YU ISSN 0372-8633 ZELEZARSKI ZBORNIK VSEBINA STRAN CONTENTS PAGE Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: VPLIV VARILNE TEHNOLOGIJE IN IZBIRE DODAJ-NEGA MATERIALA NA LOMNE LASTNOSTI EPP ZVARNEGA SPOJA NA NIZKO OGUIČNEM FINOZRNATEM JEKLU 117 Rodič Jože: KOBALTOVE ZLITINE V LESNI INDUSTRIJI 127 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: FAKTOR MEJNE INTENZITETE NAPETOSTI PRI POČASNEM NATEZANJU NAVODIČENE-GA JEKLA Z VISOKO TRDNOSTJO 139 Grozdanič Vladimir, J. Črn-k o : KOMPJUTORSKA SIMULACIJA SKRUČIVANJA ODLJEVAKA KOMPLEKSNE GEOMETRIJE 149 Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: THE INFLUENCE OF VVELDING TECHNOLOGY AND VVELDING MATERIAL SELECTION ON FRACTURE PROPERTIES OF SUB-MERGED ARC VVELDED, LOW CARBON, FINEGRAINED STEEL PLA TE 117 Rodič Jože: COBALT BASE ALLOYS IN VVOODCUTTINGINDUSTR Y 127 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: TRESH-OLD STRESS INTENSITY FACTOR AT SLOW STRETCHING OF HYDROGEN CHARGED HIGH-STRENGTH STEEL Grozdanič Vladimir, J. Črnko: SOLIDIFICATION SIMULATION OF CASTINGS OF COMPLEX GEOMETRY ? 139 149 LETO 25 ŠT.4-1991 ŽEZB BQ 25 (4) 117-164 (1991) IZDAJAJO ŽELEZARNE JESENICE, RAVNE, ŠTORE IN METALURŠKI INŠTITUT 1 •sliH Hi ■ i" < ŽELEZARSKI ZBORNIK Izdajajo skupno Železarne Jesenice, Ravne, Štore in Metalurški inštitut Ljubljana UREDNIŠTVO Glavni in odgovorni urednik: J. Arh Uredniški odbor: A. Kveder, J. Rodič, A. Paulin, F. Grešovnik, F. Mlakar, K. Kuzman, J. Jamar Tehnični urednik: J. Jamar Lektor: C, Martinčič Prevodi: A. Paulin, N. Smajič (angleški jezik), J. Arh (nemški jezik) NASLOV UREDNIŠTVA: Železarski zbornik, SŽ-Železarna Jesenice, 64270 Jesenice, Yugoslavia TISK: TK Gorenjski tisk, Kranj IZDAJATELJSKI SVET: prof. dr. M. Gabrovšek (predsednik), Železarna Jesenice dr. B. Brudar, Iskra, Kranj prof. dr. V. Čižman, Univerza v Ljubljani prof. dr. D. Drobnjak, Univerza v Beogradu prof. dr. B. Koroušič, Metalurški inštitut Ljubljana prof. dr. L. Kosec, Univerza v Ljubljani prof. dr. J. Krajcar, Metalurški inštitut Sisak prof. dr. A. Križman, Univerza v Mariboru dr. K. Kuzman, Univerza v Ljubljani dr. A. Kveder, Metalurški inštitut v Ljubljani prof. dr. A. Paulin, Univerza v Ljubljani prof. dr. Z. Pašalič, Železarna Zenica prof. dr. C. Pelhan, Univerza v Ljubljani prof. dr. V. Prosenc, Univerza v Ljubljani prof. dr. B. Sicherl, Univerza v Ljubljani dr. N. Smajič, Metalurški inštitut v Ljubljani prof. dr. J. Sušnik, Zdravstveni dom Ravne dr. L. Vehovar, Metalurški inštitut Ljubljana prof. dr. F. Vodopivec, Metalurški inštitut Ljubljana Published jointly by the Jesenice, Ravne and Štore Steelvvorks, and The Institute of Metallurgy Ljubljana EDITORIAL STA F F Editor: J. Arh Associate Editors: A. Kveder, J. Rodič, A. Paulin, F. Grešovnik, F. Mlakar, K. Kuzman, J. Jamar Production editor: J. Jamar Lector: C. Martinčič Translations: A. Paulin, N. Smajič (English), J. Arh (German) EDITORIAL ADDRESS: Železarski zbornik, SŽ-Železarna Jesenice, 64270 Jesenice, Yugoslavia PRINT: Gorenjski tisk, Kranj EDITORIAL ADVISORY BOARD: prof. dr. M. Gabrovšek (Chairman), Iron and Steel VVorks, Jesenice Dr. B. Brudar, Iskra, Kranj Prof. Dr. V. Čižman, University of Ljubljana Prof. Dr. D. Drobnjak, University of Belgrade Prof. Dr. B. Koroušič, Institute of Metallurgy, Ljubljana Prof. Dr. L. Kosec, University of Ljubljana Prof. Dr. J. Krajcar, Institute of Metallurgy, Sisak Prof. Dr. A. Križman, University of Maribor Dr. K. Kuzman, University of Ljubljana Dr. A. Kveder, Institute of Metallurgy, Ljubljana Prof. Dr. A. Paulin, University of Ljubljana Prof. Dr. Z. Pašalič, Iron and Steel VVorks, Zenica Prof. Dr. C. Pelhan, University of Ljubljana Prof. Dr. V. Prosenc, University of Ljubljana Prof. Dr. B. Sicherl, University of Ljubljana Dr. N. Smajič, Institute of Metallurgy, Ljubljana Prof. Dr. J. Sušnik, Health Centre, Ravne Dr. L. Vehovar, Institute of Metallurgy, Ljubljana Prof. Dr. F. Vodopivec, Institute of Metallurgy, Ljubljana Oproščeno plačila prometnega davka na podlagi mnenja Izvršnega sveta SRS — sekretariat za informacije št. 421-1/172 od 23. 1. 1974 1(229280 ZELEZARSKI ZBORNIK IZDAJAJO ŽELEZARNE JESENICE, RAVNE, ŠTORE IN METALURŠKI INŠTITUT LETO 25 LJUBLJANA DECEMBER1991 Vsebina Stran Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: Vpliv varilne tehnologije in izbire dodajnega materiala na lomne lastnosti EPP zvarnega spoja na nizko ogljičnem finozrnatem jeklu 117 Rodič Jože: Kobaltove zlitine v lesni industriji 127 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: Faktor mejne intenzitete napetosti pri počasnem natezanju navodičenega jekla z visoko trdnostjo 139 Grozdanič Vladimir, J. Črnko: Kompjutorska simulacija skručivanja odljevaka kompleksne geometrije 149 Contents Page Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: The Influence of Welding Technology and Weid-ing Material Seiection on Fracture Properties of Submerged Are VVeided, Lovv Carbon, Fine-grained Steel Plate 117 Rodič Jože: Cobalt Base Alloys in Woodcutting industry 127 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: Treshold Stress intensity Factor at Slow Stretch-ing of Hydrogen Charged High-Strength Steel 139 Grozdanič Vladimir, J. Črnko: Solidification Simulation of Castings of Complex Geometry 149 ZELEZARSKI ZBORNIK IZDAJAJO ŽELEZARNE JESENICE, RAVNE, ŠTORE IN METALURŠKI INŠTITUT LETO 25 LJUBLJANA DECEMBER1991 Vpliv varilne tehnologije in izbire dodajnega materiala na lomne lastnosti EPP zvarnega spoja na nizko ogljičnem finozrnatem jeklu The Influence of Welding Technology and Welding Material Selection on Fracture Properties of Submerged Are Welded, Low Carbon, Finegrained Steel Plate I. Rak1, V. Gliha2, F. Vodopivec3, M. Tavčar2 1. UVOD Razvoj nizko ogljičriih jekel, ki so izdelana na termo-mehanski način ali kaljena in popuščana in ki so uporabljana v konstrukcijah z zahtevnimi obremenitvenimi pogoji, je v veliki meri spremenil varilno tehnologijo. Zaradi nizkega ogljika in nizke vsebnosti difuzijskega vodika v zvaru izdelovalci jekel ne priporočajo več predgrevanja pred varjenjem. Pri tem ni nevarnosti, da bi v TVP nastopila razpokljivost v hladnem, kar je mogoče npr. preveriti z Durenovim in Suzuki konceptom /1, 2/. Če navedeno prenesemo na dejanske zvarne spoje (hlajenje pod 50° C po vsakem varku in začetek varjenja brez predgrevanja), ki so zvarjeni z nizko dovedeno toploto (10—15kJ/cm), da bi zajamčili dobro žilavost v grobozr-natem predelu TVP, se lahko v raztopljenem zvaru pojavi nizka žilavost kot posledica tvorbe podolgovatih M/A strukturnih faz glede na kemično sestavo jekla in dodajnega materiala. V primeru napetostnega žarjenja pri 580° C se pojavi neugodni efekt izločevanja Fe3C iz M/A faz, ki predstavljajo prenasičeno raztopino. Pri tem je žilavost lahko še nižja, če je staljeni zvar občutljiv na re-verzibilno popuščno krhkost. V tem prispevku so obravnavane samo lastnosti raztaljenega dela zvarnega spoja. Zaradi jasnega efekta vpliva znižanega dovoda toplote pri varjenju in zaradi zmanjšanja velikosti zrna in primarnega ferita po kristalnih mejah je bil izbran za raziskavo dodajni material z 0,4% Cr in 0,2% Mo ter z dodatkom Ti-B. Uporabljena je bila metoda elektro varjenja pod praškom - EPP. Žilavost zvara je bila ugotavljana z udarnim Charpijevim kladivom. Zareze v Charpy preizku-šancih so bile locirane v kovini in korenski legi X simetričnega zvara pravokotno na debelino. Razlog za to je bilo pričakovano različno razmešanje, predvsem s stališča Nb (NB = 0,01 % v krovni legi in 0,04% v korenski legi). Preizkusi so bili opravljeni na celotni debelini zvara po metodi CTOD z upogibnimi preizkušanci, zarezanimi in utrujanimi pravokotno na debelino v sredini zvara. Lomne površine so bile preiskane z vrstičnim elektronskim mikroskopom predvsem glede pojava in lokaci- 1 Tehniška fakultata Maribor, VTO-S, 2 RRI, Metalna Maribor, 3 Inštitut za kovinske materiale in tehnologije 1. INTRODUCTION Development of LC steels, produced on thermome-chanical way or by quenching and tempering and used in constructions under sophisticated load conditions, has largely changed the vvelding technology. Because of the L C and lovv diffusibie hydrogen content in the weld, the steel producers do not recommend preheating before vvelding. There is no danger of cold cracking appearance in HAZ what is possible to check up vvith Duren and Suzuki theory for exampie /1, 2/ Transferring these statements into the real vveldment (interpass temperatures under 50 degrees Celsius and no preheating before vvelding) vvelded vvith low heat in-put (10-15 KJ/cm) to ensure good toughness in coarse grain area of HAZ, can in the melted part of the weld cause the appearance of low toughness values as the result of oblongated M/A struetural phase formation de-pending on the chemical composition of steel and vvelding material, in the čase of stress reiieved heating at 580 degrees Celsius there appears the undesirable effect to Fe3C precipitation from M/A phases representing a satu-rated solution. The toughness values can be lovvered even more if the sensitivity of the melted part of the vveld on a reversible temper brittieness is present. Because of the dear effects of iow heat input at vvelding and reduced grain size and primary ferrite on crystai boundaries, the vvelding material vvith 0,4% Cr and 0,2% Mo and Ti-B additions vvas chosen in these re-searches. The SA W vvelding method vvas used. The vveld toughness vvas determined by Charpy V-notch impact test. The notehess of Charpy impact specimens vvere lo-cated in face and root location of doubie V butt type vveld, right angled on the thickness of the plate. The rea-son vvas in the expected uneven dilution, first of aH from the Cb point of vievv (Cb= 0,01% in the face of the vveld and 0,04% in root of the vveld). Further tests vvere carri-ed out on the compiete vveld thickness by the CTOD method on the bend type specimens notehed and fati-gued in the rightangled direetion to the thickness in the middle of the vveld. Fractured surfaces vvere examined vvith line EM, first of aH on the appearance of LBZ. Both, line EM and opti-cal microscope vvere used by the examinations of mic-rostruetures in the vvelded joint. The presented resuits je LKP. Mikrostrukture zvara so bile preiskane z optičnim in vrstičnim elek. mikroskopom. Rezultati dokazujejo, da staljeni del zvara kljub drobnemu zrnu in nizki vsebnosti primarnega ferita po kristalnih mejah ni vedno dovolj žilav. Torej ocena struktur z optičnim mikroskopom po priporočilu IIW /3/ ni vedno zadostna za analizo in določitev vzrokov za dobro ali slabo žilavost staljenega zvara. Dodatne metode, kot npr. uporaba vrstičnega mikroskopa, lahko pokažejo dejanski vzrok za nizko ali visoko žilavost v staljenem zvaru v izhodnem in napetostno žarjenem stanju. 2. LASTNOSTI IN MIKROSTRUKTURA NIZKO OGLJIČNEGA EPP STALJENEGA ZVARA Poznan je odnos med mikrostrukturo zvara in žila-vostjo /4/. Razmešanje raztaljenega zvara v času varjenja spreminja transformacijsko kinetiko staljenega zvara, pri čemer imajo lahko vključki znaten vpliv. Gostota, velikost in porazdelitev vključkov narekujejo razvoj velikosti avstenitnih zrn po strditvi. Oksidni in drugi vključki ter koncentracija avstenitne trdne raztopine skupaj s hitrostjo ohlajevanja vplivajo na različne feritne morfologije v času premene (y/a /5a). Puščičasti ferit se npr. pojavi v EPP zvaru, če je koncentracija 02 višja od 450 ppm; njegova rast iz primarnega ferita na kristalni meji v zrno je intergranularna. Pri nižjih vsebnostih 02 se pojavi več ugodnega acikularnega ferita, medtem ko nizke vsebnosti 02 vodijo do tvorbe bainitne strukture. Poleg vsebnosti 02 je pomembna velikost in enakomerna porazdelitev vključkov. Vključki velikosti >0,2 (im bodo povzročili pospešeno tvorbo acikularnega ferita in znižali vsebnost primarnega ferita, pri čemer se povečajo primarna den-dritna zrna s fino porazdeljeno intergranularno strukturo /5b/. V času tvorbe primarnega in puščičastega ferita se preostali avstenit znatno obogati s C ter se lahko trans-formira v faze, ki vsebujejo zaostali avstenit in martenzit ali bainit-M/A strukturne faze. Takšno strukturo često najdemo pod izrazom ferit s sekundarno fazo. Prisotni vključki so tvorci puščičastega ferita, katerega oblika je odvisna od hitrosti ohlajevanja /6/. Višje hitrosti ohlajevanja pospešujejo tvorbo Widmanstattenskega ferita, ki ga spremlja neugodna porazdelitev M/A faz. Dodatki Nb ta fenomen še pospešujejo, kar se kaže v nižji žilavosti. Na drugi strani dodatki Ti in B izboljšajo žilavost, ker pospešujejo tvorbo acikularnega ferita, ki se pojavi na drobnih intergranularnih vključkih. Večina elementov vpliva na tvorbo in hitrost rasti puščičastega ferita. Dodatki Si in Al pospešujejo tvorbo puščičastega ferita, medtem ko jo Mn in Mo zavirata. Pretvorbo iz acikularnega ferita v ferit s sekundarno fazo, kot npr. M/A fazo, pospešujeta Cr in Mo, ki tudi povišujeta mejo plastičnosti in trdnost /7/. 3. RAZISKAVA SOČELNEGA EPP STALJENEGA ZVARA 3.1. Izbira osnovnega in dodajnega materiala Raziskave so bile opravljene na EPP zvarnem spoju nizkoogljičnega poboljšanega jekla Niomol 390. Lastnosti osnovnega in dodajnega materiala so navedene v tabeli 1. Navedeni dodajni material je bil izbran, da bi poudarili razlike pri nižjem in višjem dovodu toplote zaradi njegove povišane zakaljivosti, kar je razvidno iz višjega Pcm v primerjavi z osnovnim materialom. V predhodnih raziskavah /8/, kjer je bil uporabljen komercialni dodajni material (z 1 % Ni s Pcm = 0,149 in je bilo varjenje opravljeno prav tako brez predgrevanja, so se pojavila vidna LKP v prelomu CTOD preizkušanca v show, that the melted part of the vveid is not aiways tough enough in spite of fine grained coarse and iow pri-mary ferrite on crystai boudaries. Therefore, the estima-tion of structures vvith opticai microscope, as it is rec-ommended by IIW /3/, is not aiways sufficient enough for the anaiyse and determination of good or bad toughness in the melted part of the vvelded joint. Additional methods, as line eiectron microscope for example. can shovv the real reason for lovv or high toughness values in the melted part of the vvelded joint in as-welded or stress relieved conditions. 2. THE PROPERTIES AND MICROSTRUCTURE OF THE LOVV CARBON MELTED PART OF THE SUBMERGED ARC VVELDED JOINT The relationship betvveen microstructure and toughness of the vvelded joint is knovvn /4/. Dilution of the melted part of the vvelded joint during vvelding is chang-ing the transformation kinetic of melted vveldment and the inclusions can have a significant influence on it. Den-sity, size and distribution of inclusions do dictate the size development of austenitic grains after the solidifica-tion. Oxide and other types of inclusions and concentra-tion of austenitic solid solution together vvith the cooling speed. are influencing upon different ferrite morpholo-gies during the transformation y/a /5a/. Acicuiar ferrite can appear for example in the SA vvelded joint, if the concentration of oxygen is higher than 450ppm; its grovvth is intragranular from primary ferrite on crystal boundary into the grain. A t the lovver oxygen concentra-tions more favourable acicuiar ferrite appears, vvhiie lovv oxygen concentration lead to the formation of bainite structure. Beside the lovv oxygen concentration the size and uniform distribution of inclusions is important. The inclu-sion sizes over 0,2[im will cause an accelerated formation of acicuiar ferrite and reduce the content of primary ferrite increasing at the same time the size of the den-drite grains vvith fine distributed intragranular structure /5b/. During the formation of primary and acicuiar ferrite the remained content of austenite becomes significantly rich vvith carbon and can transform in phases containing residual austenite and martensite or bainite-M/A struc-tural phases. Such structure we can often find in the ex-pression ferrite vvith the second phase. Present inclusions are authors of acicuiar ferrite vvhich shape de-pends on cooling speed /6/. Higher cooling speed ac-ceierates the formation of VVidmanstatt-ferrite accom-panied by unfavourable distribution of M/A phases. The additions of Cb accelerates this phenomena vvhat results in the iovver toughness. On the other side the additions of Ti and B improve the toughness because they accelerate the formation of acicuiar ferrite, vvhich appears on intragranular inclusions. The most of elements have their effect on the formation and grovvth speed of acicuiar ferrite. The additions of Si and Al accelerate the formation of arrovv-shaped ferrite, vvhiie Mn and Mo re-tard it. The transformation from acicuiar ferrite to ferrite vvith secondary phase, as for example M/A phase. is accelerated vvith Cr and Mo vvhich also increase the yield point and strength /7/. 3. RESEARCH OF THE BUTT — SA VVELDED MOL TEN JOINT 3.1. The choice of the base and vvelding material The researches vvere carried out on the SA vvelded, lovv carbon, guenched and tempered steel plate Niomol izhodnem in napetostno žarjenem stanju celo pri + 20°C. Preiskave na rasterskem mikroskopu so odkrile več kot 10 % M/A strukturne faze. Dodatek Ti-B je v tej preiskavi bil izbran z namenom preprečiti oz. znižati tvorbo primarnega ferita po kristalnih mejah in pospešiti tvorbo acikularnega ferita v strukturi staljenega zvara /9/. Vpliv grobega zrna in primarnega ferita na žilavost staljenega zvara je bil tako z izbiro navedenega dodajnega materiala izključen. Namen preiskave je bil ugotoviti vpliv M/A faz z ozirom na njihovo sestavo (visok C), velikost, usmerjenost in gostoto na udarno in lomno žilavost v odvisnosti od hitrosti ohlajevanja in legiranosti staljenega zvara /10, 11/. Nadaljnji namen je bil ugotoviti vpliv napetostnega žarjenja na razpad M/A faz in tako na žilavost staljenega zvara. Tabela 1: Lastnosti uporabljenih materialov Lastnosti osnovnega materiala Debelina Re a5 Žilavost CTODi, -40° C /mm/ /MPa/ /MPa/ /%/ -60OC/J/ /mm/ 30 432 528 25 198,161,149 0,51 Kemična sestava 0.08C 0,30Si 1.11Mn 0.006S 0.015P 0.049Nb 0,0 8Sn 0,18Ni 0,012As 0.047AI Pcm = 0,1- CE = 0,27- 49 0 Lastnosti čistega dodajnega materiala EPP- R„ Rm a5 Žilavost CTOD dod.mat. /MPa/ /MPa/ /%/ -60°C/J/ /mm/ OP121TT 480 520-620 24 >50, _ Fluxocord varjeno 35.22 >35, — zarjeno Kemična sestava 0.05C 0,20Si 1,2Mn 0,04Cr 0,20Mo 0.005B Tidod. Pcm = 0,1 -CE = 0,37- 75 0 Kemična sestava dejanskega EPP zvara 0,05C 0,36Si 1,67Mn 0,70Cr 0,41Mo 0.025V 0,031Ti 0.023AI 0,04Sn 0,04Sb 0.006B 0,007As 0.009S 0.020P 0,022Nb Pcm = 0.2-CE = 0,55- 02 = 227p- HD<2,6 - 46 7 pm ml /100 gr 3.2. Varjenje preizkusnih plošč in priprava preizkušancev Uporabljeno je bilo večvarkovno varjenje na simetrično pripravljenem X zvarnem žlebu. Keramični vložek je bil uporabljen z izvedbo korenskega varka. Zvarjeni sta bili dve preizkusni plošči z različnima tehnologijama, A-brez predgrevanja in B-s predgrevanjem. Podatki o varjenju: debelina plošče — 30 mm predgrevanje — brez (A), 150°C (B) vnesena toplota — 15kJ/cm At8/5 — 5,4s(A), 10,5s(B) vmesna temp. — 50°C(A), 200°C(B) pogrevanje — 200° C 390. The properties of base and vveiding material are gi-ven in Table 1. Given vveiding material has been chosen for the reason to point out the differences at lower and higher heat input because of its higher quench capabi/i-ty. what is obvious from higher Pcm value in comparison vvith base material. In previous researches /8/, vvhere commercial vveiding material (vvith 1% Ni) vvas used, vvith Pcm= 0.149 and vvhere vveiding vvas also carried out vvithout preheating, visibie LBZ have appeared on fractured CTOD specimen surfaces in as-vvelded and stress relieved conditions even at + 20 degrees Celsius. Raster microscope ex-aminations have determined more than 10% M/A struc-tural phase. Ti-B addition in this research has been chosen for the reason to prevent or to reduce primary ferrite formation on crystrai boundaries and to acceier-ate the acicular ferrite formation in the structure of molt-en weld /9/. With such vveiding material selection the influence of coarse grain in primary ferrite on molten weid toughness vvas expelled. The aim of the examination vvas to find out the influence of M/A phases concerning their composition (high C(, size. orientation and density on fracture toughness of the molten vveld depending on its cooling speed and chemical composition /10.11/. Another aim vvas to find the influence of stress relieving on the disin-tegration of M/A phases in the molten vveld. Table 1: Properties of material used Base material properties Thickness Re Rm 85 Tough- CTODI. /mm/ /MPa/ /MPa/ /%/ ness -40° C -6CPC/J/ /mm/ 30 432 528 25 198,161,149 0.51 Chemical composition 0.08C 0,30Si 1,11Mn 0.006S 0,015P 0,049Nb 0.08Sn 0,18Ni 0,012As 0.047AI Pcm= 0,1- CE= 0,27- 49 0 Weid metal properties S A l/V Re Rm 5 5 Tough- CTOD ness I/Ve Id. Mat. /MPa/ /MPa/ /%/ -6CPC/J/ /mm/ OP121TT 480 520—620 24 >50, _ Fluxocord as welded — 35.22 >35, — stress re/. Chemical composition 0.05C 0,20S i 1,2Mn 0.04Cr 0.20Mo 0.005B Ti add. Pcm= 0,1-CE= 0.37- 75 0 Chemical composition of reai weid meta/ 0.05C 0.36Si 1,67Mn 0,70Cr 0.41 Mo 0.025V 0.031 Ti 0.023AI 0,04Sn 0,04Sb 0.006B 0,007As 0,009S 0,020P 0,022Nb Pcm= 0,2-CE= 0,55-O227p- HD< 2,6- 46 7 pm ml /100 g 3.2. Welding of test plates and specimen preparation A vveiding technique vvith more passes vvas used on double-V grooved steel plate. For the root vveld accom-piishment a ceramic insert vvas used. There vvere two Po varjenju je bila polovica vsake plošče napetostno odžarjena pri 580°C v času 3 ur, nekaj izrezanih preizkušancev pa še pri 650° C ter s kombinirano termično obdelavo. Preizkušanci so bili izrezani iz zvarnega spoja v izhodnem in napetostno odžarjenem stanju. Pripravljeni so bili metalografski preizkušanci, žilavostni preizkušanci zarezani v krovni in korenski legi ter CTOD preizkušanci iz celotne debeline; vsi zarezani pravokotno na površino plošče v sredini staljenega zvara. 3.3. Udarna in lomna žilavost staljenega zvara Udarni Charpijevi preizkušanci so bili pretežno odvzeti iz krovne lege zvara in preizkušani v izhodnem in različno termično obdelanih stanjih. Namen je bil ugotoviti vpliv M/A strukturnih faz na udarno žilavost. Nekaj preizkušancev je bilo odvzetih iz korenskega dela zvara z namenom ugotoviti vpliv Nb, ki se je izcejal iz osnovnega materiala v zvar ter ugotoviti, ali je prisotna termična reverzibilna krhkost. Rezultati za udarno in lomno žilavost staljenega zvara v izhodnem stanju, zvarjenega z in brez predgrevanja, so dani v tabeli 2. Tabela 2: Udarna žilavost in vrednosti za CTOD Stanje Lokacija Tempera- Žilavost CTOD Trdota zareze tura /°C/ /J/ /mm/ /HV5/ 1. 2. 3. 4. 5. 6. Brez krovni + 20 92 pred- sloj, -20 50 — 274 grevanja. korenski + 20 88 varjeno sloj celotna + 20 0,148-au debelina 0 0,094-ac —20 0,064-oc Brez krovni + 20 72 pred- sloj, -20 24 grevanja, korenski + 20 20 zarjeno sloj -20 8 580° C celotna + 20 0,034ac debelina Brez krovni -20 23 260 pred- sloj grevanja, zarjeno 650° C Predg. in krovni -20 65 257 varjeno sloj Predg. in krovni + 20 103 žar. 580° C sloj -20 29 Predg. in krovni + 20 114 236 žar 650° C sloj -20 42 Brez predgrevanja, varjeno Ref/8/, 1 % Ni krovni sloj —20 78 celotna debelina + 20 0 -20 0,420-ou 0,062-ac 0,091-ac Brez predgrevanja, žarjeno, 580° C, Ref. /8/, 1 % Ni krovni sloj -20 65 celotna debelina + 20 0 -20 0,323-au 0,103-ac 0,062-ac test plates prepared, vvelded vvith two different technolo-gies; A-without preheating and B-vvith preheating. VVeiding data: — 30 mm - vvithout (A), 15CPC (B) — 15 KJ/cm - 5,4 s (A), 10,5 s (B) — 50oC (A), 200° C (B) - 20CPC piate thickness preheating heat input A/s/5 interpass temp. heating after vveiding After vveiding each ha/f of the plate vvas stress re-lieved at 580° C in a period of 3 hours but some speci-mens vvere heated to 65CPC vvith combinated heat treatment. Test sampies vvere cut from the vveid in as vvelded and stress relieved condition . Test specimens vvere cut out for metal/ographic examination, toughness, notched in face and root area of the vveld, and CTOD specimens form the vvhole vveid thickness, ali notched perpendicu-iary on the plate surface vvith the notch location in the middle of the molten vveid. 3.3. Notch and fracture toughness of the molten vveld metal Table 2: Notch toughness and CTOD vaiues Condition Notch Tempera- Tough- CTOD Hardness location ture/0 C/ ness /mm/ /HV5/ /J/ 1. 2. 3. 4. 5. 6. VVithout face + 20 92 preheat.. vveld -20 50 — 274 as vvelded root + 20 88 vveld -20 16 vvhole + 20 0,148-8U thickness 0 0,094-8c -20 0,064+0c VVithout face + 20 72 preheat.. vveld -20 24 heated. root + 20 20 58CPC weld -20 8 vvhole + 20 0,034-?> c thickness VVithout face -20 23 260 preheat.. weid heated. 650° C Preheatedface -20 65 257 as vvelded vveld Preheatedface + 20 103 heated weid -20 29 580° C Preheatedface + 20 114 heated vveld —20 42 236 650° C VVithout face -20 78 preheat., weld as vvelded, vvhole + 20 0,420-6 u Ref. /8/, thickness 0 0,062-8 c 1 % Ni -20 0,091-8C VVithout face —.20 65 preheat., weid heated, 58CPC, vvhole + 20 0,323-8 u Ref. /8/. thickness 2 0,103-8 c 1% Ni —20 0,062-8 c Za primerjavo so dani rezultati iz reference /8/, kjer je bil uporabljen komercialni dodajni material (1% Ni). Varjeno je bilo z isto nizko vnešeno toploto in brez pred-grevanja kakor pri tej preiskavi. Tudi v tem primeru so se pojavila vidna LKP na preloma vrednosti, dobljene pri CTOD preizkusu ob pojavu "pop-in" efekta, so uporabljene različne oznake. Za nastop "pop-in" efekta po počasni rasti razpoke je oznaka (au), za nastop "pop-in" efekta takoj za otopitvijo razpoke pa (oc). Občutljivost na popuščno reverzibilno krhkost /13/ se je določala z VVatanabe faktorjem J /14/ in z udarno žilavostjo; rezultati so dani v tabeli 3. Tabela 3: Popuščna krhkost, določana z udarno žilavostjo Tehnologi- Brez predgrevanja Predgrevanje na 150°C ja varjenja Vmesna temp. 50° C Vmesna temp. 200° C Termična Lokacija T Žilavost Lokacija T Žilavost obdelava* zareze /° C/ /J/ zareze /°C/ /J/ 550° C krovna + 20 72 krovna + 20 103 lega, lega, krovna -20 23 krovna -20 30 lega, lega, korenska -20 10 — — — lega 550° C + krovna -20 55 krovna + 20 108 750° C lega lega 550° C + krovna -20 30 krovna + 20 86 710°C + lega lega 550° C * Pri vsaki temperaturi po 4 ure; VVatanabe faktor J = (Si + Mn) + (P + S) x 104; v našem primeru je J = 512 Za J= <200 je nizka občutljivost na reverzibilno popuščno krhkost Za J= >400 je visoka občutljivost na reverzibilno popuščno krhkost 3.4. Metalografske preiskave z optičnim in vrstičnim mikroskopom Preiskave so bile opravljene na metalografskih obru-sih, odvzetih iz plošč, varjenih brez in s predgrevanjem. Sliki 1 in 2 prikazujeta stebričaste dendrite z drobno in-tragranularno strukturo in nizko vsebnostjo ferita po kristalnih mejah v staljenem zvaru, zavarjenem brez predgrevanja. Drobna struktura, posneta z vrstičnim mikro- Slika 1 Brez predgrevanja. Fig. 1 VVithout preheating. Notch toughness specimens were mostiy cut out from face area and tested in as we/ded and different heat treated conditions. The aim was to find out the influence of M/A structural phases on notch toughness. Some specimens vvere cut out from root area vvith the aim to find out the influence of Nb, segregated from the base material to the vveld and also if the reversible tem-per embrittlement is present. The result of notch and fracture toughness of the molten vveld metal in as vvelded condition and vvelded vvith and vvithout preheating are shovvn in Table 2. For comparison, the results from the reference /8/ are given, vvhere the commercia/ vvelding material (1% Ni) vvas used. Welding vvas carried out vvith the same lovv heat input energy and vvithout preheating as in this ex-amination. In this čase visible LBZ on the fractured sur-faces of the specimen also appeared at + 20 degrees Celsius Different designations are used for the CTOD vaiues at "pop-in" effect appearance. At "pop-in" effect appearance, after the slovv crack grovvth, the designa-tion (t)J is used and designation (8J at "pop-in" effect appearance immediately after the crack tip is blunted. The sensitivity on reversible temper embrittlement /13/ vvas determined vvith VVatanabe factor J /14/ and vvith notch toughness; the results are shovvn in Table 3. Table 3: Temper embrittlement determined vvith notch-tough-ness VVelding VVithout preheating Preheating 15CPC technolo-- gy Interpass temp. 50° C Interpass temp. 200° C Heat treatment' Notch location T /°C/ Toughness /J/ Notch location T /° C/ Toughness /J/ 500° C face vveld root vveld + 20 -20 —20 72 23 10 face vveld + 20 -20 103 30 550° C 750° C face vveld -20 55 face vveld + 20 108 550° C+ 75CPC+ 55CPC face vveld -20 55 face weid + 20 108 'at each temperature 4 hours; VVatanabe factor J= (Si+ Mn)+ (P+ S)x 10*; in our čase is J= 512 For J— < 200 is sensibiiity on reversible temper embrittlement lovv For J= > 400 is sensibility on reversible temper embrittlement high 3.4. Meta/lographic examinations vvith optic and line microscope Metalographic specimens vvere cut out and ex-amined from steel plates, vvelded vvithout and vvith preheating. Fig. 1 and Fig. 2 shovv columnal dendrites vvith fine intragranular structure and lovv ferrite content on crystal boundaries in the molten vveld, vvelded vvithout preheating. Fine structure taken off with a line microscope in the face vveld area is shovvn in Fig. 3; oblong M/A phases along intragranular precipitated ferrite are visible. Fig. 4 shovvs the microstructure after stress re-lieved heating; severa! cementite precipitations on the boundary betvveen M/A phase and ferrite are perceived. Microstructure of the molten vveld, vvelded vvith preheating is shovvn in Fig. 5; there are less M/A phases and or- Slika 2 Ista struktura kot slika 1. Fig. 2 The same structure as in Fig. 1. Slika 3 SEM mikrostruktura zvara, brez predgrevanj. Fig. 3 SEM microstructure of the weid, vvithout preheating. Slika 4 Ista struktura - nap. žarjeno pri 580 °C. Fig. 4 The same structure - stress - reiieved heated at 580°C. Slika 5 SEM mikrostruktura zvara, s predgrevanjem. Fig. 5 SEM microstructure of the weid. vvith preheating. Slika 6 Ista struktura - nap. žarjena pri 580°C. Fig. 6 The same structure - stress - reiieved heated at 580°C. ientation is iess distinctive. The same microstructure after stress-reiieved heating is shovvn in Fig. 6: a strong tendency of cementi te coaguiation is visibie. The examination of LBZ areas has detected a quasi brittle fracture in the fractured area of the specimen in as-vvelded condition - Fig. 7. The fracture area of stress reiieved specimen shovvs besides the quasi brittle frac-tures also the presence of intergranular brittleness along the co/umnar dendrites as it is shovvn in Fig. 8. Slika 7 LKP - kvazi krhki transkristalni prelom. Fig. 7 LBZ - guasi brittle transcrystal fracture. skopom v krovni legi. je razvidna iz slike 3; vidne so podolgovate M/A faze vzdolž intragranularno izločenega ferita. Mikrostruktura po napetostnem žarjenju je razvidna iz slike 4; zaznavni so številni cementitni izločki na meji med M/A fazo in feritom. Mikrostruktura staljenega zvara, zavarjenega s predgrevanjem, je razvidna iz slike 5; M/A faz je manj in usmerjenost je manj izrazita. Ista mikrostruktura po napetostnem žarjenju je razvidna iz slike 6; vidna je tendenca močnega skepljanja cementi-ta. Pregled površin LKP je odkril kvazi krhki lom v prelomu preizkušanca v izhodnem stanju — slika 7. Površina preloma v napetostno žarjenem preizkušancu je poleg kvazi krhkega preloma pokazala še nastop intergranular-ne krhkosti vzdolž stebričastih dendritov, kot je razvidno iz slike 8. 4. DISKUSIJA REZULTATOV Preiskave so pokazale bistveno razliko v udarni Charpijevi žilavosti med osnovnim materialom in staljenim zvarom. Kljub dodatkom Ti-B in ugodni vrednosti 02 (ca. 227 ppm) v zvaru, ki pospešuje tvorbo acikularnega ferita in preprečuje tvorbo primarnega ferita, je bila dosežena udarna žilavost nižja od pričakovane. Lomna žilavost v prisotnosti ostre utrujenostne razpoke pa je pokazala popolno krhkost že pri temperaturah pod 0°C. Vzrok za to je varjenje brez predgrevanja in z nizko dovedeno toploto. Posledica tega je tvorba drobne mikrostrukture. ki vsebuje ferit s sekundarno fazo v obliki M/A strukturne faze namesto acikularnega ferita. Dodatki Cr, Mo in Nb težijo k pospeševanju tvorbe M/A faze, v kateri vsebnost C naraste tudi nad 1% /10/. Razpotegnjene M/A faze vzdolž intrgranularno izločenega ferita so lahko potencialni izvori zgodnjega začetka loma in zato nizke žilavosti. Dodatek Nb in njegovo izcejanje (0,04% v korenu in 0,01% v temenu) imata bistveni vpliv na udarno žilavost. Kemična sestava staljenega zvara je povišala VVata-nabe faktor J >400, tako da je posledica pojava popuš-čne reverzibilne krhkosti (SI. 8). Zaradi navedenega ni iz- Slika 8 LKP - intergranularni krhki prelom. Fig. 8 LBZ - intergranuiar brittte fracture. 4. RESULTS DISCUSSION The examinations have shovvn the essential differ-ences in Charpy toughness values of the base material and the molten vveld. In spite of Ti-B additions and fa-vourable oxygen content (ca. 227ppm) in the vveld, ac-celerating the acicular ferrite formation and prevention of primary ferrite formation, the notch-toughness values achieved vvere iovveras it vvas expected. But the fracture toughness in the presence of sharp fatigue crack has shovvn the absolute brittleness also at temperatures under 0 degree Celsius. The reason is vvelding vvithout preheating and vvith lovv heat input energy. The result is formation of the fine microstructure containing ferrite vvith secondary phase in the shape of M/A structural phase in state of acicular ferrite. The addition of Cr, Mo and Nb has a tendency of M/A phase formation in vvhich the carbon content can rich the values greater than 1 % /10/. Oblong M/A phases aiong the intragranutar precipitated ferrite can be the potential sources of early fractures be-ginning and the reason for lovv toughness values. Chemical composition of the molten vveld has increased the VVatanabe factor J= > 400, so that it resuits in the reversibie temper embrittlement appearance /8/. That is the reason why there is no toughness improvement at heating above 650 degrees Celsius vvhere the causes for temper embrittlement are dissolving, on the other side an intensive precipitating of Fe3C appears on the boundary betvveen oblong ferrite and M/A phase. The size of Fe3C precipitates are increasing vvith eievat-ed temperature under the coagulation mechanism vvhat is the reason for brittle fracture at elevated temperature /15/ (Fig. 4). From the above we can conclude, that stress-re-lieved heating does not have a favourable effect on molten vveld toughness if M/A structural phases are present into it. M/A phases formation is a consequence of too high alloyment or/and too high cooling speed, so that their presence depends on the vveld material selection and vvelding technology. Using the vvelding technology vvithout preheating the content of M/A structural phase boljšanja žilavosti po segrevanju nad 650° C, kjer se povzročitelji za popuščno krhkost raztapljajo, na drugi strani pa se hkrati pojavi intenzivno izločanje Fe3C na meji med razpotegnjenim feritom in M/A fazo. Velikost izločkov Fe3C se povečuje s povišanjem temp. po mehanizmu skepljanja, kar privede pri višji temp. do krhkega loma /15/ (SI. 4). Iz zgornjega je mogoče sklepati, da napetostno žar-jenje nima ugodnega učinka na izboljšanje žilavosti staljenega zvara, če so v njem prisotne M/A strukturne faze. Tvorba M/A faz je posledica previsoke nalegiranosti ali/in previsoke hitrosti ohlajevanja, tako da je njihova prisotnost odvisna od izbire dodajnega materiala in varilne tehnologije. Pri uporabi varilne tehnologije brez predgrevanja je znašala vsebnost prisotne M/A faze okrog 38% v staljenem zvaru. Z uporabo predgrevanja se je ta vsebnost znižala na okrog 33%. Iz raziskave je mogoče sklepati, da je pri varjenju brez predgrevanja in nizki dovedeni toploti za dosego visoke žilavosti v TVP potrebno uporabiti dodajne materiale, ki ne bodo tvorili večjih količin razpotegnjenih M/A faz. Navedeno lahko dosežemo z nižjimi vsebnostmi Si, Al, Cr in Mo in z dodatki Ti-B, ki pospešujejo tvorbo aciularnega ferita. Na drugi strani je podoben efekt mogoče doseči z uporabo predgrevanja in višjo dovedeno toploto ne glede na prisotnost omenjenih legirnih elementov. Vendar je takšna varilna tehnologija uporabna le, če izberemo jeklo, ki ni občutljivo na rast zrna, ker bomo v nasprotnem primeru dobili nizke vrednosti za žilavost v TVP. Omenjena moderna jekla so izdelana na osnovi TiN in BN izločevalnih efektov s vsebnostjo raztopljenega N pod 10 ppm. Pri varjenju preizkusnih plošč se vmesna temperatura visoko dvigne in lastnosti staljenega zvara niso primerljive z lastnostmi na dejanskem zvarnem spoju, ki je narejen brez predgrevanja ali z nizkimi At8'5 časi. Namesto da bi se varilo brez predgrevanja, dejansko varimo plošče z visokim predgrevanjem (često >200°C). Takšni rezultati niso uporabni za varjenje brez predgrevanja na dejanskih zvarnih spojih, kjer se zaradi dolgih zvarov in velikih površin vmesna temperatura zniža oz. pade celo na sobno. Takšen varilni postopek lahko znatno vpliva na žilavost izhodnih zvarnih spojev in tako na varnost celotne zavarjene konstrukcije. Torej je navedeno vzrok za pazljivo in temeljito izbiro verifikacije varilne tehnologije, ki mora odgovarjati dejanskim pogojem pri izvedbi varjenja v delavnici in na montaži. 5. ZAKLJUČEK Preiskava EPP zavarjenih raztaljenih zvarov, ki so za-varjeni brez predgrevanja in z nizko dovedeno toploto z uporabo več varkov na nizko ogljičnem finozrnatem jeklu, je odkrila naslednje: — Udarna in lomna žilavost raztaljenega zvara je odvisna od izbire dodajnega materiala in varilne tehnologije. — Pri varjenju in preizkušanju zavarjenih plošč morajo biti izpolnjeni pogoji, ki bodo reprezentirali dejanske pogoje v delavnici in na montaži. — Varilna tehnologija brez predgrevanja in z nizko dovedeno toploto, da bi dosegli dobre žilavostne lastnosti v TVP, lahko povzroči v raztaljenem zvaru v zrnih s primarnim feritom po kristalnih mejah in intragranularnim feritom s sekundarno fazo tvorbo krhkih M/A faz, ki lahko znatno znižajo žilavost. Odločitev, ali uporabiti pred-grevanje ali ne, je odvisna od ekvivalenta Pcm osnovnega in staljenega zvara in pričakovanih lastnosti. in moiten weid vvas 38 %. vvhiie vvith using the preheating this content vvas decreased to ca. 33 %. From this examination we can conciude that for vvelding vvithout preheating and lovv input energy to achieve high toughness in HAZ it is necessary to use vvelding materials vvhich vvill not form higher quantities of oblongated M/A phases. We can achieve that vvith iovver contents of Si, Al, Cr and Mo and vvith additions of Ti-B, vvhich accelerate the acicular ferrite formation. On the other side we can achieve a similar effect vvith the use of preheating and higher heat input energy regardless to above mentioned al/oying elements. But such vvelding technoiogy is useful only if we choose steel vvhich is not sensible to grain grovvth because in the opposite čase we will get iow toughness values in the HAZ. The above mentioned modem steels are produced on the base of TiN and BN precipitation effects vvith the nitrogen content under 10 ppm. During the vvelding of experimental plates, the interpass temperature raise high up and the properties of the moiten vveld are not comparable vvith the properties of the actual vveld vvelded vvithout preheating or vvith lovv At8/5 time. Instead of vvelding vvithout preheating we actually vveld them vvith high preheating (often > 200 degrees Celsius). Such results are not useful for vvelding vvithout preheating on actual vvelded joints because of their long-ness and greatness the interpass temperature de-creases or even tal/s dovvn to room temperature. Such vvelding process can have a considerable effect on as vvelded vveldments and in this way it effects on security of the vvhole vvelded construction. AH that has been stat-ed is the reason for a careful and profound choose of verification of vvelding technoiogy vvhich has to respond to actual vvelding conditions in the vvorkshop and as-sembling. 5. CONCLUSION The examination of SA vvelded moiten vvelds, vvelded vvithout preheating and vvith lovv heat input energy, vvith more run technique on lovv carbon fine grained steel has discovered the follovving: — Notch and facture toughness of the moiten vveld depends on vvelding material choose and vvelding tech-nology. — During vvelding and testing of vvelded plates such conditions has to be fulfilled that can represent the actual condition in the vvorkshop and assembling. —- Welding technology vvithout preheating and vvith lovv heat input energy can in the moiten vveld in the grains vvith primary ferrite vvith secondary phase, cause the formation of brittle M/A phases, vvhich can con side r-ably decrease the toughness. Decission to use or not to use the preheating depends on Pcm equivaient of base metal and moiten vveld and on properties expected. — By CTOD method determinated fracture toughness is othervvise conservative but gives a good insight into the quality estimation of the vvelded joint. — The LBZ appearance in the moiten vveld on the fractured area of CTOD specimens has to be already estimated at room temperature vvith iarge test (l/Vide Plate Tests) vvith fatigue crack inserted on the surface area. Test has to be carried out at the lovvest operating temperature of the future construction /16/. — Stress releaving vvith heating has negative conse-quences because of the M/A structural phases pres-ence. Cementite precipitation on the boundary betvveen ferrite and M/A phase additionally decrease toughness. In the čase of stress-reteaved heating it is recom- — Lomna žilavost, določana po metodi CTOD, je sicer konservativna, vendar daje dober vpogled v oceno kvalitete zvarnega spoja. — Pojav vidnih LKP v staljenem zvaru na lomni površini CTOD preizkušancev že pri sobni temp. je potrebno oceniti z velikimi preizkusi (Wide Plate Tests) s površinsko vgrajeno utrujenostno razpoko. Preizkus je potrebno opraviti pri najnižji temp. obratovanja bodoče konstrukcije /16/. — Termično sproščanje zaostalih napetosti ima negativne posledice zaradi prisotnosti M/A strukturnih faz. Izločanje cementita na meji med feritom in M/A fazo še dodatno zniža žilavost. V primeru uporabe napetostnega žarjenja je priporočljivo preveriti možnost pojava reverzi-bilne popuščne krhkosti, ki je odvisna od količine in izcej legirnih elementov ter nečistoč v raztaljenem zvaru. mended to control the possibility of reversible temper embrittlement appearance, vvhich depends on the quan-tity and segregations of alloying elements and impure-ments in the molten vveid. LITERATURA i REFERENCES 1. Suzuki, H.: Revised Cold Cracking Parameter PHA and Its Application, IIW IX-1311-84. 2. Durren, C.: Determining the Preheating Temperature for the Field-vvelding of Large-diameter Pipe, IIW IXG-318-84. 3. Guide to the Light Microscope Examination of Ferrite Steel Weld Metals, IIW IX-1533-88. 4. Cochran, R.C.: VVeld Metal Microstructure - a State of the Art Revievv, VVelding in the World, Vol. 21, No. 1/2, p. 16-25, 1983. 5a. Abson, D.J.: Non-metallic Inclusion in Ferritic Steel VVeld Metals. A Revievv IIW IX-1486-87. 5b. Liu, S., Olson, D.L.: The Role of Inclusions in Controlling HSLA Steel VVeld Microstructures, Supplement of VVelding Journal AWS, June 1986. 6. Easterling, K.: Introduction to the Physical Metallurgy of VVelding, Buttervvorths, 1983, p. 88-103. 7. Evans, G.M.: The Effect of Chromium on the Microstructure and Properties of C-Mn All-weld Metals Deposits, Oerlikon-Schweissmiteilungen 47/89, No. 120. 8. Rak, I., Gliha, V., Sidjanin, L., Petrovski, B.: J-integral Test-ing of SA VVeldments on HSLA Steel Regarding Different Treatments After VVelding, European Symposium on Elas- tic-Plastic Fracture Mechanics, Elements of Defect Asses-ment, Freiburg, 1989. 9. Okhita, S., Homma, H., Tsushima, Mori, N.: The Effect of Oxide Inclusions on the Microstructure of Ti-B Containing VVeld Metal, IIW IX-1070-86. 10. Ikavva, H., Oshige, H., Fukada, Y.: Effects of Martensite Austenite Constituents on HAZ Toughness of a High Strength Steel, IIW IX-1156-80. 11. Komizo, Y., Furusavva, J., Fukada, Y.: Development of High Toughness in Heat Affected Zone, JOM-3 Conference, Hel-singor, 1989. 12. Suzuki, H.: Root Cracking and Maximum Hardness in High Strength Steel VVelds, IIW IX-1983. 13. Dhooge, A., Ostyn, K., Magula, V., Vinckier, A.: Temper Embrittlement in Modem Off-shore Structural Steels. 14. Komizo, Y.: A Review on Reversible Temper Embrittlement in Cr-Mo Steel VVeld Metal, IIW IX-1488-87. 15. Knott, J.K.: Microscopic Aspects of Crack Extention, Adv-ances in Elasto-Plastic Fracture Mechanics, Applied Science Publishers London, 1979. 16. Denys, R.: Incentives for Fracture Toughness Testing, IIW X-1131-86. O O ^ o o do O K o E P k 64270 Jesenice, Cesta železarjev 8 telefon: (064) 81-231, 81-341, 81-441 telex. 34526 ZELJSN. Jugoslavija telefax 83 395 Kobaltove zlitine v lesni industriji Cobalt Base Alloys in Woodcutting Industrv J. Rodič*1 1. UVOD Več kot dve desetletji je v praksi dobro poznan ugoden učinek navarjanja rezilnega dela zob s kobaltovimi zlitinami-steliti* za izboljšanje rezilne sposobnosti in povečanje vzdržljivosti različnih vrst žag v lesni industriji (slikal) =<] a Slika 1 Stelitiranje Obloga konice zoba s posebno kobaltovo zlitino, odporno proti obrabi, omogoča bistveno podaljšanje vzdržljivosti in rezne sposobnosti žag. Stelitiranje je posebno priporočljivo pri žaganju svežega lesa z mnogoštevilnimi kremenčevimi vključki. Fig. 1 Stellite Tipping Tooth tipping vvith speciai cobalt base alloy resistant against abrasion enables significant proiongation of iifetime and cutting ability of savvs. Stellite tipping is especiaity recommended for savving fresh. nontreated wood vvith many inclusions of silicon oxides. Postopek prostega ročnega navarjanja zob na žagah se kljub poznanemu in dokazanemu zelo ugodnemu učinku dolga leta ni širše uveljavil. To ugotovitev lahko v veliki meri povezujemo s potrebo zahtevnega, zamudnega in zelo strokovnega dela, s slabim materialnim izkoristkom drage zlitine in z zahtevnostjo brušenja zob. V strokovnih krogih se je za postopek nanašanja ko-baltovih zlitin na zobe žag udomačilo ime "stelitiranje". Ta postopek je bil dolgo skoraj izključno prepuščen uporabnikom za vzdrževanje in obnavljanje žag v vsakodnevni praksi. Specializirani proizvajalci žag so se za nekatere vrste žag usmerili na obloge z lotanjem trdokovinskih ploščic na zobe. Z razvojem specializiranih polavtomatskih in * Stellite je prva blagovna znamka Cabot Corporation - Stoody Deloro za kobaltove zlitine, odporne proti obrabi in povišanim temperaturam. To ime je danes za veliko skupino kobaltovih zlitin splošno uporabljeno in udomačeno v praksi. Prof. dr Jože Rodič, dipl. inž. metalurgije, direktor podjetja MIL-PP d o.o . Ljubljana za razvoj in proizvodnjo specialnih zlitin 1. INTRODUCTION Beneficial effects of savv teeth tipping vvith cobalt base alloys - stellites' for improving cutting abiiity and iifetime of savvs in vvoodcutting industry are ivel/ knovvn more than two decades (Fig. 1). Although these beneficial effects vvere knovvn for many years the stellite tipping vvas not extensiveiy used due to the disadvantages of manual vvelding procedure vvhich is tirne consuming and requires a lot of experi-ence in vvelding and sharpening. Furthermore, this prac-tice is associated vvith lovv yield of expensive material. The stellite tipping vvas therefore applied only by users for maintenance and resharpening of savvs. Speciaiized producers on the other hand introduced hard metal tipping by soldering process. The nevv progress of stellite tipping arrised vvith the development of automatic machines for vvelding and sharpening vvhich vvere recently introduced by three companies ALLIGATOR (France), ISELI (Svvitzerland) and VOL L MER (Germany). Due to the increased productivity achieved by automatic stellite tipping a nevv netvvork of servicing centers for maintainance of aH types of band, gang and circular savvs is grovving. This professionai maintainance tech-noiogy assures better quality and life tirne of savvs vvhich has direct impact on productivity and economy of savv mili produetion vvhiie at the same tirne the quaiity of cut surfaces is improved. The stellite tipped savvs enable an uninterrupted eight-hour savving vvith cut length 80-100 km and produetion over 20 m3 per shift so that interruptions and changes of savvs during one shift are an exception. This promising progress has encouraged a syste-matic appiied research. With the grovving exploitation of stellites in vvood cutting technology a need for development of speciai assortment of aHoys devoted to this application is emerging. The need for intensive research in this area is also supported vvith results of comparative studies vvhich are presented in Section 7. These studies shovv that cobalt base alloys have a high priority for this application so that optimisation is expected vvithin these grade s. Three grades of stellites (12, 1, 6) are the most fre-quently used in vvoodcutting and the most important is grade 12 (See Tab. 1). The grade 1 vvas commonly appiied for hard vvood for many years, but recentiy the use of this grade has con-siderably decreased due to experience. The grade 6 has appeared in general vvoodcutting techno/ogy very recentiy. Some savv mills appiied this Stellite is a registered trade mark of Cabot Corporation -Stoody Deloro for cobalt base ailoys vvhich are abrasion resistant at room and elevated temperatures. This designation is commonly used for a vide variety of cobalt base alloys. polno avtomatiziranih strojev za stelitiranje žag, ki so jih razvile v zadnjem obdobju tri firme ALLIGATOR - Francija, ISELI - Švica in VOLLMER - Nemčija, se je začela situacija na področju stelitiranja žag bistveno spreminjati. Logična posledica visoke produktivnosti, dosežene z razvojem avtomatskega strojnega stelitiranja, je nastajanje vse širše mreže specializiranih servisnih centrov za stelitiranje in ostrenje vseh vrst žag, tračnih, gaterskih in krožnih. Vzdrževanje in obnova žag v takih servisnih centrih dosega vrhunsko in zanesljivo kakovost, kar se neposredno odraža v izrednem povečanju produktivnosti žaganja, v ekonomiki proizvodnje z zniževanjem stroškov, ob istočasnem napredku kakovosti rezanih površin in močnem podaljšanju življenske dobe kvalitetnih žag. Neprekinjeno osemurno žaganje hlodovine z dolžino poti rezanja 80—100 km in s storilnostjo razreza nad 20 m3 na izmeno, je danes za žago kar normalni normativ in zastoji zaradi nepričakovanih menjav žag so ob rednem vzdrževanju in kvalitetni obnovi žag že kar izjemen pojav. Razumljivo je, da je ta napredek vzpodbudil tudi intenzivnejše in bolj sistematične razvojne raziskave. Dosedanja uporaba standardnih vrst stelitov že napoveduje razvoj optimirane sestave teh zlitin za potrebe stelitiranja žag v lesni industriji in to v nekaj namenskih variantah glede na vrsto rezanega lesa in tehniko žaganja s specifičnimi parametri v proizvodnji. Tudi primerjalne raziskave stelitiranih žag in tistih z zobmi iz trdih kovin, ki jih bomo na kratko povzeli v poglavju 7, so privedle do presenetljivih spoznanj, ki močno uveljavljajo pomen stelitov v nadaljnjem razvoju. Kobaitove zlitine so se v primerjalnih raziskavah11 izkazale za najboljše, zato optimiranje asortimenta dodajnih materialov za stelitiranje žag pričakujemo med njimi. Dosedanji tipični asortiment stelitnih zlitin za lesno industrijo lahko omejimo na tri standardne vrste, med katerimi daleč prevladuje poznana zlitina št. 12. (Glej tabela 1) Za nekatere trde vrste lesa je bila močno uveljavljena in je še veliko uporabljana zlitina št. 1, vendar se prav v zadnjem obdobju na podlagi izkušenj v praksi vloga in pomen stelita 1 v uporabi na tem področju očitno zelo zmanjšuje. Ob preusmerjanju razvoja, ki izhaja iz bogatih izkušenj zadnjega obdobja izgleda očitno, da so lesarski strokovnjaki v splošnem preveč pomena pripisovali samo trdoti stelitiranih zob51. Nekateri "Žagarji", ki jim notranji raziskovalni nagon preprosto ni dopuščal, da bi se povsem prepuščali tradicionalnim pravilom pri izboru stelitov, so uspešno z zelo vzpodbudnimi rezultati zamenjali standardni stelit 12 s stelitom 6 in še naprej raziskujejo nove ideje z industrijskim preizkušanjem vzdržljivosti različno stelitiranih žag. Zelo zanimivi rezultati teh industrijskih raziskav v kombinaciji z metalurškim razmišljanjem o sestavah, lastnostih uporabljenih zlitin in z metalografskimi študijami že nakazujejo nove smeri razvoja z modifikacijami k optimalni sestavi kobaltovih zlitin za stelitiranje žag v dveh ali treh variantah glede na značilnosti vrste rezanega lesa. S pilotnimi napravami horizontalnega kontinuirnega litja7891 in z razvojem novih spremljajočih tehnoloških postopkov10' že poteka projekt specializacije v proizvodnji kobaltovih zlitin za specifične potrebe v lesni industriji. Razvojne raziskave potekajo v tesni povezavi s proizvajalci strojev za stelitiranje žag, s proizvajalci žag in s servisnimi centri za stelitiranje in ostrenje žag. V tem razvoju želimo zajeti in upoštevati čimveč praktičnih iz- grade vvhich vvas not commoniy used pureiy for research interest. The encouraging results confirmed opinion of some researchers5 that the hardness itseif should not be considered as the oniy decisive property of ste/lites for cutting abiiity. The research aiong these iines vvhich is out of traditionai practice is in progress. Hovvever. those savv miils vvhich tried grade 6 instead of grade 12 prefer the former one for standard use. These observations in vvoodcutting practice in combination vvith metallurgical knovvledge of chemical composition and material properties vvith respect to metai-iographic studies of microstructures should provide opt-imal assortment of stellites for cutting of various woods. The applied research project in the area described above is being undertaken in pilot plant for horizontal continuous casting (HCC)789 and subsequent ther-momechanical treatment10. This research is performed in cooperation vvith producers of stellite tipping machines. producers of saws and vvith service centers for stellite tipping and sharpening. It is expected that this joint research programme vvhich considers expertises gained in industry reiated to vvoodcutting vvill result in metallurgical development of products for stellite tipping. A Slika 2 Trakovi žag, zvarjeni in napeti. A - surovo ozobljeni trak, B - Razperjanje in brušenje zob. C - Nakrčenje vrhov zob in brušenje, D - Stelitiranje in brušenje zob. Fig. 2 Savv bands, weided and tensioned A - Saw band vvith raw teeth. B - Setting and grinding of teeth. C - Swagging of tooth tips and grinding. D - Stellite tipping and grinding of teeth. kušenj, zbranih v zadnjih letih in le-te usmeriti v metalurški razvoj zlitine za specifične namene in potrebe. Poleg razvoja asortimenta za optimalni izbor vrste uporabljenih zlitin je pomemben tudi razvoj samih postopkov stelitiranja in posebnih oblik preseka HKL-palic, dodajnih materialov za stelitiranje zob. Ti proizvodi so poznani pod imenom FORM-STELITI in so se doslej izdelovali samo po tehnologiji metalurgije prahov (PM). Danes se kot pomembna dopolnitev asortimenta že uveljavlja tudi uporaba HCC-FORM-STELITOV. 2. VRSTE ŽAG, OBLIKE IN GEOMETRIJA ZOB Pri stelitiranju enakovredno obravnavamo vse tri osnovne vrste žag, tračne, gaterske in krožne. Na sliki 2 je shematično prikazano surovo ozobljenje trakov, razperjanje ali nakrčevanje zob pri klasičnih in stelitiranje zob pri modernih tračnih žagah. Slika 3 prikazuje geometrijo stelitiranega zoba z vsemi koti, ki so pomembni za ostrenje z ravnim in poševnim brušenjem. Slika 3 Geometrija zoba žage X - Cepilni kot, V - Prosti kot. Fig. 3 Geometry of savv tooth X - Rake angie. V - Ciearance angie. 3. RAZVOJ STELITIRANJA ŽAG Pri tem moramo omeniti dva bistevno različna pristopa k stelitiranju žag. Od tega je odvisna seveda tudi konstrukcija in način delovanja strojev za stelitiranje. Postopek I.: Navarjanje zob s TIG postopkom ali s plazmo Pri postopku stelitiranja z dodajanjem kobaltove zlitine na zob z navarjanjem preko tekoče faze uporabljamo okrogle HCC palice tankih presekov. Daleč največ se uporablja palice standardne dimenzije 003,2 mm. Po-dajalni mehanizem stroja podaja palico za taljenje nad kokilo. Pri tem postopku avtomat stroja obda zob žage z bakreno kokilo (slika 4), ki ima določeno obliko zoba in stelit se po TIG postopku ali s plazmo natali v kokilo na vrh zoba. Po strditvi je vrh zoba s stelitom ustrezno oblikovan in stelit trdno zvarjen z osnovo zoba. Uvedba plazma gorilcev pri teh strojih omogoča povečanje storilnosti stelitiranja in tudi uporabo nekoliko Besides the development of optimal assortment of appiied a/loys the development of vveiding procedures in ste/lite tipping and appiication of special shapes ofcross section for adding the material in vvelding is a/so important. These so ca/led FORMSTELLITES are mainiy produced by povvder technology but recently HCC is con-sidered as an alternative technology. 2. SAVV TYPES, FORMS AND TOOTH GEOMETR V Stellite tipping is equally treated vvith aH three general types of savvs: band savvs. gang savvs and circular savvs. Figure 2 represents schematically a saw band vvith raw teeth. setting and svvagging of teeth vvith conven-tional savvs and stellite tipping of teeth. The geometry of stellite tipped tooth vvith aH important angels for sharp-ening vvith straight and bevei grinding is shovvn in Fig. 3 3. DEVELOPMENT OF STELLITE TIPPING Tvvo essentially different procedures have been introduced in the approach to stellite tipping of saw teeth. The construction and operational characteristics of stellite tipping machines are adjusted to special require-ments of the process. Procedure /.; Stellite tipping vvith TIG or plasma vvelding For stellite tipping vvith vvelding through liquid phase HCC rods of small sections are usually used vvhere 0O3.2 mm is the most commonly appiied dimension. The stellite tipping machine automatically supplies the rod to the appropriate position above the mould for melting. Slika 41> Skica naprave, ki kot kalup daje zahtevano obliko konice zoba pri stelitiranju. Fig. 41> Schematic dravving of a j/g vvhich performs the mouiding action to g/ve the tip the desired geometry. In this procedure the b/ocks of cooper are moved to embrace the saw tooth in order to form a mould for the desired geometry. The mould is then filled up by either TIG or plasma melting. After the solidification a preform of the tooth vvhich is vvelded to the base is obtained (Fig. 4). The introduction of plasma vvelding enabied the improvement of productivity and the application of stellite rods of larger section vvhich are considerably cheaper so that plasma vvelding is reducing the overall production cost of stellite tipping. Figure 5 illustrates stellite tipping and sharpening of saw teeth. Procedure II.: Resistance vvelding in stellite tipping In resistance vvelding approach the adding material in the form of stellite tip is heated up in automatic ma-chine to melt both materials at the interface betvveen tip and saw base. The machine is then pressing the tip to the exact position on the tooth. For tips used in this approach precision čast stellite pieces or pieces of pressed povvder stellites vvere introduced by analogy vvith hard metal tipping. debelejših palic, kar se pomembno pozna pri ceni dodaj-nih materialov oziroma pri stroških stelitiranja v celoti. Slika 5 prikazuje stelitiran zob in brušenje31. Slika 6 Električno uporovno navarjanje stelitnih delcev na vrh zoba A - Vertikalno podajanje okroglih stelitnih palic, B - Vertikalno podajanje stelitnih palic trapeznega preseka, C - Horizontalno podajanje okroglih stelitnih palic. D - Horizontalno podajanje stelitnih palic paralelo gramskega preseka. Fig. 6 Electric resistance vvelding of stellite tips A - Verticai adding of round stellite rods. B - Verticai adding of form-steliite rods vvith trapeze section. C - Horizontal adding of round stellite rods. D - Horizontal adding of form-steliite rods vvith paralielogram section. Slika 53' Stelitiranje in brušenje zob A - Oblika stelitiranega zoba, B - Brušenje cepilnega kota, C - Stranske ploskve se brusijo samo na novo stelitiranem zobu, D - Večkratno prebrušenje cepilnega in prostega kota. Fig. 53> Stellite tipping and sharpening of teeth A - Form of stellite tipped tooth. B - Grinding of rake angle. C - Side ciearance surfaces are ground only after steliite tip-ping. D - Repeated sharpening of rake and ciearance angle. Postopek II.: Uporovno navarjanje stelitnih delcev na zobe Po drugem osnovnem postopku stelitiranja žag z več variantami se dodajni material-stelit določene oblike v avtomatu uporovno segreje, na stičišču z osnovnim materialom žage natali in vtisne točno v ustrezen položaj na zobu žage. Za ta postopek privarjanja so že ponudili tržišču tudi precizne ulitke stelitov, ali pa oblikovance iz stelitnega prahu, vsestransko oblikovane. S tem naj bi posnemali izkušnje iz uporabe trdokovinskih ploščic, ki se lotajo na zobe žag. Uporaba teh predoblikovanih koščkov se po začetnih korakih razvoja v stelitiranju žag ni veljavila po pričakovanjih in danes že prevladuje mnenje, da ta prvotno zelo obetajoča pot v nadaljnjem razvoju stelitiranja žag nima perspektive. Več uporabljajo palice z različnimi oblikami preseka, ki jih avtomat reže ravno ali poševno med postopkom stelitiranja pri podajanju palice. Avtomat lahko delček prej odreže, pa ga nato privari ali pa konec palice privari in jo nato avtomatsko odreže. Razvoj dodajnih stelitov v obliki palic različnih dimenzij s posebnimi oblikami preseka je odvisen od sistema dodajanja palic za stelitiranje zob pri avtomatskem ali ročnem podajanju na stroju. Najprej se je razvil postopek z uporabo palic okroglega preseka, ki jih avtomatsko ali ročno krmiljeni stroj za stelitiranje podaja z vrha ali od strani pred cepilno ploskev zoba kot kaže slika 6. * 1. varianta: Vertikalno podajanje palic Pri tem načinu stelitiranja žag s podajanjem palic okroglega preseka od vrha pred zob žage je povsem razumljivo v fazi brušenja prišlo do ideje za uporabo palic kvadratnih ali pravokotnih in končno trapeznih presekov, tako da ima zob že takoj po stelitiranju nastavljeno ravno cepilno ploskev in predoblikovani stranski ploskvi s prostim kotom. Palico se odreže poševno pod kotom, ki ustreza prostemu kotu zoba. Na ta način se doseže velik prihranek brušenja zob. * 2. varianta: Horizontalno podajanje palic Za drugo varianto postopka stelitiranja z uporovnim navarjanjem se uporablja dodajanje palic stelitov od strani. Namesto okroglih palic večjega preseka (slika 73') imajo precejšnjo prednost palice s presekom paralelograma (slika 83)). Nekaj nevšečnosti pri postopku uporovnega navarja-nja form stelitov povzročajo ostanki iztisnjenega materiala. Odstranjevanje teh brad je lahko neprijetna ovira normalnega postopka. Po stelitiranju se morajo konice zob popuščati, kar se z modernimi stroji opravi v toku samega postopka stelitiranja. Izkušnje kažejo, da med vsemi form-steliti po količini daleč prevladuje uporaba tistih s trapeznim presekom. Uporaba form-stelitov je kljub višji ceni upravičena, ker omogoča Slika 73> Uporovno navarjanje zob z okroglim stelitom A - po stelitiranju, B - po končnem brušenju. Fig. 7a> Stell/te tipping with electric resistance vvelding of round stellite rod A - after stellite tipping, B - after final sharpening. B Slika 83' Uporovno navarjen zob s form-stelitom paralelogramskega preseka A - po stelitiranju; B - po končnem brušenju. Fig. 8?> Stellite tipping vvith electric resistance vvelding of form-stellite rod vvith para/lelogram section A - after stellite tipping, B - after final sharpening. The application of these preformed tips vvas expect-ed to be successful but very recently more economic procedures are proposed in vvhich rods of various dimensions and cross-sections are automatically supplied and cut at appropriate angles during the process. There are tvvo alternatives: The piece of rod for tooth tip is cut off first and then vvelded or the end of rod is vvelded first and then cut off. The development of stellite rods in various sections and forms depends on adding system of the stellite tipping machine. At the beggining round section rods vvere supplied either from the top or side to the face of the tooth as shovvn in Fig. 6. * Alternative 1: Verticai suppiy of rods In this alternative of stellite tipping the round section rod is supplied from the top and cut at clearance angle. To reduce the vvaste of material and grinding costs square and rectanguiar sections vvere introduced but fi-nally trapeze form sections vvere accepted as standard. After stellite tipping vvith trapeze-form rods ali main angles of the tooth are close to the final sharpening geometry. Side clearance surfaces are ground only after stellite tipping vvhile repeated sharpening is performed only for rake and clearance angles. * Alternative 2: Horizontal suppiy of rods In this alternative of stellite tipping vvith resistance vvelding rods of larger sections are supplied from the side and cut at side clearance angle. Initially round rods (Fig. 7) vvere used but later parallelogram sections vvere introduced to reduce the vvaste of material and grinding costs (Fig. 8). The difficulties can arrise in resistance vvelding vvhere the material vvhich is pressed to the side needs to be removed. After stellite tipping the teeth must be tempered. Modem machines include this in automatic procedure. Production practice shovvs that number of various forms is reducing tovvards a limited number of standard forms and that trapeze section forms are preferred. Although the stellite forms are expensive their application is justified by subsequent costs such as reduced consumption of stellites, up to 60 % shorter sharpening times and iovver consumption of abrasive tools. 4. PRODUCTION TECHNOLOGY FOR ADDITIVE MA TE- RIALS IN STELLITE TIPPING Stellite rods can be produced by tvvo different technologie s: — Horizontal Continuous Casting (HCC), (Fig. 9); — Povvder Metallurgy (PM). Almost a/l round section rods are produced by HCC vvhile form stellites on the other hand vvere until now produced only by PM. The first samples of form stellites produced in pilot plant MILPP represent a new technology and are novv tested by users. The initial resuits shovv that this tech-nology vvill be introduced in practice as complementary rather then competitive to the existing PM technology. The PM form stellites can not be replaced by HCC for special applications and chemical compositions but for common stellite tipping at iovver priče level and larger quantities an additional market can be opened for the HCC form stellites. — manjšo porabo stelitov — do 60 % krajši čas brušenja, ker sta prosti in cepilni kot zoba že podana in — manjšo porabo brusnih plošč. 4. TEHNOLOGIJA IZDELAVE DODAJNIH MATERIALOV ZA STELITIRANJE ZOB Za proizvodnjo stelitnih palic sta v uporabi dve osnovni tehnologiji: — horizontalno kontinuirno litje (HCC), — metalurgija prahov (PM). Skoraj vse okrogle palice za stelitiranje so izdelane po HCC tehnologiji (slika 9). Za proizvodnjo form-stelitov v glavnem prevladuje metalurgija prahov. Prvi vzorci HKL-formstelitov iz pilotne proizvodnje MIL-PP predstavljajo novost in so že na preizkušanju v uporabi. Po prvih izkušnjah pričakujemo, da se bo uporaba HKL-formstelitov v praksi uveljavila bolj kot pomembna dopolnitev in ne toliko kot konkurenčna akter-nativa dosedanjega asortimetna formstelitov, izdelanih po tehnologiji metalurgije prahov. Cenejši HKL-formsteli-ti bodo omogočili širšo uporabo te tehnologije stelitira-nja, PM-formsteliti pa bodo še naprej nepogrešljivi za specialna področja uporabe. 5. ASORTIMENT ZLITIN ZA STELITIRANJE ŽAG Ker gre pri stelitiranju za tipično interdisciplinarno področje med metalurgijo, lesarstvom, strojništvom in kemijo (korozijo) je prav, če na kratko predstavimo sicer poznane vrste in specifične lastnosti kobaltovih zlitin pod skupnim imenom "steliti". Tudi nekaj primerov mi-krostruktur stelitov iz HKL in PM tehnologije zanimivo prikazuje značilnosti teh posebnih zlitin in obeh postopkov. Skupina stelitov obsega zlitine s precej širokim območjem variacij kemijske sestave. Skupna značilnost vseh stelitov je osnovna sestavina, kobalt. Dodatki drugih kovin v različnih kombinacijah vplivajo na značilne mehanske lastnosti, in s tem tudi neposredno na obstojnost proti obrabi ter na odpornost proti koroziji, ki je pri rezanju svežega lesa izredno pomembna. Žilavost stelitov, čeprav slaba, je mnogo boljša od izredno krhkih WC-trdin. Lastnosti stelitov so dobro obstojne tudi pri povišanih temperaturah. Steliti se odlikujejo z nizkimi koeficienti trenja. Slika 9 Shema postopka horizontalnega kontinuirnega litja (HKL) tankih palic. Fig. 9 Scheme of the proces for horizontal continuous casting (HCC) of thin rods. 5. ASSOR TMENT O F ALLOYS FOR STELLITE TIPPING Since development of stellite tipping requires inter-disciplinary cooperation betvveen metallurgists. mechan-ical and chemical (corrosion) engineers and vvoodcut-ting experts it is appropriate to recall vvell knovvn properties. chemical compositions (Tab. 1) and hardness of typical grades for this appiication. Some typical micros-tructures of steilites produced by HCC and PM technol-ogies are also presented for comparison. Ali steilites are cobait based. They form a group ofal-ioys vvith a vvide variety of chemical compositions vvhich determine their mechanical properties. cutting ability. abrasive resistance and corrosion vvhich is recently con-sidered as decisive especialiy for cutting fresh vvood. Although the toughness of steilites is iow it is aiways much higher then toughness of tungsten carbide grades of hard metals. Many properties of steilites remain a/most un-changed even at elevated temperatures. The friction caused by steilites is much tovver then friction produced by tungsten carbide grades of hard metals. 6. METALLO GRA P H Y OF HCC- AND PM- STELLITES11 At MIL-PP a modification of standard grade 12 vvas introduced vvith designation MILIT 12 W. A metailogra- Tabela 1: Kemijska sestava in trdota preizkušanih zlitin Table 1: Chemical composition and hardness of te ste d alloys Co % Ni % Cr % W % Si % B % Fe % C% HRC I. Štiri zlitine primerjane v raziskovalnem projektu FORINTEK CANADA CORP.11 I. Four ailoys compared in research project by FORINTEK CANADA CORP.1' STELLITE 12 59 STELLITE 20 45 DELORO 50 - 77 DELORO 60 - 70 II. Druge zlitine, uporabljane za rezanje lesa II. Other alloys. applied in vvood cutting STELLITE 1 STELLITE 6 (STELLITE F) 54 65 39 29 33 10 15 30 28 25 9 18 12 4 12 4 4.5 1.5 3 4 4.5 1.8 2.5 0.4 0.5 2.5 1.1 1.7 47-51 55-59 49-52 59-62 51-58 39-43 40-45 6. METALOGRAFIJA HKL- IN PM-STELITOV11' MIL-PP je za stelitiranje žag razvil specialno HKL varianto stelita pod imenom MILIT 12 W s posebno kemijsko sestavo. Za ilustracijo prikazujemo nekaj metalo-grafskih .posnetkov HKL-palice 003,2 mm v surovem litem stanju. Sliki 10 A in B prikazujeta značilno površino HKL palice, posnete z rasterskim elektronskim mikroskopom (REM). Vidi se sled koraka, katere globina znaša poprečno 0,1 mm. Slika 10 Površina HKL-palice 003,2 mm - Kobaltove zlitine MILIT 12 W (A - REM 20x, B - REM 100x). Fig. 10 Surface of HCC-rod 0O3.2 mm - Co base alloy grade MILIT 12W (A - SEM 20x, B - SEM 100x). Slika 11 Značilna strjevalna mikrostruktura HKL palice 003,2 mm -MILIT 12 W - v litem stanju (A-prečno 200x; B-vzdolžno na sredini preseka 200x). Fig. 11 Characteristic soiidification microstructure of HCC-rod 0O3.2 mm - MILIT 12 W - as čast (A - transverse 200x. B - iongitudinai in the middle of section 200x). Naslednji dve sliki 11 A in B kažeta mikrostrukturo istega vzorca na prečnem in vzdolžnem preseku. Značilna strjevalna mikrostruktura dendritnega tipa s primarnimi in sekundarnimi vejami je izredno fina, kar je posebna prednost HKL-tehnologije. Nov specialni postopek termomehanske konsolidacije HCC palic101, ki ga razvija MIL-PP v okviru posebnega projekta specializacije proizvodnega programa kobalto-vih zlitin za stelitiranje žag v lesni industriji, obeta še dodatne kakovostne prednosti. Za razliko od tipične mikrostrukture HKL vzorcev vidimo na naslednjih slikah 12 A, B in C mikrostrukturo PM-vzorca paralelogramskega preseka, ki je bil izdelan po tehnologiji metalurgije prahov. Očitna je značilna porazdelitev karbidov v matrixu. Slika 13 prikazuje značilnosti porazdelitve elementov med osnovo in karbidi PM form-stelita. Slika 12 Značilna mikrostruktura paralelogramskega formstelita izdelanega po tehnologiji metalurgije prahov (A - pov. 200x, B - pov. 500x, C - REM 2000x, D - REM 5000x, E - pov. 200x ob ostrem kotu paralelograma). Fig. 12 Characteristic microstructure of parailelogram formstellite pro- duced with povvder metailurgy (A - magn. 200x, B - magn. 500x. C - SEM 2000x. D - SEM 5000x. E - magn. 200x at the sharp angle of parailelogram). phic iilustration of HCC ste/lite 0O3.2 mm in as čast condition made by scanning eiectron and optical micro-scope is presented in figures 10 and 11, respectively. The depth of ivitness mark on the surface of HCC rods is usually 0.1 mm. The solidified microstructure of dendritic type vvith primary and secondary branches is very fine vvhich is characteristic for HCC technoiogy. A new thermomechanical consolidation of HCC rods is being deveioped by MIL-PP and is expected to enable improved quaiity of stellites for tipping of savvs. A typical microstructure of PM form ste/lite vvith parailelogram cross-section is presented in Fig. 12 vvith optical and SEM photographs vvhere characteristic distribution of carbides in the matrix is evident. in Fig. 13 presence of chemical elements in the ma-trix and car- bides of PM form stellite is presented. 7. COMPARA T/VE RESEARCH1 2 The research and development center FORINTEK CANADA CORP. published results of a comprehensive research vvhich had a decisive infiuence on further development of ste Hite tipping for vvoodcutting. ES I ETo Co Cr i Fe Ni Si i W Mn Slika 13 Elektronska slika in prisotnost elementov v osnovi in karbidih PM form-stelita. Fig. 13 Eiectron picture and distribution of eiements in the matrix and carbides of PM form-stellite. 7. PRIMERJALNE RAZISKAVE12' V laboratorijih raziskovalno-razvojnega centra FO-RINTEK CANADA CORP. so opravili obsežne raziskave", ki so odločilno vplivale na nadaljnji razvoj stelitiranja žag v lesni industriji. V prvi seriji poskusov so primerjali štiri zlitine odporne proti obrabi pri žaganju svežega lesa. (Tabela 1) Na sliki 14 je prikazan način merjenja otopitve rezalnega roba s fotomikroskopijo". Glede na otopitev rezalnega roba je dal najboljše rezultate nanos zlitine stellite 12 (slika 1511). Otopitev zob po 35 km reza, kar ustreza približno 4 uram žaganja, je bila pri jeklenih zobeh kar dvanajstkrat večja kot pri stelitiranih z zlitino stellite 12. Še pomembnejša je ugotovitev, da so jekleni zobje dosegli že po 1 km reza ali 6 minutah tako stopnjo otopitve kot steliti-rani zobje po 35 km reza ali 4 urah žaganja. Krivulja obrabe pri jeklu za žage kaže, da so standardni zobje žage že po eni uri rezanja močno obrabljeni. Zaradi zmanjševanja zastojev je žal kar običajno, da se s tako otopelimi zobmi žaganje nadaljuje. Posledica tega je slaba kvaliteta površine rezanega lesa, neenakomerna debelina in netočnost reza ter seveda močno povečana poraba energije. Obe zlitini na osnovi niklja DELORO 50 in 60 po ugotovljenih rezultatih ne moreta konkurirati zlitini stellite 12. Zlitine stellite 20 že zaradi slabih rezultatov uporabe na žagi z meritvami v laboratoriju sploh niso preizkušali. Meritev otopitve na zobu žage. Fig. 141> Measurement of cutting edge retreat on savv tooth. DOLŽINA REZA [km] CUTTING PATH [km] Slika 151' Otopitev konvencionalnih in obloženih zob na žagah. Fig. 15» Duiiing of conventional and tipped savv teeth. In the first set of experiments four grades of abra-sion resistant alloys for cutting fresh wood vvere tested (Tab. 1). in Fig. 14 a method for measuring duiiing of cutting edge vvith photomicroscopy is presented1. The best results vvith respect to duiiing of tooth cutting edge vvere obtained by tipping of stellite grade 12 (Fig. 15) The duiiing of teeth after cut iength of 35 km. vvhich corresponds to approximateiy 4 hour s of savving. vvas measured for savv steel and stellite 12. The value of cutting edge retreat for savv steel vvas found to be tvvelve times higher than for stellite 12. Even more important is the observation that the measured duiiing of stellite 12 OSTRO SHARP OTOPITEV JEKLENEGA ZOBA DULLING OF STEEL TEETH TOPO DULL OSTRO SHARP OTOPITEV TRDOKOVINSKEGA ZOBA DULLING OF CARBIDE SAW TEETH Slika 161' Otopitev jeklenih in trdokovinskih zobVVC - volframov karbid, Co - osnova, V - prazen prostor, ki je bil prej zapolnjen s kobaltom. Fig. 161> Dulling of savv steel teeth and carbide tips WC - tungsten carbide, Co - matrix, V - empty space, formeriy occupied by cobait. Zanimiva je ugotovitev, da je obstojnost žage v veliki meri odvisna tudi od sposobnosti zlitine na zobeh za kvalitetno brušenje. Zaključek te serije raziskav je bilo priporočilo uporabe zlitine stellite 12 za stelitiranje vseh vrst žag, ne samo zato, ker je ta zlitina pokazala najboljše rezne sposobnosti, najboljšo obrabno in korozijsko obstojnost, ampak tudi zaradi najboljšega obnašanja pri brušenju. Pri brušenju te zlitine je dosežena najboljša začetna ostrina zob, kar pomembno vpliva na celotno izdržljivost žage. V drugi seriji posebnih laboratorijskih poskusov z natančnimi meritvami so primerjali obrabo rezalnega roba zlitine stellite 12 z dvema vrstama karbidnih trdin in s standardnim jeklom za žage. Raziskave so opravili za tri tipične vrste lesa. Pri teh poizkusih je bila dolžina reza 80 km, kar ustreza normalno žaganju 7-8 obratovalnih ur. Slika 1611 shematično prikazuje značilnosti otopitve na rezalnem robu jeklenega oziroma trdokovinskega zoba. Obraba oziroma otopitev stelitiranih zob je po mehanizmu podobna jeklenim zobem. Stelitirani zobje so pokazali v primerjavi z obema kar-bidnima trdinama (volframov karbid - 6 % Co in volframov karbid - 18%Co) najmanjšo obrabo in bistveno boljšo vzdržljivost. To pomeni, da imajo stelitirane žage precejšnje prednosti pred trdokovinskimi, posebno pri rezanju svežega lesa in pri tankih rezih. Ta raziskovalna ugotovitev je bila potrjena tudi v industrijski praksi žaganja zadnjih let. Korozijska odpornost zlitine in spoja z osnovo v veliki meri ugotovitev dodatno pojasnjuje. Kislinski ekstrakti svežega lesa napadajo kobalt v osnovni masi in s tem poslabšajo odpornost proti izpadanju vol-framovih karbidov. Stelitirane žage bodo pri rezanju svežega lesa popolnoma izpodrinile žage s trdokovinskimi zobmi. Trdokovinske žage se bodo še naprej obdržale v pohištveni in drugi finalizacijski lesni industriji. Tudi tam ima zaradi problemov odpadanja trdokovinskih ploščic pri velikih hitrostih ter ob udarnih obremenitvah na zob vez med oblogo zoba in osnovno kovino, ki jo dosežemo pri stelitiranju določene prednosti v primerjavi z nalotano trdo kovino. 8. POVZETEK PREDNOSTI STELITIRANJA ŽAG (a) v primerjavi z uporabo standardnih jeklenih žag: — Rezna zmogljivost stelitiranih žag je v primerjavi z žagami z razperjenimi zobmi večja, ker reže vsak zob na obeh straneh. — Hitrost pomika se lahko poveča do 30 %. - Povečana obstojnost rezalnega roba zob in zmanjšanje zastojev at 35 km is reached by saw steel already after first kilometre of cut length vvhich corresponds to approxi-mately six minutes of savving. Thus the dulling of savv steel teeth is very intensive in initial stages so that savving after fevv kilometres of cut length is performed by reiativeiy duil teeth vvhich cause bad surface quality of cut vvood, nonuniform thickness and inereased energy consumption. The stellite 12 vvas found to be superior in comparis-on to stellite 20 and both nickel base alloys vvhich vvere considered in these dulling tests. On the basis of this comparative research a conclu-sion vvas made that stellite 12 is highly recommended for aH types of savvs not only because of its best cutting ability, abrasion and corrosion resistance but also because of its best grindability vvhich enables good sharp-ening of teeth. in the second set of laboratory experiments the stellite 12 vvas compared vvith tvvo grades of tungsten carbide hard metals for three typical sorts of vvood. In these tests a cut length of 80 km vvas considered. in Fig. 16 the mechanism of dulling is schematically illustrated and compared betvveen savv steel and hard metal. The dulling mechanism of stellite teeth is similar to that of savv steel. The results obtained by stellite tipped teeth vvere better then both tungsten carbide grades vvith 6 % and 18 % of cobait. Stellite tipped savvs are superior to hard meta/ savvs especially for cutting fresh vvoods and thin cut. This vvas also confirmed in practice. The corrosion resistances of stellite 12 and its vveiding junetion are both better than tungsten carbide grades and their sol-dering junetions. The acid extracts of fresh vvood are chemically interaeting vvith cobait matrix and in this way badiy influence the resistance of carbides against separ-ation from cobait matrix. Stellite tipped savvs vvill replace hard metal savvs in cutting fresh vvood vvhiie they vvill retain their position in furniture and other finaiising industries employing treat-ed vvoods. Hovvever, ei/en in these applications the stellite tipping has certain advantages especially in the čase vvhere high speed cutting causing high impact loads is present. 8. SUMMARIZED ADVANTAGES OF STELLITE TIPPING (a) Advantages vvith respect to conventional steel saws: — The stellite tipping enables higher cutting produetivi-ty vvith respect to the savvs vvith teeth setting; — The cutting speed can be up to 30 % higher; — Manjša hrapavost površin rezanega lesa in večja natančnost reza po daljšem trajanju žaganja. Povečan je izplen rezanega lesa. — Zmanjšanje potrebne energije. Pri standardnih jeklenih žagah se poraba moči po štiriurnem žaganju v povprečju poveča za 15 %, pri uporabi stelitiranih žag pa samo za 1,5 %. — Ostrenje žag je bolj ekonomično in poraba žagnih trakov ali diskov je manjša. — Poprečni proizvodni stroški žaganja so manjši. Vse naštete prednosti se dosežejo v glavnem brez povečanja stroškov, ker se stroški stelitiranja skoraj v celoti izravnajo s tem, da ni več potrebno delo z na-krčevanjem in razperjanjem zob. (b) v primerjavi z uporabo trdokovinskih žag: — Povečana obstojnost rezalnega roba zob. — Izboljšanje možnosti za žaganje s tanjšim rezom. Varjeni spoj med stelitom in nosilnim jeklom je v splošnem precej trdnejši od lota med karbidno trdino in jeklom. — Konice stelitov so manj občutljive za poškodbe in nastale poškodbe se dajo lažje popravljati. — Stroški brušenja so nekoliko manjši, zaradi možnosti uporabe cenejših plošč. — Nižja cena stelitnih konic. Poraba časa in stroški dela za trdokovinske ali stelit-ne konice zob v praksi ne predstavljajo pomembnih razlik. (c) nekaj dodatnih razlogov za uvedbo stelitiranja žag — Rezanje trdih in zelo neenakomerno rastočih lesov (exotov) in lesov z mineralnimi vključki je mogoče samo s stelitiranimi žagami. — V zadnjem času se je izkazalo, da prinaša stelitiranje velike prednosti tudi pri rezanju mehkih lesov. — Z relativno majhnimi dodatnimi ukrepi za stelitiranje se dosežejo veliki ekonomski uspehi. 9. ZAKLJUČEK Stelitiranje žag močno podaljšuje življensko dobo zob in ima posebno pri žaganju nesušenih lesov pomembne prednosti pred vsemi danes razpoložljivimi načini zmanjševanja obrabe, vključno z oblaganjem zob s trdokovinskimi konicami. Tehnologija stelitiranja omogoča precejšnje zmanjšanje proizvodnih stroškov ob istočasnem povečevanju produktivnosti in izboljšanju kakovosti. Nabava potrebne opreme za stelitiranje se v večini žagarskih obratov amortizira v kratkem času. Zelo priporočljivo tehnično, kvalitetno in ekonomsko rešitev predstavlja organiziranje servisnih centrov za stelitiranje in ostrenje vseh vrst žag. Nadaljnji razvoj na področju stelitiranja žag bo usmerjen v kompletiranje proizvodnih programov kobal-tovih zlitin s posebnim poudarkom na specializaciji ponudbe, pri čemer bo imelo optimiranje sestave zlitin za določena področja in namene uporabe prav gotovo vse večji pomen. — The life time of cutting edge is longer and therefore enabies reduction of interruptions in the production; — A lovv cut roughness and narrow toierances can be retained for a longer time; — The yield of wood is higher; — The energy consumption and povver requirements are Iovver. — The increase of povver consumption in first four hours is approximately 15 % for steel savvs and 1.5 % for stellite tipped saws; — The sharpening of teeth is more economic and consumption of savv bands or discs is Iovver; — The average production costs are iovver. AH above mentioned advantages can be achieved vvithout additional costs because stellite tipping cost are comparabie to the costs related to svvagging and setting of conventionai steel savvs. (b) Advantages vvith respect to hard metal savvs: — The life time of cutting edge of teeth is longer; — A better possibitity for savving vvith thin cut because the vvelded stellite junction is stronger than soldered hard metat junction; — The stellite tips are less sensitive to damages and can be easier repaired; — The sharpening costs are Iovver because cheaper grinding plates can be used; — The priče of stellite tips is Iovver. The time consumption and labour costs for hard metal and stellite tipping are approximately the same. (c) Some additional reasons for introduction of stellite tipping — The savving of certain hard and irregularly grovvn vvoods (exots) and vvoods vvith mineral inclusions is often possible only vvith stellite tipped savvs; — Recent production experience confirms advantages of stellite tipping also for cutting soft vvoods; — With modest investments for stellite tipping relatively high savings and improvements can be achieved. 9. CONCLUSION The stellite tipping of savvs enabies longer life time of teeth and has many advantages especially in savving un-treated vvoods. It is superior to ali today knovvn ap-proaches for reduction of teeth vvear including hard metal tipping. The technology of stellite tipping enabies considerable lovvering of production costs vvhile at the same time quality and productivity are improved. It is expected that stellite tipping and sharpening vvill be organized in servicing centers. Further developments in stellite tipping vvill be aimed to an optimized assort-ment of alloys and forms of products for specialized ap-plications. LITERATURA / REFERENCES 1. Kirbach E., T. Bona: Stellit-bestuckung von VVeichholzsag-en (prevod), Lumber Manufacturing Update - Seminar, 26-28. Oct. 1981, Seattle, WA - Technical Report No.17, FORINTEK CANADA CORP., Aug.1981. 2. Kirbach E.: Methods of improving wear resistance and maintenance of saw teeth. Technical Report No. 3/1979. Forintek Canada Corp. 3. VOLLMER - prospekti: Das Stellitieren und Sharfen von Sagezahnspitzen. 4. ISELI - privatna informacija/pnVaf communication 5. ALLIGATOR - privatna informacija/pr/Va/ communication 6. STN-W. Niggl: Priporočila za stelitiranje žag/ Recommenda-tion for steiiitizing of saws 7. Rodič J.: Skrajševanje tehnološkega postopka od taline do žice/Shortenings of Techno/ogical Procedures from Melt to Wire; Železarski zbornik 22, dec. 1988, str. 101-109 8. Rodič J., W. Holzgruber, M. Haissig: Development of a new CW&BP* - Process for Specialty Steels and Superalloys, EUROMAT 89 - Conference, Aachen. Nov. 1989 9. MIL-PP d. o.o., Ljubljana: Posebne zlitine na osnovi Co za lesno industrijo/Special Co-base a/ioys for wood cutting; 1991 10. Rodič J.: Development of a New Process-Thermomechani-cal Consolidation of HCC Rods, not published 11. Rodič A.: Internal reports of metallographic examination of samples, Institute of metals and technologies, Ljubljana 1991 MIL-PP RAZVOJ IN PROIZVODNJA SPECIALNIH ZLITIN Ljubljana DEVELOPMENT AND PRODUCTION OF SPECIAL ALLOYS HORIZONTALNO KONTINUIRNO LITJE - HKL HORIZONTAL CONTINUOUS CASTING - HCC 61000 Ljubljana, Slovenija p.p. 431 - Lepi pot 11 Tel. +38 61 151 161, 216 709 Fax. (061)-213-780 MCC PROGRAM "MILIT" POSEBNE ZLITINE NA OSNOVI Co ZA LESNO INDUSTRIJO SPECIAL Co-BASE ALLOYS FOR WOOD CUTTING MILIT MILIT MILIT MI L I T MILIT MILIT 6 H 12H 1 2WH 1 H FH 2 1 H C 1,2 1 ,<* 1,6 2,5 1 ,8 0,3 Si 1 1 1,5 1 1 0,5 Mn 0,5 2,5 2,5 0,5 0,5 0,5 Cr 29 30 29 30 25 28 Ni 1,5 1 ,5 - 23 3 U 4,5 a 10 13 12 - Mo 0,5 - - - - 6 Co 60 5<+ 53 5 O 35 60 HRC 4S 48 54 55 30 OBLIKE FORHS O dimenzije dihens10ns 2.0 •■ i 8.0 •• dolžine len6ths l = 8 •» t 16 »i 350 ■« i mo i* dobavno stanje deliverv condition lito ali brušeno as čast or gnnded 2.0 •• 7 6.5 it l = 8 »e 7 16 •§ 1000 >* brušeno - gnnded F = 8 n£ 7 50 uE l = 8 >• 7 16 »i 1000 •■ brušeno - grinded B = 2.7 •« f 6.0 n H = 2.5 •> 7 4.0 u A - 3.2 » 7 6.5 >■ A = 4.0 m 7 7.5 u B = 3.0 7 5.0 n D = 2.8 m 7 4.5 H H - 3.5 7 6.5 u a = 30' D, = U *• D? = 5 u a = 50« A = 4.0 •■ 7 6.5 n a = 60' l = 1000 ii brušeno - grinded l = 1000 ii l = 1000 ■■ l = 1000 n brušeno - grinded brušeno - gnnded brušeno - grinded Faktor mejne intenzitete napetosti pri počasnem natezanju navodičenega jekla z visoko trdnostjo Threshold Stress /ntensity Factor at Siow-Strain-Rate Tension of High-Strength Hydrogen-Charged Steel B. Ule1, F. Vodopivec1, L. Vehovar1 and L. Kosec2 Konstrukcijska jekla z visoko trdnostjo in visoko napetostjo tečenja se vedno bolj uporabljajo celo za izdelavo manj zahtevnih strojnih delov. Zaradi razmeroma nizke žilavosti tovrstnih jekel in slabo izraženega prehoda v krhko stanje postajajo toliko pomembnejše njihove lomne značilnosti. Na izgubo lomne duktilnosti močno vpliva zlasti vodik v jeklu, čeprav pri tem ne učinkuje bistveno na napetost tečenja. Poslabšanje lomne duktilnosti pa je izrazito le pri počasnem natezanju navodičenega jekla, medtem ko ga pri konvencionainem nateznem preizkusu skoraj ne zaznamo. Mate koncentracije vodika v jeklu z visoko trdnostjo torej ne vplivajo na lomno žilavost takšnega jekla, peč pa imajo za posledico pojavljanje faktorja mejh-ne intenzitete napetosti. 1. UVOD Ena od znanih oblik porušitve jekel z visoko trdnostjo je tako imenovani zapozneli lom statično obremenje- Slika 1: Tipičen zapis pojava zapoznelega loma na nateznem preizkuš-ancu z zarezo, obremenjenem s konstantno obremenitvijo. Zapis velja za vodičeno jeklo, vrisana pa je tudi trdnost ob zarezi za jeklo brez vodika (preizkušano na zraku) /lit. (1)/. r . F'9- I-' iypical delayed-failure phenomenon for hydrogen-charged notched tensile specimens at constant load. The notch tensile strength of uncharged steel (measured in air) is also shovvn (Ref. i). Inštitut za kovinske materiale in tehnologije, Lepi pot 11, 61000 Ljubljana ~ Fakulteta za naravoslovje in tehnologijo, Oddelek za montani-stiko, 61000 Ljubljana The use of structural steels vvith high-tensile and high-yieid strength is increasing even in the manufacture of less demanding machine parts. Because of their relativen lovv toughness and poorly expressed transition into brittle state, their fracture properties are of major im-portance. The decrease in fracture ductility is in particular strongly influenced by hydrogen content in steel, al-though hydrogen does not essentially affect its yieid strength. Hovvever, the deterioration of fracture ductility is distinctive only at slovv-strainrate tension of hydrogen-charged steel, vvhereas it practically cannot be detected in a conventional tensile test. Consequently, the lovv concentration of hydrogen in high-strength steel does not influence its fracture toughness, but results in the appearance of a threshold stress intensity factor. 1. INTRODUCTION Deiayed fracture caused by stress-induced hydrogen segregation is one of the knovvn types of fracture of high-strength steel. This problem is characterised by the nucieation of a microcrack, vvhich then grovvs until it achieves a critical size, resulting in an abrupt fracture (Fig. 1). The incubation period, as well as the delayed tirne to failure occurrence, are prolonged vvith the decrease in load until, at a sufficiently lovv load, the delayed failure does not occur. Therefore, we can speak of threshold of appiied stress or threshold stress intensity factor Kth, vvhich can be considerably lovver than the critical stress intensity factor or fracture toughness K/c of steel. In the čase of hydrogen embrittlement the threshold stress intensity factor is iikevvise denoted as KHE. The threshold stress intensity factor is inversely pro-portionai to the hydrogen concentration in steel, vvhich leads to the idea that the mutual effect of hydrogen and applied stress provokes the nucieation of microcracks. It vvas also found that the incubation period strongly de-pends on the hydrogen concentration in steel, vvhiie the effect of the appiied stress magnitude is coniderably lovver. Since the effect of hydrogen on the mechanicai properties of high-strength steel is manifested by the de-creased fracture ductility at slovv-strain-rate tension and since such decrease depends on crack nucieation as well as on crack propagation, it is therefore logical that a slovv-strain-rate tension test vvill by ali means prove to ' Institute of meta/s and Technology, Lepi pot 11, 61000 Ljubljana " University of Ljubljana, Fac. of natural Science, Metal, dept., 61000 Ljubljana nega jekla, ki je posledica napetostno induciranega se-gregiranja vodika v jeklu. Pri tem pojavu pride najprej do iniciiranja prve mikrorazpoke, ki nato počasi raste, vse dokler ne doseže kritične velikosti, kar povzroči hipno porušitev (si. 1)1. Inkubacijski čas kot tudi čas do loma se podaljšujeta z zniževanjem statično delujoče obremenitve, vse dokler pri neki dovolj nizki obremenitvi zapozneli lom izostane. Govorimo torej lahko o pragu delujoče napetosti oziroma o faktorju mejne intenzitete napetosti Kth (threshold stress intensity factor), ki je lahko tudi občutno nižji od faktorja kritične intenzitete napetosti, to je od lomne žilavosti jekla K|C. V razmerah vodikove krhkosti označimo faktor mejne intenzitete napetosti tudi kot Khe Z znižanjem koncentracije vodika v jeklu se faktor mejne intenzitete napetosti zvišuje, kar napeljuje na misel, da je porajanje mikrorazpok posledica vzajemnega učinkovanja vodika in delujoče napetosti. Ugotovili so tudi2, da je inkubacijski čas zelo odvisen od koncentracije vodika v jeklu, le malo pa od velikosti delujoče napetosti. Ker se učinek vodika na mehanske lastnosti visoko-trdnega jekla manifestira z izgubo lomne duktilnosti pri počasnem natezanju in ker je lomna duktilnost jekla odvisna tako od porajanja kot tudi napredovanja razpok, je logično, da bo počasni natezni preizkus vsekakor primernejši za ugotavljanje vpliva vodika na lastnosti jekla kot pa konvencionalni natezni preiskus. Če je hitrost deformacije pri natezanju tako velika, da pride do loma v času, ki je krajši od inkubacijskega časa, učinek vodika na lastnosti jekla ne bo zaznaven. Poleg statičnih preizkusov s konstantno obremenitvijo (static delayed failure test) se za določevanje občutljivosti jekla za lom, induciran z vodikom, še največ uporablja natezni preizkus s cilindričnimi preizkušanci z zarezo po obodu. O tem priča opis Pollockove metode v reviji Metals Progress3. Pollock določa občutljivost jekla za lom, induciran z vodikom, z merjenjem sile loma cilindričnih preizkušancev z zarezo po obodu pri hitrosti na-teznaja 2x 10-4 mm s~1. Trdnost zarezanega preizkušan-ca je namreč v takšnih primerih manjša od trdnosti gladkega, kar neposredno odseva izgubo duktilnosti zaradi učinkovanja vodika. V strokovni literaturi je opisanih še več različnih načinov kvalitativnega določevanja občutljivosti jekla za lom, induciran z vodikom, ki pa vsi temeljijo v glavnem na enostavnem merjenju stopnje poslabšanja kontrakcije pri natezanju jekla4 5. Ker določevanje faktorja mejne intenzitete napetosti KTH le na osnovi rezultatov nateznega preizkusa v literaturi še ni ustrezno obdelano, smo raziskali rešitev tega problema. Izkoristili smo merljivo poslabšanje lomne duktilnosti jekla pri počasnem natezanju kot nadomestku za dolgotrajni statični natezni preizkus pri konstantni obremenitvi. Ob tem smo upoštevali hipotezo, po kateri poslabšanje lomne duktilnosti pri počasnem natezanju dejansko potrjuje obstoj faktorja mejne intenzitete napetosti, če se lomna žilavost takšnega jekla - merjena pri običajnih hitrostih natezanja - le malo ali pa sploh ne spremeni. Razvoj in teoretična utemeljitev te metode omogočata določanje faktorja mejne intenzitete napetosti kar na osnovi rezultatov nateznega preizkusa, s tem pa je določevanje občutljivosti jekla za lom, induciran z vodikom, bistveno bolj objektivno, kot je pri sedaj uporabljanih metodah. 2. TEORETIČNI DEL Vodik je v železu v atomarni obliki bodisi na intersti- cijskih mrežnih mestih bodisi vezan v večji ali manjši me- te more convenient in determining the influence of hy-drogen on steel properties than a conventionai tension test. If the deformation rate at tension is so iarge that failure occurs in a period shorter than the incubation one. the influence of hydrogen on the properties of steel will not be cognizable. Besides static test at constant load (static delayed failure test), the tension test on cylindrical specimens vvith a circumferential notch is more frequently used for the determination of hydrogen induced fracture of steel. This is evident in the description of Pollock s method in Metals Progress3. Pollock determines the sensitivity of steel to hydrogen-induced fracture by measuring the fracture load at a crosshead speed of 2x 10~* mm s'1 using cylindrical notched tensiie specimens. In this čase the strength of notched specimens is tovver than that of smooth specimens. vvhich directly reflects the decrease of ductility due to the effect of hydrogen. Many more methods for qualitative determination of hydrogen induced fracture of steel have been described in professional literature. Hovvever. aH of these are main-ly based on the measurement of the decrease of reduction of area at tension test of steel4 5. Since the determination of threshold stress intensity factor KTH only on the basis of the results of a tensiie test has not been ade-quateiy treated in literature, our attempts vvere aimed at finding a solution to this problem. Instead of a long-term static test at constant load, we used the siovv-strain-rate tension test for determining the measurabie decrease in fracture ductility of steel. We follovved the hypothesis that the decrease in fracture ductility at siovv-strain-rate tension test actually confirms the existence of a threshold stress intensity factor if the fracture toughness of such steel - measured at conventionai strain rate - changes only slightly or doesn 't change at ali. The development and theoretical justification of this method prove its usefuiness in determining the threshold stress intensity factor on the basis of results at-tained in a tensiie test. In this way, the determination of steel sensitivity to hydrogeninduced fracture is essen-tiaily more objective than in currently used methods. 2. THEORY Hydrogen is present in iron either at interstitial sites at the lattice or bound as trapped hydrogen to different discontinuities of the cristall lattice - "traps" - and thus referred to as trapped hydrogen. Some hydrogen in moleč ular form is always found in the microvoids as vvell. The partial mola/ volume of hydrogen in iron as vveli as in most other metals is surprisingly high (appr. 2 cm3/mol of hydrogen or 0,33 nm3/atom)6-8. This results in a strong interaction betvveen hydrogen intersticials in the cristai lattice and elastic-stress fields of the loaded metal lattice. A thermodynamic analysis of this process, based on the assumption that hydrogen is a completely mobile component, vvas performed by Li, Oriani and Darken9, who found the follovving relation: u — /io= ct/j E;j dV (1) The distortion fieid around the hydrogen atom is described by the deformation tensor E:j; is the stress tensor vvhich determines the stress state originating from externai loads acting on cristai lattice. Ho is the chemical potential of hydrogen in the absence of exter-nal stress, vvhile u represents the chemical potential of hydrogen under externai stress. The difference betvveen ri na različne diskontinuitete kristalne mreže, ki jih imenujemo s skupnim imenom pasti in od tod v pasteh ujeti vodik (trapped hydrogen). Nekaj vodika je v železu vedno tudi v porah v molekularni obliki. Parcialni molski volumen vodika v železu in večini drugih kovin je presenetljivo velik (približno 2 cm3/mol vodika, oziroma 0,33 nm^atom)6-8. Posledica tega je močna interakcija med vodikovimi intersticijami v kristalni mreži ter polji elastičnih napetosti v obremenjeni kristalni mreži kovin. Li, Oriani in Darken9, so s termodinamično analizo tega problema, pri čemer so vodik v železu obravnavali kot povsem mobilno komponento, prišli do izraza: u—H0=a,, Ejj d V (1) Deformacijski tenzor Ey opisuje deformacijsko polje okrog intersticijskega atoma vodika, nc^ je napetostni tenzor, ki opredeljuje napetostno stanje, izvirajoče od zunanje mehanske obremenitve kristalne mreže. Z |i0 je v zgornjem izrazu (1) označen kemijski potencial vodika v neobremenjeni kristalni mreži kovine, ji pa je kemijski potencial vodika v mehansko obremenjeni mreži kovine. Razlika potencialov je zato enaka delu, ki je potrebno za vgnezdenje intersticijskega vodika v polje delujočih napetosti. Gradient napetosti torej povzroči gradient kemijskega potenciala vodika, le-ta pa predstavlja gonilno silo za difuzijo intersticijsko raztopljenega vodika. Rezultat tega je segregiranje vodika v neenakomernem polju napetosti: vodik se zaradi reverzibilne dilatacije kristalne mreže s pripadajočo pozitivno spremembo volumna, ki spremlja vgnezdenje vodikovih intersticij, koncentrira v področjih prevladujočih nateznih napetosti, medtem ko se področja s prevladujočimi tlačnimi napetostmi z vodikom osiromašijo. Prerazporejanje vodika v obremenjeni kristalni mreži poteka toliko časa, dokler ni dosežena v vseh točkah mreže ravtežna koncentracija vodika, določena z izrazom: a E dV [H] = [H]0 exp ^uJaHjl n I (2) pri čemer je [H]a koncentracija enakomerno porazdeljenega vodika v neobremenjeni kristalni mreži. Če upoštevamo le volumsko spremembo v okolici vrinjenih vodikovih atomov, lahko izraz (2) zapišemo v obliki: [H] = [H]0 exp ■ RT (3) kjer je z am označena hidrostatična komponenta napetostnega tenzorja [am=1/3 (ax + ay + arz), VH pa je parcialni molski volumen vodika v železu. Z enačbo (3) je mogoče izračunati koncentracijo vodika v lokaliziranem področju, na primer v zoženem vratu nateznega preiskušanca, kjer deluje hidrostatična napetost am. Ko koncentracija vodika [H] na tem mestu doseže kritično vrednost [H]cr, ko je torej K| = KTH, moramo računati z iniciiranjem mikrorazpok in zapoznelim lomom jekla. Problem je analitično rešil Gerberich10, ki je za faktor mejne intenzitete napetosti izpeljal izraz: Kth = M m M [H]0 (4) aVH 2 a pri tem ima a eksperimentalno ugotovljeno vrednost 2/5 mm"1'2. Odvisnost (4) je eksperimentalno dobro potrjena, vendar pa pri napetostih tečenja, ki so nižje od 1200 MPa, pogosto prihaja do neujemanja med enačbo (4) in rezultati eksperimentov. To neujemanje lahko deloma razložimo z odvisnostjo razmerja [H]cr/[H]0 od napetosti tečenja jekla. Farrell and Ouarrell4 sta namreč ugo- the potentials is the vvork needed to ptace the hydrogen into the active stress field. The gradient of chemical potential of hydrogen is therefore caused by the stress gradient and represents the driving force for the diffusion of interstitially dissolved hydrogen, resulting in hydrogen segregation in non-uniform stress field. Hydrogen concentrates in the areas of predominantly tensiie stresses due to the rever-sibie diiatations of the cristal tattice vvith the correspond-ing volume changes accompanied by the insertion of hy-drogen interstitiais, vvhile the compressiveiy strained re-gions become impoverished vvith hydrogen. The redis-tribution of hydrogen in the strained cristal iattice takes plače until an equilibrium concentration of hydrogen is achieved in aH points of the cristal Iattice. This is expressed by: [H]=[H]0expa«EJdV HT (2) vvhere [H]0 is the concentration of hydrogen, uniformly distributed vvithin the unstrained cristal Iattice. If only the volume change around the inserted hy-drogen atoms is considered, the equation (2) may be ex-pressed as: [H]=[H]0exp dVH RT (3) vvhere om is the hydrostatic component of stress tensor om= 1/3 (ax+ ay+ oz) and VH is the partiai molal volume of hydrogen in iron. Equation (3) may be used to calculate the concentration of hydrogen in a localised area, as for example in the narrovved neck of tensiie specimens vvith hydrostatic stress <7m. Microcracks nucieation and delayed fracture of steel can be expected vvhen the hydrogen concentration [H] in this region achieves the critical value [H]cr, i.e. vvhen K,= Kth. This problem vvas so/ved analyticaily by Gerberich10, who expressed the threshold stress intensity factor in the form of: KTH=*JLln[±!lcr. a VH [H]0 2 a (4) vvhere factor a reaches the experimentally determined value of 2/5 mm~1/2. The reiation (4) is experimentally vvell confirmed, ai-though some discrepancies can often be observed betvveen Eq. (4) and the experimental resuits at yield strength belovv 1200 MPa. These discrepancies can be partly explained by the dependance of the [H]cr/[H]0 ratio on the yield point. Namely Farrell and Ouarrell4 ascer-tained that larger concentrations of hydrogen are needed to produce embrittlement in steel vvith Iovver yield strength, vvhich they expressed vvith the reiation [H]cr °o 1/ays. Kim and Loginow11 proved that the content of solu-ble hydrogen in steel vvas proportionai to the yield strength, thus [H]0°o cry5. If both statements are taken into account this can be vvritten as: (5) [HL_ P [H]o Pys vvhere p is a constant for specific types of steel vvith determined hydrogen concentration. By substituting (5) for (4), we arrive at the vvell-knovvn Gerberich eguation for threshold stress intensity factor in its final form: Kt, Min-L■ aVH avs 2a (6) tovila, da so ze doseganje krhkosti v jeklih z nižjo napetostjo tečenja potrebne višje koncentracije vodika, kar sta zapisala kot [H]croo1/ays. Kim and Loginovv" pa sta dokazala, da se v jeklih z višjo napetostjo tečenja topi več vodika, torej [H]0ooays. Z upoštevanjem obeh navedenih ugotovitev lahko zapišemo: [H]Cr_ P [H]0 Pys pri čemer je p konstanta za posamično vrsto jekla in za določeno vsebnost vodika v njem. Ko substituiramo (5) v (4), dobimo znano Gerberic-hovo enačbo za faktor mejne intenzitete napetosti v njeni končni obliki: (5) aV„ 2a (6) /0,05 E, n2 E crvs K|C= |/-o-(MPa m12) (7) P = CTysexp aVH f 0,05 E, n2 E Gy, 2a (10 Since hydrogen in steel mostly affects the fracture ductility at slovv-strain-rate tension, it would therefore seem adequate for further analysis to determine the re-lation betvveen fracture toughness Klc and the parameters of tensiie test. Such a relation, knovvn as the Hahn-Rosenfield correlation12 13, is given by: K, 0,05 E, rf Eov. (MPa m1/2) (7) vvhere E, is the fracture ductiiity, calculated from the ac-tual reduction of area Z using the eguation: Ef= in [1/(1 -Z)] (8) vvhereas the strain hardening exponent n can be calculated from the uniform elongation eu using the eguation: Ker vodik v jeklu še najbolj vpliva na lomno duktilnost jekla pri počasnem natezanju, je za nadaljnjo teoretično analizo smiselno poiskati soodvisnost med lomno žila-vostjo Klc in parametri nateznega preizkusa. Takšno soodvisnost poznamo pod imenom Hahn-Rosenfieldova korelacija1213, ki ima naslednjo obliko: n= In (1+ ej (9) Pri tem je E, lomna duktilnost, ki jo izračunamo iz znane kontrakcije jekla Z po formuli: Ef = In [1/(1—Z)] (8) medtem, ko eksponent deformacijskega utrjevanja n izračunamo iz enakomernega raztezka eu po formuli: n = ln(1+ej (9) Ker pri običajnih hitrostih obremenjevanja, kakršne uporabljamo pri merjenju faktorja kritične intenzitete napetosti KIC, ne zaznamo opaznejšega poslabšanja duktilnost) jekla, ki bi ga sicer lahko pripisali vplivu majhnih koncentracij vodika v jeklu (okoli 1 ppm)4 14, se zdi utemeljena hipoteza, da poslabšanje lomne duktilnosti jekla pri počasnem natezanju dejansko odraža eksistenco faktorja mejne intenzitete napetosti KTH. V skladu s to hipotezo bi lahko Hahn-Rosenfieldovo korelacijo (7) uporabili kar za izračunavanje faktorja KTH, potem ko bi v enačbo (7) vstavili vrednosti, izmerjene pri počasnem natezanju. Ker pa je poznana tudi teoretično izpeljana Gerberichova enačba za KTH (6), v kateri je neznana le vrednost p, je mogoče po izenačenju enačb (6) in (7) vrednost p izraziti eksplicitno: Since the evident vvorsening of fracture ductiiity, vvhich may be attributed to the small amounts of hydrog-en in steel (appr. 1 ppm), cannot be detected4 14 by con-ventional strain-rate used in measurements of critical stress intensity factor KIC, it therefore seems that the hy-pothesis according to vvhich the decreased fracture ductility at siovv-strain-rate tension reflects the exis-tence of threshold stress intensity factor is justified. In accordance vvith this hypothesis. the Hahn-Rosen-field correlation (7) could be used for the calculation of Kth after the vaiues measured at slovv-strain-rate have been inserted into equation (7). Since Gerberich s theor-etically developed equation for KTH (6) is also knovvn (P being the only unknovvn value). it is therefore possible to express the value of p explicitly after the balance of Eq. (6) and (7): P= cTyS exp oVh R T f-t 05 E, n2 Ea, is Oj 2 a (10) RT1' 3 V izrazu (10) so veljavne vrednosti za ays, E, ter n, kot že rečeno, izmerjene pri počasnem nateznem preizkusu. Verificiranje postavljene hipoteze se bo torej reduciralo na ugotavljanje konstantnosti veličine p, ki mora biti neodvisna od napetosti tečenja jekla cys. Konstantna vrednost p pomeni, da je postavljena hipoteza pravilna in da s podatki počasnega nateznega preizkusa lahko izračunamo Kth kar z enačbo (7). 3. EKSPERIMENTALNI DEL Za eksperimentalno delo smo izbrali jeklo Č.4751 z naslednjo kemijsko sestavo: 0,38% C, 0,99% Si, 0,38% Mn, 0,012% P, 0,010% S, 5,19% Cr, 1,17% Mo ter 0,23% V. Po homogenizacijskem žarjenju in normalizaciji smo iz kovanih palic s premerom 16 mm za natezni preizkus /4s already mentioned, the relevant vaiues of oys, E, and n in Eq. (10) are measured in a slovv-strain-rate tension test. The verification of the postulated hypothesis vvill thus be reduced to the measurement of constancy of the p value. vvhich has to be independent of the yield stress oys of steel. The constant p value means that the hypo-thesis is correct and that KTH can be calculated using the Eq. (7) on the basis of siovv-strain-rate tensiie test data. 3. EXPERIMENTAL Steel Č.4751 containing (wt-%) 0.38% C, 0,99% Si, 0.38% Mn. 0,012% P, 0.010% S. 5.19% Cr, 1,17% Mo and 0,23% V has been chosen for experimental vvork. Cylin-drical tensiie specimens vvith a diameter of 10 mm, gauge length of 100 mm and total length of 250 mm vvere machined from the forged rod after it had been homoneously anneaied and normalized. Specimens vvere thermally treated in a vacuum annealing furnace. They were austenitised at 980° C for a short period. quenched in a flovv of gaseous nitrogen and then tempered at temperatures of 620° C, 640° C and 670° C re-spectively. Thus. three separate and distinct classes of yield strength - 1220 MPa. 1020 MPa and 900 MPa re-spectively vvere achieved. The cathodic charging of thermally treated tensiie specimens vvas carried out for 1 hour in 1 N sulfuric acid at a current density of 0,3 mA/cm2. The experimentai set-up for cathodic polarisation of tensiie specimens, as shovvn in Fig. 2 is composed of a potentiostat and corrosion celi vvith electrodes. izdelali 250 mm dolge cilindrične preiskušance s premerom 10 mm in dolžino 100 mm. Preiskušance smo toplotno obdelali v vakuumski žarilni peči, tako da smo jih po kratkotrajni austenitizaciji pri 980° C kalili v toku plinastega dušika, nato pa popuščali pri temperaturah 620°C, 640° C oziroma 670° C. Na ta način smo dobili tri ločene in dobro definirane trdnostne razrede z napetostjo tečenja 1220 MPa, 1020 MPa oziroma 900 MPa. Toplotno obdelane preiskušance za natezni preizkus smo navodičili z enournim katodnim polariziranjem v 1 N raztopini žveplene kisline pri gostoti toka 0,3 mA/cm2. Eksperimentalni sklop s katodnim polariziranjem preizkušancev, sestavljen iz potenciostata in korozijske celice z elektrodami, je prikazan na sliki 2. Natezne preizkuse smo opravili na nateznem trgal-nem stroju INSTRON, potem ko smo natezne preizku-šance po končanem navodičenju 24 ur zadrževali na zraku, da so se koncentracije vodika v jeklu približale resi-dualnim vrednostim (približno 0,7 ppm), ki se nato časovno skoraj niso več spreminjale. Za hitrost natezanja smo izbrali tako hitrost 1 mm/ min, značilno za običajni natezni preizkus, kot tudi hitrost 0,1 mm/min, značilno za počasno natezanje. Merili smo napetost tečenja ays (MPa), natezno trdnost ctts (MPa), maksimalni enakomerni raztezek eu ( x 100%) ter kontrakcijo jekla Z ( x 100%). Lomno duktilnost E, in eksponent deformacijskega utrjevanja n smo izračunali z enačbama (8) in (9). Mikrofraktografske preiskave prelomnih površin vo-dičenih in pri hitrosti natezanja 0,1 mm/min obremenje-vanih nateznih preizkušancev smo opravili s scanning elektronskim mikroskopom JEOL JSM-35 (SEM). 4. REZULTATI Izmerjene mehanske lastnosti navodičenega jekla kot tudi jekla brez vodika (pod 0,05 ppm) so zbrane v tabeli 1. V tej tabeli so zbrane še lomne žilavosti K,c jekla, izračunane s Hahn-Rosenfieldovo korelacijo (7) na osnovi rezultatov običajnih nateznih preizkusov pri hitrosti natezanja 1 mm/min ter nadalje še faktorji mejne intenzitete napetosti KTH vodičenega jekla, izračunani z isto enačbo (7), vendar na osnovi rezultatov počasnega natezanja pri hitrosti 0,1 mm/min. Slika 2: Eksperimentalni sklop za vodičenje nateznih preiskušancev s katodnim polariziranjem. (P-potenciostat, D-natezni preizkušanec, G-grafitni protielek-trodi, K-kalomelova elektroda in R-rotometer). Fig. 2: Experimentai set-up for hydrogen charging of tensiie specimens with cathodic poiarisation. (P-potenciostat. D-tensiie specimen. G-graphite eiectrodes. K-kaiomei eiectrode. Ft-rotameter). The tension tests vvere made on an INSTRON testing machine. after hydrogen charging of specimens vvas compieted and the specimens exposed to air for 24 hours. This enabied the concentrations of hydrogen in steel to approach the residuai vaiues (appr. 0,7 ppm), vvhich remained near/y time-independent. The tension tests vvere performed at conventional strain rate i.e. at a crosshead speed of 1 mm/min as well as at iovver-strain-rate i.e. at a crosshead speed of 0,1 mm/min. The yie/d strength cjys (MPa), tensiie strength aTS (MPa), max. uniform elongation eu (x100%) and the reduction ofarea Z (x100%) ivere measured. The fracture ductiiity E, and the strain hardening exponent n vvere calculated using equations (8) and (9) respectiveiy. The fracture surfaces of tensiie specimens tested at a crosshead speed of 0,1 mm/min vvere examined in the scanning electron microscope JEOL JSM-35 (SEM). R Tabela 1: Mehanske lastnosti jekla brez vodika in istega jekla po navodičenju Table 1: Mechanicai properties of uncharged and hydrogen-charged steel Hitrost natezanja 1 min/min Crosshead speed 1 mm/min Napetost tečenja Yield strength ays(MPa) Enakomerni raztezek Uniform elongation e„ x 100% Kontrakcija Reduction ofarea Zx 100% Lomna žilavost Facture toughness Hitrost natezanja 0,1 mm/min Crosshead speed 0.1 mm/min MPa. m1' Napetost tečenja Yield strength oys(MPa) Enakomerni raztezek Uniform elongation eux 100% Kontrakcija Reduction of area Zx 100% Faktor mejne inten. napet. Threshold stress inten. factor MPa • m1 Konstanta Constant /En (10)/ /Eq. (10)/ MPa Jeklo brez vodika, uncharged steel 924 8,7 52 126,9 910 8.5 51 1010 7,4 51,3 112,5 1027 6,5 50,3 1270 6,4 50 107,5 1214 6,2 50,3 Navodičeno jeklo, hydrogen charged steel 885 8.4 50.3 117,2 899 8,1 47,7 109,8 4005 1082 7,2 49,3 110,1 1078 6,5 42,7 90,1 4223 1209 6,1 47,3 96,3 1226 6,0 27,3 67,4 4037 V tabeli 1 so prikazane tudi vrednosti za konstanto p, izračunane s pomočjo enačbe (10) na osnovi rezultatov počasnega natezanja. Slika 3: Diagram sila-deformacija pri počasnem natezanju jekla trdnostnega razreda 1300 MPa; a) brez vodika (pod 0,05 ppm) in b) 24 ur po vodičenju (ca. 0,7 ppm vodika). Fig. 3: Load-deformation diagram obtained at slow-strain-rate tension test of steel vvith yieid strength of 1300 MPa. (a) vvithout hydrogen (iess than 0.05 ppm) and (b) 24 hours after hydrogen charging (appr. 0.7 ppm hydrogen). V diagramu sila-deformacija na sliki 3 je prikazana odvisnost med silo in raztezkom pri počasnem natezanju jekla s trdnostjo ca. 1300 MPa. Odvisnost, označena z a), velja za jeklo brez vodika (pod 0,05 ppm), z napetostjo tečenja 1070 MPa. trdnostjo 1286 MPa, enakomernim raztezkom eu = 6% in kontrakcijo Z = 49%, medtem ko velja odvisnost, označena z b), za jeklo istega trdnostnega razreda, ki pa je bilo natezano 24 ur po vodičenju (ca. 0,7 ppm vodika). V tem primeru smo namerili napetost tečenja 1090 MPa, trdnost 1284 MPa, enakomerni raztezek eu = 5,6% ter kontrakcijo Z = 39%. Mikrofraktografske preiskave prelomnih površin na-vodičenih nateznih preizkušancev kažejo, da z vodikom inducirani lom pri počasnem natezanju takšnih preizkušancev ne ostane povsem duktilnega tipa, celo pri preiz-kušancih z relativno nizko napetostjo tečenja ne (slika 4). Pri višji napetosti tečenja so prelomne površine navodičenega ter počasi natezanega jekla mešane narave; poleg kvazicepilnih ploskev najdemo na prelomnih površinah tudi jamičasta duktilna področja ter številne grebene, nastale s trganjem (slika 5). 5. RAZPRAVA Analiza rezultatov mehanskih preizkusov (Tabela 1) kaže, da je lomna žilavost K,c, s katodnim polariziranjem navodičenega jekla z visoko trdnostjo, le malo manjša od lomne žilavosti enakega jekla brez vodika, kot o tem tudi sicer lahko sklepamo iz diagrama na sliki 1. Počasnejše natezanje pri jeklu brez vodika ne povzroči kakšnih opaznejših sprememb, medtem ko se pri Slika 4: Jamičasta duktilna prelomna površina s posameznimi kvazice-pilnimi detajli (B) pri vodičenem in počasi natezanem jeklu z napetostjo tečenja ca. 900 MPa. Fig. 4: Dimpied ductile fracture area vvith some quasicleavage details (B) in hydrogen-charged and slow-strain-rate tested steel vvith yield strength of 900 MPa. Slika 5: Mešana oblika preloma vodičenega jekla z napetostjo tečenja ca. 1070 MPa. Poleg cepilnih oziroma kvazicepilnih ploskev (B) je na prelomni površini moč zaslediti tudi jamičasta duktilna področja ter številne grebene, nastale s trganjem. Fig. 5: Mixed fracture mode on hydrogen-charged steel vvith yield strength of 1070 MPa. Besides cleavage and quasicleavage facets (B). dimpied ductile areas and many tear ridges can also be observed. navodičenem jeklu močno poslabša lomna duktilnost, t.j. kontrakcija jekla, ne pa tudi enakomerni raztezek, trdnost in napetost tečenja takšnega jekla. Poslabšanje lomne duktilnosti navodičenega jekla pri počasnem na-tezanju dejansko kaže na obstoj faktorja mejne intenzitete napetosti KTH (KHE). S Hahn-Rosenfieldovo korelaci-jo (7) izračunane vrednosti KTH dajejo namreč, po substituciji v Gerberichovo enačbo (6), za konstanto [} vrednost približno 4100 MPa. Ta vrednost je neodvisna od napetosti tečenja jekla in zato v okviru eksperimentalne natančnosti merjenja res konstantna količina, skladno z Gerberichovim modelom. Eksperimenti so nadalje pokazali, da je bila uporabljena hitrost natezanja 0,1 mm/min že dovolj majhna, da smo lahko iz poslabšanja lomne duktilnosti navodičenega jekla izračunali take vrednosti faktorja mejne intenzi-tete napetosti KTH. za katere je p konstanta. Če upoštevamo, da velikost plastične cone v trenutku loma navodičenega preizkušanca dosega velikost približno polovice vratu preiskušanca (I = 3 mm), dobimo za Ec, upoštevaje hitrost natezanja v=1,6x10~3mm s-1 (0,1 mm/min), vrednost Ec = v/I = 5,3 x 10"4 s-1. V strokovni literaturi15 navajajo za nerjavna jekla nekoliko višje vrednosti Ec, približno 10_1 s-1. Raziskave Nakana in sodelavcev16, opravljene s počasnim natezanjem vodičenega jekla z napetostjo tečenja 500 MPa, pa kažejo, da se pri zadostni koncentraciji vodika v jeklu kontrakcija jekla asimptotično približa neki znižani vrednosti že pri kritični hitrosti deformacije Ec= 10-4 s-1, to pa je že velikostni red naših izmerjenih vrednosti. Ta hitrost deformacije je namreč že dovolj majhna, da Cottrellovi oblaki vodikovih atomov lahko potujejo skupaj z disloka-cijami globoko v plastično cono nateznih preizkušancev. Diagram na sliki 3 potrjuje, da vodik ne vpliva bistveno na mobilnost dislokacij v zgodnjih fazah deformacij-skega procesa pri natezanju, saj skoraj ne učinkuje na napetost tečenja, trdnost in enakomerni raztezek jekla, pač pa le na kontrakcijo jekla. Vodik torej spreminja obliko diagrama sila-deformacija šele od pojavljanja plastične nestabilnosti dalje, kar se dobro ujema z navedbami iz različnih literaturnih virov17-25. Po teh navedbah vodik ne vpliva niti na zgodnje nukleiranje mikropor, niti na gostoto mikropor, ko se stopnja deformacije približuje lomni deformaciji. Očitno je zato vpliv vodika zaznaven šele v fazi rasti mikropor in/ali fazi njihovega združevanja. Do pospešene rasti in koalescence mikropor v tej fazi pa lahko pride tudi z mehanizmom ločevanja prostih površin, na katerih je adsorbiran vodik26. Na sliki 6 je shemat-sko prikazana rast in koalescenca mikropor vzdolž meje dveh kristalnih zrn. Mehanizem koalescence mikropor z ločevanjem prostih površin, na katerih je adsorbiran vodik, prične delovati, ko se oblikuje troosno napetostno stanje v zoženem delu nateznega preizkušanca. Posledica tega je že opisano "zgoščevanje" zadnje faze plastične deformacije pri počasnem natezanju navodičenega jekla. Mikrofraktografske preiskave samo še ilustrirajo pravkar opisani mehanizem loma. Pojasnjujejo namreč lome jamičaste duktilne vrste, pri katerih pa kažejo stene in dna jamic posamične mikromorfološke značilnosti cepilnega oziroma kvazicepilnega loma. Res smo tudi pri naših raziskavah vodičenega in počasi natezanega jekla višjega trdnostnega razreda opazili poleg duktilnih jami-častih področij še trganja (tearing), ki so sicer značilna za jekla z zadostno duktilnostjo in dovolj majhno napetostjo tečenja, da do porušitve lahko pride s plastično deformacijo. Poleg detajlov takšne vrste pa smo opazili na prelomnih površinah tudi področja kvazicepilne narave, često na samem obrobju večjih in globjih, lijakasto 4. RESULTS The mechanical properties of hydrogen-charged as vvell as hydrogen uncharged steel (/ess than 0,05ppm) are presented in Table 1 The fracture toughness of steel K,c, ca/culated according to the Hahn-Rosenfieid correlation (7) on the basis of conventional tension tests made at a crosshead speed of 1 mm/min, as vveil as the thresho/d stress intensity factor KTH of cathodic charged steel. also calculated using eguation (7), but on the basis of results obtained at a crosshead speed of 0,1 mm/ min, are also shovvn in Table 1. The vaiues of constant ji are also given in Table 1. These vvere calculated using equation (10), on the basis of slow-strain-rate tensile test data. The /oad-deformation diagram (Fig. 3) shovvs the re-lation betvveen load and elongation at siovv-strain-rate tension of steel vvith a tensile strength of appr. 1300 MPa. Curve a) denotes the uncharged steel (less than 0,05 ppm hydrogsn), having a yield strength of 1070 MPa, tensile strength of 1286 MPa, uniform elongation eu=6% and reduction of area Z=49%, vvhereas Curve b) denotes hydrogen-charged steel (appr. 0,7ppm hydrogen) of the same strength class. being tested 24 hours after it had been charged. In this čase the yield strength of 1090 MPa, tensile strength of 1284 MPa, uniform elongation eu = 5,6% and reduction of area Z= 39% vvere measured. The microfractographic axaminations of fracture sur-faces of hydrogen-charged specimens confirm that hy-drogen-induced fracture at slovv-strain-rate tension does not remain predominantly ductile. even in specimens vvith a relatively lovv yield strength (Fig. 4). The fracture surfaces of hydrogen-charged steel vvith higher yield strength, stretched at slovv-strain-rate tension, are of mixed mode. In addition to quasicleavage details, ductile dimpled areas and numerous tear ridges have also been observed (Fig. 5). 5. DISCUSSION The anaiysis of mechanical testing data (Table 1) shovvs that the fracture toughness K/c of hydrogen-charged high-strength steel is only slightly lovver than that of the uncharged steel. vhich can also be con-cluded from the diagram in Fig. 1. The slovv-strain-rate tension of uncharged steel does not provoke any noticeable changes, vvhereas that of hy-drogen-charged steel strongly decreases the fracture ductility. i.e. the reduction of area. but does not affect the uniform elongation, tensile strength and yield strength of such steel. The deterioration of fracture duc-tility of hydrogen-charged steel at slovv-strain-rate tension indicates the existence of the threshold stress in-tensity factor KTH (KHE). Namely, the KTH - vaiues, calculated vvith the Hahn-Rosenfieid correlation (7) after being substituted into Gerberichš equation (6), give an approximate value of 4100 MPa for the /)-constant. With-in the experimental error of the measure- ments, the obtained value is constant and, in accordance vvith Gerber-ich 's model, independent of the yield strength of steel. The experiments further shovved that the applied crosshead speed of 0,1 mm/min vvas sufficiently low to enable the calculation of KTH - vaiues from the decrease in fracture ductility of hydrogen-charged steel, i.e. the calculation of KTH - vaiues for vvhich f) is a constant. Considering that the size of the plastic zone of a hy-drogen- charged specimen is approximately half of the neck diameter (1=3 mm) at fracture, the crosshead speed is v= 1,6x 10~3 mm s~' (0.1 mm/min), then a value of Ec= v/l= 5,3x 10-4 s'1 is obtained. Troosno napetostno stanje zaradi pojavljanja vratu Slika 6: Shematski prikaz nastajanja por, njihove rasti in koalescence vzdolž meja zrn, na katerih je adsorbiran vodik /lit. (26)/. Fig. 6: Schematic representation of microvoid formation. growth and coaiescence aiong grain boundaries where hydrogen is ad-sorbed (Ref. 26). oblikovanih jamic (slika 4, detajl B), pa tudi kot povsem samostojna plitvejša področja. 6. SKLEPI Na osnovi opravljenih raziskav smo ugotovili, da izgubo lomne duktilnosti pri počasnem natezanju navodi-čenega jekla z visoko trdnostjo lahko uspešno izkoristimo za kvantitativno določevanje faktorja mejne intenzitete napetosti KTH (KHE)- Ugotovili smo namreč, da majhne koncentracije vodika (pod 1 ppm) v jeklu, ki je bilo navo-dičeno s katodnim polariziranjem, ne vplivajo bistveno na lomno žilavost, merjeno pri običajnih hitrostih nateza-nja (1 mm/min). Počasno natezanje (0,1 mm/min) navo-dičenega jekla z visoko trdnostjo pa poslabša lomno duktilnost, kar nakazuje obstoj faktorja mejne intenzitete napetosti KTH. Z izračunanjem faktorjev KTH s pomočjo Hahn-Rosenfieldove korelacije (7) in vstavljanjem teh vrednosti v Gerberichovo enačbo (6) smo ovrednotili parameter P /enačba (10), tabela 1/. Dobili smo konstantno vrednost okoli 4100 MPa, neodvisno od napetosti tečenja jekla, kot to tudi zahteva Gerberichov model za KTH. Mikrofraktografske preiskave prelomnih površin počasi natezanega vodičenega jekla z visoko trdnostjo ka-čejo, da je prelom lokalno še vedno tudi duktilne vrste. Kljub detajlom kvazicepilne narave smo v vseh primerih našli še duktilne grebene, nastale s trganjem, in jamiča-sta področja duktilnega tipa. Kvazicepilna oblika loma na obrobju večjih in globjih lijakasto oblikovanih jamic dokazuje, da so le-te rasle in se medsebojno zlivale tudi z mehanizmom ločevanja prostih površin, na katerih je bil adsorbiran vodik. Professionai literature15 quotes somewhat higher Ec -values for stainless steels, approximately 10~1 s~'. How-ever, the investigations performed by Nakano and co-workersw on hydrogen-charged steel vvith yield strength of 500 MPa using siow-strain-rate measurements shovv that at sufficient concentration of hydrogen in steel the reduction of area asymptotically approaches the re-duced value already at a criticai deformation rate of Ec= 10-4 s"', vvhose magnitude is of the same order as found in our investigations. This deformation rate is, in fact, iovv enough to enable the Cottrell atmosphere of hydrogen atoms pinned on dislocations to penetrate deep into the piastic zone of tensiie specimens. The diagram on Fig. 3 confirms that hydrogen has no essential influence on the mobility of dislocations in ear-iier phases of the deformation process at tension. It has aimost no effect on the yield strength, tensiie strength and uniform eiongation of steel, but only on the reduction of area. Thus hydrogen changes the shape of a ioad-deformation diagram only after the appearance of piastic instability, as aiready found by a number of au-thors,7~ss. According to these sources. hydrogen has no effect on the early nucieation of microvoids, nor on microvoid density, when the deformation approaches the fracture deformation. Therefore, the effect of hydrogen becomes obvious only at the stage of microvoid grovvth and/or during their coaiescence. The grovvth of microvoid and their coaiescence can also be acceierated by a mechanism of separation of internat interfaces vvhere hy-drogen is adsorbed26. The grovvth and coaiescence of microvoids along the grain boundary are schematically shovvn in Figure 6. The mechanism of microvoid coaiescence and the separation of internaI interfaces due to adsorbed hydrogen becomes operative vvhen the triaxial stress state in the narrovv neck of the tensiie specimen is formed, resulting in the previousiy described "conden-sation" of the tast stage of piastic deformation at siovv-strain-rate tension of hydrogen-charged steel. The microfractographic investigations are an addi-tionai illustration of above-mentioned fracture mechanism. They explain the ductile-dimpled types of fracture in vvhich the vvalls and bottoms of dimpies exhibit individual micromorphological characteristics of the cleavage or quasi-cleavage type of fracture. It is true that our investigations of hydrogen-charged high-strength steel at slow-strain-rate also shovved, in addition to ductile-dim-pled areas. tearing regions typicai for steels vvith sufficientiy high ductiiity and yield strength iovv enough that the fracture may be the result of piastic deformation. Be-sides the detaiis of such ductiie types. we also observed quasicieavage areas on fracture surfaces. most often on the very periphery of iarger. deeper funnel-type dimpies (Fig. 4, detaii B) and also as entirely independent shal-low regions. 6. CONCLUSIONS On the basis of the performed investigations. we have ascertained that the ioss of fracture ductility at slow-strain-rate tension of hydrogen-charged high-strength steel can be successfully used for the quantita-tive determination of the threshold stress intensity factor Kth (KHe>- It was estabiished that small concentrations of hydrogen (less than 1 ppm) in cathodically-charged steel have no substantial influence on the fracture toughness, as measured at conventional strain-rate (1 mm/min). Hovvever, the slow-strain-rate (0.1 mm/min) of hydrogen-charged high-strengthsteei vveakens the fracture ductility, which refiects the existence of the threshold stress intensity factor KTH. The parameter fi /Eq. (10), Table 1/ vvas determined by Inserting the KTH -values, calculated using the Hahn-Rosenfield correlation (7), into Gerberich s equation (6). We thus obtained a constant vaiue of about 4100 MPa, independent of the yield strength of steel, as requested by Gerberich s model for KTH. Microfractographic investigations of fracture sur-faces of highstrength hydrogen-charged steel tested at siovv-strain-rate indi- cate that. locally. the fracture is stili of ductile type. Despite the quasicleavage details, ductile ridges as a result of tearing as well as ductile dimpied areas vvere found in ali cases. The quasicleavage type of fracture on the periphery of larger. deeper and funnel-type dimpies proves that the grovvth and coales-cence of voids are also the consequence of the mechan-ism causing the separation of internal interfaces vvhere hydrogen is adsorbed. LITERATURA / REFERENCES 1. C.S. Kortovich in E.A. Steigervvald, Eng. Fract. Mech., 4, 637 (1972). 2. G.L. Hanna. A.R. Troiano in E.A. Steigervvald, Trans, of the ASM, 57.658-671 (1964). 3. Nevv hydrogen-embrittlement test. Advanced Materials and Processes inc., Metal Progress, 7, 10-11 (1988). 4. K. Farrell in A.G. Ouarrell, J. Iron Steel Inst., 202, 1002 (1964). 5. H. Morimoto in Y. Ashida. Trans. ISIJ, 23. B-352 (1983), 6. H. VVagenblast in H.A. VVriedt, Metali. Trans., 2, 1393-1397 (1971). 7. J.OM. Bockris, W. Beck, M.A. Genshavv. P.K. Subramanyan in F.S. VVilliams. Acta metali,, 19, 1209-1218 (1971). 8. H. Peisl v Hydrogen in Metals, Topics in Applied Physics, Ed.: G. Alefeld and J. Volkl. vol. 28, 53-74 (1978). 9. L.C.M. Li, R.A. Oriani in L.S. Darken. Z. Phzsik. Chem., 49, 271 (1966). 10. W.W. Gerberich v Effect of Hydrogen in high-strength and martensitic steels, Hydrogen in Metals. ASM-Ohio, 115 (1974), kot tudi: W.W. Gerberich in S.T. Chen, Metali. Trans., 6A. 271 (1975). 11. C.D. Kim in A.W. Loginow, Corrosion, 24, 313 (1968). 12. G.T. Hahn in A.R. Rosenfield, Sources of Fracture Toughness: The Relation betvveen KIC and the Ordinary Tensile Properties of Metals, Applications Related Phenomena in Titanium Alloys, ASTM STP 432, 5-32, Philadelphia, (1968). 13. G.T. Hahn in A.R. Rosenfield, Metali. Trans., 6A 653-668 (1975). 14. G.T. Hahn in A.R. Rosenfield, Trans.. ASM, 59, 909 (1966). 15. M.B. VVhiteman in A.R. Troiano, Corrosion, 21, 53-56 (1965). 16. K. Nakano, M. Kanao in T. Aoki, Trans, of National Research, Institute for Metals, 29, No. 2, 34-43 (1987). 17. R. Garber. M. Bernstein in A.W. Thompson, Ser Metali 10 341-345 (1976). 18. A.S. Argon in J. Im, Metali. Trans., 6A, 839-851 (1975). 19. J.R. Rice in D.M. Tracey, J. Mech. Phys. Solids 17 201-217 (1969). 20. D.M. Tracey, Eng. Fract. Mech., 3, 301-315 (1971). 21. A.S. Argon, J. Im in R. Safoglu, Metali. Trans., 6A, 825-838 (1975). 22. C.D. Beachem, Metali. Trans., 3, 437-451 (1972). 23. T.D. Lee, T. Goldenberg in J.P. Hirth, Metali. Trans 10A 199-208 (1979). 24. In-Gyn Park in A.W. Thompson, Metali. Trans., 21A 465-477 (1990). 25. D. Kwon in R.J. Asaro, Metali. Trans., 21 A, 117-134 (1990). 26. H. Cialone in R.J. Asaro, Metali. Trans., 10A 367-375 (1979). PRIKAZ Izdelava jekel v elektro obločni peči, sekundarna rafinacija v vakuumski napravi, kontinuirno vlivanje jekla, vlivanje jekla v kokile, vlivanje odlitkov v livarni, valjanja gredic, slabov in predtrakov na valjalnem stroju bluming. valjanje žice in profilov, valjanje debele pločevine PROIZVODNJE Toplo valjanje trakov na valjalnem stroju (štekel). hladno vlečenje žice, hladno vlečenje profilov, hladno valjanje trakov, proizvodnja žebljev, proizvodnja dodajnih materialov, izdelava hladno oblikovanih profilov, izdelava vratnih podbojev SLOVENSKE ŽELEZARNE P 64270 Jesenice. Cesta železar|ev 8. telefon: (064) 81 231. 81 341. 81 441 teleks: 34526 ZELJSN. Jugoslavija, telegram: Železarna Jesenice V proizvodnem programu so naslednji izdelki: gredice, toplo valjana debela, srednja, in tanka pločevina. hladno valjana pločevina in trakovi toplo valjana žica, hladno vlečena žica. hladno vlečeno, luščeno in brušeno paličasto jeklo, hladno oblikovani profili, kovinski vratni podboji, dodajni materiali za varjenje, žeblji, jekleni ulitki, tehnični plini Poleg navedenih izdelkov pa nudimo tudi storitve: valjanje v pločevino ali trak. vlečen|e v žico ali paličasto jeklo, toplotne obdelave, raziskave oziroma meritve lastnosti jekla računalniške obdelave psihološke sociološke in ekološke študije tehnološki inženiring Kompjutorska simulacija skručivanja odijevaka kompleksne geometrije Computer Simulation of Solidification of Compiex Geometry Castings V, Grozdanič*1, J. Črnko*1 1. UVOD Pri proizvodnji odijevaka složene geometrije od velike su pomoči modeli kojima se simulira njihovo skručivanja jer je več za kompjutorskim terminalom moguče odrediti tok njihova skručivanja, vrijeme skručivanja, mjesta moguče pojave defekta u odljevcima, te utjecati na njihov dizajn kako bi se proizvedli zdravi odljevci. Medutim. raznolikost i kompleksnost oblika odijevaka umnogome otežavaju simulaciju skručivanja jer je potrebno primije-niti kompleksan matematički aparat da bi se proces opi-sao matematičkim modelom koji uz osnovnu diferencijal-nu jednadžbu provodenja topline sadrži početne i gra-nične uvjete. Diferencijalna jednadžba rješava se nume-rički primjenom metode konačne razlike ili konačnog elementa. Metoda konačne razlike može biti eksplicitna i implicitna a po svojoj matematičkoj formulaciji nešto je jednostavnija od metode konačnog elementa, što je uv-jetovalo da se ta metoda prije koristila pri rješavanju kru-čivanja odijevaka. Ne ulazeči u prednosti ili nedostatke navedenih metoda opčenito je prihvačeno da se kod modeliranja skručivanja odijevaka manje složene geometrije za numeričko rješavanje parcijalne diferencijalne jednadžbe koristi metoda konačne razlike. Primjer od-Ijevka relativno složene geometrije predstavlja kučište ventila (slika 1), čiji matematički model skručivanja sadrži numeričko rješenje parcijalne diferencijalne jednadžbe implicitnom metodom promjenljivog smjera. 2. MATEMATIČKI MODEL Pri operacionalizaciji matematičkog modela potrebno je riješiti parcijalnu diferencijalnu jednadžbu provodenja topline koja odgovara geometrijskom uzoru kučišta ventila (slika 1)m: SJ= a (52T 15T j 82T 8t 8r2 rSr 8z2 (1) Kako u horizontalnoj osi simetrije kučišta ventila vri-jedi da je r = 0, jednadžbu (1) potrebno je modificirati pomoču L'Hospitalovog pravila nakon čega se dobije diferencijalna jednadžba slijedečeg oblika: 8t 8r2 Sz2 (2) Početni uvjeti. Temperatura kalupa i njegove vanjske strane jednaka je Ts, dok je temperatura metala u kalupu jednaka temperaturi lijevanja TL. Početna temperatura na Institut za metalurgiju, Sisak 1. INTRODUCTION Computer simulation of solidification is very helpful for castings of compiex geometry since it enables reii-able determination of the course of solidification. the tirne required for complete solidification and the points of potential casting defects. The simulation can also con t rib u te to improve casting design in order to assure sound castings. However, the simulation is very difficult in the čase of complex gecmetrical shape because it re-guires the elaboration of a sophisticated mathematical model vvhich beside basic differentiai equation of heat fiow must include initiai and boundary conditions also. 4 ioo Initial conditions. Mold temperature and temperature of its outer side is equai to Ts vvhereas the temperature of metal is equai to casting temperature TL. Initial temperature at the mold/casting boundary interface can be obtained by solving Fourier's differential equation for heat flow through the contact area of two semiinfinite medias: TL-TS T,= Ts+- 1+- (3) Derivation of eq. (3) is given in Appendix 2. Boundary conditions. Outer mold surface maintains constant temperature Ts. On contact mold/metal, metal/ core and moid/core area there is a continuous heat fiovv for vvhich boundary condition of the fourth sort holds: 8Tm _ k 8Tk Sn k8n STm_kSTi " 8n 8n k 8n 'Sn (4) (5) (6) Thermophysical properties of material, it has been asummed that thermal properties of mold, metat and core are temperature dependent'2'. (Appendix 3). 3. IMPLICIT METHOD OF VARIABLE DIRECTION Differential heat flow equations (1) and (2) vvith corresponding initial and boundary conditions have been numerically solved by impiicit method of variabie direction'3'. The method utiiize division of tirne interval into two steps. In the first half of tirne interval the equa-tion is solved implicitly for z and explicitly for r direction. The procedure is reversed in the second half of tirne interval. — za drugi At/2 vrijedi OTI1 + 1 OTn + 1 Tn + 1/2 o Tn + 1/2 , xn + 1/2 2 f | i, 2 ~ ^ 1 i, 1 j 1 i-1, 1 ~ ^ 'i.1 + 'i+1. 1 (Ar)2 (Az)2 b,v1 + c1v2 a2v, + b2v2 + c2v3 a3v2 + b3v3 + c3v4 + bivl + civl + 1 = di = d2 = d3 = dj (11) Vn = Yn v( = y, - i = N-1, N-2, ..., 1 (12) (13) gdje se p i y računaju iz rekurzivnih formula Pi = b, (14) ri = d,/p, (15) P, = b,--3-Pm, =2,3.....N ci-1 dnzMhl, i = 2,3,...,N P, (16) (17) Consequently, for differential equation (1) and first half of tirne interval At/2 we have: Tn _P Tn -+- T" Tn _T" ' i. H * ' i. n ''.;+' ' i. H (Ar)2 2 /- Ar+ , Tn +1 -rn + 1/2 1 11 (10) a, i,„ At/2 Primjenom implicitne metode promjenljivog smjera dobije se sistem simultanih linearnih algebarskih jed- nadžbi, čije su nepoznanice v,, v2.....vn, a koje imaju tri- dijagonalni oblik: + T"+ »/■?_OTn+ 1/2-L T"+ 1/2 * T"+ T" ' i-1. j i. j + '/+>./__1 ' i. i —I i. i ai j. n At/2 (Az)2 VVhereas for the second t/2 we obtain: an-1vn-2 + vn_! + cn_,vn =dN_1 aNvN_, + bNvN = dN Posebno efikasan algoritam za rješavanje tridijago-nalnog sistema jednadžbi je: Tn+ 1_p Tn+ 1 , -rn+ 1 *rn+ 1 _-rn+ 1 ' i. j-1 i j "r,i.y+r 'i. j+1 (Ar)2 2 rjAr jn+ i/2_2 Tn+1/2 + Tn+ 1/2 i Tn+1/2_Tn+1/2 | ' ' i_'_[__■'' 1 i___'__i_i__' ! /o\ (Azf ~ ai.i. n At/2 Numericai solution of the differential equation (2) of heat flovv for first At/2 is: 22TJ£-2TJi T°X\/2-2T°y/2+TX;/2 (Ar)2 + (Azf 4 T"+ 1'2 T" i 1 i. 1 ~ ' 1. 1 a>. 1. n At/2 (9) and for second At/2: 2 £77+'-2 77+' 77_t f - 2 T?+ 1/2 + 77+,"? (Ar)2 + ~(Alj2 = 7-n+ 1 Tn+ 1/2 t ' i. 1 ~ ' 1.1 a, m At/2 (10) U Dodatku 4 navedeni su tridijagonalni koeficijenti koji daju algoritam procesa skručivanja kučišta ventila u pješčanom kalupu. 4. DIJAGRAM TOKA Na temelju prikazanog algoritma napisan je program u programskom jeziku ASCII FORTRAN koji je riješen na računalu SPERRY 1100. Detaljan dijagram toka prikazan je na slici 2. Osnovna karakteristika programa je da se koriste dvije matrice temperatura T i T*. Prva matrica sadrži temperature na početku i kraju vremenskog koraka, a druga matrica sadrži temperature na kraju prvog At/2. Na početku programa pridaju se početne vrijed-nosti pojedinim varijablama i konstantama. Potprogra-mom TYP pridaju se početne vrijednosti temperature pojedinim mrežnim točkama, a takoder se tipiziraju sve točke u kalupu, odljevku, jezgri i na njihovim medusob-nim granicama. Nakon ispisa početne temperature ras-podjele pomoču potprograma ISPIS1, sistem algevarskih tridijagonalnih jednadžbi rješava se prvo redak po redak (potprogram RED), a zatim stupac po stupac (potpro-gram STUP). Rezultati se periodički ispisuju po cijeloj geometriji odljevka, kalupa i jezgre (potprogram ISPIS1) ili samo po geometriji odljevka (potprogram ISPIS2) do unaprijed zadanog vremena tmax. 5. DISKUSIJA REZULTATA Simulacija skručivanja ventila od čeličnog lijeva s oko 0,25% C u pješčanom kalupu provedena je uz prostorni korak Az = Ar = 1 cm i vremenski korak At = 10 s do vre- The use of impiicit method of variable direction results in a system of simultaneous iinear algebraic equa-tions vvith variables v1: v2...vn of tridiagonal form: bM+CtV? a2v,+ b2v2+ c2v3 a3v2+ b3v3+ c3v4 ayM+ b,Vi+ c,vi+1 = d, = d2 = d3 = d, (11) an-1v n-2 + bN_1VN_1+ , l/N = dN_ j aNvN_,+ bNvN = dN Specially efficient algoritm for the solving of tridiagonal system of eguations is: vn=Yn (12) v(= y, - BlZ+i, /= n-2, .... 1 (13) vvhere /5 and 8 are calculated from recursive formulas Pi= b, (14) Yi— d,/p, (15) P<=b,--— Bh1, =2,3,...,N (16) c 1—1 Pi Tridiagonal coefficients vvhich give algorithm of valve housing solidification in sand mould are presented in Appendix 4. Slika 2 Dijagram toka. Fig. 2 Fiow diagram. mena tmax = 200 s. Temperatura lijevanja bila je 1580° C a početna temperatura kalupa i jezgre 25° C. Na temelju sukcesivnih temperaturnih ispisa za pojedine mrežne točke dobije se vrijeme skručivanja desnog toplinskog centra od 136 s i lijevog toplinskog centra 181 s, što ujedno predstavlja vrijeme skručivanja cijelog odljevka. Na temelju pomicanja izosolidusa (slika 3) može se zaključiti da su mjesta moguče pojave defekta (lunkera) upravo mjesta gdje završava skručivanje toplinskih cen-tara, odnosno u bližini unutarnjeg ugla odljevka. Iz slike 3 uočava se da jezgra bolje odvodi toplinu od kalupnog materijala jer su izosolidusi pomaknuti više prema periferiji odljevka. Točnost simulacije skručivanja čeličnog ventila u pješčanom kalupu ograničena je s tri aspekta: predpostavkama uvedenim pri definiranju matematičkog mode- 4. FLOW DIAGRAM Based on the presented algorithm a computer program was written in ASCII FORTRAN and sotved on SPERRY 1100 computer. A detailed ftow diagram is seen on Fig. 2. The main feature of the program is its use of two temperature matrixes name/y T and T*. First matrix contains temperatures at the start and end of tirne step. The other contains temperature at the end of first At/2. Initial values are assigned to program var-iables and constants. Program module TYP provides for initial values of temperatures of particular net points as well as for standardization of ali points in mold. casting. core and theirs boundary interfaces. Module ISPIS1 prints out initial temperature distribution. The system of tridiagonai equations is then sotved firstiy row by row (module RED) and then cotumn by column (module STUP). Resuits are periodicaily printed over the whole geometry of casting, mold and core (module ISPIS1) or over the casting geometry oniy (module ISPIS2) untill the prescribed tirne tmax. 5. DISCUSSION Simulation of the soiidification of 0.25% C steel valve housing casted in sand mold is carried out by space step Az=Ar = 1 cm and tirne step Af= 10 s till tmax=200s. Casting temperature vvas 158CPC vvhereas the initial temperature of mold and core vvas 25> C. On the basis of successive temperatures print out for particular net points, soiidification tirne 136 s for the right and 181 s for the teft heat centre is obtained. The later value is also soiidification tirne for entire casting. By isosolidus shift as seen in Fig. 3 it can be conciuded that potentiai defect sites are obviously the points of final soiidification of heat centres i.e. in the vicinity of inner corner of the casting. From fig. 3 it can be seen that cooling rate of the core is higher than that of the mold since isosolidus curves are shifted outvvard. The accuracy of the simulation is iimited depending on the assumptions utiiized in mathematical model, method of numerical analysis and the utiiized values of thermophysical material properties. Severa! assumptions have been used in the e/aboration of the mathematical modeI. The most important are: the heat transfer rate is comptete. the casting temperature is equal to initial temperature of metal in the mold and the mold/cast-ing interfacial thermal contact is perfect. The first assumptions restrains the analysis to the mold-casting- Slika 3 Napredovanje izosolidusa (1449°C) u odljevku za vremena 20, 80 i 140 s. Fig. 3 Progress of isosolidus (1449° C) after 20. 80 and 140 s. la, primjenjenoj metodi numeričke analize i korištenim vrijednostima toplofizičkih svojstava materijala. Pri postavljanju matematičkog modela krenulo se od više pretpostavki od kojih su najvažnije pretpostavka o pot-punom provodenju topline, pretpostavka da je temperatura lijevanja jednaka početnoj temperaturi metala u kalupu i pretpostavka o savršenom toplinskom kontaktu na graničnoj plohi kalupa i odljevka. Prva pretpostavka og-raničuje razmatranje skručivanja na sistem kalup-odlje-vak-jezgra u kojem se toplina prenosi samo provode-njem, što znači da se ne razmatraju parcijalni procesi prijenosa topline vezani uz vlagu u kalupu i jezgri. Druga pretpostavka predstavlja pojednostavljenje uvedeno da se izbjegne kompleksno razmatranje protjecanja metala kroz uljevni sistem i kalupnu šupljinu povezano s prijela-zom topline. Pretpostavka o savršenom toplinskom kontaktu na graničnoj plohi odljevka i kalupa prihvatljiva je iz razloga što se na graničnoj plohi tek djelomično javljaju plinski zazori, pa se pri matematičkoj formulaciji uzima da vrijedi granični uvjet četvrte vrste. Parcijalna diferencialna jednadžba provodenja topline riješena je nume-ričkom metodom konačne razlike - implicitnom metodom promjenljivog smjera koja je odabrana iz razloga što ima veliku točnost pri aproksimaciji i prostora i vremena. Metoda je drugog reda s obzirom na diskretizaci-ju prostora i vremena. Nedovoljno poznavanje toplofizičkih svojstava materijala posebno pri visokim temperaturama značajno utječe na simulaciju skručivanja. To se posebno odnosi na toplinska svojstva kalupnog materijala i jezgre, koja je moguče odrediti samo eksperimentalnim putem a pri visokim temperaturama pokazuje širok dijapazon rasipanja. 6. ZAKLJUČAK U radu je provedena numerička simulacija skručivanja kučišta ventila od niskougljičnog čeličnog lijeva na formuliranom matematičkom modelu. Modelni sistem je kompleksan jer se sastoji od triju materijala: kalupa, jezgre i odljevka relativno složene geometrije. Matematički model skručivanja odljevka postavljen je uz pretpostav-ku da u sistemu postoji samo provodenje topline, što predstavlja fizikalno realnu pretpostavku. Diferencijalna jednadžba provodenja topline koja odgovara geometrij-skom uzoru kučišta ventila odgovarajuče je modificirana i riješena numerički implicitnom metodom promjenljivog smjera s time da je uzeta u obzir temperaturna ovisnost toplofizičkih svojstava pojedinih materijala. Na temelju dobivenog algoritma napisan je program u program-skom jeziku ASCII FORTRAN za računalo SPERRY 1100. Na temelju simulacije konstatirano je vrijeme skručivanja od 181 s, a na temelju pomicanja izosolidusa moguče je odrediti tok skurčivanja te mjesta moguče pojave defekta u odljevku. Dodatak 1 Popis oznaka korištenih u radu a — temperaturna vodljivost materijala a„ b„ cs, dt — koeficijenti uz nepoznanice u tridijagonal- nom sistemu algebarskih jednadžbi cp — specifična toplina pri konstantnom tlaku k — toplinska vodljivost materijala n — normala r — prostorna koordinata t — vrijeme T — temperatura v, — nepoznanica u sistemu simultanih algebarskih jednadžbi z — prostorna koordinata core system vvith heat conduction oniy, i.e. partia/ heat flovvs associated vvith mold and core moisture are not considered. The second assumptions is simplification in-troduced to avoid complex consideration of metal flovv through gate system and mold cavity and matching heat transfer. The assumption of perfect thermal contact on the interface is acceptable since on/y partial appearance of gaseous ciearance therefore in mathematical formula-tion the boundary condition of fourth kind is usually taken as valid. Partial differential equation for heat flovv is solved by numerica/ method of finite difference - the im-plicit method of variable direction vvhich vvas chosen due to its high accuracy at approximation of both time and space. The method is of the second order in respect to discretion of time and space. Insufficient knovvledge of thermophysical properties of material specially at high temperatures has a strong influence on simulation of the solidification. It holds specia//y in respect to thermal properties of mold and core material vvhich can be determined only by experiment. Moreover vaiues for thermal properties at higher temperatures shovv considerable dissipation. 6. CONCL USIONS Numerica/ simulation of solidification of lovv carbon steel casting (valve housing) has been carried out on the basis of a suitable mathematical model. The complex model system is composed of three materials: mold, core and casting of comparatively complicated geome-try. Mathematical model of solidification has been elabo-rated assuming thermal conduction as only heat flovv in the system vvhich is considered as a physically real assumption. Differential equation for heat flovv suited to the geometry of valve housing has been modified and numericaiiy solved by the use of implicit method of variable direction. Temperature dependence of thermophy-sical material properties has been taken into account. Based on the obtained algorithm a computer program vvritten in ASCII FORTRAN for SPERRY 1100 computer ivas used for simulation of the solidification. It has been determined that complete solidification takes 181 sec-onds. The progress of solidification as vveil as hot spots i.e., sites of potential shrinkage cavities can be determined by shift of isosolidus. Appendix 1 Abbreviat/ons used: a - temperature conductivity a j, bi: C/, d, coefficients adjoining to unknovvns in tridia-gonal system of algebraic equations, cp - specific heat at constant pressure, k - thermal conductivity, n - verticai direction r - space coordinate t - time T - temperature v, - unknovvn in system of simultaneous algebraic equa-tions z - space coordinate Indices i - space coordinate z, value for boundary mold/metal surface j - space coordinate r, core k - mold L - casting m - metal s - room, interfacial Indeksi i — prostorna koordinata z, vrijednost na graničnoj plohi kalup-metal j — prostorna koordinata r, jezgra k — kalup L — lijevanje m — metal s — sobno, površinski Dodatak 2 Pri izvodenju jednadžbe za početnu temperaturnu raspodjelu na dodirnoj plohi kalupa i metala potrebno je riješiti parcijalnu diferencijalnu jednadžbu provodenja topline s odgovarjajučim početnim i graničnim uvjetima 8T_ 8t 3 Sx2 T(x, O) =TS T(0,X) = T, I T(x, t) I 1499° C k = 25,96 1499° C > T > 1449° C k = 207,54-0,12114T 1499° C > T > 893° C k = 26,6+ 0,00374 T 893° C > T k = 50,31 - 0,0225 T Specifična toplina, J/kgK -kk£f*)x=0=-km(^)x=0 (29) ox 5x By including proper temperature gradients: lc - n)* = 0 (29) 8x 8x Uvrštavanjem odgovarajučih gradijenata temperature dobije se (30) (31) T i-Ts k^nakt = kr TL-T, (30) Finatiy, initiai temperature distribution on the boun-dary mold/metai surface is obtained: t^T.+ML Na graničnoj plohi u kontaktu su dva polubeskonač-na medija (kalup i metal) pri čemu vrijedi granični uvjet četvrte vrste: |/nakt ' l/xcamt Sredivanjem jednadžbe (30) dobije se početna temperaturna raspodjela na graničnoj plohi izmedu kalupa i odljevka /+*kl/s (31) km" ak Appendix 3 Thermo-physical properties of the materials a) Low carbon (0.25% C) castad steel Thermal conductivity, W/mK T> 149SPC k= 25,96 1499P C> T> 144SP C k= 207,54-0,12114 T 14990 C> T> 893P C k= 26,6+ 0,00374 T 893° C> T k= 50,31-0,0225 T Specific heat, J/kgK T> 1499"C cp= 879,2 149SP C> T> 1474° C cp= 652273,5- 434,585 T 14740 c> T> 14490 C cp= 436,258 T- 631251,9 1449PC> T> 982°C cp= 421,36+0,28712 T 982°C> T> 704°C cp= 1502,8-0,81391 T 704° C> T> 427° C cp= 143,76+ 1,11535 T 427° C> T cp= 458,86+ 0,37681 T b) Core Thermal conductivity, W/mK k= 3.246x 10~6 T2 - 3.894x 10~3T+ 3.052 Temperature diffusivity, m2/s a= 1.689x 10~'2 T2 - 2.174x 10~9T+ 1.785x 10~6 c) Mold Thermal conductivity, W/mK k= 3.937x 10~6T2 - 3.57x 1Cr3T+ 3.421 Temperature diffusivity, m2/s a= 1.957x 10~12T2 - 1.8x 10~9T+ 1.913x 10~6 Appendix 4 Constants vvhich appear in tridiagonal coefficients a At Pr- 2 (Ar)2 aAt 4tjAr P3 = P i P2 p4=Pi+P2 P5 = At (kA+ kB) 2 c (Ar)2 ^ At (kA+ kB) 4 c r/Ar c=!lA + ! 1499° C c0 = 879,2 1499° C > T > 1474° C 1474° C > T > 1449° C 1449° C > T > 982° C 982° C > T > 704° C 704° C > T > 427° C 427° C > T cp = 652273,5-434,585 T cp = 436,258 T — 631251,9 cp = 421,36 + 0,28712 T cp= 1502,8-0,81391 T cp=143,76 + 1,11535 T cp = 458,86+ 0,37681 T b) Jezgra Toplinska vodljivost, W/mK k = 3,246 • 10-6 T2—3,894 ■ 10~3T +3,052 Temperaturna vodljivost, m2/s a =1,689- 10~12T2—2,174 ■ 10"9 T+1,785- 10"6 c) Kalup Toplinska vodljivost, W/mK k = 3,937 ■ 10"6 T2—3,57 • 10"3 T+ 3,421 Temperaturna vodljivost, m2/s a =1,957 ■ 10"12T2— 1,8 • 10"9T+ 1,913 • 10"6 Dodatak 4 Konstante koje se javljaju u tridijagonalnim koefici-jentima Pi = p2 = a At 2 (Ar)2 a At 4r.Ar P3 = Pi — p2 P4 = P, + P2 Ps = Pe = At (kA + kB) 2 c (Ar)2 At (kA + kB) 4 c r,Ar c = kA+kg aA 3B Qi : q2 = q3 = aAt 2 (Az)2 KAAt c (Az)2 kAAt Q4 = q5=- q6 = c (Az)2 At (kA + kB) 2 c (Az)2 KAsa40D' c (Ar)2 kBAt c (Ar)2 Tridijagonalni koeficienti 1. Točka (i,j) u kalupu, metalu odnosno jezgri — prvi At/2: q2= aAt 2(Azf KAAt c (Az)2 kAAt q4= Qs=- <7s= c (Az)2 At (kA + kB) 2 c (Az)2 K.sa40D' c (Ar)2 kgAt c (Ar)2 Tridiagonal coefficients 1. Point (i,j) in the mold. metal or core; first At/2: a,= c,= —q, b,= 1+2 q, d,= p3 T:h+ (1-2Pl) 77,+ p477/+, (32) — second At/2: a~ —p3 b,= 1+2 p, §1 q,nX!/!+ (1—2qJ 1^+ q, r+lf (33) 2. Point (i,j) on the boundary surface para/le/ to r axis separating the material A (left) and B (right). - first At/2: a~ -q2 b,= 1+ q2+ q3 c,= q3 d,= (P-rPe) T"h+ (1-2p5) rij+ (p5+ p6) T°J+, (34) T,- '-1.J AZ Ti,j T'*1.j Z Tu-, Slika 5 Mrežna točka (i,j) na vertikalnoj graničnoj plohi. Fig. 5 Net point (i,j) on verticai boundary surface. r < Ar Az B Ti-i.J Ti.j Ti«1,j Z c Ti.j-1 A Slika 6 Mrežna točka (i,j) na horizontalnoj graničnoj plohi. Fig. 6 Net point (i,j) on horizontai boundary surface. ai = Ci = —q, b, = 1+2q1 di = p3 Tni.j-i + (1 —2p,) T" j + p4Ti j + 1 — drugi At/2: a, = — P3 b,= 1+2 Pl d) = q, 1/2 + (1 -2q1) T/1/2 + q, TF+1f,j — second At/2: a,= -(p5-p6) bj= 1+2 p5 cj— -(Ps+Pe) dj= q2 T?_y/i+ (1~q2~q3) 7?/ q3 77/,f (35) 3. Point (i,j) on the boundary surface parai/ei to z axis separating the material A (down) and B (up). — first At/2: a- C/= - q4 b- 1 + 2q4 d,= (q5- q6) r>j_ ,+ (1-2p5) Tl+ (q6+ p6) T»J+ , (36) — second At/2: ai= Pe~ Qs b,= 1 + 2ps Cj=- (q6+p6) d,= q4 r/,'f+ (1-2qjri*1/2 + q4 Tftjf (37) 4. Point (i,1) out of boundary surface; first At/2: a,= c,= - q, b;= 1+2q, (1— 4p,) T",+ 4p1 T"g (38) — second At/2: bj= 1+4p, Cj= 4p, d,= q, 7/,'f + (1-2q,) T* '/2+ q, 77/,ff (39) (32) 5. Point (i, 1) on the boundary surface which separate material A (ieft) and B (right) — first At/2 : a~ - q2 b-2q4+ 1 dl=(iq-4p5)T"n+4psT"l2 (40) (33) 2. Točka (i,j) na graničnoj plohi paralelno r osi koja odvaja materijal A (lijevo) i B (desno) — prvi At/2: a, = -q2 b;=1 +q2 + q3 c, = q3 d, = (p5-p6) T" H + (1-2ps) T", + (p5 + p6) "^.h (34) — drugi At/2: a,= — (Ps—Ps) b,= 1+2 p5 c,= —(p5 + p6) d, = q2Tj/, + (1-q2-q3) T/1'2 + q3TT/lf — second At/2: b~ 4p5+ 1 d= q2 7T/f + (1-2qJ '*+ q3 77/,'f (41) (35) 3. Točka (i, j) na graničnoj plohi paralelno z osi koja dijeli materijal A (dolje) i B (gore) — prvi At/2: a, = c,= -q4 b, = 1 + 2q4 d, = (q5 - q6) T7(_, + (1 - 2p5) T" + (q6 + p6) , (36) — drugi At/2: aj = P6-q5 b, = 1 + 2p5 c,= - (q6+p6) -i cl4 T".*,1 j + (1 -2q4)T"j" 1 i,2 Ti-i.i ' i,1 Slika 7 Mrežna točka (i,1) na vertikalnoj graničnoj plohi. Fig. 7 Net point (i. 1) on vertical boundary surface. d, = q4 TL+,1;2 + (1 -2a4)T",+ 1/2 + q4 T/,"2 (37) 4. Točka (i, 1) koja nije na granici — prvi At/2: a, = c,= bi = 1+2q, di= (1 —4p,) T", +4p, T,n2 (38) — drugi At/2: bj = 1+4Pl c, = 4pi d, = q, T[!+1/2 + (1 —2qi) T"^1/2 + q1 T?+1/2 (39) 5. Točka (i, 1) na graničnoj plohi koja dijeli material A (li-jevo) i B (desno) LITERATURA / REFERENCES 1. E. B. G. Eckert, R. M. Drake, Analysis of Heat and Mass Transfer, McGravv-Hill Kogakusha, Tokyo, 1972. 2. R. D. Pehlke, A. Jeyarajan, H. Wada, Summary of Thermal Properties for Casting Alloys and Mold Materials, University of Michigan, Ann Arbor, 1982. 3. J. Douglas, H. H. Rachford, Trans. Amer. Math. Soc. 82 (1956), 421. — prvi At/2: a, = - q2 b, = 2q4+1 d,= (1-4p5)^1 + 4p5T[,2 (40) — drugi At/2: b, = 4p5+1 Cj= — 4p5 d, = q2 ^«+(1 -2q4) Ttf1* + q3 (41) Vsebina — I. Rak, V. Gliha, F. Vodopivec, M. Tavčar Vpliv varilne tehnologije in izbire dodajnega materiala na lomne lastnosti EPP zvarnega spoja na nizko ogljičnem fin-ozrnatem jeklu Železarski zbornik 25 (1991) 4s 117-125 Žilavost, lom, M/A-faza, COD meritev, nap. žarjenje V sestavku so predstavljene lastnosti raztaljenega dela zvarnega spoja narejenega v laboratoriju pod dejanskimi pogoji na nizkoogljičnem finozrnatem jeklu kvalitete Niomoi 390. Preiz-kave prikazujejo nizko lomno žilavost in pojavo lokalnih krhkih področij-LKP zaradi tvorbe neugodnih M/A strukturnih faz, ki jih kljub drobnozrnatosti povzročijo dodatki Ti-B. Termično sproščanje napetosti je še dodatno znižalo žilavost zaradi izločevanja cementita iz M/A strukturnih faz na meji z intragranularnim feri-tom. Contents I. Rak, V. Gliha, F. Vodopivec, M. Tavčar The Influence of Welding Technology and Welding Material Seleetion on Fracture Properties of Submerged Are Welded, Lovv Carbon, Finegrained Steel Plate Železarski zbornik 25 (1991) 4P117—125 Toughness, Fracture, LBZ, M/A-constituents, COD-measure- ment, Stress relieving Properties of aii-vveid metal made in the laboratory under the actuai fieid conditions on lovv carbon steel grade Niomoi 390 are presented in this paper. The examinations shovved re-duced impact toughness and appearance of LBZ caused by the formation of the unfavourabie M/A constituents in spite of the presence of fine grain formed by Ti-B addition. Thermal stress relieving due to precipitation of cementite on M/A constituents and ferrite interfaces lovvered the toughness considerably. Prof. dr. Jože Rodič, dipl.inž. - MIL-PP d. o. o., Ljubljana Kobaltove zlitine v lesni industriji Železarski zbornik 25 (1991) 4s 127—137 Kobaltove zlitine, lesna industrija Izboljšanje rezne sposobnosti in vzdržljivosti žag v lesni industriji, ki se doseže z navarjanjem kobaltovih zlitin - stelitov* na rezalni del zob je že dolgo poznano, vendar je šele v zadnjih letih razvoj avtomatskih strojev za "stelitiranje" vseh vrst žag omogočil industrijsko uporabo tega postopka. Opisan je razvoj tehnologije, strojev in dodajnih materialov za stelitiranje. Posebna pozornost je namenjena rezultatom primerjalnih raziskav, izkušnjam v praksi ter nadaljnim usmeritvam razvoja za optimiranje materialov in tehnologije. Stelitirane žage imajo številne pomembne prednosti in izpodrivajo iz uporabe konvencionalne, pa tudi trdokovinske, predvsem pri rezanju svežega lesa in exotov. Tehnologija stelitiranja omogoča pomembno zmanjšanje proizvodnih stroškov ob istočasnem povečevanju produktivnosti in izboljšanju kakovosti. Prof. dr. Jože Rodič - MIL-PP d.o.o., Ljubljana Cobalt Base Alloys in Woodcutting lndustry Železarski zbornik 25 (1991) 4 P 127—137 cobalt base alloys, vvoodcutting industry Beneficiai effects of savv teeth tipping vvith cobalt base al-ioys - stellites for improving cutting ability and iifetime of savvs in wood cutting industry are well known more than two de-cades, but not before recent years the development of auto-matic stellite tipping machines enabied an industriai application of this important process. The development of technology, machines and additive ma-terials for stellite tipping is shortiy presented. Particular attention is dedicated to the results of compara-tive research, practicai experience and the aims of development for optimizing of materials and technology. The stellite tipped savvs shovv many advantages in application and therefore conventional and hard meta/ savvs will be re-placed especiaiiy for cutting fresh untreated wood and exotes. The technology of stellite tipping enabies considerabie tow-ering of production costs vvhile at the same time quality and produc-tivity are improved. B. Ule, F. Vodopivec, L. Vehovar in L. Kosec Faktor mejne intenzitete napetosti pri počasnem natezanju navodičenega jekla z visoko trdnostjo Železarski zbornik 25 (1991) 4 s 139—147 Metalurgija - vodik v jeklu - mehanske lastnosti - prelom jekla Na izgubo lomne duktilnosti močno vpliva zlasti vodik v jeklu, pri tem pa ne učinkuje zaznavno na napetost tečenja. Poslabšanje lomne duktilnosti je še posebej izrazito pri počasnem natezanju jekla, medtem ko je pri običajnih hitrostih deformacije manj izrazito. Male koncentracije vodika v jeklu z visoko trdnostjo zato ne vplivajo bistveno na lomno žilavost takšnega jekla, pač pa imajo za posledico pojavljanje faktorja mejne intenzitete napetosti KTH. Eksistenco tega faktorja pri počasnem natezanju navodičenega jekla smo določili z merjenjem konstante P v Gerberichovi enačbi. Konstanta (3 je ime/a od napetosti tečenja skoraj neodvisno vrednost, kar pomeni, da se faktor mejne intenzitete napetosti res lahko enostavno izračuna s pomočjo Hahn-Rosenfieldove korelacije. B. Ule. F. Vodopivec. L. Vehovar and L. Kosec Threshold Stress lntensity Factor at Slovv-Strain-Rate Tension of High-Strength Hydrogen-Charged Steel Železarski zbornik 25 (1991) 4 P 139-147 Metaiiurgy - Hydrogen in steel - Mechanical properties - Fracture of steel The decrease in fracture ductiiity is strongiy infiuenced by the presence of hydrogen in steel, although hydrogen does not es-sentially affect the yield strength. The deterioration of fracture ductility is especiai/y intensive at slovv-strain-rate tension, vvhereas at conventional crosshead speed it is less pronounced. Consequently, lovv concentrations of hydrogen in high-strength steel do not substantiaiiy affect its fracture toughness. but re-suit in the appearance of the threshold stress intensity factor Kth■ The exsistence of this factor at slovv-strain-rate tension of hydrogen-charged steel vvas proven by measuring the B constant in Gerberich's equation. This constant is practically inde-pendent of the yieid strength, vvhich means that the threshold stress intensity factor can be calculated simply vvith the heip of the Hahn-Rosenfieid correlation. Grozdanič V., J. Črnko Kompjutorska simulacija skručivanja odljevaka kompleksne geometrije Železarski zbornik 25 (1991) 4s 149—158 Numeriča simulacija, skručivanje, niskougljični čelični lijev, ku- čište ventila U radu je provedena simulacija skručivanja odljevka relativno kompleksne geometrije — kučišta ventila od niskougljičnog čeličnog lijeva na formuliranom matematičkom modelu. Postavljeni matematički model temelji se na fizikalno realnim pretpos-tavkama i riješen je numeričkom metodom konačne razlike — implicitnom metodom promjenljivog smjera. Simulacija skručivanja provedena je tako da je latentna toplina kristalizacije in-korporirana u jednadžbu za specifičnu toplinu metala s time da je uzeta temperaturna ovisnost toplofizičkih svojstava svih ma-terijala u sistemu kalup-odkljevak-jezgra. Simulacija skručivanja kučišta ventila, odnosno odljevka relativno složene geometrije koji je često primjer u Ijevaoničkoj praksi, omogučuje da se unaprijedi naše saznanje o procesu skručivanja takvih odljevaka te da se na suvremen i znanstveni način ukaže na mjesta moguče pojave defekta a da se pri tome ne odlije ni jedan od-Ijevak. Grozdanič V., J. Črnko Solidification Simulation of Castings of Compiex Geometry Železarski zbornik 25 (1991) 4 P 149—158 Numericai simulation, solidification, iow-carbon steel, valve housing The solidification simulation of the casting of reiativeiy com-plex geometry, iow-carbon steel valve housing, has been pre-sented in the paper. Mathematical model has been based on physicai reaiistic assumptions, and has been so/ved by finite difference method — implicit a/ternating directions method. Simulation of solidification has been performed so that the latent heat of solidification has been incorporated in the eguation for specific heat of meta/, and the temperature dependence of thermai properties of aH materials in the system mould-casting-core has been taken. The solidification simulation of valve housing, that is the casting of relativeiy compiex geometry which is a common example in foundry practice, enabies the improvement of our knovviedge about solidification process in these castings, and points in a modern and scientific way to the points vvhere defects occurrence is possible and in this way the casting is not necessary. Železarski zbornik, 25, 1991, 1-4 1. KRONOLOŠKO KAZALO Vehovar Leopold: Mehanizmi delovanja vodika v kovinah in vodikova krhkost................ŽZB 25 (1991) 1, 1-12 Vodopivec Franc: Topologija rasti rekristalizacijskih zrn v jeklu z 1,8% Si 0 3% Al in 0,02%C v razponu temperature 700—800°C . , . . . . ŽZB 25 (1991) 1, 13-20 Bolčina Marjan: Reševanje stacionarnega in nestacionarnega temperaturnega polja po metodi končnih elementov na PC računalnikih . . . . ŽZB 25 (1991) 1, 21—24 Risteski Ice B.: Primena ERROR funkcije u difuzionom hromiranju . , . . ŽZB 25 (1991) 1,25-28 Kmetič Dimitrij, J. Žvokelj, B. Ralič, M. Jakupovič, L. Jovanovski: Lastnosti s CaSi obdelanega konti litega jekla C4830 pri dinamičnih obremenitvah ............ŽZB 25 (1991) 2, 49-55 Mikec Darko, D. Finžgar, P. Sekloča, B. Glogovac, T. Kolenko: Rekonstrukcija koračne peči CUSTODIS v Valjarni žice in profilov . . . . . . ŽZB 25 (1991) 2,57-61 Kolenko Tomaž, M. Debelak, B. Glogovac: Ugotavljanje začetnega temperaturnega stanja vročih plošč pri zalaganju v potisno peč . . . . . , ŽZB 25 (1991) 2, 63 - 68 Koroušič Blaženko: Vpliv sestave žlindre na dezoksidacijo . . . , . . ŽZB 25 (1991) 3, 77-81 Jenko Monika, et al.: Študij segregacij Sb na površini jekel . . . ŽZB 25 (1991) 3, 83-87 Arzenšek Boris, et al.: Hladno preoblikovanje superferitnega jekla . . , . ŽZB 25 (1991) 3, 89-96 Torkar Matjaž, B. Šuštaršič: Mikrostrukturne značilnosti vodno atomiziranega prahu ........ŽZB 25 (1991) 3, 97-103 Todorovič Gojko, et al.: Uporaba odbruskov pri proizvodnji jekla , . . ŽZB 25 (1991) 3, 105-110 Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: Vpliv varilne tehnologije in izbire dodajnega materiala na lomne lastnosti EPP zvarnega spoja na nizko ogljičnem finozrnatem jeklu . . ŽZB 25 (1991) 4, 117—125 Rodič Jože: Kobaltove zlitine v lesni industriji . . . . . . ŽZB 25 (1991) 4, 127-137 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: Faktor mejne intenzitete napetosti pri počasnem natezanju navodičenega jekla z visoko trdnostjo ..............ŽZB 25 (1991) 4, 139-147 Grozdanič Vladimir, J. Črnko: Kompjutorska simulacija skručivanja odljevaka kompleksne geometrije , . . ŽZB 25 (1991) 4, 149—158 2. AVTORSKO KAZALO Arzenšek Boris, et al.: Hladno preoblikovanje superferitnega jekla . ŽZB 25 (1991) 3, 89—96 Bolčina Marjan: Reševanje stacionarnega in nestacionarnega temperaturnega polja po metodi končnih elementov na PC računalnikih . . ŽZB 25 (1991) 1,21-24 Kolenko Tomaž, M. Debelak, B. Glogovac: Ugotavljanje začetnega temperaturnega stanja vročih plošč pri zalaganju v potisno peč . . . . . . ŽZB 25 (1991) 2, 63 -68 Grozdanič Vladimir, J. Črnko: Kompjutorska simulacija skručivanja odljevaka kompleksne geometrije . . . ŽZB 25 (1991) 4,149—158 Jenko Monika, et al.: Študij segregacij Sb na površini jekel , . . . . . ŽZB 25 (1991) 3, 83 - 87 Kmetič Dimitrij, J. Žvokelj, B. Ralič, M. Jakupovič, L. Jovanovski: Lastnosti s CaSi obdelanega konti litega jekla C4830 pri dinamičnih obremenitvah ............ŽZB 25 (1991) 2, 49-55 Koroušič Blaženko: Vpliv sestave žlindre na dezoksidacijo . . . . . , ŽZB 25 (1991) 3, 77- 81 Mikec Darko, D. Finžgar, P. Sekloča, B. Glogovac, T. Kolenko: Rekonstrukcija koračne peči CUSTODIS v Valjarni žice in profilov . . . ŽZB 25 (1991) 2, 57 -61 Rak Inoslav, V. Gliha, F. Vodopivec, M. Tavčar: Vpliv varilne tehnologije in izbire dodajnega materiala na lomne lastnosti EPP zvarnega spoja na nizko ogljičnem finozrnatem jeklu . . . ŽZB 25 (1991) 4, 117—125 Risteski Ice B.: Primena ERROR funkcije u difuzionom hromiranju . . . . , ŽZB 25 (1991) 1,25-28 Rodič Jože: Kobaltove zlitine v lesni industriji . . . . . . ŽZB 25 (1991) 4, 127-137 Todorovič Gojko, et al.: Uporaba odbruskov pri proizvodnji jekla . . . . , ŽZB 25 (1991) 3, 105-110 Torkar Matjaž, B. Šuštaršič: Mikrostrukturne značilnosti vodno atomiziranega prahu ŽZB 25 (1991) 3, 97—103 Ule Boris, F. Vodopivec, L. Vehovar, L. Kosec: Faktor mejne intenzitete napetosti pri počasnem natezanju navodičenega jekla z visoko trdnostjo ..............ŽZB 25 (1991) 4, 139-147 Vehovar Leopold: Mehanizmi delovanja vodika v kovinah in vodikova krhkost.............ŽZB 25 (1991) 1, 1-12 Vodopivec Franc: Topologija rasti rekristalizacijskih zrn v jeklu z 1,8% Si, 0,3% Al in 0,02%C v razponu temperature 700—800°C . . . . . . ŽZB 25 (1991) 1, 13—20 SLOVENSKE ŽELEZARNE ŽELEZARNA RAVIME n.sol.o RAVNE NA KOROŠKEM SLOVENIA - YUGOSLAVIA Železarna Ravne kot proizvajalec kvalitetnih in plemenitih jekel nenehno razvija in izpopolnjuje tehnološke postopke s ciljem povečevanja finalizacije, kvalitete, avtomatizacije in humanizacije dela. Izgradnjo novih tehnoloških naprav v jeklarni, kovačnici, termični obdelavi in širjenje proizvodnje finalnih izdelkov je spremljal intenziven tehnološki razvoj podprt z uvedbo procesnih računalnikov, numerično krmilnih enot ter avtomatizacije. Računalniško vodenje procesa Jekleni valji za valjanje kovin Različna industrijska rezila iz plemenitega jekla SLOVENSKE ŽELEZARNE ŽELEZARNA ŠTORE ŠTORE PROIZVODNI PROGRAM Toplo valjani profili — kvalitetno in plemenito ogljikovo jeklo ter — kvalitetno in plemenito nizko legirano jeklo v okrogli, ploščati in kvadratni obliki, — specialni profili po načrtih Hladno oblikovani profili — vlečeno in brušeno jeklo v vseh kvalitetah v okrogli, ploščati in specialni izvedbi Livarski proizvodi — ulitki iz sive litine, — ulitki iz nodularne (KGR) litine, — kontinuirno liti profili, — litoželezni valji, — jeklarske kokile, — priklopna sedla, — mehanski sklopi, — strmoramenska platišča Industrijska oprema — industrijski gorilniki, industrijske peči za ogrevanje, žarenje itd., — indukcijske peči, — rekuperativna toplotna tehnika, — plinski oskrbovalni sistemi za ZP in zamenljive mešanice Vlečene palice kakovostnega jekla 64270 Jesenice, Cesta železarjev 8, teleks: 34526 ZELJNS, Jugoslavija telefon: (064) 81 231, 81 341, 81 441, telegram: Železarna Jesenice SLOVENSKE - ŽELEZARNE