

What is the largest rotor weight that can be produced by this process with

sufficient quality and adequate delivery times?

What are the optimum properties that can be expected from the internal struc-ture of this largest forging with respect to 0.2 yield strength and fracturetoughness?

For this purpose, not only the standard test plan was altered and addi-

tional examinations performed, but also some of the more recent rotors were

provided with axial bores in order to obtain comparative sample material inorder to evaluate the initial rotors. In the delivered condition, the firstrotor 2.7901 was subjected to all non-destructive tests (ultra sonics, magnetic

particle, penetrant test, borescopic examination) as well as Baumann-Prints,

structural examinations, microfractography, 50 tensile tests, 219 notch impact

tests, 65 NDT tests and 21 fracture toughness tests. In addition, a block wasforged from surplus material which was obtained from the top end of the same

ingot. This block had a forging weight of 14 ton and a dimension of 1020 mmdiameter by an 1160 mm length. It was subjected to a simulated heat treat

process and served as a source for 26 fracture toughness samples up to a sizeCT 5 Figure 11.

COMPARISON ROTOR

Prior to discussing the JSW-forgings, the test results of generator rotor7642 will be covered in order to establish a so-called basic reference. This

rotor was fabricated from a 300 ton ingot and was up to now the largest forging

which we have inspected in similar manner. According to Figure 46, this rotorhas a 100 mm smaller body-diameter, a delivered weight of 139 ton and was manu-factured in 1968 from ASTM A469 Class 4 steel, which was customary at that timeand corresponds to type 24 Ni Mo V 14 5.

This chromium free steel displays a magnetizability at the upper limit ofthe total spread. The narrow band, however, applies only to 1 producer is).As already discussed, the more favorable magnetizability of the chromium freesteel is obtained at the expense of reduced tempering characteristics and

- 26 -

consequently adverse fracture-toughness (Figure 15). For chromium containingsteels with 2% Ni (26 Cr Mo V 85) and with 3 1/2% Ni (26 Ni Cr Mo V 14 5) acommon scatterband occurs, which agrees closely with international evaluationsof 3 1/2% Ni Cr Mo V-steel. The average value of the 3 1/2% Ni Cr Mo V-steel,however, is somewhat more favorable than that of the 2% Ni Cr Mo V-steel. Thetest results of 3 each JSW-rotors of both steel grades are plotted with theirindividual values on Figure 47. In all cases they are located above thecorresponding average value.

Figure 46 indicates that the radial trepans k and f could not be removedfrom rotor 7642 at a length of 0.6 X D in accordance with the standard test plan(Figure 23). Their length of only 500 mm (20") does not suffice to cover thezone of maximum segregation. The trepans k' and f' of 125 mm (5") lengthcover only the segregation free outer zone. Accordingly, fracture I of alltrepans also occurred with good deformation characteristics (6/4).

The ultrasonic inspection yielded indications from the rotor center, whichlead to the decision of boring to a 325 mm diameter. Subsequent to tempering,the extracted axial trepan was again subjected to an ultrasonic test whereby thesample AXM (Figure 46) was positioned at the zone of greatest indications.As expected, the removed axial samples yielded also good deformation values,however not very good fracture toughness. (NDT-temperature + 10°C, Kic seeFigure 15). The radial samples displayed zones of locally reduced deformationas a result of inclusions in an otherwise ductile matrix. Refer to the micro-fractography on Figure 48, Figures 48 a/c also show porosity at the axialcenter with an associated Mo-concentration of 0.59% with a corresponding in-crease in hardness at the "V" segregation as compared to 0.36% in the melt. Thetest results of rotor 7642, which was fabricated from 24 Ni Mo V 14 5 steel,may be considered as normal, however with the important exception that no radialtrepans were tested which had a length greater than 1/4 body diameter.

LARGE CAPACITY ROTORS

The melt analysis and final heat treatments of 9 generator and 2 LP-turbine rotors are compiled in Table 6. The quality level which applied to1500 rpm was obtained for the first 4 pieces by way of conventional pouring of26 Ni Cr Mo V 85. In accordance with the development program, rotors 5 and 6were poured in 26 NiCrMoV 14 5 per the "JSW-A.P. Process", rotor 5 with Si-deoxidation and high tempering temperature, rotor 6 with vacuum deoxidationand low tempering temperature. On the basis of prior experience, rotors 7through 9 were poured in 26 NiCrMoV 11 5 steel. The following specificationsapply to the axial trepan sample AXF: a 0.2 yield strength a52 kp/mm2, for theslot bottom ?56 kp/mm2 and an NDT-temperature of .S-100C. These specificationswould also suffice for 1800 rpm, such that a double back up is not required.

A 0.2 yield strength of >13.5 kp/mm2 at 300°C was specified for the LP-turbine shafts with only 1000-rpm, Figure 5. This yield strength had to beobtained by means of a normalizing process for the first shaft 10.47377, sincespray cooling was not possible at that time for the large diameter of 2780mm.The second shaft was provided with a higher Ni content in order to obtain better

- 27 -

toughness values. In the meantime, however, equipment has been procured forspray cooling of these shafts.

The essential test results of the first 6 generator rotors are compared inFigures 49 through 56. Since all the individual values could not be listed,only the results of fracture I were tabulated for the radial trepan samples.The most and least favorable values were listed each time for the axial trepansamples Figure 55/56.

The data of greatest interest in Figure 4'9 are those of the residual stressmeasurements. Together with the producer we performed 2 measurements each inthe tangential and axial direction. No measured value exceeded -5kp/mm2. Thespecified value of .5...-6kp/mm2 can be met even for these large dimensions.

The data comparison also indicates that only the first two rotors, whichwere produced from conventional ingots, received an axial bore on the basis ofthe ultrasonic inspection. As mentioned already in conjunction with Figure 31,slight porosities and MnS-inclusions were detected. Rotor 4,7933 displayed asomewhat smaller maximum equivalent flaw size, which corresponded to theultrasonic indications. For this reason a bore was not necessary. In itsplace, in order to gain further knowledge, 6 radial trepans were removed fromzones of the highest incidence of ultrasonic indications. The location of thetensile specimen fractures was discussed already in conjunction with Figure 27.The fourth rotor also had indications up to the specified maximum limit.Since the properties of the individual forgings were compatible with each otherand sufficient test results were available, especially of fracture toughness,no bores were made thereby preventing unnecessary stress concentrations. Onlya shallow bore was made to obtain the sample material; this, however, was ac-complished at the opposite body end, where only minor indications were dis-covered (Figure 50). For comparison testing an axial trepan was again removedfrom both rotors 5.8154 and 6.8189, since these were manufactured in accordancewith the "A.P. Process" with and without vacuum-deoxidation at varying tensilestrength levels.

Figures 51 and 52 contain the results of the first tensile fractures ofthe 540 mm (21-1/2") long radial trepan outer halves. When the fractureelongation values are compared with the results of the tangential outer samplesin Figure 49, there is practically no difference, which means; these largeforgings do not possess greater inhomogeneities up to a depth of approximately1/4 body diameter, not even in the highest stressed zone of four-polegenerator-rotors.

Results of the greatest interest to the development program were to beexpected from the inner halves of the long radial trepans. These results offractures I are compiled in Figures 53 and 54. Fractures with reduced deforma-tion received special emphasis. The respective fractographic examinationresults were already discussed in conjunction with Figures 28, 29, 30. Theinhomogeneities, which were discovered here, are of the same type as are knownfrom other ingots; their extent, however, is even less than that of many, very

- 28 -

much smaller units. This result confirms that starting at a certain ingot sizebelow 100 tons, the inhomogeneities by type and extent do not necessarily haveto become greater in large ingots than are present in small ones.

Finally, the results of the axial trepans of 2 rotors each, produced fromconventional and "A.P." -ingots are compared in Figures 55 and 56. It must berecognized here that varying Ni-contents and tensile strength values are pre-vailing. This was done intentionally, but it affects the comparison. As ex-pected, the fracture elongations are affected in the radial direction by the"V" - segregations. The decisive factors are fracture toughness and NDTtemperature, which were obtained from the interior of these large cross-sections.The worst case value in the radial direction for the 2% NiCrMoV steel correspondsto the respective value of rotor 7642 (Figure 46), which was fabricated from3-1/2% NiMoV steel. The 3-1/2% NiCrMoV steel, especially the one obtained fromthe "A.P." ingot with vacuum-deoxidation, displays excellent fracture toughness,even at high tensile strength values. Consequently, there are no difficultiesin obtaining the necessary quality characteristics for forgings of these ultimatedimensions, even within the internal structure, when suitable pouring - anddeoxidation - processes are applied.

SUMMARY AND FUTURE PROSPECTS

Forgings, which we test today and release for generator and turbine rotors,have an entirely adequate quality, thanks to the developments which occurredduring the past approximately 15 years. With this, we confirm a statement madeby Mr. Creenberg 23) at the 1970 session in Terni. The attainment of thisquality was considerably supported by the introduction of vacuum-degassing,ultresonic testing and fracture mechanics. We also confirm, however, therequest by Newhouse and Deforest 24) at the same session, whereby forgings willbe needed in the 80's with weight up to approximately 250 tons from 500 toningots for 4-pole generators, while forgings for 2-pole generators will requirea higher tensile strength and a higher fracture toughness with improvedhomogeneity.

The prerequisite for attaining this goal is a realistic evaluation ofinhomogeneities and stress concentrations and a comparison in their effect onthe actual stresses occurring during long-term machine operation. On the otherhand, a full utilization of all metallurgical sources is needed to reduce thesegregations. Our investigation has shown that normal single inclusions andporosities, which are detected ultrasonically, had in practice no effect on thefracture-toughness, the crack-growth behavior and the NDT-temperature.

Local concentrations of alloy elements, especially Mo, were also detectedwith associated increased hydrogen content and inter-crystalline fracture forma-tion at room temperature. These cannot be detected by ultrasonics. Their effecton the operational characteristics has to be further investigated. Their origin-ation can evidently be effected by the ingot geometry, pouring sequence, deoxida-tion and other metallurgical means. At any rate, there are large conventionalingots (400 tons) at present, which display less segregations than small ingots(100 tons) having the same overall alloy composition.

- 29 -

Ingots which have been manufactured by the electro-slag-remelting process(ESU), display an improved homogeneity. The present state of the art iscapable of producing ingots of approximately 1000/1500mm diameter with a weightof 20 to 50 tons. A research installation is presently being operated in theFRG (Federal Republic of Germany), wherein the technology is being developedfor ingotSiZ8S of up to 2300/2500mm diameter and weights of 80 to 160 tons.In the event that this effort is successful, installations will probably beplanned for ingots of approximately 3000mm diameter and a weight of 300 to 350tons. On the other hand, ESU-ingots of medium size can be welded together andthen be forged into larger rotors. In Japan, large and more homogeneous ingotshave already been manufactured by the "A.P. Process". Based on the presentlyavailable results for generator and turbine rotors, which were presented herepreviously, this process is very promising.

In order to arrive at a more realistic evaluation of forgings, a change inthe conventional specifications, presently based on standard test values, isneeded with a view towards evaluation of true operational characteristics. Aspecimen should be developed within the dimensions of standard radial trepansfor the determination of fracture toughness, thereby reducing the operationalstresses by complete elimination of axial trepans below the body.

-30 -

Literature References

Literaturverzeichnis

1.) Leukert, W. und Raymund, H.:GroBstromerzeuger.- Die Entwicklung der Starkstromtechnikbei den Siemens-Schuckertwerken.Siemens-Schuckertwerke AG., Berlin-Erlangen, 1953.

2.) Krick, N. und Hiebler, H.:Generatoren far Kernkraftwerke.Technische Rundschau, Heft 52 vom 10. Dezember 1971.

3.) Irwin, G. R.:Crack extension force for a part-through crack in a plate.Journal of Applied Mechanics, Vol. 84E, No. 4,December 1962,pp. 651-654.

4.) ASTM-Committee on Fracture Testingof High-Strength Materials,5th Report 1964:Progress in measuring fracture toughness and using fracturemechanics. Materials, Research and Standards 4/1964, March,pp. 107-119.

5.) Greenberg, H.D., Wessel, E.T., Clark, W.G. und Pryle, W.H.:Criticalflaw sizes for brittle fracture of large turbinegenerator sotor forgings.Atti Del 5 Convegno Internazionale della Fucinatura,Terni/Italien, 6.-9. Mai 1970, S. 71-105.

6.) Wolf, H. und Bdhm, W.:Das Ring-Kern-Verfahren zur Messung von Eigenspannungen undseine Anwendung bei Turbinen- und Generatorwellen.Archiv fUr das EisenhUttenwesen 42/1971, Heft Marz, S. 195-200.

7.) Greenberg, H.D., Wessel, E.T. und Pryle, W.H.:Fracture Toughness of turbine-generator rotor forgings.Engineering Fracture Mechanics1/1970,April, pp. 653-674.

8.) Schinn,R. und Schieferstein, U.:Eigenschaften warmfester 1% CrMoV-StUhle fur HD- und MD-Turbinenwe;len.Atti Del 5 Convegno Internazionale della Fucinatura,Terni/Italien, 6.-9. Mai 1970, S. 25-70.

9.) Paris, P.C.:The fracturemechanics approach to fatigue.Proceedings of the 10th SagamoreArmy Materials ResearchConference, August 1963 (Syracuse University Press, 1964).

10.) Clark, W.G. und Trout, H.E.:Influence of temperature and section size on fatigue crackgrowth behavior in NiMoV-alloy steel.Engineering Fracture Mechanics2/1970,November, pp. 107-123.

11.) Pellini, W.S. und Puzak, P.P.Fracture analysis diagram procedures for the fracture-safeengineering design of steel structures.U.S. Naval Research Laboratory, NRL-Report 5920, March 15, 1963,and Bulletin of the Welding Research Council 88/1963, May.

12.) Loss, F.J. und Pellini, W.S.Coupling of fracture mechanics and transition temperatureapproaches to fracture-safe design.U.S. Naval Research Laboratory, NRL-Report 6913, April 14, 1969.

13.) Pouillard, E.:Quelques Aspects Des Problemes Metallurgiques Poses Par LaFabrication Des Rotors D'Alternateurs.Journees Internationales de la Grosses Forges Francaises,27.-30. Mai 1963.

14.) Bastien, P.G. und Rogues, C.:Etude de l'Heteregeneite des Gros Lingots de Forge.Atti Del 1° Conyegno Italiano Della Grossa Fucinatura,Terni/Italien, 26.-29. September 1961, S. 29-65.

15.) Kawaguchi, S. und Kudo, K.:Segregatioe and Heat-Treatment of Large Forgings.Atti Del 5 Conyegno Internazionale Della Fucinatura,Terni/Italien, 6.-9. Mai 1970, S. 215-265.

16.) Ammareller, S.:Seigerungen und Festigkeitseigenschaften eines Schmiede-stUckes aus einem 56 t-Block.Stahl und Eisen 70/1950, S. 125-133.

17.) Knorr, W.:Beeinflussung von Blockseigerungen durch die Vakuum-Kohlenstoff-Desoxydation.Unver6ffentlichter Bericht der Fried. Krupp AG.,24. Februar 1972, S. 1-15.

18.) Smith, H.C. und Hartman, G.3.:Manufacture of Large Generator Rotor Forgings over 135 MetricTons.Atti Del 5° Conyegno Internazionale Della Fucinatura,Terni/Italien,6.-9. Mai 1970, S. 679-710.

19.) Scalise, V.:Unver6ffentlichter Bericht der ITALSIDER, Febr. 1970.

20.) Opel, P.:Unyer6ffentlichter Bericht der Stahl- und R6hrenwerke Reis-holz GmbH., DUsseldorf 30.1.1963.

21.) Hochstein, F.:Discussion on the papers presented at the

Atti Del 50 Convegno Internazionale Della

Terni/Italien, 6.-9. Mai 1970, S. 123-126

22.) Kawaguchi, S. und Mitarbejter:

Forgings from Gigantic Ingot with 3550 mmDiameter and 400 t (881,000 pounds) WeighPart 1: Production and Metallurgical DeveSixth International Forgemasters Meeting,1.-6. Oktober 1972.

23.) Greenberg, H.D.:Discussion on the papers presented in theAtti Del 50 Convegno Internazionale DellaTerni/Italien, 6.-9. Mai 1970, S. 888.

24.) Newhouse, D.L. und Deforest, D.R.

Meeting Requirements for Larger Generator

A Metallurgical Challenge.

Atti Del 5 Convegno Internazionale DellaTerni/Italien, 6.-9. Mai 1970, S. 739-757

1st session.Fucinatura,

(140 inch)t.lopment.Cherry Hill/USA,

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62

TABU 1 PRODUCTION OF GIGANTIC ROTORS

( ). QUOTED ONLY

PRESSCAPACITY

MAXIMUM INGOT

AVERAGE TOTAL0 WEIGHTmm

MAXIMUM ROTOR

SHIPPED OUTERWEIGHT 0

mm

750 1.0 a

875 1.5 a

LP-ROTOR t MAX AT 3000 RPM

TYPE WITHOUT WITHBBORE ORE

TABLE 2:SIZE AND NECESSARY STRENGTH FOR LARGE GENERATOR MONOBLOCKROTORS1) MULTIBLOCK 2) 79 kp/mm

2FOR 75 % UTILIZATION

3) 1955 ROTOR 4318 FIGURE (PAGE ) 4) WITHOUT BORE

MATERIAL

2.0 a 34 NiCrMo 74

26 NiCrMoV 85

26 NiCrMoV 145

3.0 a 34 NiCrMo 74

26 NiCrMoV- 85

26 NiCrMoV 145

TABLE 3: STRESS-4ATERIAL-DEFECT SIZE-RELATION

LLOWABLE FLAW SIZE+)

WITHOUT WITHBORE BORE

6 D

30 D

98 D

3 D

13 D

45 D

1 D

4.7 D

16 D

0.3 D

1.4 D

4.7 D

+) RATIO OF DIAMETERS D OF EQUIVALENT PENNYSHAPE DEFECT SIZE

Symbol1) Rotor Body KIc-Test Location2) 0,2% Y.S.

Dia. at 20°C(mm) (kp/mm2) rad.

LS 72276 1000 core 59 +10

7776 It core 58 -15 -20/-25

2.7901 1800 T,tang.,rim,top 59 -15

Ci " AXM,ax.,core,middle 40 +0/+50 +0/+5--115/+10

" AXK, " " ,top 50 +15/+20 +15

7901 core,rad. + tang. 53-59 +5 +5Test block3)

3.8153 1800 T,tang.,rim,bottom 63 -45

" AXF,ax.,core " 43 -10

AXK, " ",top 42 +15 +15

1) see Figure 16; 2) see Figure 49 and 55; 3)see Figure 14

3,5 % NiCrMoV-STEEL 26 NiCrMoV 14 5

NDTT (°C)

ax. tang.

:35

CD

CONV. = CONVENTIONAL

A.P. = AFTER POURING

TABLE: 6

1970

1915

1970

1930

1972

1928

-rt I -1 r s 1I g. 4.LJ

13700

14 5295

FIGURE 1: FORGINGS FOR GENERATOR ROTORS WITH 3000 RPM

01390

10760

13660

14880

MULTIBLOCK

10350

0770

01000

0 1150

0'1800

01800

- a

FIGURE 2: FORGINGS FOR GENERATOR ROTORS WITH 1500 RPM

9.4 MVA9.5 t

80 MVA[33 t]

1000 MVA80 t

—011

100 MVA

][421]

1070 MVA172 t

1500 MVA197 t

Nr MVA

7901 15008153 118183 118307 11

7933 105079348154 11

8309

8504 2000

400t

1971 1972 1973 1974

FORGING PRODUCTION

MED GENERATOR PRODUCTION

FIGURE 3: TIMING OF GIGANTIC FORGINGS

1500 /1750 MVA

-.1800- - -

1000/1475 MVA

16 780

500t INGOT

FIGURE 4: GENERATOR ROTOR 240 t

11111111

EZ:Z=

50/60 HZ

PZZ

r2

14 ACCEPTANCE t= TRANSPORT

500 tBACK UP AND DEVELOPMENT FORGING

2000 /2300 MVA

7780

F I

F2

oN TEMPERATURE READING

• RESIDUAL STRESS MEASURING

FIGURE 5: LP TURBINE SHAFT FOR 1000 RPM

TEMPERING TEMP IN °C

-1 7 - I 5 - I 8 -I 3

600 - 4 - 2 - I - 1_.+2 7 oc 2.5°/H

o - 5 - 4.2 - 4.3 - 3.7

500\te83°C\ 2 5°C/H

e400 e

TEMP IN CENTER6°C/H

SPEC <8°C/H 11 ON SURFACE

3000 20 40 60 80

COOLING TIME IN H

FIGURE 6:RESIDUAL STRESSES IN ROTORS

0 2780

F3 •=1> 0 600

T

400t INGOT WEIGHT

I65t SHIPPING WEIGHT

RESIDUAL STRESSES IN KP/MIM2

1 11a.AX aTANG a'AX cYTANG

SPEC. < 6 < 6 < 6 < 6

-19 -17 -19 -146°/H

h2c0 2a

-f

Gti at2

r > > ra:

ari=°' ati

at2--= 2'ati

0 a3 = atati

2.5...-../..- ...- 1 •

/ I •

r / ari Crr2 \ r r

2c -01±

C/t1

a3

FIGURE 7a: FIGURE 7b:

STRESSES IN ROTATING SHAFTS

rj « ra 012c 0.11 Q zconst

0 0

mE2c

1-2a

2 2 2

1 Ka _1.

Klc'Q lc'Q a -1Kic • Q

crt ir . 0.2 a cr2 ' c r3

ti1.21 w•cou

3

1 1a - --- a a Cr3 - — a

cr2 - CM4.84

- cri6.25

FIGURE 8:CRITICAL FLAW SIZE IN ROTATING SHAFTS

STR A I NGAGE

44, -012 lc-- 0 14/ / / / / / / 77/ /

non I 9 RELAXATION GROOVE FOR STRESS MEASUREMENTS

RE

SID

UA

L S

TR

ES

SE

S I

N K

P/M

M2

- 8

fr98

h95

h

6 M

o =

1.1%

A

A

A4

0 0

4•

00

•fi

9•

A54

h-4

•..-

-4,

• •

•.6

•

•0

•*15

h A

•

E.

o.

I

-24

A

0 0

I.

610

640

670

710

900

1200

15

00

1800

2600

TE

MP

ER

ING

TE

MP

ER

AT

UR

E IN

°C

B

OD

Y D

IAM

ET

ER

IN

M

M

GE

NE

RA

TO

R R

OT

OR

S .

AX

IAL

ST

RE

SS

:

• N

i(Cr)

MoV

- S

TE

EL

o C

- S

TE

EL

TU

RB

INE

RO

TO

RS

T

AN

GE

NT

IAL

ST

RE

SS

:A

Ni C

r h4

oV-

ST

EE

L 0

CrM

oV-S

TE

EL

FIGURE 10a

FIGURE 10b

RESIDUAL STRESSES IN ROTOR FORGINGS

generator rotors, axial stress

: S Ni(Cr)MoV

steel

0

C-steel

turbine rotors, tangential stress : A

NiCrMoV - steel

0CrMoV - steel

FIGURE 11: COMPACT TENS ON SPECIMENS FOR FRACTURE TOUGHNESSKIc.

FRACTURE TOUGHNESS TEST INSTALLATION.

0 990BOTTOM

TEMPERATURECONTROL

0 1020

[ 550°C /0.o. + 850°C 600°C/F.C.]1250 4240

I I II I I

G=7.41

-± 0- 10

0

01120

[ 4670°C /F. C.]

WO 200 WO 400 WOMM FROM ROTORSURFACE

CT-SPECIMENS

FIGURE 13: FRACTURE TOUGHNESS TEST MATERIAL, ROTOR 7776,2 % NiCrMoV-Steel 26 NiCrMoV 8 5.

,

1160 °C

+ 40

+ 20

0'02x575 x59.5 x 55.2

-

TOP

Grz5E3t

x575 kpimm2FATT rad.

NI:ITTrod.

4— -----------100 200 300 400 500

MM FROM ROTORSURFACE CT-SPECIMENS

FIGURE 14: FRACTURE TOUGHNESS TEST MATERIAL, TESTBLOCK FROM 400t-INGOT FOR ROTOR 2.7901,2 % NiCrMoV-STEEL 26 NiCrMoV8 5.

KlcK5141rn. kp/mm3/2

3.5 NiCrMoV610.2=79.5-97.7 kp/rnm2

°x x

/

/0970)

150 x

1

0 x xx 1 2 NiCrMoV

x x I a02. 54-63kp/mm2

400

° ox x x0

+ /+ / 1(1961 0)

x , /,

100i 1

+ i iNDTT ,+ 3 NiMoV

00 + % ,0027 60 kp/mm 2,,/

xo

x (§+ .4»:

+ ,

x /

---+ 4,01.<1DTT

0-x ± Wessel et al.,1969---------o n KWU

2050 x

OXX

0 0-180 -120 -60

Temperature in

tO .60 +120 °C

-240 -120 tO +120 .240 °F

FIGURE 15: FRACTURE TOUGHNESS KIe OF SEVERAL STEELS FOR LP-TURBINE

SHAFTS AND -DISCS AND FOR GENERATOR-ROTORS.

3,5NiCrMoV = 26 NiCrMoV 14 5SHAFTS MAX.1500 mm DIA.(SPECIMEN POSITION 1/4 x DIA.BELOW SURFACE) AND TURBINEDISCS WITH 250 TO 680 mm HUBTHICKNESS.

3 NiMoV - 24 NiMoV 14 5 (ASTM-A 469, CLASS 4)ROTORS MAX. 1700 mm DIA., INGOTWEIGHT MAX. 300 t,SHIPPED WEIGHT MAX. 139 t.

2 NiCrMoV : ROTORS MAX. 1000 mm DIA.,1960-PRODUCTION : STEEL 34 NiCrMoV7 4,1970-PRODUCTION : STEEL 26 NiCrMoV8 5.

Temperature in0 0

-180 -120 - 60 ± 0 +60 +120 °C

-240 -120 ±0 +120 +240 °F

FIGURE 16: FRACTURE TOUGHNESS KIo OF

2 % NiCrMoV-STEEL 26 NiCrMoV 85

KKSIATIR kp/mm312

150

100

50

Kic

kp/rnm3/2

160 ROTOR L OC AT ION

A 72 226 0 1000, CORE

A 7 776 " II

19 2 7901 tb 1800. TANG ,RIM. TOP

0 . II , A X CORE MIDDLEO u u n 0 . TOP

O 7901 TEST BLOCK,

TANG L RAO , CORE <I>

400 0 3 8153 0 1800 u RIM,BOTTOM

O .1 A AX CORE, il 0 0 A10 o . il II II .TOP

eO tifr

a&

A

II A A 0 S O A

AO GI 0 / i NDTT :s. +15°C. - o...,,,,,,• //

ai A .200 00 0

50 _t: jr,e-r. A 0 0

0 o 0 ---0

— 24NiMoV 14 5(ASTM-4469,CLASS 4)

150

600

400

200

o No

--------

0 0-180 -120 -60

-240 -120

0 9 0 0

NDTT

—A

0 ib o

o /oe o 26NiCrMoV 85 15. 018001

CORE,MIN Kjc24NiMoV 145 951700

Temperature in

±0 +60 +120 °C

±0 +120 +240 °F

FIGURE 17: FRACTURE TOUGHNESS EIcOF

3,5 H.%NiCrMoV-STEEL26NiCrMoV14 5

NOTT -0°C

ROTOR LOCATION

• 5.8154 551600, TANG.,RIM, BOTTOM

O it i AX, CORE MIDDLE

• 6.8189 it . TANG.. RIK TOP

o ti it , AX. CORE, BOTTOM

•

o n , MIDDLE

, TOP

DISKS 0 3000

A 44096 385mm THICK. HUB

44351 450 it tt

.144344 440 it ti ti

47 075 250 it tt It

20/c,Ni Cr MoVda [10°6mn1dNL LW

3000

0 ROTOR 2 7901AXK

1000 dol =45-51 kpirnm2

300

100

30

10

mr3

30 50 100 2050 100 200 500 30 50 100 200 500 1000aK[kp/mm312]

FIGURE 18: FATIGUE CRACK GROWTH RATE OF 2-3,5 % Ni—STEELS.

COHERACYr .g0.005 mm

FIGURE 19:

SPECIMENS : CT1- 2 , WOL IT- 3T / ROOMTEMP (AIR) / f =10-160 HZ

NATURAL CRACK

3•/. NiMoV 3.5%NiCrMoV

mr. 3

it(8g

000

0

I 6°.44 *10

ROTOR a0.2 [kplaim2] /

/ 0

ROTOR(10.2

[kp/mm2]• 8887 52 0 3178 79G 124 J357 57.5 / 0 ZV3037 110O DISC B 59.5 / / • 6.8189AXF 63

/0 /7704 744- il (INDICATIONS)

400pm,

150-V0.25 mm

MACHINED NOTCHES(IMPACT TEST SPECIMENS)

DVM- R1 ram

°(rn=3)

qr.°/ 0

CRACK AND IMPACT SPECIMEN NOTCH.

DVM-F4 mm

Kawaguchi/Hu &

INGOT

FIGURE20a:SEGREGATION PATTERNOF INGOTS

k /%Diluted

/I1 Core 1I Iit /I /I /% /\ /‘ /.„

Shell

TOP

BOTTOM

C= 0.3119 (MAX.)P=0.012% (MAX.)

C=0.24% (MIN )

C=0.30%(MAX.)

A A AJ V 1../

\IV .....

VP=0.005% (MIN.)

tSURFACE tCENTERROTOR

LADLE ANAL.: C=0/7% P=0.009%

FIGURE20b:C/P-SEGREGATION IN A ROTOR

JAPAN STEEL WORKS

FIGURE 21SCHEMATIC PATTERN OF INGOT MADE BY AFTER POURING PROCESS

TOP

K2 F2

Xl F1

CHARPY—V AT-20°C0.2% Y S.

/

U.S T 1 2

NDTT0.2% Y S.

7750

FIGURE 22 : C/P—SEGREGATION IN ROTORWITH EXTREME HIGH C—CONTENT

0 200 400 600 800 0 200 400 600 800

DISTANCE IN MM FROM ROTOR SURFACE

to.

1+2 FAT T

FIGURE 23 : ACCEPTANCE TEST FOR GENERATOR ROTOR

BOTTOM0 2°Io V. S

7.0

/ FRACTURE SURFACEINSPECTION

0.2% Y. S.

16

20

16

60

61

54

FR

AC

T.:

II H

I IV

24:RADIAL TREPAN PIN POINTED

1ST

FR

AC

TU

RE

W

EA

KE

ST

PO

INT

115

18

54

57

4III

FR

AC

TU

RE

LO

CA

TIO

NP

RE

DE

TE

RM

INE

D

Rot

or

7557

E

CCORDING m0 MA

'I-INDICATION

a

FRACTDR—

x 5

x 100 c

CTOGRAPHY OE FRACTURE 1: FLAK

STUFF Y1 ULD o , ULT.TENcD.STR.y

26 kp/mm2

±2-,p/mm'' 77 kp/mm2

ROTOR 3599, 1155 mm 0, 44 t

INGOT 100 t

FABRICATION 1951

Si P 5 Cr Ni Mo

.31 .3 .019 .013 1.4 2 .44

x 400 d x 10 00

jUi‘L 25- /b/c/11: TREPAN PIN POiNTED ACCORDING MAX UT-INDICATION

ELCN1. 6_

7

RED.AREA

10

a

MnS AT SURFACE OF FRACTURE I

FIGURE26a/ b:

ROTOR 7557E0 1000 mm, 34 tINGOT 100 tFABRICATION 1970

MICROFRACTOGRAPHYOF FRACTURE IINCLUSIONS INDUCTILE MATRIX

x 600TREPAN PIN POINTED ACCORDING MAX UT-INDICATION(SEE FIGURE24)

FRACTURE YIELD a0,2 ULT.TENS.STR. aB ELONG. 55I - 73kp/mm2 1 %

II 74 kp/mm2 87 kp/mm2 15 %

Si F S Cr Ni Mo V

.23 .03 .004 .014 1.4 3.4 .42 .14 %

x 100

RED.AREA T

5 %54%

cs—

AR

EA

O

F U

T-I

ND

ICA

TIO

NS

a

SURFACE AT LOCATION OFSILICATE INCLUSION

FRACTURE YIELD a0,240 Np/mm2

40 Xp/mm2

x5

FRACTURE I OF TREPAN F 22 PIN POINTED ACCORDINGMAX. UT-INDICATION

x 200

MICROFRACTOGRAPHY

DUCTILE FRACTURE NEARINCLUSION

FIGURE28 a/D/c: ROTOR 4.79)5, 0 1800 mm, 172 t, INGOT 400 t

Si P S Cr Ni Mo V.25 .2 .009 .010 1.4 2 .4 .11 %

ULT.TENS.STR. aB ELONG. 88

58 Np0mm2 11 %62 Np/mm2 25 %

x 400

RED.AREA

37%64%

x400

FIGURE29a

/b/c:

FIGURE30 a

/b:

ROTOR

4.79

33TREPAN K

14 INNER HALF, TEST RESULTS IN FIGURE54AND TABLE6

FRACTURE I

FRACTURE II

400

FIGURE 31a: ROTOR 2.7901

FIGURE32: ROTOR 3.8153

INCLUSIONS IN CENTER OFBOTTOM JOURNAL CAUSINGUT-INDICATIONS

x 5

x 5

FIGURE31b: ROTOR 3.8153

CENTER POROSITY UNDER BODY CAUSING UT-INDICATIONS

C Si.25 .24 .010 .010

Cr Ni Mo V

1.4 2.1 .4 .12 %

200 p .

x 100FIGURE33: ROTOR 2.7901

MnS TYPE 2 CAUSING LOWDUCTILITY TENSILE FRACTUREIN RADIAL SPECIMENS

FIGURE34

4 mm

50

100

200

300

400

DIS

TA

NC

E F

RO

M S

UR

FA

CE

IN

MM

RADIAL TREPAN - FRACTURE IV IN SEGREGATION

RO

TO

R : 7

805

0108

5 m

m

38.5

t

ING

OT

10

0t

C

Si

Mn

Cr

Ni

Mo

.30

.22

.44

1.3

1.9

.34

a x 5a/b: FRACTURE F II

C .26 00 S •012i CrMIC-20FRACTOGRAPHY

.2 Ni

a °m

e:Mo d:HVO15

Mo

o CDCO

...! .1 cr, c.,• a a a c:xlFIGERF35 a/b/c/d/e: ROTOR 7718, 0 973 mm, 15 t, INGOT 85 t

TREPAN F F K K

FRACTURE I II II III

YIELD GO 2 65 68 69 69 kP/ HULT.TENS.STR. 81 82 84 85 Bp/ iruc-GBELONG. 12 oP6

5RED.AREA T 29 17 16 48 ;6c/d/e: SULPHUR PRINT, HARDNESS AND Me DISTRIBUTION

BETWEEN FRACTURE K II - III

x 260

FRACTURE

II

III

c/d/e: 25

aFRACTURE II

SULPHUR PRINT

x 5

YIELD c

58 kp/mm2

57 kp/mm2

58 kp/mm2

ROTOR 4318

0 1155 mm

45 t

INGOT 96 t

x 20 e

FIGURE 36a/b/c/d/e: ROTOR 4318, FABRICATION1955ULT.TENS.STR.

67 kp/mm2

69 kp/mm279 kp/mm2

MICROFRACTOGRAPHYC Si S Cr NI

.33 .27 .008 1.1 2.5

B

46

ELONG. 65

SY...

RED.AREA T

mm FROM FRACTURE II, Mo = .92 HV I - 290 VERSUS 270

x 2CG

No.40

x 200

a

FRACTURE III

FIGURE37 avb/ c

FRACTURE YIELD G0,2

II 26 kamm2

III 28 kp/mm2

x 5

ROTOR 78)8

0 1560 mm

t

INGOT )00 t

c: MnS TYPE 2

ULT.TENS. STR. (Y8

48 kp/mm2

49 kp/mm2

x 400

INCLUSIONS IN DUCT. MATRIX

x 100

1) V.

C .25

SI .)0

.009

.017cr .36

.34

Ro .0)

.o6V

ELONG, 8_

18 % 27 %

RED.AREA T

12 %

Mo

.5

.4

.3

.2

10 20

I.

N.,. RN

Ms HV5 ink p isnas

30 i 40mm

270

250

FIGURE 38ROTOR 5.8154SULPHUR PRINT, HARDNESS AND MO DISTRIBUTIONRfa: X-RAY FLUORESCENT ANALYSISMs : MICRO ANALYSERHV : VICKERS-HARDNESS

a x 10

x 100

FIGURE39 a/b/c/a: ROTOR 5.8154 TREPAN F 4 INNER HALF

b: IN SEGREGATION d: IN MATRIX

c: TRANSITION TO MATRIX

x 200

a: Mo-SEGREGATION BETWEEN FRACTURE I - IIVICKERS HARDNESS NV 1 - 266 VERSUS 290

b/c/d: MICROFRACTOGRAPNY OF FRACTUREIN Mo-SEGREGATION

x 2000

F1— —

400 1 2007 .25 .23 .21%C

0 200 400 600 800 1000 1200mm

DISTANCE FROM BODY SURFACE

LADLE ANALYSIS IN 1.,

C Si Mn P S Cr Mo Ni V

.24 .07 26 .008 .012 1.54 39 3.52 .11

FIGURE 40 C-SEGREGATION IN A ROTOR BY "AP PROCESS"

50

40

30

20

41)

CRACK LENGTH ( mm)

-till- 0.5mm

3

6.6_4

4

• STANDARD SPECIMEN P2, ASTM-E208

NDTT = -25°C

.1-200 -180-140 -120 -80 -40 ±0

TEMPERATURE (°C)

FIGURE 41: INFLUENCE OF ARTIFICIAL FLAWS ON NDTT.

40

30

±0

19mm --41.

_

CRACK LENGTH Onmp

50 (40) 2.IMPACT

19 min

X-RAY INDICATION FRACTURE SECTION

(POROSITY ) SURFACE AXM 3/1

20 1IMPACT

o STANDARD SPECIMENS P2, ASTM-E 208

NDTT = 0°C

P2-SPECIMEN

WITH POROSITY

+10 +20TEMPERATURE (°C)

FIGURE 42: I NFLUTENCE Q- pQQc3jrpy ON RESULT OF DROP-WEIGHT TEST.

a.)

b.)

50 m

m

FIGURE

43:

ARTIFICIALLY FLAWED FATIGUE CRACK-GROWTH SPECIMENS.

SU

LPH

UR

PR

INT

,D

IRE

CT

ION

„A

"

AX

IAL

FIGURE

SEGREGATION AND FATIGUE ORACK-GRNNTE

FA

TIG

UE

CR

AC

K

CU

T

kr-s

Om

m -0

1

30 50 100 200 30 50 100 200AK Dcp/mm3/2]

FIGURE 215: INHOMOGENEITY AND FATIGUE CRACK-GROWTH RATE.

Testa

2,1

2,0

1,9

1,8

1,7

FIGURE46: TEST RESULTS OF ROTOR 7642

NIJI 22 NiMoV 14 5

26 NiCrMoV 14 5 n=27 — 53=1=

26NiCrMoV 85 riz29

..... 7642 n=10

FIGURE47: MAGNETIZABILITY OF ROTOR STEEL GRADES

BSC = Bethlehem Steel Corporation 18)JSW = Japan Steel Works

o n=3

o flS+

1,60 10000 20000 30000 Amps / rn

ILI 11.1 S

0z

9T0

810 6T

O lo 0

161

0, Z

C..-- r--1

H0.1

re \ 0

H

LC 1

Pr \

C \I

gal

al 0

0 0

rc \ tf

Hge

0 SZ

<rr\

rx4

09Zg.-1

0c0

[Li ca

0>

OL

ZZ

UJW

/Cbi

St

A H

0 E

-f>

.< Z

\•—•-• E

-1E

-10

>-1

0 0

Lflo

CV

H

P

q 0

1;4 Z

grl=

4->

ZZ

ZH

Z

CC

EHrat, )(1=

H

<>

4 a)'.E

HE

xgt

E

H

oo o

eN

-P-

EH

Ho

Z2---4 rE

T3i

cb

CC

cd '0

00

FIGURE 49 : RESULTS OF TANGENTIAL SURFACE TESTS

13660

14860

7600

x PIN POINTED ACCORDING TO MAX. UT-INDICATIONS

x x x

AREA OF UT- INDICATIONS

6390 ri

FIGURE 50 : TREPANNING ACCORDING UT-RESULTS

x) RADIAL

TOP

TOPK18

951800-

KI4 F22

F2111-°-

BOTTOMh_

540

4-

ROTOR: 2.7901 3.8153 4.7933 5.8154 6.8189

71

81

15

64

495

BOTTOM540

_T

F22

51

71

21

65

300

PRE 52 : PELTDTP IA-DOH RADIAL TREFANS OUTER EADIES

1ST FRACTURE MINIMUM VALUES

FIGURE51 : RESULTS FROM RADIAL TREPANS OUTER HALtiTES

la- 180044

15T FRACTURE t;' MINIMUM VALUES

TOP

-0 1800-

7-1.:iuf:1:: 54 : .:(tit, • tt t•t'.-,t tut tt

BOTTOM

460

YIELD STR. (0.2)

ULT.TENS.STR.

E LONG.

RED.AREA

6.8189

.22 .23

72 71

81 80

15 18

56 52

DISTANCE FROM 570 840

AXIS1ST FRACTURE MINIMUM VALUES

FI u Litt: 53 : _F-Ctit; hA Dl A1 T.E PAIL:t i t‘INER HALVES

TOP BOTTOM:11(14 1F22F2 :t

--t-EsLr---_:.4 .0 _t— 5404

RI8 I, KI9 I1F3I

id-1800 -laX

F2 F3 F22I I I

.23 .20 .19

42 40 40

62 60 58

22 22

64 67 37

50 120 160

125 HO 90

X PIN POINTED ACCORDING TO MAX. UT-INDICATIONS 1ST FRACTURE IL-I MINIMUM VALUES

GRADE: 26 NiCrMoV 85, TABLE .6.

TOP BOTTOMAXKS AXK AXM AXE AXF

(6 1800

ROTOR AXKS AXK AXM AXF AXFS

5.8154 C V. .25 .25 .23 .23 .24

YIELD STR.(0.2) KP/MM2 AX. 64 65 63 64 64 64 62 63 63 63

II it R4D. 64 64 63 64 64 67 62 63 63 63

ELONG. "/,, AX, 22 23 21 21 20 20 23 23 22 23

n II RAD. 13 19 8 13 9 10 19 22 19 20

NDTT °C AX. -40 -25 -30 -45 -55

II II RAD ± 0 -1 5 tO tO -40 -50

6.8189 C 0/0 .24 24 .25 .23 .21 .22 .23

YIELD STR.(0.2) KP/MM AX. 74 74 73 72 72 71 72 72 72 72

RAD. 75 74 74 73 71 72 71 71 72 72

ELONG. AX. 17 19 18 21 18 18 18 20 19 19

RAD. 13 14 14 18 13 15 9 18 18 19

NDTT °C AX. -55 -65-65 -60 -55-45 -85

RAD. -40 -40-40 -50 -35-40 -80

FIGURE 56 : TEST RESULTS (MIN. MAX.) FROM AXIAL TREPANS OF 2ROTORS MADE FROM INGOTS BY "AFTER POURING" PROCESSGRADE: 26 NiCrMoV 14 5, TABLE .E

Contents

1. Introduction 3

2. Properties of large forgings

2-1 General 3

2-2 Properties of gigantic 20i-Cr-Mo-V generator rotor forgings 6

2-3 Properties of gigantic 3.50i-Mo-V generator rotor forgings 10

2-4 Properties of gigantic 3.50i-Cr-Mo-V generator rotor forgings 13

3. Developments of internal properties of gigantic forgings

3-1 Internal homogeneity of forgings 16

3-2 Developments of manufacturing process for gigantic forgings 21

3-2-1 Ingot production 21

3-2-2 Forging process 25

3-2-3 Heat treatment 26

4. Future of gigantic forgings - Conclusion 29

References

Appendix I List of forgings produced from 400 ton ingots

Appendix II Conversion tables

FORGINGS FROM GIGANTIC INGOT WITH 3,550 MM (140 INCH) DIAMETER AND 400 METRICTON (881,000Las) WEIGHT

PART 1 - PRODUCTION AND METALLURGICAL DEVELOPMENTS

Dr. Eng. Saburo Kawaguchi, Ryosuke Homma, Kazumi Takahashi and Tateo Jin

The Japan Steel Works, Ltd.Chatsu 4, Muroran, Hokkaido, Japan

Abstract

Since 1969, Japan Steel Works has utilized vacuum degassed 400 metric ton(881,000 lbs.) ingots for the successful production of gigantic generatorrotors, heavy plate mill back-up rolls and forgings for a 30,000 ton forgingpress.

Four types of steel, 20i-Cr-Mo-V, 3.5%Ni-Cr-Mo-V and3.5%Ni-Mo-V, have been utilized for gigantic generator rotor forgings. Con-sideration of the basic problems associated with 400 ton ingot production wererequired along with careful evaluation of the metallurgical characteristics ofthese four types of steel.

Four massive 270Ni-Cr-Mo-V steel generator rotor forgings were manufac-tured. In order to obtain the specified internal yield strength and goodfracture toughness, special attention was given to the control of carboncontent and heat treatment cycles.

The 3.50i-Mo-V steel is widely used for generator rotor forgings in theU.S.A., United Kingdom, France and also in Japan. At the previous Forge-masters Meeting held in Terni, an excellent paper was presented by Messrs.Smith and Hartman on the massive generator rotor forgings of this steel.

Two massive 3.50i-Cr-Mo-V steel generator rotor forgings were manufac-tured. This steel, which has been widely used for low pressure turbine rotorforgings in the world, has advantages, such as excellent toughness andstrength. However, serious considerations on metallurgical problems of thismaterial, such as temper embrittleness, coarse austenitic grain, and so on,in the massive forgings produced from 400 ton ingot were required.

Internal homogeneity, such as porosity, solidification pattern, alloysegregation, and mechanical properties is being improved by developments of

manufacturing techniques. The production facilities and some technical devel-opments ere described.

AP process of tapping of 400 ton ingot minimizes alloy segregation andreduces porosities in the ingot. JTS forging process is effective 'to closeingot porosities. The internal properties of gigantic forging are improvedsuccessfully by developed preliminary heat treatment after forging and qualityheat treatment including pressurized water spray quenching.

1. Introduction

During the Sth International Forgemasters Meeting in 1970, several paperswere presented regarding production of large forgings from ingots in excess of300 tons 1) 2) 3). Messrs. DeForest and Newhouse reported on the anticipatedneeds for still larger four pole generator rotor forgings; much larger in dia-meter and weight without any sacrifice in quality and reliability 4).

Since 1969, Japan Steel Works has utilized vacuum degassed 400 metric toningots for the successful production of massive four pole generator rotors,heavy plate mill back-up rolls and forgings for a 30,000 ton forging press.The majority have been generator rotors.

The ingot is poured sequentially from five furnaces and then tapped offfrom a sixth, Figure 1. The main dimensions of the ingot are as follows:

Mean DiameterBody LengthTotal WeightH/D Ratio

3,550 mm (140 in.)4,180 mm (164 in.)Approximately 400 ton (881,000 lbs.)1.2

Forging is accomplished by means of a 10,000 ton hydraulic press as shownin Figure 2, Figure 3 shows the preliminary heat treatment furnace (verticaltype) and the resultant forging. After preliminary machining and inspection,the forging is water spray quenched and tempered in the new heat treating faci-lities, Figure 4. After final machining, as shown in Figure 5, the forging isshipped to the purchaser.

In manufacturing such massive forgings, we have been confronted withseveral metallurgical problems; two of which will be discussed in Part 1 ofthis paper:

"Properties of large four-pole generator rotors made of variousalloy steel compositions and their metallurgical evaluations."

"Alloy segregation and internal homegeneity of massive forgingsand related development of manufacturing processes."

In Part 2 of this paper, Dr. Schinn and Dr. Schieferstein will discussthe development of 1,500 and 1,800 RPM turbine generator units with emphasis ongenerator rotor forgings.

2. Pro erties of lar e for in s

2-1 General

Since the first production of 400 ton ingots in 1969, four types of steel

have been utilized for massive generator rotor forgings. In this chapter the

properties of each steel will be discussed. Table 1 shows the typical analyses

General features of these steels are as follows:

20i-Cr-Mo-V steel •... This steel is commonly used in Europe and many

generator rotors with diameters up to about 1,000 mm have been produced with

good strength and ductility 5) 6).

2.80i-Cr-Mo-V steel .... This steel is also used in Europe and is being

utilized for the higher strength and ductility improvements over that of

20i-Cr-Mo-V steel.

3.50i-Cr-Mo-V steel .... This steel has been used widely for low pres-

sure turbine rotors since the early part of the 1960's because of its prominent

combination of high strength and good ductility at low temperatures. At the

present time, most low pressure turbine rotors and some generator rotors are

made of this steel in the U.S.A. and Europe, and also in Japan 6) 7) 8).

3.5%Ni-Mo-V steel .... This steel is widely used in the U.S.A. and Japan,

and has high strength and good ductility 7) 9).

The C-C-T diagrams for these four steels are compared in Figure 6. In

general, the microstructure at the center section. of large rotor forgings is

completely bainitic, however, it sometimes includes a small amount of ferrite.

The presence of ferrite is not desirable, because of its detrimental effect on

mechanical properties. The 3.50i-Cr-Mo-V and 2.80i-Cr-Mo-V steels have the

desirable transformation characteristics for large rotor forging material,

because, as indicated by the C-C-T diagram, these steels are considered to be

free from ferrite with commonly used cooling rates from the austenitizing

temperature. On the other hand, the 2/0Ni-Cr-Mo-V steel is considered to be a

material which is sensitive to the influence of mass and cross sectional area

upon cooling rates.

In order to obtain cooling rate data, temperature distributions duringwater spray quenching of a test rotor forging with an 1,800 mm diameter weremeasured by thermocouples inserted at various depths below the surface. Thetest results are shown in Figure 7. In this figure, cooling curves for 1,020and 1,370 mm diameter rotor forgings are also shown. With water spray quench-ing, the most effective cooling method, the cooling rate at the center portionof an 1,800 mm diameter rotor forging is only about 63 C/hr. (800 - 300 C).On the other hand the cooling rate at the center portion of a 1,020 mm dia-meter forging, typical of a conventional fossile fired generator rotor, isabout 167 C/hr. (800 - 300 C). This cooling rate corresponds to that whichoccurs at a location 200 mm below the surface of an 1,800 mm diameter rotorforging. The cooling rate at the center portion of an 1,800 mm diameter waterspray quenched forging corresponds nearly to the cooling rate at the centerportion of a 1,020 mm diameter air cooled rotor forging.

From the above data it is quite apparent that the effect of mass on heattransfer of a gigantic rotor forging is considerably greater than that of thesmaller conventional generator rotor forgings. The transformation behaviorsof the massive generator rotor forgings utilizing the steels shown in Table 1can be approximated from the C-C-T diagrams shown in Figure 6 and the coolingcurves shown in Figure 7. However, in actual large rotor forgings, the in-fluence of micro-alloy segregation on the transformation behavior must also betaken into consideration. Consequently a series of experiments were performedto simulate the transformation behavior at the center portion of giganticrotor forgings in order to predict the associated micro-structure at thiscritical location. These experiments were carried out using specimens fromactual large rotor forgings.

The results of the above experiments, as related to the cooling curve for1,800 mm diameter rotor forgings, are summarized in Figure 8. The 20i-Cr-Mo-Vand 3.50i-Mo-V steels show two stages of transformation, whereas the othertwo steels show only one transformation stage.

From the above test results it is apparent that the following generalfeatures of these four steels must be considered when they are used formassive

generator rotor forgings:

2/0Ni-Cr-Mo-V steel The presence of ferrite and pearlite is inevitableat the center portion of the forging and consequently this steel is sensitiveto alloy segregation.

2.130i-Cr-Mo-V steel .... The hardenability is better than that of2/0Ni-Cr-Mo-V steel.

3.50i-Cr-Mo-V steel .... The hardenability is the best, and the trans-formation temperature is the lowest of these four steels.

3.50i-Mo-V steel .... The hardenability is good, however ferrite may bepresent within the center portion.

When a massive rotor forging is quenched from the austenitizing tempera-ture, the forging should be cooled down to a low temperature as possible toprevent cracking and then reheated for tempering. In this case, the transfor-mation must be complete prior to tempering. Table 2 shows the influence ofnickel content on the difference between Bf and Ms temperatures. In order toaccomplish complete transformation, center of a forging must be cooled to andheld at a temperature between Bf and Ms. Consequently, as the differencebetween the Bf and Ms temperatures becomes greater the selection of holdingtemperature becomes less critical. From this point of view, it is apparentfrom Table 2 that the required holding temperature for 3.5%Ni-Cr-Mo-V steel isconfined within narrow limits. In the case of 2.0%Ni-Cr-Mo-V and 3.5%Ni-Mo-Vsteels, the selection of the holding temperature is quite easy due to thelarge difference between the Bf and Ms temperatures.

Table 2 - Influence of nickel content on Bf and Ms temperature:

SteelBf(C)

2.0%0Ni-Cr-Mo-V 409 361

2.8%Ni-Cr-Mo-V 348 340

3.5%Ni-Cr-Mo-V 336 340

3.5%Ni-Mo-V 464 376

Ms Bf - Ms(C) (C)

Note: a. Chemical compositions: As shown in Table 1

b. Calculation of Bf and Ms: By Steven and Haynes 10)

2-2 Pro erties of i antic 20Ni-Cr-Mo-V enerator rotor for in s

- 6

48

8

- 4

88

Four gigantic generator rotor forgings were manufactured of this steel.All were made of the heats produced in the early phases of 400 ton ingot pro-duction. Therefore, consideration of the basic problems associated with 400ton ingot production were required along with careful evaluation of the metal-lurgical characteristics of the 20i-Cr-Mo-V steel.

In 1960's common practice in the production of conventional 2%Ni-Cr-Mo-Vrotor forgings was to aim for a carbon content of 0.3% or higher. In 1969 thecarbon content was specified as 0.3% or lower in order to improve the tough-ness of rotor forgings. The aim of carbon content of rotor forgings made fromgigantic ingots is important. Experience shows that the tendency for crackingduring heat treatment increases as carbon content exceed 0.4% in this steel.On the other hand, carbon content in the negative segregation zone of thebottom of the rotor forging has to be sufficiently high to assure meeting thespecified yield strength. It is noted that the internal yield strenath of

2%Ni-Cr-Mo-V steel rotor forgings is greatly influenced by carbon segregation.In general, the carbon content is the highest at the top end of the ingotwhich leads to the highest yield strength, and the highest NDTT as well. Con-sequently, in order to lower the NDTT the carbon content at the top end mustbe maintained at a practical low level.

From the above considerations combined with other fundamental investiga-tions the carbon content for massive generator rotor forgings was selected as0.2% minimum to guarantee meeting the specified yield strength. It was alsodetermined that the carbon content should not be allowed to exceed 0.4%. There-fore, the final carbon content of ladle was selected as 0.26% in accordance withthe specified carbon content, 0.24 - 0.30%, by Kraftwerk Union AG.

As previously stated, the inevitable presence of ferrite and pearlite inthe center portion of massive 2%Ni-Cr-Mo-V rotor forgings will have a detri-mental effect on NDTT. In order to improve the NDTT, such actions as refiningof austenitic grain size and maintaining the yield strength as low as permittedwere attempted. Experience shows that the austenitic grain size of large rotorforgings is influenced by the thermal history prior to final heat treatment,such as the annealing process used after forging. Therefore, proper actionswere taken before bringing the forgings to the stage of final heat treatment.The austenitizing temperature for the final heat treatment was selected at theallowable lower limit to prevent unnecessary grain growth.

The physical dimensions, chemistry and heat treatment of four mammothrotor forgings made from 2%Ni-Cr-Mo-V steel are shown in Table 3. The materialtest results are shown in Table 4. The test results satisfy the specifiedrequirements. The differences in yield strength between the body surface andthe center core, considered to be associated with the presence of ferrite andpearlite in the center portion, is noticeable. The yield strength o thebottom end center core is just above the specified value of 40 kg/mm . TheNDTT of the top end center core is lower than 20 C.

Figure 9 shows the properties along the axial center core of rotor CAB andGAC from Table 3. The extent of carbon segregation of rotor GAB is ratherlarge, and the difference between maximum and minimum carbon content is 0.11%.Consequently, the yield strength and NDTT are higher at the top end of theforging. Comparison of rotor GAB and CAC shows clearly that the mechanicalproperties of this steel are markedly influenced by carbon segregation. Themicrostructure of their center cores consist of bainite, ferrite and pearliteunder the significant influence of micro alloy segregation as expected by theresults of previously mentioned simulation cooling tests.

The distribution of hardness and microstructure from the rotor surface tocenter core is shown in Figure 10. The hardness decreases abruptly at around400 mm from the surface because of the appearance of ferrite and pearlite,

Table3

Four gigantic 2:-?0Ni-Cr-ko-V

generator rotor forgings mace from400ton ingots

Body

Total

Chemical composition ()

4uality heat

Rotor

dia.

length n'eight

treatment

*(aim)

(mm) (ton)

Si

Ni

Cr

o LV

900 0.46hr.- Fan

GA

».

1 008

15,600

172

0.26

0.21 0.54 0.0100.0112.08 1.47

0.400.11

640 0.48hr.- Spray

660 C.56hr.- FC

900 0.34hr.- Fan

GAB

1,808 14,800

196

0.25

0.24 0.34 0.010 0.0102.10 1.44 0.410.12

840 C.51hr.- Spray

660 C.56hr.- FC

900C.41hr.- Fan

GAC

1,808

14,600196

0.25

0.20 0.35 0.011 0.0101.95 1.47 0.440.12

840 0.51hr.- Spray

650 0.60hr.- FC

900 C.46hr.-Fan

GAD

1,808 13,600

172

0.25

0.21 0.36 0.009 0.0101.98 1.42 0.420.11

840 0.53hr.- Spray

650 C.64hr.- FC

LO

Location

Rotor

0.2%Y.S

(kg/mm2)

vE25 C

(kg -m)

NDTT

(C)

Table 4

Mechanical

properties

of 20i-Cr-Mo-V generator

rotor forginge

ST

GA

A

B

C

Ta

-

59

Lb - 16 -

-

Body surface SB

GA

DABC

D

A

58

•

59 59 - 51

42158

50

441

96;

180

- 15

CT

Body center

CT

CM

CB

GA

GA

GA

BCDABCDABCD

25

444

- 40

42

42

43

20

5-

- 5

510 15 20

-

-

5

0

a: Tangential

specimen,

b: Longitudinal

specimen,

c: At -20C

ST

SB

Top

Bottom

CM

CB

- 41

43

42

44

41

3 4

4 6

-

4 5

- -

7 7

9

O -

10

-

O -5

-

Before beginning to manufacture gigantic 20i-Cr-Mo-V rotor forgings, manystudies were conducted to determine proper heat treatment cycles consideringthe large mass effect on heat transfer which was predicted by the C-C-T diagramand simulation cooling tests. Actual qualities obtained were almost as expect-ed. The carbon segregation in these rotor forgings was slightly greater thanexpected in this early stage of 400 ton ingot production. Since then, improve-ments on reducing carbon segregation in 400 ton ingots have been made as shownby the following data.

2-3 Pro erties of i antic 3.50Ni-Mo-V enerator rotor for in s

The 3.50i-Mo-V steel is widely used for generator rotor forgings in theU.S.A. and also in Japan. At the previous Forgemasters meeting held in Terni,an excellent paper was presented by Smith and Hartman on the generator rotorforgings of this steel made from 318 ton ingots 1).

Thus far, Japan Steel Works has made six gigantic generator rotor forgingsusing this steel. Pertinent data, representative of these forgings, are shownin Table 5, Table 6 shows the mechanical test results of these forgings.These test results satisfy the specified requirements completely.

Figure 11 shows the relationship between tensile strength and FATT ofcenter cores. The data reported by Smith and Hartman are also included. Nosignificant difference is observed when comparing the data of gigantic rotorforgings with those of the smaller conventional generator rotor forgings.

Table 7 shows temper embrittlement sensitivity estimated by step-coolingheat treatment of rotor GBB and GBF shown in Table 5. The results confirmcommon knowledge that this steel is not sensitive to temper embrittlement.

Figure 12 shows the axial distribution of alloy segregation, mechanicalproperties and microstructures in the center core of rotor CBE. In this case,lower alloy segregation was obtained than that observed for 2/0Ni-Cr-Mo-V steel.It is believed that this was accomplished by the corrective actions taken,based upon early experiences in 2/0Ni-Cr-Mo-V ingot production. Variations ofyield strength between the top and bottom end is not so large; however that ofFATT is noticeable. This variation of FATT cannot be explained by alloy segre-gation alone. However, from the relationship of tensile strength and FATTshown in Figure 11, the FATT of this gigantic rotor forging can be consideredto be almost the same as that of conventional generator rotor forgings.Ferrite, which appears by influence of micro alloy segregation, was observedas predicted from the simulation tests for the center portion of 1,800 mmdiameter forgings.

- 10 -

Table5

Six gigantic

3.5%Ni-MO-V

generator

rotor forgings

made from400ton ingots

BOY

Total

Chemical

composition

(%)

Quality heat

Rotor

dia.

lengthWeight

treatment

(mm)

(mm)

(ton) ' C

Si

Mn

PS

Hi

Mo

V

900C.39hr.-

Fan

GBA

1,706

14,200

158

0.21

0.24

0.400.0080.0104.01

0.46

0.12

820 C.41hr.-

SPIAY

640 C.57hr.-

FC

900C.32hr.-

Fan

GBB

1,632

16,480

155

0.17

0.22

0.270.0110.0133.62

0.27

0.13

810 C.39hr.-

Spray

620 C.54hr.- FC

920 C.40hr.-

Fan

GBC

1,714

14,926

175

0.21

0.06

0.280.0080.0083.53

0.20

0.11

820 C.51hr.-

Spray

630 C.60hr.-

FC

900 C.38hr.- Fan

GBD

1,706

14,550

158

0.20

0.23

0.410.0090.0083.50

0.36

0.09

820 C.47hr.-

Spray

635 C.58hr.-

FC

900 C.43hr.-

Fan

GBE

1,651

12,655

127

0.19

0.21

0.430.0070.0083.54

0.31

0.13

820 C.31hr.-

Spray

630 C.58hr.-

FC

900C.35hr.- Fan

GBF

1,685

15,310 ! 176

0.22

0.23

0.330.0090.0083.55

0.36

0.12

820 C 42hr.- Spray

635 C.58hr.-

FC

1

ST

CT

SM

SB

CM

CB

a: Radial specimen, b:

Tangential

specimen, c:Longitudinal

specimen,

d: 0.04 off set

Top

Bottom

Table 6

Mechanical

properties

of six gigantic

Body surface

Location —

ST

SM

SB

GB

GB

GB

3.50i-Mo-V generator

CTGB

rotor forginge

Bod

y center

CM

GB

CB

GB

Rotor

AB

CD

EF

AB

CD

EF

AB

CD

E

FA

BC

DE

FA

BC

DE

FA

BC

DE

Fd

d

d

d

d0.2%Y.S Ra59 69 561 60 66 - 58 - 56 61 64

5959 63 57 6064

(kg/mm2)

Tb Lc -66 ;

64 - - -

1 1- 53 63 60 63

51 5363 63 62- - 50 59 57 60

- 54 49 58 59 61- - 48 59 59 59

- 50 47 60 61 59

vE25 C

R- - 1421 30 - 7

- - 32 12 - - - - 33

(kg-m)

T15 13

5 6 9 4 4

- 4 9

85

46 9 10 14 9

7 9 7 10 7

- 8 17 119 8

- 818 14 18 13

FA

TT

--15 -5-43 -25

-9

(C)

17

10.

-13

- 26 17 23 43 34

-29 7 25 22 32- 24 -1 18 10 21

28 8 29 28 25- 28 0 23 25 22

- 22 -8 8

3 10

NDTT

-25-40

-30

-40-50 --

_

Table 7 - Temper embrittlement sensitivity of center core of 3.5%Ni-Mo-Vgenerator rotor forgings made from 400 ton ingots

Rotor

GBB

GBD

If

Location Direction TTWQ TTSC AFATT(C) (C) (C)

Middle Long. 17 12 -5

Trans. 23 34 11

Top Long. 5 11 6

Trans. 19 28 9

Bottom Long. 8 10 2

Trans. 13 22 9

Note: TTWQ FATT after water quenched from 595 CTTSC FATT after step cooled from 595 C 11)

AFATT = TTSC minus TTWQ

2-4 Pro erties of i antic 3.50Ni-Cr-Mo-V enerator rotor for in s

Thus far, two 3.5%Ni-Cr-Mo-V massive rotor forgings have been produced.This material has been widely used for low pressure steam turbine rotor forg-ings for fossil fuel power plants. This steel has advantages, such as excel-lent toughness and good hardenability, which are important for large rotorforging applications. However, there are also some disadvantages, such assensitivity to temper embrittlement and the tendency to form coarse austeniticgrains. Therefore, serious considerations were required prior to the applica-tion of this steel to gigantic rotor forgings.

As mentioned previously, the temperature difference between Bf and Ms inthe T-T-T diagram is very small in 3.50i-Gr-Mo-V steel. With this and theinevitable segregation in large rotor forgings, the completion of transforma-tion during quenching was considered to be difficult. So, it was important toinvestigate the transformation behavior of retained austenite in this steel.

Some results of investigations on decomposition of retained austenite in3.50i-Cr-Mo-V steel are shown in Figure 13 and Figure 14.

Figure 13 shows the isothermal diagram of the decomposition of retainedaustenite. In this case, samples were quenched to room temperature from 1,000C and then followed by reheating to isothermal transformation temperatures.It is apparent that holding at around 300 C is effective for decomposition ofretained austenite.

- 13 -

Retained austenite may also be decomposed by a conditioning treatment.In conditioning, the quenched material is heated to a temperature below Ac1and cooled after holding at the temperature. Decomposition of the retainedaustenite occurs during cooling from the conditioning temperature. Figure 14shows the temperature-time relationship required for decomposition of auste-nite by conditioning. It is shown that the holding time required, becomeslonger with lower holding temperature. At temperatures below about 520 C, nodecomposition is expected.

Figure 15 shows an example of the heat treatment cycle for large lowpressure steam turbine rotor forgings of 3.5%Ni-Cr-Mo-V steel, in which decomp-osition of retained austenite is tried by techniques such as holding at 300 Cbefore reheating.

The fact that 3,510Ni-Cr-Mo-V steel has a tendency to form coarse austenitegrains cannot be ignored, especially with regard to attenuation of sonic energyin ultrasonic inspection. Considerable investigations have been performed onthe behavior of austenitic grains in this steel. The following are someresults of these investigations.

It has been determined for this steel that the heating rate in austenitiz-ing markedly changes grain size. Figure 16 shows the influence of heating rateto the austenitizing temperatures upon grain size. It is believed that slowheating, which is inevitable in large forgings, is one of the most importantfactors for the formation of coarse grains in this steel.

Figure 17 shows the influence of prior austenitic grain size on grain sizeafter reaustenitizing. It has been known by past experiences that the austen-itic grains in large forgings of this steel become finer by repetition ofaustenitizing. Investigation results shown in Figure 17 suggest that therefining effect of an austenitizing cycle is inherited accumulatively to thenext austenitizing cycle.

It has been determined that the change in grain size shown by Figure 16and Figure 17 is related to the initial grain size formed in the early stage ofaustenitizing, but it is not related to the so-called grain coarsening tempera-ture.

Figure 18 shows the influence of nickel content on grain size. Obviously,the tendency to form coarse grains in 3.50i-Cr-Mo-V steel is related to itshigh nickel content. This suggests that, in the case of 210Ni-Cr-Mo-V steel,grain refinement can be easily achieved. In addition, from another investiga-tion on the influence of carbon, chromium, molybdenum and vanadium on grainsize, it was found that the combination of these elements has a close relation-ship with the tendency to form coarse grain size in 3.5%Ni-Cr-Mo-V steel.Further investigation suggests that these phenomena must be explained essenti-ally from the standpoint of crystallography.

- 14 -

Sensitivity to temper embrittlement is one of the most critical aspects of3.5%Ni-Cr-Mo-V steel. Recently, an excellent report about the influence ofresidual elements on temper embrittlement in this steel has been published bythe ASTM Special Task Force on Large Turbine and Generator 12). Our resultscoincide with the data described in this report. In Figures 19 and 20, theinfluence of residual elements on temper embrittlement in 3.50i-Cr-Mo-V steelis shown. Results in Figure 19 show temper embrittlement induced by slow cool-ing from the tempering temperature. The A FATT in Figure 20 was determined bythe step cooling 11). From these investigations, it is clear that as far asthe practical range of their content is concerned, phosphorus and tin havesignificant effect but antimony and arsenic have little influence on tamperembrittlement.

The AFATT determined on step cooled center core samples of large3.5%Ni-Cr-Mo-V rotor forging also has good correlation with the combined amountof phosphorus and tin as shown in Figure 21.

It has been known that this steel when vacuum-carbon deoxidized tends toshow less sensitivity to temper embrittlement than when silicon deoxidized.Table 8 shows some examples of temper embrittlement in center core tests of3,5%Ni-Cr-Mo-V turbine rotor forgings. It clearly demonstrates that vacuum-carbon deoxidized steel has less sensitivity to temper embrittlement.

Table 8-Influence of deoxidation methods on temper embrittlementof center core of 3.50i-Cr-Mo-V turbine rotor forgings.

Note: TTWQ FATT after water quenched from 595 CTTSC FATT after step cooled from 595 C 11)

A FATT TTSC minus TTWQ

The facts mentioned above show the very special nature of 3.50i-Cr-Mo-V,steel and it is assumed that they may be explained by some basic characteristicwhich belongs to this steel. In the meantime, manufacturing of massive generatorforgings from 3.50i-Cr-10-V steel 400 ton ingot is to be extremely interesting

- 15 -

Deoxidation TTWQ TTSC A FATTRotor method Location Direction (C) (C) (C)

TC SiliconTop Long.Bot,

-50 119-45 64

iI

169109

Vacuum- Top -71 -34 37TD 11

carbon Bot. -80 -38 42

TE 11TopBot,

ti-66-63

-22-26

4437

Top -81 - 5 76TF Bot.

11

-62 20 82

to forgemasters because very careful considerations of various factors arenecessary.

In Table 9 the size, chemical analysis and heat treatment conditions oftwo massive 3,5%Ni-Cr-Mo-V steel generator rotor forgings are listed. Therotor forging designated as GCA is silicon deoxidized, whereas GCBis vacuum-carbon deoxidized. As discussed above, with respect to temper embrittlement,it is better to choose vacuum-carbon deoxidized steel. However, the first 400ton ingot of this steel was made by silicon deoxidization in order to avoidextra problems associated with vacuum-carbon deoxidizing treatment of such alarge ingot. The forging designated as GCB was manufactured from vacuum-carbondeoxidized steel, a step for technical development of 3.50i-Cr-Mo-V giganticgenerator forgings.

In Table 10, mechanical properties of these massive generator rotor forg-ings are shown, All mechnical test results satisfied the specification require-ments. Of the two rotor forgings, the silicon deoxidized GCA has a higher FATT.Figure 22 shows the distribution of mechanical properties and carbon content inthe center care of the two rotor forgings. Higher FATT in GCA is mainly theresult of its higher sensitivity to temper embrittlement. Yield strength,carbon content and microstructures show little variation.

So far, twelve gigantic generator rotor forgings of 20i-Cr-Moje,3.50i-Mo-V and 3.50i-Cr-Mo-V steels have been manufactured. In addition,three massive rotor forgings are now being produced with 2.80i-Cr-Mo-V steel.As in the cese of the other steels, various investigations have also beenperformed on 2.8%Ni-Cr-Mo-V steel.

All past experiences on 400 ton ingots including vacuum-carbon deoxidiza-tion have been applied in the production of 2.80i-Cr-Mo-V massive generatorrotor forgings. Careful follow-up and investigations are continuing in thecourse of manufacturing of these newly applied steel forgings.

These results of investigations shall be reported in future.

3. Develo ments of internal ro erties of i antic for in s

3-1 Internal home eneit of for in s

These forgings were ultrasonically inspected with high sensitivity at threemanufacturing stages, that is, before final heat treatment, after final heattreatment without bore and after finish machining. The test apparatus usuallyare Sperry UM type or Krautkraemer USIP type and the sensitivity is adjusted soas to have a noise level of 5 mm or less. Frequency is usually 2.25 or 2.00 MHz,By using the above ultrasonic method, we are able to detect discontinuitieshaving calculated sizes of 1.6 or 2.0 mm diameter at the center of forgings witha 1,800 mm diameter.

- 16 -

Table9

Two gigantic

3. i-Cr-Mo-V

generator

rotor forgings

made from400ton ingots

Body

Total I

Chemicalcomposion(%)

4uality heat

Rotor

dia.

length Weight

; treatment

(mm)

(mm)

(ton):C

Si

Mn

PS

Ni

Cr

Mo

V

900 C.47hr.-

Fan

Goa

1,808 13,600 172

0.230.200.400.0110.0093.481.540.390.11 840 8.468r.

-spray

635 C.64hr. -FC

900C.42hr.-Fan

GCB

1,808141920200

0.240.070.260.0080.0123.521.540.390.11840 0.458r.

-Spray

610C.63hr.-

FC

Location

0.2%Y.s

(kemm2)

FATT

(C)

NDTT

(C)Table 10 Mechwnical properties

of two giganti

i -Cr -MC -V generator

rotor forginge

Rotor TopTa Lb

1vE25 C

T

(kg-m)

Body surface

ST

SM

SB

GC

GC

GC

ABAB

- 76 -

1

Body center

CT

ccA

B

A

B

62 -

63 74

63 73

17

4 10

7 14

-85 -

57 -16

18 -30

15 -40

-25 -65

a: Tangential

specimen,

b: Longitudinal

specimen,

cs At -20C

ST

SE

SB

CT

CM

CB

Bottom

In general, we can find flaw indications in gigantic forgings by the aboveultrasonic test methods. The typical test results of three generator forgingsmade from 400 ton ingots are shown in Figure 23.

The generator forging "GAC" in Figure 23 was manufactured at an earlierstage. The majority of ultrasonic indications located in the body and top sidejournal were caused by small ingot porosities. We found 59 ultrasonic indica-tions in the body. Ninety-two percent of the calculated flaw sizes were smallerthan 3,0 mm and the maximum calculated flaw size was 3.3 mm,

The generator forging "GCA" in Figure 23 was bored to a diameter of 396 rm.The majority of internal defects were removed and only a few indications re-mained in the forging.

In the bottom end journal of the generator forging GBF in Figure 23, theultrasonic indications were caused by non-metallic inclusions.

These flows mentioned above were naturally the same as those observed inthe conventional rotor forgings. The typical porosities and non-metallicinclusions, which appeared in the forgings made from 400 ton ingot, are shown inFigures 24 and 25, respectively.

It is very important to examine the characteristics of the material aroundthese defects, such as strength, fracture toughness, solidification patterns(macrostructure and sulphur print), segregated streak and alloy segregation.The evaluations of these defects must be performed after the above examinations,

In the 5th International Forgemasters Meeting in 1970 3), we discussedsolidification behavior and segregation of large forging ingots. The periodfor complete solidification of 400 ton ingots is naturally longer than forothers, As mentioned in Section 3-2-1 of this paper, we are developing the"AP process" (After-Pouring process) in order to control the solidification ofthe 400 ton ingot.

Figure 26 shows the sulphur prints of both ends of the forging made from400 ton ingots. Because of the lower solidification rate, more segregatestreaks were found here than in conventional forgings. It is evident that the"AP process" improved upon the internal segregate streak condition.

Figure 27 shows the sulphur prints and macrostructures of cross sectionsof center core material of a forging made from a 400 ton ingot.

In general, there are some typical patterns of sulphur prints and macro-structures of center core material from large forgings made from ingots weighingover 100 tons. Figures 28 and 29 show the typical pattern of sulphur prints andmacrosegregations of center core material.

- 19 -

Non-metallic inclusions locate generally in the area shown as Type A inFigures 28 and 29. In this area, the harmful segregates are not found andfracture toughness properties are improved.

Porosities locate generally in the areas shown as Type B and C. In Type B,porosities locate at interdendritic zones. In Type C, porosities locateoccasionally at the network segregates. In Part 2, these porosities are men-tioned in relation to fracture toughness of these area.

As a result of the above considerations, the suitable bore diameter ofgigantic generator forgings, except the exciter bore, will be 300 mm, whichcorresponds to 17%of the outer body diameter of the forging. If required, thebore could be enlarged up to 400 mm or so, which corresponds to 22% of the outerdiameter of the forging. As shown in Figure 8 of the previous paper 3), thesegregated area with porosities may be removed by boring as mentioned above.

As a result of examination of defects appearing in the forgings made from400 ton ingots thus far, and anticipated future developments, it is feasiblethat gigantic forgings may be produced without a bore.

The alloy segregations in large forging ingots are generally considered tobe influenced by ingot weight or diameter and shape of ingot especially height/diameter ratio and taper. However the significant factors which influence thealloy segregations in gigantic ingots are as follows:

(1) Tapping schedule in multi-pouring(2) Control of segregation by "AP process"

Figure 30 shows the axial carbon segregation of generator forgings manu-factured from 400 ton ingots, in comparison withconventionaland developedpouring techniques (AP process).

It is clear that the carbon segregation was satisfactorily reduced by thedeveloped pouring techniques.

If the 400 ton ingot would be multi-poured at the same time and with thesame chemical analysis (for example, carbon content: 0.26%), the assumed carboncontent of the top and bottom of the ingot should be more than 0.40% and lessthan 0.20%, respectively. This carbon content at the top end may cause crackingof the forging by severe water quenching. The carbon content of the bottom endmay be cause to select a lower tampering temperature in order to obtain theminimum yield strength in the center of bottom end, and as a consequence of thetempering operation the fracture toughness in the center of top end may belessened.

In general, molybdenum has a tendency to segregate. As shown in Figure 8

- 20 -

of the previous paper 3), molybdenum content is higher in the top end of theingot and lower in the bottom end. This segregation of molybdenum affects theproperties of the center core of large forgings. Therefore, it is necessary tocontrol the molybdenum content of each heat, in the same manner as for carboncontent, in the tapping schedule of massive 400 ton ingots.

Figure 31 shows the distribution of various elements in a large generatorforging manufactured from a 400 ton ingot. This forging, manufactured from a3.5%Ni-Cr-Mo-V ingot, was poured by a developed tapping schedule followed by"AP process". The alloy segregation in the forging was evidently reduced.

As stated above, segregation of molybdenum in forgings may be reduced.But it seems more important that molybdenum has a tendency to be enriched inthe segregate streak of the ingot 13).

Dr. Scalise 14), has shown the possible degree of segregation of the mostimportant elements to be as follows:

Mn Cr Ni Mo Cu P S As

1.2-2.0 1.2-1.6 1.1-1.4 2-7 1.3-1.7 4-10 4-10 18

The degree of segregation of molybdenum in segregate streaks is assumed as1.5-3.5. Therefore, the molybdenum content of a segregate streak, when themolybdenum content of ladle analysis is 0.50%, is assumed as 1.0%. This contentis too high and causes embrittlement of the segregate streak. Phosphorus,antimony and other enriched elements in the segregate streak emphasize theembrittlement of the segregate.

The behavior of molybdenum and other elements in segregates of massiveforgings is to be studied.

3-2 Develo ments of the manufacturin rocess for i antic for in s

3-2-1 In ot roduction

The 400 ton ingot is made by a multi-pouring process. It seems rathermetallurgically beneficial that we could not apply a single-pouring processbecause of the unavailability of a furnace with a capacity of 400 tons. Thedevelopments of a multi-pouring process have possibilities to reduce consider-ably the porosities and to minimize segregations in gigantic ingots 1).

At Japan Steel Works, the 400 ton ingot body is multi-poured from fivefurnaces and then tapped off from one or two furnaces. Table 11 shows themelting facilities of Japan Steel Works.

- 21 -

Table 11 Melting and ingot production facilities

Melti facilities

I ot roduction facilities

Vacuum equipment : Steam ejector

Vacuum tank

No. 1

No. 2

No.3

No. 4

No. 5

500 ton tank

250 ton tank

140 ton tank

90ton tank

45 ton tank

Note : * BOH : Basic open hearth

BEF : Basicelectricfurnace

- 22 -

Furnace Nominal ca cities

No.13 BOB (*) 90 tonNo. 4 BEF (*) 120 ton

No. 3 BEF 100 ton

No. 2 BEF 27 ton

Nu. 1 BEF 10 ton

As mentioned above, the developed tapping schedule is required to improvethe internal qualities of the ingot. Table 12 shows an example of the tappingschedule for 400 ton ingots at Japan Steel Works. The carbon and molybdenumcontents of each heat are controlled to reduce the alloy segregation in the in-got.

Table 12 Tapping schedule of 400 ton ingot.

Preferred ta i schedule a

Tapping order

Vacuum de assi schedule

Vacuum tank : No. 1 tank

Degassing method : Stream degassing

Pony ladle : 60 ton ladle

Vacuum schedule : 0.10-0.15-0.18-0.20 mm Hg

Vacuum carbon deoxidizing : Yes

Stripping time

1 No. 15 BOH 932 No. 4 BEF 132

3 No. 3 BEF 100

4 No. 1 & 2 BEF 35AP No. 1 & 2 BEF 40

100 hrs. after pouring

The preferred carbon and molybdenum contents of each ladle should bedetermined in consideration of the locations of each heat within the ingot mold.Figure 32 shows schematically the positions of each heat within the ingot moldin accordance with the tapping schedule of 400 ton ingots as shown in Table 12.Naturally the boundaries between each heat are not finitely delineated becauseof violent agitation of molten steel caused by pouring of the additional heats.

We are not able to explain the exact mechanism by which the control ofelements of each heat in multi-pouring contributes to reduce the alloy segrega-

- 23 -

Furnace Weight

tions in gigantic ingots. The physical behaviors - of mechanical and thermalagitation, diffusion, etc. - of molten steel in the ingot mold after multi-pouring are complex and difficult to study.

Nevertheless, reduction of alloy segregation by a controlled tappingschedule has been observed.

In the previous paper 3) (page 247), we explained the after-pouring methodfor huge ingots for rolling mill rolls. Because of the high carbon low alloysteel, the tendency of severe segregation in the center of conventional multi-poured large ingots was observed. These internal properties may cause breakageof the huge roll forgings in service In order to improve the internal fracturetoughness of these forgings, the "AP process" ("After-Pouring process") has beenapplied for making huge ingots for roll forgings.

After a sufficient thickness of a solidified shell is formed, the "AP pro-cess" for roll forging ingots is to be performed. That is, the lower carbonlower alloy molten steel is poured into and mixed with the residual molten metalin the mold. The solidification which has progressed before the AP processoperation is stopped temporarily because of the rising temperature and thereduced melting point of the residual molten steel. And then new solidification,emanating from the surface of the solidified shell, proceeds. Therefore, theboundary between the solidified shell and the molten steel as mentioned above isevidently observable.

The "AP process" for massive ingots for rotor forgings should evidentlydiffer from that applied to roll forgings. The boundary zone appearing iningots for roll forgings should not be allowed in rotor forging ingots. Somedefects have been found along the boundary zone of roll forging ingots.

The most important purpose of the "AP process" for massive rotor forgingingots is to minimize alloy segregation and to reduce porosities in the ingot.It seems that the "AP process" has successfully improved the internal propertiesof massive ingots.

The solidification behavior of the 400 ton ingot was calculated by computeras shown in Figure 33. When the axial solidification is accelerated, the alloyelements - especially carbon, phosphorus, etc. - and gases in the residualmolten steel are enriched. Prior to the above solidification stage, the newmolten steel is gradually poured into the sink head. It seems undesirable topermit the molten steel to penetrate mechanically into the center of the ingot.The control of the depth of mechanical penetration into the residual moltensteel is most important.

Therefore, the intermediate arrest of solidification in the ingot should beavoided. The after-poured molten steel is mixed mechanically with only moltensteel in the sink head and the resultant melt penetrates into the residualmolten steel in the center of ingot by physical movement and diffusion.

- 24 -

In the earlier phases of manufacturing 400 ton ingots, the control of theAP process was unsatisfactory and alloy segregation and porosities wereobserved in the massive forgings. The further developed AP process as men-

tioned above, has improved the internal properties of massive forgings.

3-2-2 For in rocess

At Japan Steel Works, the massive forgings from 400 ton ingots are hot

worked by the 10,000 metric ton forging press. The main dimensions of the

forging presses of Japan Steel Works are shown in Table 130

200 ton

100 ton

50 ton

30 ton

- 25 -

Overhead crane Hooking heights Unit

17.0 m 326.5 m 112.0 m 2it 1

An example of a typical forging process for a generator forging (shippingweight: 200 metric tons) is shown in Figure 34. The ingot was transported fromthe melting shop, gascut to remove the sink head and put into the large reheat-ing furnace. The reheating temperature is usually 1,200-1,270 C. Some consid-erations on heating rates and holding time for ingots with large diameters havebeen published 15) 16). It is very important to heat sufficiently to obtainrequired temperature at the center of the ingot. At the first stage, the topand bottom ends of the ingot are hot worked in order to make prolongations forhandling. At the second stage, the ingot is square forged using wide anvils(1,300 mm) with full power of the press and then further forged using the "JTSforging" technique which was published in 1963 17). These square forging(andJTS forging operations are performed twice. It is very important that closingof ingot porosities occurs during the second stage of forging.

We have observed by destructive examination, that deformations in thecenter of ingots with large diameters are considerably less than those occur-ring at the surface during square forging, even by using wide anvils. Thismeans that closing of ingot porosities by only powerful square forging is notsufficient.

The internal temperature gradient in the forging before JTS forging seemvery essential to effect closing of ingot porosities. In conventional JTSforging, the suitable temperature gradient in the forging is achieved by stillair cooling. But in the case of the gigantic forging made from the 400 toningot, accelerated controlled cooling is required.

At the third stage, the forging is hot worked to the final forged shape.

We have experience manufacturing two gigantic forgings with 2,700 mm(110.2 in.) diameter. Because of the capacity of the forging press we are notable to upset the 400 ton ingot. Therefore, the forging ratio of the aboveforgings was only 1.7. This means that reduction of ingot porosities mustoccur during casting of the ingot.

The process for forging the 2,700 mm diameter forging is shown inFigure 35. The ultrasonic test results of the two forgings show that one ofthe forgings was good, whereas the other was found to have ultrasonic indica-tions resulting from ingot porosities.

3-2-3 Heat treatment

In Section 2-1 of this paper, the transformation properties of fourmaterials for gigantic generator forgings were discussed. It is important thatthe heat treatment of gigantic forgings should be performed carefully in consi-deration of the transformation properties of the material, the alloy segrega-tions and internal temperature distributions during heat treatment.

- 26 -

Japan Steel Works has three oil-fired horizontal type furnaces for prelim-inary heat treatment, after hot working of gigantic generator forgings asfollows:

Furnace No. Width Height Length

507 4.040 mm 4.500 mm 18 mK14 3.800 mm 3.500 mm 19 m1<16 2.600 mm 2.620 mm 18 m

The temperature in the furnace is controlled within 10 C.

- 27 -

Control

AutomaticManualManual

The preliminary heat treatment cycles after hot working are shown inFigure 36. The forgings must have a fine grained structure after the abovetreatments. It is very important that the massive forgings made from 400 toningots should be completely ultrasonically inspected. Generally, the loss ofsonic energy due to grain structure of forgings with 1,800 mm diameter shouldbe less than 15 dB.

The 20i-Cr-Mo-V and the 3.50i-Mo-V steel forgings are double austenit-ized and tempered. As mentioned in Section 2, it is easy to obtain fine grainstructure in these steel forgings.

The 3.50i-Cr-Mo-V and 2.8%Ni-Cr-Mo-V steel forgings are to be tripleaustenitized and tempered. The lower temperature treatments for bainitictransformation after cooling from the austenitizing temperature are verysignificant. After the first austenitizing treatment, the forgings are to bethoroughly cooled down to 180 C and then held at 300 C. After the second aus-

' tenitizing treatment, the forgings are to be thoroughly cooled down to 180 C,heated up to 600 C and then cooled to 100 C. As mentioned in Section 2, thisconditioning treatment at 600 C appears effective to complete the bainitictransformation of these very large 3.50i-Cr-Mo-V steel forgings.

In 1970, Japan Steel Works constructed the 4th Heat Treatment Shop forquality heat treatment of gigantic generator forgings. The new shop was re-quired to obtain sufficient handling capacity, fully controlled furnaces andquenching equipment. The main dimensions of the 4th Heat Treatment Shop areshown in Table 14 and Figure 4.

The automatically controlled electric furnaces were designed and construc-ted by Japan Steel Works with the cooperation of an excellent specialist,Swedish heating element manufacturer.

The quenching equipment is designed to water spray or immersion quench,and fog or fan cool. At the present time, we prefer the violent water sprayquenching for gigantic generator forgings because of the following reasons:

(1) The great amount of pressurized water spray violently attacks thescaled forging surface produced during austenitizing.

(2) Steam produced during quenching is easily dispersed.

(3) The severity of quenching is easily controlled to be compatible withthe forging diameter.

Table 14 Main dimensions of the 4th heat treatment shop

Max. weight of products to be handled; 265 tons

Max. length of products to be vertically handled; 201000 mm

Vertical heating furnaces;3 units

No. 4E1 Furnace for gigantic generator rotor forging.

No. 4E2 Furnace for gigantic generator rotor forging.

No. 4E3 Furnace for low pressure turbine rotor forging with large

diameter.

Type; Electric furnaces

Control; Automatic

Inner diameter; 2,500 mm

Height; No. 1 Furnace 20,000 mm

No. 2 Furnace 15.500 mm

No.3 Furnace 12,000 mm

Vertical quenching equipment; 1 unit

Available quenching method; Water spray quench, water immersion

quench,fog cooling and fan cooling.

Max. length of products to be water sprayed; 19,000 mm

Note: The length would be made longer in the future.

Max. depth of water immersion tank; 17,000 mm

Water volume to be used;

4uenchingpit; 420ton

Reserve tank; 1,000 ton

- 28 -

The quality heat treatment cycles for gigantic generator forgings areshown in Figure 37. All forgings are to be double austenitized, spray quenchedand then tempered.

Before final heat treatment, all forgings are to be metallurgicallyexamined by the following three methods:

(1) Ultrasonic test(2) Sulphur prints on both ends(3) Check analysis on both ends

After metallurgical evaluation of the forging by the above tests, thefinal heat treatment schedule is determined.

In about 10 hours after the beginning of the spray quench, the surfacetemperature of an 1,800 mm diameter forging is to be lower than 100 C and theinner temperature is assumed to be about 350 C. The inner temperature is to belowered to less than 150 C by continued cooling with a mild water spray, fog orfans. The total quenching time of the 1,800 mm diameter forging is 16-20 hours.

The tempering temperature of these massive forgings should be greater than610 C. The higher tempering temperature is required to increase the fracturetoughness and to reduce the residual stress as explained in Part 2.

The residual stress at the forging surface should be less than 6 kg/mm2.

The measuring method and test results are described in Part 2 of this presenta-tion.

4. Future of i antic for in s - Conclusion

Since the first pouring of 400 metric ton ingots in 1969, the developmentsin minimizing segregation, reducing internal defects and improving both strengthand fracture toughness have been progressively accomplished. From the metallur-gical point of view, the selection of chemical composition, application ofvacuum carbon deoxidizing, AP process and improved forging techniques, alongwith the study of heat treatment to obtain fine grain structure have beendeveloped. Based upon the results of the above developments, the manufacture ofgigantic forgings is now successfully being performed.

It is necessary to continue to further investigate the character of segre-gate streaks as influenced by alloy elements, deoxidation and heat treatment.

Based upon these successful experiences with 400 ton ingots, we shall pro-duce still larger ingots in the near future.

- 29 -

References

1) B.C. Smith and G.S. Hartman, " The manufacture of large generator rotor

forgings over 135 metric tons " 5th 1114-1970, Terni

2) C. H. Meyer, " Erfahrungen bei der Herstellung von Schmiedestrdcken

aus Bldcken grdsser als 250 t " 5th 1FM. 1970, Terni

3) S. Kawaguchi and K. Kudo, " Segregation and heat treatment of large

forging " 5th 1FM. 1970, Terni

4) D.L. Newhouse and D. R. DeForest, " Meeting Requirements for larger

generator rotors I a metallurgical challenge " 5th IFM. 1970, Terni

5) R. Schinn, " Beitrag zum Stand der Anwendung und der Prdfung von

Schmiedestdcken fdr Dampfturbinen und Generatoren " Stahl und

Eisen, 81(1961) p. 1632

6) P. Opel, C. Florin, F. Hochstein and K. Fisher, " Aussichtsreiche

Nickel-Chrom-Molyban-Stdhle fibir Turbinen- und Generatorwellen "

Stahl und Eisen,90(1970)p. 465

7) C. J. Boyle, R. M. Curran, D. R. DeForest and D. L. Newhouse, " Further

progress in the development of large steam turbine and generator

rotors " ASTM 68th Annual meeting, 1965, W. Lafayette

8) R. M. Curran, " Progress in the development of large rotor forging "

5th 1FM. 1970, Terni

9) H. Greenberg, " Summary of mechanical properties of large Ni-Mo-V

generator rotor forgings (1955 - 1964) " ASTM 68th Annual meeting,

1965, W. Lafayette

10) W. Steven and A G. Hynes, " The temperature of formation of marten-

site and bainite in low-alloy steels " 3.I.3.1, 183(1956) p. 349

- 30 -

11) G. C. Gould, " Temper embrittlement in high purity 3.5Pli, 1.75Cr,

0.20C steel " ASTM STP407,ASTM, 19661 P. 59

12) " Effect of reeidual elements on temper embrittlement of NiCrMoV

rotor steels. " Research Sub-group on Temper Embrittlement, ASTM

Special Task Force on Large Turbine and Generator Rotors, Sub-

committee VI on Forginge of ASTM committee Al on steel, 1971.

13) E. Pouillard " Some metallurgical problems encountered in produc-

tion of generator rotors (French) " Jounées Internationales de la

Grosse Forge, 1963, Paris p. 43

14)V. Scalise " The problem of segregations on large forging " nn-

published, 1970, Italy.

15) J. Javelle and M. Common, " Evolution des températures dens le

grosses pièces " Jounées Internationales de la Grosse forge, 1963,

Paris, p. 463

16) E. A. Reid) " Forging method and tooling for production of large

forging " International Forging Conference, 1967, Sheffield

17) M. Tateno and S. Shikano, " Study on closing of internal cavities

in heavy forgings by hot free forging " Jounées Internationales de

la groese forge, 1963, Paris, p. 473

- 31 -

Figure 3 Preliminary heat treatment

igure 4

Quality heat treatment

Figure 5 Machining of gigantic generator forging

1000

800

600

2,0%Ni-Cr-Mo-V- - - -''' -... ... '"' .... -

... a%.. N. ..‘. .. -" '' \---. F+P '

. .. . .

400 me \ .

\ \200 \ \ \

Hy 325 10 270

1000

800

600

400 s

0 200

k0

M 1000

800

(-4600

4008

200

1000

400Ms

200

2.8%Ni-Cr-Mo-V

3.5%Ni-Cr-Mo-Y

1 lo 103lo 2

Time (sec)

235

c ..... .... ..... ............unc.,. ....,.... ..... ".- ."' ...

... ... ..... •... ....3/4. ... .... ...‘....... ... '.. %.... -.. -... \ ‘

%. -N.'`.

\ \ \ \‘ \

\ \ \ \

\ \\ \

\ B ' \\ \

\\ \

1 \ \ \

Hy 405 397 327331 328

...... , , ........ •.., ., \..... ....% N

....\ NN \ \ \

\ \ \—\ \\ \ \ \

\ \ \ \\ \ \

\ \ B \\ \

\\ - \ \ 1

Hy 475 420 405 380

3.50i-Mo-V

800 ---" e , -- --, , . .., . . .

a a '.. . a

600%.N. %..s. \ -

-. \

\ N-

\\ N N \

N B \ \

\\ \ \ \\ \ \

‘ \ --%

% X \Hy 430 375 2 0 260 250

10

Figure 6 C-C-T diagrams of four steels

105

0 2 4 6 8 10 12 14

Time (hr)

Figure 7 Cooling curves of 1,800mm diameter rotor forging

350

C2.00i-Cr-Mo-V

700 C

60o C

325

C

2.801-Cr-ho-V

3.5%Ni-Cr-Mo-V

45I5C

300

C

Temperature

380

C

375

C

6 0

c

590

C

500

C

Figure

8 Dilatometric study on transformation during cooling from austenitizing temperature

( Austenitizing temperature ; 840 C

Cooling rate ;60C/hr/

F., 0.30

bt 0 0.20

0.4o

0.10 30

10

-20

Top

GAB

GAC

Location

Location

Bottom

Figure

9Influence of segregation along axial center core material

on mechanical properties (Longitudinal specimen)

.M 0

250

220

210

200

190

160

Top

170

160

150

0 10

0 20

030

0 40

0 50

0 00

070

0 B

OD

90

0

Bottom

Distance

from

body

surface

)

Cen er

Figu e 10

Distribution

of hardness and microstructure from surface

to center

x 100

-5055 60 65 70 75 80 85 90

T.S (kg/mm2)

Figure 11 Tensile strength and impact propertiesat center section of 3.5% Ni-Mo-V generatorrotor forgings

Top

Chemical composition 043)

Location C Si Mn P S Ni Mo V

.22 .26 .52 .011 .013 3.73 .35 .15

.21 .26 .52 .011 .014 3.69 .34 .15

.19 .24 •49 .009 .012 3.65 .32 .14

••• ••

20

E-410

0

Microstructure ( x 100)

Location

Bottom

Figure 12 Properties along axial center core of rotor GBL(Longitudinal specimen)

0

+a

0• •••

cl7 300

I.

a.

700

••••

500

400

E. 200

10010 10 10

Figure 13 Isothermal decomposition diagram of retained

austenite of 3.5%Ni-Cr-Mo-V steel

•

Decomposition start

•

No decomposition occurduring cooling

Holding times (sec)

Decomposition finish

10

• Decomposition occuduring cooling

gl 600 • •

a)+3 •

0-.4 500004.2 0 0V0

0

•

4001 lo 102 10

Holding time at conditioning temperature (sec)

105

105

Figure 14 Decomposision of retained austenite of

3.5% Ni-Cr-Mo-V steel by conditioning treatment

Preliminary heat treatment

1000

1000 C

900

C0

Air cool.

800

Fan cool.

cl63

0 C

O60

04.

)

400

IDa O

200

Furnace cool.

1000

900 C

840

c

•

800

Fan cool.

Water spray

T59

0 -

640

CO•

600

4.) O•

/too

a

200

Control cool.

F

Quality heat treatment

Time

Figure 15 Heat treatment cycle of 3.5%Ni-Cr-Mo-V steel

low pressure rotor forgings

8

3.5% Ni-Cr-Mo-V steelAlistenitizing temperature: 90 C

1 10 102

10 3 104

105

Heating rate (C/hr)

Figure 16 Influence of heating rate to austenitizingtemperature on austenitic grain size

0

E. 2cn

0

0

3.5% Ni-Cr-Mo-Y steelX Reaustenitizing

temperature: 900 C1.00

0

a-rc Heating rate

50 C/hr

20

4.ID

0

0LI) 4

14

+1

0

//Heating rate of400 c/ r

Ott

4 2 0 -2

Prior austenitic grain size (ASTM No.)

Figure 17 Influence of prior austenitic grain sizeon grain size after reaustenitizing

04-1 4

0 Austeni izing temperature:9-1 900C

61 2 3 4 5

Ni (%)

Figure 18 Influence of nickel content on austenitic grainsize of Ni-Cr-Mo-V steel

50

0

.0040.As A

_ Sb

0 0.01 0.02 0.03 0.04 0.05 0.06P, Sn, As and Sbx10 (90

Note: AFATT = TTFC - TTWQ

TTFC FATT after furnace cooled (50 C/h)from 600C

TTWQ FATT after water quenched from 600 C

Figure 19 First laboratory data on influence of residualelements on temper embrittlement of 3.50i-Cr-Mo-Vsteel

••••••

a.

15 0

100o o P

50 Sn0'3/44

0

.0.0 ,a As Sb

0 0001 0,02 0003 0,04 0.05 0.06 0.07

P, Sn, As and Sbx10 (%)

Note: ,AFATT = TTSC - TTWQ

TTSC FATT after step cooled from 595 C 11)

TTWQ FATT after water quenched from 595 C

Figure 20 Second laboratory data on influence of residualelements on temper embrittlement of 3.5%Ni-Cr-14o-Vsteel

200

150

50

0

100

-4fa.

0 000

•

02,0

oo

o 8o 00o

0 0OD

0 0.010 0.020 0.030 0.040 00050

P + Sn (%)

Figure 21 Influence of residual elements on temperembrittlement of center core material of3.50i-Cr-Mo-V low pressure turbine rotorforgings by step cooling

0.20

(;) •••••••••C I 70Y•

btc,f .%.""hp 600

0.30

80

5080

4o FATT

NDTTEi

-4o

—80

GCA

GCB

Top Bottom

GCA GCB

Microstructures ( x 100 )

FATT

NDTT

Location

Figure 22 Properties along axial center core of two gigantic3.594Ni-Cr-M0-V generator rotor forgings

( Longitudinal spocimen )

Top(T.S) .

Top(T.S)

Top(T.S)

Forging: GAC

Forging: GCA

Forging: GBF

MaterialBody diameterShipping weight:Bore diameter



,

20i-Cr-Mo-V1,808 mm196 ton400,220, 250 mm

3.5%Ni-Cr-Mo-V1,808 mm172ton400, 220, 250 mm

3.5%Ni-Mo- V1,685 mm176 ton300, 205 mm

Bottom(E.S)

Note 1. Number of flaw indications in body: 59 indications2. Maximum calculated flaw size 3,3mm (0.13 in.)

dia.

Bottom(E.S)

Note 1. Number of flaw indications in body: 6 indications2. Maximum calculated flaw size : 2.4 mm (0.10 in.)

dia.

Bottom. . (E.S)

Note 1. Number of flaw indications in body: 4 indications2. Maximum calculated flaw size : 2.0 mm (0.08 in.)

dia.

Figure 23 Typical ultrasonic test results of generatorforgings made from 1400ton ingots

Figure 24 Typical porosities in gigantic forging

x /too

x 100

Component

Component

Figure 25 Non-metallic inclusions observed in giganticforgings

x 100

0 .0 p.

0

X

-I

fl

•

00 4csio

Type A

Type B

Type C

c.. k '1\ 1).(14.0.!

Z

+ :["(et

,;2'. ' r

, „•

4',t. .

. -t" e• 't8 N>:." vni"'10 •-•,*

Figure 28 Typical patterns of sulphur prints of centercore materials of large rotor forgings

Type A

Type B

Type C

Figure 29 Typical patterns of macrostructures of centercore materials of large rotor forgings

Carbon (%)

.4o

. 30

.20

.10

.40

. 30

.20

.10

Bottom Top

AP Process: Conventional

Forging A : 2% NiCrMoV-172 ton

•

B : -196 ton

•

D

•

-196 ton

AP Process: Developed

Forging E : 3.5% NiCrMoV-172 ton

•

F 3.5%NiMoV -127ton

x_ ________ --x-ow/

AA

D

_ x ______-x-- ---- - F

-

Figure 30 Axial carbon segregationat center core of

gigantic forgings

Figure 31 Alloy segregation of center of generator forgingmade from 400 ton ingot

!4?,,1-2BEF)

No. 1-2 BEF• • (C, 0049.

No. 3 BEF(C : 0.23)

No. 15 BOHC : 0.30)

Figure 32 Schematic positions of earch heat withinthe 400 ton ingot mould

4,000

3,000P0bk0

o 2,000

Izo4

1,000

10 20 30 4o 50

Time (hr)

Figure 33 V rtical solidification curve of 4o0 ton ingot

Ingot

Forming ofhandling stems

JTS forging

Rough forging

Finishingof body

•Finishing oftop side journal

Finishing

.><

1,880

•••• ••••• ••• •

- 3,55 0 -

4,965-4- 7,780-4- 3,345

Forging ratio : 3.5 in body

As forged weight: 288 tons

Figure 34 Forging process of gigantic generatorforging weighing 200 metric tons(Rotor forging: GCS)

Ingot: 400 ton

Reheating tempera-ture:1,270 C - 1,220 C

JTS Forging:twice

Ingot

Forming ofhandling stem

Rough forgingJTS forging

Rough forging

Finishing ofbody

Finishing oftop sidejournal

Finishing

1,680

3,550

r•-• 0cy.N

C..

3,020 2 ,140

Forging ratio : 1.7 in body

As forged weight : 239 tons

Figure 35 Forging process of gigantic forging with

2,700 mm diameter(Rotor forging: TA)

Ingot: 400 C

Reheatingtemperature:1,270 - 1,220 C

1000

800

\

•

600

7 O•

400

A •200

1000

o 800

•

600

0400

A200

1000

Air cool

200

20

200

1050

2.8%Ni-Cr-Mo-V & 3.50i-Cr-Mo-V steel

1000

Air cool

80

300

900 950

Air cool

2‘Ni-Cr-Mo-V & 3.5%Ni-Mo-V steel

600/650

900/950

Air cool

Air cool

600

600/650

Air coo

Furnace

cool

100

300Furnace cool

Figure 36 Preliminary heat treatment cycle

180

Time

Time

1000

800

0

600

4.) 0 O

400

a F•

200

1000

900 Fan cool

zo

820/840

spray

quench6Io 640

2%Ni-Cr-11110

-V &

3.50

i-M

o-

V •teel

400

900

84o

0800

Fan cool

600

Spray quench

Si

610/640

•

600

ItAir cool

400

a

ooE" 200

180 10

100/

150

Figure

37Quality heat treatment cycle'

Time

2.80i-Cr-No-V

& 3

,50i

-Mo-

V steel

Tin

Appendix I

List of for i $ roduced from ton i ots

Designation Material Body diameter Total length Weight Note

GAA

Four poles generator rotor forging

20i-Cr-Mo-V 1,808 mm 13,600 mm 172 ton

GAB u 1,808 14,800 196

GAC II 1,808 141800 196

GAD II 1,808 13,600 172

GBA 3.50i-Mo-V 1,706 14,200 158

GBB tI 1,632 16,480 155

GBC li 1,714 14,926 175

GBD II11706 14,550 158

GBE II1,651 12,655 127

GBF II1,685 151310 176

GBG II 1,685 151310 176 *

GBH II11706 14,550 158 *

GBI H 1,708 14,750 177 *

GBJ II 1,708 14,750 177 *

GCA 3.50i-Cr-Mo-V 1,808 13,600 172

GCB II 1,808 14,920 200

GDA 2.8%Ni-Cr-Mo-V 1,808 13,660 169 *

GDB It 1,808 16,780 239 *

GEC II 1,808 14,880 196 *

GEA Carbon steel 1,557 15,055 151 *

Turbine rotor forging

TA 10i-Cr-Mo 2,710mm 7,760mm 158 ton

TB 20i-Cr-Mo 2,710 7,760 158

Back-up roll forging

BA Cr-Mo 2,100 mm 10,250mm 150 tonBB 2,100 10,250 150

Designation Material Body diameter Total length Weight Note

PA Ito 4,597 x 4,597 x 1,211 mm 150 ton

PB 4,369 x 3,962 x 1,524

170

PC 51486 m 3,302 i 1,219 139

* s Now inproduction

Parts of forging press

Appendix II

Conversions tables

Length

Weight ( fl )

mm in mm in mm ft in mm ft in

3 0.12 1,632 64.3 3,800 12 6 14,880 48 103.3 0.13 1,651 65.0 3,962 13 0 14,920 48 115 0.20 1 p680 66.1 4,040 13 3 15,055 49 615 0.59 1,685 66.3 4,180 13 8 15610 50 2100 3.9 1,704 67.1 41369 14 4 15600 50 10200 7.9 1,714 67.5 4,500 14 9 16,480 54 1205 8.1 1,720 67.7 4,597 15 1 16,780 55 1220 8.7 1,800 70.9 4,800 15 9 17,000 55 9250 9.9 1,808 71.2 41965 16 3 18,000 59 1300 11.8 1,880 74.0 5,040 16 6 19,000 62 4400 15.7 2,000 78.7 51486 18 o 20,000 65 7450 17.7 2,075 81.7 6,000 19 8 26,500 87 0500 19.7 2,100 82.7 7,045 23 1600 23.6 2,140 84.3 7,760 25 6700 27.6 2,290 90.2 7,780 25 7800 31.5 2,500 98.4 8,000 26 3900 35.4 2,600 102.4 10,250 33 8

1,000 39.4 2,620 103.2 12,000 39 41,020 40.2 2,670 105.112,655 41 61,210 47.6 2,710106.713 160o 44 71,220 48.0 2,760 108.713,660 44 101,300 51.2 3,000 118.1 14,200 46 71,370 53.9 3,545131.714,550 47 81,524 60.0 3,500 137.814,750 48 41,557 61.3 3 650 139.814,800 48 7

ton lb ton

Weight I )

lb ton lblb ton

10 22,000 127 2801000 170 345,000 265 584,00027 60,000 132 291,000 172 379,000 288 635,00035 77000 139 306,000 174 584,000 300 661,00040 88,000 140 309,000 175 386,000 315 694,00045 99,000 150 331,000 176 388,000 400 882,00060 132,000 151 333,000 177 590,000 420 926,00090 198,000 155 342,000 196 432,000 5001 0.02 •00093 2051000 158 348,000 200 441,000100 220,000 165 364,000 239 527,000120 265,000 169 373,000 250 551,000

metric ton 1,000 2,000 3,0004,000 6,000 101000short ton 1,102 2,205 3507 4,409 6,614 11,023

Stress

Temperature

-110 -148 -10 14 80 176 450 842 800 1,472-90 430 -5 23 100 212 455 851 810 1,490-85 -121 0 32 120 248 465 869 820 1,508-80 -112 5 41 130 266 500 932 840 '1;544-70 -94 10 50 180 356 520 968 900 1,652-65 -85 15 59 200 392 590 11094 920 1,688-60 -76 20 68 300 572 595 1403 930 1,706-55 -67 25 77 325 617 600 1,112 950 1,742-50 -58 30 86 335 635 610 1,1301,0001,832-45 -49 35 95 340 644 620 1,148 1,0501,922-40 -40 40 104 350 662 630 1,166 1,220 2,228-35 -31 45 113 360 680 635 1,175 1,2702,318-30 -22 50 122 375 707 640 1484-25 -13 55 131 380 716 650 1,202-20 -4 60 140 400 752 660 1,220-15 5 70 158 410 770 700 1,292

Tem erature difference

1 1.8 9 16.2 0 86.4 88 158.4 200 360.02 3.6 10 18.0 50 90.0 100 180.0 400 720.04 7.2 11 19.8 60 108.0 109 196.25 9.0 37 66.6 63 113.4 150 270.06 10.8 42 75.6 76 136.8 167 300.68 14.4 44 79.2 82 147.6 169 304.2

MANUFACTURE OF

LARGE HEAD FORGINGS

FOR

NUCLEAR REACTORS

Prepared by

H. C. SmithAssistant Chief Metallurgist

Bethlehem Plant

G. S. HartmanMetallurgical Engineer

Bethlehem Steel CorporationBethlehem, Pennsylvania

Preprint of paper to be presented byH. C. Smith before the International Forgemasters' Meeting

to be held in Cherry Hill, New Jersey, USAin October 1972

TABLE OF CONTENTS

Page

ABSTRACT 1

INTRODUCTION 2

PLANNING 5

STEELMAKING 5Melting 6Pouring 6Stripping 6

MANUFACTURING 7Forging 7

LWBR Closure Head 7FFTF Closure Head and Hydrostatic Test Head Halves 12

Initial Machining and Ultrasonic Examination 16LWBR Closure Head 16FFTF Closure Head and Hydrostatic Test Head 17

Torch Cutting and Machining to Contour 19Heat Treatment for Mechanical Properties 20

LWBR Closure Head 20FFTF Closure Head and Hydrostatic Test Head 203-1/2% NiCrMo Analysis 24

Finish Machining and Final Inspection 26LWBR Closure Head 26FFTF Closure Head and Hydrostatic Test Head 27

Fabrication of the FFTF Closure Head 30

DISCUSSION 32

APPENDICESI Conversion Tables 35II Mechanical Properties (LWBR Closure Head) 36III Mechanical Properties (FFTF Closure Head and 37

Hydrostatic Test Head)IV Determination of Critical Cooling Rate 38

for Embrittlement of 3-1/2% NiCrMo

Determination of Minimum 38Reversible Embrittlement Temperature

REFERENCES 39

Number

II

III

IV

V

LIST OF TABLES

Sub'ect

Chemical Analyses

Mechanical Properties(LWBR Closure Head)

Page,

6

22

Mechanical Properties 25(FFTF Closure Head and Hydrostatic Test Head)

Determination of Critical Cooling Rate 26for Embrittlement of 3-1/2% NiCrMo

Determination of Minimum 26Reversible Embrittlement Temperature

Number

LIST OF FIGURES

Title 22E2.

1 Nuclear Reactor Vessel - Bolted Closure 3

2 3300 mm Diameter 310,000 kg Ingot After Stripping 7

3 Ingot on Carbottom Furnace After Heating to 1260 C 8

4 Rejectable Condition in First LWBR Closure Head 10

5 Forging Sequence of Second LWBR Closure Head 11

6 Slabbing with 1220 mm Wide Flat Dies 13

7 Forging Sequence of FFTF Closure Head 14

8 Chipping Surfaces of FFTF Slab 15

9 Machining FFTF Slab on 4.35 m Ingersoll Milling Machine 17

10 Ultrasonic Examination of FFTF Slab 18

11 Torch Cutting FFTF Closure Head Half from Slab 19

12 Mechanical Property Test Location (LWBR Closure Head) 21

13 Mechanical Property Test Location (FFTF Closure Head 23and Hydrostatic Test Head)

14 Machining FFTF Closure Head Half on 254 mm Horizontal 27Boring Mill

15 Inspection of FFTF Closure Head Half 28

16 Magnetic Particle Inspection of FFTF Hydrostatic Test Head 29

17 Final Ultrasonic Examination of FFTF Hydrostatic Test Head 29

18 Gas-Fired Equipment and Electrical Heater Used for Welding 31FFTF Closure Head

19 Preheating Furnace Used for Welding FFTF Closure Head 31

20 Completed Weld (FFTF Closure Head) 32

ABSTRACT

A major challenge exists in producing large forged closure heads fornuclear reactors. The combination of thickness and diameter of these forgingsnecessitates a thoroughly planned, well-controlled and properly executedmanufacturing sequence to assure the required quality in these massiveconfigurations.

A closure head for a prototype Light Water Breeder Reactor (LWBR)measuring approximately 3930 millimeters (mm)* diameter x 1295 mm thick wasproduced from ASTM A508 Class 4 analysis (.18 C, 3.5 Ni, 1.75 Cr, .50 Mo).Both a closure head and hydrostatic test head were made for a prototype SodiumCooled Breeder Reactor. These heads were produced as two semicircular discswhich were welded together by Combustion Engineering, Inc. (CE) using itsSub-Vert process - the approximate overall dimensions are 7670 mm diameter x560 wm thick. The grade of material was ASTM A508 Class 2 (.22 C, .70 Ni,.35 Cr, .60 Mo).

Steelmaking, vacuum degassing and ingot pouring operations are described.All forgings were produced from 3300 mm ingots.

The open die forging was accomplished on a 6800 metric ton press. Theplanning and execution of this phase of manufacture were the most critical ofall the manufacturing steps. The size and configuration of the forgingspresented considerable risk in obtaining the required quality. A modificationof the standard forging technique was used to successfully produce the headfor the LWBR.

Larger forgings, such as generator rotors, have been produced. However,ingots for those rotors are blocked out and ultrasonically examined beforebeing committed to a drawing. In the case of the large head forgings, theyare forged to their final shape before a quality evaluation can be made.

Various machining operations, as well as torch cutting, to achieve shapepreparatory to heat treatment for mechanical properties are described. Non-destructive examinations, both ultrasonic and magnetic particle, wereperformed in accordance with agreed upon specifications and standards atpertinent steps during the manufacturing sequence.

The heat treatment was straightforward except that two quenchingtechniques were used. The LWBR head was immersion quenched in a 113,000 litertank with circulating water, whereas the Sodium Cooled Breeder Reactor headswere spray quenched. Tensile and impact tests were removed from designatedtest locations and results are reported. Data on temper embrittlement arereported on the ASTM A508 Class 4 material.

* - Conversion Tables for all metric units included in this report areshown in Appendix I.

INTRODUCTION

At the Fourth United Nations Conference on the Peaceful Uses of AtomicEnergy held in September of 1971, some startling predictions were made. Thepresent world energy requirement was estimated to be .15 Q (Q = .25 x 1018kilocalorie). By the year 2000, this requirement is expected to be 1 Q, by2070, 9 Q; and finally by 2100, the energy needed to service an estimatedworld population of 16 billion is expected to be 16 Q. It is shocking tothink that in the next 28 years we may have to increase our energy outputover six times its present value.

The result of this increased energy requirement, the Conference was told,would be that the world's fossil fuel resources would come under increasedpressure in the next century. However, it was stated that nuclear fission inbreeder reactors would provide an ample energy supply for many, many centuries.Obviously, a necessary accelerated shift toward nuclear fission as a source ofpower is indicated. In the United States this shift has started to take place.In 1970, there were 90 nuclear power plants, 175 megawatts (mw) or over,operating or under construction. The total rated capacity of these plants was70,899 mw. In 1971, there were 104 plants with a total rated capacity of86,794 mw.

Scientists claim that the absolute energy supply available will not be alimiting factor in the future fulfillment of our energy requirements. However,economic cost will become critical. With nuclear power plants, increased unitsize can result in a downward trend in costs. Understandably, as the plantsincrease in size, the component parts of the plants and the reactor vesselswill also increase in size. This places a burden on manufacturing facilitiesfor producing forged steel reactor components, such as closure heads, vesselflanges, shells and support equipment.

This paper will discuss the manufacture and processing of two large, forgedsteel, nuclear reactor closure heads. In certain designs, the forged closurehead is the most massive component of the reactor vessel. A sketch of a typicalbolted closure nuclear reactor vessel is shown in Figure 1. The closure headis located at the top of the vessel and is usually bolted to the vessel. Thecontrol rods, instrument ports, and access ports for fueling pass through thepenetrations in the head.

It should be stated, at this point, that most nuclear reactors in operationtoday do not have forged steel closure heads. Closure heads are usuallymanufactured from plate material which is formed into a dome and attached tothe vessel shell. However, material property and quality requirements maydictate the use of forged steel.

control rod

Figure 1Nuclear Reactor Vessel - Bolted Closure

closure head

bolt

shell

The size of a closure head forging is as varied as the size of thereactor vessel itself. They may vary in thickness from 560 mm to 1420 mmwith diameters from 1950 to 7620 mm. The first closure head which will bediscussed was used in a prototype IWBR. It measured approximately 3930 mmdiameter x 1295 mm thick.

The second closure head to be discussed will be used in a prototype FastFlux Test Facility (FFTF) and measured 7620 mm diameter x 560 mm thick. Aforging of this size cannot be made in one piece on existing equipment.

Therefore, two rectangular pieces, 680 mm thick x 4440 mm wide x 7860 mm longwere forged. A semicircle, 3810 mm x 7620 mm, was burned out of eachrectangular piece. These half discs were submerged arc welded by CE toproduce the circular closure head. Since this head contained unusually largepenetrations, a head of the same diameter and thickness dimensions was neededto perform the hydrostatic testing of the vessel. This hydrostatic test headwill also serve as a backup head for the FFTF vessel. Hence, four half discswere manufactured.

The FFTF, a Sodium Cooled Breeder Reactor, is the testing laboratory forAmerica's next generation of nuclear power plants. Advanced breeder reactorsshow promise for extending the world's present uranium resources by hundredsof years. These advanced breeders are expected to produce more fuel thanthey consume. The FFTF closure head represents the largest forged head evermade and, after welding the halves together, it is expected to weigh193,000 kilograms (kg).

Components of this size represent a challenge to the forger because theirmassiveness strains the capabilities of existing forge equipment. In addition,not only must the closure heads be forged to the proper size and shape, theymust also be made to very high quality levels.

To manufacture the extremely large closure head forgings discussed inthis paper, ingots weighing in excess of 272,000 kg were used. Prior to theprocessing of the above two types of closure heads, Bethlehem had made a headforging 4320 mm diameter x 680 mm thick. This forging was made from a216,000 kg ingot of a 3-1/27 NiCrMo analysis. The excellent qualityexperienced with this closure head inspired confidence for moving into largersizes. With larger and heavier ingots, it becomes more difficult to attainthe required quality level. Therefore, optimal forging practices are neededto insure meeting the quality. Anything less than optimum may not be goodenough as witnessed by the experience in forging the LWBR closure head. Aswill be discussed shortly, the first LWBR closure head was rejected becauseof unsatisfactory internal quality. Prior to reviewing the manufacture ofthe LWBR and FFTF closure heads, it may be advisable to briefly describe theforging operations used in their production.

The forging of closure heads can be accomplished by two methods. Thefirst and most common method is to upset a blocked ingot to form a large discand then finish to the desired configuration by spreading. However, closurehead forgings may be of a size that cannot be formed by upsetting. In thiscase, the second method is to forge the ingot into a slab of the desiredthickness and the closure head or closure head segments can be torch cut fromthe slab and individually handled. The segments are then welded to form thedesired head.

The LWBR closure head was formed by upsetting and spreading and the FFTFand Hydrotest head was made from halves which were burned from rectangularslabs.

Design and strength requirements govern the material selection for closureheads. The LWBR closure head was made from a 3-1/2% NiCrMo analysis, ASTM A508Class 4, (.18 C, 3.5 Ni, 1.75 Cr, .50 Mo); the FFTF heads were made from a leanNiCrMo analysis, ASTM A508 Class 2, (.22 C, .70 Ni, .35 Cr, .60 Mo).

The remainder of the paper will present a general description of overallquality control procedures, discuss details regarding forging technique, andgive a brief synopsis of the balance of the fabrication sequence. Inspectionand test procedures will be discussed and pertinent results will be highlighted.

PLANNING

As mentioned earlier, both closure heads were for prototype nuclearreactors and these designs were unique. Lines of communication wereestablished between the forge, heat treatment, and machine shops, as well asthe reactor designers, fabricators, contractors, and all other concernedparties involved with the nuclear reactor project. Massive forgings requirea cooperative effort and there were constant negotiations between the designersand producer. The designers had to be made aware of the capabilities of theproducer, the producer had to interpret the designer's requirements andunderstand the needs of the fabricators.

The first step in assuring quality is thorough planning and this planningwas undertaken before the customer's order was received. The melting practicewas chosen, the maximum ingot size available was selected, and sufficient ingotdiscards were planned. The method of forging and heat treatment cycles wereselected. Procedures for nondestructive examination were developed and agreedupon by the designer. Forging and machining allowances and machine sequenceswere incorporated into the planning. In addition to metallurgical quality,mechanical quality, i.e., size, shape, and straightness had to be considered.

By the time the orders were received, the planning was essentiallycomplete. On initial orders, such as these closure heads, an effective plancannot be developed without discussions with the customer. The importance ofthis communication cannot be overemphasized.

STEELMAKING

Steel for the two LWBR closure heads and the four halves of the FFTFclosure head and hydrostatic test head was melted in basic electric arcfurnaces. Each forging was processed from an ingot requiring the product offive furnaces. Four of the furnaces yielded approximately 70,000 kg each andthe fifth yielded approximately 29,500 kg. A double slag practice with a1.0% Fe() maximum on the refining slag was used on each of the furnaces.

Melting

Temperature control is critical during the melting and refining periodsand a number of determinations were made on each of the five heats to assureconformance to the standard melting practices.

The average weighted chemical analyses and weights of the six ingotsare shown in Table I.

Pouring

Stripping

Table I - Chemical Anal ses

Ingot Weight Wei hted Anal ses (%)

Number (kg) C Mn P Si Ni Cr V Mo Cu Sn Co_ _

A 225,000 .17 .32 .008 .008 .25 3.46 1.79 .02 .50 .08 .01 .006B 288,000 .17 .31 .010 .008 .20 3.42 1.66 .02 .50 .06 .01 .017C 310,000 .21 .65 .007 .008 .27 .73 .36 .01 .60 .06 .01 .02D 310,000 .21 .61 .006 .009 .25 .71 .35 .01 .59 .06 .01 .01E 300,000 .23 .62 .009 .012 .24 .73 .31 .01 .61 .06 .01 .01F 300,000 .23 .56 .009 .013 .27 .72 .38 .01 .58 .07 .01 .01

(Note - Ingots A and B were for the first and second LWBR closure heads,respectively; C and D - FFTF closure head halves; E and F -Hydrostatic head halves.)

Each of the six 3300 mm diameter ingots was stream degassed. The 29,500 kgcapacity pony ladle was approximately 75% filled before the 57 lam diameterzirconium nozzle was opened to start pouring directly into the mold. The levelof metal in the pony ladle was kept sufficiently high to insure an uninterruptedflow of metal while pouring the ingot body at a rate of approximately 7650 kgper minute. The time required to pour each of the ingots was approximately44 minutes.

Vacuum was achieved by a five-stage steam ejector system. During pouring,a pressure of 1.5 mm or less was maintained after the initial surge at thestart of the pour.

All ingots were allowed to solidify a minimum of three days beforestripping from the mold. After stripping (Figure 2), each of the ingots wascharged into the forge furnace within seven hours.

MANUFACTURING

Forging

Figure 23300 firm Diameter 310 000 k In ot After Stri in

All of the forging operations on the LWBR and FFTF closure heads wereperformed on a 6800 metric ton steam hydraulic press.

LWBR Closure Head

The configuration of the LWBR closure head required that it be forged byupsetting. Because of its size, 3930 mm diameter x 1295 mm thick, the majorconcern was to achieve adequate forge work in the center region of the pieceto assure consolidation of porosity inherent to the center core of the ingot.It is the objective of the forging operation to produce sound forgings byconsolidating the inherent internal porosity. However, with extremely largeforgings, modifications and innovations in forging procedures are sometimesnecessary if sound forgings are to be produced.

Juretzek and Richter(1) state that a small degree or token upsettingaggravates porosity and they advocate a minimum upset ratio of 2.25 to 1.With the LWBR, the upset ratio would not approach this minimum because ofthe limitations of the press. In forging very large diameter pieces betweenupsetting plates, working of the metal is limited by press capacity.

Although it is Bethlehem's practice to obtain consolidation by blockingthe ingot before upsetting, again because of the size involved, the amount ofblocking, prior to upsetting, would not be sufficient in itself to assurecenter consolidation.

After upsetting to as large a diameter as possible, the forging is spreadbetween flat dies to the required diameter and thickness. Normally, thisoperation does not provide satisfactory die penetration to achieve significantconsolidation in a piece of this thickness.

Based on past experience, it was postulated that the combination ofblocking, upsetting, and spreading would provide adequate consolidation ofthe inherent porosity in the ingot.

Ingot A

This ingot was heated for forging in an oil or gas-fired carbottomfurnace, 5.2 meters (m) wide, 4.5 m high and 11.5 m deep. The surfacetemperature of the ingot, when charged, was 540 C. The time required to bringthe ingot to a uniform temperature of 1260 C was over 90 hours (Figure 3).

Figure 3In ot on Carbottom Furnace After Heatin to 1260 C

The 3300 mm diameter corrugated ingot was blocked using 1220 mm flatdies as follows:

After removing the top and bottom discards, the 2500 mm diameter x 4320 mmlong block was recharged into the heating furnace and held at 1260 C forapproximately 41 hours.

Ingot B

1-1/2 hours > 2500 mm diameter xingot 4320 mm long

3300 mm diameter

The blocked ingot was upset between the crosshead and pit plates:

2500 mm diameter x4320 mm long

3720 mm long

3720 mm diameter x17 minutes 3720 mm long

The upset ratio was 1.16 to 1. The capacity of the press was reached and nofurther upsetting was accomplished.

The forging had to be reheated four additional times to obtain thedesired configuration. The forging operations were a combination ofspreading and edging, using 1220 mm and 305 mm flat dies:

2720 mm diameter x 4190 mm diameter x4 hours 1400 mm high

After the forging operation was completed, the closure head wastransferred to the heat treatment furnace and double normalized and tempered(normalize - 1010 C, 925 C; temper - 690 C) . After this preliminary heattreatment cycle, the top face of the forging was machined to enable apreliminary ultrasonic examination to be performed. This examinationrevealed that continuous clusters of indications existed in the forgingfrom 305 mm to 910 mm from the top surface in an approximate 910 mm diameterconcentric circle (Figure 4). The condition was determined to be a core ofunconsolidated metal associated with segregation and the forging wasrejected.

From the experience with Ingot A, it was determined that two areas ofthe forging process should be modified. First, the need for a more effectiveupset ratio required that the blocked ingot should be soaked at the forgingtemperature for a longer period of time. Secondly, the need for moreeffective penetration at the center of the forging during spreadingnecessitated a different configuration for the final spreading operation.

The time required for Ingot B to be heated to a uniform temperature of1260 C was over 93 hours. The ingot was blocked as follows:

3300 mm diameteringot

1 hour5 minutes

2500 mm diameter x4320 mm long

After removing the top and bottom discards, the block was rechargedinto the heating furnace and held at 1260 C. The time to temperature andon temperature was approximately 77 hours (Ingot A was 41 hours). Figure 5illustrates the sequence of forging operations for making the second LWBRclosure head.

The blocked ingot was upset between the crosshead and pit plate:

2500 mm diameter x 3175 mm diameter x4320 tutu long 24 minutes > 2620 mm high

The forging upset ratio was 1.65 to 1. Ingot A had only been reducedto 3720 mm long.

The upset forging was recharged into the furnace, reheated, and heldat 1260 C for about 22 hours prior to being spread using 1220 mm flat diesas follows:

3175 mm diameter x 3650 mm diameter x2620 mm high 10 minutes 1930 tutu high

Figure 4Re'ectable Condition in First LWBR Closure Head

-10-

UltrasonicCondition

3300 mm

Step No. 2 Step No. 3

1Ingot Axis

Upsetting Between Crosshead and Pit Plates Spreading on 1220 mm Wide Flat Dies

3175 mm diameter x 2620 mm high. 3650 mm diameter x 1930 mm high.

1Ingot Axis

"\....

Bottom Discard


Stepping the Forging on 1220 mm Wide Flat Dies

Central Area 1930 mm high.

Overall length 4950 mm

1

Blocking on 1200 mm Wide Flat Dies I

2500 mm diameter x 4320 mm long. I

Figure 5

Forging Sequence of Second LWBR Closure Head

Step No. 1

Ingot Axis

I Top Discard

Spreading on 1220 mm Wide Flat Dies

4135 mm diameter x 1420 mm high.

At this point, an innovation was made. Two diametrically opposite sidesof the piece were forged to 147011/111high leaving the central area at 1930 mmhigh. An illustration of the stepped disc appears in Figure 5 as Step No. 4.This configuration would allow more effective penetration by enabling the useof full 1220 mm die bites and heavy drafts in the central area whereconsolidation is critical. This adaptation was based on the work reportedby Reid(2) demonstrating the improvement in consolidation at the center withincreasing die width. This operation took 30 minutes. The overall length ofthe stepped disc was 4950 mm. This length exceeded the distance of 4625 mmbetween the press columns and would not permit the disc to be rotated 90° forthe final spreading operation. Therefore, approximately 215 mm had to betorch cut from each end of the disc.

The stepped disc was then reheated to and held at the forging temperatureof 1230 C for 21 hours. The piece was removed from the furnace and rotated90° from the previous working direction. The spreading operation was completedusing the full width of the 1220 vu flat dies and heavy drafts. Immediatelyafter the spreading operation, excess metal was torch cut from the periphery:

Stepped Disc4135 mm diameter x

20 minutes 1420 mm high

The closure head forging was transferred hot to the heat treatmentfurnace and given a double normalize and temper cycle. (Normalize - 1010 C,925 C; temper - 690 C.)

A preliminary ultrasonic examination indicated that the new practicesfollowed on the second LWBR closure head were successful.

FFTF Closure Head and H drostatic Test Head Halves

The size and shape of the half heads necessitated that these be forgedby a combination of slabbing and cross forging. Slabbing reduces the ingotcross section to a rectangular shape instead of octagonal, square, or roundas is done in blocking. Flat dies are used to reduce the cross section andsimultaneously elongate the ingot body. To influence mechanical propertiesor develop greater width, cross forging with the dies set parallel to theingot axis is employed after the initial lengthening operation.

All four FFTF ingots were heated for forging in a carbottom furnace.The surface temperature of each of the ingots, when charged, was approximately595 C. The ingots were heated to the forging temperature of 1260 C and thetime required to bring each of them to a uniform temperature was a minimumof 95 hours.

-12-

Figure 6Slabbin with 1220 mm Wide Flat Dies

The initial slabbing operation was accomplished using 1220 mm wide flatdies (Figure 6). The 3300 mm diameter ingots were reduced as follows:

2290 mm thick x3300 l[u diameter 2790 mm wide x

ingot 1-1/2 hours 4060 ItsEl long

After removing the top and bottom discards, the slabs were recharged intothe furnace, reheated and held at a temperature of 1260 C for approximately45 hours. The sequence of forging steps is illustrated in Figure 7.

The second forge operation was to cross forge (turn 90°) using 450 mmwide flat dies from:

2290 mm thick x2790 mm wide x4060 mm long

45 minutes 1470 mm thick x3680 ium wide x4060 mm long

Each forging was reheated to 1260 C in about 50 hours and cross forgingcontinued using 450 mm dies from:


23 minutes

-13-


3300 mm

•Ingot Axis

Bottom Discard

Ingot Axis

Slabbing on 450 mm Wide Flat Dies

680 mm x 4440 mm x 7860 mm.

Step No. 1

///

Ingot Axis

Slabbing on 1220 mm Wide Flat Dies

2290 mm x 2790 mm x 4060 mm.


Step No. 4

Ingot Axis

Figure 7

Forging Sequence of FFTF Closure Head

-14-

Top Discard

Cross Forging on 450 mm Wide Flat Dies Cross Forging on 450 mm Wide Flat Dies

1470 mm x 3680 mm x 4060 mm. 1320 mm x 4700 mm x 4060 mm.

The forgings were then reheated for the third time to 1260 C in 40 hours.The 450 mm flat dies were used to finish by slabbing and edging:


680 mm thick x4440 mm wide x

1 hour >

7860 mm long

During the final slabbing operation, a problem may occur with pieces ofthis size. The available forging dies may not allow full penetration of thepiece with the result that the surface metal of the piece may flow more thanthe center. As a consequence, the edges may develop an hourglass or fishtailshape. Care must be taken to provide sufficient usable metal so that afterthe edges have been trimmed a satisfactory forging remains.

After forging, the pieces were transferred hot to the heat treatmentfurnaces. The preliminary heat treatment cycle consisted of a normalize at880 C and a temper at 620 C.

After preliminary heat treatment, the surfaces of the slabs were preparedby chipping as shown in Figure 8. The fishtail appearance on the edges can beseen. The forging was preheated at 320 C for 29 hours and the edges weretrimmed by torch cutting which resulted in a satisfactory forging. Aftertorch cutting, the forgings were stress relieved at 595 C.

Figure 8Chi in Surfaces of FFTF Slab

-15-

Initial Machinin and Ultrasonic Examination

LWBR Closure Head

The forging was transferred to the machine shop where the top face wasmachined to bright metal by removing a minimum amount of metal on a 4.35 mIngersoll milling machine. Some peripheral milling was performed tofacilitate setup. At this point, an ultrasonic examination was made with a2.25 Mhz ceramic transducer using a Sperry Model UR Reflectoscope. The gainlevel was established by setting up to a 190 mm sweep to peak (S/P) backreflection (BR) in an indication free area through the thickness as thereference standard. The purpose of this examination was to obtain the earliestpossible evaluation of the forging to assure the absence of an unacceptablecondition. As previously reported, the first LWBR closure head contained anunacceptable ultrasonic condition. The replacement forging was consideredsatisfactory and released for completion of the machining preparatory to theofficial preliminary ultrasonic examination prior to heat treatment formechanical properties.

Since twelve penetrations or holes were to be subsequently trepanned inthe forging, the importance of the preliminary examination is obvious. Ifthere were isolated unacceptable conditions in the head forging, it would belikely that these could be located in a penetration by careful layout of thefinished closure head in the forging.

The forging was machined on the opposite face on the 4.35 m Ingersollmilling machine to 1310 mm thick and the diameter was turned to 3950 mm on a7.6 m vertical boring mill. The surface finish was equal to or better than6.3 micron to satisfy the procedural requirements for the ultrasonicexamination.

The official preliminary ultrasonic examination was performed inaccordance with the approved procedure using a Sperry Model UR Reflectoscopewith a 2.25 Mhz Quartz transducer for the axial examination and a 1.0 Mhzceramic transducer for the radial examination. The 1.0 Mhz transducer wasused for the radial examination due to the inability to adequately penetrateat 2-1/4 Mhz because of a structure pattern. Such a structure condition isnot uncommon on massive forgings particularly when examining them prior tothe quench and temper operation.

The forging was examined through the top and bottom faces and theperiphery after establishing a distance amplitude correction (DAC) curve froma 8.75 mm diameter flat bottom hole machined in acoustically compatiblereference blocks. Any discontinuity causing an indication equal to orexceeding 50% of the DAC reference curve was to be reported. However, onexamination, no reportable indications were noted in the closure headforging.

-16-

Nine of the twelve penetrations were trepanned on a 254 mm horizontalboring mill. The three remaining penetrations were to be trepanned followingthe quench and temper operation to enable determination of deep-seatedmechanical properties by testing the core bars.

FFTF Closure Head and H drostaticTest Head

The forgings were transferred to the machine shop where the top face ofeach was machined to bright metal on a 4.35 m Ingersoll milling machine.Some end milling of the edges was performed to facilitate setup. At thispoint, a preliminary ultrasonic examination was made with a 1.0 Mhz ceramictransducer using a Sperry Model UR Reflectoscope. The gain level wasestablished by setting up to a 190 mm S/PBR through the thickness as areference standard. All of the forgings were considered satisfactory andreleased for completion of the machining operation prior to the officialultrasonic examination before heat treatment for mechanical properties.

All of the forgings were machined on the 4.35 m Ingersoll milling machineto 605 mm thick x 3885 mm wide x 7765 mm long (Figure 9). This thicknessprovided 19 mm excess metal, per surface, over the finished size. The surfacefinish was equal to or better than 6.3 micron to satisfy the requirements forthe ultrasonic examination.

-17-

Figure 9MachininFFTF Slab on 4.35 m In ersoll Millin Machine

Since the forgings were to be subsequently torch cut to a semicircularcontour with several major cutouts in the closure head, the importance ofthis ultrasonic examination is obvious. If there should be an unacceptablecondition in the rectangular pieces, possibly the condition could be removedby careful layout of the finished closure head within the rectangular forging.

The ultrasonic examination was performed with a Sperry Model URReflectoscope using a 1.0 Mhz transducer (Figure 10). The 1.0 Mhz transducerwas used due to the inability to adequately penetrate at 2-1/4 Mhz because ofa structure pattern.

Figure 10Ultrasonic Examination of FFTF Slab

The ultrasonic procedure for the FFTF forgings required that theexamination be made by both the back reflection and flat bottom holetechniques.

The forgings were scanned on both the top and bottom faces afterestablishing a 38 mm S/P BR through the thickness as the reference standard.The top and bottom faces were then examined after establishing a DAC curvefrom a 12.7 mm diameter flat bottom hole as a reference standard. Finally,the longitudinal edges were scanned after establishing a 38 mm S/P BRthrough the width of the forgings.

-18-

Discontinuities causing indications which showed a back reflection lossof more than 20% and equal to or exceeding 100% of the DAC curve were to berecorded. No reportable indications were noted during the ultrasonicexamination of the halves of the closure head and hydrostatic test headforgings.

Torch Cuttin and Machinin to Contour

The FFTF forgings were laid out to size for quench and temper with 19 mmallowance over finished dimensions, except for designated test areas. Thesemicircular portion of the periphery was torch cut approximately 25 mm overfinished dimensions.

Prior to torch cutting the periphery and penetrations, the forgings werepreheated to 315 C and held for 37-1/2 hours. Cutting was performed at a speedof approximately 38 mm per minute (Figure 11) . There were no penetrations inthe hydrostatic test head. After torch cutting, the forgings were chargedinto a carbottom furnace, heated to 595 C, held for 23 hours and air cooled.

The forgings were machined to contour on the 4.35 m Ingersoll millingmachine and a 254 mm horizontal boring mill. The joint faces and theperiphery were milled, with the periphery being milled in a series of 24flats and the corners hand ground. The penetrations were machined on thehorizontal boring mill. These operations were performed prior to heattreatment for mechanical properties.

Figure 11Torch Cuttin FFTF Closure Head Half from Slab

-19-

Heat Treatment for Mechanical Pro erties

LWBR Closure Head

The LWBR closure head was transferred to the heat treatment departmentand charged into a gas-fired carbottam furnace. The forging was heated to855 C and held for 47 hours. The forging was then withdrawn from the furnaceand completely immersed in water which was being circulated at approximately34,000 liters per minute. The closure head was immersed in the quench foreight hours and then recharged into a carbottom furnace for equalizing at150 C. After 26 hours, the forging was heated to a temperature of 615 C andheld at this tempering temperature for 52 hours before being air cooled.

Mechanical properties were evaluated at the official peripheral testlocations located 51 mm x 152 mm from the nearest quenched surfaces.Additional testing was performed, using the three trepanned cores, whichwere removed from the penetrations. The location of the test areas isshown in Figure 12.

All test coupons were given a simulated stress relief treatment byheating to 580 C at a rate not exceeding 28 C per hour, held for 50 hours,and cooled to room temperature at a rate not exceeding 28 C per hour above320 C. After thermal cycling, the test specimelas were machined from thecoupons. Mechanical properties are listed in Table II.

FFTF Closure Head and H drostatic Test Head

The forgings were transferred to the treatment department and chargedinto gas-fired carbottom furnaces. They were heated to a temperature of855 C and held for 18 hours. The forgings were quenched to ambienttemperature in a spray quench unit which consisted of a stationary bottomgrid of spray outlets and a moving top grid. The unit measured 4.9 m wide x11.6 m long x 3 m high. The spray outlets on the bottom grid delivered8530 liters per minute and those in the top grid 9890 liters per minute fora total of 18,420 liters per minute. The closure head forgings were sprayquenched for 4-1/4 hours and the hydrostatic test heads were spray quenchedfor six hours. Upon completion of the quench, the forgings were rechargedinto a carbottom furnace, heated to 150 C and equalized for 12 hours.Subsequently, they were heated to the tempering temperature of 665 C,held for 36 hours and air cooled.

Mechanical property tests were removed from the locations shown inFigure 13.

-20-

Heat-treated

Finish-machined

r• v•V****•44::44/***WiN••••••..4••••••st, fra•-•A

Test Area

Figure 12Mechanical Pro ert Test Location LWBR Closure Head

-21-

Table II - Mechanical Pro erties LWBR Closure Head)

Tensile ResultsStrengths(kgf/mm2)

Yield(0.2%) Tensile Elongation Reduction

68.2 79.565.8 77.7

23.0 72.023.5 71.6

Im act Strength Transition Tem erature (5.2 k fm/cm2

Impact Strengths (kgfm/cm2)Test Temperature (Average of three test results)

(°C) Test Area 1 Test Area 2

-107 1.2 0.5- 84 2.8 1.5- 62 11.9 8.9- 40 14.7 11.3- 7 19.9 17.738 23.5 22.8

ISTT5.2 -77 C ISTT5.2 -65 C

20.119.121.216.718.416.8

Results Taken from Trepanned Cores

-22-

Impact ResultsTestin Performed at -29 CImpact LateralStrength Expansion(kgfm/cm2) Fibrosity (mm)

858695767474

2.01.92.11.61.81.7

Distance - from

Tensile ResultsStrengths(kgf/mm2)

to of for in )

Im act ResultsDistance Yield 7. Distance ISTT5.2

(sus) (0.2%) Tensile ElongationReduction (mm) Direction (°C)

25 65.4 75.6 23.0 71.4 100 Longitudinal -62152 66.1 76.7 22.0 70.8 100 Transverse -68330 66.8 78.1 22.0 70.3 610 Longitudinal -60660 67.9 80.2 23.0 69.7 610 Transverse -68980 68.2 80.2 23.0 71.4 1220 Longitudinal -601168 66.8 79.5 23.0 71.8 1220 Transverse -71

HYDROSTATIC TEST HEAD

*M ONr•••••••••••••••ab tail Test Area

FFTF CLOSURE HEAD

Figure 13Mechanical Pro ert Test Location

-23-

All test coupons were given a simulated stress relief cycle by heatingto 620 C at a rate not exceeding 55 C per hour, held for 40 hours and cooledto room temperature at a rate not exceeding 55 C per hour above 320 C. Afterthermal cycling, the official test specimens were machined from the coupons.The test depth of the specimens removed from the FFTF closure head was19 mm x 38 mm. The depth of the specimens from the hydrostatic test headwas 38 mm x 38 mm. Table III lists the mechanical properties. After passingthe required mechanical properties, additional test coupons containingdeep-seated (178 mm) V-Notch Charpy specimens were removed from the testareas. These coupons were thermal cycled as previously described and theresults are also shown in Table III.

3-1/2% NierMo Anal sis

A comparison of the mechanical property data in Tables II and III showsthat the 3-1/2% NiCrMo analysis provided increased impact energy levels at ahigher yield strength than the lean NiCrMo analysis. In addition, the 3-1/2%NiCrMo analysis exhibited uniform properties throughout the 1310 mm thicknessof the LWBR closure head. This grade offers outstanding potential for use inlarge closure head forgings where higher strength and improved deep-seatedproperties are required.

Since the 3-1/2% NiCrMo analysis is known to be susceptible to temperembrittlement, material was evaluated from a 4320 mm diameter x 680 mm thickupset forged closure head. After the quench and temper operation, mechanicalproperty tests which had not been given a simulated stress relief cycle showedan average impact strength of 21.2 kgfm/cm2 at -29 C. Material from the sametest location was given a simulated stress relief cycle of 72 hours at 595 Cfollowed by slow cooling 10 C per hour and showed an average impact strengthof 8.2 kgfm/cm2 at -29 C. These results indicated that a significant amountof embrittlement was present in the test material on slow cooling (10 C perhour) from the 595 C stress relief temperature. The same test material,when heated to 595 C for 1-1/2 hours and water quenched, showed an averageimpact strength of 25.8 kgfm/cm2 at -29 C.

In order to determine the cooling rate at which a minimum amount oftemper embrittlement would occur, i.e., the critical cooling rate, anexperiment was conducted. Specimens were heated to 595 C for 72 hours andfurnace cooled at rates of 17 C, 22 C, 28 C, and 55 C per hour. The impactspecimens from each cooling rate were then tested at -12 C and the resultsappear in Table IV. The data in Table IV shows the critical cooling rate tobe 28 C per hour.

If temper embrittlement has been induced in 3-1/2% NiCrMo material, itseffects can be erased by reheating to a specified temperature and then coolingat a rate faster than the critical cooling rate. To determine the minimumreheating temperature, embrittled samples were heated to a series oftemperatures and then cooled at a rate of 28 C per hour. Charpy V-Notchimpact specimens were tested at -12 C. with results shown in Table V.

-24-

Table III - Mechanical Pro erties (FFTF Closure Head and H drostatic Test Head

-25-

TensileStrengths(kgf/mm2)

Results Impact Results(Testin Performed at -12 C)Impact Lateral

Forging Yield % Strength % ExpansionNumber (0.2%), TensileElongationReduction(kgfm/cm2)Fibrosity (mm)

C 43.3 59.4 27.5 73.1 16.5 40 1.945.7 60.5 26.5 71.8 21.3 58 2.3

18.2 52 2.019.6 52 2.215.1 35 1.814.1 35 1.7

D 44.3 60.1 27.0 74.3 16.8 52 1.944.3 59.8 27.0 73.7 17.9 52 2.0

20.8 63 2.022.9 76 2.422.2 74 2.320.8 64 2.2

E 42.9 58.7 27.5 74.3 12.8 33 1.342.6 58.4 27.5 73.3 10.9 29 1.1

9.5 29 1.114.7 35 1.521.2 58 2.019.3 46 1.9

F 44.0 60.5 26.0 73.1 19.6 49 1.845.4 62.6 25.5 72.4 22.0 55 2.1

19.1 57 1.814.1 38 1.418.4 40 1.819.6 46 2.0

Dee -Seated Im act Results (Testin Performed at -12 C 178 mm Below Surface)

Impact Lateral Impact LateralForging Strength Expansion Forging Strength ExpansionNumber (kgfm/cm2) Fibrosity (wn) Number (kgfm/cm2) Fibrosity (mm)

8.3 23 1.0 12.3 29 1.212.0 35 1.4 12.7 29 0.811.6 29 1.4 10.9 29 1.117.7 46 1.9 5.7 27 0.72.6 21 0.4 10.2 33 1.214.6 35 1.7 10.1 29 1.0

I5TT5.2-54 C ISTT5.2-35 C

16.1 46 1.8 12.3 35 1.314.2 35 1.7 18.2 46 2.04.2 16 0.5 14.4 35 1.612.8 35 1.5 17.7 46 2.013.9 35 1.7 17.9 46 1.99.5 16 1.0 16.0 40 1.7

I5TT5.2-31 C I5TT5.2 -38 C

Table IV - Determination of Critical Cooling Ratefor Embrittlement of 3-1/2% NiCrMo

Cooling Rate Impact Strength at -12 C(°C per hour) k fm/cm2

17 6.922 13.728 16.055 16.0

Table V - Determination of MinimumReversible Embrittlement Temperature

Reheating Temperature Impact Strength at - 12 C°C

Finish Machinin and Final Ins ection

LWBR Closure Head

-26-

k fm/cm2)

Unembrittled Material 18.6455 13.5470 10.8480 10.2495 12.1510 10.8525 17.3540 19.1550 19.6565 19.6580 15.6

The data in Table V indicates that 525 C is the minimum temperaturerequired to erase previously induced embrittlement.

While the 3-1/2% NiCrMo analysis provides superior mechanical properties,an awareness of its susceptibility to temper embrittlement is necessary forproper usage of the grade.

The forging was machined to final drawing size, again on the 7.6 mvertical boring mill and the 254 mm horizontal boring mill. At this point,the forging was inspected for thickness, flatness, and surface finish.Magnetic particle examination was performed on all surfaces with 102 mm gridspacings using the continuous method with an amperage of 125 amps per 25 mm.The examination was performed in accordance with the ASME Boiler CodeSection III requirements. No reportable indications were found.

A final ultrasonic examination was performed following the same basicprocedure previously described. In this case, due to the refinement of thequench and temper operation, a 2-1/4 Mhz Quartz transducer was employed forthe axial examination. A 1.0 Mhz ceramic transducer was employed for theradial examination. The forging was scanned on both faces and all edges.No reportable indications were found.

The shipped weight of the LWBR closure head was 91,400 kg.

The forgings were machined to final drawing size, again on the 4.35 mIngersoll milling machine and the 254 mm horizontal boring mill. The topand bottom joint faces were milled as was the periphery. In this case, theperiphery was milled in a series of 48 flats. At this point, the forgingswere then laid out for final machining of penetrations and semicircularperiphery by locating centerlines across the width and length. The closurehead forgings were then placed on the 254 mm horizontal boring mill wherethe surfaces of the penetrations were machined to contour as shown inFigure 14.

FFTF Closure Head and H drostatic Test Head

Figure 14Machinin FFTF Closure Head Half on 254 mm Horizontal Borin Mill

-27-

The corners of the flats on the periphery were ground to blend theperiphery into a semicircle. Magnetic particle examination was performedon all surfaces using the same procedure as with the LWBR closure head.Figure 15 shows a closure head half being inspected for conformance to thedrawing. The 102 mm grid spacings used for the magnetic particle examinationcan also be seen. Figure 16 shows a hydrostatic test head half beingmagnetic particle examined. No reportable indications were found on anyof the forgings.

A final ultrasonic examination was performed utilizing the same basicprocedure previously described. At this time, due to the grain refinementof the quench and temper operation, a 2-1/4 Mhz Quartz transducer wasemployed. The forgings were scanned on both faces and all edges and noreportable indications were found. (See Figure 17.)

The shipped weight of each of the FFTF closure head halves was78,300 kg. The hydrostatic test head halves weighed 102,000 kg each.

Figure 15Ins ection of FFTF Closure Head Half

-28-

Figure 16Ma netic Particle Ins ection of FFTF H drostatic Test Head

Figure 17Final Ultrasonic Examination of FFTF H drostatic Test Head

-29-

Fabrication of the FFTF Closure Head

Welding of the closure head forging halves together to make the 560 mmthick, 7670 mm diameter flat head was performed by CE at Chattanooga,Tennessee. A pressure weld of this thickness and length is unique in thenuclear industry and CE had to determine the best method for welding,preheating, fixturing, and positioning to produce a weldment to morestringent requirements than the ASME Boiler Code Section III.

To keep the weldment from distorting into a 'V' type configuration,CE had to consider the type of weld groove (straight sidewall, single V orU, double V or U), and the location of the weld, if V or U type designswere used. If the normal submerged arc technique were used, specialfixturing and preheating equipment would have to be designed to allow forthe weldment to be turned from one side to the other as the weld progressed.

Because of the potential problems associated with submerged welding ofvery thick sections, it was decided to utilize the CE Sub-Vert method(contraction of submerged arc vertical progression)(3). CE had usedSub-Vert previously on light water nuclear reactors up to 305 mm thicksections. The Sub-Vert process simplifies the fixturing and preheatequipment design becausetheweld can be made from one side only withoutany turning and with a minimum of distortion.

The groove design was straight sidewalled with a nominal 32 mm groovewidth. Start and stop tabs increased the total as-deposited weld thicknessto 622 mm. A 150 C preheat was provided by placing gas-fired equipment onthe back side and electrical heaters on the front side (see Figures 18 and 19).

The welders were qualified by welding test blocks of the same thicknessas the head. The weldment was evaluated for mechanical properties andnondestructively examined using radiography and ultrasonics. Fitting up andpositioning the head, attaching stop and start tabs, along with the actualwelding, took approximately one month. Figure 20 illustrates the completedhead assembly.

After the welding was completed, the stop and start tabs were removed andthe weld was evaluated by magnetic particle examination. The head was thenplaced in a furnace for an intermediate post weld heat treatment (620 Cstress relief). When this treatment was completed, the weld was examinedradiographically and ultrasonically with acceptance standards morerestrictive than the ASME Boiler Code Section III. No unacceptableindications were noted by either nondestructive examination techniques.

The same welding procedure and equipment was used to fabricate thehydrostatic test head.

-30-

Figure 18Gas-Fired E ui ment and Electrical Heater Used for Weldin FFTF Closure Head

-31-

Figure 19Preheatin Furnace Used for Weldin FFTF Closure Head

DISCUSSION

Figure 20Com leted Weld FFTF Closure Head

Increasing world energy requirements will undoubtedly result in a growthin the size of nuclear reactors and their component parts. Close cooperationbetween designers and producers has made the task of arriving at the presentsize nuclear reactors feasible. As reactor sizes increase, the need forcooperation and communication will become even more pressing. The manufactureof large reactor components, such as closure heads, represents a great riskto the producer and close liaison with all of the involved parties aids inminimizing these risks.

Present reactor component sizes are approaching the physical limitationsof existing manufacturing equipment. Innovative forging techniques must beused to circumvent these limitations and assure acceptable quality.Modifications in heating practice, choices in die width selection,developments in intermediate forging configuration all offer avenues for thereduction of production risks.

The production of all large forgings is associated with high potentialhazards. The condition of the electric furnaces, mold and sinkheadpreparation, uniformity in heating, adequate forge work, and effective heattreatment to produce the best mechanical properties are just some of thephases in the manufacturing sequence that require detailed attention.

-32-

In the production of large generator rotors, a technique has beendeveloped, at Bethlehem, to increase the likelihood of success. Asreported by Smith and Hartman,(4) the maximum weight, 3300 mm diameter ingotis blocked to 2030 mm octagon full length and normalized and tempered. Theblock is then evaluated ultrasonically on a flame descaled surface. Thegenerator rotor can be carefully positioned in the block to minimize theeffect of any discontinuities present. If the heavier size generator rotorcannot be made from the block, application to a maller size generatorrotor is made.

Unfortunately, this technique cannot be used on large closure headforgings. First, with a 2540 mm diameter block, an ultrasonic evaluation ona flame descaled surface is difficult. Second, since the amount of reductionfrom the ingot is minimum, it would be expected to have many discontinuitiesin a block of this size. The consolidation is accomplished during theupsetting and spreading operation so the block evaluation would be almostworthless. Finally, the possibility of applying the block to anotherrequirement is negligible.

Since it is imperative that an early evaluation of the forging be made,the piece is ultrasonically examined immediately after forging and preliminaryheat treatment. After the quality level of the forging has been determined,the closure head is laid out in the forging so as to provide the optimumoverall quality after final machining. Early evaluation is necessary toinsure the promised delivery to the customer and to minimize processingcosts to Bethlehem.

Interwoven with the whole manufacturing sequence of large nuclearreactor components is quality control. From the monitoring of the scrapmetal charge in the electric furnace to the control of forging and heattreatment practices, and, finally, to the inspection of the finished product,quality control is paramount. The essence of an effective quality controlprogram may be found in an old Arabian proverb - "Think of the going-outbefore you enter.". If proper time and care is taken in planning all phasesof the production of a large nuclear reactor forging before the steel ismelted, then quality control falls right into place.

It may be worthwhile to list some of the trends that appear to be onthe horizon for forged nuclear reactor parts.

1. Lower amounts of phosphorous and copper will berequired in certain forged components of the reactor.Studies have shown that these elements enhanceradiation embrittlement and those components whichbecame irradiated during service will have to be madefrom steel with lower copper and phosphorous contents.To the producer, this means closer control of scrapand more effective electric furnace practices.

-33-

2. Increasing sizes of nuclear reactors will continueto burden existing manufacturing equipment. Shortof major capital investments, the producer mustcontend with increased size by innovating.

3. Superior mechanical properties will be required inthe design of future nuclear reactor components.The need for better deep-seated properties isbecoming evident. The test results presented inthis paper show that the 3-1/2% NiCrMo analysiscan provide excellent properties throughout theentire thickness of large nuclear reactorforgings.

-34-

Appendix I

Conversion Tables

-35-

97.0 113.0 23.093.5 110.5 23.5

Appendix II

Mechanical Pro erties (LWBR Closure Head

Tensile ResultsStrengths(ksi)

Yield(0.2%) Tensile Elongation Reduction

Test Temperature(°F)

72.071.6

-36-

Lnpact Results(Testin Performed at -20 FImpact Lateral

Strength 7. Expansion(ft-lb) Fibrosity (mils)

1161101229610697

858695767474

Impact Strength Transition Temperature30 ft-lb)

Impact Strengths (ft-lb)(Avera e of three test resultsTest Area 1 Test Area 2

-160 6.7 3.0-120 16.3 8.7- 80 68.7 51.3- 40 84.7 65.020 115.0 102.3100 135.3 131.3

ISTT30 -107 F I5TT30 -86 F

787584637066

Distance

Results(Distance

Tensile ResultsStrengths(ksi)

Taken from Trepanned Cores- from to of for in

Im act Results7. Distance I5TT30Yield

(Inches)(0.2%)TensileElongationReduction(Inches)Direction (°F)

2 93.0 107.5 23.0 71.4 4 Longitudinal-886 94.0 109.0 22.0 70.8 4 Transverse -9113 95.0 111.0 22.0 70.3 24 Longitudinal-7726 96.5 114.0 23.0 69.7 24 Transverse -9039 97.0 114.0 23.0 71.4 48 Longitudinal-7846 95.0 113.0 23.0 71.8 48 Transverse -96

Appendix III

Mechanical Pro erties (FFTF Closure Head and Hydrostatic Test Head)

Dee -Seated Impact Results (Testin Performed at +10 F -- 7" Below Surface)

Impact Lateral Impact LateralForging Strength % Expansion Forging Strength % ExpansionNumber (ft-lb) Fibrosity (mils) Number (ft-lb) Fibrosity (mils)

C 48 23 4069 35 5767 29 55102 46 7415 21 1484 35 68

I5TT30 -64 F

D 93 46 7082 35 6624 16 2074 35 6080 35 6655 16 41

ISTT30 -23F

-37-

71 29 4973 29 3263 29 4433 27 2659 33 4758 29 40

ISTT30 - 32 F

71 35 53105 46 7783 35 63102 46 77103 46 7592 40 66

I5TT30 -37 F

Appendix IV

Determination of Critical Cooling Ratefor Embrittlement of 3-1/2% NiCrMo

Cooling Rate Impact Strength at +10 F(°F per Hour) (ft-lb)

30 4040 7950 92100 92

Determination of MinimumReversible Embrittlement Tem erature

Reheating Temperature Impact Strength at +10 F(°F) (ft-lb)

Unembrittled Material 107850 78875 62900 59925 70950 62975 1001000 1101025 1131050 1131075 90

-38-

REFERENCES

(1) Juretzek, G. and Richter, G., "Study of the Process of Drawing DownHeavy Forgings With and Without Previous Upsetting of the Ingot",Translation, Kuznechn.-Shtamp, Proczvodstvo, No, 10, October 1964,pp 4-8.

(2) Reid, E. A., "Forging Methods and Tooling For Production of LargeForgings", presented at the International Forging Conference,September 1967, (Sheffield, England).

(3) Bunn, W. B. and Berger, C. F., "Reactor Welding With Bead AxisPerpendicular to Joint Axis", presented at the ASME Winter AnnualMeeting, November 1969, (Los Angeles, California).

(4) Smith, H. C. and Hartman, G. S., "Manufacture of Large GeneratorRotor Forgings Over 135 Metric Tons", presented at the InternationalForgemasters' Meeting, May 1970, (Terni, Italy).

ACKNOWLEDGMMENTS

The authors gratefully acknowledge the assistance and outstandingcontribution of M. J. Weinberg, Experimental Engineer, Bethlehem SteelCorporation in the preparation of this paper.

Thanks are also extended to N. S. Wamack, Project Supervisor,Combustion Engineering, Inc. in providing the information on the weldingof the closure head halves.

-39-

METALLURGICAL EVALUATION OF NUCLEAR HIGH PRESSURE ROTOR FORGINGS

ABSTRACT

Seven vacuum carbon deoxidized Ni-Cr-Mo-V high pressurenuclear rotor forgings weighing up to 54,400 kg (as shipped fromthe supplier) and measuring up to 1680 mm in diameter wereanalyzed for mechanical and physical properties and metallur-gical structure. Four of these seven rotor forgings were heattreated to obtain 56.3 kg/mm2 minimum yield strength and threewere heat treated to obtain 63.3 kg/mmi minimum yield strength.

Specimens were taken from the rotor surface (test prolong-ations), ends of rotor shafts and from the rotor center (borebars) in the axial and radial directions relative to the rotoraxis. The mechanical properties were determined by tensile,impact and rotating beam fatigue tests. Susceptability toembrittlement was determined by isothermal aging tests. Studieswere made of the homogeneity and isotropy of the bore core bars,shaft ends and the prolongs of the rotor by chemical analysis,microstructure observations, and mechanical testing in two orien-tations.

The chemical analysis and tensile and impact test resultsthroughout the rotor, i.e. the values from shaft ends and borebars, were treated statistically and compared with the ladlechemical results and the prolong impact and tensile resultswhich are the basis, for the rotor acceptance. The "t" testwas used as a basis for comparison.

Applying Barsom and Rolfe's correlation between the fracturetoughness parameter, Kw ' and charpy impact energy, the fracturetoughness of the rotors at room temperature was determined and usedto find the critical flaw sizes at a stress equal to the yieldstrength.

Based on the test results from the rotor center (bore core bars),the rotor surface (pro1ongations) and the rotor shaft ends it appearsthat the impact properties and tensile ductility are dependent uponthe radial distance from the rotor center and the specimens" orientation.The yield and tensile strengths and microstructure are dependent onthe radial distance from the rotor center, while the fatigue strengthis independent of location and orientation. Chemistry gradientswere found to exist through the rotor's cross section. The critical

1

crack dimension, as obtained from the fracture toughness equa-tion, is more than three orders of magnitude (at the appliedstress equal to the yield strength of the material) of the crackwhich can be detected by our ultrasonic inspection.

INTRODUCTION

The nuclear high pressure rotor forging, as shown in Figure1, is the central component of the nuclear steam turbine. It isone of the heaviest integral rotor forging weighing up to 55,000kilograms as shipped by the supplier.

A test program was commenced on seven Ni-Cr-Mo-V rotors,four made to 56.3 kg/mm2 yield strength, and three to 63.3 kg/mm

2

yield strength. The object of this program was twofold:

1. The determination of the uniformity of the metallurgicalproperties throughout the rotor.

2. The effect of orientation on the metallurgical prop-erties of the rotor.

Specimens were taken from the rotor surface (test prolonga-tions), ends of rotor shafts and from the rotor center (borecore bars) in the axial and radial orientations relative to therotor axis. The mechanical properties were determined by ten-sile, impact and rotating beam fatigue tests. Susceptibilityto embrittlement was determined by isothermal aging and sub-sequent Charpy impact tests. Studies were made of the homogen-eity and isotropy of the rotors by chemical analysis, microstructureobservations, and mechanical testing in two orientations of the borecore bars, prolongations and the shaft ends of the rotors.

The test results are treated statistically and the effectsof radial distance and orientations are determined by the "t"test, (1) i.e. the ratio of the difference between the meansdivided by the standard error of the difference.

Barsom and Rolfe's correlation(2)

KIc Y.S.

2 —= 5

Y.S.CVN Y.S.

20

where:

Y.S. = yield strength

CVN = Charpy V notch energy

is used to determine fracture toughness of the cross-section of therotor (bore core bars, prolongations and shaft ends) from charpyimpact results.

2

MANUFACTURING fiISTORY

The seven rotors were manufactured from electric furnace,vacuum carbon deoxidized NiCrMoV steel. The steel was tappedin the 1590 to 1600C temperature range, and the vacuum was main-tained at the pressure of one to two millimeters of mercury.

The ingot weight varied between 175,000 kg and 212,000 kg,while the corresponding forging weight of the seven rotors wasin the 74,000 kg to 93,000 kg range.

The forgings were austenitized in the 800 to 850 C range formore than 45 hours and either water quenched or water spray quenched.They were tempered at 600 to 630 C range for more than 45 hours.

CHEMISTRY

Neither supplier deliberately varied chemistry to achievethe different strength levels of the rotor. The differencein strength (56.3 kg/mm2 versus 63.3 kg/mm4) was achieved byvarying the heat treatment. The composition of these sevenrotors will be considered, therefore, to constitute samplesfrom one statistical group. Ladle analyses are shown in TableI (3), while Westinghouse's check analyses are shown in TableII (4).

The results from both ladle and forging chemical analyseswere shown to be normally distributed. The "t" test was used todetermine the degree of variation of chemical composition through-out the rotor.

Specifically, the following information was desired:

1. The degree of variation in chemical composition be-tween the prolong and the bore core bar results,and its effect on the mechanical properties of theprolong and the bore core bar.

2. The degree of variation between the ladle results,used as an acceptance criteria, and both the borecore bar and the prolong results. This informationwould facilitate the evaluation of the rotor as awhole in terms of chemical composition and itseffect on mechanical properties at the time of itsacceptance.

In order to obtain this information the following "t"test comparisons were made:

a. Bore core bar results with prolong results.

b. Ladle results with prolong results.

c. Ladle results with bore core bar results.

3

The results of these comparisons are presented in Table III.In order to evaluate the results of the "t" test comparisonstwo levels of significance are established. Since the variationsin carbon, phosphorus, sulfur and silicon are considered to bemore significant from the viewpoint of strength and toughness a95% level of significance is assigned to these analyses. A 97.5%level of significance is assigned to the othor constituents.

Based on these levels of significance the following con-clusions may be made:

a. There is a statistical difference in carbon, sulfurand chromium levels between prolong and bore corebar results. The difference in means are, however,small; 0.03% for carbon, 0.002% for sulfur and0.13% for chromium. Carbon, manganese, sulfur andchromium are generally higher in the bore corebars, whereas nickel is lower.

b. With the exception of silicon and manganese thereis a good agreement between ladle and prolongchemistries. The differences in means for thetwo exceptions are small: 0.02% for silicon,and 0.03% for manganese.

c. There is a statistical difference in the carbon,manganese, silicon and nickel levels between ladleresults and bore core bar results. The differencein means are, however, small: 0.03% for carbon,0.04% for manganese, 0.03% for silicon and 0.13%for nickel.

d. The rotor as a whole, independent of sample lo-cation, meets the specification requirements atmore than 99.9% probability level.

These chemistry differences will be later evaluated in termsof mechanical properties.

MICROSTRUCTURE

Since the determination of the homogeneity of the rotorforgings was one of the primary objects of this program, a seriesof microstructure studies were made of the rotors. The objectof these studies was to determine the differences in microstructure,if any, throughout the cross section of the rotor and to relate,if possible, these microstructure differences to the subsequentmechanical variations.

4

In light of the above, two approaches were taken:

a. A thorough microstructure study of two rotors.

b. A sampling of other rotors for the purpose ofdetermining the microstructure variations fromrotor to rotor.

Specimens for microstructure examination were taken fromtest prolongs and bore core bars. Examinations were made inboth radial and axial orientations. Rotors No. 3 and No. 4were examined thoroughly, that is: microstructure specimensfrom X-1, X-2 and X-3 prolongs and from B-1, B-2 and B-3 borecore bars were examined and the structure photographed at 100Xin both radial and axial orientations. The representativephotomicrographs (5) are shown in Figures 2A through 5. Theprevalent structure present consists of upper and lower bainite.There is some evidence in one rotor prolongation of Widmanstattentype structure as shown in Figures 2A and 2B. Summarizing thestructure, the following conclusions may be made:

a. The rotor is homogeneous in structure along itsaxis, as is evident from Figures 3A through 3C.

b. There is a definite relationship between micro-structure and the radial location, i.e. rotor sur-face (prolong structure) versus rotor center (borecore bar structure) , as can be seen when Figure 3Bis compared with Figure 5. Lower and upper bainitepredominate in the prolong (Figure 3B), whereas thebore core bar structure, as is shown in Figure 5, isessestially upper bainite.

c. At any specific location in the forging, the micro-structure is essentially independent of the orienta-tion, as indicated in Figures 3B and 4.

TENSILE CHARACTERISTICS

Presnetly, the mechanical properties, as derived from thetensile test, serve as one of the major design criteria ofturbine rotors. Since the mechanical aspects of the rotor ac-ceptance are based on the tensile test made on a specimen takenfrom the rotor prolongation, it was considered imperative toinvestigate the cross sectional mechanical properties of therotor and establish their relationship to the prolong mechanicalproperties.

5

Radial Bore Bars vsAxial Bore Bars

Radial Prolongs vsAxial Bore Bars

Radial Bore Bars vsAxial Bore Bars

% Probability that theTest Positions Differences are significant

T.S. Y.S. El

99.3 72.0 93.8

96.5 89.1 10.0

40.0 36.0 97.0

*In these "t" test comparisons the means were identical resulting,therefore, in 0 probability.

On the basis of Table IV and the "t" test comparisons, andusing 95% probability as our criteria, the following generali-zations are made:

a. The tensile and yield strengths of the rotor areindependent of orientation and dependent on theradial distance from the rotor axis. The differ-ences due to radial distance from the rotor axisare small, on the order of two or three ksi andwell within the specification acceptance criteria.

6

R.A.

99.6

65.0

98.8

b. The elongation and the reduction of area are de-pendent on both orientation and the radial distancefrom the rotor axis. The difference betwecn themeans are 2% in the case of elongation and 12% inthe case of the reduction of area.

c. All rotors met the specification acceptancecriteria independent of test location and specimenorientation.

The differences in mechanical properties between the pro-longations and the bore core bars may be attributed to thestatistical differences in chemistry and microstructure dif-ferences. These differences are the results of:

a. Temperature and composition gradients duringsolidification. Carbon and chromium are higherin the bore core bars, while nickel is higher inthe prolongs.

b. Variations in forging deformation.

IMPACT CHARACTERISTICS

c. Temperature gradients during heat treatment.

The toughness of the rotor, as measured by Charpy "V" notchimpact energy, is directly related to its resistance to cata-strophic brittle failure. Since the bore area of the rotor ex-periences the highest steady and transient stresses imposed uponthe turbine element, it is of major concern to us.

And since the toughness acceptance criteria is based on thecharpy "V" notch specimens obtained from the rotor prolongationsit was considered essential to determine the toughness of therotor throughout its cross section. Accordingly test specimenswere taken from the rotor surface (test prolongations), ends ofrotor shafts and from the rotor center (bore core bars) in theaxial and radial orientations relative to the rotor axis.

The data from these tests were treated statistically andare presented in Table V.

Using the "t" test the following comparisons are made:

For the 56.3 kg/mm2 rotors:

% Probability that the DifferencesTest Position are significant

Radial Prolongs vsRadial Bore Bars

FATT

>99.9

7

Kgm at R.T.

90.0

Examining the above relationships and analyzing statisti-cally the data from Table V the following conclusions may bemade:

a. The 56.3 kg/mm2 rotors met the FATT and the ab-sorbed CVN impact energy at room temperature re-quirementsnat 99.7% probability level, and the63.3 kg/mre rotors at 96% probability level, in-dependent of the specimen location or orientation.

b. The FATT and the impact energy are dependent onboth orientation and the radial distance from therotor center.

The temperature and the composition gradients during thesolidification of the ingot; the directionality associatedwith the forging; and the temperature gradients during heat treat-ing contributed to the chemical and structure gradients which,in their turn, affected the impact properties throughout thecross section of the rotor.

8

Test Position % Probability that the DifferencPsare significant

FATT Kgm at R.T.

Radial Prolongs vsRadial Shaft Ends 64.0 97.5

Radial Prolongs vsAxial Bore Bars >99.9 97.5

Radial Prolongs vsAxial Shaft Ends 99.9

Radial Bore Bars vsAxial Bore Bars >99.9 96.6

For the 63.3 kg/mm2 rotors:

Radial Prolongs vsRadial Bore Bars >99.9 >99.9

Radial Prolongs vsAxial Bore Bars )99.9 69.0

Radial Bore Bars vsAxial Bore Bars 65.0 94.0

Specifically, it is seen that the differences in chemistrybetween the prolong and the bore core bars (higher chromium,carbon and sulfur in the bore core bars and higher nickel inthe prolongs), as well as the differences in microstructure,influence the impact characteristics of the rotor.

FRACTURE MECHANICS

Since room temperature plain strain data are not availablefor Ni-Cr-Mo-V steel heat treated to the 56.3-70.3 yield strengthrange, the Charpy "V" notch impact energy at the upper shelf isused to produce, by way of Barsom and Rolfe relationship, atable of KIc for this rotor at room temperature. These valuesare shown in Table VI.

It is felt that the Harsom and Rolfe technique is realisticfor these reasons:

a. When the actual KIc data, which are available at-18C and below (6) are extrapolated conservatively,the results are approximately equal (within thenormal expected test scatter,) to the values ofKIC's as shown in Table VI

b. When the Barsom and Rolfe relationship is appliedat room temperature to Ni-Cr-Mo-V steel heattreated to a relatively high yield strength level(105-113 kg/mm2), or when it is applied to lowstrength Ni-Cr-Mo-V below -18C, the results cor-relate very well with the actual Kic data.

Taking the values of KIes from Table VI and making thefollowing assumptions:

a. A one to ten depth to length ratio surface crack,

b. An imposed stress equal to yield strength of therotor:

A table of critical cracks (acr) of the rotor istabulated from the equation:

KIc2 = 1.21 <2 (a/Q)

9

CRITICAL CRACK DIMENSIONS IN CENTIMETERS

Considering the radial bore bar results from 63.3 kg/mm2rotors to be both most severe and the most applicable in ourconcern of catastrophic rotor failure, at 95% probability levelwe are dealing with acr of 1.8cm by 18cm which produces a re-flective are of 32.4 square centimeters.

All discrete cathode ray tube reflections of 50% amplitudeor more are recorded, and correspond to the reflective are of.0097 square centimeters.

Taking the radial bar bore crack dimensions at two sigmalimits as the most critical (1.8 cm x 18 cm) the following areacomparisons are made:

1.8 cm x 18 cm = 3350. This shows that the crack area

necessary to cause a catastrophic failure of the rotor for a singleload cycle (under the conditions as outlined in this report) is3350 times larger than is detected and recorded by present ultra-sonic acceptance criteria.

FATIGUE CHARACTERISTICS

In order to determine the fatigue characteristics of therotor a series of fatigue tests were made. Twelve fatiguecurves (a total of 65 rotating beam specimens) were establishedfrom tests run at room temperature and 260C on both axial andradial specimens taken from the bore core bars and prolongation.

10

The specimens were taken from both 56.3 kg/mm2and 63.3 kg/mm

2

rotors. After the results were analyzed (7) it was evidentthat they constitute two distinct statistical groups by temp-erature. The fatigue limit appears to be independent of bothorientation and radial distance from the rotor center.

The following fatigue strengths are obtained (at 107 cycles):

Yield StrengthLevel

56.3 kg/mm2

56.3 kg/mm2

63.3 kg/mm2

Fatigue Strength kg/mmTemperature, C Mean Sigma

R.T. 39.1 2.02

260 34.2 2.81

R.T. 40.8

The ratios of the means of fatigue strength to the meansof tensile strength are:

a. For 56.3 kg/MM2 rotors at room temperature,

39.1 kg/mm20 - . 49878.5 kg/mm,.

b. For 56.3 kg/mm2 rotors at 260C, 34.2 = .52

65.8

c. For 63.3 kg/mm2 at room temperature,

40.8 kg/mm2 = .5180.4 kg75E2

EMBRITTLEMENT

Embrittlement characteristics of No. 1, 2, 6, and 7 rotorswere studied by both isothermal aging and step cooling, (8) andthe results are presented in Table VII.

11

2

An interesting phenomenon in No. 1 rotor is the relative changesin FATT between the X2 prolong and the B-2 bore core bar. In 12months of aging at 399C the X-2 prolong shifted 114C, whereas theB-2 bore bar shifted only 18C.

It is appropriate to bring out the fact that although theforging exhibits high degree of toughness, that toughness isonly beneficial as long as the high pressure element of the tur-bine is not operated in the embrittling temperature zone.

Presently, nuclear high pressure elements do not operate inthe embrittling temperature range.

CONCLUSIONS

a. The nuclear high pressure rotor forgings heattreated to either 56.3 kg/mm2 or 63.3 kg/mm2exhibit a high degree of structure and propertyuniformity. All rotors exhibited extremelyhigh fracture toughness at room temperature(as derived from the CVN energy) with the cal-culated critical crack area more than threeorders of magnitude higher than the areathat can be determined by ultrasonic inspectionprocedure.

b. The more than 55C difference in the FATT betweenthe prolong and the bore core bar results is sur-prising. This difference in FATT underscores theeffects of melting, forging and heat treating onthe impact characteristics of the rotor forging.The temperature and chemistry gradients duringsolidification, the amount of working during forg-ing, the inclusion density and distribution aswell as the temperature gradients during heattreating cause a shift in the impact energycurve. However, even with the above conditionsthe 56.3 kg/mm2 rotors met the FATT and theabsorbed CVN impact energy at room temperaturerequirements at 99.7% probability level, and the63.3 kg/mm2 rotors at 96% probability level, inde-pendent of the specimen location or orientation.Furthermore, the upper shelf energy remained thesame for both core bore bars and the prolongations.

c. All four rotors which were isothermally aged exhi-bited some degree of embrittlement. This is incon-sequential for the present operating temperatures,but it is definitely a condition to be taken seri-ously if and when higher operating temperatures arecontemplated.

d. The excellent properties obtained with these sevenrotors open design possibilities for larger diameter,more massive rotor forgings. The possibilities forlarger rotor forgings are limited only by the capa-bilities of the suppliers to produce ultrasonicallysound forgings to present turbine quality levels.

12

Bohdan Hasiuk

REFERENCES

1. Duckworth W. E., Statistical Techniques in TechnologicalResearch, Methuen & Co. Ltd., (1968).

2. Sarsom, J. M. and Rolfe, S. T., "Correlations Between

KIc and Charpy V Notch Test Results in the TransitionTemperature Range", ASTM Technical Publication 466,p 281-302.

3. Private communications

4. Materials Engineering Laboratories Reports: 8320, 8570; 8902;9875; 20,973; 20,975; 20,993; 20,994.

5. Materials Engineering Laboratory Reports: 9875, 20,973,20,975, 20,993, 20,994.

6. Greenberg, H. D., Wessel, E. T., Tryle, W. H., "FractureToughness of Turbine - Generator Rotor Forgings,Westinghouse Scientific Paper 68-1D4-MEMTL-P1,68-1D7-MEMTL-P1, April 5, 1968.

7. Material Engineering Laboratory Reports: HT 2165;HT 2167 (MEL 8813); HT 2168; HT 2358 (MEL 9682);HT 2384 (MEL 9875).

8. Private communications

13

TABLE I

SUPPLIERS" LADLE ANALYSIS (%)

Rotor

No.

C

Mn

P

S

Si

Ni

Cr

Mo

V

Sb

As

Sn

1 2 3 4 5 6 7

0.22

0.21

0.23

0.23

0.24

0.25

0.24

0.30

0.29

0.32

0.30

0.26

0.31

0.28

0.008

0.007

0.009

0.010

0.008

0.006

0.008

0.012

0.015

0.008

0.013

0.011

0.013

0.011

0.04

0.03

0.03

0.04

0.07

0.06

0.06

3.70

3.50

3.40

3.45

3.36

3.41

3.43

1.75

1.75

1.78

1.73

1.74

1.72

1.73

0.52

0.40

0.41

0.39

0.37

0.37

0.35

0.12

0.11

0.13

0.11

0.13

0.12

0.14

0.0012

0.0022

0.0007

0.0080

0.0008

0.0006

0.0007

0.005

0.008

TABLE II

Notes: 1. Xl, X2, X3; official acceptance test prolongations.

2. Bl, B2, 33; bore core bars from under the Xl, X2 and X3 prolong locations.

TABLE III

% PROBABILITY THAT THE DIFFERENCES ARE SIGNIFICANT

Notes: 1. Ladle results represent all seven rotors.

2. Bore core bar results are from eight points.

3. Prolong results are from nine points.

TEST POSITIONS

Mn

CHEMISTRY

P

S

Si

Ni

Cr

Mo

V1.

Bore Core Bars vs Prolongs

99.4

<0.1

<0.1

99.1

40.0

88.5

99.0

93.6

<0.1

2.

Ladle vs Prolongs

56.0

97.6

87.5

<0.1

95.5

16.0

75.0

44.0

81.0

3.

Ladle vs Bore Core Bars

99.5

99.0

(0.1

<0.1

99.7

97.8

56.0

<0.1

<0.1

TABLE IV

STATISTICAL ANALYSIS OF TENSILE TEST RESULTS

56.3 kg/mm2 Rotors 6.3 kg/mm2 RotorsTensile Specimen Specimen Number NumberProperty Location Orientation Tested Mean 6." Tested Mean .6-

Tensile Strength Prolongsksi (X1,X2,X3) Radial 14 77.1 1.23 9 82.1 1.72

Tensile Strengthksi Shaft Ends Radial 8 77.1 1.22 2 79.6

Tensile Strength Bore Barsksi (131,B2,B3) Radial 6 79.2 0.85 9 84.6 1.79

Tensile Strengthksi Shaft Ends Axial 8 77.6 1.12 2 80,4

Tensile Strength Bore Barsksi (b1,B2,B3) Axial 18 78.5 1.44 9 77.8 2.07

Yield Strength Prolongsksi (X1,X2,X3) Radial 14 64.1 2.18 9 70.4 1.90

Yield Strengthksi Shaft Ends Radial 8 65.2 1.62 2 67.9

Yield Strength Bore Barsksi (B1,B2,B3) Radial 6 65.4 0.98 9 71.3 1.29

Yield Strengthksi Shaft Ends Axial 8 64.4 3.02 2 68.3

Yield Strength Bore Barsksi (B1,B2,B3) Axial 18 65.4 1.62 9 71.7 2.11

Elongation Prolongs% (X1,X2,X3) Radial 14 22.6 1.1 9 22.0 0.8

ElongationShaft Ends Radial 8 20.6 1.5 2 22.0

Elongation Bore Bars(B1,B2,B3) Radial 6 21.2 1.4 9 19,8 2.3

ElongationShaft Ends Axial 8 23.1 1.1 2 22.0

ElongationBore Bars Axial 18 23.3 1.0 9 21.7 1.0

Reduction of ProlongsArea % (X1,X2,X3) Radial 14 68.2 2.6 9 67.8 1.7

Reduction ofArea % Shaft Ends Radial 8 59.4 7.8 2 64.2

Reduction of Bore BarsArea % (B1,B2,B3) Radial 6 59.7 4.3 9 54.5 10.9

Reduction ofArea % Shaft Ends Axial 8 72.2 0.8 2 70.6

Reduction of Bore BarsArea % (81,B2,B3) Axial 18 68.4 2.3 9 65,6 4.4

TABLE V

STATISTICAL ANALYSIS OF IMPACT RESULTS

56.3 Kg/mm2 ROTORS

FATT,C

kgm at FATT

No.

kgmat RT

Test Positions

Points

Mean

elr

Mean

Mean

C;

Prolongs

10

-100

19

6.2

1.4

13.7

1.7

Radial Orientation

Shaft Ends

Radial Orientation

8- 89

28

6.6

1.4

12.0

1.2

Shaft Ends

Axial Orientation

818.0

1.7

Bore Bars

Radial Orientation

6- 29

712.4

1.2

Bore Bars

Axial Orientation

18

- 39

88.7

1.4

15.2

1.7

63.3 Kg/mm2 ROTORS

Prolongs

Radial Orientation

6- 93

15

6.4

1.0

14.0

1.2

Shaft Ends

Radial Orientation

2-101

5.2

12.1

Bore Bars

Radial Orientation

- 22

18

6.6

1.5

10.6

1.8

Bore Bars

Axial Orientation

9- 30

13 --

9.1

1.4

12.8

8

TABLE VI

DERIVED KIcFROM CVN AT

M N

ROOM TEMPERATURE,

56.3 Kg/mm2

NO.

7T-772

ROTORS

LOCATION

ORIENTATION

POINTS

MEAN

MEAN+ 26

MEAN - 26'

Prolongs

Radial

10

229

264

191

Shaft Ends

Radial

8216

243

186

Bore Bars

Radial

6219

244

192

Shaft Ends

Axial

8267

304

230

Bore Bars

Axial

18

244

278

209

63.3 Kg/mm2

ROTORS

Prolongs

Radial

6247

281

218

Bore Bars

Radial

9218

256

180

Bore Bars

Axial

9237

294

180

TABLE VII

ISOTHERMAL EMBRITTLEMENT RESULTS

Aging

Aging

Rotor

Specimen

Specimen

Temperature,Time,

FATT

Kgm

Kgm

No.

Location

Orientation

CMonths

at FATT

at R.T.

1X-2

Radial

399

0

-68

6

16

7.2

8.2

12

44

10.0

6.6

1B-2

Radial

399

0

-23

13.0

6

-17

5.1

9.0

12

- 5

5.8

9.5

2X-2

Radial

399

0

-34

9.3

14.0

6

-11

8.0

11.0

12

10

6.1

8.0

6X-2

Radial

399

0

-50

6.9

13.0

6

-37

6.2

13.0

12

-43

6.9

13.0

6B-2

Radial

399

0 6

-28

6.6

12.0

12

-20

6.4

11.0

B-2

Radial

315

0

-23

5.0

12.0

6

-45

2.9

9.1

7B-2

Radial

399

0

-23

5.0

12.0

6

7

4.2

6.2

7B-2

Radial

454

0

-23

5.0

12.0

6

22

4.6

4.7

100

mm

2540

m

m10

0

127

229

mm

Dia.

470

mm

Dia

7620

ra

m

3556

m

m

rn

L L

- -J

L

-.J

914

mm

Dia

1674

mm

Dia

/16

74

mm

Dia

B-2

B

-1B

-3

Axial

X-3

X

-2

X-1

Figure

1 - Nuclear

high pressure

rotor forging.

127

mm

mm

— —

— —

— —

—

—

.7.7

:11

I I

I

I I

LiL.

_

305

itml_

sj

1524

m

m

law—

10

0 m

m

I12

7 m

m

100 ra

m

152

mm47

0 m

m Dia.

Radial

Figure 2A - 100XProlong Axial Orientation (No. 1 Rotor)

Figure 2B - 100XProlong Radial Orientation (No. 1 Rotor)

Figure 3A - 100XNo. 4 Rotor, X-1 prolongAxial Orientation

Figure 3B - 100XNo. 4 Rotor X-2 prolongAxial Orientation

Figure 3C - 100XNo. 4 Rotor X-3 prolongAxial Orientation

Figure 4 - 100X

No. 4 Rotor, X-2 prolong, radial orientation

Figure 5 - 100X

No. 4 Rotor, B-2 bore bar, radial orientation

CENTRO SPERIMENTALE METALLURGICO

R 0 M A

PROPOSED ULTRASONIC CLASSIFICATION OF FLAWS IN LARGE FORGINGS

F. Baldi (1) - G.Canella (1) - T.Tili (2)

ABSTRACT

The results obtained by ultrasonic examination, usually performed on the basis of inspection specifications, provide indic:itions that cannot readily be correlated with manufacturing vari-Ebles. This paper suggests a scheme for classifying all the fauitsrevealed in a forging by ultrasonic inspection.Through this scheme it is possible to derive fault indices which,taken together, provide a classification of a forging on the basis of the number, size and location of the flaws. From this —classification, statistical processing may furnish a possiblecorrelation between flaws in the forging and manufacturing variables, thus providing the quality control metallurgist with a very useful tool.

The classificat on scheme has been applied by Terni Steel-works (see Annex) to rotors over a period of twentynine months.From the results, histograms have been plotted for the faultindices in relation to the location - axial or radial - of rotorflaws and the period of manufacture.

(1) Centro Sperimentale Metallurgico - Rome(2) Terni Steelworks - Terni

1. - INTRODUCTION

For something like twenty years, nondestructive testing byultrasonics has been used systematically for ascertaining thequality levels of steel forgings.

Over this period the results obtained from this type of inspection have become increasingly more significant because of tFiebetter understanding that has been gained of the basic principlesof ultrasonics and because of improved ways of applying taese tothe testing of forgings.

Particularly worthy of note in this regard are the studiesmade at the General Electric Company's laboratories by Rankin andMoriarty (1), those at the Westinghouse laboratories by Renner,Greenberg and Clark (2) , those by international study groups suchas the Club Europeen des Gros Forgerons, and those by international agencies such as the Verein Deutsche Eisenhattenleute (3), —together with others by a number of authors, among whom one mayname Schinn and Wolff (4), Pignet and Kerversan (4), Liversidge,Fearn and Dodgson (9) , and Schiebold and Tietz (10).

These studies have led to the preparation of practical re-commendations (11, 12, 13, 14) and specifications for ultrasonicinspection of forgings. Though the specifications most general-ly followed leave some points still uncertain, they do howeverprovide a reasonable guarantee on the acceptability of the product.

Manufacturers, however, are not faced solely with the pro-blem of ensuring that their products meet the acceptance specifications, but also with the very basic one of exercising control—over quality during the various phases of manufacture, in orderto be able to ascertain with some degree of certainty why faultsoccur, so that these may be eliminated thus reducing rejects andthereby improving costs.

It was precisely with this objective in view that the authorsdecided to tackle the question of ultrasonic classification of inlarge forgings, since it seems that little has been done on thismatter to date.

A survey of available literature indicates that the few ef-forts that have been made in this respect concern the classification of flaws, but always for the purpose of inspection. Here,—reference is made particularly to the studies carried out by theVerein Deutsche Eisenhattenleute, jointly with the German Steel-founders' Association (3). This classification includes tenclasses concerning five parameters: type of flaw, size of flaw,frequency and distribution of flaws, distribution in the crosssection, distribution on the longitudinal section. The ten classes are associated with the seriousness of the flaws (higher fi7-gures generally indicate greater magnitude or greater danger).

The classification reported by Rauterkus was adopted bySchiebold (15) as a basis for application at VEB Schwermaschinenbau "Ernest Thalmann" of Magdeburg, in an attempt to standardize-ultrasonic inspection of forgings. The classification adoptedby Schiebold differs from Rauterkus's mainly in the number ofclasses (eight instead of ten) , but as no further information hasbeen published on this application, it may well be that the results

- 2

were not sa istactory.

For a flaw cla sifica 'fln_aimed exclusiyely_aj resolvingualit control problems„as is the case—de-all with here 'cer-ain charac eristics--6i the flaw sucli—as Its nature - iistribution and orientation, assume muc reater importance thanIn the case of ac e t n , ein hi hly significant asregards the metallurgical origin of the flaws tiemseifej:--The—

c ass' ' a ion at las een app ie prac lEariri7a -s--eTOlved precisely to single out and highlight these characteristics.

To attain this objective, it is of course necessary tohave an ultrasonic control method whose strict application ena-bles the results to be expressed in a concise, uniform manner,involving the very minimum of subjective interpretation.

2. - METHOD OF ULTRASONIC CONTROL

Though the object of control is different from that of in-spection, the problem may be approached in such a way that thedata required for classification may be derived from the normalmanufacturing control procedures, including final inspection,with some specific modifications to suit the purpose in question.Hence the test methods must be established only after carefulexamination of the inspection specifications followed in the workswhere it is wished to apply the classification.

The test methods must include instructions concerning bothfrequency and adjustment of the equipment. This may be done withthe help of reference blocks (in applying the classification atthe Terni Steelworks a block complying with ASTM E 127 - 64 wasutilized).

)1;In the case of rotors, sensitivity must be adjusted takingdue account of their diameter, for both bored and solid rotorsalike.

Regulations must also be laid down for radial and for axialexamination.

To determine the dimensions of a flaw set_Realpendicular tothe ultrasonic beam, accaunt_mube taken of its depth, by referei_i_c_e—te a ensatin curve relating the amplitude-75fl1T&—di-stontinuity signaland the distance of the flaw tromth6—TTansdu_cer.

In the case of flaws whose orientation is not perfectlyperpendicular to the ultrasonic beam, as evidenced by travellingindications through small shifts of the probe, the apparent di-mensions must be corrected by applying a factor which depends ontheir inclination; this is determined through detailed inspectionat three different frequencies.

The type of flaw (inclusions, porosity, cracks, etc.) isidentified by taking account of the morphological characteristicsand behaviour of the corresponding echo, and the metallurgicalhistory of the forging; the experience of the most expert opera-tors is also very useful in this respect.

- 3 -

2.1. - Classification ro osedThe classification scheme (appl ed in this case particularly

to rotors) is given in Fig. 1.

There are eleven parameters in the classification, each ofwhich has been allocated a letter of the alphabet. From four toseven flaw classes are associated with each parameter; there isalso Class 0 for flaw-free rotors. The classes have numbers, generally denoting the seriousness of the flaws detected (except —for parameters E, F and M, the higher the number, the greater themagnitude or the danger).

Following ultrasonic inspection and the application of theclassification table, every forging is assigned one of eleven indices, each consisting of a letter and one or more numbers.

The table is divided into two parts, the first concerns discontinuities located outside the axial zone (six parameters) and—the second (five parameters) relates to those in the axial or thecase of rotors of this type.

It was considered advisable to keep the indices for theaxial zone quite separate because, since some rotors have to bebored, from the metallurgical control aspect any flaws existingon the axial direction must be clearly highlighted.

In order to better define the discontinuities, it was decided to utilize two parameters (A and B for non-axial flaws, and—G and H for axial flaws) , the first of which takes account of theindications and the second of the interpretation. The numberof discontinuities (parameter C for non-axial flaws and I foraxial flaws) is defined within very accurate and quite narrow limits. In the case of non-axial defects, account must also be ta-ken of the volume in which they are distributed. Thus, the co7responding class is determined not only by the number of discontinuities but also by the weight of the casting in question. Thregweight intervals were established: less than 5 Mg (5t) (parameterC1); between 5 and 20 Mg (5 to 20 t) (parameter C2); over 20 Mg(20 t) (parameter C3).

The magnitudes of the discontinuities (parameter D for non-axial flaws, and L for axial flaws) are also defined within verycareful and quite restricted limits: 6 mm2 for class 1, between6 and 10 mm2 for class 2, between 10 and 15 mm2 for class 3, be-tween 15 and 32 mm2 for class 4, and over 32 mm2 for class 5.It was felt that further classes for flaws larger than 32 mm2could not be considered, since these would have corresponded toexcessively defective production.

The location of flaws is given by parameters E and F forthose lying outside the axial zone and parameter M for those inthe axial zone.

For the purpose of metallurgical control, the position offlaws must be fixed as accurately as possible, especially theirdistribution in the longitudinal section. For this reason thereare seven classes for parameters F and M, which refer to longitudinal distribution, while for the rest there are only five (foul:in the case of parameters A and G) . It should be noted that not

- 4

OU

TS

IDE

A

XIA

L Z

ON

E

Fig. la - Classification scheme - discontinuities

outside axial zone

C AXIAL ZONE

L DI SCON T I NU I TY I N TERPRETAT I ON NUMBER Of OF LCCAT I ONA INDICATIONS OF DISCONTINUITY DISCONTINUITI DISCONTINUITY OF DISCONTINUITYS (11.2)S

G H I L M

0 NO __,....INDICATION

INDIVIDUAL NECK OF ROTOR,1 INCLUSIONS 1 4 5 < 6FLAWS TOP SIDE

2 OF FLAWS

TRAVELLING3 INDICATIONS

4 CLOUD-FORM***CRACKS 214 30

> 15 1/3rd ROTOR BODYI NDI CAT I ONS •S 32 BOTTOM SIDE

5

7

CLUSTERSPOROSITY 6 4 10

ICLUSIONSTRINGERS"

11 20

RADIAL CLOUD-LIKEFLAWS

6 NOTE:

" - In case 113 the "LX" dimensions are in mm

In case 64 the °IX" numbers indicate the number of

el ouds„

Fig. lb - Classification scheme- discontinuities In axial zone

- 6 -

> 6$ 10

> 10S. 15

> 32

1/3rd ROTORBODY TOP SIDE

1/3rd ROTOR BODYMI COLE SIDE

NECK OF ROTOR,BOTTOM SIDE

ON ENTIRE BODYOF ROTOR

ON ENTIRE ROTOR

only are the varous parts of the rotor taken into consideration,but indications are givenin the case of each as to whether itis located in the top part (head) or bottom part of the originallingot. Naturally this presupposes that during the various phasesof manufacture the necessary markings are made.

- DISCUSSION AND CONCLUSIONS

The classification described provides the metallurgist witha means of following the quality of the product over the courseof time.

It is felt that its merits lie in the fact that certain features are singled out, highlighted and carefully classified.These are:- Type of flaw- Magnitude- Number of flaws compared with weight of forging- Location of flaws in forging, separating axial-zone defectsfrom non-axial-zone defects.

The classification may be used in a var ety of ways forquality control, depending on the main points of interest.

If it is wished to keep a particular type of flaw underobservation, once the parameters characterizing it have been identified, the corresponding classes are plotted on the ordinate of—a quality control chart and time on the abscissa.

If, instead, it is wished to highlight the difference inthe flaws occurring during two different periods of time, whendifferent production processes are in use, it is as well to plottwo diagrams, one for the first and the other for the second period,showing total flaws as a function of their location in the for-ging.

The results of a preliminary application of this kind to alimited number of forgings made at the Terni Steelworks is givenin the Annex.

As will be appreciated, the scheme put forward refers inparticular to the type of product and methods of manufacture pe-culiar to one company. However, it is considered that the sche-me can be used as a basis for other types of forgings.

- 7 -

- REFERENCES

1. - RANKIN A.W.,MORIARTY C.D. - Acceptance guides for ultraso-nic inspection of large rotor forgings. Transaction of theASME - Oct. 1956 - 1063/622.

2. - RENNER R.W., GREENBERG H.A., CLARK Jr W.G. - Ultrasonic andmetallurgical evaluation of flaws in large rotor forgings.Proc. 5th Int. Conf. on Nondestructive Test Canada, 1967,37/42.

3. - RAUTERKUS W., Vorschlag zur zahlenmassigen Kennzeichnungvon Ergebnissen der Ultraschallprdfung an SchmiedestdckenArchiv Eisenhdttenwesen vol. 8, 1963 601/604.

4. - SCHINN R., WOLFF U. Einige Ergebnisse der Uberschallprafungschwerer Schmiedestdcke mit dem Impulsecho - Verfahren.Stahl und Eisen Vol. 72 (1952) n. 12 5, Juni, 695/702.

5. - PIGNET J., KERVESAU (de) E. - Contreile ultrasoniscopiquedes grosses pi6ces forgées et moulée. Position américainevis-A-vis des clauses de réception dans ce domain.Revue Metallurgie LIV, n. 3 1957, 169/174

6. - MICHALSKI A., KRACHTER H. Kennzeichnung von Uberschallanzeigen an Stahlerzeugnissen.Archiv Eisenhilttenwesen, vol. 4, 1957, 213/222

7. - MICHALSKI F. - Détermination de la grandeur des d6fout dé-tectés aux ultrasons dans les pièces forgées.1969 - C.I.T. n. 12, 2695/2718.

8. - OPEL P., IVENS G. - Fehlergrössenermittlung mit Ultraschallan SchmiedestUcken.Archiv Eisenhattenwesen 33, n. 5, Mai 1962, 311/316

9. - LIVERSIDGE D.G., FEARN G.A. DODGSON M.W. - Ultrasonic asse-ment of unbored rotor forgings.

10. - SCHIEBOLD K., TIETZ H.D. - AVG Vorsatzkalen zur Fehlergrös-senabschtzung bei der Ultraschallprdfung.Neue Witte 15 Jg - Heft 9 - Sept. 1970 557/560.

11. - Chambre Syndicale de la Gross Forge Frangaise. Specifica-tion de recette aux ultra-sons des pi6ces de forge, CSGFFn. 2-66

12. - ASTM A 388-67 Recommended Practice for Ultrasonic Testingand Inspection of Heavy Steel Forgings

13. - ASTM A 418-64 Standard Method of Ultrasonics Testing andInspection of Turbine and Generator Steel Rotor Forging.

14. - BS 4124: Part. 1 : 1967Methods for Nondestructive Testing of Steel Forgings Part.1 : Ultrasonic flaw detection.

15. - SCHIEBOLD K. Beitrag zur definierten Fehlerklassifizierunghei der Ultraschallprilfung von SchmildestlickenNeue Hutte vol, 12, 1967,R. 11, 601/693.

- 8 -

-ANNEX

TERNI STEELWORKSTERNI

The classification has been applied by the NondestructiveTesting and Control Dipartment of Terni Steelworks for a periodof twenty-nine months. During that period 247 rotors were checked.Each of these was characterized by one of eleven indices of theclassification scheme and by the month of manufacture.

The data obtained were then processed and histograms wereplotted showing the trend of the defective area for each type ofdiscontinuity.

About eighteen months after application of the classifica-tion commenced, changes were made in the manufacturing parameters,and as these were not made simultaneously, two types of histogramhave been plotted.

The first type shows the axial orradial distribution of thedefective area of all rotors, for each type of flaw. The histo-grams are divided into two groups: those concerning rotors madein the first nineteen months (Period I) and those made in the following ten months (Period II).

In the second type of histogram, for each kind of flaw andits location, the defective area (normalized on a per-rotor basis)is plotted against the month of manufacture.

Axial and radial distribution of defective areas, for each t eof flaw

Figs. 2,3,4,5 and 6 contain the histograms of the firsttype. The abscissas correspond to the Ei and Fi (outside theaxial zone) and Mi (axial zone) boxes of the scheme in Fig.1,i.e. the zones into which the forgings were divided radially andaxially to facilitate localized analysis of flaws.

The ordinates represent the defective area AD for each lo-cation, obtained by multiplying the values of the pairs Ci.Di andIi.Li for each forging and then adding together all the forgingsin each of the two periods.

Since classes Ci, Di, Ii and Li are not represented in theclassification table by single values but by intervals (which,inthe case of the Ci values also vary with the weight of the for-ging) , to obtain the numerical valuation required to compile thehistograms, it was necessary to take an avarage value for each.The values used are reported in Table I.

From January 1968 to May 1970

- 1 -

The average values which emerge from all combinations ofproducts (Ci.Di) and (Ii.Li) of the classificationformthe elements of the determinant:

As there were 185 forgings in the first period and 62 inthe second, the values of the first period have been divided bythree so as to obtain comparable results.

In addition, the total defective value AT was calculatedon each histogram to permit comparison of total flaws in the twoperiods.

The AT values for the individual flaws for periods I andII are as follows:

Flaw AT (ram2 AT (mm2) II°

Inclusions 490 272

Porosity 91 0

Inclusion stringers 850 0.1

Inclusions on axis 62 47

Inc.string. on axis 253 125

* The factor 1,6 was introduced to simplify the table and thesubsequent calculations, but was eliminated in the histograms.

AD

200

100

AD

200

100

I period II period

1/16AC-14..

Fig. 2a - Axial distribut on of inclusions (B1)

Fig. 2b - Radial distribution of inclusions (B1)

I period II period

AD

50

25

AD

50

25

I period II period

• I I;ITI —I

Fig. 3a - Axial distribution of porosities (B2)

I period II period

(16W 'In If1(1'gC-r,Fig. 3b - Radial distribution of porosities (B2)

- 4 -

Ao I period II period

300-

200

100

Fig. 4a - Axial distribution of inclusion stringers (B3)

AD

400 I period

300

200

100

Fig. 4b - Radial distribution of inclusion stringers (B3)

- 5 -

I II

1 I I 1

WODKV1

II period

A i)

60

30

Fig. 5 - Distribution of inclusions on axis (HI)

Al)

60

30

I period

I period

-'

- 6 -

II period

II period

Fig. 6 - Distributionof inclusion stringers on axis (H3)

Defective areas as a function of the eriod of manufacture

The defectives trend, normalized on a per-rotor basis, isshown in the histograms in Figs. 7, 8, 9, 10, 11, 12, 13 and 14,for some kinds of flaws in given locations, by month of manufac-ture.

The AD (defective area) ordinates have been calculated onthe basis of the determinant set forth above, and then normalizedin the following manner.

Since the number of rotors manufactured in a month was notconstant, in order to make comparisons, the average monthly number n was calculated

ri 24729

- 7 -

8.5

and then the factor k was calculated from the ratio between Kand the actual number of rotors made in any month neff.

k - neff.

Hence the average monthly defective figure AD per rotor is

kAD TD -8.5

where AD is the actual monthly defective figure.

The principal aim of this paper is to report the defectosity classification system used for the influence evaluation of dif—ferent manufacturing variables.

Therefore, the processing results of a third period, inwhich remarkable further changes where made in the manufacturingvariables, are not related with here. We inform solely a remar-kable further improvement of the general defectosity until toget a minimum level of rejects.

A 1)

5 -7- - 10 15 20 25 29months

Fig. 7 - Trend of inclusions located near surface (Bl-E1)

-

A D

72

48

21

18

36

24

12

Fig. 8 -

102

r5 M - 13 20 25 29

monthsTrend of inclusions located near rotor axis (Bl-E4)

- 8 -

-

AD

Fig. 9 - Trend of inclusions distributed over entire thickness(Bl-ES)

72

48

24

AD

28

21

14

cl1155

15

15

- 9 -

-20

1120

25 29months

20 25 29

monthsFig. 10 - Trend of inclusions located on rotor neck, bottom side

(B1-F5)

AD

60

:i6

21

12

I 0 15 20 25 29months

Fig. 11 - Trend of inclusions distributed over ent re body ofrotor (Bl-F6)

A114

L

In I

F g. 12 - Trend of inclusions located over

- 10 -

190 r25

monthsentire rotor (Bl-F7)

An173

1.10

105

70

.15

r I I TT2 0 " " '25' "29'5 111

monthsFig. 13 - Trend of inclusion stringers located on rotor neck,bottom

side (B3-F5)

28 -

21 -

Fig. 14 - Trend of inclusions distributed over entire body ofrotor (H1-M6) axial zone

15 20 25months

REMARKABLE CASES OF DEFECTS IN THE MANUFACTURE OF LARGE STEEL FORGINGS

3.1) Remarks

4. INNER NITRIDE DEFECTS

Foreword

1. NON METALLIC INCLUSIONS

1.1) Micrographic aspects, dimensions, chemical composition and locationin the ingot

1.2) Origin and formation model

1.3) Possibility of reducing the presence of slag inclusions

1.4) Separation of deoxidation products from the bath

1.5) Remarks

2. INNER FORCING DEFECTS

2.1) Plastic deformation in stretching forging

2.2) Experience data

2.3) Quality checks

2.4) Interpretation of experiment results

2.4.1) Distribution of deformation

2.4.2) Penetration depth

2.4.3) Shape and dimensions of dies

2.4.4) Effect of temperature

2.4.5) Capability of open die forging presses

2.5) Remarks

3, EXTERNAL FAILURES CAUSED BY FORGING AND HEAT TREATMENT

5. ORGANIZATION OF THE METALLURGY AND QUALITY CONTROL DEPARTMENT IN "TERNI"STEEL COMPANY

U.S. = Ultrasonics

MIT = Millimetres

ppm = Parts per million

V = Diameter

= Tons

MN = Mega Newton

HV = Vickers Hardness Values

S.A. = Without etching

= Yield point

= Tensile strength

A = Elongation %

= Reduction of area

ABBREVIATIONS

ToC = Temperature (degrees centigrades)

= Forging reduction %

= Deformation on the axial zone

= Distance from the axis of the forging

= Width of dies

r.l.f.= Slow cooling in the furnace

REMARKABLE CASES OF DEFECTS IN THE MANUFACTURE OF LARCE STEEL FORGINGS

T. Salinetti, T. Tili, M. Massarini

Society Terni, Italy

In this paper some of the most serious defects still leading to the rejec-tion of large steel forgings for the electromechanical industry are reviewed.They are defects of the following types: non metallic inclusions, axial poros-ities, nitrides, forging and heat treatment cracks. Besides, techniques andmanufacturing processes having greater influence on defect formation are dealtwith. Also an outline of Quality Control organization is given, for we believeit has an important function in the modern manufacture of large forgings.

Foreword

At least 25 years have now passed since the use of ultrasonic examinationwas applied to investigation of defects in large steel forgings. This method,which is always being improved has made it possible a remarkable qualitativeprogress of such products also with the help of more adequate forging and heattreatment techniques and processes, vacuum degassing of steel, reduction of im-purities and non-metallic inclusions.

As everyone knows, inspection procedures for large rotor forgings normallyinclude visual examination, magnetic particle control of the whole surface andhigh sensitivity ultrasonic examination of the whole forging. If no anomaliesare found with the above tests there are no problems. Yet, at times, some forg-ings may clearly show anomalies at ultrasonic or magnetic particle tests. Inthese years many of such anomalies were submitted to particular investigations.On the ground of the results of such examinations and of an exact evaluation ofservice stresses, designers and metallurgists, because of lack of exact accept-ance standards, expressed in the past their own personal judgement on the des-tination of these forgings.

These people were necessarily extremely careful in judging, which has ledto more rejections than was necessary. In the manufacture of large forgings,existing a direct relationship between manufacturer and buyer, it is possiblefor the former to know the requirements of the latter and consequently to makeuse of all suitable resources just from the designing stage. The quality may bedistinguished into "designed quality" and "technical quality obtained"; they areconnected with the two essential production stages; i.e. design and manufacture.They both subsist for current production and single pieces.

In the design stage the qualitative properties of the product are defined,this should lead to a theoretical quality we could define as "degree of satis-faction" that the properties have in comparison with the customer's requirements.

The quality in the manufacture stage may be found in the "degree ofconformity" of the product and therefore of its properties compared with theones fixed in the design stage.

All industrial steels show some defects; they may create sources ofstress concentrations which can sometimes cause brittle failures of parts ofsteel even perfectly designed in accordance with the specifications in use.This is so true that a clearer and clearer definition of the acceptancelimits for such a defectivity is required. For reaching this, as a firstphase, an ultrasonic examination technique suitable to define the defectivityof the forging with the best reliability was set up by various specialists.

As regards this subject we mention "The reference Block Technique",which according to our direct experience, considering all the parametersexamined, together with the position and orientation of the anomalies, showsa good conformity between the dimension of the defect evaluated by ultra-sonics and the actual ones. Later on and in these years there has been adevelopment into very interesting dynamical tests aiming at measuring thefracture toughness degree of steel, taking the presence of a certain defectiv-ity for granted. It is plain that the most feared type of failure, which canlead to real disasters, is the "brittle" one. Charpy V impact tests wereinserted in the specifications and the principles of transition temperaturewere considered in the design.

As it is well known the conditions for the brittle failure of a materialare three:

a) an agent, resulting from design or manufacture defects, generating over-stresses;

b) a local yielding occurred near the overstress;

c) the material which is below the transition temperature from ductile-brittle.

"Pellini" (1) refers to three critical temperatures the effect of temper-ature on the behaviour of steel:

NDT - Nil Ductility Transition.FTE - Fracture Transition for Elastic Loading.

At this temperature, which is higher than NDT, plastic deformation isnecessary in order to start breaking.

FTP - Fracture Transition for Plastic Loading.Above this temperature, which is higher than FTE, brittle failurescannot grow even through a material with serious plastic deformationscaused by high overstresses.

"Pellini" (1) has found that FTE falls approximately 100C I 60C over NDT,FTP 38°C ± 110C over NDT.

The definition CAT "Crack arrest temperature" of an acceptable orunacceptable defect noted by ultrasonics now results improvable thanks to thetechnology of linear elastic fracture mechanics which has rapidly developedand is now so advanced as to offer a promise of more exact methods to

establish acceptance standards for large forgings. The essence of the methodlies in comparing the field of stresses developed near the tip of a crack typedefect, with the rated stress applied to the structure, with the propertiesof the material and with the crack dimension necessary to cause breakdown.Such a criterion is based on the rapid growth of the crack when the stresseeat its tip exceed a certain critical value to which corresponds a criticalvalue Kc of the stress intensity factor. Such Kc results then to be anintrinsic property of the material and defines its breaking strength in thepresence of a crack type defect. Any combination of applied load, structuralconfiguration, geometry and dimension of the crack resulting in a K valueequal to or higher than Kc will cause the rapid failure of the structure.Yet, while in a structure submitted to a constant stress a safety criterionagainst failure may be based on the limit dimension of a possible existingdefect, i.e. such that K value is always lower than Kc, in presence ofalternate loads it is necessary to introduce the concept of working life ofthe structure, depending on the growing rate of the crack from a sub criticalto a critical dimension creating rapid failure. Studies on the growing rate CLNof a fatigue crack on rotors such as fracture mechanics, have been recentlyundertaken by several researchers. This is certainly a remarkable step for-ward for a better definition of inspection specifications for forgings but wewould say it is not all.

If we want to follow the principle according to which a product must notbe "as good as possible", but above all "as good as necessary" also to have aminimum production cost, it would be extremely useful not only to undertakeresearches on fracture mechanics, presuming the worst defect such as cracks,but presuming the presence of various types of defects which can really occur(non metallic inclusions, porosities, grain size etc.) about which we all know,though from tests of an indicative nature, how their influence on fatigue fail-ures is certainly less serious than the one of crack type defects.

If it is true that knowing the exact nature of an ultrasonic indicationis not always possible, experience has certainly shown that by making use ofall the forgemaster's information concerning manufacture and checks, it isalmost always possible to have reliable data on the kind of defects remarkedby ultrasonics. In order to achieve a more rational utilization, anotherimportant parameter the designer may supply the metallurgist with, is made upnot only by the maximum acceptable equivalent dimensions of a defect detectedby ultrasonics, but how these may increase as function of a better toughnessof the material.

After what we have just said, we think it is extremely useful to supplyinformation on nature and origin of some kinds of defects in special alloysteel forgings.

It is a matter of defects which can still lead to an internal rejection inforging shops or by customers, or some other defects now already eliminated.Such survey of ours has the double purpose of supplying designers with betterinformation on the nature of some ultrasonic anomalies and to frankly exchangeour experience with the one of other specialists and firms. Though looking

with the utmost interest at the future developments in the manufacture oflarge forgings of high quality level with the use of ingots produced by theESR process, it seems to us also important to continue a deeper and deeperanalysis of the present manufacture techniques and processes. We do notintend to illustrate the whole series of defects which can occur during themanufacture of a forging and which will obviously vary in time and from firmto firm, but only the ones concerning our more recent past experience in themanufacture of large steel forgings. Some of them are more frequent andtherefore they are cause of worries for rejections or low yield of ingots,some others are less frequent but they are still important because of theirseriousness.

1. NON METALLIC INCLUSIONS

Apart from the acceptability or not at the final ultrasonic inspec-tion, everybody has now realized, as recently R. Kiessling (4) said,that non metallic inclusions exist as "inseparahle components of steel".

In these years, particularly after the elimination of defects dueto gas by means of vacuum degassing, non metallic inclusions, almostinevitably, have become the main subject attracting the metallurgist'sattention. Many researchers undertook the study from the pure metal-lurgic point of view and the results published up to now are of consid-erable value. What we are going to review here makes up our experiencefor the limitation of the content of these heterogeneities to valuesacceptable at ultrasonic inspection on forgings and to the individuationof when inclusions have a very slight influence on strength propertiesof the material:

1.1) Micro ra hic as ects, dimensions, chemical com osition and locationin the ingot

On the ground of the results of the ultrasonic examination onabout 130 alloy-steel rotor forgings weighing from 15 to 100 tabout, we have come, considering dimensional variations and discardscut away during the hot plastic working, from the location of ultra-sonic anomalies found in the forgings to the respective one in theoriginal ingots. We point out that the ingots had bottom discardsfrom 15 to 20% and top discards of 3% of the weight of the ingotbody. Fig. 1 shows the results of the ultrasonic examination of aforged billet.

We show the locations of the anomalies remarked, all judged asderiving from non metallic inclusions, and the position of two corestaken for ascertaining the exact nature of the defects and their in-fluence on the strength properties of the material.

Figures 2 and 3 show macro and micrographic aspects which aretypical of the type of inclusions found in forgings, and the valuesof the mechanical properties obtained. On the ingot outline ofFig. 4 the more recurrent positions of the above anomalies areproportionally shown. Some specimens were sent to the "Centro

Sperimentale Metallurgico" of Rome for a qualitative analysis withmicroprobe. Dividing the zones of the ingot according to solidifi-cation isothermal lines, as shown on the diagram of Fig. 4, theanomalies found at the ultrasonic examination may be distinguishedinto two groups:

A) Inclusions with basis of CaO,Si02,Al203, Mg0 sometimes contain-ing Ti and V with dimensidns varying from 0.1 mm to 0.5 mm indiameter for a length which sometimes reaches 50 mm in thezone which is normallyllinterested by sulfides (150 to 300 mmabout from the periiery of the ingot).4 \

fB) Inclusions with baSis of MnO,Al203 and Si02 generally with re-

markable and very extended diTensions found at a little depthat the ingot top near the riser and in the central part of thebottom.

1.2) Ori in and formation model

Inclusions of type A) may be likely considered as coming fromparticles of furnace slag fallen into the ingot mold and imprisonedin the ingot during solidification, while inclusions of type B)may be of exogenous origin and due to oxidation of steel duringpouring in air-furnace-ladel-pony ladle. They are also caused bythe washing of the refractory material of the ladle and pony.ladleby liquid steel during transfers, they are afterwards kept by theoxidized veil almost always existing in the pouring "front" when theladle is being filled. The origin of the inclusions of group A)due to furnace slag, has brought to completely re-discuss theadvantage or disadvantage of continuing the practice of tapping thesteel mixed up as far as it is possible with slag so as to facili-tate a more strict contact between slag and steel in order to havea further reduction in S and P content. Nevertheless before givingup the practice of mixing steel and slag at tapping, we undertooka control of the inclusions*by means of microscopic examination ofstandard specimens taken at different stages of pouring from thefurnace up to the'pony-ladle. Such control has first of all made itnecessary the setting up of a standard for the classification of thevarious types of possible inclusions, a procedure for taking thespecimens and a micrographic evaluation method considering not onlyaspects and nature of inclusions, but also their dimensions andquantity.

The diagram of Fig. 5 shows the different stages and takinglocations of specimens including the ones taken from the furnacejust before tapping.

Fig. 6 shows the main data of the standard we adopted for theclassification of inclusions together with the dimensions of thecast specimens employed for the control of inclusions.

Apart from type of inclusions and controls performed, itclearly appeared that the average inclusion content with regard to

the specimens taken from the furnace, increased in the ladle andslightly decreased in the pony ladle though it remained alwayshigher than it was in the specimens taken from the furnace.

All this, besides, seemed to be in contrast with Stockes'slaw concerning the ascent rate of non metallic inclusions in theliquid steel.

As a matter of fact we think the above law is still valid butwe must consider what happens in practide.

The liquid steel in the ladle in fact, is not all at the sametemperature also in the best case when an effective pre-heating ofthe ladle is performed. The steel coming into contact with theladle bottom and walls will undergo a certain loss of temperaturereducing the ascent rate of inclusions: moreover the colder metal,because of it increase in weight will descend towards the bottomcausing rabbling and increasing the inclusion content in the metalnear the ladle or pony-ladle bottom. Such a situation, as experienceproves, grows in the ingot mold. Therefore presuming that the mainfunction of the mold is above all the one of allowing the metal topass from the liquid to the solid state, and that ladles are an un-avoidable means to permit the transfer of the metal, one of the wayswhich seems to more effectively reduce the inclusion content is theone of limiting as much as possible the rabbling of furnace slagand steel at tapping, realizing preheating and cleanness conditionsof ladle and pony ladle so as to make the ascent of other unavoid-able inclusions easier. As far as inclusions of group "8" areconcerned, their elimination at the origin is somewhat more diffi-cult, while their elimination when the ingot is poured or when itundergoes its successive transformation stages has appeared com-paratively easier.

They are non metallic inclusions of predominantly exogenousnature; perhaps also some of endogenous nature according to theoxidation degree of the steel pouredintOthe ladle and to thequality of its refractory are also added.

The refractory is more or less corroded according to quantityand types of oxides present. The strong etching by Fe and Mnoxides is well known. This kind of inclusions is not new: it wasminutely described by F. Hartmann since 1930. They are non metal-lic inclusions the chemical composition of which, already found byseveral metallurgists, generally varies within these values:

Si02 A1203 Fe0 Mn0

50/60% 5/20% 5/25% 10/28%

From some researches carried out it clearly appeared thatbecause of a mechanic action due to the movement caused by the pour-ing jet, decantation etc. when the metal is poured into the ingotmold these inclusions group more or less in large quantities by thewalls of the ingot-mold itself. They are visible as black slag whenthe ingot mold is being filled. It has been noted that their dis-

tribution with regard to the longitudinal section of the ingot ison an average the one shown on the diagram of Fig. 7. When the levelof the liquid steel approaches the top of the ingot mold they groupbecause of the pouring jet and are pushed towards the ingot-moldwalls. As soon as the steel reaches the riser step, the inclusionsare kept there according to the diagram of Fig. 8. Of course thenon uniform distribution of said inclusions, with regard to thecross section of the ingot is due to centering and compactness of thestream during pouring into the ingot mold.

1.3) Possibilit of reducin the resence of sla inclusions

On the ground of several investigations concerning number andtype of inclusions present in the various stage of the heat fromsteel melting in the furnace, tapping and pouring into the ladle, aseries of provisions concerning aboveall tapping temperature and theseparation of deoxidation products from the liquid steel were taken.The production of steel in basic electric furnaces with the doubleslag process includes, as everyone knows, the oxidation stage start-ing from a carbon content of 0.65 to 0.85% followed by the reducingstage or refining performed only with basis of limestone, calciumsilicide and silicon when the steel is not vacuum carbon deoxidized.No alluminium is added.

The knowledge of the oxidation degree of the steel bath.by meansof the determination of 02 content is certainly a valuable helpbecause it is possible to more exactly determine the quantities ofdeoxiding elements necessary to reach the minimum 02 contents insteel and consequently to obtain the best yields from deoxidingalloys added. Another indirect way of knowing the 02 contentpresent in the bath is that of measuring the quantity of 02 blown inor added as mineral in the oxidation stage.

As some steelmakers have pointed out, the yield of deoxidatingadditions depends not only on 02 content, but also on externalfactors, such as possible reactions with the refractory material ofthe ladle and oxidation caused by air coming into contact with thesteel during tapping and pouring into the ladle. Opinions onelements and modalities to be used for a better steel deoxidationare contrasting.

Of course the most discussed alloy elements are Si,A1,Mncalcium silicide and iron alloys of Si and Mn. Nevertheless ourpractice foresees as already touched upon, deoxidation in furnacewith Si or Si-Ca or vacuum deoxidation. Fig. 9 shows the behaviourof 02 (ppm) content as function of the quantity of deoxidizer (%)added. The behaviour of the two substances is analogous, thoughthere is a slight advantage for Si-Ca with additions superior to0.4%. The quantity of 02 remaining in the steel is in any caselower than 80 ppm and above 0.5% it is almost constant. The deoxi-dation rate of silicon is increased if higher quantities of Mn areadded. This results from Fig. 10 in which are shown the percentagesof 02 removed from the bath after 30 seconds from the addition of

deoxidisers as function of different quantities of Mn (continuous-line curve).

Experiments carried out on the deoxidating effects using Alalone or combined with Mn, show that as the quantity of Mn increasesthe deoxidation rate with alluminium also increases remarkably.

1.4) Se aration of deoxidation roducts from the bath

The ascent of deoxidation products to the surface is regulatedby two different physical laws: from one side by the mutual actionbetween strength of the upwards push of oxidation products and bathviscosity; from the other side, by the surface tension betweendeoxidation products and melted material.

The strength of the upwards push of an oxide particle isdetermined by the difference in specific gravity between oxide andbath, which is invariable for the deoxidizing metal employed, andit is also determined by the oxide particle dimensions.

The bath viscosity is a constant depending on temperature andchemical composition and therefore, after what we have said, nonaffectable as well. Therefore, the ascent rate of the oxide parti-cles may be accelerated only through their growing. If two partic-les in a liquid meet, they merge and generate a bigger particle:according to this point of view the separation of deoxidationproducts found at the liquid state in the range of the meltingtemperature of steel is favorable, while solid oxides, because oftheir high melting temperature cannot join together and grow.

Moreover the ascent of deoxidation products is influenced bythe surface tension existing between them and the bath. Thisstrength is originated by the atomic structure in the limit layers.The more these structures are similar, the lower the surface tensionis.

With the reduction of the surface tensions, wettabilityincreases; in this case one can suppose that the molecules locatedon the limit surfaces of both substances, increasingly penetrateone into the other. In this way the oxide ascent in the bath may bemade difficult and sometimes it may result even impossible. Byadding certain materials one can increase or reduce surface tension.The substances reducing it are considered as surface-active;because of energy reasons they preferably concentrate or increasein the limit layer. For this reason, even small quantities ofsurface-active substances greatly reduce the surface tension.

In steel melting, oxygen nitrogen and sulphur are particularlysurface-active; in oxide baths they are silicon dioxide (Si02)manganese oxide (Mn0) and iron oxide (Fe0). The increase in alumin-ium oxide (Al203) on the contrary, increases surface tension.Because of the low surface tension of mixed Si-Mn-Fe oxides, withsilicon and manganese deoxidation, the separation of deoxidationproducts is more or less incomplete. That is the melting tempera-

ture of these mixed oxides falls, according to the composition,down to 120000, so that comparatively big oxide drops form; theirdiameter reaches 1 mm and may have an upwards push strength ofcorresponding entity. Nevertheless a good quantity of these oxideparticles remains in the bath. On the contrary alluminium oxide,forming with alluminium deoxidation, determines, because of itshigh melting temperature of 20500C, the formation in the steel bathof solid crystals which in spite of their limited size separate ina little while. Because of the high surface tensions the crystalsof pure alluminiumoxideare not "wetted" by the liquid steel. Alsoduring the transport of the melted steel for pouring into the ingotmold we noted a remarkable reduction in non metallic inclusioncontent if ladle and pony-ladle have been carefully pre-heated andthe steel is left for at least 15 minutes in the ladle and 5 minutesin the pony-ladle.

We think that the total variation of the per cent oxygencontent mentioned by E. PlBckinger (2) on the diagram of Fig. 11 isvalid. In order to obtain as pure steels as possible, we thinkthat the following four conditions must occur:

A. When deoxidating, the steel must have the lowest possible in-clusion content.

B. Deoxiders must have the highest possible affinity with 02 sothat even with small additions one has the minimum quantity ofresidual oxygen in the steel.

C. Deoxidation products must be such that their separation fromthe liquid bath be as quick as possible.

D. After deoxidation, the liquid steel must be protected againstfurther secondary oxidation by external agents.

1.5) Remarks

Investigations on several specimens have brought to the formu-lation of the following remarks:

I. During the oxidation stage of the steel bath, calcium sili-cates are picked up by the bath itself because of slagturbulence.

II. During the refining period, the composition of inclusionsreflects the previous additions, the presence of Al in addi-tions caused a remarkable increase of Al203 inclusions. As therefining time is protracted, they decrease.

III The method for reducing Al203 in inclusions is the one of add-ing only little Al just at the end of decarburization. Ca0 mayalso be reduced by limiting the addition of calcium silicide.

IV. During tapping, slag-steel mixing must be avoided as far as itis possible.

The classification of inclusions in endogenous (i.e. due to re-actions taking place in liquid or solidifying steel) and exogenous(i.e. coming from mechanical reactions between steel and furnace,steel and ingot mold) should no longer be considered exact: in

fact, one can say that the majority of inclusions in a steel are theresult of an endogenous precipitation or crystallization on anexogenous nucleus. Stating beforehand that, as Prof. R. Kiessling(4) affirmed, it is technically impossible to fabricate steel with-out inclusions, the number of inclusions contained in a steel isoften more important together with their classification in microand macro-inclusions if we consider their influence on the mechani-cal properties of steel. As far as this matter is concerned itwould be more useful to consider the deformability of inclusions.

Making use of the same Kiessling index the inclusions in steelmay be classified in five categories:

1. Those of Al203 and calcium aluminates which are not plasticallydeformable at any temperature.

2. Double oxides of the spinel type which are not deformable atthe normal forging temperature ranges, but at 12000C.

3. Silicates which are generally deformable at forging temperatureexcept some of particular composition.

4. Fe0, MnO, which are plastic at room temperature but not attemperatures over 400°C.

5. MnS quite deformable at room temperature up to 900°C.

Unavoidable inclusions should have dimensions inferior to theones which are critical with regard to fatigue; they should have adeformability almost equal to that of the steel matrix and theyshould not generate microflaws.

The critical dimensions should be nowadays calculable with theintroduction of steel Kic. Inclusion detection methods should bethus set and finally the best technology for production and use ofsteel should be set up.

2. INNER FORGING DEFECTS

The knowledge of the behaviour of deformation during forging and ofthe direct influence of shape and use of dies, make up only part of thewhole knowledge required for a technically exact planning of all forgingoperations. According to the forgemaster's point of view all measurestaken during forging aim at achieving the maximum production and a goodquality, that is to say having pieces with suitable properties andmaterials free from defects. It is necessary however to consider someaspects concerning the behaviour of deformation during forging, thestresses which occur and the conditions for the formation and/or elimina-tion of inner defects. Stretching and upsetting forging may be consideredthe basic elements of open die forging. The separate study of these twotypes of deformation allows to draw the conclusions for all the forgingworks one can carry out. Especially in forging large pieces the directinfluence of the conditions deriving from the shape of dies, behaviourof deformation and forging section are not satisfactorily recognizablebecause they are masked by a series of components which can hardly bemaintained constant. Such as, for instance, the metallurgic characteris-tic of the ingot, the use of the best forging temperature, the maintenance

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of the temperature from the beginning to the end of forging. A scrupu-lous study of the defects that take place in forgings may neverthelesssupply, in part, important notions on their direct cause.

2.1) Plastic deformation in stretchin for in

In stretching forging, deformation phenomena were, up to now,scarcely clear because of variety in shapes and cross-sectionsobtained and because of differences in dies and in forging tech-niques made possible by presses of various power. Also the materialcondition has its importance, yet it cannot be discussed,in suchconsiderations regarding the forging technique only.

Deformation in stretching forging has to perform three mainduties (6)

A. Reducing the surface cross section and stretching the bar inthe longitudinal direction. This may be considered as theexterior characteristic of stretching forging.

B. Acting, as far as it is possible, on the whole cross section,so as to close all possible empty spaces existing in thestarting material also in the central zone of the piece.Such a duty is of the utmost importance when the central zonemust have particular homogeneity and quality properties, suchas, for instance, in huge turbine and generator forgings.From theexteriorstateoftheforging operationit is not alwayspossible to realize if such an action has been achieved.

C. Maintaining in all stages deformation conditions in the crosssection of the piece to be forged, so that no inner defectsoccur. For this reason it is important to avoid inner tensilestresses during deformation.

Experiments carried out by Marimer and Cook on steel andplasticine, with the use of tools of varying breadth for square startingcross sections, show the state of deformation in the longitudinaldirection. Fig. 12 shows the results obtained. The deformationhas been obtained from the variation in thickness of the lamels inthe deformed zone, clearly shown on the upper part of the picture.

The wide die, having a ratio between diameter or thickness ofthe piece to be forged and breadth of the die = 2 on the one handstrongly contrasts deformation in the contact surface, and on theother it causes, with its deep pressure action, a strong localdeformation the state of which is similar to the one obtained atthe center in disc upsetting. In logitudinal deformation thecentral part greatly flows forwards, especially when a strong reduc-tion is impressed, thus originating a strong shearing stress.

We are convinced that forging conditions must be duly studiedin order to obtain forgings also with sufficiently sound innerzones. As everyone knows, in such zones the material shows microand macro-segregation whose melting temperatures are quite near theforging ones and so the material has scarce hot deformability.

We believe that forging should be carried out in an adequateway in order to consider also these unavoidable anomalies.

It is clear that shape and width of dies are not the onlyelements leading to a good forging, but temperature and reductionin section obtained at every pressing with regard to the diameteror thickness to be forged are important as well.

Analyzing the parameters: die width, forging reduction,temperature and diameter to be forged, one realizes their greatimportance.

E. Siebel studied the influence of die width and pressures onforging confirming the experience with his qualitative deductions.

Fig. 13 shows that, with narrow dies, forging conditions areirregular just because of the limited die width in comparison withthe height of the forging. With smaller forgings, using the samedies, conditions are better. The two systems of flow lines from upand down merge and reach also the material at the periphery of thedie: an adequate pressure forms a closed, plastic and stressedzone.

By_forging a round bar with too narrow a die, only the part ofmaterial-ne-ar the surface Undergoes deformation while the ma rial_ _ into contact with dies remains stressed. Remarkable transversal

_stresses capable of leading to failures are generated at core.

Nowadays it is not possible to measure the stresses generatedin a piece under press deformation. An indirect means for such ameasurement is the geometrical change of shape of the piece and theforging degree. For low deformability steels also other considera-tions come into play.

Various are the systems of getting from initial diameter orthickness to the final one. Comparison among these systems isdifficult because an average criterion of forging degree definitiondoes not exist.

Diagrams of Fig. 14 show the influence of die width andforging reduction entity on the different shape of ingot end orbillets in forging.

Fig. 15 shows two examples of billet ends forged with narrowdies and insufficient reduction and with wide dies and remarkablereduction respectively. Flat dies allow the forgemaster the bestfreedom and besides they allow the material to flow during hotplastic deformation; we can also say that they lead to the cheapestforging conditions especially with products of non repetitive shape.

Dies were given a shape different from the flat one only whenan economic advantage was expected, above all in cases of severalidentical pieces with cylindrical or different shape.

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Economy generally consisted in evaluating the best compromisebetween the forging cost necessary to give the piece a shape asclose as possible to the final one, considering the normal allow-ances and defects coming from hot plastic working together withmachining cost and swarfs.

The only qualitative aspect conditioning the forgemaster wasthe realization of a forging process favorable to the taking direc-tion of test specimens. That is the so called "flow of fibers"deriving from heterogeneity existing in the ingot.

With the use of ultrasonics in forging inspection, more orless serious solutions of continuity were found along the metal-lurgical axis of the examined forgings.

Such defects were originally called "central defects".Owing to the limited experience on the interpretation of ultra-sonic examination results, it was impossible to define the nature ofsuch defects with this only test. Nowadays almost e/pryhndy_aaPeeson affirming that almost all the defects have their origin in the_ingot and that forging may only exalt or not the dimensions of suchdefectST As regar&S-this niaIter, we would remark that whfle-fn the. range of small ingots, several investigations in order to checktheir compactness were carried out; in case of large ingots theseinvestigations are quite rare or inexistent. For this reason,large ingots have been generally proportioned on the ground of theexperience acquired on the small ones, some theoretical considera-tions and indirectly on the ground of the results acquired on thefinished product. This, we believe, should make us more carefulwhen affirming that inner defects always have their origin in theingot.

Just for information we wish to point out an experiment per-formed about15 years ago at Terni on rolled billets of 160x160 mmcoming from a square cross section ingot of 380x380 mm. Thesebillets had inner pipings as shown on Fig. 16.

Several investigations carried out on ingots having differentdimensions showed that it was possible to guarantee absence ofaxial defects with well proportioned ingots cast with duly killedsteel.

From ingots belonging to the same heat of the sectionized onesafter ultrasonic examination with satisfactory result, billetsrolled by a semiblooming train were obtained. On these piecesaxial defects of the type shown on the above mentioned figurewere discovered. This and what we are going to §aY snows_thatdefects may form and grow even and only because of the hot plastic_working. Nevertheiess we agree on saying that the ideal hot plasticworking cycle should be chosen not only to prevent defect formationbut also, as far as it is possible, to facilitate the elimination ofinner defects if already present in the ingot. As far as the forg-ing conditions determining the welding of possibly existing defectsare concerned, the most important of them are: total deformation

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degree (forging ratio) in upsetting and stretching operations,deformation speed, stress distributions and entity of stresses inthe defective zones; specific pressure, temperature etc.

On the other hand, obstacles towards forging welding ofdefects are found in the high susceptibility that alloyed steelshave towards defects, in the presence of heterogeneity due to segre-gation, impossibility at times, of performing the total reductionat the required speed and in one time. We shall further on reviewthe experiment carried out at Terni about the influence of shape anddimensions of dies on the formation of inner defects.

2.2) Experience data

For about six months three experiments aiming at clearing upthe influence of shape and/or dimensions of dies on the formationof inner defects were carried out with the use of three NiCrMosteel ingots having a total weight of 152 t, 136 t and 59 t. Theywere transformed by forging into billets with diameters of 1650 mm,1200 mm and 1700 mm respectively; then they were examined by ultra-sonics.

The three ingots were made of vacuum degassed electric steelproduced with the double slag method in perfectly normal conditions.

Table 1 shows the chemical analysis of each of the threb heatsemployed. The forging cycle foresaw, in each case, the use of flatupper and lower dies in certain forging phases, in contrast with thenormal practice which always foresees the use of upper and lower"V" dies of the type shown on Fig. 17.

Table II shows the complete cycle followed in the three cases.As one can see, the flat dies were employed in the 6th heat foringot no. 1 (152 t), in the 5th heat for ingot no. 2 (136 t) and inthe 3rd heat for ingot no. 3 (59 t).

After annealing the billets were rough machined and examinedby ultrasonics; the results appear in the reports of Table III. Itcan be seen that all the three cases show lozenge-shaped defectivityof remarkable entity distributed in the axial zone; it is attributedto probable porosities. The billets no. 2 and 3 after the checksperformed on the billet no. 1 and mentioned hereinafter, were re-forged according to the modalities also described on the cycle shownon Table II. Only from the defective zones of billet no. 3, beforere-forging, in order to be sure about the nature of the defects, twocores with a diameter of about 30 mm were taken. Their aspect isshown on Fig. 18.

Then the two holes of the cores were filled and hot repaired bywelding before re-forging. After this last operation performed withthe normal upper and lower "V" dies we obtained two forgings com-pletely free from defects.

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2.3) Quality checks

From forging no. 1 we took an axial core trying to take awaythe whole defective zone. The draft of Fig. 19 shows the dimensionsof the axial core bore and the specimens obtained from it for microand macro-examinations. The core was taken from the bore alreadybroken into three pieces. The fracture appearance is shown onFig. 20a and on Fig. 20b in detail.

The surface was absolutely free from any hot oxidation,macroscopic non metallic inclusions and overheating phenomena.From the part shown on Fig. 19 we took a specimen in the longitudi-nal direction. It was examined by sulphur print and acid etching.Fig. 21 shows the relative aspects. The material appears withnormal segregation, considering that it is the ingot axis, and withother smaller breakings near the main one.

On Fig. 21 we have also shown the positions A and B where thetest pieces for microscopic examination have been taken; theiraspect is shown on Fig. 22.

One can see that, except for some sulphides, the material hasa limited inclusion content, and failures have probably occurredalong the boundary of the austenitic grain existing at forging.The small failures shown on the micrographics of Fig. 22 might con-firm this.

2.4) Inter retation of ex eriment results

2.4.1) Distribution of deformations

Hot plastic deformation of a huge forging ingot is chieflyperformed by stretching. Upsetting is at times performed duringa certain stage of the cycle for reasons of quality and shape.During stretching the dies come into contact with a part of thematerial to be deformed with their whole width or only with apart of it and penetrate for a certain depth into the material.Shape and dimensions of dies generally are chosen according todimensions of the forging system, shape and dimensions offorgings, and quality of the material. Fig. 23 shows diagramaccording to H. Mattel (11) where it is possible to remark thevariation in width of the die and starting dimensions of thepiece to be forged as function of press power.

It must be noted that the ratio between diameter orthickness of the piece to be forged and die width variesbetween 0.25 and 0.5; in the final forging stage this ratiomay become 0.5-0.7. The irregular distribution of local defor-mations during forging is attributed to cooling on dies andalso to cooling of the material. The technique of allowing ex-ternal zones to cool before forging has been adopted for a longtime in forging systems, but it sometimes finds its limits inpower of presses.

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Very interesting appears the work by M. Vater and H. P.Heil on the distribution of deformations with the varioustypes of dies. The authors (9) mention a comparison of de-formation on the SedfIdff-ef-rd:Ies-t_piee-eSlairing-stret-ching'withAiTferent die .shapes Fig,_24., end demonstrate that wi-th- -the [Fee-of-round or "V"_diea having an angle of 1350 oneTanh-6- iethe maximum deformations in the axial zonewkiile ai gt-bh-ingildtisqat die and lgwer .1w."--die -with an PD91„8,9171200 one has deformations in the axial zones inferior COothers, confirming Ç Coupette's previous opinion One canIsafënIaTh'ilkatwith flat upper die and "V" lower die with anangle of 90°C, we have deformations in zones which are distantfrom the axial one even of 1- r (radius) up to 3/4 r duringstretching. /in such a way it is possible to have failuresbetween the more external zone and the axial one,/ The authorsshow that with "V" dies with angle of 1200 the risk of failureis greatly reduced and the mpximum deformation_at core is_obtained_with_ angles of_1350. On sueh resulte We also agreeand when radii of angles of 1250 and 1350 are great, one has auniform distribution of deformations such as with round dies,with maximum values also at core.

2.4.2) Penetration depth

During stretching of square sections and perhaps also ofoctagonal sections, the axial zone of the piece may be forgedalso with slight penetration depths. Round sections insteadshould be stretched with penetration depths superior to 10% inorder to have good forging also at core. Octagon stretchingis better not only to give good core forging, but also toavoid inner failures. The axial zone of the piece may beforged also with small reductions but in order to realize thenecessary forging conditions dies must have shape and widthaccording to ingot dimensions and type of steel for a givenpower of the press.

2.4.3) Sha e and dimensions of dies

The first condition for a good core forging is to befound in the use of "V" dies with an angle of 125-135° so thatwith their opening they can include on four sides a somewhatoctagonal section, see Fig. 17.

The breadth of dies must no,t he_inferiar_to 0,4 _times thediameter-t6 be-Thiged. No doubt that increasing the die-piece--surface one can meet difficulties because of oreater powerbecoming necessary. If the power_of_aVailable.presses_is.notsufficient narrower dies must be used but in these casesforging_penetration is no MOT,Q„ Sufficient at the first stages,_the forging effect arrives at core only in the final forgingstages.

In these phases the defects will weld only if the reduc-tion is remarkable and if they are of small entity.

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2.4.4) Effect of tem erature

The deformation temperature of the material changes dur-ing forging; there is a strong temperature gradient betweenthe surfaces into contact with the dies and the core zones ofthe forging. Of course also these temperature conditionshave important effects on forging and on possible growth ofinner defects.

In the usual open die forging process the ingot or inter-mediate forging is taken from the heating furnace and under-goes hot plastic forging until the piece has reached the"forging end" temperature with reference to power of the pressand cross-section to be forged. This generally means that westart at a temperature of 1230-1150°C and end at about 8000Con the surface.

The loss in temperature on the surface has a remarkablepositive effect on a good forging penetration.

In the limit case of an external temperature of 700-750°Cand inner temperature of 12000C, the axial zone undergoes aforging effect almost as if one operates with a round dieconsisting of the external shirt of the colder material.This naturally requires a powerful press and material whichmust be hot plastically workable up to lower temperatures.Diagram of Fig. 25 shows cooling curves relative to ingots orintermediate forgings heated up to a temperature of 1200-12300C and air cooled.

The curves relative to diameters of 1250 mm, 500 and 800mm were experimentally plotted while the ones relative todiameters of 1000 and 2600 mm were calculated.

The forgemaster, knowing such variations in temperaturemay have a better knowledge of the core effectiveness especi-ally in final stages of forging when one has low surfacetemperature and insufficient capability of the press.

2.4.5) Ca abilit of o en die for in resses

With regard to the capability of the large pressesavailable in the various countries, the upsetting operation isperformed or not on large ingots intended for large rotorforgings for the electromechanical industry. The suitableload of a press is determined by practical experience andapproximative calculations. According to Haller (12) twofactors regulate it: the cross-section of the largest ingotand the mechanical properties of the material at forging temp-erature.

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2.5) Remarks

In order to calculate the press load the following formulais generally used:

P = A Rf A = area of contact of diesRf = deformation strength of the

material at forging temperatureE = deformation effectiveness

The actual A value depends on the type of forging opera-tions to perform: upsetting or stretching.

The deformation effectiveness E considers friction lossesbetween pieces being worked and dies and the resistance to theinner flux which chiefly depends on the type of deformation.Nevertheless the deformation effectiveness greatly varies fromone case to another. For this reason deformation strength isgenerally given by Rd = lifwhich may be obtained from thediagram of Fig. 26 (12).

We believe that in order to meet the greater demands ofhuge forgings for the electromechanical industry, modernpresses should have powers not inferior to 20,000 t in orderto forge ingots weighing up to 450 tons.

Publications concerning the choice of a press capable offorging ingots weighing up to 500 tons are rare or do notexist. Sanderson and Frith recommend the following data forforging ratios from 1.4 to 3.1.

A thorough analysis of the experience data acquired up to nowon heavy forging leads to consider, above all, flye_interdependentparametersia_order_to haVe_9.09“99i89of axial_zenes: tempera-tu-re, diameter to be forged, width and sh-6-0-6-bt dies, amount offorging reduction.

On the ground of the data found in the bibliography derivingfrom direct experience and obtained from the former by extrapola-tion, we have tried to synthetize on the diagrams of Fig. 27-28-29an example showing how the various above parameters might inter-dependently vary in order to always have good forging on the axialzones of intermediate forgings of 2000 mm diameter.

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3. EXTERNAL FAILURES CAUSED BY FORGING AND HEAT TREATMENT

We are convinced that the kind of defects we are going to deal withhave their origin chiefly in hot plastic working and inevitably developin the heat treatment stage. Other defects on the contrary, as everyoneknows, may chiefly have their origin in this last operation.

We are normally used to consider as hot forgeability or hot plasticworkability of a steel, a characteristic which should define: strengthopposed to deformation and fracture ductility. The former determines theentity of the involved stresses and therefore the power of the systemwith which a certain process must be realized; the latter determines themaxim, deformation the material may undergo without generating fracturein some service conditions. The critical factor for the success of manyhot plastic workings, especially in the field of heavy forging in specialsteels generally is fracture ductility because the available systems areoften insufficiently over dimensioned for meeting unexpected increases inthe required power, which makes it necessary to make use of artificessuch as increasing forging temperature or adopting of inadequate dies.In the various works carried out for many years by various researchersaiming at determining the best start and end forging temperatures forsteels the reason for a limitation in the highest temperature is clearbut we do not think there is as much evidence on the reasons for minimumtemperature limitation. We must point out that in rolling systems thereis no problem of minimum temperature while this is nowadays very importantin heavy forging systems.

--- We are going to review some serious failures which sometimes take

Vplace on large alloy-steel pieces after annealing and that, in spite ofappearances, we think they must be ascribed to the final forging opera-tion.„.---

The need of using very wide dies with an ample seizing surface onthe piece in order to obtain good core penetration is certainly very goodbut we think we must also keep under control what takes place on thesurface of the forging in contact with the dies. This area in fact, inthe first forging phases, undergoes limited cooling because of the greatamount of scale adhering to the piece and acting as an insulating body.In the final phases instead, when the temperature is near the end forgingone it is possible to have pellicular temperatuPes of 200°C especiallyon sections with minimum diameter when the dies come into contact with thepiece. Crack cannot be normally seen by visual examipption, but,puttingthe piece-,again inte-the_furnace_at_a_temperature of 1000-1100°,C_and,pro--Piedwith forging the said zones with remarkable-reductions; it is possi-,ble-to see small pelliallar failurps whichloften coMPel-f-o}gPMBsters-tomake use of furt-De'r-fr6m-e16,asDings. In case no further forgingfs per-=_'rd—rmedand the pieCe is air cooled or put into a furance for annealingwith the above smaller section at a temperature of 500-6000C with a stopat 600°C, slow cooling in furnace down to 300°C, one or more heatings ataustenitizing temperature preceded by stops at 3000C and final tempering,serious failures of the type shown on Fig. 30-31-32 may occur. It isknown that in large forging shops, where mechanization and modernizationof facilities is not yet remarkable at least because of high investmentexpenses, for various reasons forging is finished at low temperature

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especially when the piece is characterized by great differences in dia-meter or thickness.

We think that in such a critical stage failures may develop becauseof phase transformation of steel according to chemical composition,temperature and deformation. The phenomenon is generally more frequentin high-hardenability steels. On Table IV we show the chemical composi-tion of the most typical ones. On Table V we schematically show endforging temperatures relative to the forgings of Fig. 30-31 respectivelyand their annealing cycle. The visual aspect of breaking surfaces wasthe same for almost all the cases (see photo of Fig. 33).

On Fig. 34, where the magnified aspects of the same surface aredocumented, it is possible to note a zone next to the surface, more orless deep, and with remarkable oxidation starting in the radial directionfrom the external surface, the successive inner zone is less oxidized andat last one or two zones distinguished for oxidation degrees becausethese have occurred at temperature not higher than 65000.

The micrographic examination performed on several radial corestaken in correspondence of the failures shows in all cases always analo-gous appearances, except for the different structure dapending on the typeof steel. On Fig. 35 to 39 we show the micrographics relative to the coreof Fig. 34 at different depth and in positions with different oxidationcoloring. It is possible to remark that failure has grown in differentphases and that the deepest zones of failure with decarburation reflectthe annealing cycle followed (Table V). From the oxidation condition ofthe failure surface, it is possible to affirm that the last 3/4 offailure has taken place because of thermal stresses in the final coolingstages of annealing. The measures adopted, such as decreasing final cool-ing rate in order to reduce stresses did not reveal any effect.

3.1) Remarks

We believe that big failures_which_may sometimes be found onalloy steel for_gings after annealing or after go5lin2_ing_teqperature are to be cconsidered in onneatiOn with the lowtemPerature involved and with deformations taking place in the finalforging stage. In such zones cooling under pressure and deformationwith consequent-phase transforMaions ganarating etressea raading topellicular failures take place: -That--Maa-na-Ehat-fadI-t-hermomeehant=-car -EFaai-M-e6ts15YTTI-ghTaf-IEICTemperature may take place, see Fig. 40.

Well, in the final forging operation it is possible to haveboth processes because the metal in contact with dies undergoes cool-ing and deformation at the same time. Cooling and deformation areremarkable on the piece-die contact surface and will progressivelydecrease in the zones below.

An important difference which must be considered is that whilein the real thermomechanical treatment the deformation of materialis performed only once, in the final forging operation one zone maybe deformed several times, which we think causes failures.

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We are carrying on a research at the "Centro SperimentaleMetallurgico" in Rome, for the determination of the influence thatminimum forging temperature, as function of the number and speedof deformations, has on failures. This research is being carried onby means of a special apparatus called "plastometer" studied and madeby C.5.M. itself. The first results of the researches alreadycompleted appear very interesting and seem to confirm what we havejust said.

The diagram of Fig. 41 shows a curve (13) relative to NiCrMosteel hot deformability as function of test temperature. Beforestarting the tests, all the specimens were heated at 1000°C for 15minutes. The curve shows an apparent fall in ductility at 780°C.

In the final forging phases the pellicular metal undergoesthree different stresses: cooling stresses, deformation stresses andstresses caused by phase transformation. All these strains beingcontrasted by the much hotter inner metal, may generate more or lessdeep failures which the successive annealing cycle, because of itscomplexity, may help developing.

4. INNER NITRIDE DEFECTS

Though these defects are more or less well known, we wish to expresssome considerations on their aspect, very similar to "burglf_material,and on the danger they cause.

k'They may sometimes be misunderstood if no microanalyses are per-

formed and if one does not ascertain whether the material surrounding thedefect comes up to the surface or not.

Nitride_dsfects_may chiefly occur in zones_q1Stal_mith greaterthickness and which solidify the last, i.e. upwartis_near the axial zonesahretung-thb-Mpbrehite grain-boundary. This- defectivaly ma-7 have macro6-n-a-microscopig-ba-ffbhbrons. On fd-rgings it is rarely macroscopic becausethe forging operation can disperse nitrides and also reduces their dimen-sions sometimes to microscopic level. The defect may be found in a moreor less serious degree according to the particular oxidation conditionsof the steel bath in the refining stage, and to the quantity of deoxidersuch as A1,Ti and V added in the presence of an excessive amount ofnitrogen in the bath. Nitride defectivity more often affects Cr-Mo-Vteels. The materialci;hteinin-rnitircre— Eref- e-Cfs'always-hab-VbFrgir--ductility and toughness-properties; i sometimes in the cast conditighFigg-Ohly goodArengthi propertree; while in the forged cOhdition it shOws'chiefly a decay of'toughness. In most cases even repeated heat treat-Mbnts cannot improve the poor mechanical properties of a material withsuch defects.

Another aspect we think worth emphasizing is that this type ofdefect cannot be detected by ultrasonics; it can be found by magneticparticle test if it reaches the surface if it has undergone good finish-ing machining; the defect must also show macroscopic solutions of con-tinuity caused by thermal and/or mechanical transformation stresses.

- 21 -

Fig. 42 to 45 show macro aspects of the breaking surface of a castspecimen containing nitrides, its section after Oberhoffer and 50%hydrochloric acid hot etching, the micrographic aspect and the appearanceof a tensile test specimen. The visible coarse grain, makes up the boun-dary of the austenitic grains of the material along which nitrides areplaced. fjt can be seen that the material has a fracture appearanceremarkably similar to the one of overheated or burnt steel.

On Fig. 46-47 the same type of defect and its micrographic aspectare documented; this was obtained from a forged specimen after tensileand impact tests respectively. It has smaller dimensions and on somepoints only it shows an appearance similar to the one that may be foundin the as cast condition.

As regards micrographic examination, nitrides may be easily observedat 500 and 1200 magnifications. The presence of nitrides was confirmedalso with microanalysis.

As we have already touched upon, said defect is to be considered asperhaps the most dangerous because it cannot always be detected with theinspection means available nowadays. We think the only good way to avoiddangers from nitrides is to act on steel manufacturing in the furnace,using well selected iron-alloys charges with controlled nitrogen content,controlling oxidation and deoxidation in the steel bath, limiting and/oravoiding deoxidation with Al and Ti if not immediately afterErie ena—trfthe—oxidation studerr------

_

5. ORGANIZATION OF THE METALLURGY AND QUALITY CONTROL DEPARTMENT IN "TERNI"

STEEL COMPANY

The introduction of new and more developed inspection means such as

electronic microscope, microprobe, ultrasonic and magnetic particle testsetc. have permitted a better knowledge of steels features and how to

improve them; at the same time transformation and utilization are in

continuous progress.

For such reasons the organization of the Quality Control Department

has been characterized by remarkable evolution in these last time. It

aims at being based on "equipes" of specialists independent from produc-

tion and grouped in a centralized department at present called "Serviziodi Controllo Qualita" (Quality Control Department) or "Servizio

Metallurgico" (Metallurgy Department).

These specialists will certainly bring all the results coming from

their experience of laboratory studies, manufacture data gathered in the

workshops and customers' requirements. In reorganizing the Metallurgy

and Quality Control Department of Terni Steelworks which dates back tothe month of May 1969 many organization criteria applied in other simi-

lar firms were considered. This has taken place after considering the

various high quality products, the actual availability of resources and

the need of developing the department, defining responsibilities andcreating the conditions for the best exchange of experiences.

- 22 -

The department consists in four metallurgical "lines" coveringhomogenous products and in a group of specialists for applied research;they act in connection with external research bodies and with all theinternal control laboratories.

We must also mention the establishment of a group for statisticalanalyses intended for a widespread and systematic application of suchmethods. Each metallurgical "line" is organized so as to perform andcontrol all the operations relative to the whole manufacturing process ofthe products it covers. Each line is responsible for issuing quotationspecifications, issuing and bringing up to date of standards, metallurgi-cal assistance to the relative area, phase controls and inspections onthe finished product. The Statistic Centre gives its assistance to each"line" for the issuing of periodical qualitative reports and for thestatistical formulation of short and long term metallurgical investiga-tions. This centre has its own computer and may also make use of thecentral computer.

Also a Quality Control Handbook has been prepared; it deals withorganization, specialists' duties, all control procedures of purchasedmaterials, qualification of suppliers, personnel qualification, controland setting up of equipments and control procedures.

In order to ascertain the rigorous and systematic compliance withthe above synthetized procedures, an inspection function is being devel-oped in the Quality Control Department.

This inspection on the various controls is performed with systematicchecks based on a monthly program, devised by the responsible for suchfunction, only known by the Quality Control Management and daily communi-cated to the people in charge of the inspection. The program has thepurpose of ascertaining the compliance with control modalities and inparticular;

- the knowledge of specifications relative to the performed control bythe personnel employed in such function.

- the efficiency of the equipment used and the performing of setting upand periodical maintenance to the prescribed frequency.

- the correct use of the prescribed control materials with regard totheir type and preparing modalities.

- the respect of security principles.

- the exclusive employment of skilled personnel.

- 23 -

BIBLIOGRAPHY

1 - Pellini (W.S.), Pinzac (P.T.) - N.R.L. Report 5932 Washington D. C.Marzo 1963

2 - E. PlOckinger - Influence of deoxidation practice on cleanness of steel

3 - G. Pompey, B. Trentini - Quelques considerations sur le proprietedesaciers. Revue de Metallurgie 1971

4 - Prof. Kiessling - Inclusions in steel, vol. III

5 - Morgan - Formation of non-metallic inclusions in electric arc steel-making. Journal of the Iron and Steel Institute, Oct. 1968

6 - M. Kroneis e Th. Skamletz, Kapfenberg - Contributo alla conoscenza deifenomeni di deformazione plastica durante la fucinatura.Berg, U. Hatenniannische Monotshefte gennaio 1964

7 - E. Siebel - Stahl u. Eisen (1925) 5.12 bis 14

8 - Cook-Metal Treatm.Pro.Forg. 1953 November 5541 bis 548

9 - M. Vater u. H. P. Heil - Unformbedingnngen und Gestaltung derWerkzeuge beim Frieformen - Stahl u. Eisen 91-1971-22 July

10 - Gerhard Richter u.Hans, Georg Lotze - Zur Froge der Durchschmiedungbei der Herstellung grosser SchmiedestUcke Neue Hate - Januar 1965Heft 1

11 - Mattel H. - Vorgetragen aulgsslich der 12 - Vollsitzung desSchmiedeausschusses des VDEham 26 April 1961

12 - Hans W. Haller - Handbuch des Schmiedens - Principles of Forging Presses

13 - P. Brozzo - Rapporto interno C.S.M. Roma

Mantichi Tateno, Shoichi Shikano - Study on closing of internal cavitiesin Heavy Forgings by Hot Free forgings

040

ANOMALIE RILEVATE

Distanza stela Aspetto DimensioniIle indicazioni empo delta Oil etti

a "X" mm, indicazioni mmq

ApperecchioFrequenzaTn sduNoreAttains eco fondoMezzo di accoppiamentoStab° superficial* del fucinato

CONDIZIONI 01

3593337

54208625154815884820511 0516 0524 06900

mg +76707918795841376235

6600+ 66206793684571087511779215372140350235475154I 70219125434561,96603810837205278

4630 +465065987016804516253141329354045685570657355868

.•

11.10 iforme0.50 20.250,500.25 061.301.501. 30 1100 21.031. 10 ..I. 55I. 20 Run gata 5I. 05 Roane 1.21.00 1.52.252.45 „ 12. 47 Ilungata 22. 25 Worm*2.282.422.552.15 /.53.303,35 23.20 23.00 TS3,404.154.304.25 1.54.30 1.54. 17 1.54.02 •6.30 .. 15.30 1.26. 40 Ilungata6. 35 ntiforme 26.15 .•7.55IL 3511.3511.3511.32II. 18 2II. 2011.2511.30 5

F ro Pos.none dal ditto. tipo inclusioni non matailiche in tanluainaio per rotor. tr acctato el Ni Cr-Mo • tt modaliti dell'esameU S

L AVORO

KRAUTKRAMER USIP 10 W.2 MHz.132S UrnmSS mm.olio SAE 20sgrosaato di macchina100 % deli* superf ici

1 2CA.ROTA A 37 -Milli/

x 50 senza attacco x50 senza attacco

x 10 senza attacco

Fig. 2 - Aspetto micrografico del difetto prelevato nellaposizione A di cui al la Fig. 1 .

Fig. 2 — Micrographic aspect of a defect taken in the position A

x10 - Sezione longitudinale provetta di trazione

Fig. 3 - Aspetto delta provetta di trazione su inclusioni( difetto N213 ) .

Fig.3 — Appearance of a tensile test piece with inclusions

x2 - Sezione di rottura

provetta di trazione.

Risultati prove di trazione su provetta "a" contenente

difetto N2 13 e su provetta "b" prelevata dallo stesso fucina_

to su zona esente da difetti.

Fig.5— Location scheme of specimens

for the control

of inclusion

content

1= 1:

a •

t

4W

'

t.Z

-'O

s

£3

a

Fig. 4 — Locations in the original ingot corresponding to anomalies

found in the forging shown on fig.1 With solidification

isothermal lines

fl

fl 5' 00

Fig,6— Some types of non metallic

inclusions

more recurrent

on specimens

shown onfig.4

Scoria durante it riempirneto delta lin ottiera.

Battente tra materozzalingotto phi accentuatoIi s i oil

Fig. 7 - Distribuzione delle inclusioninon metalliche tipo• 8) secondo lasezione longitudinale del I ingotto.

Fig.7 — Distribution of non metallic inclusions of type 5 on thelongitudinal section of the ingot

Fig. S - Portico lare dell'estremittsuperiore del lingotto che pud trot.tonne le inclusioni non metalliche&warn ii riempimento in lingottiera.

Fig.8 — Detail of the ingot top which may holdnon metallic inclusionsmoldis tieingfilled

Getto di col4ta.

Fronte di colata.

Inclusioni che si trastormanoin • Iineelongitud.nali dopofucinatura.

when the ingot

x—x Ca - Si

—0 100 0—iso Si Metallico

a 1600 °C

200

0 • x

Zog 100

00 0,5 1,0 1,5

Disossidanti aggiunti in %

Fig. 9 - Andamento del tenore di 02 (ppm) in

tunzione della percentuale di' disossidanteaggiunta.

Fig. 9 — Behaviour of the 02content (p.p.m.)

as function of deoxidizer percentage added

100

90

80

500

are cameo

*OMB* 011• •

Disossidazione con Mn +Si

Disossidazione con Mn +Al

70

60

05 10Aggiunt a di Mn %

Fig. 10 - Velocita di disossidazione delSi e AI in funzione del tenore percentodi Mn aggiunto.

Fig.10 — Deoxidation rate of Si and Al asfunction of the Mn percentage added

Tempo diSpil le M Tempo di st In sin swim

4010

4080

4070

x‘.gopw

Pe. 0

4030

4020

-

oCeSiSi

AAl Mstellico

'7 Er

}

0 TiSiMs 47:11•

f.C Si

11—1

oingie none peniera

SemPO di cologgio in l•ogotliera

a---4D ' — --t:L., , \I A-1'27 ---17

• Valori otlemtli4010 ' ' rAimil lirooilo

2 4 6 8 • 10 2 14 M . 18 20

TEMPO. m'n.

Vaniezione del lenem dell'ossigeno Wale, col dilleremi

disossidam. e durame le gaMe fasi dell. Colata, Second.)

E. PLOCKINGER.

Fig. Il

Fig.11 - Variation of the per cent Oxygen content with different deoxidizerduring the different stages of the heat according to PAckinger (2)

A

C-- 0

•

• /1.21 1-28 /1-42IA' IA A

Fig, 12 - Deformazione nella fucinaturaper stiramento: influenza della larghez _za degli stampi ( second o P. M. Cook.

Fig.12 - Deformation in stretching forging. Influence of die breadthaccording to P.M. Cook

Fig. 13 - Penetrazione della fucinatura con stampi stretti (7).

Fig.13 - Forging penetration with narrow dies

8

a) Riduzione 5 %

10%

Riduzione 11 %

20%

c) Riduzione 18 %

A

A

a) Riduzione 4 %

b) Riduzione 10 %

20%

40% N%

Fig. 14 - Rappresentazione schematica deWinfluenza

della larghezza degli stampi (1) ed entital di ri

duzione sulla forma diverse delle estremiti dello

sboztato.

c) Riduzione /9

Fig.14 - Schematic representation of the influence of die breadth andreduction entity on the different shape of the forged piece ends

Fig. 15 a

Fig. 15 b.

Aspetto di due estrernitå di sbozzati in acciaioal Ni CrMo fucinati con stampi stretti (Fig. 15 a)

e larghi ( Fig. 15 b ).

Fig.15 - Appearance of the end of two intermediate forgings forged withnarrow and wide dies

12 •

e.)

Fig. 17 - Stampi a "V" per la fucinaturadi grossi diametri.

Fig.17 "V" dies for forging huge diameters

Fig. 18 - Aspetto dei difetti u asse del la billetta- esperienza N19 3e un dettagl io dei difetti .

Fig.IM - Appearance of the defects on the axis of the hi I et of ON riment

110.3 and relative detail

2530

1160

15

30

Fig

. 19

-

Pos

izio

ne

della

ca

rota

as

sial

e su

lla

bille

tta

espe

rienz

a N

! 1

.

Fig.19 — Location of the axial core on the billet of experiment

no.1

Fig, 20 - Aspetto di rottura della carota assiale

di cui alla Fig.19.

Fig.20 — Fracture appearance of the axial core of fig.19

a

:2 Attacco con NCI al 50 % a caldoFig.21 - Macrogmphy and sulphur print on the

longitudinal section of a part of the coreshown on fig. 19. Positions A and B of thesamples for the micrographic examination

x 6 x 6

Irnpronta Baumann,

sezione longitudir

:1,5 Attacco con reattivo Oberhoffer

Fig. 21 - Aspetto dell' impronta Baumann e macrografico sulla sezione

longitudinale di uno spezzone della carota di Fig.19 . Posizione A e B

dei campioni per l' esame rn crografico (v. Fig. 22) .

CT

X

a)

if

C)

a)

100BO60

Z 40

10

4t to

6

200 400 600 600 1000 1100 1100

Larghezza stamp* a dimension* iniz ia, dal luc Mato in anal.

Fig.23 — Variation of die breadth and diameter of the intermediate forgingto be forged, as function of the power of the press (11)

Fig. 13 - Varian ion* delta laeghezza dello *tempo a Marnelmdell* abozzato da NCinart icr lunzione dells pol*nzadelta mess.. (II /

d)

b)

Fig. 2i - Contronto deleandamento della def ormazione sulla sezione di provinitondi durante la stiratura con differenti forme di stampi. (9)

Fig.24 — Comparison of the course of the deformation on the section ofround test pieces during stretching with different shapes of dies(9)

1230nmumLI moogoo

' 800 fit.600 )11.750 0=1000 0.125091230

€252

20

5

60 120 180

Tempo , min.

Fig. 25 - Curve di rat freddamento in aria. di sbozzati

di diametro diverso, delta teinperatura di 1230 °C .

Fig.25 - Cooling curves in air from temperature of 1230°C relative toforgings having different diameters

ts •\V CL<S•cC

. 0.22m

OA?' •1.

600 700 803 900 1000 1100 1200Temperature, •C

Fig. 26 - Resistenza all a delormazione dagli accieiel carbonic in lunzione delta temperature. (fe)

240

0.2500

Fig.26 - Deformation strength of carbon steel as

function of temperature.(10)

Fig.27 — Interdependence among surface temperature of the forging per cent

reduction and diameter to be forged

..0elormazione suite tizone assist.:

= 300 mm.

(0-)eoo.Larghezza degli d-stampi: L=1200 mm. 2

600

E

-0 4090

E 300

250

E200

fl

;- ISO

9 100

1200

t so0

00 5

//WM

Fig. 27 - Interdipendenza Ira ternperatura in superlicie del tucinato,

percentuale di riduzione e diarnetro sbozzato da Iucinare

to 15

P-Riduzione

5 10 iS N 2e 0P- Riduzione

20 25

Fig. 28 - Interdipendenza Ira detormazione

fucina.to e percentuale di riduzione.

suits

250 056..y

b'

zone ass

woo of

750 •-å>

500

ale , raggio del

20013

- Temperature sulfa sup6.

fide det lucinato: 1200•C

-Larghezza stamp:1200[16/n

Fig.28 — Interdependence among deformation on the axial zone, radius of

the forging and per cent reduction

-Detormazione sulla zonaassiale: 0=300

-Oiametro da tucinare:id =2000 mm

1200

1000100

CO

%% .5 800

0.%

0 S 10 15 20 25P-Riduzione %

600

Fig.29 - Interdipendenza fra temperature in superticie deltucinato, percentuale di riduzione e larghezza :tempi.

mpo.f

800

1200

Fig.29 - Interdependence among surface temperature of the forging, per centreduction and breadth of dies

Fig. 30 - Rottura rilevata su una billetta di sezione 250 x 250 mm.

in acciaio Prestem dopo fucinatura e raffreddamento in aria.

Fig.30 - Failure found on 250x250 mm. Prestnm steel billet after forgingand air cooling

Fig. 31- Fucinato per rotore generatore in acciaio al Ni-Cr•o-V

del peso di 50 t. ca. con una rottura su un collo (indicata con

la freccia).

Fig.31 - Ni-Cr-Mo-V steel egnerator shaft forging weighing about 50 tonswith failure (pointed out by the arrow) on a Journal

Fig. 32 - Aspetto in dettaglio della rottura nel fucinato di cui

alla Fig. 31.

Fig.32 Detail of the failure appearance of the forging shown on fig.31

ci

7 0

fl

73 14 Is Is II

20

Fig. 33 - Aspetto delle rotture sulle carote prelevate sulfucinato di cui alla Fig. 31.

Fig , - Fracture a ppea rance on sampl es taken from the forp-ing shown on

, 3 1

Fig. 34 Fig.34 — Appearance of the fracture surface of fig.33b with detailsmagnified x 15 at different depthst .

e

LO

r 0 tn0

0 m

W

Oss

idaz

ion

e

Bo

rdo

del

la

rott

ura

Fig.35 - Micrography

performed

on the edge of the failure on positionF

relative

to the specimen

of fig.34

Attacco:Nital 2 % x BO

00

Fig. 36 Micrografia effettuata sul bordo delta

rottura in posizione G del campione di cuialla Fig. 34.

!z. HI cut!LtHv po Fro rmod art t Ito th It' al ha lit I lira on pos i ti on 6

I ti VU tt t [iI/OH III0/1

fl c

Fig.37 -Micrography

performed

on the edge of the failure on position

Hrelative

to the specimen

of fig.34

HV

=17

616

4 20

8 19

7

Bor

do

delta

ro

ttura

218

225

229

218

co

II

Attacco:Nital 2 VD x 80

Fig. 38 - Micrografia effettuata sul bordo della

rottura in posizione 1 del campione di cui

al la Fig. 34.

Fig.38- Micrography performed on the edge of the failure on positionrelative to the specimen of fig.34

en

en(N

(N4CN

CO

Attacco Nital 2 % x 80

Fig. 39 - Micrografia effettuata sul bordo della

rottura in posizione M del campione di cuialla Fig. 34.

Fig.39 - Micrography performed on the edge of the failure on position Mrelative to the specimen of fig.34

AO

30

20

15

Oeformazione

Tempo --e-

Temperature divicristal I izzazione

Fig. LO - Rappresentazione schematica del

processo del trattament o ter momeccanico

ad alta e bassa temperature.

Deformazione

w Per entrambi i processi

il raft reddamento deve

essere rapido per evita_

re la ricristallizzazionet

Fig.40 - Schematic representation of the thermomechanical treatmentprocess at high and low temperature

800 900 1000 1100 1200

Fig. 01 • Andarnonin dent, dultilild in ploy. di 1cdsione continu•

su materiale lugineto in acciaio al Ni•Cr•Mo. 1131

Fig.41 - Course of forgiability continuous torsion tests on Ni-Cr-Moforged steel (13)

Fig. 42 - Aspetto della superficie di frattura di un saggio fusoin acciaio Cr-Mo contenente nitruri di Al .

.42 — irk ihc a Cr-Nlt Lot.. I con hi In Ilg A I ni i dos

Fig. 43 a - Aspetto macrografico dopo attacco a caldo

con HCI al 50 % del saggio di cui alla Fig. 42,Fig. Macrographig aspect of the specimen

Fig. 43 b Aspetto macrografico dopo attacco con reattivoOberhofter del saggio di cui alla Fig. 42.

F14,-.43h- Macrographic aspect of the specimen of i

etching

g.42 after HCL 1k.

.42 after Oherhoffor

x 2

etch

x 2

S.A.

S.A.

ax 100

x 630

Fig. 44 .- Aspetto micrografico (a) e relativo

particolare (b) del saggio di cui alla Fig. 42.

Fig.44 — Micrographic aspect of the nitrides on thespecimen of fig.42

x 6

Fig. 45 - Aspetto di rottura di una provetta di trazione ricavata

dal saggio di Fig. 42.

Fig.45 - Fracture appearance of a tensile test specimen obtained fromthe sample of fig.42

x 6

, t

e.

k.

trtt ;rb-.131/4,

ephs, )1,4!

2 48 4At; t

NOS/

.• t 'Vt5",`ikt:r ,

Fig, 46 - Aspetto di rottura d i provette di trazione e di resilienza

ricavate da un saggio tucinato in acciaio al Cr-Mo -V contenentenitruri di Ti .

Fie.4b - Fracture appearance of tens i le andinpa ct t es t piece obtainedfrom Cr-Mo-V forgedsteel s pe ci men containing Ta nitri des

Fig. 47- Aspetto micrografico dei nitruri di Ti

rilevati sullo stesso saggio fucinato di cui alla

Fig. 46.

a

Table I - Ladle analysis of the three ingots relative to ee experimentsno.1-2-3

TAG. 1 Analisi chimica di colata dei tre lingotti dicui allot esperienze NS 1-2-3 .

ANAL IS! CHIMICA 14

TAB. Ill - Condiziont e risuttati degli esami ultrasonori

suite billette , in ammo tucinato, relative alto noniron 1-2-3.

ESPERIENZA NQl • Gillette diametro 1620 Fran.

Condieioni di *same: frequenza a 264Hz tipo sands e diame_

tro = 8 25 , Ø 24 mm altezza eco tondo (su zone nente da

anornalie)e 50 mm.; estensiono dell'esame a 100 % delta super _

I ici in dirnione .( drools sgrossatura )

ORisultato dell'iname anomalie conlenute antra la WO

ri tratteggiata provenienti da difetti

tipo porositt

ESPERIENZA Nt 2 - Billelta diarne ro 1170 mm.C`0 Condizioni di name : in tutto uguali a quell* relative

•H 1 alresperienza Nt 1 .ce

o •u) coco 0S-41-1 (A

a)E

•r-i4->

a)c - o a) Risultata dell'esame: anomalie da ditetti tipo porositi canto_

nute entro la zona tratteggiata.

ESPER1ENZA Ns 3 - Gillett& diametro 1670 mm.1-11-1 Condizioni di esame in tutto ugualt a quell< relative

all'esperienza NS1.

rdoT

Risultato dell'esame anomatie contenute entro la zone

tratteggiata provenienti da (Weill

!Ito porositi.

ESFEWIENZA 1"22I

II RItable 1115°C 60 h.

21101scatde 1220+1230°C it 62 ItSkoloatoon

3)1315c.111, 1220 4.12X° C tOOk.

4) RINcebtak 1220

Siontzeoure S oliolonc can On, 41444

SI FtIscoIde ¶220°C i 62 It I 51RIsc Ida 1210°C &nth<4 tangs con lump

ofiasoconaa cea •lenp.

1111L.:LH4) Maul& 1220a1220°C )12511,

1

-Sbcolanots .16eno•

Pin

•14•Volerts • 4.4,0 4444 4 14441411• V

TM. II CICLO 01 FUCINATURA DEI YDE 1.1N10111 DI WIAL6O ESPIRIV426 N2 1 2 3

ESPERIEW/A 1402 ESPEPIENZA 1423

11 RIsealdo 1100•C 5 10 h. 11 Rinaldo 1150 C • 2511.

2/RIsca169 1110°C c-Sbeactivn

3) R14caldt) 1220°C i 36-FhcaOcancra

\

r.a.

°C 1211. 4)(dIstaIde 1100°C I 40 h.o Stanch. an stamp

5) filessIdo 1200°C :42 h. 5101m:11a 1200 - C 26 I.

o'Stornivio can •tomp: ontni. .nonotwil. ton clomp

E-77-

• no, •••-•

.104" .1122

•Igeolsatua 0.1170 conisclro

21 RIsceItlo 1180 °C tub.-Soc /aka.

R/seldo 1220°C i 20 h.

4/RIscaldo

-54c4-n4t4. 0.nr oma, • •••• ••i•

oSgiongoin 4 0010 •••• •

auziats

Mkt

RIFUCINATIMA

71 RIses14. 1230°C 54011,

-11n0nkolola

aticdcoint.

IDRIscalclo 1220°C i22 h.

11) Riscalde 1200•C 4211.

-14.11Mort

- Steatitic. cc* starnop

101RIscaldo 1220°C 11411.

•Sbomtucli 10A .411Ingi

•Ricoliwo:

Nit

•Ssonicauca • coetnallo

itt

-

chimica

Table IV — Chemical composition of some steels with more critical endforging temperatures

Operazione

Analisi

Temperature

di tine

tucinatura

Ciclo di

Ricottura

TAB. IV - Composizione chimica di alcuni acciai con temperature

di tine tucinatura piu critiche.

TAB. V - Temperature di fine tucinatura e ciclo di ricottura

seguito sui tucinati di cui alle tigg. 30 e 31

Billetta da 250 x 250 mm. Fucinato per rotore generatore in

in acciaio PRESTEM . acciaio Ni-CbMo4 del peso di 50 T.

C= 0,23 IC; 5i=0,37 Sr0,014 ;

chimica P=0,008 %;Mn=0,78%;Ni=345%;Mo=3,20%•

850 ° C 770 ° C

1050°C

s0°C/Rattreddamento in

°C/

aria tino a tern._ 620°C

peratura ambiente

C=0,25 04 ; Si =0,04 04 S=0,009 04;

P=0,008 %; Mn=0,38 04 ; Cr=0,84 %

Ni=3,56 04 ; Mo=0,50 04; V=0,10 %

sscpc

300

900°C

670°C

"It

Table V — End forging temperatures and annealing cycle relative to theforging of fig. 30 and 31

SOURCE OF INCLUSIONS IN FORGING INGOTS

by

R. B. SnowU. S. Steel CorporationResearch Laboratory

Abstract

A review of the literature suggests that the relatively largeinclusions occasionally found in the body of forgings consist ofthe products of oxidation of the steel combined with some refrac-tory material. A study of the inclusions in a large as-cast ingothas revealed no such inclusions in the body of the ingot. However,inclusions large enough to cause the diversion or rejection of aforging were found in the frozen top crust and in a band near thebottom of the ingot. This study reveals that when the initial topcrust of the ingot has remained in place during the entire solidi-fication period, the forgings have been free of sonic indicationsof inclusions.

Many of today's problems with forgings have plagued steel-makers and forgemasters for years, and the problem of oxideinclusions in the body of ingots has caused many diversions andrejections. Leach,l)* in a paper delivered before the Clean SteelConference at Balatonfured, Hungary, in 1970, reported that approxi-mately half of forging rejections are caused by inclusions.Howorth,2) in 1905, reported fibrous tensile fractures that repre-sented reduced ductility in both the muzzle and breech portion offorged gun barrels. With the aid of a petrographic microscope,particles removed from the fractured surfaces with a needle wereidentified as manganese silicates. To explain the banding, or whatwould now be called inverse-V segregation, Snead3) postulated in1907 that currents formed in liquid steel during solidification andthat the flow was upward along the solidifying faces of the ingotand downward along the central axis. It is presently believed thatthe opposite is true, that the currents are downward along the faceof the ingot and upward in the axial center.4) Although the problemof the inverse-V segregation still persists, this paper will be con-cerned with oxide inclusions, which include silicates and aluminates.

Although Hibbard5) reported the presence of silicate scum onthe surface of ingots in 1910 and McCance6) determined the rate ofrise of silicate particles in steel in 1918, it was Dickinson,7)in 1926, who showed exceptional insight into teeming and solidifi-cation processes that might be related to the presence of silicatescum on the surface of the steel during teeming. He pointed outthat scum could be carried downward again by the steel stream andthat the recirculated particles could be entrapped in the growingdendrites to later appear as subsurface defects. He also postu-lated that the scum was the result of silicate agglomeration duringteeming and that air oxidation of the teeming stream probablyincreased the amount of inclusions in the steel. In addition,inclusions that floated upward after the ingot was teemed werecaught in the underside of the freezing top crust and in thebridging dendrites that formed beneath the top crust as solidifi-cation continued; he stated that inclusions in the bridging dendriteswould give rise to inclusion concentrations when there was an insuf-ficient top discard. Dickinson also extracted inclusions from thesegregated portion at the bottom of the ingot, and by their chemicalcomposition, showed them to be essentially composed of Mn0 and Si02with little contamination from ladle refractories.

* See References.

Introduction

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During the past 45 years, although steelmakers have attemptedto control inclusion formation in steel by furnace and ladle deoxi-dation practices, inclusions have persisted. In the bottoms ofbig-end-down ingots, the persistence of inclusions suggested thatoxygen dissolved in the steel precipitated as oxide inclusionsduring solidification and that these inclusions were then pushedto the center of the ingot by growing dendrites, where they agglom-erated to form large particles. Such a mechanism for macroscopicinclusions appeared to have merit because the inclusions usuallyhave the microscopic appearance of the products of deoxidation.

In the Clean Steel Conference in 1962, Richardson8) describeda radioactive tracer technique which showed that 10 to 20 percentof the large inclusions were of refractory origin. The occurrenceof refractory in inclusions has been confirmed by Charles andSalmon-Cox9) and more recently by Pickering.10) Leach1) reportedthat the major type of inclusions responsible for the rejectionof forgings were exogeneous and represented metal and/or slag-metalreactions with ladle refractories. Comon and Bastien,11) in 1967,plunged a capsule of radioactive iridium through the freezing topcrust of a vacuum-cast forging ingot and subsequently found a non-radioactive piece of the top crust lying on the nonradioactivecolumnar dendrites on the bottom of the ingot. These investigatorsalso found a group of macroscopic inclusions associated with den-drites that had formed in the cooler regions of the ingot and postu-lated that the dendrites had then fallen to the bottom. On theother hand, Nepper and Laubin,12) using the same techniques, found"oxidized" macroscopic inclusions in the bottom of one vacuum-castingot and concluded that these particles represented the combina-tion of the products of deoxidation and silicates from near theladle bottom which entered the mold during the initial stages ofteeming.

Because most of the inclusions in the body portion of ingotsthat subsequently cause the rejection of the forging are so largethat they should have floated out of the steel soon after teeming,a petrographic study has been made to determine (1) the identity,(2) the composition, and (3) the distribution of the inclusionsfound along the center line of a 134-inch-diameter, 150-inch-highingot. These results have been correlated with inclusions found inforgings and forged blooms.

Ex erimental Procedure

The ingots in this study required the product from fiveelectric-furnace heats that were teemed in succession at an optimum

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temperature into a pony ladle while an uninterrupted stream of steel

flowed into the mold in the vacuum chamber. The specimen identifi-

cation for this study is given in Table I with the pertinent pro-

cessing variables; the nominal steel composition (in percent) was

as follows:

Mn Si Ni Mo V Al

0.24 0.40 0.24 3.5 0.2 0.1 residual

Table I

Specimens Examined From 134-Inch-Diameterb 150-Inch-Hi h Vacuum-Cast In ots

Time FinishHot To Pour to

Covering Covering,S ecimen Lining Material minutes

During and after the finish of teeming of Ingot A, specimens

of the pony-ladle slag and the mold scum were collected for chemical

determinations. After the ingot had cooled, the 36-inch as-cast

top-discard portion was saved, and seven 1- by 3/8-inch sectionswere cut from the center 7-inch-thick top crust to determine the

composition and distribution of inclusions that floated out of the

steel during the initial portion of the approximately 50-hoursolidification period.

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The as-cast 134-inch-diameter Ingot B was sectioned alongthe vertical axis, and a 3/4-inch-wide strip for examination wascut from within 8 inches of the center line of the ingot, Figure 1.In addition, three sections were cut from the top crust near theedge of the ingot. The top crust was missing at the center of thetop surface of the ingot and in its place a layer of burned hot-topping compound containing thin layers of steel rested directlyon the steel surface. The sections in the center of the hot-topsection were therefore cut from the underlying continuous steellayer. The sections examined from Ingot B were oriented as follows:

Sample Location, inchesabove bottom of in ot

* Approximate location on Figure 1.

Dimensions of SurfaceNo. of Examined, inchSections Vertical Horizontal

In the body of Ingot B, 14-1/2 to 148 inches above the bottom, the49 polished sections were taken at selected locations along theslice; the distance between polished sections did not exceed 3 inches.The location of the sections examined and the number of inclusionsin each size range for the 88 sections examined are shown inFigure 1.

After ultrasonic inspection showed a serious discontinuity inthe bottom journal of a forging from Ingot C, a 1-inch-thickslice (a) was cut from the 31-1/2-inch-diameter journal. Aftersurface grinding, a single macroscopic inclusion was found approxi-mately 4 inches from the axis of the forging. The inclusion was cutout, and the top surface as well as longitudinal sections adjacentto the inclusion were examined. The general cleanliness of thejournal was determined by examining thirty 3/4- by 1-inch sectionstaken from a 3/4-inch-deep cut through slice (a) that passed throughthe center of the journal and within 2 inches of the inclusion; thesections represented the entire cross section of the journal.Ultrasonic inspection also revealed a defect region 15 to 20 inchesbeneath the side and 90 inches from the bottom of a 370-inch bloomthat was forged from Ingot D. This portion of the bloom was sec-tioned to determine the nature of the defect.

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This study included the petrographic identification of theinclusions and electron-probe microanalysis to establish theirchemical compositions. Inclusions in selected specimens wereextracted by the methanol-halogen technique, and their compositionswere determined by x-ray fluorescence.

The segments between the polished sections from the body ofIngot B were surface-ground, examined for macroscopic inclusions,and then etched in a 2 percent nital solution to reveal any macro-scopic dendrites. To determine the size distribution of inclusions,the number of oxide inclusions in each of four size ranges werecounted visually at a magnification of X200 on 0.375-square-inchpolished sections from Ingots A and B. The four size ranges were(1) greater than 100 microns, (2) 50 to 100 microns, (3) 15 to50 microns, and (4) under 15 microns. The inclusion size did notinclude the manganese sulfide rim, and manganese sulfide inclusionswere not counted.

The reproducibility of the techniques was tested by a redeter-mination on 10 of the sections from the body of Ingot B, whichshowed a change of less than two inclusions in the smallest sizerange. The size distribution was also determined on ten duplicatespecimens cut adjacent to the original slice and on 12 sectionsafter resurfacing. The data for the duplicate and resurfaced speci-mens agreed closely with the first data set except for local varia-tions that will be referred to later in the text. In addition, thesize distribution of inclusions within the steel matrix was deter-mined on nine sections after etching in a 2 percent nital solution.

Petrographic Examinationof In ots A and B

Observations and Discussion

The petrographic examination of the inclusions in Ingots Aand B, Figure 1, showed the occurrence of two distinct types. Allthe inclusions larger than 50 microns and some of the smaller inclu-sions contained medium-gray crystals of a spinel in a complexsilicate matrix, Figure 2. The electron-probe microanalysis showedthe spinel crystals to contain only Mg0 and Al203 and the matrixto contain CaO, Si02, and some Al203 with little Mg() and no K20or Na20. All these inclusions showed a nearly continuous manganesesulfide rim. However, most of the less than 50-micron inclusionsappeared in the sections as a dark-gray silicate with a continuousmanganese sulfide rim, Figure 3. The electron-probe microanalysisshowed these inclusions to contain Si02 and Al203 with trace amountsof other oxides.

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In.ot A. The occurrence of Ca() and Mg0 in the large inclusionsin these ingots is unusual because calcium- or magnesium-bearingalloys were not added to these heats. One method to determinewhether such inclusions are process-oriented is to compare theircomposition with that of the pony-ladle slag and the silicate moldscum that often floats on the surface of the steel after teemingof the ingot. The process-oriented inclusion composition inIngot A was obtained by extracting the inclusions in the 1/2- to2-inch layer beneath the top skin of the ingot. The compositionsof the scum and the inclusions were similar, Table II.

Table II

Composition of Pony-Ladle and Mold Scum FromIn ot A and Extracted Inclusions From In ots A and B

Visually the top crust of the ingot appeared to be in placeafter the ingot had cooled, and the following distribution ofinclusions in the top 7-inch-thick top crust of the ingot confirmedthe visual observation.

Number of Inclusions inDepth Below Each Size Ran e, micronsthe Surface, Less Greaterinches Than 15 15 to 50 50 to 100 Than 100

S ecimen Si02 Al203 Ca0 Mgt) Mn0

Ingot A, pony-ladle scum 30 6 50 9 1

Ingot A, mold scum 31 30 20 8 6

Ingot A, extracted inclusions -top crust

37 30 22 11 0

Ingot B, extracted inclusions - 38 23 26 13 0A-1

Ingot B, extracted inclusions - 35 31 23 11 0A-4

0 to 1 60 77 101 51 to 2 72 107 67 22 to 3 53 85 68 3

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Number of Inclusions inDepth Below Each Size Ran e, micronsthe Surface, Less Greaterinches Than 15 15 to 50 50 to 100 Than 100

These data suggest that the pony-ladle slag represents ladleslag that was carried into the pony ladle with the steel streamnear the end of teeming of one or more of the five heats. Some ofthe pony-ladle slag was then probably carried down into the steelin the pony ladle by the action of the steel stream, and some ofthe slag probably entered the mold with the steel. This mechanismwould account for the Ca0 and Mg() found in the mold scum, as wellas in most of the inclusions found in Ingots A and B. Prior to thisinvestigation, Ca0 and Mg0 had not been found as major constituentsof inclusions without additions of calcium or magnesium, and thesedata indicate that the composition of inclusions may be dependenton the teeming practice. These data further suggest that the dis-solved oxygen in the liquid steel may have caused the smallersilica-alumina inclusions to precipitate from the steel during thesolidification process.

Ingot B. Comparison of the ladle analysis of the heats teemedinto Ingot B and the compositions of steel drillings from the edge,quarterpoint, and center of the ingot body showed little positiveor negative chemical segregation, although the center bottom of thehot-top section showed a positive segregation of 0.06 percent carbon,0.03 percent manganese, and 0.006 percent sulfur.

Distribution of Inclusions. An exceptionally large number ofinclusions was found in the 1-1/2-inch-thick top crust from theedge of the ingot, which remained attached to the top sidewall ofthe ingot, Zone A-1, Figure 1. The absence of top crust at thecenter of the hot top was verified by the fact that no inclusionswere found in the 9-inch-thick top layer of steel, Zone A-2, fromthe center of the ingot and just beneath the interface layer ofoxidized exothermic material and steel. The steel from the bottomcenter of the hot-top region, Zone A-3, contained only a few inclu-sions. A second narrow 2-inch band of inclusions was found atZone A-4, located 10-1/2 to 12-1/2 inches from the ingot bottom,and a deep-etch test showed this band to extend horizontally beyonda 12-inch-wide slice. Some of the inclusions in Zones A-1 and A-4

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had a maximum size of 300 microns, and all inclusions in these twozones were petrographically similar. The compositions of theextracted inclusions, Table II, were similar to the compositions ofthe mold scum and the inclusions found in the top crust of Ingot A.

Evidently the band of inclusions in Zone A-4 represented themissing piece of top crust, which probably broke into relativelyflat pieces when the exothermic layer was placed on the frozen topsurface of the steel some 35 minutes after the end of teeming.These flat pieces probably sank rapidly to the bottom of the ingotwithout remelting. This appears to be the only mechanism thatexplains such a narrow concentrated band of inclusions that containmore Ca0 and Mg0 than would be expected from the steel composition.

Plus-50-Micron Inclusions in In ot Bod . Excluding the inclu-sion in Zone A-4, only four sections of the 65 body specimens con-tained an inclusion larger than 100 microns, and the maximum sizewas 120 microns, Figure 1. A macroscopic dendrite that containedone of the plus-50-micron inclusions is shown in Figure 4. Whenthree sections that each contained one of the plus-100-microninclusions and nine additional sections were resurfaced and re-polished, no inclusions in this range were observed. Nor were anyinclusions in this size range found in the ten duplicate specimensthat were studied. Only 17 of the same 65 specimens containedinclusions in the 50- to 100-micron size range, and specimens con-taining more than one were taken from the region 15 to 45 inchesabove the ingot bottom. In the etched specimens, all the inclusionslarger than 50 microns were found in ferrite grains, Table III andFigure 5. Visually these ferrite grains appeared to be the remainsof original macroscopic dendrites that had sunk to the bottom of theingot and had been only partially decomposed during the long coolingperiod. The inclusions were composed of spinel crystals in a com-plex silicate matrix. If these inclusions had precipitated andagglomerated as the liquid steel solidified, they would have con-tained little Ca0 and Mg0 because calcium and magnesium were notalloying elements.

Minus-50-Micron Inclusions in In ot Bod . The data in Figure 1incidate that the distribution of inclusions in the 15- to 50-micronsize range was relatively constant throughout the body of the ingot,excluding Zone A-4, and that from 1 to 15 such inclusions were foundin each section. However, the number of inclusions smaller than15 microns varied from 9 to 115 per section, and most of these werein sections located 25 to 90 inches above the ingot bottom.Stringers of 3 to 6 inclusions in essentially a straight line werenumerous in sections containing large numbers of inclusions smallerthan 15 microns. In the etched sections, those stringers wereusually found in the ferrite grains.

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The petrographic examination showed that, in general, thelarger inclusions in the 15- to 50-micron range, Figure 3A, anda few of the smaller inclusions were composed of spinel crystalsin the complex silicate matrix, as were the inclusions largerthan 50 microns. Thus, only the smaller inclusions were composedof essentially silica and alumina, Figure 3B, the expected generalcomposition for inclusions precipitated from the steel duringsolidification. The etched sections showed that many of thesealumina-silica inclusions were also in ferrite grains, Table III.

Table III

Size Distribution of Inclusions Found in Ferrite Dendrites,the Steel Matrix, and the Grain Boundaries of In ot B*

TotalNumber ofInclusions

* Etched in 2 percent nital solution.

The finding of more inclusions in the steel matrix in the top100 inches of the ingot while more inclusions were in the ferritegrains in the lower 50 inches, appears to support the theory ofinclusion entrapment in sinking dendrites. If the inclusions werethe controlling factor for ferrite grain development, there shouldbe no difference in the location of inclusions in the structure ofthis very slowly cooled ingot.

The many small alumina-silica inclusions, the relatively few15- to 50-micron alumina-silica inclusions, and the absence of

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even relatively large alumina-silica inclusions indicates littleagglomeration during the final solidification process of inclusionsthat precipitate from residual amounts of dissolved oxygen in theliquid steel. This mode of inclusion formation is probably not animportant source for harmful inclusions in the body portion ofingots. Furthermore, the growing dendrites probably did not forcelarge numbers of small inclusions into shrinkage cavities becausesonic tests of forgings from other ingots of this series have notshown the occurrence of silicate inclusions along the central axisof the forged steel.

Inclusion in Bottom Journalof For in From In ot C

After a macroscopic discontinuity was located by ultrasonicinspection in the bottom journal of a rejected forging fromIngot C, a 1-inch-thick transverse slice was cut from the 31-1/2-inch-diameter journal. After surface grinding, a single macro-scopic inclusion was found approximately 4 inches from the axialcenter; the inclusion thinned but extended entirely through the1-inch-thick slice. The remainder of the etched surface of theslice showed no unusual defects. The polished inclusion, Figure 6,showed 3Al203 2Si02 crystals in a complex silicate matrix, andthe group of small subscale-type inclusions around the large massindicates that oxygen diffused into the steel after the steel hadsolidified. The chemical composition of the inclusion was deter-mined by (1) averaging 11 determinations made at various positionson the polished surface of the inclusion with a scanning electronmicroprobe, and (2) inclusion extraction techniques on the lower3/4-inch segment of the inclusion. These determinations showedthe following composition:

Constituents, ercentSi02 Al203 Mn0 Mg0 Ca0 Ti02 K20

1

* Alkali cannot be determined by th's method.

The inclusion had a silica-to-alumina ratio of 1.5, which issimilar to that of fireclay refractories, and the presence of K20also suggests fireclay refractory. The Ca0 content indicates thatsome mold scum may also be associated with the inclusion, but thehigh Mn0 content and the surrounding subscale inclusions indicatean oxygen level that should not occur in vacuum-cast steel.

Scanning method 56 27 10 1 3 2Extraction method 45 31 17 1 3 3

Teeming records revealed that the upper "straight" section ofthe hot top, Figure 7, was filled to "T" which was 35 inches above"E." Shortly after teeming was completed and within a few minutes,the top crust showed the normal concave contour, "C." However,when the ingot was stripped, the original top crust was located ata point "Y," 20-1/2 inches below the original pour line. Therewas no visible evidence that the top crust had been lost and nometal skull was found against the fireclay hot-top lining. However,there could have been some disruption of the skin adjacent to point"E." These data suggest that the top crust of the ingot slid down-ward and caused some of the thin skull of frozen steel against thefireclay hot-top lining to loosen and fall into the body of theingot. Such a skull would be associated with surface scum andaltered fireclay brick, and subsurface oxide inclusions in thesurface layer of the skull would contain considerable MnO. Incontrast to the high Mn0 content of the inclusion, the mold scumtaken from the surface of Ingot A contained only 6 percent MnO.

The steel surrounding the inclusion both in a transverse andlongitudinal direction did not show any additional large inclusions.The 30 sections in the longitudinal cut through the center of the1-inch-thick transverse journal slice showed only 37 stringer inclu-sions greater than 100 microns in length. This steel was relativelyclean because no more than two stringers were found in any singlelocation and the 1000-micron stringer inclusion averaged 20 micronsin diameter, which would be the equivalent of an 84-micron-diameterinclusion in the as-cast steel. All the other stringers were lessthan 15 microns in diameter and would represent 25- to 50-micron-diameter particles in the as-cast steel. Additional inclusionsin each section averaged only 18 inclusions in the 5- to 25-micronsize and an average of 55 inclusions were less than 5 microns insize. The 3/4- by 1-inch specimens are twice the size of thesections examined along the axis of Ingot B and on the latter basisthere would be one stringer for each 1.6 section (1/2 by 3/4 in.).The high Mn0 content, the associated envelope of small subscaletype particles, and the sinking of the top crust of Ingot B madeit possible to pinpoint the probable source of this inclusion.

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Bloom Forged From Vacuum-Cast134-Inch-Diameter In ot D

Table IV

Number and Length of Stringer Inclusions Found in a 3/4- by31-1 2-Inch Slice Throu h the Bottom Journal For ed From In ot C

the forging).

Sonic inspection after the preliminary forging of a bloomfrom Ingot D showed a defect region 15 to 20 inches beneath the sideand 90 inches from the bottom of the 370-inch bloom. The hot tophad been lined with insulating brick, and a 3000-pound exothermiclayer had been used to cover the steel, Table I. A saw cut throughthe defect region showed numerous cracks, Figure 8, and the presenceof macroscopic silicate inclusions in steel adjacent to the defectsuggests a condition similar to that of Ingot B where large inclu-sions were found in the top crust at the edge of the ingot.Polished sections showed alumina particles in the steel along thecracks, Figure 9, and many of the cracks showed a thin continuouslayer of alumina particles.

Although the aluminum content of the steel was ordered asresidual and aluminum-bearing alloys were not used, a portion ofthe clean steel in the defect region contained 0.011 percent totalaluminum, whereas a sample taken outside of the defect region con-tained 0.004 percent total aluminum. Methanol-halogen extractions

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of the inclusion layers along two cracks revealed the followingrange of composition (in percent):

siO2 Al203 Ca0 Mg0 Mn0

1 to 4 80 to 85 7 to 10 4 to 5 1

An electron-probe microanalysis confirmed the presence of Al203particles and showed traces of Si02, CaO, and Mg0. A scanningelectron-probe examination showed no Na20 or K20. The refractoriesused in the hot top and the roof of the vacuum chamber containedmuch less Al203 than the inclusions, and fragments of refractorieswere not observed along the cracks. The most probable source forthese particles is the exothermic hot-topping material, which hadthe following composition range (in percent):

Si07 Al203 Ca0 Mg0

6 to 12 68 to 84 5 to 10 5 to 10

Since the top crust at the edge of Ingot B was up to 2 inches thick,a similar piece of top crust on Ingot D could have sunk into theliquid portion of the ingot late in the solidification process.The composition of the inclusions, the greater aluminum contentof the steel in the defect region, and the associated macroscopicsilicate inclusions indicate that the inclusion region is relatedto the addition of the exothermic mixture and the sinking of aportion of the top crust of the ingot.

General Discussion

Large forging ingots have been free of ultrasonic indicationsof nonmetallic matter when the top crust and sidewalls of theingot have remained intact. On two occassions visual observationsindicated the loss of the center region of the top crust eventhough the ingot had been covered with a brick-lined lid. Specimensfrom one of the, top crusts showed clean steel rather than theexpected inclusion layer, and inclusions were later found in theforging. Two types of inclusions were found and it was determinedthat the inclusions formed from dissolved oxygen in the steelprobably do not agglomerate readily during solidification. Inaddition, the inclusions in the bottom portion of the ingotprobably were contained in dendrites, and individual inclusionsmay have become entrapped in one or more dendrites that remeltedbefore reaching the bottom cone region of the ingot.

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The conclusions reached on large forging ingots have beenn further verified by the use of zircon (zrSiO4), vermiculitemica (complex silicate of K20, Mg0, Al203, and Si02) and zirconia(zr02)as tracer compounds on open hearth and BOP fine-grainedingots.12) These latter data have shown that the hot-toppingmixture and/or a layer of the top crust has been found as macro-scopic inclusions by the ultrasonic inspection of the plates orblooms rolled from the ingots. The findings were verified bypetrographic, extraction, and electron-probe techniques.

In the case of the plate steels, an increase in the hot-top volume and the use of very coarse vermiculite particles hasessentially eliminated diversions for nonmetallic particles. How-ever, when fine vermiculite was used as a hot-topping material,a number of macroscopic silicate particles were found in the fine-grained steel and the composition of the silicate was essentiallythat of the vermiculite.

In the BOP steel, ultrasonic inspection indicated macroscopicinclusions composed of aluminum and silica with an occasionalsmall concentration of K20 in blooms when a 1-1/2-inch nozzle wasused, but these inclusions were not found when the steel was teemedthrough a 2-1/2-inch nozzle. Thus, the increased air oxidationof the teeming stream, an increase in nozzle-bore erosion, andpossibly a lower effective temperature in the ingot caused theagglomeration of mold scum and its retention in the ingot. Ingeneral, the occurrence of macroscopic inclusions in the bodyportion of an ingot has been minimized when the steel is teemedat an optimum temperature and with a sufficiently large nozzleto produce an optimum rate of rise in the mold. This latter teemingpractice should also minimize surface defects in the steel. How-ever, a relatively long hesitation at the hot-top junction duringteeming and then filling the more constricted hot-top region witha reasonably full stream can drive most of the collected scum backdown into the steel, where it can become incorporated on a rela-tively thick layer of steel dendrites. The top crust of ingotscan also be lost during the movement of ingots from the teemingplatform.

Mold scum incorporated within an ingot is often difficultto identify because of its conglomerate nature. This materialoften contains small rounded droplets of steel, and the compositionmay vary from that expected from the steel composition and theladle additions to particles of pure refractory, as in the caseof the large vacuum-cast ingot where the Ca0 and Mg0 content prob-ably result from the nature of the teeming procedure. In general,

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the finding of 1<20 in the inclusion suggests the presence of ladle

refractories, but in top-poured ingots 1<20 is readily volatilizedat steel temperatures and the oxide is readily reduced to an evenmore volatile metal phase. Thus, in top-poured ingots, K20 and

Na20 (from the liquidizer) probably indicate mold scum.

Any relatively large particles of nonmetallic matter thatshould have floated rapidly out of the steel probably resultfrom the collapse of the top crust or sidewall of the hot top.Smaller particles may represent minute particles that floatedslowly out of the steel, perhaps because of a relatively loweffective temperature in the steel after teeming. Such an inclu-sion can agglomerate by sinking and rising until it finally reaches

a resting place at the base of the liquid portion of the ingot.This work suggests that large inclusions are not a result of theagglomeration of inclusions formed from dissolved oxygen duringthe solidification of the steel.

Summary

Inclusions in a large as-cast ingot were found to be relatedto the sinking of dendrites or pieces of the top crust of the

ingot. Inclusions in blooms and forgings have been related tothe collapse of the top crust or the sidewall of the hot-top region.

Ultrasonic inspection of forgings has not revealed the presence of

inclusions when the top crust and the sidewalls of the ingot haveremained intact. Small inclusions that may have precipitated from

the liquid during cooling of an as-cast ingot did not agglomerate.

In top-poured ingots the presence of 1<20 or Na20 in an inclusionprobably indicates mold scum rather than the agglomeration ofparticles during the solidification of the ingot.

References

1. J. C. C. Leach, "Macro-Inclusions in Forging Ingots," ReviewJournal ISI, 209, 944, 1971 (Clean Steel Conference,Balatonfured, Hungary, to be issued as a special report, ISILondon, 1972).

2. H. G. Howorth, "The Presence of Greenish Colored Markings inFractured Surfaces of Test Pieces," Journal ISI, 68, 301, 1905.

3. J. E. Snead, "Segregation in Steel," Engineering Conference ofthe Institution of Civil Engineering (England) , Paper No. li,1907.

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References (Continued)

4. R. J. McDonald and J. D. Hunt, "Fluid Motion Through thePartially Solid Region of a Casting," Trans. AIME: 245,1993:1969.

5. H. D. Hibbard, "Solid Nonmetallic Impurities in Steel," AIME,41, 803, 1910.

6. A. McCance, "Nonmetallic Inclusions: Their Constitution andTheir Occurrence in Steel," Journal ISI, 97, 239, 1918.

7. J. H. S. Dickinson, "A Note on the Distribution of Silicatesin Steel Ingots," Journal ISI, 113, 177-211, 1926.

8. H. M. Richardson, "The Use of Radioisotopes to Trace the Originof Oxide Inclusions in Steel," Clean Steel Conference, SpecialReport No. 77, ISI London, 1963.

9. P. H. Salmon-Cox and J. A. Charles, "Further Observations onthe Analysis and Distribution of Nonmetallic Inclusions in a0.26 Steel Ingot," Journal ISI, 203, 493, 1965.

10. F. B. Pickering, "Effect of Process Parameter on the Origin ofInclusions," Review Journal ISI: 209, 943, 1971 (Clean SteelConference, Special Report, ISI London, 1962).

11. J. Comon and P. Bastien, "Experimental Study of the RelationshipBetween the Solidification and Heterogenity of Steel IngotsFrom 3 to 30 Tons," Rev. Met., 65, 13-24, 1968, BISI Translation6732.

12. M. Nepper and M. Laubin, "Heterogeneity of Hard Steel ForgedIngots," IBID, 1968, 65, 25-34, 1968; BISI Translation 6736.

13. R. B. Snow, "Inclusion and Their Relation to Solidification inthe Hot-Top Region," (to be published in the Proceedings ofthe Open Hearth - Basic Oxygen Steelmaking Conference, AIME,Chicago, Illinois, 1972).

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FIG. I SIZE DISTRIBUTION OF INCLUSIONS IN A LARGE AS-CAST INGOT --DATA

FROM

88 0.375-SQ-IN.

(242-SQ-MM) SECTIONS LOCATED WITHIN

8INCHES

(200 MM) OF CENTER LINE OF THE INGOT.

FIG. 2. Inclusions averaging approximately 20 and 100 microns in sizewere found in a section taken 63 inches above the bottom ofIngot B. The inclusions comprised angular MgO.Al203 crystalsin a silicate matrix of CaO, Al203, and Si02 surrounded by asulfide rim. Section etched in 2 percent nital solution toshow the ferrite and transformed bainite. X350.

FIG. 3. Typical silica-alumina inclusions found in Ingot A and B,less than 15 microns. X700.

FIG. 4. A 100-micron inclusion was found in this macroscopic dendritefrom the body of Ingot A. Etched in a 2 percent nitalsolution. X2.

FIG. 5. Inclusions were entrapped in ferrite grains in a steel matrixof coarse bainite, which formed during the slow cooling of theas-cast ingot. Section etched in 2 percent nital solution.X50.

FIG. 6. A silicate inclusion with silicate subscale inclusions in thesurrounding steel found in the bottom journal of a forgingfrom Ingot C. The lighter grey needle-like crystals aremullite and the black areas are holes; a few steel particlesare visible within the silicate mass. Xl00.

520 mm(20-1/2 inches)

FIG. 7. Sketch showing metal level after pouring to point T and theconcave surface C that formed during the first hour ofcooling. After complete solidification, the top crust wasformed at Y. Some skin distortion was observed at point E.

FIG. 8. Successive saw cuts through the defect area in Bloom Crevealed cracks and round silicate droplets (arrow in lowerpiece). X1/2.

‘S

FIG- 9. Alumina particles were found in one of the thinner stringersin Bloom C. The black areas are holes. X500.

INFLUENCE OF "A" SEGREGATIONS IN THE MECHANICAL PROPERTIESOF FORGINGS OBTAINED FROM VACUUM POURED INGOTS.

S. FERNANDEZ

Astilleros Españoles S.A.-Factoria de Reinosa.-SPAIN.

ABSTRACT

Forgings obtained from vacuum poured ingots show "A"segregations which are similar, and sometimes more noticeable,than those which are observed in non-degassed material; themain difference is that discontinuities are not detected byultrasonic tests in the former case.

Sulphur prints and acid etchings are presented showing "A"segregations in castings and forgings. The morphology of "A"segregations in castings suggests an explanation for theirintensity in machined forgings from vacuum poured ingots and fortheir ocurrence nearer to the surface of the ingot. Thepredominance of type II sulphides in the segregations may alsomake them more intense.

Measurements have been made of mechanical properties inthe "A" segregation zone and out of it, both with longitudinaland transverse test bars. A decrease of mechanical propertiesfrom the surface to the "A" segregation zone has thus beenquantitatively determined for a total of 49 forgings fromvacuum poured ingots.

The loss in mechanical properties is always greater fortransverse test bars. In the case of carbon steels losses inelongation, reduction of area and impact strength are studiedvs. tensile strength in the range 45-77 kg/mm2. Bigger lossesrecorded are of 50% for reduction of area, 40% for elongation,and 30% for impact strength at Al 70 kg/mm2.

- 1 -

INTRODUCTION

Of the metallurgical problems posed by large forgings,one of remarkable interest is the lack of uniformity inmechanical properties due to segregations in the ingot: thesegive rise to a decrease in ductility as measured by elongationand reduction of area in tensile tests and by impact tests.There is one particularly interesting type of segregation withwhich we shall deal here: "A" segregations in ingots and,consequently, in forgings.

"A" segregations are visible to the naked eye in partswhich have been subject to a deep machining process (such astwisted crankshafts or the area where the flange joins theshaft in large forgings) . It is generally accepted that theyproduce extremely undesirable effects on mechanical properties,and such was our experience with forgings from non-degassedsteel without special heat treatment for hydrogen removal(1).But we also found them in pit-cooled forgings obtained fromvacuum poured ingots(*); "A" segregation lines were againclearly visible and, occasionally, even wider than before.There was, however, an important difference:with non-degassedsteel, small discontinuities were detected by ultrasonic testswith a 4 MHz probe in the "A" segregation region(even for lowcarbon steels of 0,20% C) , whilst with vacuum degassed steeldefects were no longer detectable by this method. One wouldthus conclude that "A" segregations were now free frommicroshrinkages or hair line cracks,because of being welded inthe forging process or because the content in hydrogen was verylow. But, even though "A" segregations do not involve cracksin vacuum degassed steel, it is not uncommon that doubtsarise in the customer's mind as to the quality of the part,when segregations are readily observed in deep machineditems such as the above mentioned.

This work was undertaken, precisely, to ascertain theimportance of the loss in mechanical properties when "A"segregations are present in forgings from vacuum poured ingots.We hope it may mean a small contribution to a better knowledgeof "A" segregations and their influence on the mechanicalproperties of forgings; we are thankful for the opportunity ofbeing able to divulge our results to you and would certainlywelcome any suggestions that you may have to offer.

(*) All our steels, before and after vacuum facilitieswere available, have been killed basic electric steels. Ingotswere practically the same in either case, with similar contentsin residual elements such as sulphur and phosphorus. Vacuumequipment was installed in our factory in 1966. We were thusable to pour ingots of up to 140 Tm in a vacuum chamber. Morethan 1000 vacuum poured ingots in the 35-60 Tm range havebeen obtained since 1966.

- 2 -

"A" SEGREGATIONS AND THEIR INFLUENCE ON MECHANICALPROPERTIES.

"A" se re ations

"A" segregations have the appearance of dark ribbons insulphur prints and are clearly visible after etching with cupricreagents; they occur as ring zones, the diameter of these ringsdecreasing towards the top of the ingot or casting. Fig. 1shows the location of different types of segregations in aningot. Most ingots and castings are affected by "A" segregations,but these are only readily visible for sizes bigger than, say,400 mm. in diameter.

In longitudinal sections, they show up as ribbons ofvariable width and length, depending on the size of the ingotor casting; they may be of several millimeters in width andcentimeters in length, and are wider and longer in the upperthird (*) . In transverse sections they have an almost circularshape of several mm. in diameter (Figs. 2 & 3).

COLUMNAR ZONE

Fig. 1

SEGREGATIONS IN AN INGOT

--SHRIN KAGE

"V .SEGRAGAll pON

- SEGREGATION

NEGATIVE

GEGREGRAT ION

- 3 -

Macroscopically,the widthof "A" segregations does notshow a great variation alongits length, but they tend to bethinner at their base and widerat their upper part. Theoutline is more clearly definedand smoother towards the outerside than towards the innerside of the ingot or casting.(Fig. 4).

Methods of observationsuch as sulphur prints oretchings with cupric reagentsare only qualitative; othermethods become necessary inorder to gain insight intothe chemical and structural

(*) The actual length is,however, not accurately knownand most of the segregationsare probably longer than theyseem to be: it would becertainly difficult to cut aningot in such a way so as toprecisely intercept the pathfollowed by the segregations.

0.1

01-4

1-0

1-4 ts

Z4

$40H

Z

ellE

l 0 4

C.) H

Fr.1Urzi o

I-1co4

a)z rta

roH

Cl)

0 r»

E9

)-1H

(1)

•C

_D4-1

H

Q.)

it 4

4-1<

14

(a)Fig.4

- 6 -

(b)

(a) SULPHUR PRINT AND (b) STEAD REAGENT ETCHING OF AN "A"SEGREGATION SHOWING HOW OUTLINE IS BETTER DEFINED TOWARDSTHE OUTER SIDE OF THE INGOT (LEFT). (x 1.)

form of the elements present. Figs. 5 & 6 show manganesesulphides in longitudinal and transverse sections of an "A"region; in the latter the eutectic origin of the sulphides ismore clearly appreciated. In all the forgings of which mechanicalproperties were studied we observed that, in "A" segregations,sulphides belonged almost invariably, using Sims' classification(2) , to type II; this was perhaps due to relatively high localcontents in sulphur and to the level of residual oxygen. In thesurrounding material most of the sulphides were also of type II,but there was some 20% of type III. It should be noted that allthe steels tapped were killed with ferro-silicon in the furnace,and that no aluminium was added either to the furnace or to theladle. We would thus conclude that, in the procedure followedby us, the residual content in oxygen was adequate for theformation of eutectic sulphides.

Another aspect is hardness. Vickers hardness tests showedthat "A" segregation regions were always harder than thesurrounding material; and carrying out microhardness tests

Fig. 5

TYPE II SULPHIDE INCLUSIONIN "A" SEGREGATION(LONGITUDINAL SECTION).(x100)

within segregations we found noticeable differences in hardness,due perhaps to differences in segregation thickness. However,some of the readings revealed zones which were softer than theregions without segregations (Fig. 7): we suppose that this isdue to the presence of sulphide inclusions. We also observedthat whilst for 0,45% C steels the average difference in Vickershardness between "A" segregations and the matrix was of 20 RV,in quenched and tempered NiCrMo and CrMo steels the differencewas of 10-15 HV.

Microprobe analyses revealed differences in compositionbetween regions with and without segregations; an analysis,showing slight differences in composition in the case of aNiCrMo steel is the following:

Mn(%) Ni(%) Cr(%) Mo(%) Si(%)

- 7

Fig. 6

TYPE II SULPHIDE INCLUSIONIN "A" SEGREGATION(TRANSVERSE SECTION).(x100)

"A" segregation: 0.60 3.4 0.65 0.41 0.20Out of the seg.: 0.71 3.1 0.60 0.41 0.20

260

260

270i

270

265

260

NiCrMo STEEL

- 8

270

x5

25570

26

290280285290

Fig. 7

VICKERS TESTS WITHIN AN "A" SEGREGATION. (x5)

(a) Fig. 8 (b)MICROSHRINKAGES IN AN"A" SEGREGATION.(a) NITRICACID ETCHING (x 1). (b)STEAD REAGENT ETCHING.

Together with "A" segregations, andoccur next to the feeder head, there mayvisible to the naked eye or with the aid8 & 9 show, at different magnifications,which appeared in the segregations which

- 9 -

Fig. 9MICROSHRINKAGES IN AN "A"SEGREGATION. STEAD REAGENTETCHING (x 5),

especially when theybe microshrinkagesof a microscope. Figs.some of the shrinkagescan be seen in Fig. 2.

The study of segregations in castings is interesting, notonly as far as soundness is concerned but also from the point ofview of their proximity to the surface. In cases like the oneshown in Fig.3, in which solidification was symmetrical, "A"segregations were also symmetrical; but in Fig. 2, whichcorresponds to a ring of 1900 and 550 mm. external and internaldiameters respectively, cast with a central core of 550 mm. (*),

(*) Average chemical composition, in samples taken at 100mm. from the surface, was the following:

C S P Mn Si Ni Cr MoPercentages .19 .021 .020 1.38 .37 .30 .30 .06

H2=3 p.p.m.; 02=29 p.p.m.; N2=77 p.p.m.

outer segregations can be seen to be more intense, more tiltedtowards the feeder head and more numerous than inner segretations(the also involve more microshrinkages) ; inner segregations are,in turn, wider, nearer to and more parallel to the surface. Thiscould very well be due to slower solidification because of thecore; and it suggests an explanation to the fact that in forgingsfrom vacuum poured ingots segregations appear as visible as -oreven more visible thanin non-degassed material: There may be asimilarity between the solidification of a casting, such as theone in Fig. 2, and the solidification of vacuum poured ingots,in which segregations are also wide and near to the surface.Apart from that, if type II sulphides predominate, as in thecase of our forgings, "A" segregations will be more intensethan if we only had type I or III sulphides.

* * *

"A" segregations have been considered, for many years now,as important internal defects; a considerable amount of researchwork on their origin has been carried out, and a number oftheories has been proposed concerning their growth mechanism.Many authors have reviewed the subject, and new bases for anexplanation are frequently established.

Mechanisms have been suggested in which the ultimate causefor "A" segregations would be the evolution, through the regionof dendritic growth, of gas bubbles trailed by impure metal(3-5) ; but, as we already pointed out, "A" segregations arealso present in vacuum poured ingots.

Other theories base their explanation on differentassumptions (6-10) , and many agree in assigning a decisive roleto convection currents in the molten liquid; according tothem, convection currents give rise to nucleations ahead ofimpure regions, which thus become entrapped. In any case, anyexplanation of the growth of "A" segregations requiresconsidering the nature of the solidification front, becausethese defects seem to be intimately related to the arrest ofcolumnar growth and the outset of equiaxial growth.

"A" segregations and mechanical pro erties

For the comparison of mechanical properties in "A"segregation zones with those at the surface (which we shallcall the "S" zone for simplicity)we used pit-cooled forgingsobtained from vacuum poured ingots, with a mean diameter of1500-1700 mm. and a length-to-diameter ratio of 1.7. Piecesof 600-800 mm. diameter were taken from the forgings aftercutting off the feeder head and some 5% of the body. Forgingreduction ratios ranged from 3.5 : 1 to 6.5 : 1.

- 10 -

Fig. 10SULPHUR PRINT SHOWING S AND A REGIONS FROM WHICHLONGITUDINAL (L) AND TRANSVERSE (T) TEST BARS WERETAKEN.

"A" segregations do not occur at the same position inforgings coming from different ingots, but differences inposition are slight. The exact location was determined in eachcase by sulphur prints of transverse sections of the forgingsof which mechanical properties were to be determined.

Fig. 10 shows "S" and "A" regions from which test barswere taken for tensile and impact tests. Rough machined testbars were normalized (carbon steels) or quenched and tempered(alloy steels).

Elongation, reduction of area, and impact strength dependon tensile strength, and at the "A" segregation zone tensilestrength is somewhat higher because of positive segregation.This variation is usually higher for normalized steels thanfor quenched and tempered steels.Our aim being to compareelongation, reduction of area, and impact strength between"A" and "S" zones for a given tensile strength, we discardedthose materials in which the difference between tensilestrengths at the "A" and "S" zones was higher than 6 %.

A total of 49 forgings has been studied, from which 41are carbon steels and the rest CMn, CCr, CrMo, and NiCrMosteels. Sulphur percentages ranged from 0.015 to 0.025, andsteels with a lower o higher content in this element werediscarded.

The number of test bars for each forging, in order todetermine longitudinal and transverse properties both at thesurface and at the "A" zone, was always higher than 5 and 10for tensile and impact tests respectively.For each forging wethus tested more than 20 tensile test bars and more than 40impact test pieces. Tensile test bars were of the L = 5d type,and impact tests were carried out using Charpy V notch.

Results for normalized carbon steels are given in fig.11,in which ratios of properties in the "A" region to propertiesin the surface ("S") have been represented vs. tensile strength.Results for quenched and tempered alloy steels are summarizedin Table I.

There is indeed a loss in mechanical properties due tosegregations, and it is always higher in transverse tests. Inthe case of carbon steels, a tendency is observed for higherlosses at higher tensile strengths.We thus have,at Rm=45 kg/mm2,losses of 5 % for elongation, 6 % for reduction of area, and12 % for impact strength in longitudinal tests; and 20 %, 25 %and 17 % respectively in transverse tests. At Rm = 7_kg/mm2losses are of 8 % (elongation) 11 % (reduction of area) ånd-20 % (impact strength) in longitudinal tests; and 31 %(elongation) , 43 % (reduction of area) and 29 % (impactstrength) in transverse tests. It will be noted that resultsfor elongation and reduction of area are more scattered in thecase of transverse tests; this is due to the fact thattransverse test bars sweep regions in which heterogeneitiesexist (Fig. 10).

TABLE IMechanical properties in "A" zone Same properties in S zone

Allo steels

- 12 -

.9

AA7 A A

KCVA/K0,06to- 0

9

8

7

6

oA 0 a,

A A AL A A

0A A

A A A0 A

0 0 0 0 @

0A A

A

.ci50 60 70 R m(Kg/m m2)

ZARSto

00 0 00

A 00 @A 6

oA A

7 A A A A AA

A A AA AAA A AA

A A T

o 0 0 0 0 0

5 AA

so 60 70 R r„(Kg/mni)

50 60

0

- 13 -

A

0

A A A- A

70 R,(1<g/rnm2)

Fig. 11

RATIO OF MECHANICAL PROPERTIES IN THE "A" SEGREGATIONZONE TO THOSE IN THE SURFACE FOR CARBON STEELS OF DIFFERENTTENSILE STRENGTHS (R ).A = elongation; Z = Teduction of area; KCV = Charpy V notchimpact strength; L = longitudinal tests; T = transverse tests.

0.5

0 4

0.3

0 2

0.1

0.5

0.4

0.3

0.2

0 1

0.5

0,4

0.3

0.2

0,1 /-

1-51-15

0.5

0.4

0.3

0.2

0.1

AA/As =09

A (%)

HA/As 0 7-1

20

15

LONGaUDINAL TESTS

rZA/Zsr.--

0.5

0.4

0.3

0.2

0.1

0.5

0,4

0.3

0.2

0.1

b.30.2

0.1

0.6

0.5

0.4

0.2

--4

5

10

A

4

60

TRANSVERSE TESTS

ZA/Zs as

30

0.5

0.4

0.3

0.2

0,1

0.50.4

0.3

0.2

0,1

0.6

0.5 - +- 0.5

04 - ri 0,4

0.6

0.5

0.4

0.3

0.2

0.1

KCVAMCV=obsi

0 1

KCV(Kimftm2)

KCVAMCVS 0.9

3 5 7

A

2

10 30 3 5 7

Z (%) KCV(Kgsric m2)

Fig. 12ELONGATION (A), REDUCTION OF AREA (Z), AND CHARPY VIMPACT STRENGTH (KCV). NiCrMo STEEL.Relative frequency histograms; longitudinal and transverse

tests in the "A" and "S" zones. Mean values and standarddeviations are shown by vertical segments and arrows respect.Number of tests: 10 (tensile) and 23 (impact), both forlongitudinal and transverse tests.

Data in Table I are insufficient, and do not enable us toinfer a "law" of variation as in the case of carbon steels,butit may be observed that, except for 40Cr4, losses in mechanicalproperties are lower than for carbon steels of the same tensilestrength.

Each of the points in Fig. 11 and each of the entries inTable I is in turn the summary of a number of measurements.This is made clear in Fig. 12, in which results are given ofthe measurements made on a NiCrMo steel: 10 traction testsand 23 Charpy tests were performed, both for longitudinal andtransverse properties.

Standard deviations are some 5% (20%) of mean values forlongitudinal (transverse) tests; and 90% confidence limits formean values (i) are ey (1 + 0.03)Tc and (1 + 0.12)Rrespectively. Therefore, de-viations from the itraight lineswhich describe the general trend in Fig.11 should not besurprising.

Let us also note that Charpy notch impact tests which aresummarized in Fig. 12 were carried out with notches inso-to-say random locations. Tests were made with notches notonly in the "A" region but also in such a way that theyintercepted a segregation; the mean value of KCVA was then4.2 kg.m/cm2, instead of 5.5 kg.m/cm2: a 30% insEead of 10%loss in impact strength.

* * *

SUMMARY AND CONCLUSIONS

In forgings obtained from vacuum poured steel ingots,machining processes reveal "A" segregations which are nearerto the surface and, sometimes, more clearly visible than inforgings of non-degassed steels. A possible similarity ispointed out between the solidification of castings and thatof vacuum poured ingots. Type II sulphides were predominantin our forgings, and this would also contribute tocomparatively intense segregation lines in sulphur prints.

Elongation, reduction of area, and impact strength havebeen measured in the "A" segregation and surface zones, bothwith longitudinal and transverse test bars, for a total of49 forgings from vacuum poured ingots.

There is always some loss in mechanical properties fromthe "S" zone to the "A" zone; which is lower in longitudinal

- 15 -

test bars. In the case of normalized carbon steels, with tensilestrengths (Rm) between 45 and 77 kg/mm2, losses increase withincreasing Rm; greater losses were of some 50% in reduction ofarea, 40% in elongation and 30% in impact strength for Ruin.' 70kg/mm2. In the few cases in which quenched and tempered alloysteels were tested, losses were generally lower than fornormalized carbon steel of the same tensile strength.

Impact strengths show greater reductions when notches aremade to intercept the "A" segregations proper.

* * *

REFERENCES

(1) S. Fernandez; Inst.Hierro y Acero, 86 (Oct. 1963)

(2) C. E. Sims and F.B.Dahle; Trans.A.F.S.,46, 65 (1938)

(3) R.Ferry and C.Roques; Rev.Metal., 3, 175 (1957)

(4) A.G.McMillan, J.E.Russell, E.H.Watson, F.W.Jones, andA.Stanfield; International Forging Conference, Sheffield.September 19-21, 1967.

(5) A.Ambruz and G.Bicego; Convegno Internazionale dellaFucinatura, Terni. May 6-9, 1970.

(6) E.Marburg; J.Metals, 5(2), 157 (Feb. 1953)

(7) K.W.Andrews and C.R.Gomer; ISI Pub. 110, 363 (1970)

(0) J.R.Blank and F.B.Pickering; ISI Pub. 110, 370 (1970)

(9) E.H.Watson and H.Moore; Seventh Conf.on Forging,BISRAOct. 1970.

(10) K.Narita and M.Taniguchi; Tetsu-to-Hagane, 56, 212(1970)

* * *

- 16 -

HETEROGENEITY IN LARGE INGOTS.

INVESTIGATION OF THE INFLUENCE OF

IMPURITIES AND ALLOYING ELEMENTS UPON SEGREGATION

ABSTRACT

At the International Forgemasters Meeting held at Terni, in 1961, Le CreusotLaboratory presented a contribution about heterogeneity in large ingots. Thispaper was based on the analyses made on a large number of plain carbon steelingots, and a lesser extensive number of alloy steel ingots.

Since then, further investigations were conducted mainly on axial coresextracted from Cr-Mo, Ni-Cr-Mo and Ni-Cr-Mo-V steel rotor forgings. Such in-vestigations clearly demonstrate that in alloy steel ingots, the carbon segre-gation coefficient is significantly lower than is found in plain carbon steelingots, and that this is essentially related to the presence of molybdenum andvanadium : the ingots with the least segregation are those containing thesetwo elements. On the other hand, low sulfur, phosphorus and silicon contentsresult in a reduction of carbon segregation. An equation is proposed, relatingcarbon segregation coefficient to ingot dimensions and chemical composition.This enables to understand why the turbine and generator rotor forgings - whichin most cases contain molybdenum and vanadium - are affected by much lesssegregation than it can be assessed from our 1961 investigations.

To support those assumptions, two large forge ingots were vacuum pouredand cut off into slices. The first one is a 30 ton plain carbon steel ingotof standard quality. The second one is a 98 ton 3.5 Ni-Cr-Mo-V steel ingot ofhighest purity. Quite comprehensive investigations conducted on those twoingots (copper reagent etching, sulphur printing, micrographic examinations)and chemical analyses in large number made on longitudinal and cross sectionsreveal that the 30 ton ingot has much more segregation than the 98 ton ingot.The total carbon segregation coefficient is 75 % in the former, whereas it is28 % only in the latter.

This evidences that segregation in large ingots, sometimes considered asunavoidable can, to a certain extent, be minimized.

I. Introduction

At a time when all forgemasters are investigating methods to be adoptedfor the manufacture of rotors of 200 or 250 tons unit weight, we felt itopportune to publish this article about the segregation in large forging ingots.We hope it will contribute towards a better knowledge of a phenomenon which hasnot yet been completely mastered.

May we remind you that during the Forgemasters Meeting 1961 in Terni,Le Creusot presented a paper regarding the heterogeneity in large forge ingots.This was a relatively detailed report covering all aspects of the problem: Theelements of segregation at the ingot stage, ghost lines, minor segregationsetc.... However, this study covered essentially plain carbon steel ingots andonly a few alloy steel ingots.

The paper we present today should be considered as an extension of theone presented in Terni. We shall this time give details of our study of themajor carbon segregation and shall also try to go further by sifting out thefactors which determine the amplitude of segregation by means of a statisticalanalysis. This has been made possible by carrying out a considerable numberof new tests and new analyses on a large number of alloy steel ingots.

Eventually, we shall see that the presence of certain alloy elements inthe steel acts differently on the carbon segregation rate.

As a confirmation, we shall briefly give the results of tests made on twoforging ingots of 30 T and 98 T which constitute typical cases vis-a-vis thestatistical analysis obtained.

A large part of this work was carried out with the financial aid of theCommission of the European Communities.

1 Statistical anal s s of the carbon se re ation

1 0 pescri tion of statistical material reviewed

The segregation rates were measured either on "drilling cores" of hollowpieces, on axial bore cores or other similar pieces, or, in some exceptionalcases, on ingots which were split along the axis and destroyed for the purposeof completing the investigations. In all these cases, only firmly establishedresults were taken into consideration and all partial or doubtful analyseswere eliminated.

In this way we havecollectedthe segregation rates of 152 ingots. The

princ pal features of the studied statistic are given in the first plot oftable 1. You can see that the ingots are of variable dimensions, elaboration,teeming methods and chemical analyses. In the following text, we shall call"carbon segregation rate" the expression tC

AC = CM - Cm x 100Co

where CM = max. carbon content in the ingot body (except hot top)

Cm = min, carbon content in the ingot body

Co 7 ladle carbon analysis

The fact that the carbon segregation can be characterized by this segre-gation rate means that we assume the max./min. carbon content to be pro-portional, all other things equal, to the average carbon content of the ingot.Justification of this assumption was already given previously.C11

1 Preliminar ra hic anal s s

It is well known that the size of the ingot is the essential factor forthe segregation rate. ThisSize may be characterized by either the weight[2]or the diameter of the ingot. It can be noted that a sensibly linear relationexists between AC and the diameter D (note 1).

The graph on table 2 plots all our results in the AC D plane. The

first imperative remark is that for a given diameter, the carbon segregationrate can vary considerably.

A more detailed analysis showed that two alloy elements seem to play animportant role: molybdenum and vanadium. The first has already been givenattention in the pastel). The second seems to have a just as remarkable aninfluence. Actually, when considering in graph 1 the points representingingots without molybdenum and vanadium, it can be noted that these points arelocated in a scatter band showing the linear influence of the diameter. More-over, the regression straight which characterizes this relation passes verynear the origin which is absolutely normal - segregation will tend towardszero for a very small diameter.

Note 1: We call "ingot diameter" the average inscribed circle diameter in theoctogonal or 24-sided section.

The points standing for ingots containing molybdenum or vanadium arelocated in a band which is significantly below the first one, and the ingotscontaining at the same time molybdenum and vanadium are represented by thelowest points on the graph.

It can therefore be logically assumed that the influence of the diameteron AC can be drawn as a straight passing through the origin.

With growing molybdenum and vanadium content, this straight's inclinationdeclines. The most simple equation susceptible to reflect this pattern is

or

ac = D (a b % C %)

A.E. • 1 = a+bMo % +CV%

Eventually, we have also studied more into detail the factors susceptibleto have an influence on the size AC . 1 , which we shall call "standard segre-

gation rate" of whatever ingot is actuaYly the segregation rate the ingot wouldhave if it was 1 meter in diameter.

Before starting the total statistical analysis, we must point out thatthe sulphur content appears to have a particularly noticeable influence. Thisinfluence is revealed in the graph on table 3 where we have plotted AC . 1 as

a function of S %, distinguishing three ingot statistics: 1) without eitherMolybdenum nor Vanadium. 2) with either molybdenum or vanadium. 3) with molyb-denum and vanadium. The higher the sulphur content, the higher the carbonsegregation becomes.

1 2 C m lete statistical anal sis

Although the influence of the principle factors can be shown by means ofsimple correlations - as we did for molybdenum, vanadium and sulphur - this isnot possible for the secondary factors which can only be revealed by multipleregression calculations. These calculations were made according to the methodof least squares by adopting the following pattern:

n mAC . 1 = a0 +

n = 1

an Fn

Fn stands for the different factors which might intervene.

Several successive calculations led to the follow ng equation:

AC • 1 u 2.81 + 4.31 H + 28.9Si%+805.86%+235.213%-9.2%%-38.2V%C D

Details of calculation and intensity of signification are given in table 4.It can be seen that the influences of the already mentioned elements are con-firmed. In addition we must take into consideration:

- the H (height/diameter) ratio of the ingot body

- the phosphorus and silicon contents

Nonetheless, it must be said that the phosphorus content and the H/D ratioare secondary factors: the standard segregation rate can go down by 5 % as aconsequence of one of the following actions:

- addition of 0.55 % molybdenum- addition of 0.13 % vanadium- lowering by 0.006 % the sulphur content- lowering by 0.17 % the silicon content- lowering by 0.021 % the phosphorus content- reducing H/D ratio by 1.15

It is easily understood that the last two "modifications" are generallydifficult to realize, whilst the first three are quite possible in some cases.

We wanted to show how important the effect of an action on the molybdenum,

vanadium and sulphur contents can be and have calculated (see table 5) the se-gregation rate of several ingots with variable composition and diameter (1.5 -2.5 and 3.5 m. Naturally, the results for the last diameter are extrapolative.)

The considerable segregation rate which must be expected in a very largeingot of 3.5 m is prohibitive. But a 2.5 diameter ingot made from a carefullyelaborated molybdenum - vanadium alloy steel has a segregation rate which canbe qualified as reasonable. It should be noted that the influence of molyb-denum, to which referred the paper of Boques at the Terni conference in 1961,is confirmed by our calculations. Actually, table 6 of the paper of Roguesshows that a molybdenum addition (of 5 %) diminishes ,AC by about 15 % in a2,5 m diameter ingot.

Our calculations show for the same cese a 12 % decrease. The two evalu-

ations have quite close results.

3 C m ementar examinations

A certain amount of additional calculations and graphic analyses weremade in search for other factors. It would be tiresome to go into detailsabout these investigations. We tried to introduce the following elements intothe multiple regression calculation: carbon, nickel, chromium, manganese,copper, tin. None of these elements was revealed to have any significant in-fluence on the standard segregation rate.

Moreover, we also tried to introduce the "vacuum cast" material into ourcalculations in case degassing would have an influence. It appeared that nosignificant difference existed between the air-cast and the vacuum-cast ingots.

Neither did we find a traceable influence of the pouring temperature,which Stanfield et al [33 considered likely at the Sheffield Meeting in 1967.Table 7 shows that there is no relation between the pouring temperature andthe difference between the observed and the calculated standard segregationrate. Correlation is also inexistent when the teeming temperature is replacedby the "overheat", i.e. the divergence between the teeming temperature and theliquidus temperature of each steel.

Finally, there are some factors, such as the wall gradient of the ingotmould in respect to the vertical (3.6 to 4.8%) which might have an influence,but which do not vary enough in our statistic.

1 4 Inter retation

It is obviously difficult to know in which way the mentioned factorsintervene. The interpretation we are proposing now is exclusively based onhypotheses. The influence of the H/D ratio should very likely be consideredas a corrective term for the diameter. Actually, when you compare two ingotswith the same diameter, the ingot with the high N/D ratio will take longer tosolidify than the short ingot, i.e. it will have a similar behaviour as aningot with a larger diameter. It is not surprising therefore that a somewhatstronger segregation is noted in this instance.

The influence of sulphur and phosphorus is probably correlated with theiraction on the solidification interval: these elements increase the temperaturedifference between liquidus and solidus of which depends the partition co-efficient of the different liquid elements at the solidification front level.It is not abnormal that the major segregation of carbon, one of the final con-sequences of this partition, be more intense in a steel containing much sulphurand phosphorus.

More surprising is the fact that silicon plays a similar role. But itmust be noted, that its influence is feasible.

Molybdenum and vanadium act through a mechanism which supposedly is dif-ferent, This mechanism could be correlated with the strongly carburatingproperties of these elements. It might certainly be surprising to talk aboutcarbides in relation with phenomena occurring at the solidification temperature.Remember that during solidification, the crystals in formation push a liquidfilm in front of them, which is enriched with segregating elements, and parti-cularly with carbon.

Under these conditions, it is not inconceivable that in this liquid filmthe activity of the carbon be such that carbide precipitation of vanadium andmolybdenum could occur, impovering thus the solution in dissolved carbon. Thesecarbides could then play the role of solidification germs and thus artificiallyincrease the carbon content of the crystallites to the detriment of those ofthe liquid. This mechanism could compensate the effects of the classical segre-gation phenomenon.

2 Investi ations n tw in ots

We felt it would be useful to illustrate the above report and thereforegive certain results of investigations made on two fore ng ingots.

The first ingot was 1.3 m in diameter, 30 tons in we ght, in an ordinarycarbon steel. Analysis (14):

Si Mn S P Cu Sn As 00335 0.36 0.69 0,026 0.014 0.16 0.025 0,050

The second ingot was 2.15 m in diameter, 98 tons in weight, in 3,5 Ni CrMo V steel, with a very low impurity level. Analysis (10:

Si Mn S P Ni Cr Mo V Cu 0.245 0.095 0.36 0.008 0.009 3.70 1.85 0.45 0.10 0.067

Sn As0.012 0.009

Both ingots were made in a basic electric steel and vacuum cast (BochumerVerein process). They were cut through the axis as per drawing and were sub-mitted to a certain number of tests which already served and continue to serveas a basis for studies and reports. We shall confine our report tot

- carbon dosage made on a great number of points on these ingots,- sulphur prints made on axial faces.

Table 8 shows the carbon content variations along the axis of the 30-T ingot.

Not only the extent of segregation rate is remarkable, attaining 75 %, but alsothe irregularity of its variation: near the ingot's gravity center, a zonehas a clearly lower carbon content than the adjacent regions.

On the other hand, the results of the second ingot, although three timesheavier, are really better: the carbon content profile along the axis isregular, and only the upper end of the ingot body is affected noticeably bysegregation (see table 9). The segregation rate does not go beyond 28 %.

The other elements (alloy elements and impurities) behave in an absolutelysimilar way: sulphur and phosphorus, for example, are affected by a definitelylower segregation in the 98-4 ingot than in the 30-T ingot. It appears there-fore that the indicated factors not only rule the carbon segregation, but alsothe segregation of the other elements. It is of course difficult to be categor-ical on this point since the evaluation of segregation rate of an element suchas sulphur is always unprecise because of the low content of this specificelement. Nonetheless, the quantity and intensity of the sulphurous segregations(ghost lines and straight Vs) are considerable in the 30-T ingot, but very lowin the heavy 3,5 Ni Cr Mo V ingot (see table 10).

3, Cnnclusions

Hereafter you find the conclusions which can be drawn from this study:

1. It can be confirmed that - at a constant chemical composition - thesegregation rate is sensibly proportional to the ingot diameter.

2. The segregation rate of carbon, and apparently of the other elementstoo, is increasing with a higher H/D ratio and a higher sulphur, phosphorusand silicium content, and is decreasing with a higher molybdenum and vanadiumcontent.

3. This explains why the segregation rate found on rotor forgings insteel Ni Cr Mo V or Cr Mo V would be lower than can be deduced from theRogues study presented in Terni in 1961. This was particularly presented byStanfield et al at the Meeting in Sheffield in 1967,

4. The equation which has been established leads to an evaluation of thesegregation rate level which must be expected on an ingot with given dimensionsand analysis. In a 2,5 m diameter ingot, vanadium and molybdenum additionswill maintain carbon segregation below 50 %, whilst this rate can reach valuesof 100 to 150 % on 3,5 m diameter ingots in ordinary steel.

REFERENCES

[1] C. ROQUES, P. BASTIEN - Etude de l'heterogeneite desgros lingots de forge.

Convegno italiano della grossa fucinatura, Terni 1961

Revue de Metallurgie, 1960, 57, Dec., pp. 1091 - 1105

[2] J.W. HALLEY - Ingot structure and segregation -Basic Open-Earth Steelmaking -

E

AIME (1944) pp. 434 - 458

[33 A.G. MACMILLAN, J.E. RUSSELL, E.H. WATSON, F.W. JONES,A. STANFIELD - The production of ingots for the manu-facture of forgings.

International Forging Conference, Sheffield 1967.

J. COMON, P. BASTIEN - Etude experimentale des re-lations entre la solificationet les heterogeneites de lingotsd'acier de 3 a 30 tonnes.

Revue de Metallurgie, 1968, 65, Janv. pp. 13 - 23.

0•-1U)

Number of ingots studied

NOMBRE DELINGOTS ETUDIES

.cii CO lliiFour s

1 `)) a l'air

kiMarHn coules

0 basi ue sous vide0 Four cou iss, a I'air0.,)

ri 0 electrique cou sto 001 (I basique sous vide

TOTAL

C L

d/UU0steel stee

ACIER ACIERau C all ié

44 23

0 0

26 8

15 3685 67

TOTAL

67

0

34

51152

Aircast

Vacucast

Air Icast

Vacucast

DESCRIPTION PlancheDE LA POPULATION 1STATISTIC DEt;CRIPTICC

10

90

80

70

60

50

40

30

20

10

egrega ion ra e

taux de segregation,%

Steels without either Mo nor V

• Aciers sans Mo ni V (n:101)Steels with either No or V

o Aciers cu Mo ou au V (n : 27)Steels with Mo and V

Aciers au Mo et (n :24)

A 0

A

AA

A

A&S,

00 0,5 1,0 1,5 2,0 Z5

Ingot diameter diamètre du lingor, m

CL INFLUENCE DU DIAMETRE PlancheSUR LE TAUX DE

SEGREGATION DU CARBONE 2LlFLUZ10E OF DIAMET1 ON THE CARBON

SiDRZEATION RATE

Taux de segregation L\C

90

80

70

60

50

40

30

20

10

CL

• Aciers sans Mo ni V (n:101)ii Aciers au Mo ou au V (n: 27)A Aciers au Mo et V (n: 24)

•.• •

• ••141118 •

01.• I.

4•60 A

n •L A

AA

•

••

So

0 0,010 0,020 0,030Teneur en souFre , %Sulphur content

INFLUENCE DU SOUFRE SUR PlancheLE TAUX DE SEGREGATION

STANDARD DU CARBONE 3INFLUENCE OF SULPHUR ON THE STANDARD

S FORMATION RATE OF CARBON

•

•

Factor

EQUATION

Minimum Average Maximum Direction Significa-Value value Value of tion

de ee

Facteur Valeur Valeur Valeur Coef t- Sens Degremini moyenne maxi influence signif.°n

3,200 4,31 A r 97, 7%

0,450 28,90 7*. 99,5 Z

0,033805,8 A 99,97.

0,036 235,20 A 974 %

1,280 -9,20 'N*„. 99,9%

0,370 -38,20 N99,9 %

AC I _ 2,81 +4,31 it + 28,9 Si %+ 805,8 %C D

+ 235,2 P — 9,2 Mo % —38,2 V °X

(-± 2 G 15,6)

CL RESULTAT DU CALCUL Planche

7 DE REGRESSION MULTIPLE 4

R=EIPS OF THE ULTIPLE REOREBSION ANALYSIS

VALEUR DE kg- EN FONCTIONC L C Planche

7.3rDE LA COMPOSITION POUR

5DI FFERENTS DIAMETRE5

VA IHU: OF AC AS A Aq CTION OF 'ME

0 C;;:l PCS TT I(31 Fra“Afi IOUS DIA:':LLTERC

AC

0

60

40

20

•Acier avec Mo (n :16)

oAcier sans Mo (n :95)

molybdenum to'

00°doot Acier contenant

du molybdene0 Steel with molybdenum

0 500 1000 1500 2000 mm

diametre du lingot

Relation entre le taux de segregation du carboneRelation between the carbon segregaCion rate and the

et le diametre des lingots d "acier moyennementdiameter or low alloy steeJ. ingots containing molybdenum and

allié contenant du molybdenewithout molybdenum

Comparaison avec les lingots d'acier sans molybdene

REFERENCE Planche

t ETUDE DE L' HETEROGENEITEDES GROS LINGOTS DE FORGE 6

P. BASTIEN - C. ROQUESConvegno Italiano Della Grossa Fudnatura- TERNI 1961

STUDY 01',' THE H2FERrYJENEITY OF _LARGE L.:COTS

Acier sans60"

.000

to/".

Ecart- AC 1 mesure .... AC I calculé•C D

+30 • Coulee a l'air

o Coulee sous vide

+20 •

+10

C D

••

• •• •

• o •• o • • •

• io• gm•it, isEo 0 oi itipA •r :II CD• r • 0 • 00• o

• ••

STUDY CD THE IrUFLUIFI4CE OF POURlliGTiF2FIRMATURE

•

-20 01540 1560 1580 1600

Temperatures de coulee , °CPouring temperature

ci ETUDE DE L'INFLUENCE DE Planc

T. LA TEMPERATURE DE COULEE 7

C L

MasselotteHot top

%Pied in 11)eighl from bottom

A c %1900,2 0,3 OA 0,5 06

CARBONE DANS L AXET cAraia:3TEil.. INGOT

nT;TR “-771: -)H. Ai ‘'ninz' Trin, AXIS

C MINI0,258 %

TAUX DESEGREGATION

AC 75 %c

Tadi e analysis

C.COULEE0,335 %

0,2 0,3 0,4 0,5 0,6LINGOT 30 T ACIER AU CARBON Pkinch

REPARTITION DU 8

Pied

%Pied40

1.30

20

See above

C z0,2 0,3 0.4 0,5

•

a C. MAXI

II0,300%

TAUX DE

C.MINI0,230Z

10 SEGREGATIONAC_ 28 °Z,c

0 1•0,1 0,2 0,3 0,4 0,5

C.COULEE0,245%,

CL LINGOT 98T. ACIER 3,5 Ni Cr Mo V Planch

175REPARTITION DU 9

CARBONE DANS L' AXE

Lingol- 1300 de 301Acier au Carbone

oct. ingot of 30 Tin carbon steel.

CL EMPREINTE BAUMANN PlancheSUR SECTION AXIALE 10SULPHUR FRIIIT ON AXIAL 3"ICTI00

Lingo'. 2150 de 98TAcier 3,5 Ni Cr MoV, 1-3. p.r ngo OLGT

in alloyed steel

NEW SPECIAL ALLOY FORGING PLANTCOMPRISING 2600T OIL HYDRAULIC OPEN DIEFORGING PRESS INTEGRATED WITH 40T41MANIPULATOR

Dr. Hideo OhsawaDaido Steel Co., Ltd.Shoji FukudaIshikawajima Harima HeavyIndustries Co., Ltd.

ABSTRACT

A new forging plant mainly consist ng of a 2600 ton oil hydraulic open dieforging press and 40t-m manipulator was completed a• the Shibukawa factory ofDaido Steel Co., Ltd. in September 1971 and started operation immediately.Thus the Shibukawa factory has established the production capacity of forgingsto 48,000 ton per year by operation of the new forging shop together with theold one. Forging materials are stainless steel, heat resisting steel, toolsteel, high and low alloy steel and carbon steel.

A steel melting shop consisting of two electric arc furnaces, one vacuumarc remelting furnace and other equipments were newly built with forging shop.

For planning of the plant and introduction of facilities, such specialconsiderations were paid as high productivity, low production cost, labour saving,safty of work and prevention of public nuisance. The forging press having acapacity of 2600 ton is two column pull down type of plate welded frame, drivenoil hydraulically in high speed, suitable design for forging of special steel.

A 40T-M rail traveling manipulator integrated with the forging press, asit has big capacity of driving power with accurate control system, achieving100 mm feeding of max. ingot per each cycle at 80 S.P.M. planishing.

Though operation running of the plant is in short period, they arecontributing to rise up the production capacity of the Shibukawa factory asone of main facilities, being operate in high performance and producing highquality forgings.

INTRODUCTION

Daido Steel Co., Ltd. was etablished as an integrated maker of special steelin 1916 and has been producing hot and cold rolled products, steel castings,free and stamped steel forgings, narrow steel strips and industrial furnances in8 factories located all over Japan.

The Shibukawa factory was estdblished in 1937with steel shop, small andmedium section bar mill shop and forging shop, now being one of the main factoriesof Daido Steel Co., Ltd., and produced 70,000 tons of stainless steel, tool steeland structural alloy steel in 1969. About 15,000 tons of said annual output areoccupied by forgings produced by means of air hammers having capacities of 3 tons,2 tons and 1 ton. These products were made by old forging facilities using 300kg to 4 ton ingots, being bars and shafts having a diameter of 150 mm to 300 mm.

In 1970 a blueprint of increase of forging equipment was made in order tomeet the demand from the market for enlargement of product dimensions andincrease of product output. In infroducting new facilities there was made thefolloving productive plan.

(1) Total production target of old and new forging shop; 48,000 t/year(2) Kinds and rates of steel products

Stainless steel and heat resiting steel (20%), tool steel and highspeed tool steel (20%), low and high alloy steel and carbon steel (60%).

(3) Kinds of productsRound bar and shaft 100-754), square bar (100-750 squares), disc(max. 1,400/), ring max. 2,504) and the like.

(4) Kinds of ingots25 tons - 2 tons

(5) Ratios of forging worksDrawing and spreading work 70%Forming work 30%

The factors to be considered in introducting new facilities based onaforesaid plan are productivity, economization and in particular safety of works,good surrounding of workshop and prevention of public nuisance having beenemphasized recently.

The matters to be considered in productivity are as follows:

1) High speed forging to permit one heat production.2) Improvement in working efficiency due to organic combination of material

handling between facilities.(3) Minimum maintenance enough to avoid stoppage of facilities owing to

trouble.(4) Wide working space enabling to forge large sized ingots for the purpose

of raising production ratios.(5) Forging accuracy serviceable to improvement in yield rate of products.(6) Good management of steel making process and pertinent forging plan to

improve the qualities of forgings.

The matters to be taken into consideration as to economization are as mentionedbelow:

1) Low initial cost of equipment2) Labor saving3) Cheapness of utility and mainte ance costs

As for safety the following matters are to be considered, that is,

1) Release from dangerous works2) Safety of maintenance works

MPrevention of equipment from breakdownPrevention of accidents such as fire and so forth

For the prevention of public ni sauce the following matters are required, namely

(1) Low level of vibration and noise(2) Prevention of release of poisonous substances.

Taking these factors into consideration to the full, there were newly establishedin 1971 an steel melting shop including tNw electric arc furnaces and a forgingshop mainly comprising 2,600 t open die forging press made by Ishikawajima HarimaHeavy Industries Co., Ltd.

OUTLINE OF NEW E UIPMENT

1. Newly-built steel melting shopA new forging shop has been built to the south of the old shop, and further,

a new steel melting shop has been constructed in the new site across the publicroad to the south of the said new forging shop. Photo. 1 shows a panoramic viewof the Shibukawa factory and Fig. 1, the layout of the factory site.

Illustrated in Table 1 is the steel melting equipment before and after thenew installation.

Table 1 Main equipment in steel works

Moreover, 30-ton electro slag remelting furnace is scheduled to be built within2 or 3 years.The steel melting shop and forging shop are connected with the girder bridgewhich runs across the public road of the City. Steel ingot up to the maximwb25 tons is degassed and a hot ingot of about 1000°C is carried as is in theforging shop by the trailer. For the trailer, a trailer truck equipped withthe heat retaining cover, shown in Photo. 2 is used in order to retain heat aswell as to maintain heat uniformly during transportation.The electro arc furnace to be installed newly is of DAIDO-LECTROMELT type andthe vacuum arc remelting furnace and electro slag remelting furnace are ofDAIDO-Heraeus type. Both of them have been manufactured at the Takakura Worksof Daido Steel Co., Ltd.

2. Layout of forging shopFig. 2 displays the layout of the new and old forging shop.

The main equipment of new forging shop is as follows, T-M(1) IHT-SCHLOEMANN 2600-ton oil hydraulic open die forging press and 40

Manipulator2) 50 aind 36-ton heating furnaces: one each, and 25-ton one: 2 units3) 30T/51and 95T/5T overhead travelling cranes: one each(4) Frame cutting machine

Hot ingot is carried in the new forging shop through the east entrance, chargedin the heating furnace and after heated up to the specified temperature, it istransferred to the turntable at the front of the forging press by the overheadtravelling crane.This turntable (Photo. 3) is remote-controlled by the press operator and hasbeen so designed as to permit easy grip by the manipulator. •The material forged by the 2600-ton press is cooled down gradually on the coolingbed, and a part of the material is conveyed from the west exit and the rest istransported to the old forging shop via the northern exit to carry out re-forging.The hammer in the old forging shop is still used for forging the product havingdiameter of 300mm or less; however, since heavy vibration and noise are generatedand results in the working site situation being bad, replacement of the hammerwith a small forging press is now under consideration.As to the heating furnace to be installed in the new forging shop, it isconsidered that the materials to be heated will be various types of the weightand shape, therefore, a truck type furnace has been decided to be employed tosmoothen the charging and extracting of the materials. All heating furnaceshave been made by the Takakura Works of Daido Steel Co., Ltd. and the mainspecifications are as shown table 2. Photo. 4 shows the outside view of furnaces.

Heating Capacity Furnace Roomton Hight mm Width mm Depth mm

50 1,850 4,000 7,20036 1,850 3,500 6,70025 1,850 3,500 3,500

Heating Temp. 1,250°C regularly)1,350°C Max.)

Allowable deference of heat distribution ±25°C

Table 2. The S ecification of Heatin Furnace

For the fuel of the heating furnace, crude oil is used, which is veryeconomical and besides, LP gas pilot burner has been provided taking account ofstability of burning condition of fuel in the case of low temperature.For raising temperature, the automatic control by.pre-setting program isavailable and also, with the object of stabilizing burning, an automatic roompressure control device has been installed in each furnace. These devices areremote-controlled at the centralized control room. Consequently, operation ofthese furnaces axe controlled by one man. Operation of the furnace door, drawingout and returning of the truck are to be performed by means of the operatingdesk located at the front of the furnace and also, remote-controlled through theoverhead travelling crane. As a result, the crane is operated efficientlyaccording to the forging work, permitting the time required for taking ingot outof the heating furnace to be shortened to the minimum. Moreover, the furnace

- 4 -

door is to be operated by the crane operator and results in it being un-necessary to arrange the worker usually near the furnace. Therefore, thefact that it has been any longer unnecessary to arrange workers at thefront of the furnaee and at the rear of the manipulator is desirable from thestandpoint of the safety management.

In the building of the new forging shop, utilizing fully the advantagethat the press is of pull-down type, the crane rail has been installed atthe lower position (8.5m above the floor) and besides, in order to eliminatehot air and dust generated during the forging process, a large monitor has beenprovided on the roof and at the same time, full consideration has been givento the lighting and illumination apparatus too.A high speed crane has been adopted to enable the high-speed fnrging press tobe operated efficiently, and in order to smoothen charging and extracting thematerials and other handling procedures, a grab tongs has -been provided whoseturning and adjustment of opening angle grip can be performed by the craneman. (Photo.5)

Slow cooling of the materials after forging is an important taetorobserving from the quality control, and taking account of economy andworkability, a simple slow cooling equipment has been devised. The method is;to place forgings on the porous volcanic rock or pwnice laid on the floor andput an steei case on them. It has resulted in efficient slow cooling beingpossible by ffleans of the adiabatic property of pumice and glass wool providedon the inner face of the steel case.

FORGING PRESS AND MANIPULATOR

1. Dec sion of type of forging press

As the result of the examination of various items required for the fo g ngequipment stated in "Introduction", the following have been decided to beadopted as the fundamental type; that is a two column pull-down type havinga wide working space together with better workability and oil hydraulic pumpdirect drive system in which the initial cost is less along with easy maintenanceand high efficiency can be assured with respect to a high-speed forging work.This forgiJig press has of course been provided with the thickness control deviceby a digital control system in order to ensure high quality and accuracy of theproduct and at the same time, a high-speed manipulator with a powerful drivingforce which matches the high-speed operation of the press and the modernestcontrol system has been integrated with this forging press with the object ofensuring a high productivity.

Under such target, the forging press has been manufactured through thejoint design of IHI and SCIILOEMANN Co. of West Germany and is now beingoperated fully with its expected performance since September, 1971. Photo.bshows the press and manipulator and fig.3 and 4 are drawings of them.

2. Main Specificat ons of the forging press and manipulator.

(1) Forging Press

Power Step;Stroke;Daylight;Inside spate of columns;Capacity of table shifting cylinder;Capacity of die shifting cylinder;Hydraulic pumps;

Working pressure;Delivery volumn;No. of pumps;

(2) Manipulator

Load capacityMax. load momentTongs openingTongs liftTravelling speedTongs turningCar travel

3. Features of press design

-6-

2,600 ton and 2,000 ton1,550mm3,200/mn1,000x2,20Hun25 ton15 ton

550 k cm645 in6 sets existing)8 sets in future)

2

20 ton40 T-M400-1,220m750mn0-60 m/minmax 16 r.p.m.12m

(1) Adoption of steel plate welded frame.The main frame of the press is of a steel plate welded structure.

'The advantages of Said welded structure will be enumerated as follows.(a) h/iployment of one body frame of fewer divided parts permits obtainment

of a high rigidity in spite of its weight is ruther less.(b) The welding structure undergoes less structural restrictionsin design

as compared with assembled frame with forging columns or cast steelframe, thereby a suitable structure can be employed depending uponload distribution with a wide range of allowable eccentricity.

(c) As the weight of movable parts can be lessened in comparison withcast steel structure, the shock exerted on the foundation is also less.

(d) It is highin fatigue resistance due to its material reliabilityhighterthan cast steel.

In manufacturing this welded frame, even I.H.T. satisfactory results inmanufacturing and planning large sized welded structures and having wide experienceespecially in welding techniques of thick plate structures, they has conductedseveral kinds of experiments including the fatigue test with maximum eccentricload repeated over ten million times on the scale model of the frame (show photo.7), thereby to lead to actualization of 630 ton press (photo.8), and based onthus obtained experience has made necessary improvements to manufacture thispress. In the case of this press it is divided into upper and lower framesand two column parts for transportaion's sake, but may be made in one bodywelded structure unless inconvenient for transportation purposes.The connecting portion between column parts and upper and lower frames has agradually widened surface so as to be connected with a smooth, curved surface.This configuration was determined to be best according to the result of

experiments aforementioned and calculation by finite element method by meansof computer. Consequently, the stress concentration on this portion can bereduced to a minimum degree, thereby removing possibility of fatigue failuredue to stress concentration as seen in the conventional forging columnconstruction with colunm nuts.In the case of this machine the portion connecting upper and lower frames withcolumns, where bending moment is minimum, is provided with a shrinkage ringat its tension side, and is clamped at its compression side by a tie rod witha pre-compressed sleeve.The column part is of a hollow, rectangular section and has a sectional areabeing wide for the weight. In comparision with a non-hollow column identicalin sectional area as shown in figure5,ours is high in flexural rigidity,thereby allowing to put thereon more than two times eccentric load in cooperationwith advantages of its guide system.Figure6 shows stress distribution in terms of actual measured values.

(2) Two column pull down type pressThe advantages of two column pull down system have been referred to in

a great many documents, but reference will be made again thereto, namely,a Wide spaced working room which permits to make larger sized forgings.b The columns are disposed on a diagonal line, thereby letting workers

have a wide field of vision, and further handling of forgings ismade easy because a crane is well approached due to the press heighton the floor is low.

(c) The cylinder is located under the floor and oil hydraulic system areall in cellar with high safety against fire.

Photo.9 shows the new 2,600 T open die forging press.

(3) Fixed ram type guide systemIn the case of two column pull down type the distance between columns is

large to undergo thermal influence readily, and press frames are likely to shakevigorously owing to the presence of clearance in the guide part.

Therefore, it is an important problem to be solved the guide mechanismfor movable parts and to regulate the clearance. In the press, the main ramfixed under the press bed is made a lower guide for movable frames and thecolumn guide portion of the bed is made an upper guide, being the so-calledthree-point guide system, and therefore the distance between guides is great.According to this system, owing to possibility of reducing the load exerted onthe guide liner the life of liner can be prolonged and in addition thereto astability of movable frames during operation can be obtained because of lessenedshaking there-between.

As regards the plane liner in the column guide portion, furthermore,its clearance can be easily regulated by means of wedge.

(4) Oil hydraulic unitAs oil hydraulic pump there are now urd 6 sets of axial plunger pumps

having normal discharge pressure 350 kg/cm and each discharge 645 1/Min.,butconsideration is paid with reference to piping and the size of valves so thatadditional two pumps of the same capacity may be utilized in future.The characteristics of oil hydraulic direct driving system consist in that itis high in efficiency, easy to maintain and little in shock, thereby permitting

- 7

high speed operation. The press is provided with a main cylinder of maximumpower 2600 tons, but the power may be switched to 2000 tons by using a liftingcylinder of 600 ton output as a differential circuit to thereby increase penet-ration speed by 30%.In the automatic cycle for planishing and cogging purpose it is not effectivein obtaining a high pressure cycle only to increase the speed the time of no-load. It is surely necessary for obtaining a high pressing cycle to reduce thelength of over run to the fixed stroke length in a manner of minimizing time-lag by means of high performance control valves superior in response, and thepertinent way to obtain a quiet operation with necessary and minimmn stroke isTo operate at an apprapriate non-load speed.In addition what is especially considered for oil hydraulic circuit is to disposea decompression valve apart from a filling valve, said decompression valvebeing adopted to minimize the shock at the time of decompression for thepurpose of setting the period of time for decompression individually and mostsuitably. In view of a high S.P.M. it is deeply concerned with the compressi-bility of oil and for this purpose the press is designed so that the volumeof oil to be compressed in the cylinder and piping system is minimized.All the oil hydraulic unit and piping are accomodated in the cellar (photo.10),and the cellar is partitioned into three rooms, that is valve room, pump roomand tank room. A large capacity of main tank is isolated from others, whichis effective in prevention of fire, and the pump chamber is also effective forthe purpose of arresting noises. By way of precaution against accidentalfire the cellar is provided with a fire extinguishing equipment using CO2 gasand also with emergency exits at appropriate locations.

(5) Control of the forging pressThough the product thickness control according to a degital control system

have been cmmnonly used in the recent high speed open die forging pressreference will be made to features of our thickness control system, that is,

(a) One man operation system partially integrated with manipulator.(The operating desk is shown in photo.11)

(b) As detection, calculation and regulntion of the upper die positionare all made according to a degital control system, namely a simpleregulating circuit free from A-D or D-A conversion is employed, ourregulating system has little causes of trouble and is high inreliability.

(c) Though a difference between the presetting position of reversesignal and actual operation reversing position caused by the lag ofelectric and hydraulic controls, that is overrun is unavoidable,in our system there is employed an overrun compensation mechanismto automatically correct irregularity in the overrun length dependingupon conditions of load or operation speed.This mechanism has been developed by Schloemann Co., wherein isemployed a system to shift reverse signal transmitting position fromproduct dimension setting position by the actually measured overrunlength, and operates accurately even in the case of high speed runningwith a very short stroke. In case of this degital control system,due to several sorts of auxiliary counters provided therewith for thepurpose of making correction operation against the irregular motionof press there can be guaranteed high degree of forging accuracy.

(d) As for the operating mode, there may be three ways of manually-operation, semiautomatic and automatic operation, and since the

4. Manipulator

electic c rcuit in case of manually-operation is completely isolatedfrom that in case of automatic operation, even when the automaticcircuit is out of order the system can be manually-operated. Asthe manually-operated circuit comprises a simple relay circuit,some trouble even if occurred, can be eas ly found and repaired.

Previously, the manipulator was considered an auxiliary equipment forthe forging press. However, as the forging press has recently come to beoperated at high speed, therefore the relative decrease in operability andworking speed of the manipulator has prevented the operation of the high-speedpress.This manipulator was examined at the time of the fundamental planning that themanipulator should be one of the major components of the forging plant havinga systematic mutual relation with the forging press, and it has been sodesigned thdt it can follow up the forging press operated at high speedproperly on the performance. Photo. 12 shows the manipulator.

(1) Features on construction

(a) Peel guide mechanismThe peel with a tonis for clamping ingot is suspended on the manipulatorcarriage which travels on the rail via the link mechanism, and it isso constructed as to permit the parallel movement in vertical andhorizontal directions and inclined movement. The peel is guided bythe carriage frame and is always kept at a fixed position by acertain force. Therefore, comparing with the conventional type alwayssuspended by gravity and possible to swing freely, more accurateforging work is possible. Also, it permits a correct handling to aspecial work including forming or ring forging.

(b) Hydraulic spring of peel suspensionIn the case of forging, ingot is forced down according to the penetra-tion of upper die, however after the upper die has lifted, the ingotshould be returned to its original position. For this, the tongswhich has gripped the ingot moves up and down repeatedly during forgingprocess. This movement of the tongs is carried out by the spring actionof the cushion cylinder, and the front of the peel is suspended bythe cylinder, resulting in the peel moving with inclination centeringaround the rear fulcrum. Comparing with the parallel ascending/descen-ding system, this system has the following advantages (i) a force appliedfrom the press is absorbed by the front cushion cylinder and is nottransferred to the rear section, and further, (ii) since the levershaft does not rotate, the bearing service life lasts long.

(c) Inclination of peel axis and parallel descendingFurther mentioning the advantages of the inclined type peel in theordinary forging work, it is the fact that the bending moment whichworks on ingot and each part of the tongs becomes very small. ShownFig.7 is a comparison in forgeable length without bending steelingot in both cases of the parallel and inclined systems. This lengthis further extended by providing the cusion power automatic controldevice that is special regulating mechanism for balancing to the ingotload moment gradually incleasing according to the ingot lengthened

- 9

by forging. There is an opinion that ingot is bent during forging

process due tn inclination of the peel. But it is negligible, andthe main causes are; irregular distribution of temperature, wear ofdies or irregular shape dies, incorrect position of tongs and so forth.

Bending of the ingot is caused when the end is free during forging.Consequently, uhen forging is carried out in the direction where

the manipulator pushes ingot, it will bend regardless of the parallelor inclined systems of peel. Conversely, when forging is made inthe direction where the ingot is pulled by the manipulator, almostno bending is caused since the end of the ingot is kept restrained.

The parallel system requires more strength than the inclined oneobserving from the construction because of the relation with the bending

moment stated previously and as a result the weight also increases.

In this equipment, however, another manipulator is to be provided atthe opposite side to the press and two manipulatnrs are to be inte-

grated with each other and the press in the future. For forgingwith both ends of ingot gripped with tun manipulators, a parallelascending descending system is necessary, hence it is so designed

that even the manipulatnr of this system can easily be changed to theparallel descending system one by merely switching the hydrauliccircuit-

(d) Other featuresThe centnr of jaw which grips ingot can be brought to the center of

the press and also, it is possible to pick up ingot on the floor bythe manipulator itself, thus permitting a upset or forging work tobe done easily. The oil hydraulic control device including thehydraulic pump is installed at the rear of the manipulator tosecure the safety against fire and at the same time, it provides

for easy maintenance and inspection. It has also been arranged notto put any obstacle out of the frame side face so that the work can

be carried out safely.As the consideration on maintenance; the tongs and the peel aresupported one side on the ground and other side by the crane, and

they can be removed by giving back the truck.

(2) Control of manipulator

A digital control device is provided for turning the tongs and feed-

ing the carriage and one-man control system has been employed which isintegrated with the press.The operating mode is;

Feeding of ingotManual feedingCar travel preset feedingCar shot feedingTongs shot feeding

Turning of ingotManual turningConstant-angle turningIntermittent -curing

Horizontal movement and suing of peelAscending/descending Inclined

Grip Open/close

- 10 -

By the car-travel preset feeding, the stepping position or cutting positionis to be decided. By this, it is unnecessary to measure the ingot lengthone after another by the scale in forging process.When the forging length is set while performing the car shot feeding or tongsshot feeding, the press and manipulator stop automatically at the set value.This is a control at the previous stage of the program control of forging.The car position and tongs turning position are always given on the operat-ing desk by digital indication and also, the up-and-down position and in-clining condition, and horizontal position and swinging condition of thepeel are given by analogue indication.

Car shot feedingSince the car is driven by means of the driving device with a powerful

hydraulic motor through mesh of the rack and pinion, the steady operation isguaranteed.The driving force has a power of 2-4 times that of the conventional mani-pulator along with sufficient acceleration deceleration abilities, thereforeit permits feeding the specified quantity of ingot without giving impact tothe press dies and with no interference to the motion in the high-speedoperation of the press. Fig. 8 indicates the oscillograph of the forgingwork by this feeding.

Tongs shot feedingTongs shot feeding means that moves forwards and backwards only the tongs

which has gripped ingot to the car by the hydraulic cylinder, with the car runat a fixed speed and stop the ingot relatively to the press while the dies andingot contacts with each other. And, acceleration or deceleration of heavycar running is unnecessary, so it is possible to feed ingot at a long distancein short time. Consequently, forging of simple stepless round or squarebars can be carried out most effeciently. In the conventional tongs shotoperation, the tongs is accelerated by gravity, spring or oil pressure; however,as a measure has ever been taken that stops the tongs with the ingot being heldby the dies, there was a limit to the feeding speed because of impact given tothe dies and the high-speed operation was impossible.Further, a vital disadvantage of the conventional system is; when the penetra-tion of ingot is large, the ingot feeding is interruped by the dies, result-ing in operation being impossible.The tongs shot system employed in this equipment can be used even to themaximum S.P.M. during automaticplanishing cycle of the press and it permitsa large feeding. That is; since the tongs which has gripped ingot stops justbefore it touches the press dies and ingot, the high-speed operation ispossible without giving impact to the dies. Displayed in Fig.9 is the oscillo-graph showing the condition of 100mm feeding of 18-ton ingot at 80 S.P.M.Also, Fig. 10 shows the relationship between press S.PM. and amount of ingotfeeding by car shot and tongs shot.

Turing of tongsThe tongs turning is driven by the hydraulic motor, and this driving

force too is 2-4 times that of the conventional manipulator. For the forgingmode in which round bar is fed while turning, it is possible to have the high-speed press disply its function fully. This turning

torque also can resist the twisting force to ingot transferred from the pressduring forging process and as a result, it is unnecessary to provide anothermechanical brake for the rotary shaft of the peel and a correct shape canbe assured even in forging of polyhedral part.

Various feeding processes of these manipulator are given the individualextremely high acceleration deceleration abilities, and the speed controland feeding control are carried out. Further, in order to prevent the apparatusand piping from being damaged due to abnormal impact pressure generated onthe hydraulic circuit, a brake valve having high performance has beenprovided especially in addition to that of the acceleration ciruit so thatacceleration and deceleration can freely be adjusted independently.

(3) Integration of press and man pulatorA digital control device has been provided for the press and manipulator

which integrated with each other, and one-man control system is employed.By adding the input device, a program forging using the tape or punch card isalso possible.The combination of operation of the press and manipulator is as give below:

Press

o Hand Operation

o Automatic cogging

o Automatic planishing{

o Hand Operation

.r---------... o Automatic car shot

1'E---- --- o Automatic tongs shot

5. Anxiliary equipmentIn order to improve efficiency of forging work as well as to secure the

safety of the work, the following devices have been provided; upper dies quickchange and rotary device, dies shifting device and turn table with ingot lifter.These device are all remote-controlled on the operating desk so as to permitthe work to be done safely and quickly and as a result, workers are kept awayfrom the dangerous position.

6, Safety of forging equipment and consideration to public nuisanceA forging work was a field whose mechanization or automation were hard to

be accomplished, and it was considered a work involved in various dangersalong with ,a bad working situation. In this equipment, a serious considerationhas been given to the safety precautions and programs. For the sake of thecentralized control of the heating furnace and remote control of the furnacedoor by crane operator, it is unnecessary to arrange the workers around theheating furnace any longer, thus permitting the safety handling of ingot throughthe cooperative action of the overhead travelling crane with turning tongs,turn table and manipulator. Quick change and rotation of the dies and also,even correction and cutting of the forged material can be performed safelythrough the remote control on the operating table.Because of employment of two-column pull down press, a wider visibility is

- 12 -

Manipulator

o Automatic car shot and tongs turning

assured at the working site and the overhead travelling crane can also beoperated easily. For the aske of the above, labour-saving and workingsituation have been improved and the safety has been enhanced. As the hydraulicdriving and control equipment of the press has been arranged under the floor,safety against fire has been raised despite the hydraulic system being employedand also, even on the scaffolding used in maintenance, the safety has been takeninto account.The safety on the device has also been examined carefully, e.g. for over-runof the manipulator, a triple preventive device, that is, electric, hydraulicand mechanical stopper has been provided.Further, a perfect measure has been taken even to the safety on interlockbetween devices, liquid level control of each tank, pressure control of prefilltank and so on.The public nuisance generated from the forging works is vibration, noise,discharge of filthy liquid, air pollution, etc. The works is situated at thelocation relatively far from the residential district, however, vibrationgenerated from the hammer operation in the old equipment has recently broughton public discussion. Therefore, a careful consideration has been given tothe preventive measure of vibration and noise from the newly-installed press.Vibration of the press is generated from the speed change when starting pressurepushing of ingot and when decompression at the reversing point up and down.Increasing the working speed of the press unnecessarily in order to raise theS.P.M. in the case of planishing is a main cause of generation of vibrationand noise. We have succeeded in minimizing generation of vibration with theS.P.M. raised by obtaining the minimum stroke required at the optimum speedwith no impact.

Also, to impact caused during decompression of main cylinder decompressiotime has been adjusted freely by means of the valve exclusively used fordecompression in order to set to the optimum condition.

CONCLUSION: Actual results of operation of equipment

The operation of the newly-built forging plant was started on a full scalein October, 1971, and at the present stage, about 35011r operation is beingcarried out monthly in 2 shifts. Meanwhile, the downtime of the machines dueto accident including initial trouble of the equipment is about 10 hours monthlyon average, and routine inspection and maintenance are being performed once aweek for 8 hours. Three maintenance men take care of all the equipment.The monthly production target is 550 hours in 3 shifts and 3,000 tons; it is nowat about 2,000 tons/month in 2 shifts, which is almost as scheduled.The standard number of workers is; 2 or 3 operators of press and manipulator,2 crane operators, 1 or 2 workers for conditioning of forgings (one of them isalso engaged in press work) and 1 worker for heating furnace, totaling7.Only half as many workers are needed as compared with the hammer old equiped.A good quality, shape and accuracy can be ensured on the forgings, and accuracyof the product is within ± 2.5mm. The shape is also good. Particularly, atthe corner part of the square product, very sharp corner is generated.

It is only a short time since the equipment was operated and there are someincomplete respects on skillfulness of workers and preparation of the forgingprogram; however, it may be judged that the productivity and economy intendedat the initial stage have almost been achieved as expected. However, at the

- 13 -

present stage, the market of the forged products is in a dull tone and as atarget, the work done which meets the number of working man-hour is short,but we believe that the production intended can be secured fully in the eventof improvement of the market. Also, at this time, another installation of themanipUlator almost equivalent to the existing 40T44one at the opposite side isscheduled to further enhance the productivity by integrated two manipulator witheach other and the press.

In Japan, there is few actual results of installation of a high-speedhydraulic forging press as a high alloy steel manuafacturing plant, however,as many new data will be obtained through the actual operation of thisequipment, rationalization of the forging works will be greatly promoted,we sure.

- 14 -

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PIS. 2 The Layout of the New and the 01 d. Forging Shop

PHOTO. 3 The Turn Table

Furnaces of the f4ew Forging Shop

PHOTO.6 2600 ton Po

PHOTO.5r e Turning Grab Tongsof the Crane

P ess d 40T41 Manipulator

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FIG. 3 Assembly Drawing of 2600 ton Forging Press

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FIG. 4 Plan View of the Forging Press and the Manipulator

PHOTO. 7Load Test of the PressFrame Vtidel

Hollow Rectangular Column

Ram Rigidly Connected withBed

PHOTO. 8E30 ton Oil 'HydraulicForging Press

Solid Column

Ram Free from Bed

Compari ISo Scorp of Allowable AccentricLoad on. Follow Peotonyular and Solid Column.

CENTRALLY LOADED

STRESS DISTRIBUTION DIAGRAMON

STEEL WELDED CONSTRUCTION,ED

STRESS

TENTION

e COMPRESSION

C)r—

STRESS OCCURED IN PRESS MAIN FRAME

EXCENIRALLY LOADED

OPERATDR

FIG. 6 Stress distribution Diagram on SteelWelded Press Frame

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PROTO. 9 2600 ton 0f1 Hydraulic Forgino Press

170 10 Th Pumg Poom of 2600 tonForging Press

PHOTO. 11 The Operation Desk of theForging Press and theManipulator

PHOTO, 2 40T-M Manipulator

PENDING MOtMENA IZADED ONINGOT AT CRAMPED POSITION

Ingot Length 00

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FIG. 9-2Performance of the Tongs Shot Feeding IntegratedWith Press Planishing Operation

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FIG. 11Performance of the Car Shot FeedingWith Tongs Turning Integrated WithPress Planishing Operation

NEW FORGING EQUIPMENT AND SOME STUDIES

OF CLOSED DIE FORGING METHOD OF LARGE FORGINGS

by

Dr., Eng. S. Shikano, T. Oikawa and T. Iwasaki

The Japan Steel Works, Ltd. / Muroran Plant

Abstract

The greatest defect of the open die forging method oflarge forgings is poor formability which is inherent of thismethod. We have made effort to develop the closed die forgingmethod by a conventional vertical press as one of the meansto correct this defect for the parts of marine diesel engine.The new forging press suitable for both open and closed dieforging has been built as the result of these development.

In this paper, the closed die forging metnods by boththe conventional and new forging presses, the metallurgicaleffect of the closed die forging and the new forging equipmentare described.

Some closed die forging methods have been developedusing the vertical press but there are many problems on it,for example, larger die necessary for the method, complicatedforging process etc. Both to work out the solution to theseproblems on the closed die forging method and to improve theforging method itself, the new forging equipment was planned,and the new multiple ram press with the necessary accessorieswas built after settlement of the problems on its design forthe combination between closed and open die forging operations.

To study the metallurgical effect of the closed die forg-ing, closed die forged products were subjected to some examina-tion on the quality. Furthermore, the genuine fatigue test.of the open die and the closed die forged solid crankshaftsfor marine diesel engines was done.

With pressing without transformation on the closed dieforging method, it is expected to close the internal cavitiesand make the toughness uniform. The grain flow of the closeddie forgings is generally continuous along the outline surface.From the genuine fatigue test of unit throws of both closedand open die forged solid crankshafts, it is clear that theclosed die forged throw was higher than the open die forgedone by 19.3%, in fatigue strength against the maximum stressat the fillet. Thus, the closed die forged products are betterin quality and from an economical point of view.

Content

Page

I. Introduction ....... ........... . ......... • 6

II. The Closed Die Forging byConventional Vertical Press 2

2.1 The examples of the closeddie forging method 2

2.2 Problems of the closed die forgingby conventional vertical press

V. Forging Effect of Closed Die Forging 8

5.1 Grain flow 8

5.2 Strength and toughness 9

VI. Conclusion 9

NEW FORGING EQUIPMENT AND SOME STUDIESOF CLOSED DIE FORGING METHOD OF LARGE FORGINGS

by

Dr., Eng. S. Shikano, T. Oikawa and T. Iwasaki

The Japan Steel Works, Ltd. / Muroran Plant

Abstract

The greatest defect of the open die forging method oflarge forgings is poor formability which is inherent of thismethod. The effort has been made to develop the closed dieforging method by a conventional vertical press as one of themeans to correct this defect for the parts of marine dieselengine (crankthrow, piston rod, center piece etc.), and havebuilt the new forging equipment suitable for both the closeddie forging and the open die forging, making effective useof this experience.

In this paper, the closed die forging methods by theconventional vertical press, the new forging equipment, theclosed die forging methods by the new forging equipment andthe test results of the closed die forged products are described.

It is clear that the new multiple ram press with thenecessary accessories can satisfactorily work for both openand closed die forging operations as required, and that closeddie forged products are better in quality and from an economicalpoint of view.

I. Introduction

The greatest defect in the open die forging method is poorformability and difficulty in repetition which depends on thehuman factor, Forgings are certainly better in quality thancastings, welded structures, etc., but are more difficult toproduce. Therefore, due to the low productive yield and longmachining time, the manufacturing cost of forgings is high.

As a method improving the above defects, there is theclosed die forging method which has been used for some timefor the mass production of small forgings such as motor carparts. However, the closed die forging method has seldom beenemployed for the larger forgings because the development ofthe closed die forging method for larger forgings requiresa special technique to cover the many different shapes whichmay be necessary. However, products of the same shape arepretty numerous in the long term, and therefore there is thepossibility of growth of the closed die forging method for suchlarge forgings.

- 1 -

Recently, the development of the closed die forgingmethod with good formability and ease of repetition has comeunder review. The Japan Steel Works, Ltd. (J.S.W) has forsome time made a study on the development of special forming,and have obtained good results in some closed die forgingmethods using the conventional vertical press. As the resultof these developments, J.S.W has planned and designed a newpress which can be used in the closed die forging method forlarger forgings and also for open die forging, and have builta new multiple ram press with the necessary accessories.

The development of this new forging equipment and someclosed die forging methods are described hereunder.

II. The Closed Die Forging by Conventional Vertical Press

First, some examples of the closed die forging method byconventional vertical press are stated and then, the troublesof those method caused by use of the vertical press aredescribed below.

2.1 The examples of the closed die forging method

2.1.1 Crankthrow for semi-built-up large crankshaft

The closed die forging process for a crankthrow is shownin Fig. 1. This closed die forging method is the first onewhich was developed and patented by J.S.W. This methodattains the best formability, and the quality of the forgings,as compared with former methods, i.e. the block forging method,the ring forging method, the bending forging method etc.

The outline of the process is given as follows:

1) Second process is light forging before heavy forgingin order to tighten the weak corner of steel ingot.

2) In third and fourth process, the bloom is forgedequally on both the inside and outside by upsetting and solidforging, to give a good homogeneous structure with good mech-anical properties. Further, the top and bottom portions ofingot are cropped to get rid of harmful segregations, non-metallic inclusions etc.

3) In the fifth process, longitudinal direction of bloomis changed for better metal flow in the next closed die forgingprocess. Moreover, the bloom is given such forging effectsfrom six directions as are also given in the third and fourthprocess, so that the mechanical properties of both the insideand the outside of the bloom are improved.

- 2 -

4) The sixth process is the final process and is theclosed die forging operation, which has the following twooperations.

First stage: a billet is set on the open die and coveredwith a pushing die to form a closed die. Then, the billet inthe closed die is given the heavy compression forging effectwith the 10,000 ton forging press. This operation is themethod of performing effectildely the J.T.S. forging with theclosed die to close the internal cavities which can be scarcelyclosed by the open die forging, by concentrating the forgingpressure at the center of billet. Thus, as the high forgingpressure concentrates at the center of the billet without anytransformation of the forging, this operation closes completelythe internal cavities, besides this operation makes for a goodhomogeneous structure and improves the strength and toughnessof the steel.

Second stage: the dividing die is pushed into the billetand the crankpin is formed with high pressure. The metal flowwhich is perpendicular to the axis of billet is moved down bythe dividing die along the outline of crankthrow. The crankpinand its adjacent parts are therefore given heavy forgingeffects.

2.1.2 Piston rod for marine diesel engine

Only the top of the piston rod, which cannot be preciselyformed by the open die forging method, is forged by this method.The rod, being simple shape, is formed by the open die forgingmethod. The outline of this process is shown in Fig. 2,

The billet from which the rod is formed with the open dieforging, is partially heated at the top, set into the die, andcovered with a punch and cap die. Then the billet is pressedwith the vertical press to form its hollow by punching andits flange by upsetting with the cap die.

2.1.3 Center piece of marine diesel engine

The center pieces for marine diesel engines are used forcrankshaft bearings and have hitherto been manufactured inseveral ways; fully welded construction of steel plates,welded construction of steel castings, forgings and plates,and also monoblock castings. The shape of a center piece isshown in Photo. 1. The closed die forging technique for thecenter piece is as follows:

1) Rolling of slabs

2) Forming of bearing saddle and heavy compressionforging in a closed die

3) Forming of concave part

3 --

These processes are done in one heating. The featuresof this forging technique are as follows;

First is the fact that closing up of the internalcavities, increases the uniformity of mechanical properties,and improvement of the strength and toughness can be expectedby the heavy compression forging as in the closed die forgingof the crankthrow.

Second is the fact that the accuracy of dimensions, themetal flow, and the productive yield are higher, and theclosed die forged center pieces are used in the as forgedcondition except for machining of the circumambient portion.

2.2 Problems of the closed die forging by conventionalvertical press

Some examples of the closed die forging method by theconventional vertical press as mentioned above, are excellenton both formability and quality. However, there are someproblems caused by the use of the vertical press, and solutionsto these problems must be found in order to further developthe closed die forging method using the vertical press. Thefollowing are these problems;

1) Die becomes complex and large

The die is generally divided for putting over thebillet and taking out the forging, and especially, the divi-sion of the die for the vertical press is more complex.Therefore, to keep these dies as a closeddie during theforging operation, larger outer die is needed, and therefore,the weight of the die becomes great and its production costincreases.

2) Forging operation becomes intricate

As the billet is given pressure in one direction only,transformation is also subject to restriction. For preciseforming of forgings whose shape is complex, the forging workmust be done by dividing the forming operations. (For example,the foming of bearint, saddle and concave part of center piece)

3) There is much restriction to the shapes of forgings

It is impossible with the vertical press to do theclosed die forging of products, which require transformationsin two directions or more at a time.

4) The shape of billet is complex

As the transformation is subjected to restrictionbecause of one direction pressure, the billet must have a shapethat is easy to fit into dies. Therefore, the shape of the

billet is apt to be complex and the forging of the b llettakes much time.

It is considered that there is a limit to the developmentof the closed die forging by the vertical press due to theseproblems, and the press suitable for the closed die forgingis hoped. The productivity of a press suitable only for closeddie forging would be low, therefore the new forging equipmentwhich is suitable for both the closed die forging and the opendie forging has been developed.

III. The New Forging Equipment Designed by J.S.W

3.1 Equipment

1) Press: 4 column, push-down type multiple rams press

Capacity: Vertical 4,0006,000tonsHorizontal 4,000 tons

Stroke of main rams: 3,000 mm

Sliding platen: 4,000 x 10,000 mm

2) Rapid forging crane (Wireless controlled)

Capacity: 200 tons

3) Forging manipulator (Remote controlled)

Capacity: 120/360 ton-m

4) Tool manipulator (Remote controlled)

Capacity: 2.5 tons

3.2 Structure of new press

The structure of the new press is fundamentally similar tothat of the common four column push-down press. However, thedifferences are that the new press has side cylinders and thatits working conditions are more severe.

Structural features of this press are as follows.

1) As the side cylinders are fixed on the lower frame,the working space in the press is large as on the conventionalvertical press. The installation and dismantling of the sidecylinders is very simple and the cylinders can be removed forrepair etc. when necessary thus allowing the press to continuein use for vertical forging operations--mainly open die forgingoperations.

5

2) To endure high speed and pressure on open die forgingoperations and to ensure the accuracy on closed die forgingoperations, the press needs high strength and rigidity.Especially, because of eccentric load in addition to high speedoperation on open die forgings, the press operates under severeconditions, so that the press frame is given not only simpletensile and bending strains but also the energies in the shapeof the complex unsymmetrical strains combined tensile and bend-ing. Accumulation and release of these energies causes swayof the press and is a factor of the obstacle against accuracyof the closed die forgings. The structure of the frame, thedisposition of the columns and cylinders etc. are so designedto minimized this effect, and the press can satisfactorilyoperate on both open and closed die forging operations asrequired.

3) As the oil hydraulic installation to mount the toolswhich is newly developed, is attached to the press, changingof open and closed die forging tools can be carried out easilyand quickly.

3.3 Features

The features of the new forging equipment are summerizedas follows;

1) Both open and closed die forging can be performed bythe new forging equipment.

2) As the stroke of each cylinder is well-balanced withthe working space in the press, both open and closed die forg-ings can covera wide range of size.

3) Employing high pressing speed and automatic control,few operatos are required and the productivity is very high.

4) By use of thickness control, the accuracy of productsis improved on both open and closed die forgings.

5) The press is equipped with adequate safety devicesincluding the interlocker between the table cylinder and theknockout cylinder.

IV. Forging with the New Forging Equipment

4.1 Open die forging

Pressing speed of new forging press is higher in proportionto its capacity, and eight operating forms can be selected,employing automatic and thickness control. Thus, all kinds ofopen die forgings can be done precisely and efficiently.

Employingthe remote control system, operators of the

- 6

forging crane, manipulator and tool manipulator, can controlthe equipment at a place near the press operator, and there-fore, the efficiency of operations is markedly high.

These effects of the new equipment are notable. Namely,the improvement of ingot yield rate reaches 10 15%, and thatof efficiency 7 10%, as compared with the former open dieforging methods.

4.2 Closed die forging

4.2.1 Solid crankshaft for marine diesel engine

This method is for making forged solid crankshafts formedium size marine diesel engines and the closed die forgingprocess is outlined in Fig. 3.

R-R method, which has been developed by C.A.F.L in France,is well-known as an essentially fine forging method for mediumsize solid crankshafts, and this method as well as the R-Rmethod are approved by Nihon Keiji Kyokai as the ContinuousGrain Flow Process (C.G.F. Process).

This method has essentially developed from the closed dieforging method of large crankthrow, and has the same featuresas the forging method of large crankthrow.

These features are as follows.

1) During the last process, the closed die forging methodof large crankthrows is applied.

2) Forming of journals is done by forging and is notsubject to machining.

3) As the billet is fully and continuously heated in allprocesses without cooling to room temperature, there is nounstable operation concerning heat cycles during the forgingprocesses.

4) Formability is good, productive yield is high andmachining cost is low, because of the unique forging method.

5) Internal cavities are perfectly closed by the J.T.Sforging in the third process.

6) Pin and its adjacent part which are the most importantparts of the crankshaft, are fully forged at high pressure withextruded forging effect, and the metal flow becomes ideal withcontinuous grain in all sections of the journal, web and pin.

4.2.2 Crosshead pin for marine diesel engine

7

This forging process is shown in Fig. 4. The finalprocess is the closed die forging and is outlined as follows.

The billet is placed in the louer die and the upper diemounted on the vertical ram is set on billet. Then, sidepunches carrying the required shapes enter the closed die andthe forming is completed with a high pressure compressionforging effect.

This method is the forging process that makes the best useof the characteristic features of the new forging equipment.In comparison with the open die forging, the quality and theproductive yield are improved and the forging time is shortened.

This closed die forging process can be applied to theforging of products like pipefittings (tees, laterals etc.).

4.2.3 Connecting rod for marine diesel engine

Motion of the press in this forging is outlined in Fig. 5and is the same as the forging of crosshead pin. This is themethod that side segments (carrying the required shapes) whichenter the closed die upset billet in a horizontal direction,fit it into the die and complete the forging. The major aimof this method is the improvement of formability. The totalforging time is shortened, the productive yield is higher, andmachining costs are reduced.

V. Forging Effect of Closed Die Forging

5.1 Grain flow

As the elongation, reduction of area, and impact valuesof mechanical properties in a vertical direction to the grainflow are generally lower than those in a parallel directionthough there are some differences of values in proportionto degrees of forging effect, it is hoped that the grain flowis continuously along the shape of the forgings, depending onthe stresses under working condition of the forgings.

The major purpose of the closed die forging is to attainhigh repetition and precise forming of the products, and theclosed die forged products meet these demands more than theopen die forged ones because the forged shape by the closeddie forging method is near the final dimensions with consequentrepetition in machining time. Macrostructures (etching withcopper choloride ammoniac) of the closed die forged solidcrankshaft and crosshead pin are shown in Fig. 3 and 4.Both of them show clear continuous grain flow.

On the other hand, it is well-known that the conditionof grain flow has a great influence upon the fatigue strength.

- 8 -

This influence has been investigated by the genuine fatiguetest of unit throws of both closed and open die forged solidcrankshafts. The results of this test are shown in Fig. 6.In fatigue strength against the maximum stress at the fillet,the closed die forged throw was higher than the open dieforged one by 19.5%.

Reason for this difference is that, in the closed dieforging, the grain flow at the fillet is parallel with themachining surface, but, in the open die forging, the grainflow at the fillet is discontinuous by being machined as shownin Fig. 5, which forms a large angle between the grain flowand the machining surface. Concerning relation between fatiguelimit and inclination of grain flow, it is known that thefatigue limit does not decrease and is constant up to 220 ofinclination, decreases remarkably with increase of inclinationfrom 450 and is minimum at 900.

The grain flow of the closed die forgings is generallycontinuous along the outline surfaces and, as the grain flowis not cut in large inclination after machining, it may besaid that the closed die forgings are very good products fromthe point of view of fatigue strength.

5.2 Strength and toughness

Mechanical properties of the locations of closed die forgedcrankthrow, solid crankshaft and crosshead pin are shown inTable 1, 2 and 3. All values satisfy the requirements of theeach standard, and are nearly uniform. Elongation, reductionof area and impact of solid crankshaft of the test specimenstaken vertically to grain flow are lower than those takenparallel to the grain flow but the decrease of toughness issmall against degree of forging effect, showing a differenttendency to the open die forged one. Such uniform toughnessof the closed die forged products is due to the pressing in theclosed die without transformation. Pressing without trans-formation effects is expected to close the internal cavities,and makes the structure of forgings dense. It is a well-knownfact that the density of steel forgings is larger than thatof steel castings, and as the result of our measurement, thedensity of forginGs with closed die is larger in few mg/cm3order than that with open die. It may be said from this pointof view that higher quality obtained by the closed die forging,as compared with the open die forging.

VI. Conclusion

The above description outlined the new forging equipment,some closed die forging methods and its effect. The new forgingequipment is designed to be suitable for both the closed dieforging and the open die forging from the point of view of theproductivity and forging technique.

-- 9 -

It is clear that closed die forged products are better inquality and from an economical point of view, but the develop-ment of those froging method is very difficult. However, thereis room and possibilities of 0-owth of the closed die forgingmethod in the field of the open die forging of large forgingsas mentioned above.

- 10 -

Photo. 1 Center piece

- 12 -

Photo. 2 The now forging press

3

?ho to . 3 hacro-structure of closed die forged solid crankshaft

Photo. 4 Macro- tructure of closed die forged crosshead pin

A2

82 B3

- 15 -

A.

A3

Table 2 Mechanical properties of closed dieforged solid crankshaft

Chemical composition

Si Mn p Ni Cr Cu Mo

0.37 0.26 0.63 0.010 0 0012 0.16 1.12 0.23 0.20

Me chanical properties

Location of specimens

450

<k5,450 45°

wt. %

Direc-tion

Loca-tion

kg/14112

Tension test

T.S. EL.

kg/litn2

R.A.

Charpy impact

2mm U 5mm Ukg-m/okg-m/cm2

(1) 60.3 77.7 25.4 61.9 13.1 10.1

w(2)

(3)

61.861.1

79.578.1

23.7

24.5

60.8

61.2

13.4

13.2

9.2

11.5

) 59.8 77.8 23.7 61.6 13.4 11.5

ti (5) 59.6 77.2 25.2 62.4 10.8 8.8

0 (6) 58.3 75.7 24.4 62.1 13.4 10.2

H (7) 60.2 73.0 23.7 60.8 12.8 9.0

r-4 (8) 59.6 77.1 25.2 61.3 13.4 10.3

ci(9) 58.8 75.8 25.4 62.4 14.0 11.5

(10) 60.3 77.8 25.3 62.0 10.4 10.3

1 0d

0S-1 I-1 d

<-4

(11)(12) 57.9

ie,2

74.4

20.8

21.0)3.054.3

7.s8.1

6.i

6,9

Table 3 Mechanical properties of closed dieforged crosshead pin

Chemic 1. composition wt. %

0 Si Mn 13 S Ni. Cr

0.55 0.38 0.78 0.012 0.010 0.41 0.41

Mechanical properties

Location of sneoinens

=Li

T,

-17 -

c== L

0 T2

c=o L3

01.3

Process

Steelingot

Upsetting

Roughforging

Top &bottomend cut-in off

Squareforging

10,000 tonpress

push-ingdie

Finishing

(Closeddieforging)

10,000 tonpress

Forging shape Remarks

Light Light forgingforging (To tighten weak

Bottom end corner of steelcutting ingot)off

10,000 tonpress

-18 -

a) To forge theingot to behomogeneous upto the innerpart

b) To discard bothends

c) To change thematerialdirection

a) Heavily forgedby 10,000 tonpress

b) To close inter-nal cavities byso called compres-sion forgingmethod

10,000 ton a) Pin forgingb) Metal flowpress

4, divid-ing die

c) To forge thepin heavily andit's adjacent part

Fig. 1 Closed die forging process of crankthrow

Process -:Forging shape Remarks

Steelingot

Roughforging

Roughforgingandcutting

Finishing

(Closeddie forg-ing)

Y? (<7 /1.

Fig. 2 Closed- die forging process of piston rod

ProcesS. Forging shape Remarks

Steelingot

Roughforging.

J.T.S. Special forging method patented by To clbse .forging I the Japan-Steel Works, Ltd. internal

cavities

Roughforging To discard

both ends

Settingdown

Finishing

(Closeddief'orging)

"ID

Fig; 3 Closed; die, forging Trocess of solid trankahaft

- 20 -

To form thejournals byforging

rci_eet tneangle andeccentricity

Process Forging shape ! Remarks

Steelingot


Finishing(Closeddieforging)

c=0.%L.

— - -t

- 21 -

era

t=t>

Fig, 4 Closed aie forging process of crosshead pin

Process

Steelingot


Finishing

(Closeddieforging)

Ctri

Forging shape Remarks

== >

- 22 -

Fig. 5 Closed die forging process of connecting.rod

Fig.

6Results of fatigue test on solid crankshaft

Ei 4-1

CO co

t

2-

C e

sed

-rev

-O

pen

-a-

A

A

7ac

i 106

N.6.3

25

A23

0

a—

o• 1 2-11

01

1/40

<

(11

2 -4®

042

0.4-

0

0.38 0.36

2-2®

471

2-3

2 -4

(f)c

2--

II

1I

2I

die

die. I Without

2

Dri

vflS

0

I 3

F crac

k

45

IIBoi

ance

6

—

2

0.3`

18

1 1Open die forged

"

Ws'

cycle

-

3 4

5 6

7

—

4.3 f3.0

Ii

050

0.48

0A-6

3-21II I

III

—

01"

0 I-o 42

0

-0 -

3-4

3

Closed

die

forged

3-I

L41

- 3

0475

ON-FORGE HEATING ON THE OPEN DIE PRESS

T. W. Johnson

Corporate Laboratories of the British Steel CorporationSheffield, England

ABSTRACT

The loss of heat in open die forgingan unfortunate but inevitable part of theintervals during working and give rise tohas a substantial bearing on productivity

has until recently been regarded asprocess. Reheatings are required ata jobbing type of production whichlevels in forges.

To reduce cooling ratee techniques and equipment have been developed forsupplying heat while working is in progress, based on powerful infra-redheaters strategically placed round the workpiece. A prototype heating planthas been designed and installed on the experimental forge in the CorporateLaboratories in Sheffield and data applicable to a wide range of forgingsituations have been derived.

Trials on the prototype have shown that infra-red heating is effective inextending working times and that large forging reductions are feasible in asingle working period. Additionally, the development has particularimplications for the forging of low workability materials. Substantial costsavings could accrue for higher speed tool steels which have shown increasedmaterial yield through reductions in cracking, oxide scaling and decarburiza-tion. The intractable nickel based superalloys would also benefit; theirmanipulation is eased considerably by employing infra-red heating methods.

On-for e heati on the o en die ress

T. W. Johnson

The past two decades has seen significant increases in many sectors ofthe steel industry. Economies of scale have been achieved in the majorprocesses. Steadily increasing automation has resulted in increased outputand improved quality at the same time reducing overall manning requirements.In the steel finishing process of rolling and extrusion progress in theutilising of these newer aspects of technology has been rapid.

In open die forging progress has been less marked, this not necessarilyas a result of reactionary management but more because of the nature of aforge complex product mix. Small product batches having a wide range ofcomplexity of shape and size do not readily lend themselves to automatedmass production techniques. Nevertheless, wherever possible the forge-master has utilised improved techniques. Press position control is nowregarded as normal. The use of manipulators with the resultant improvementin presentation giving greater uniformity of working and increased through-put has steadily increased. Integration of press and manipulator operationhas now been amply demonstrated and accepted as a production technique andin the last two or three years automatic computer control has been installedin a few forges.

It is therefore clear that management in the forging industry is alertto and will utilise as far as possible developments which lead to greaterautOtation in their industry. It is perhaps also clear that changes in thestructure of the industry, for example rationalisation in product areas willmake''possible greater benefit from automation. Be this as it may thereremains one very significant obstacle to a straight through forging flowlineand that is heating.

There is no need to spell out the fact that many forgings requireseveral reheats. The resultant need to withdraw the forging from the press,transfer it back to the furnace and then yet again to the press, reducespress availability, and involves costs in more handling of the workpiece.Wasteful heat losses are high. All these add significaatly to the overallforging costs. Thus a method which permits of one heat forging shouldpromote significant cost reductions. In what follows research and develop-ment work carried out at the Corporate Laboratories of the British SteelCorporation to achieve this aim is described.

The choice of a heatin mediumIt was considered at the outset of the work that the most rational

approach to the Troblem was to heat the stock in a furnace in the normal wayto the maximal forging temperature and then to supply heat during forging toretard cooling sufficiently to allow forging to be completed in one period.It was therefore necessary to obtain data on the cooling of stock overtypical forging periods and thus obtain estimates of the heat input requiredto balance the loss of heat. The approach to and results of these calculationsare described elsewhere by Beard and Bly1. In summary they concluded that heatinputs of the order of 5 to 10 W/cm2 would be required. From these data and

1

consideration of practical forging conditions a wide variety of heating methodswas considered and infra-red radiant heating chosen. Again the details leadingto this choice are summarised by Beard and B1y1.

Infra-red Heating This is a non-contact form of heating in which energy from a high-intensity

heat source is radiated, using reflectors, on to the stock to be heated. Hither-to this type of heating had not been used for heating steel to hot workingtemperatures because of the lack of suitable heating elements. As the heattransfer process is entirely by radiation it is necessary to have a heat sourceat a temperature considerably higher than the desired stock temperature. Heatingelements capable of meeting these requirements were introduced in the late 1950's,originally of 1 kW power but now available up to 20 kW with a life of 4,000 hrs.

The make up of a typical heating unit which has been developed to housethese high power infra-red elements is shown in figure 1. The.heating elementitself consists of a heavy, spirally wound tungsten element enclosed in a quartzenvelope. This is backed by a water cooled reflector of anodised super-purityaluminium which may be of parabolic or elliptical profile to direct a parallelor focused beam of radiation on the stock. The reflector is fitted with aquartz window in front of the element to prevent deterioration of the anodisedsurfaces. Element working temperatures are currently in the region of 3000°Kand thus considerable visible glare is emitted. This fact raised potentialproblems with regard to the forge operators but was not considered insurmountable;a solution is referred to later in the paper.

In finalising on the choice of the infra-red radiant heater a number offactors were considered. Heating efficiency appeared to be favourable. Earlierwork carried out at BISRA2 on the development of these heaters had shown thatefficiencies in excess of 50% were possible when used for primary heating ofsteel. On the basis of this work it was estimated that efficiencies around 30%could be expected in the forging situation. The use of reflectors to produce aparallel or focussed beam of energy meant that the distance between the stockand heaters was not critical. This feature offered considerable flexibility indesign, in that both a fairly wide range of stock sizes and dimensional changes,inherent in any forging situation, could be accommodated. Finally, because theheat - purely radiant - could be focussed on the stock there would be no dangerof any of the forging equipment becoming overheated.

Feasibilit of on-for e heatinIn the foregoing sections the considerations on which a development could

be based have been outlined. Experimental data to substantiate calculationwas necessary, methods of assembling the heaters had to be established,potential operating problems had to be identified and the potential cost/benefithad to be established. In order to answer some of these questions a first andsimple assembly was manufactured.

The heating equipment shown in figure 2 was installed on a 2 MN experimentalpress and comprised of a central heater bank of 64 kW total power fixed withinthe columns of the press and a moveable bank of 48 kW power. 4 kW heaters wereused throughout arranged around a 50.8 cm diameter pitched circle. The moveablebank was mounted on a rail bound trolley and could thus be aligned along theforging axis. Both heater banks were water cooled.

2

The potential efficiency and performance was first obtained from non-forging experiments. Bars were heated to forging temperature and thenallowed to cool down whilst being held stationary along the centre line ofthe heater bank, to within a few degrees of their equilibrium temperature.At this stage the heat supplied by the heaters exactly counterbalanced theheat loss from the bar surface. From a knowledge of the equilibriumtemperature it was possible to calculate the heat incident on the surfaceof the bar and hence the efficiency of the process in terms of the propor-tion of the energy supplied to the heaters which arrives at the bar surface.

Table I summarises the results of these experiments. In addition tosubstantiating the original assumptions the data shows that substantialretardations of cooling can be obtained. These initial series of experimentsprovided satisfactory indications concerning efficiency but did not reflectconditions obtaining in practical forging procedures. They were followed byfurther series in which the bars were moved through the heater banks tosimulate the movements occurring during a forging operation. The results ofthis work indicated that it should be possible to design equipment on whichsubstantially increased forging time should be made available. Furthermore,the temperature gradients along the forging length should be less than 50°C.

Further valuable information was obtained from these series of trials.Operation of the manipulator with hot stock indicated that with good handlingpractice the heaters should not be damaged. The quantitative data on heatrequirements coupled with the indicated engineering costs enabled a preliminaryand favourable cost/benefit to be established.

On the debit side the problem of glare was confirmed as being very realand a question of how the life of the heaters might be affected by vibrationwas raised. Subsequent work has solved the glare problem and furtherexperiments and experience have shown that the vibration problem can beovercome.

As a result of the feasibility experiments it was apparent that theconstruction of a more advanced unit was justified. A unit on which fullforging trials would be possible from which industrial design parametersand operating procedure could be established and a firm cost/benefit derived.This constituted the next stage of the development.

Protot e on-for e heatin unitA preliminary examination of the form that the heating equipment might

take indicated that at one extreme there could be a powerful single bank ofrelatively short length situated at the centre of the press and at the othera continuous tunnel of heaters such that the forging was always within theheat flux. Between these two extremes were systems using a number of bankswhich although not covering the complete length of the forging could possiblyproduce uniform heating by moving them along the forging axis. A balancehad to be struck bearing in mind the need for uniform heating, cost of theequipment and difficulty of operation.

3

Initially a theoretical examination was carried out utilisingexperimental data on the heat flux distribution from a 4 kW heater unit.Using these data and assumed relative movements of stock and heaters thetemperature gradients for various systems were predicted. It was concludedthat fairly complicated systems of movement were required for a one or twobank system as compared to a three bank system. Moreover calculationsindicated that with a three bank system greater and acceptable uniformityof temperature along the stock could be expected. The calculated energydistributions for various stock sizes in the two and three bank systemsare shown in figure 2. It was therefore, from considerations of temper-ature, not necessary to go to a completely enclosed system with itsattendant operating difficulties in terms of accessibility and viewabilityand higher initial capital cost.

As a result of the above analysis a prototype heating unit employingthree independent moveable banks was designed, manufactured and installedon the 2 MN experimental forging press. Views of the unit are shown infigures 3 and 4. The three banks each comprise of twelve 4 kW heatersmounted in the form of two arcs and making up a total power of 144 kW. Forexperimental purposes provision was made for overvolting to give total powersof 180 kW and 216 kW. To follow the movements of the stock during forging theheater banks are motor operated and able to move within the presscolumneonrails. To protect the heaters from excessive shock the rails are floor mountedusing anti-vibration mounts. Shock levels as high as 140 had been measured onthe experimental press but by mounting the heaters independent of the press theshocks felt by the heaters are reduced to an acceptable level of less than 1G.Electricity and water are brought in on overhead cables.

Performance of the rotot e heatin unit

1. Tem erature levels and distributionThe performance of the prototype heating unit has been established over a

lengthy period of time embracing a considerable number of trials and can besummarised as follows:-

4

Forgings of finished section 10 cm square and whose length does not exceedthe total length of the heater banks (approximately 140 cm) can be maintainedindefinitely above 900°C at the 144 kW rating. Increasing the plant sower to180 kW increases the equilibrium temperature to rather more than løse C.Temperature gradients along the forged stock are generally less than 30°C.

For longer forgings (200 cm) equilibrium temperatures are usually around800°C and 900°C for plant power ratings of 144 and 180 kW respectively. Againthrough the development of the correct system of heater bank movement temper-ature gradients of less than 30°C can be obtained.

These indicative performance figures show a large measure of agreementwith the design figures and thus there can be considerable confidence inextrapolation to other specifications.

2. Heater units and elementsThe manufacturers quoted life expectancy of the heater elements is 4000 hrs

and there is now sufficient evidence to show that this may be achieved for theforging application. The simple anti-vibration precautions have proved adequate.The disposition of the heaters has, as was expected, prevented damage from scale.Premature failures of the elements have been few indeed and possibly no morethan could be expected from typical production considerations.

The overall frame design, mounting, rail carriage and services have provedremarkably trouble free having regard to the fact that the unit is an in-housedesigned and manufactured prototype.

3. Heater bank movement and controlClearly a considerable amount of experimental work has been devoted to

understanding and developing appropriate movement patterns, for the heaterbanks. As shown by the results on temperature levels and gradients, correctoperation has been achieved.

The physical movement of the heater banks is, as explained earlier throughmotor driven wheel drive. In the experimental work the movements have beenremotely operator controlled. For development and experimental purposes theremote control system was operated by an extra crew member. With the appropriatepatterns of movement established it is a comparatively easy matter to install atthe pulpit a simple programme controller. Alternatively where programme controlledforging is being used the control of the heater bank movements can be added to theoverall control system.

4. GlareAt the outset of the work glare from the heaters was considerable and the

operators had to wear protective eye-shields. During the course of the workeffort has been applied to the development of interference films deposited onthe quartz windows and now a successful technique is established. At the costof less than a 5% decrease in heating efficiency the glare has now been reducedto an acceptable level. This is not really as surprising as it might seemsuperficially as the visible light constitutes about7% only of the total radiantenergy from the heaters.

Forging Notwithstanding the performance of individualfacetsof the heating unit

the ultimate test of acceptability lies in the ability to forge stock andproduce a satisfactory forging. On a small forging press there is clearly asize limitation and tcet some indication of performance for a commercial productmust be gained. To this end it was decided to carry out the most comprehensivework on tool steels where there are severe problems in maintaining temperaturein a relatively narrow band and where decarburisation is a problem. Such stocknormally requites several reheats and surface quality is important.

Trials have been carried out on various high speed steel compositions asshown in table 2. The majority of the trials were with 10 cm square ingotswhich were reduced 4:1 in area to 5 cm square billets; in some cases thebillets were further reduced to4 cm square.

5

The results of the trials are summarised in table3. It will be seenthat a very wide range of forging temperatures was examined and in every casethe forging was done in one heat. It is not the purpose of this particularpaper to discuss the forward metallurgical implications of the results ofthese and other trials but rather to indicate the application of the newtechnique. From a production point of view the potential is adequatelyillustrated in the table which shows a combination of high yield and lowdecarburisation in the finished billets. Indeed it is worthy of note thatthe ingots in1415quality which were forged had been scrapped fromproduction because of inferior surface quality. Using the on-forge heatingunit these were recovered for further processing.

Cost/Benefit No forging plant is typical and individual production figures and costs

given in confidence cannot be used to justify a generalised case. In theassessment of a new plant or technology there is always in this area a dilemma.However, production management does need some indication to justify deeperinvestigation for its individual requirements. To this end an assessment,but not we hope a typical situation has been examined. As in every casewhere it is attempted to put costs in print it should be noted that thefigures are historical and that the ratios are more important than theactual figures.

Basis of AssessmentQuite a number of companies are engaged in tool steel manufacture and

there is considerable diversity in practice. Ingot sizes vary widely anddifferent forms of equipment are currently employed i.e. presses, hammersand rolling mills, with the result that different material yields areobtained - a reflection of the low workability of these complex steels.

In view of the diversity in practice the economic assessment is basedon an assumed product mix and output rate of high speed steels which mightreasonably apply to an 800 ton prese unit. It is also assumed that timesavings would be utilised for extra production and that prime materialresulting from yield improvement could be sold.

To assess possible benefits :through the use of infra-red heating ithas been necessary to use typical rather than definitive production data.The aim of this present note is, therefore, to sketch in a framework whichindividual companies can modify to suit their particular views and productionpractice.

Assumed Product Mix and PlantAn 00 ton press unit has been assumed, producing a range of carbon and

special steels, which for the present exercise includes 750 tons of high speedsteel comprised of the following:

Composition Tonnwge

M2 and similar 430 (60g)

Tl 150 (20%)

High dutymaterialse.g.cobalt & 150 (20%)vanadium bearing

6

Description

(ingots mostly8 in - 12 in(square, with some 14 in(square and occasional 16 in(square ingots.

Ingots up to8 in square

An average production rate of 1 ton/hr has been taken resulting in anoverall production time of 750 hours with existing practice. Average pressutilisation is taken at 6096 because of the high handling time associatedwith conventional multi heat working for high speed steels.

Calculations of infra-red power requirements for critical products inthe three classes above show the following:

Material Heating I.R.Requirement Power kW

142 or Tl, A threefold retardation14 in square between 1100° and 7290kWingot 1000°CHigh duty Equilibrium temperaturematerial, of 1000°C 480kW8 in squareingot

Based on these calculations a heater unit of 720 kW power is indicatedfor the product mix chosen.

Appendix I itemises the capital cost and on the basis of a five yearlydepreciation derives annual capital charges. Annual maintenance costsare also given. For comparison purposes the overall figures for a smaller48o kW unit are also included.

Appendix II summarises the expected benefits from the use of on-line heating.The savings are expressed in financial terms, based on percentage savings oftime and material. Benefits are considered under the following headings.

i) Productivity.

ii) Improvement in material yield, through reductions in cracking, oxidescaling and decarburisation.

iii) Reduced dressing costs.

Appendix III provides information on fuel costs.

Potential Cost enefitThe principal results of the cost/benefit assessment in Appendices I and II

are summarised in table4.

These figures indicate a wide margin in favour of an infra-red heatinginstallation, Moreover, in certain instances the benefits assumed are believedto be conservative while in other instances possible savings have been omittedentirely, because of lack of definitive data. The resultant margin shouldprovide sufficient leeway for downward adjustment which could arise fromdifferences in accounting procedures and market approach adopted by individualcompanies.

7

In respect of fuel costs it has not proved feasible to make a specificcomparison but it is believed that the adoption of infra-red heating willnot lead to an adverse situation. The main difficulties in comparing fuelcosts are, firstly, the actual forging time (and hence infra-red usage)cannot be forecast with sufficient accuracy, and secondly, it has not beenpossible to get a meaningful figure for fuel consumption during reheatingonly (separate from the initial heating). A general calculation inAppendix III, based on the efficiency measured on the pilot infra-red plantand current electricity, gas and oil prices, indicates that for furnacereheating to break even with infra-red heating, efficiencies ranging between53%and 73%would be necessary. These are difficult requirements for theinefficient topping-up operation of reheating where most of the heat is lostin the waste gases. The requirements may well be met in a modern furnaceinstallation operated under ideal conditions but in the authors' view aconsiderable proportion of furnaces, because of the practical conditions ofusage, will be less efficient. In the longer term, there will be a potentialsaving in total furnace capacity installed for some existing installations.This has not been quantified but should be noted.

There are additional advantages which will stem from the use of infra-redheating, and which have not been taken into account in assessing the cost/benefits. These include improved production procedures, one example of whichis the forging of M2 quality, isothermally at 800°C. This has been shown toreduce oxide scaling and decarburisation to negligible proportions resultingin further material savings and reduced machining costs. Infra-red workingshould also lead to improved mechanical properties through continuous andprogressive breakdown of the metallurgical structure. The reheating intraditional practice can dissipate the improvements brought about by workingdue to growth of the matrix grains and the second phase carbides.

Additional to high speed steel production, there are further applicationswhere improved heating will be of considerable immediate benefit. Theseinclude the difficult stainless steels and heat resisting alloys, which areprone to cracking and coarse variable metallurgical structures. High carbontool steels which have a tendency to decarburise heavily represent a furtherarea of use.

A lication of on-for e heatinAs indicated in the introduction to this paper the diversity of forging

plant, and product has always led to difficulties in introducing new technol-ogy aimed at the flow-line type of production. No less will management viewon-forge heating It is accepted that there is no panacea. Each applicationwill have to be individually considered having regard to the product mix,the existing equipment and plant layout. Each installation will have to bedesigned and tailored to suit the current plant situation.

These considerations should not cause automatic dismissal of the technique.Forward looking management will see the need to closely examine their owncircumstances.

8

The results to date have shown the sound scientific basis on which thedevelopment is based and which gives confidence in extrapolation to situationsbeyond those developed on the prototype unit. The Corporate Laboratoriesdevelopment team has examined a very wide range of press installations andidentified the chief features of design for the various situations. Thebroad features of operational techniques have been established and amplydemonstrated.

The results of the cost/benefit analysis confirm that in the field ofthe more difficult to work steels, on-line heating should bring substantialreductions in costs and further improvements in product quality.

Acknowledgements The author wishes to acknowledge the substantial contribution to the work

described in this paper made by Mr. A. Tomlinson, Section Head, MechanicalWorking Department, Corporate Laboratories, British Steel Corporation, who hasthroughout led the research and development team. He also wishes to thankMr. A. Tomlinson and Mr. D. Beard for assistance in preparing the paper.

The assistance and support of the British Forging Industry, through itsmembers of the former Forging Committee of BISRA is also gratefully acknowledged.

The author thanks the British Steel Corporation for permission to publishthis paper.

References

D.E.Beard & R.I.Bly. Seventh Conference on Forging. Proceedings 1970. 104.

2. W.R.Laws. Steel Times. Feb.26. 1965. 302.

9

Appendix I

Infra-red heati costs

Plant Cost

Annual capital chargeServicing of capital at 11%

Depreciation to 10% value in 5 years(See Note 2)

Annual I.R. Plant Cost: CapitalMaintenance

10

720 kW Unit

4 Bank heater installation1 spare bank. 45 heater units 15,750

Chassis 1,300

Heater Bank Control 450

Power Supply: cables and transformer 3,000switchgear 1,750

Cooling Supplies: Air 1000 cfm 240Water80 gpm 1,100

Installation Costs. Fitting equivalent to1 man year plus local overheads 2,000(See Note 1)

Contingency 10% 2,500

Total 28,090

3,09 0

5,05 6

48o kW Unit

20 ,530

Annual Charges 8,146 5,953

Maintenance (See Note 3)

25% Heater element replacement i.e.9 1,01725% Window replacement i.e. 9 54050%Anti-glare filter replacement i.e. 18 45025% Reflector renovation i.e. 9 90

5% Reflector replacement i.e. 2 325Labour, 4 man weeks with local overheads 160

Total 2,582

8,1462,582

Total 10,728

1,813

5,9531,813

7,766

Notes

1. Design costs have not been included.

2. Capital - the heating unit has been written off more quickly than the majoritems of forging plant owing to the innovatory nature of the equipment.

3. Maintenance - allowances made for replacements are generous and with goodhouse-keeping actual expenditures would be considerably less.

Appendix II

Ex ected savin s from infra-red for in

Productivity

i) Handling Time

There are several features (see note 1) contributing to increasedproductivity in infra-red forging. One of the most important isreduced handling time between furnace and press.

Taking3 heatings as an average procedure, a typical handling timewould account for 20% of overall processing time.

Eliminating the two reheatings would give a saving in processing time = 13%i.e. 750 hours reduced to 650 hours.

On the basis that the time saved was utilised for producing a similarproduct mix the increased13%production through the plant would be:

M2 57tons typically @ £560 ingot ton plus forging E32,500

Tl 19 tons typically @ £930 ingot ton plus forging = E17,700

High DutyCompositions 19 tons typically @ E870 ingot ton plus forging = £16,500

ii) Plant Ca ital Char es

With increased production the plant capital chargeswould be sustained by 845 tons of steel instead of750 tons. The capital charges attracted by theincreased production can be regarded as contributionto profit.

Taking the plant capital charges at C8/hr. (see note 2)the total for 100 hours production Moo

Total Saving Due to Reduced Handling Time £7,470

Material Yield

i) M2 and Similar

Average billet yield improvementdue to infra-red heating = 7.5%(see note 3)Saving = 7.5% of 450 tons at C400/ingot ton E16,500

Less scrap value which would otherwise havebeen obtained on the 10% material saved (see note4)

Solid scrap, 2.25% recovery.10.0 tons at C320/ton 3,200

Swarf,2.75%recovery.12.3 tons at C245/ton 3,030

Scale,0.75%recovery.3.4tons at C100/ton 340

Net Saving

11

Increased turnover = £66,700Assumed 10% profit E6,670

6,570

9,930

ii) Tl and Similar

Average billet yield improvementdue to infra-red heating = 7.5% (see note 3)

Saving = 7.5%of 150 tons at £860/ingot ton = 9,650

Less scrap return (see note4)Solid scrap, 2.25% recovery.3.4 tons at t600/ton 2,030

Swarf,2.75%recovery.4.01 tons at t490/ton 2,003

Scale,0.75%recovery.1.1 tons at t130/ton 143 4 176

Net Saving 5,474

iii) Hi h Dut Materials

Average billet yield improvementdue to infra-red heating = 12.5% (see note 3)

Saving = 12.5%of 150 tons at £800/ingot ton . 15,050

Less scrap return (see note4)

Solid scrap,3,25%recovery.4.9 tons at £560/ton 2,740

Swarf, 4% recovery.7.5 tons at £460/ton 3,450

Scale,1.25%recovery.1.9 tons at t120/ton 228 6 418

Net Saving

Total Material Saving all Qualities

Billet Grindin (See Note )

i) Automatic

Average cost taken at £5.50/ton comprising setting up andtwo grinding passes.

One pass at t1.85/ton eliminated

Saving . 750 tons x 1.85 £1,388

ii) Swing Grinding.

Average cost taken at t12.50/ton comprising setting up andtwo grinding passes.

One peas at t4/ton eliminated

Saving . 750 tons x C,4 £3,000

Assume saving is a mean of the abovesaving on billet grinding t2 194

12

8,63224,036

Notes

1. Increased Productivit in infra-red working would result from reductionsin the following:

a) Handling time between furnace and press15% and 25% have been observed.

b) Rectification time during forging i.e.for this have not been included becauseavailable due to wide variation in thisit can account for up to 50%of forging

- figures ranging between

gouging. Estimated savingsdefinitive data are notprocedure. In extreme casestime.

c) Production delays. These include unavailability of cranes or furnacesand alow reheating because of low furnace temperature or overloading.No account has been taken in the assessment of these unforeseen delays.

2. Plant Ca ital Char es - £0.25 m capital serviced at 11%, depreciation to1 value over 15 years on 3 shift working system.

3. Material Yield In pilot plant trials cracking has been virtually eliminatedand scaling and decarburisation have been halved compared with industrialvalues. A comparison of dressed billet yields is as follows:-

Infra-Red Forging Current IndustrialPractice

To allow for variation in the figures of yield from industrial practiceand because optimal procedures for infra-red working might not always bepossible in production plant, the yield improvements due to infra-redforging assumed for the purpose of assessment have been trimmed back to7.5% for 142 and T1 compositions and 12.5% for the high duty compositions.

4. Scrap Returns 100% recovery has been assumed for solid scrap;75% forgrinding swarf and50%for scale.

5. Billet Grinding Costs ranging between Lk - Lei/ton for automatic grindingand £10 - £15 ton for swing grinding have been indicated.

13

Appendix III

Fuel Costs

Electricit Cost for Infra-Red For in of tons Hi h S eed Steel

Heating plant usage is taken as equal to press operation time

Press operating time = 750 shift hoursat 60% utilisation = 450 hours

Heating cost = 720 kW for 450 hours at 480 kW Unit0.60p/unit £1945 1280

Air fan motor, 0.6 kW )Water pump motor, 3 kW ) 11 7

£1956 £1287

Com arison of Infra-Red Heatin with Conventional Furnace Reheatin

Electricity consumption for infra-red heating of 1 ton ingot= 720 kW for i hr. = 360 kWh

Cost at 0.60p/kWh = 216p

Heat input into 1 ton ingot during reheating(see note) = 3.5thermAssuming favourable industrial tariffs,cost equivalent, gas fired furnace= 3.5therms x 4.5p/therm 15.7p

Cost equivalent, oil fired furnace= 3.5therm x 3.5p/therm = 12.3p

For breakeven cost with infra-red forging:% efficiency required from gas fired

furnace = 15.7 x 100%216

% efficiency required from oil fired

12.3furnace = x 100

216

Note

Average of three reheating periods is required for forging a 1 ton HSS ingot.Account has been taken of the temperature gradients that would exist throughthe forging and the calculated heat input is that required to raise a 1 tonforging from a mean temperature of approximately 960°C to a uniform forgingtemperature of 1150°C.

14

= 7.3%

= 5.7%

Equilibrium

erature

Size of bar

Type of test 1100- 1100- 1000- 1100- 1100- 1000-

tem-p

900°

C

1000

°C

900°

C

900°

C

1000

°C

900°

CTe °C

(approx.)

TABLE 1

COOLING TIMES AND RETARDATIONS

FOR STATIONARY BARS IN CENTRAL HEATER BANK

Surface Cooling Time

min/

Retardation Factor

Calculation

of rate of

heat supply

to bar

W/c

m2

Percentage of

power supplied

to heaters

which arrives

at bar surface

8in square

Air cool

Cooled in

9.9

3.6

6.3

heaters

23.1

8.4

14.7

2.3

2.3

2.3

750

5.1

56

Air cool

9.2

3.9

5.3

62 in dia.

Cooled in

heaters

22.5

8.5

14.0

2.5

2.2

2.6

780

5.4

45

Air cool

5.1

2.2

2.9

4in square

Cooled in

heaters

18.1

7.5

10.6

3.6

3.4

3.7

840

5-8

42

3 3/

8in

Air cool

3.5

1.5

2.0

Cooled in

heaters

16.2

5.5

10.7

4.6

3.7

5.4

850

6.5

28

High-speed Steel

TABLE 2

NOMINAL ALLOY COMPOSITIONS

Cr

V

W

Mo

Co

M2

0.85

4.o

2.0

6.0

5.0

-

m4

1.30

4.

0 4.

o 5.

54.

5 -

m15

1.

5 4.

o 5.

0 6.

5 3.

5 5.

0

Tl

0.75

4.

o 1.

0 18

.0

--

H CI‘

T6

0.8

4.5

1.5

20.0

-

12.0

T42

1.30

4.

o 3.

0 9.

0 3.

0 9.

5

High-Speed

Steels

TABLE3 -SUMMARY OF FORGING TRIALS

Temperature

°C.

Ref. Compo-

Surface Surface Time for

No. sitionFurnaceat

at

forging, Forging Response

start

finish min.sec.

forge

forge

Decarburisation

Oxide Depth Equiv- Depth Equiv- Scale

Ingot

scale to

alent to

alent and

to

loss, 90%

metal full

metal decarb. bi3et

core, loss % core

loss% lÖss %

yield,

carbon,

carbon,

in

in

1

M2

1175

1060

975

10.00)

2.0

Not tested

75

2

M2

1175

1060

930

10.10)occasional

minor

1.25

0.019

3.8

0.025 5.0

6.7

8o

cracks

3M2

1175

1080

850

12.00)

1.0

0.018 3.6 0.030 6.0

7.0

81

4M2

1175

1080

912

20.15occasional

minor

1.0

0.019 1.9

0.025 2.5

3.5

cracks

(NB 6in

ingot)

5M2

1175

1000

900

16.15

No cracking

-0.013

2.6

0.038

7.6

-6

M2

820

835

920

15.00

Moderate cracking

during final passes<0.4

Examination

suggests slight carbur

surface.

onof

7

M2

800

800

800

6.30Moderate cracking

during final passes<0.3

No decarburisation

detected<0.3

8

m4

1175

1075

825

17.00

Minor cracks

2.0

No decarburisation

detected

2.0

8o

9

m4

1175

1050

850

10.20

Minor cracks

1.75

Not tested

86

lo m15

(Substandard

ingots)

1175

1010

895

13.00

Occasional

minor

to large cracks

1.85

Not tested

11

M15

1180

---1100-950

11.30

Segregated

areas of

large cracks

Not tested

12

Tl

11754-1100

920

6.10

No cracking

1.4

0.014 2.8 0.023 4.6

6.0

13

T6

1175

1075

890

12.20

Minor cracking

2.25Nottested

81

14

T6

1175

1070

900

7.00

Minor cracking

0.6

0.006

1.2

0.034 6.8

7.4

15

T42

1175

1040

945

13.10

Minor cracking

4.5

Nottested

82

16

T42

1175

1070

910

12.30

Minor cracking

1.75Nottested

78

CO

TA

BL

E4

Annual Costs

Anticipated Benefitg

Annual capital charges

8,14

6Due to increased productivity

7,47

0

Annual maintenance charges

2,58

2Due to increased material yield

24,0

36

Total Annual Costs

10,7

28Due to reduced grinding cost

2,19

4

Total Annual Benefits

33,7

00

i.7H

EID

IA

FIG

UR

E

2

-ti fl = fl

FIGURE 4

Subject: Automatic

Author: Dipl.-Ing.

Flame Cutting in

Hans Hirschberg,

the Forging Industry

Frankfurt/M. (W.Germany)

The process of flame cutting is playing a very importantrole in many phases of steelmaking and steel processing. Auto-matic flame cutting machines have been widely accepted forcutting steel plates in various industries, while the applica-tion of automatic flame cutting machines in the Open Die ForgingIndustry is still relatively limited. It has only been in thelast several years that they have been recognized.

Various types of flame cutting machines have been specifi-cally designed for diversified jobs in the forging industry. Ma-chines for straight line cutting are used to cut forgings ofround, rectangular or octagonal shapes with thicknesses varyingbetween4 in. and 80in. sometimes up to 100 in. with additionalautomatic equipment to set the distance between multiple cuts.

Lately efforts have been spent to develop shape cuttingmachines for applicationsin the Open Die Forging Industry.Photoelectrical control systems from full scale drawings con-trolling cross carriage type cutting machines with coordinatedrives designed for heavy cutting have become available in diffe-rent categories. Three typical machines are now in use for variousfields of applications:

1. automatic flame cutting machines with photoelectrical con-trols for cutting cold or medium-hot forgings (300 to40000 with thicknesses up to 600 mm or 24 in.

2. for larger thicknesses up to 48in. and temperatures ofup to 8500G. These machines are specifically protectedagainst the heat radiation from hot forgings and heat gene-rated by the cutting process.

3. machines for straight line and shape cutting of thicknessesup to 2000 mm or 80 in. with additional equipment to achieveperfectly smooth cuts in this range.

Special torches have been developed for the various rangesof thickness of either injector type, nozzle-mixing type or ex-ternal mixing type, each one for its proper application.

Two wellknown European companies firmly established in theopen die forging business have contributed interesting informa-tion and data demonstrating the technical and economical advan-tages of using an automatic flame cutting machine in the forgingindustry. Various examples are being discussed with the savingsachieved by flame cutting at finished forging temperature,cutting to much closer tolerances than known before, thus re-ducing mechanical machining time drastically. Cost savings,accuracies of geometrical shapes and surface quality of flamecut parts are discussed.

2 111•••

Title: Automatic Flame Cutting in the Forging Industry

Author: Dipl.-Ing. Hans Hirschberg, Frankfurt/Main (Germany)

1. Introduction

The process of flame cutting is playing a more and moreimportant role in many phases of steel making and steel pro-cessing. Automatic flame cutting machines have, for some time,been accepted for cutting steel plates, for example in theshipbuilding industry, in the general engineering industry,in the production of pressure vessels or in the bridge buildingindustry. The development of larger and more automatic flamecutting machines has resulted in increased production capacity,improved part accuracy and quality of cut, in other words, inreduced cost per part.

As a result of such improved accuracy and quality, advan-tages in the further stages of production, such as reducingtime and cost in mechanical machining, reducing fit-up timeand welding time and reducing the amount of weld deposit mate-rial, were achieved. Major developments for plate cutting ma-chines are now contributing in other areas, like for automaticflame cutting machines in the forging industry.

In the forging industry, the process of flame cutting hasalso been used for some time. More often than not, the processhas been used on a manual basis and not very extensively. Inconjunction with automatic machines specifically designed forthe forging industry, the process is now resulting in cuttingcosts by saving time, heat and furnace, as well as machinecapacity. How this is achieved will be the subject of thispaper.

2. Flame Cuttin Machines for the For in Industr

There are various types of flame cutting machines specifi-cally developed for application in the forging industry. Thefollowing basie machine types are used:

2.1 Machines for Strai ht Line Cuttin

These machines are used mainly to cut forgings of round,rectangular or octagonal shapes (picture 1) with thicknessesvarying between4 inches and80inches, in some cases even upto 100 inches. The cuts may be made from markings on the forgings,or the distance between multiple cuts may be set automaticallyby means of adjustable stops on the machine or digital electronicmeasuring devices.

1

2.2 Machines for Sha e Cuttin with Manual Control

Most of these machines in use in the past were modifiedstandard cutting machines. Generally they were not very welladapted to the job. In most cases the shapes were marked outon the forging and, by means of manual guiding systems, thetorch was guided along the marked-out line. It is quite ob-vious that with the heat of the forging and the additionalheat of the preheat flame and the light intensity generatedby the preheat flame and the actual cutting process, it isquite difficult for the machine operator to achieve a guidingaccuracy better than 1/4 of an inch. For better results thefollowing device has been introduced.

2.2.1 Shape Cutting Machines Controlled Manually byLi ht Cross

A light cross steering device enables the machine operatorto guide the machine along a predescribed path in the form ofa drawing by rotating a projected light cross in such a mannerthat one bar of the cross is always tangent to the drawing lineat the point of cutting. With this device manual control be-comes easier and much more accurate than with the method des-cribed above. Still the accuracy of the cut is greatly depen-dent upon the skill of the machine operator and his particulardisposition at the time of guiding the machine. Also a secondoperator is required to watch the performance of the torch,and raise and lower the torch according to the shape of theforging.

With forgings of stainless steels and other materials,that may not be cut with the classical oxygen cutting process,powder cutting is widely (picture 2) used where the newer plas-ma cutting, because of its limited cutting depth, cannot beapplied. It should, however, be noted that it is inherent inthe process of powder cutting that large amounts of dust aredeveloped. This dust does not readily permit the use in cuttingstainless steels of the ultimate in machine controls for theapplication in forging shops: photoelectric control systems.

2.3 Shape Cutting Machines with Automatic PhotoelectricControl S stems

Cross carriage type cutting machines with co-ordinatedrives specifically designed for applications in the forgingindustry are now-a-days mainly equipped with photoelectrictracers. These use a spot of light which follows a line drawingor a silhouette placed on the black table plate of the cuttingmachine as their guide to cutting a forging to shape. With thisautomatic control system, the machine operator's attention cannow be concentrated on the progress of the cut in the forging,including raising and lowering the torch in order to followthe contour of the forged part.

- 2 -

3. Different Types of Cutting Machines for theFor in Industr

A number of different designs of shape cutting machineshave been developed in recent years for the forging industry.They are generally all of the cross carriage type, equippedwith co-ordinate drives. The older method of driving a crosscarriage type flame cutting machine by means of friction wheelhas limitations with respect to the satisfactory performanceof such a machine. In particular, limitations exist with respectto the size and weight and heat shield equipment for applicationin cutting heavy forgings. These problems have been eliminatedby powering both longitudinal and transverse movements by meansof co-ordinate drives. The invention of such drives dates backto 1938 at the Adolf Messer GmbH in Frankfurt, Germany. Separatemotors for the longitudinal and transverse motion, electroni-cally controlled from a sine-cosine resolver permit almost un-limited power for traction in both the longitudinal and trans-verse directions. With such co-ordinate drives very large flamecutting machines with working width of up to 70 feet and up to20 torch stations have been built. It is also due to the co-ordinate drive that flame cutting machines using photoelectrictracers scanning reduced scale drawings could be built. Mostimportantly the co-ordinate drive is a necessary prerequisitefor numerical control systems.

Automatic cutting machines for the forging industry havebenefitted from this development since machines can now bebuilt tough enough to stand up under forging shop conditions(picture 3). Exposed parts may be water cooled to withstandthe high heat radiation'from both heavy cutting and the heatof large forgings. Any other necessary design feature can readi-ly be applied to these powerful machines. The following threetypical machines are in use:

3.1 Automatic Flame Cutting Machines with PhotoelectricControl for Cutting Cold or Medium Hot Forgings00 to4000C with Thicknesses u to 600 mm

Picture4 shows a STATOSEC K 1 installed at a Swedishforging shop cutting crank-shafts for diesel motors for marineand railroad locomotive service. The thickness of this parti-cular crank-shaft varies up to 500 mm with the length of theshaft being8 metres. Both the throw and the web of the crankare flame cut. The crank pin is shaped by flame cutting to re-duce the machining time on the crank-shaft lathe considerably.Picture5 shows a completed crank-shaft for a 2800 horsepowerlocomotive diesel as it leaves the flame cutting machine ina Swiss forging shop. Again not only the throw and webs of thecrank-shaft are flame cut but also the edges of the webs arebevel cut to reduce machining time on the lathe. When this

flame cutting machine was introduced, lathe machining timewas almost cut to half. At an hourly rate for this large crank-shaft lathe of DM 100,--, the investment for the flame cuttingmachine was recovered in less than six months.

Details of the cutting practice are shown in the photo-graph. One is the support arrangement for the crank-shaft, whichcomprises a double roller system at each end of a movable carri-age supporting the round ends of the crank-shaft. By an ingeniousmechanism the machine operator can rotate the shaft manually intothe required position for flame cutting at various cranks. Thecrank-shaft in this particular case, made of high tensile steel,is brought to the flame cutting machine at a temperature ofapproximately 7000C on the outside with inside temperatures ofprobably a little over 8000C. The complete machining time forthis crank-shaft on the flame cutting machine is just under6 hours. The temperature on the surface of the crank-shaft bythis time is 4800c. After flame cutting the crank-shaft istaken to a controlled temperature furnace for stress relievingand hydrogen diffusion.

In order to protect the flame cutting machine and the ma-chine operator from heat radiation, the machine has a heatshield under the cantilever arm with an additional air pipedesigned to remove heat between the copper-asbestos heatshield and the cantilever arm. The machine cutting torch isof the nozzle mixing type and is water cooled. The preferredfuel gases for heavy cutting are slow burning gases such aspropane, town gas or natural gas. Machines shown in both pho-tographs are equipped with a photoelectric tracer for both edgeline tracing or silhouettes and centre line tracing for linewidth of .5 to 1 mm.

.2 Cuttin u to 1200 mm thick

For larger thicknesses that usually must be cut at highertemperatures than 400°C, a much heavier machine is available.It is capable of cutting forgings up to 1200 mm thick with theforgings either only preheated or as taken from the forgingpress usually at 8500C. The heat radiation from the hot forgingadded to the heat generated by the flame cutting process it-self make it necessary to protect the cantilever arm extendingover the forging in a much more efficient way than by simplyproviding a heat shield. Therefore for these hot forging cuttingmachines the cantilever arm and the wheel truck adjacent tothe flame cutting area are completely water cooled. This pre-vents any heat distortion in the machine frame and cantileverarm, making it possible to maintain the high degree of accuracydemanded. The heavy duty torch capable of cutting thicknessesup to 1200 mm is of the nozzle mixing type or injector type.The torch, of course, is also thoroughly water cooled.(Picture 6) Such specially designed torch cutting machineshave been used for both straight cuts to produce individual

4

multiples from large forgings or for shape cutting hot for-gings using any of the advanced tracer devices previouslydescribed. A number of these machines are now in service inheavy forging shops on the continent and a large number arepresently being built.

3.3 Machines for Straight Line and Shape CuttingThicknesses u to 2000 mm

Machines just described are used for steel thicknessesup to 1200 mm. For forgings with thickness greater than this,effective flame cutting requires certain additional equip-ment. The carriage and structural features are identical withmachines previously described. However, it has been learnedthat additional fuel gas is needed for the following reasons:

The flame cut surface produced with a good heavy dutytorch will be smooth from the top to the bottom of the cut upto approximately 1200 mm in thickness but deep irregularitieswill develop above that thickness. Extensive research hasshown that if additional fuel gas is blown into the kerf be-hind the cutting torch, then clean, smooth cuts can be achievedup to 2000 mm (Picture 7). For straight line cuts, it is simpleto do this. An additional nozzle is located in such a way thata stream of fuel gas is blown into the cutting kerf. However,when shapes are being cut, the cutting torch has to rotatearound its own vertical axis which means that the additionalnozzle introducing the extra fuel gas also has to swivel aroundthe same axis in order to keep the additional fuel gas blowinginto the cutting kerf. To do this, a mechanism, either mechani-cally or electronically controlled, rotates the torch synchro-nously to the rotation of the photoelectric tracer, maintainingthe direction of the torch tangent to the cut at all times.When it is necessary to use powder cutting on heavy forgings,photoelectric tracers often cannot be used because of the largeamount of dust generated which abscures the drawing. In suchcases, it is quite simple to attach a light cross steering deviceto the machine. (Picture8)Just as with the photoelectric tra-cer, the rotation of the projected light cross by the machineoperator also controls the rotation of the cutting torch andthe slave nozzle injecting fuel gas into the cutting kerf.

The heavy duty torch used is either of the internal mixingor injector type or of external mixing type. With external mixing,fuel gas and preheat oxygen are mixed after having left the nozz-le of the cutting torch. This latter type has the advantage whencutting forgings of very high temperatures of eliminating thedanger of backfiring entirely. External mixing has a slightlyhigher noise level than nozzle mixing or internal mixing typetorches.

It must be emphasized again that not only must the heavyduty torch be water cooled but also the other working partsbeing exposed must be protected against heat radiation. Theseinclude particularly the raise and lower motor which remotelypositions the torch as the thickness and the aspect of theforging vary. A typical machine of that special design isshown in picture9.

4. Re ort on Practical Results

A major and wellknown forging company in Germany has contri-buted some very interesting information and data to report inthis paper which demonstrates the technical and economical advan-tages of using an automatic flame cutting machine in the forgingindustry. The machine in question has been in service since Janu-ary of 1969 without interruption. The machine was installed toreduce the cost producing medium and heavy forgings.

Before the machine was installed all forged parts were flamecut either manually or with conventional cutting machines modi-fied to withstand the heat generated by cutting heavy forgings.The method of operation was as follows:

After the completion of the forging, all parts which wereto be subsequently flame cut were treated just as any other for-ging, by suitable cooling corresponding to the steel quality inair or in a controlled temperature in a furnace. The parts to beflame cut were then checked and marked, preheated before flamecutting and manually flame cut in accordance with the marking.As a final step the parts to be flame cut were again cooled eitherin controlled heat or air.

With the"STATOSEC K 1of operation.ging press atmachine. They

introduction of machine flame cutting by usinga2000", it became possible to introduce a new methodMany forgings are now taken directly from the for-finished forging temperature to the flame cuttingare then handled as follows:

At finished forge temperature of 800 to 9000C,the part istaken to the flame cutting machine. Suitable holding times areemployed so that the temperature drops, depending on the steelquality, to 300 to 60000. The part is then placed on the supporttable parallel to the machine's longitudinal axis. The template,in the form of a drawing produced in the works Planning Depart-ment is placed on the template table. Bench marks on both permitcorrect position of the drawing parallel to the machine's longi-tudinal axis. The photoelectric tracer is manually moved to afew important points of the drawing guiding the torch over thework piece to check work piece and drawing alignment. This prode-dure also checks the width and length of the forging against thedrawing. If the desired finished part can be produced from theforging, flame cutting will proceed.

6

In this method of operation two points have to be parti-cularly considered. When preparing the template, the amountof shrinkage has to be taken care of. This in turn depends onthe cutting temperature required by the steel quality. There-fore, in order to maintain direct final dimensions, the forgingtemperature must be closely controlled during cutting. Conscien-tious personnel are needed both in making templates and incarrying out the automatic flame cutting operation. This forgingcompany in Germany employs two operators for their forgings,which are up to 10 metres in length and weights up to 20 metrictons. For certain of their steels during the flame cutting pro-cess, a large amount of dust is generated. When cutting theseforgings a blower with a nozzle directing a stream of air ontothe drawing ahead of the photoelectric tracer removes the dustfrom the template drawing, thus avoiding mistracings. The costof a second man in this operation is minor compared to the re-heating costs in the old method of operation. Added to thesavings in reheating, additional furnace space is available formore productive work.

It must be stressed that this operation must be well plannedso that the forging being flame cut, at no time should reach atemperature below 30000. A major reason for that is that with thedirect postforge flame cutting the heat treatment for hydrogendiffusion is postponed until after the flame cutting operation.No difficulties have been experienced with this procedure sincethis machine was installed.

Hydrogen is being used asthat it produces the smoothestdegree of accuracy. Experienceties for thicknesses up to 500thicknesses up to 1200 mm they

a fuel gas since it has been foundflame cut surface with the highesthas shown that surface irregulari-mm are no more than .5 mm and forare only 1 mm.

The flame cutting machine is supplied with oxygen by a pipeline from a cold evaporator while hydrogen is supplied to themachine from a permanent bundle of cylinders. Each bundle con-tains 28 cylinders containing altogether 280 cubic metres of hy-drogen (Picture 10).

The flame cutting machine is equipped with two types of tor-ches. For cutting thicknesses up to 300 mm a machine torch ofthe MS 613 type is used while for thicknesses between 300 and1200 mm a heavy duty torch of the MSTW type with injector mixingis used. The small torch is attached to the heavy duty torchwith a quick disconnect holder so that changing from one torchto the other is only a question of minutes.

At thicknesses of about 700 mm the top of the cutting kerfis 12 to 14 mm , while at the bottom it is between 20 and 22 mm.Statistics from this company with respect to the accuraciesachieved are:

7

For thicknesses up to 100 mm tolerances of between .1 to.3 mm are experienced. Other data are tabulated below:

These tolerances have been consistently achieved when cleannozzles, correct flame setting and cleanliness of the machinehave been maintained.

Some interesting examples were supplied by this companysuch as the forged heads shown in picture 11. The thickness ofthese heads is 410mm. The material is a carbon steel with.22% carbon. Five to seven of these heads are forged as a mul-tiple and then flame cut at finish forging heat. Machiningallowance for the outside dimensions is 10 mm, for the bore15 mm. The following comparison of previous and today's resultshas been supplied by this German forging company:A six multiple forging was previously flame cut after coldlay-out. The forging weight was approximately 16,000 kilopond.The method of operation was as follows:

1. Mark and lay-out the shape

2. Drill a 30 mm diameter hole for the bore

3. Preheat the part

4. Carry out flame cutting

5. Cool parts off

Today again with a six multiple forging and a weight ofabout16,000kilopond operation is as follows:

The forging goes direct from the press to the flame cuttingmachine. Torch pierce hole and automatically flame cut bpre.Torch cut outside shape of each head. Multiples cool to ambienttemperature.

Manufacturing costs for this head before automatic flamecutting are given as 100 %.

Today's cost to produce these heads is only 30 % of the pre-

vious cost, with a resulting 70 % cost saving.

Picture 12 shows two forgings, a crushing head and aframe part. Both were being flame cut using finish forgingheat. The part thickness is 200 mm. The crushing head againis a carbon steel of .35% carbon which is flame cut at 3000C.The outside shape is cut with an accuracy of plus or minus.5 mm while the internal hole has a machining allowance of10 mm. The frame part is cut from a chrome nickel molybdenumquality steel being cut at 60000. The outer contour is cutto a tolerance of plus or minus .4 mm, very close, when consi-dering that the height of this part is almost 1 metre.

Very interesting parts are shown in picture13. "Y"-piecesfor high pressure superheated steam, one part being600mm thickand two of them being 700 mm thick. The material is a molybdenumvanadium steel of the designation: 14 MoV 63 which has to beflame cut at a temperature of 60000. The outer contour was cutwith a machining allowance of 40 mm while the bevel cuts left amachining allowance of 60 mm. The forging ingots were forgedthree-dimensionally into heavy discs, one 1400 0 by 700 mm thick,the other 1200 0 by 600mm thick. The main shape was flame cutautomatically from a template photoelectrically controlled. Thebevel cuts were done from a chalk lay-out on the part. The costsavings for these parts are quite impressive. The Y-piece, part A,previously forged as a round and then mechanically machined bymeans of milling machines and lathe are now being flame cut at84 % saving. In other words, the cost of the present-day methodis only 16 % of the previous cost. Part B, the 1400 mm round,with the new method of operation cost only 18.5 % of the previouscost. For the smaller disc of 1200 mm diameter and 600 mm thick-ness,the forging weight was 5,300 kilopond, and after flamecutting the weight was 1,600 kilopond. These 3,700 kilopond pre-viously had to be removed by mechanical machine tools. Thecorresponding figures for the larger disc were 8,500 kilopondforging weight, and 3,050 kilopond finished cut weight.

Another example, shown in picture 14 are connecting rodsfor a reciprocating saw. These rods are 200 mm thick and 4 metres

long. The material is a carbon steel of .35 % carbon. The outershape was flame cut to a tolerance of plus or minus .4 mm anddid not need subsequent machining. The hole was flame cut witha machining allowance of 15 mm. The templates are made to corres-pond with the shrinkage taking place to 300°C.

Picture 15 shows two fork-shaped pieces as flame cut. Thematerial is a chrome molybdenum steel with the designation34 CRMO4, which was flame cut at a temperature of 5000C. Thethickness is 740 mm while the total length of each part isapproximately3 metres. The fork and the holes were flame cutwith a machining allowance of 15 mm. The photoelectric tracerfollowed two different templates, each one representing theshape of the part to be cut in the two planes shown in the pic-ture.

9

The picture 16 shows various parts forged and flame cut,thicknesses varying between 100 and 450 mm. There are connec-ting rods, hooks, piston heads, articulated joints. The arti-culated joints for example are flame cut out of a C 45 qualityhigh carbon steel. The thickness is 450 mm with a machiningallowance of 15 mm. The stamper is made of C 45 again beingflame cut at 300°C. The thickness is 440 mm. The shape wascut to plus/minus 1.5 mm while the connecting rod with a bea-ring attached is 100 mm thick and the outer contour was cutto a tolerance of plus or minus .3 mm. The machining allowancefor the bearing hole was 10 mm.

Picture17shows the machine with which all these partswere flame cut. It is a co-ordinate drive cutting machine"STATOSEC K 1tflIn the view shown the machine is cutting afloor plate, at finish forging heat. Three of these pieceswere forged together as can readily be seen from the drawing.The thickness of the forging is 200 mm, the material is ahigh manganese steel of the 30 MN5 designation. The holeswere pierced with the machine cutting torch, using an automa-tic hole piercing device.

This equipment allows the operator to set the preheat timein a timer. In using the device, after igniting the preheat flame,he pushes a button. The complete hole piercing cycle thereafteris fully automatic.

Another very wellknown forging company in Austria suppliedsome additional interesting examples. A large flange400 mm thick,1700 mm o.d. and 600 mm i.d. was also flame cut at finish forgingtemperature. The cost comparison between using an automatic flamecutting machine and the previous method of using a machine latheis quite interesting. Taking the cost of the previous method as100 %, the cost using an automatic flame cutting machine is only70 %. With yet another part,a steam turbine shaft,even more drasticsavings could be achieved. If the cost of the previous method inthe mechanical workshop to cut off the 600 mm diameter shaftswas 100 %, the cost of performing the same job on the automaticflame cutting machine is now reduced to 25 % of the previous cost.The third example supplied by this forging company in Austria isthe most impressive one. Picture 18 shows the part with its di-mensions in mm. The method before applying an automatic flamecutting machine was as follows: The two ends of the 300 mm dia-meter shafts were flame cut, while the cut out of 635 by 645 mmof the 500 mm thick central part was mechanically machined. Thecost of this operation taken as 100 %, the savings by applyingthe new machine tool,the flame cutting machine, were most impres-sive. The cost was now only 13 1/2 % of the previous cost.

- 10 -

The machine used in this particular Austrian forging shopis a STATOSEC HK. This machine is capable of cutting forgingsup to 9 metres long, 2000 mm wide and 2000 mm thick. It isequipned with two different torches, one capable of cuttingforgings between 100 and 1200 mm thick of the SAH external mi-xing type, the second torch being of the GIGANT type for cuttingforgings up to 2500 mm thick for straight line cuts and 2000 mmthick for shape cutting. As mostly high alloy steels are beingcut, the machine is equipped with the light cross steering de-vice for manually tracing full scale drawings. A trough to posi-tion heavy forgings is arranged immediately next to the machinewith the following dimensions:3.9m deep, 2.5 m wide and 9.5mlong. In order to fulfil requirements against industrial pollutionan effective fume exhaust system has been installed. In order tokeep the power of the fume exhaust system within reasonabletechnical limits, covers for the big trough have been designedto close up the open area of the trough as much as possible.The cover has been divided into two parts, which will remotelycontrolled from the machine panel provide coverage of thetrough from the front end and the rear end respectively onlyleaving such a portion open as is necessary for the machinemotion when cutting either straight lines or shape cuts.

Summary

Various types of automatic flame cutting machines have beendescribed that serve various applications in the forging industryfor cutting forging thicknesses between 100 and 2500 mm. Prac-tical shop experiences with automatic flame cutting of heavyforgings prove that impressive savings can be achieved. Not onlycan more accurate cutting be done than with previous methodsthus reducing mechanical machining time on expensive machinetools, but further savings will be gained by cutting forgingsof different thicknesses and sizes directly from forging heat,thus reducing heat cost and making expensive furnace spaceavailable for more production. Automatic flame cutting machineshave gained importance in the forging industry on the Europeancontinent and will contribute further to improve the economyof this industry serving so many other industries.

FIGURE 1: BSE 3 K, with heavy-duty cutting torchGIGANT, cutting lot forging

FIGURE 2: Stainless steel cut with powder process

FIGURE 3: Operator's panel of cutting machinefor heavy forgings

FIGURE 4: STATOSEC K 1

FIGURE 5: Completed crankshaft for a 2800 h.p. locomotive dieselleaving the flame cutting machine

FIGURE 6: Rudder shaft being flame cut invarious planes

FIGURE 7: Rudder shaft 1200 mm thick, flame cut

FIGURE 8: STATOSEC HK co-ordinate flame cutting machine forthicknesses up to 2000 mm

FIGURE 9: STATOSEC HW heavy-duty torch for cutting forgings

FIGURE 10: Hydrogen supply for cutting machine fuel gas

FIGURE 11: Forged heads 410 mm thick, flame cut in forging heat

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FIGURE 12: Crushing head and frame part, 200 mm thick, flame cut in forging heat

FIGURE 13: "Y"-piece, 600 mm thick, for high-pressure super-heated steam

FIGURE 14: Connecting rods for reciprocating saw, 200 mm thickand 4 metres long

FIGURE 15: Fork-shaped pieces, 740 mm thick, flame cut

FIGURE 16: Various parts, forged and flame cut, of thicknesses varyingbetween 100 and 450 mm

FIGURE 17: Co-ordinate drive cutting machine, STATOSEC K 1, for heavy cutting

‘45

FIGURE 18: GFM crankshaft flame cut forging heat

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ABSTRACT

Initially flame cutting was not readily accepted by the Heavy ForgingIndustry.

The craftsman saw it as a challenge to his skill, whilst the forgemasterand his customer were deterred by early failures due to heat stress cracking.

Removal of surface defects by deseaming during forging, enabling theforgeman to complete his work satisfactorily, overcame his opposition andpersuaded him to accept the benefits of other oxy-flame techniques. Firstthe removal of discards to avoid the dangerous and time wasting cut-off atthe press, followed by the more accurate profiling to shape and size.

The solutions of problems previously causing failure and the widerange of product now dealt with, indicate the need for strict technical control,Each item has to be studied individually and given the same attention todetail normally given to the forging process•

The apparent limit of 1,370 mm ( 54 " ) for flame cutting has beenextended by the use of lance cutting, which appears to have no size limitations.

This method of cutting has greatly assisted the introduction of casthollows for the manufacture of hollow forgings, making savings in both timeand cost.

The method of deseaming normally used to detect and remove surfacedefects has been modified for the removal of severe surface tears fromlarge hot forgings.

Although more difficult to deal with than ordinary alloy steels, manyqualities of stainless steel have been flame cut, both with and without theuse of powder and a combination of processes can be advantageous.

Success with flame cutting can only be achieved by providing adequatefuel and oxygen supplies and using skilled operators with the best equipmentand suitable protective clothing.

FLAME CUTTING PROCESSES IN THE HEAVY FORGING INDUSTRY

The introduction of flame cutting into the heavy forging industry has beena slow and not always successful process. Neglect of precautionary preheatingand postheating led to early catastrophic failures through heat stress cracking.

This fed the ego of the skilled craftsman who saw this method of shapingas a threat to his skill, but at the same time, it prejudiced the forgemasterand his customer against pursuing the economic advantages of the techni ue.It is now, however, a well established process in most progressive f rges andis proving almost indispensable in the fight to increase output and reduce costs.

It is probable that the use of deseaming to remove surface defectseventually opened the way for the more accurate flame cutting processes. Thedeseaming torch, unlike the normal cutting torch, has a very large preheatingnozzle and the surface metal is progressively melted and blown along in frontof the torch throughout each stroke by the main stream of oxygen from thecentre of the nozzle.

It has the ability to remove surface defects quickly and completely,whilst still retaining heat in the forging, and so enable the forgeman tocontinue and complete his work with increased confidence.

Deseaming has proved so helpful in overcoming what had hitherto been afrustrating problem, that there has been little further resistance to theintroduction of other flame cutting equipment and processes.

First the removal of discards by flame cutting, not only eliminating thedangerous and time wasting pr actice of cutting off at the press, but preventingdamage to the hot forging tools by the narrow and hard cutters. This in turnmeant better shaped forgings and reduced costs in both forge and machine shop.

Flame cutting to shape and size followed, either reducing forging costsby eliminating a number of operations or expensive tool changes, or reducingthe amount of surplus material to be removed by subsequent machining. Profilecutting machines with different methods of control have been produced to dealwith this and are no doubt dealt with more fully in Herr Hirschbergs paper.

Before leaving the subject however, there are one or two matters worthattention, particularly the causes and cures of some of the failures.

.....Among the first items to be cut to profile were webs for built-up

crankshaftS for steam locomotives, and later those for slow speed dieselcrankshafts, and with both there were failures due to heat stress crackswhich resulted in temporary condemnation of the flame cutting process.

In the case of the locomotive crankshaft webs, these were flame cut toshape from roughly shaped slab forgings, approximately 125 - 150 mm thick,in .45C steel either in the untreated or annealed condition, and then oilhardened and tempered. Preheating prior to flame cutting was ofteninadequate and post heating ignored completely with the result that heat stresscracking occurred and was intensified by the subsequent oil quench. Thecure was oil hardening and tempering after forging, producing a fine grainstructure more acceptable for oxycutting, adequate preheating and a retemperor stress relief immediately after flame cutting. Preheat temperature must

be below the tempering temperature but at the same time sufficiently highto ensure that cooling during transport both before and after flame cuttingdoes not generate stress cracks. Preheating to 450°C minimum and rechargingbefore cooling below 250°C is recommended. Flame cutting in an areaadjacent to the preheating and stress-relieving furnace is an advantage ifthis is possible, as it eliminates delays caused by transport.

Although apparently wasteful in oxygen consumption our practice has alwaysbeen to use higher pressures than recommended by equipment and fuel supplierswho are mainly concerned with oxygen and fuel economy. Our method permitsfaster cutting speeds and consequently less heat loss and temperature drop,essential in preventing thermal stressing and cracking.

With the larger webs for slow speed diesels, the material being mild steel,many manufacturers considered preheating and post heating totally unnecessary,even though flame cutting at 300 - 350 mm thick. For a time they had littleor no trouble, but eventually there were many rejections at a late stage inassembly, resulting from delayed heat stress cracks. The subsequent costsfar outweighed the costs of preheating.

Our original practice was to normalise as a slab forging and flame cut toprofile during cooling from normalising, followed by a slow cool. The samemistaken ideas about mild steel permitted some laxity in transport arrangementsincluding multiple loading and long delays in returning to the furnace forcooling after flame cutting. At the same time the termination of the cut wasvery close to the commencement, where surface hardening due to cooling hadalready occurred, and the combination of these faults resulted in stresscracking and rejections. (Fig. la) Modification of the cutting route as shownin figure 1(b), preheating the untreated slab to 500°C and charging into a warmfurnace for normalising immediately after profile cutting, with stricter controlof transport, has completely eliminated stress cracking.

We consider that flame cutting should only be employed where cost savingsor non-availability of suitable other equipment justify it. With this product aswe have ample facilities for cold trepanning the holes, at no greater cost thanmachining from flame cut holes, this is our normal practice. It reduces thetime taken flame cutting and thus the need for a higher preheat.

Cutting narrow shapes in thick sections necessitates prior considerationof distortion due to expansion and contraction, e.g. if the inner face of athin ring segment is cut first, the straightening out as the outer face is cutresults in lack of material on the inner face as shown (See Fig. 2).

Movement of either the forging or the offcut can produce a similar effect,at the same time being a danger to the operators. Full support should beprovided to both forging and offcut throughout the flame cutting process.

Production of parts for the built-up crankshafts for the large Doxfordtype diesel engine has always been expensive. Because of the large section,short lengths and complicated shapes, forging to shape and size is difficult,sometimes impossible, with a high rejection rate if persisted in. On the

other hand, machining from relatively simple forgings is expensive due tothe excessive amount of surplus material which has to be removed.

On a forging similar to that shown in Figs. 3(a) and (b) machining awaythe surplus material beside the pins would require in excess of 150 hours ona large lathe and cost $1,750 - 2, 000. Removal by flame cutting costsapproximately $125 including rough shaping of the pins and the preheatingprior to flame cutting. This cost saving, whilst appreciable, is not the onlyfactor. Increased availability of such a lathe is in many works just as vital.

It is essential withthis type of forging that the steel is adequately worked,if the resultant forgings are to be entirely suitable for the duties they perform.

The heavy and varied sections on this forging make it difficult to handle,both before and after flame cutting and the consequent increase in handlingtime, combined with the length of time actually flame cutting, make it necessaryto preheat to 650°C. Adequate supervision of cutting and transport areessential to ensure there are no problems due to cooling off before return toa warm furnace for the necessary heat treatment.

Even with relatively simple forgings, machining can be considerably reducedby flame cutting. The Flywheel shown in Fig. 4 was forged as a plain upsetcheese with a large peripheral bulge. This was removed by flame cutting inaddition to the quadrants between the arms.

Generally speaking, it is more satisfactory to complete a forging beforeflame cutting, but there are cases where flame cutting at an intermediate stagecan simplify manufacture. This is ideally illustrated in the manufacture of arudder frame.

A relatively simple shaft and slab forging is flame cut to size and shape(Fig. 5) before bending the arms to their correct positions (Fig. 6).

Progress towards cutting greater thickness was steady up to 1, 370 mmwhen the limit appeared to have been reached.

High pressure cutting not being successful, low pressure cutting withreduced cutting speeds was attempted. The only result was a wider kerfor cut and a wider wash at the bottom of the cut, still without penetrationbeyond 1, 370 mm. Every combination of pressure, cutting speed andvariation of fuel mix had a similar result, a maximum penetration of 1, 370 mm.

The solution was eventually provided by solving another problem.

Certain 500 mm diameter alloy steel billets were called for in increasingquantities, demanding a high rate of production. Flame cutting to length wasnot expected to be a problem but, after com letion of for in , the heat contentat the axis was so great, that the additional n ut of flame cuttin formea ool o mo ten meta w a ooned, and prevented further cuttin .obvious solution was to orge a a ower empera ure and to allow the billetsto cool to 450°C or less on the exterior before flame cutting; Not only would

this have impeded production but it could have resulted in metallurgicalproblems with the steel.

As an experiment, cuttin with an oxygen lance was tried, using thelance like a saw. This was an ins a s ccess, ressure rom e ance

ng e mo en e • e ig er eat conten ecame avan age instead of a disadvantage. The quality of cut was not good at first,

but with very little practice a skilled operative was soon able to produce aclean straight ru .

Lancing had long been used in the scrap metal industry for piercing deepholes to insert blasting charges. Oxygen at 12.6 kg/sq. cm. in pressurethrough a 6mm bore mild steel tube will blow a bead of molten metal throughseveral metres of steel in a matter of seconds, particularly if the steel is hot.

In lance cutting, the consumable lance is used like a hand saw, startingfrom the top of the section and working diagonally across it. (Fig.7). Tospeed up the process, even it o for in s an ordin r s rai h tintorch is used to melt a ead of metal at the be innin of each stroke. Theoxygen s ream rom t e ance projects t is ea an bea s or its own meltingtip through the steel ahead of it and each stroke (like the stroke of a saw)deepens the cut in the section. Skilful handling of the lance is essential toavoid undercutting whilst still maintaining the maximum depth of cut.

As the melting steel and lance provide their own molten beads the depthof cut is unlimited, the length of lance and the skill and stamina of theoperator being the controlling factors.

A portable adjustable height platform is a great advantage, enabling theoperator to work on a familiar footing. During the majority of the operationhe is unsighted and should therefore know the limits of hia movements. Heshould also be assured of a constant supply of oxygen at a fixed pressure.Any sudden variation could upset his balance and prove dangerous.

The consumable lance tubes used are standard 6 mm nominal bore tubing(BS 1387) in 7 metre lengths screwed and socketed at the ends and fitted intoa lance holder (with on/off tap) connected direct to the oxygen pipe line. Ona large diameter billet, say 2,300 mm, 15 - 20 such lances will be consumed,so that changing must be simple and quick.

Whenever possible on large diameters, it is preferable to employ severaloperators working in relays. This permits more or less continuous working,reduces operator fatigue and consequently produces a better cut. As well asbeing strenuous, this work calls for extreme concentration. Although themain stream of metal flow is away from the operator, the slightest obstructionor false stroke by him will result in a fly-back of molten metal and he must beadequately protected against metal splash. Protective clothing, however,should be as light and loose as possible to reduce fatigue. An open back tothe jacket reduces its weight and allows free air circulation.

At Japan Steel Works they use a lance applied at the bottom of the cut toassist their large flame cutting machine (Fig.8) and achieve cuts of up to3,260 mm diameter. This method ensures good control of the cut but losesflexibility because the forging has to be carried to the machine.

+

Our method relies more upon the skill of the operator for the quality ofthe cut, but, because the operator goes to the job, the increased flexibilityis well worth while and even at 2, 300 mm diameter and over, waviness andout of square are limited to 50 mm.

1........Cuttin at the ress has now been eliminate com letel resultin in

longer press too i e d or in s wit square cut end faces entirel suitableor presenting to a machine face plate arti 1 in r izes.

ee ig. .

Our first really big cut was 2, 350 mm diameter, taking exactly 58 minutesand consuming approximately 142 cubic metres of oxygen. The strokes of thelance can be clearly seen on the photograph (Fig. 10).

Lancing is a strenuous occupation calling for special abilities. Strongcapable men who are adaptable and prepared to work in unusual and uncomfortableconditions are essential. Because they must be able to go to the job ratherthan bring the job to them, it is necessary to make adequate provision for this.

In our plant oxygen at 16.8 kg/sq. cm. is carried in a 150 mm diameteroverhead main, with 50 mm downtakes at every point where the operators arecalled upon to work. Propane is similarly carried in a 75 ram diameter mainat 4.2 kg/sq. cm.

As previously mentioned, when lance cutting the oxygen is fed direct fromthe supply line, without intervening valves or gauges, except the on/off tap.

It has also become our standard practice with orthodox flame cutting todispense with gauges on the oxygen supply; pressures and cutting speeds beingadjusted to suit the performance entirely from experience. Only with deseamingdo we insert a regulator between the oxygen supply line and the torch, settingthe pressure at 12.6 kg/sq. cm. This may seem strange but conditions varyso much from day to day and job to job that laying down specific pressures andcutting speeds has not consistently produced the best results. For normalflame cutting, propane is controlled at 1,75 kg/sq. cm.

Thicknesses up to 750 mm are cut using a straight nozzle mix torch with4 rnm bore nozzles, the resultant kerf or cut width being approximately 12 mm.Above 750 mm - 9 mm and 12 mm bore nozzles are used, the kerf widthincreasing to 25 mm to 38 mm. All nozzles have Copper outer and Brassinner sections to prevent welding together during operations. Although thelargest cutter is equipped for water cooling, this is not used as we have neverfound it to be necessary.

The successful use of lance cutting and in particular its ability to bridgea gap, such as the cavity in an ingot head, has been a major factor in theintroduction of cast hollows for the making of hollow forgings. See Figs. 11and 12.

Previous to their introduction our practice was to anneal the solid ingot,cold part the discards and trepan a hole through the billet. Carefully controlledpreheating followed, before the large alloy steel billet could be charged intothe forge reheating furnaces for normal heating.

With a cast hollow, the discards can be removed by lance cuttingimmediately after stripping from the ingot mould, and the resultant hollowbillet charged into a hot forge furnace for normal heating.

Not only are the savings in process time and cost* significant but so isthe elimination of risk attached to the reheating of a large steel billet fromcold - always a hazard with alloy steels.

*When using a 100 ton Cast Hollow (in 1% Cr. Mo. ) the saving insteel alone more than compensates for the cost of lance cuttingthe discards ($125), and the additional costs involved in thecast hollow process, and the process time is reduced by atleast three weeks.

The costs involved previously for

1. Annealing of ingot 2, 2002. Parting discards and trepanning hole 1, 1503. Preheating from cold before charging

into forge furnace 300

are all saved completely $ 3, 650

At the beginning of this paper reference was made to the deseaming offorgings. As stated, the function of a deseaming torch with its largepreheating nozzle is to melt the surface metal and to blow this along in frontof the nozzle making a continuous groove across the forging for the completestroke of the torch. With shallow surface defects this is entirely satisfactoryand a defect or series of defects can be removed with a few sweeps of thetorch.

With deep cracks, the technique mast be different. A deep narrow groovewould lap when reforged, so obviously the groove formed must be shaped toavoid this effect.

Following the actual line of the surface tear or crack can actuallyaggravate and extend it and the approach should be to remove the materialeither side of the defect whilst forming a groove of acceptable shape (Fig. 13).This allows the defect to be undercut without penetrating deeper than necessaryand at the same time it means melting a lesser volume of metal, some solidmetal falling away. As the groove is kept smooth and open, the molten metaland slag flow freely and do not become trapped, as they would if the torchattempted to follow the defect.

A problem is using a torch long enough to get the operator away from thehot forging, but not too long that he cannot control its movement. Torches1, 500 mm long with two adjustable hand grips appear to be an acceptablecompromise, but 12. 6 kg/sq. cm. through a 16 mm bore still requires skill,as well as strength, to manipulate. Our operators never use starting rods,

6

preferring to remove the rod feed mechanism and reduce the weight ofthe torch.

Deseaming can also be of great assistance when dealing with a steel whereimpurities and surface defects are anticipated but are difficult to detect bynormal visual examination. During deseaming, impurities break up the flowof molten metal and can easily be detected, if present, through colouredglasses, when lightly deseaming the apparently clean and defect free surfaceof an ingot or forging.

It is difficult to estimate the exact savings achieved by deseaming.Some forgemasters even claim that, since deseaming was introduced, steelmakershave tended to ignore ingot surface conditions and that deseaming has becomethe practice rather than the exception, consequently increasing overall costs.

This is a negative approach, and if it is allowed to happen, merelyhighlights lack of liaison between forge and melt shop. Immediate remedialaction is required.

Deseaming can establish the existence and extent of surface defects andeffectively remove them, and it ensures that the forgemaster is reasonablyconfident when his forging goes for heat treatment that it is satisfactory. Itis certainly a great improvement upon the previous indeterminate practice ofleaving excess material wherever surface defects were evident.

Deseaming can also be used as a shaping process. For example, askilled operator can produce an acceptable shape on the sneck of a rudderstock, avoiding a very difficult and expensive machining operation which mayhave to be carried out for weight removal and appearance purposes only. (Fig. 14).

Flame cutting and lance cutting of many qualities of stainless steel havebeen successfully undertaken, without the use of powder, including

A1S1, 301, 302, 304, 316, 317, 321, 347A1S1, 410, 416, 420A1S1, 430and 15/5 PH

Although powder cutting is undoubtedlyunless there is a continuous demand for it,site, make it an unnecessary refinement.achieved without its use.

18/8 Austenitic SteelsMartensitic Chromium SteelsFerritic Chromium SteelPrecipitation hardening steel

faster 4nd produces a smoother cut,the rig-up time, and inflexibility ofQuite satisfactory cutting is being

Recent problems with a "difficult to shape" stainless steel (A1S1 410)have been overcome by the use of "powder deseaming" or "powder washing".

Like many stainless steels, when shaping tools are used to produce a taper,the dwell of the tool causes surface chilling and results in severe surface tearing.Even rough tapering with normal flat tools causes similar tearing and meansleaving excess material which is very costly to machine away. Normal powdercutting produces a kerf or cut which is very wide at the base and still means

leaving more material to machine away than desirable.

The solution has been to reduce the section in a series of gentle steps toavoid severe surface tearing (Fig. 15). Lance cutting roughly to shape(without powder) has been followed by powder washing to achieve a smoothersurface nearer to finished size (Fig. 16).

This is the only process where acetylene is used in preference to propane -its higher heat content being an advantage.

The special deseaming torch differs from the normal in that it has anadditional tube through which the powder is blown by compressed air. Thenozzle is rectangular with a flat slot shaped central orifice, in front of whichthe powder is continuously fed throughout each stroke. Once again the skillof the operator is the major factor in the success of the process.

By combining flame cutting and deseaming many single purpose forgingtools can be provided quickly and cheaply, permitting economies in materialand machining.

The tool shown in Fig. 17 was made within a few days of receiving theorder for the vessel. Flame cutting to shape from stock plate and cornersradiused by deseaming, its use produced a bottled end which required aminimum of machining. Previous similar shaped forgings, produced inordinary forging tools, have been very roughly shaped, have required excessivemachining and have even been rejected for failing to conform to the shaperequired.

Reference has already been made to the need for protective clothing.This should be as light as possible but adequate to resist metal penetration.Thermofoil covering in addition to reflecting heat forms an effective shieldagainst metal splash.

Various methods of keeping the operator cool have been proposed,including the blowing of cold air into his suit. These methods all impedethe free movement of the operator and have had to be abandoned. In additionto reducing weight, the open-back jacket allows the operator to remove itquickly in an emergency.

Adequate protection of the hands is difficult. Shields fitted on deseamingtorches can only provide partial protection. Larger shields would increasethe weight and impede the manipulation of the equipment. Gloves or mittensmust be pliable to allow the operator full control of valves, etc. Thermofoilbacked mittens 400 mm long have proved the most acceptable to date. Anythingshorter exposes the wrists and forearms, while increased length restricts armmovement. The practice of operators quenching their gloves in an adjacentbucket of water is dangerous because of the danger of scalding and should bediscouraged.

- 8 -

Whilst goggles provide adequate eye protection during flame cutting andmost lancing, thermofoil covered hoods are preferred for deseaming. Heavierand more tiring to wear than goggles, the free movement of air over the faceand eyes, with rapid restoration of normal vision by the lifting of the vizordo make them more suitable for long periods of working in the heat. Thethermofoil covers are secured by press-studs for ease of replacement.Because of the problem of metal splash all eye protection is by double glass-plain toughened glass in front - with coloured toughened glass behind to protectagainst the intense light. For flame cutting and lancing B.S. 679G. OW lensesare adequate but deseaming requires a denser lens to B.S. 679-11EWP. Ourpreference for green rather than blue glass is because it is less tiring.

The hot and dirty conditions, in which operators have to work, call forthe provision of facilities for washing and resting, and for drying their clothes.These facilities should be adjacent to their working area to avoid thepossibility of chills.

Fume control is a difficult problem, especially when working on steelswith a high chrome content. If the forging can be taken to a special cuttingarea, pits, for collecting the swarf, can be fitted with powerful suction fansto remove the fillies fairly effectively.

With lancing and deseaming being normally carried out adjacent to thepress, little can be done, except the installation of powerful roof fans. Ahigh well ventilated forging shop obviously has a tremendous advantage, andthis factor should be taken into consideration, when laying out a new forge.

The handling of forgings, after flame cutting, can be hazardous. Wherecutting has taken place, edges are square and sharp, with a hardened surface,and wire slings and chains can be severely damaged, particularly if lifts arenot slow and steady. It is a wise precaution to provide some protection forlifting equipment, either wooden supports if the forging is relatively cool,or bent steel plate corners if dealing with hot forgings. Off-cuts shouldalways be cleared first, to allow free access to the heavier items, wherebalance may be difficult because of the altered shape.

All cutting equipment should be examined daily, before use, and dry warmstorage facilities should be provided to avoid deterioration.

I understand that in the U.S.A. you have problems due to purity of oxygenand quality of fuel gases.

Our oxygen supplies are received from a bulk oxygen plant situated6 Kilometres from the works. It is produced to a guaranteed 99. 5% purityat 42 kg/sq. cm. and throttled back to 16.8 kg/sq. cm, for our internal pipelines. Previously an evaporator in the works was filled from a tanker, asrequired, and operated at a similar pressure.

The fuel gas is stored outside the forge buildings in 1,830 ram tanks,being topped up at regular intervals.

We use propane enriched with 39% propylene with traces of butane, etc.and find this enrichment essential to obtain the quality and depth of cut we require,

9

If your supplies are not up to this standard, you must do something aboutit. Convince your suppliers that the demand is only there if they can matchyour requirements.

In conclusion I would like to express my appreciation to Japan Steel Worksfor information supplied and to British Steel Corporation for allowing me topresent this paper. At the same time, I would pay tribute to a retiredcolleague, .Albert Revitt. A silver smith by trade, flame cutting anddeseaming presented a challenge to him. A man with the constitution ofan ox, he has continually attempted to prove that the impossible only takesa little longer. Much of our advancement has been due to his courage anduntiring efforts. He has been fortunate in having operators who followedhis exarrple.

- 10 -

English Equivalents of Approximate Metric Figures Quoted in Text.

4 mm6 mm9 mm

12 rnm16 mm25 mm38 mm50 mm75 mm

125 mm150 ram300 mm350 mm400 mm500 mm750 mm

1370 mm1500 mm1830 mm2300 mm2350 mm3260 ram

7 metres

6 Kilometres -

142 cubic metres

1.75 kg/sq. cm.4.2 kg/sq. cm.

12.6 kg/sq. cm.16.8 kg/sq. cm.42 kg/sq. cm.

5/3 21/ 43/81/25/81"11 "2aa3 II

5.11

612 II

141620 tl

3054 "60729092 II

128

22 ft.

4 miles

5000 cu. ft.

APPENDIX 1

25 lb/sq. in.60 lb/sq. in.

180 lb/sq. in.240 lb/sq. in.600 lb/sq. in.

TEST

STRESSCRACKS

FIG.Ia ORIGINAL FLAME- CUTTINGROUTE

TEST

FIG. I b MODIFIED FLAME-CUTTINGROUTE

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IGURE 3(b), MACHIN INC, A 2-JOURNAL FORGING

FIGURE 4. FLAME CULT ING A FLYWHEEL

FIGURE 5. RUDDER FRAME

FLAME-GUT BEFORE BENDING

FIGURE 6. RUDDER FRAME

AFTER BENDING

-

FIGURE 7. LANCE-CUTTING

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Mill

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FIGURE 9. WATER TURBINE SHAFT WITH FLAME CUT ENDS

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FIGURE 10. 92" (2300m.) BILLET

CLEARLY SHOWING STROKES OF LANCE

FIGURE 11. CAST HOLLOW

FIGURE 1 2. SECT IONED

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TO AVOID SURFACE TEARING

FIGURE 16.

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ABSTRACT

"New Integrated 4000 Mp Forging Press Plant" by H. Hojas and Associates;Bohler Bros. & Bo., Ltd., Kapfenberg, Austria

The high grade steel forge shop at Kapfenberg has been re-equipped accord-ing to an entirely new concept. A fully integrated 4000 Mp forging press plantand a continuous forging line, comprising the biggest bar forging machine inthe world, have been erected.

In this report, a description is given of the 4000 Mp forging press plant.The forging press plant has been planned for the manufacture of heavy forgingsand big bars. The plant consists of a water hydraulic 4000 Mp 3-stage push-downpress of 4 column design, a railbound 80 Mpm forging manipulator with pressureoil drive and accumulator, a mobile turntable, an oxy-acetylene cutting station,and five bogie-furnaces with a total hearth area of 104 m2. The forging plantis controlled by one man from a control pulpit. A digital measuring and controlsystem permits fully automatic drawing-out and finishing operations. Auxiliaryequipment such as a die storage bogie, a die changing device and a mobile turn-table permit uninterrupted forging and thus, add decisively to a high produc-tivity of the plant.

CONTENTS

1. Historical survey regarding the development of the high grade steel forgeat Kapfenberg.

2. Description of the complete forging plant.

3. Control and operation.

4. Design of 4000 rip hydraulic press.

5. Design and operation of 80 Mpm manipulator.

6. Summary.

1. Introduction

NEW INlEGRATED 4000 Mp FORGING PRESS PLANT

by

H. Hojas, H. Tarmann, M. Kroneis*

In November last year, the newly equipped high grade steel forge shop ofMessrs. Bohler Bros. & Co. ltd. at Kapfenberg, has been introduced to thegeneral public. The new forge shop stands on a site which has a long traditionregarding the manufacture of iron and steel. There exists documentary proofof the existence of water hammers in this area since 1610. Every new develop-ment in respect to forging equipment resulted in a conversion of the existingplant. The water hammers were displaced by steam hammers. At the beginningof the 20th century, there existed at the site of the present forge shop, asteam hammer with a tup weight of 20 tons. Water hydraulic forge presses of630, 1200 and 2400 Wp of forging pressure as well as seven forge hammers withtup weights from 1,5 to 2,5 tons were in operation until the final conversionof the forge shop. Deliberations and planning regarding the erection of a new4000 Mp press commenced during the second half of the Sixties and were sup-plemented by the planning work for a continuous forge line which was initiatedin 1969. The previous equipment was not any longer capable to meet the existingand future demands in respect to economy and quality for the production offorged bars and other forgings.

It was the aim to create a plant capable of producing, in the most econ-omical and most up-to-date manner, components for general engineering applica-tions, plants and tooling purposes. With a total expenditure of 250 millionAustrian Shilling, the antiquated 2300 Mp press and 40 Mpm manipulator werereplaced by a 4000 Mp forging press plant, whereas the 630 Mp press and theseven pneumatic hammers were replaced by a bar forging machine, which up tillnow is the biggest of its kind in the world. The production of the new forgeshop comprises bars in the range from 100 mm square or round up to 1200 mmround and 1000 mm square. The weight of the ingots being forged varies between500 kg and 60 tons. The new forge shop comprising a 4000 Mp press, a 1200 Mppress and the bar forging machine has been planned to attain double theproduction capacity of the previous shop in two-turns per day operation. Atthe same time, the number of workers in the forge shop could be reduced by 50 %.

* Bohler Bros. & Co. Ltd., High Grade Steel Mills, Kapfenberg, Austria

2. The 4 MfrinPesPant

2. 1. jay,:saltaf_21Aat

As can be seen from Fig. 1, the plant consists of a water hydraulic 4column push-down forging press, an BO Mpm forging manipulator, a mobile turn-table, an oxy-acetylene cutting station and 5 bogie hearth furnaces with atotal hearth area of 104 m2., The plant layout was determined by the realizationof the following demands:

Avoidance of idle times to reach optimum plant availability,High rate of forging and manipulation,Highest possible dimensional accuracy,Low personnel requirements.

In order to meet the first requirement, the plant layout was planned insuch a manner, that the press was reserved for the actual forging. The cuttingof crop ends was to be carried out by means of the oxy-acetylene cutting station.This permitted the press to be fitted with a mobile die storage bogie runningparallel to the axis of the forging plant, but preventing the press from beingused at a right angle from the plant axis. By means of the die storage bogiewhich can accommodate 4 pairs of dies, the time consuming changing of the dieswas reduced to 1 1/2 minutes. Since both the upper and bottom tools arechanged at the same time, the ram is fitted with automatic locking devices forthe dies. All actions such as the release of the upperdie, the removal of thepair of dies from the press, the travelling of the die storage bogie into itsnew position, the positioning and changing of the new dies are initiated fromthe control pulpit.

The utilization of the hydraulic press can be influenced decisively by asmooth exchange of the forgings. For this reason, we have developed a mobileturntable with a carrying capacity of 50 Mp (see Fig. 2). This turntable isdesigned to remove the forging as fast as possible from the working area afterforging. While the forging is removed by means of the turntable, the cranebrings along the next ingot and hands it over to the manipulator. The secondfunction of the turntable comprises the turning of the forging through 180°after one end has been forged so that the other end can be forged withoutreheating.

2. 2. Control and 0 eration of Plant

In order to attain a high rate of work, the press has been fitted with anoil hydraulic valve control. In this manner, shortest control times and ahigh rate of strokes are assured. Exact determination of the dimensions andrapid response of the electronic and hydraulic control equipment coupled withadequate reproduceability yield forgings of good dimensional accuracy. Thehydraulic press, therefore, is equipped with a measuring device for the forgings.The movements of the ram are transferred by means of a rack and pinion arrange-ment to an electronic impulse generator. These impulses are processed by adigital counter and indicated by luminous figures. By means of the measuringequipment in connection with the automatic drawing out and finishing control,

a forging accuracy of t 1 mm is obtained. Deviations of the bottommost posi-tion of the ram from the selected desired value are indicated at the controldesk and adjusted automatically.

The digital measuring and control system permits full automatic coordina-tion between the hydraulic press and manipulator. After preselection of thedesired forging dimension and length of the press stroke and preselection ofthe advancing steps of the manipulator, the forging process takes place in afully integrated manner.

The movement of the manipulator is controlled by the upper and lower deadcentre of the ram. The movements of the manipulator must take place duringthe period while the die has been lifted from the forging (see Figs. 3 and 4).The fully automatic operation can be employed also for the forging of rings.In this case the arc through which the forging is turned per stroke of thepress is preselected. The various control possibilities for the hydraulicpress and the manipulator have been adapted to the range of products. Thisrange reaches from simple bars to the manufacture of intricate forgings.

The operator of this plant has the possibility to actuate all individualcomponents of the various items of equipment by manual control. Moreover, hehas the possibility to switch one part of the plant, such as the hydraulicpress, on automatic control and to operate the manipulator by manual control,On the other hand, the operator may preselect all movements of both the pressand the manipulator for their automatic execution (see Fig. 5), The plant hasbeen supplied with all equipment for the installation of a punched tape controlat a later date.

The aforementioned equipment permits most of the forging manipulations tobe carried out by mechanical means and thus requires only a crew of 3 workerscomprising the press operator, the leading forger and a helper.

2. 3. The 4000 M -H draulic Press

In the past years the forging experts have had many discussions regardingthe developments and tendencies of the design of forging presses. The opinionsin respect to the expediency of the push-down or draw-down press, or oil orwater hydraulic operation have produced many heated arguments. An attentivelistener to these arguments, finally, was bound to come to the conclusion thatas far as presses of more than 3000 Mp are concerned, the two designs weretheir equals in regard to economy and rate of work. The question, which ofthe two systems should be preferred, is certainly to be decided from case tocase, At Kapfenberg we have selected a water hydraulic 4 column push-downpresswithaccumulator. For this decision, the following reasons were thedetermining factors:

Since the headroom of the existing press bay was sufficient for apush-down press no savings could be expected from the installationof a draw-down press.

There existed also a pressure water accumulator plant which only

had to be supplemented by two triple pumps with a capacity each

of 1500 litres/minute.

The geological situation in the region of the press foundation

is characterizedby inclined waterbearing strata, which, onaccount of a directly adjoining mountain slope not only deliver

considerable anounts of ground water but are subject to constant

shifting. The considerably deeper excavation and foundation for

a draw-down press would have brought about serious difficulties.

The water hydraulic operation'of the press is much simpler and

repairs can be carried out by our own maintenance section. The

maintenance personnel possesses longstanding experience in the

maintenance of water hydraulic equipment.

In a forge shop concerned with the forging of medium and high

alloyed steels, it is necessary during the forging operation to

remove surface defects by flame scarfing. Experience has shown

that the highly sensitive oil hydraulic motors and pumps do not

fully stand up to pollution caused by flame scarfing and thus

may result in costly maintenance work or accidents.

The press (see Fig. 6 and Table 1) is designed as 3-stage press. The

available forge pressures of 1750, 2750 and 4500 Mp can be well adapted to the

hot working exigencies regarding the size of forging and the type of steel.

Further important design features are:

For the transmission of bending forces during excentric forging

the centre plunger is rigidly connected to the slide. In this

manner, the wear of the slide bushings during excentric forging

is considerably reduced. The two side plungers act on the slide

via spherical bearings so that seizing of the plunger due to

excentric forging or thermal expansion of the slide, is avoided.

The post bushings are suspended in the slide by means of spherical

supports and in case of excentric forging adjust themselves to

the inclination of the slide.

To avoid the fracture of theposts immediately above the press bed,a risk push-down presses are likely to run, the posts are mounted

in conical bushings in the press bed. In this manner, the counter

nuts above the press bed can be dispensed with. Excessive notch

stresses in the thread of thecounter nut, occurring in olderpresses are thus avoided.

The press stands on column supports and is not secured by anchor

bolts, In this way damage of the column foundations by the pulling

action of the anchor bolts is avoided. To improve the stability of

the press, supporting brackets are provided on both sides of the

under side of the press bed on which the press will support itself

if an excessive inclination is caused by severe excentric loading.

2. 4. The Manipulator

Fully integrated forging plants require some special design features farthe manipulators. In the past, several designs have been developed, whichdiffer mainly by the control of the advance and return steps. They may be sub-divided into three main types:

In the first type, the tong support is suspended in the manipulatorframe so that it can move along the longitudinal axis. During theforging, the manipulator frame moves at a constant speed while thetong support follows the movement of the frame in jerks accordingto the rhythm of the forging strokes.

In the second type, the tong support and drive move together, themanipulator frame is joined to the tong support by means of a springsystem and thus follows the movements of the tong support.

In the third type, the manipulator carries out definite stepsaccording to the rhythm of the forging press.

The advantage of the first two types lies in the smaller masses which are beingmoved. For fully integrated plants, however, complicated and expensive designsand control equipment are required. The third design corresponds to theprinciple employed at Kapfenberg. This principle demands a powerful drive butthe design is simpler. The control of the advance and return steps can bemade according to the time or according to the distance travelled. The designwhich offers better safety in service is the control according to time but theaccuracy of the distance passed through per step cannot be maintained with thesame accuracy as with the control according to the distance travelled. In thecase of our plant, it has been found, however, that it was less important tocover a certain distance per step than to maintain a uniform distance per stepduring a working cycle. It is only by this uniformity of the manipulatorsteps; that the desired tolerance of ± 1 mm can be maintained (see Fig. 7 andTable 2). The opening range of the tong has been made very wide to meet therequirements. It covers the range from 500 - 1700 mm and thus permits theclamping of the foot of even the biggest ingot which is used in Kapfenberg.The manipulator can carry out the following movements:

The advance and return takes place along ra ls whereby a definiteapplication of the driving force is assured by two racks which areanchored in the foundation.

The tong support can'be lifted and lowered in a parallel movement.However, there exists the possibility to swivel the tong supportby 120 upwards and downwards from its horizontal position. In thismanner, forgings can be taken up from the shop floor. The tong canbe turned continuously in both directions. An electrohydrauliccontrol permits the turning through an indefinitely preselected angle.

The manipulator described here has a driving power of 260 kW and is thus themost powerful in Europe and the biggest, fully integrated manipulator in theworld,

2. 5, The Ox -Acet lene Cuttin Station

In the past years oxy-acetylene cutting stations have become indispensablefor big forges. The plant erected at Bohler permits straight cuts up to twometres thickness and conture cutting up to 1200 mm thickness. (See Fig. 8)A light tracing equipment permits the conture cutting from drawings at thescales 11 and 1:10.

3, Summary

Within the complete re-equipment of the high grade steel forge of Kapfen-berg a fully integrated 4000 Mp forging press plant for the forging of bigforgings and big bar sections has been erected. The plant consists of a waterhydraulic three-stage 4000 Mp push-down press of 4-column design, a railboundBO Mpm forging manipulator with pressure oil accumulator drive, and a mobileturn table.

A digital measuring and control system permits fully automatic drawing outand finishing.

Auxiliary equipment such as die storage, a quick changing device for thedies and a mobile turntable guarantee continuous production and thus assureproductivity.

Table 1

Technical Data of 4000 Mp Forge Press

Operating PressureMax. Press Force (static)Press Force of Central PlungerPress Force of Side PlungersSlide Lifting Force

Plunger Stroke

Plunger DiametresMax. Pressing Speed

Max. Descending Speed

Max. Lifting Speed

Diameter of ColumnsDistance Between Centres of Columns 2400 x 4000 mmClearance Between Die MountingsMobile Press BedTravel of Press Bedin Both Directions

Height of Press AboveFloor Level

Max. Depth Below Floor LevelWeight

200 bar

4500 Mp (44,2 MN)2750 Mp (27,0 MN)1750 Mp (17,2 MN)

450 Mp ( 4,4 MN)2000 mm

0 750/750/1320 mm100 mm/S300 mm/S200 mm/S

0 540 mm

4000 mm2400 x 5500 mm

3 000 mm

10.225 mm

3850 mm

750 t

Table 2

Technical Data of 80 Mpm Forging Manipulator

Tong Clamping Moment

Lifting Capacity

Max. Opening of Tong

Min. Opening of Tong

Max. Turning Cycle of Tong

Max. Height of Tong CentreOver Upper Edge of Rails

Min. Height of Tong CentreOver Upper Edge of Rails

Thus, Parallel Adjustmentof Height

Angle of InclinPtionof Tong Support fromHorizontal Position

Max. Turnipg Speed of TongLifting and Lowering Speedof Tong Support

Max. Travelling Speed

Max. Operating Pressure

Drive Consisting of 2Electric Motors of 130 kWPower at

The Electric Motors Drive2 Pumps With a Capacity of

Weight

80 Mpm (0,8 MNm)

50 Mp (0,5 MN )

1700 mm

500 mm

3000 mm

23500 mm

1550 mm

800 mm

12° (0,2 rad)

15 rpm

150 mm/second

60 m/minute170 bar

1000 rpm

430 litres/minute

239 t

-;43kk

L:44pit.f

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6th international forgemasters meeting, cherry hill 1972

Documents