Influence of Fracture Toughness on Vacuum Hardened HSS
Vpliv lomne žilavosti na vakuumsko toplotno obdelano hitrorezno jeklo
Leskovšek V.,1 B. Ule, A. Rodič, IMT Ljubljana
Fractures. macro-chipping and micro-chipping are ali effects by which cutting edges are de-stroyed. The ability of a steel to resist these phenomena is knovvn as its toughness HSS hovvev-er. possess an appreciable ductility, although the notched or even unnotched specimens tested m the pendulum test are not sensitive enough to discriminate betvveen high and low levels of toughness. Therefore, it becomes important to use a method of testing vvhich can detect small vari-ations m ductility. To establish the fracture toughness, the round-notched tensile specimens vvith a fatigue crack at the notch root ivas used. Fatiguing vvas done in as soft annealed condition After that. the vacuum heat treatment for the achievement ofoptimal vvorking properties vvas carried out and the fmal testing vvas performed. Our experiments confirm that the correlation based on the round-notched tension test can be successfully used to caiculate the critical fracture toughness On the basis of the above-mentioned experimental results. we vvere abie to compose a diagram vvhich simultaneously scoops the technological parameters of vacuum heat-treatment the mechanical properties and the micro structure of vacuum heat-treated HSS M2.
Key vvords: fine blanking tool. fracture toughness, hardness, vacuum heat treatment
Lomi, makrookruški in mikrookruški so vzrok propadanja rezilnih robov. Sposobnost jekla da se upira tem pojavom, pa je poznana kot žilavost. Hitrorezno jeklo ima upoštevanja vredno duktilnost četudi preizkušanci z zarezo ali celo celo brez zareze pri Charpyjevem preizkusu niso dovolj selektivni, da bi nam omogočali določitev krhke oz. žilave narave loma. Za krhke materiale med katere spada hitrorezno jeklo, je pomembno, da izberemo metodo preizkušanja, ki zazna že majhne spremembe duktilnosti jekla ter je selektivna in reproduktivna. Poleg standardnega načina merjenja lomne žilavosti na preizkušancih, ki so dovolj debeli, da je izpolnjen pogoj ravninskega deformacijskega stanja, uporabljamo tudi nestandardni način merjenja lomne žilavosti s cilindričnimi nateznimi preizkušanci z zarezo po obodu. Problemi pri ustvarjanju razpoke v korenu zareze, so nas navedli na idejo, da metodo za določevanje lomne žilavosti s pomočjo cilindričnih preizkušancev z zarezo po obodu modificiramo. Doseženi rezultati so pokazali, da je modificirana metoda tudi dovolj selektivna. Osnovni namen modifikacije je, ustvariti razpoko kontrolirane globine v korenu zareze na mehko žar jenih cilindričnih preizkušancih z zarezo po obodu Pred pulz i rane cilindrične preizkušance zatem vakuumsko toplotno obdelamo, temu pa sledi natezni preizkus. Na osnovi rezultatov dobljenih s pomočjo modificirane metode, smo uspeli na istem diagramu zajeti mehanske lastnosti, tehnološke parametre vakuumske toplotne obdelave in mikrostrukturo vakuumsko toplotno obdelanih preizkušancev iz hitroreznega jekla M2.
Ključne besede: orodje za precizno štancanje, lomna žilavost, trdota, vakuumska toplotna obdelava
1. Introduction
A carefully selected vacuum heat treatment process improves the basic characteristics of HSS M2. The required vvorking hardness and fracture toughness of HSS is determined mainly by the hard-ening and tempering temperatures, depending on the alloying1. With the optimai vacuum heat treatment process, the best possible combination of fracture toughness and hardness, and therefore, wear resistance, is reached.
Vojteh LESKOVŠEK. dipl. inž,, IMT. Lepi pot 11, 61000 Ljubljana
The design calculations of HSS tools must consid-er the material strength, vvith a special emphasis on fracture toughness, because of the danger of brittle tool fracture. Fracture toughness is defined as the ability of a material under stress to resist the propa-gation of a sharp crack. To establish the fracture toughness of HSS in hardened and tempered conditions, a non-standard testing method vvith small-scale specimens vvas developed. This method in-volves the introduction of a sharp crack at the notch root, in our čase, by pulsating round-notched tension specimens, thermal treatment and tensile testing. A high level of hardness makes round specimens
greatly sensitive to notches, so the test can fail due to unsuccessful pulsating. When successful prepul-sating, a fatigue crack is performed at the notch root of the specimen. The method was modified vvith the formation of a circumferential crack of defined depth at the root of the machined notch on soft annealing specimens, than a tensile test was performed after vacuum heat treatment. Our experiments confirm that the measurements based on the modified round-notched tension test can be successfully used to determine the fracture toughness.
2. Basic characteristics of high speed tool
steel M2
Due to the higher vvear resistance of HSS, they are nowadays used also for fine blanking, cold vvorking and deep dravving tools, especially in long series. Tool steels must vvithstand compressive stresses and abrasive or adhesive vvear, vvhile have a suffi-cient toughness to resist chipping and failure. HSS have better resistance to vvear in comparison to cold vvork tool steels because of the increased hardness of the matrix, and of the carbide phase.
The carbide phase in the matrix of HSS increases the vvear resistance vvhich is relative to the total vol-ume of carbides, and also to their hardness. The vvear resistance in HSS is mainly determined by vanadium carbides vvhich have a micro-hardness of 2200 to 2400 HV2 3, (Fig. 1).
—i_i ■_i_i i i
VC (OPel-jCj (FeWMo)6C. FČ3C. Matrixhigh Matnx
speed steel carbon steel
Figure 1: The comparative hardness of carbides found in tool steels2 Slika 1: Primerjalne vrednosti trdot karbidov, ki jih najdemo v hitroreznih jeklih
Hovvever, it must not be forgotten that HSS have a greater hot hardness. Even if the vvork pieces are plače into the tools vvhile cold, the vvorking tool sur-faces become hot.
Fractures, macro-chipping and micro-chipping can destroy the cutting edges. The ability of a steel to resist these phenomena is knovvn as toughness. The toughness that can be achieved by HSS is limited by the defects in the steel (carbide segregations and bands inclusions etc.). When the steel is subjected
to a load, stress concentrations can appear around the defects and cause a tool fracture, unless the stress concentrations are relieved by a local plastic flovv on the micro scale. The ability of the matrix to undergo plastic flovv can be altered vvithin wide limits by varying the hardness. Thus, the defects in the steel determine the maximum toughness vvhich can be achieved. On the other hand, the heat treatment determines the toughness degree actually achieved vvithin the limits set by the defects.
Vacuum heat treatment is one of the most important operations in the manufacturing of tools. Therefore, vvhen HSS are used for cold vvorking processes, the situation is met by choosing low hard-ening temperatures and tempering temperatures be-lovv the peak secondary hardening temperature, to improve fracture toughness, cutting edge strength, vvear resistance and dimensional stability. It is possi-ble to exert a positive influence on ali the parameters by vacuum heat treatment vvhich is carefully select-ed to suit the HSS is determined by a choice of vari-able tempering temperatures, it is often impossible to optimise the mechanical properties, e.g. fracture toughness. A general recommendation for tools that require good impact strength, such as fine blanking tools, is that they should be hardened from temperatures as low as 1050°C1. By this treatment, resistance to tempering is diminished. For tools subjected to high pressures in service, a previous tempering at about 600°C1 is recommended.
3. Experimental procedure
3.1 Material and treatment parameters
The test material selected was a conventional high-speed steel (HSS) of the AISI M2 type of the same melt. The chemical composition of the steel ex-amined is listed in Table 1.
Table 1: Chemical composition of HSS examined (in wt.-%)
Material	C Si	Mn	Cr	Mo	W	V	Co
AISI M2	0.87 0.29	0.30	3.77	4.90	6.24	1.81	0.53
Cylindrical round-notched tensile specimens vvith a diameter of 10 mm vvere machined from soft an-nealed bars vvith a Brinell hardness of 255. Specimens vvere fatigued to produce a sharp circumferential crack at the notch root, then austeni-tized in a vacuum furnace at temperatures of 1050°C, 1100JC, 1150°C and 1230 C respectively, quenched in a flovv of gaseous nitrogen at a pressure of 5 bar abs. and double tempered one hour at temperatures 510°C, 540°C, 570DC and 600"C respec-tively.
3.2 Mechanical tests
The geometry of cylindrical round-notched pre-cracked tensile specimens, prepared according Dieter's recommendation4 is shovvn in Fig. 2.
Our previous investigations56 confirmed that such small-scale specimens can be successfully used for the analysis of the relationship betvveen the mi-crostructural variations and the fracture toughness of the investigated steels. Accordingly to Grossmann's concept of hardenability, the formation of the uniform microstructure along the crack front is possible because of, the axial symmetry of the cylindrical specimens, in comparison vvith the conventional CT-specimens, vvhere this condition is not fulfilled.
vvhere oys is the yield stress of the investigated steel. This requirement (2) vvas fulfilled on ali our measurements. The fracture surface of the cylindrical round-notched and precracked specimens vvas ex-amined in SEM at low magnification. As is shovvn in Fig. 3, the fatigue crack propagation area vvas sharply separated from the circular central part of the fast fracture area, so that the diameter d of this area vvas easily measured.
- P
Figure 2: The geometry of a cylindrical round-notched and precracked tensile specimen Slika 2: Nestandardni cilindrični natezni preizkušanec za merjenje lomne žilavosti z zarezo po obodu ter utrujenostno razpoko v korenu zareze
For a round-notched precracked specimen, the stress intensity factor is given by Dieter4 as
KI =-^-(-1.27+ 1.72 D/d)	(1)
vvhere d is the radius of the uncracked ligament after fatiguing, P is the applied fracture load, and D is the outer diameter of the cylindrical specimen. In the ex-periments, it is essential for the outer diameter of the specimen in order to obtain a state of plain strain at fracture.
In order to apply linear-elastic fracture mechanical concepts, the size of the plastic zone at the crack tip must be small compared vvith the nominal dimen-sions of the specimen. The size requirement for a valid KIC test is given by Shen VVei et. al.7 as
D > 1.5 (K|C/crys)	(2)
Figure 3: Fracture surface of cylindrical round-notched and precracked tensile specimen vvith the circumferential fatigue crack propagation area sharply separated from the circular central fast-fractured area Slika 3: Prelomna površina cilindričnega nateznega preizkušanca z obodno zarezo, s kolobarjastim področjem propagacije utrujenostne razpoke, ki je ostro ločeno od osrednjega, naglo zlomljenega dela. Premer (d) naglo zlomljene prelomne površine lahko izmerimo z optičnim mikroskopom
4. Results and discussion
4.1 Microstructural characterisation
The microstructure develops in dependence on the hardening temperature, as vvell as the austenite grain size and the residual austenite content of the initial samples. Metallographic examination of specimens shovv that the austenite grain size of ali specimens vvhich vvere gas quenched from the austeniti-zation temperature 1050 to 1230°C vvas 21 to 8 SG, (Fig. 4).
The content of residual austenite in as queched condition vvas determined by X-ray diffraction. The absolute accuracy of the determination of the residual austenite contents vvas ± 1 vol%. The HSS AISI M2 steel is fine-grained, right up to high hardening temperatures, and exhibits a residual austenite contents betvveen 21 and 30 vol%.
- Tj,
The microstructure after hardening at 1050 C, SG 21. Mag. 500x.
Mikrostruktura po kaljenju s temperature 1050°C, SG 21. pov. 500x.
The microstructure after hardening at 1100 C. SG 18. Mag. 500x.
Mikrostruktura po kaljenju s temperature 1100 C. SG 18. pov. 500x.
The microstructure after hardening at 1150 C, SG 13. Mag. 500x.
Mikrostruktura po kaljenju s temperature 1150 C, SG 13. pov. 500x.
The microstructure after hardening at 1230 C. SG 8. Mag. 500x.
Mikrostruktura po kaljenju s temperature 1230 C. SG 8. pov. 500x.
Figure 4: Microstructures vvith a different austenite grain size from vacuum hardened specimens from M2 steel Slika 4: Mikrostruktura in velikost austenitnih zrn, vakuumsko kaljenih vzorcev z različnih temperatur austenitizacije
Figure 5: The microstructure shovvs carbides and tempered martensite vvith an austenite grain size of 13 SG
(TA:1150 C) and 8 SG (TA:1230°C)
Slika 5: Mikrostruktura karbidov in popuščanega martenzita z velikostjo avstenitnih zrn SG 13 (TA:1150 C) in SG 8
(Ta:1 230° C)
The carbide particles in ali the specimens were alike in size and position, vvhich vvas due to their ori-gin: ali the specimens issued from the same metal-lurgical melt vvhich vvas submitted to the same hot plastic transformation. The carbides looked like
strips, and had a size of 1 -20 (.im, (Fig. 5). The resid-ual austenite contents are, vvith reference to temper-ing parameters, belovv 1 vol% in ali samples. After metailographic etching, a stronger marking of the austenitic grain boundary could be noticed. L. Calliari
Austenitizing temperature
Figure 6: The microstructure of vacuum-hardened and tempered specimens examined by SEM Slika 6: Mikrostruktura vakuumsko kaljenih in popuščenih vzorcev, posnetih na SEM pri 10 000 kratni povečavi
Table 2: Vacuum heat treatment parameters and mechanical properties of prepulsating round-notched tesion specimens
	Vacuum heat treatment			Fracture
			Hardness	toughness
				
Spec.	Hardening( C)	Tempering( C)	HRc	KIC (MNm32)
No.	2 min.	2 x 1h		
01-02	1050	510	60.0	18.78
03-04	1050	540	60.5	18.26
05-06	1050	570	58.7	15.80
07-08	1050	600	52.8	16.43
09-10	1100	510	61.8	17.28
11-12	1100	540	62.2	15.69
13-14	1100	570	61.3	15.49
15-16	1100	600	55.0	16.99
17-18	1150	510	60.7	18.26
19-20	1150	540	63.3	13.14
21-22	1150	570	63.2	14.70
23-24	1150	600	57.8	15.63
25	1230	510	62.5	17.77
26	1230	540	65.0	10.55
27	1230	570	65.5	12.08
28	1230	600	63.0	12.95
et al.8, compared vacuum and conventional heat-treated samples of AISI M2, and found that the re-sults of over 100 tests did not point out noticeable dif-ferences among the samples treated vvith the two different procedures. Neither systematic data nor re-lationship vvith the treatment parameters are yet available on this subject.
The microstructure of the specimens examined by SEM at a higher magnification (Fig. 6) confirmed a carbide precipitation on the austenite grain bound-aries for HSS M2 at different austenitizing and tem-pering temperatures. The quantity of fine carbide particles decreased vvith the increase of austenitizing temperatures. In addition, it vvas also noticed that at higher austenitizing temperatures, particularly at 1230 C (last column in Fig. 6), the larger carbide particles in contacts of austenite grains seemed to covering the boundaries of the neighboring grains because of variable surface tension on the matrix-carbides boundary. The microstructure of the specimens vvas of martensite type. The eventual presence of small quantities of retained austenite (1 to 5 vol%)8 examined by optical microscopy, vvere too small to estimate vvithout fail in such a heterogeneous microstructure. This phenomena can be attributed to the fact that the heating rate vvas lovver in the vacuum furnace than in the salt bath. By heating the pre-pulsating round-notched tension specimens betvveen 1050 C and 1230 C, diffusion processes in the vacuum furnace took longer than in the salt bath, vvhich can possibly explain why, after metallograph-ic etching, a more intensive marking of the austenitic grain boundary can also be noticed.
4.2 Mechanical tests
Experiments9 vvere performed on 28 prepulsating round-notched tension specimens, (Table 2), heat-treated in an IPSEN VTTC-324 R single chamber vacuum furnace vvith uniform high pressure gas quenching.
In the follovving, the assumption is made that the values KIC are determined by the above-mentioned method. The obtained values of KIC are very similar to those obtained by G. Hoyl10, vvho determined the fracture toughness KIC for HSS M2 steel, e.g. 18,3 MNm3;' for sample austenitized 4 minutes at 1220 C and tempered 1 hour at 510 C, by conventional methods. Belovv a hardness level of about 50 HRc, fracture toughness is dependent only on the hardness of the sample10. At higher levels of hardness, the fracture toughness for M2 varies in-versely vvith the austenitizing temperature, as shovvn in Fig. 7. G.Hoyl10 discovered that above a hardness level of 60 HRc, fracture toughness is in-sensitive to most metallurgical factors.
The effects of tempering betvveen 510 and 600 C on fracture toughness are shovvn for M2 in the same figure. As expected, there is a minimum of toughness values corresponding to the hardness peak. The net effect of tempering is attributed to a combination of stress relief and a reduction in ductility due to the secondary hardening effect.
F T1050 + H1050 ' FTIIOO o H1100 * FT1150 0 H1150 a FT1230 * H1230
65
H
d
55 n
e
45 H
R
35
Figure 7: The effect of austenitizing and tempering temperature on fracture toughness and hardness of M2
steel (FT-fracture toughness; H-hardness) Slika 7: Vpliv temperature austenitizacije in popuščanja
na lomno žilavost KIC in trdoto, vakuumsko toplotno obdelanega hitroreznega jekla M2 (FT-lomna žilavost, H-trdota)
In examining the evolution of tempering10, it vvas found that there vvas a peak value of fracture toughness for low tempering temperatures, (belovv 500r C). The obtained values are similar to those obtained in tempering at the conventional temperature, 25 C above the peak of the secondary hardening temperature. This is considered as advantageous, but as the effect is due to retained austenite that could transform later, the under-tempered tools could be susceptible to dimensional instability in ser-vice, vvhich is unacceptable for fine blanking tools.
On the basis of the experimental results, it vvas possible to dravv the diagram shovvn in Fig. 8 vvhere the technological parameters of the vacuum heat-treatment, mechanical properties and microstructure of the vacuum heat treated specimens are simulta-neously combined.
From the diagram in Fig. 8. it is also evident that the fracture toughness for the tempering temperatures 540 C, 570 C and 600 C, respectively, in-creases vvith the decrease of hardening temperature in agreement vvith observations in reference 10. On the other hand, for the tempering temperature of 510' C, it vvas found that the fracture toughness values vvere very close, though slightly higher than for 600 C, irrespective of the hardening temperatures. On the basis of the curves in Fig. 8. it can be reliably assumed that the HSS M2 hardened from low austenitizing temperatures and tempered at 510 C can achieve the optimal combination of hardness and fracture toughness.
Fracture toughness (MNm-3/2)
500 5Ž0 540 560 580 600
Tempering temperature (°C)
		---
TT	510 °C	—t— TT 540 °C
-*— TT	570 °C	-o— TT 600 °C
Austenite grain size SG
Figure 8: The influence of austenite grain size on the fracture toughness of HSS AISI M2, (TT-tempering temperature, TA-hardening temperature, HRc-hardness at 510 C)
Slika 8: Vpliv velikosti austenitnega zrna na lomno
žilavost hitroreznega AISI M2, (TT-temperatura popuščanja, TA-temperatura kaljenja, HRc-trdota po pop. na temperaturi 510 C)
The relationship betvveen fracture toughness and austenite grain size, f.i. SG grade 8, shovvs us that at the tempering temperatures betvveen 510 and 600 C, the obtained values KIC are from 17,77 to 10,55 MNm3 2 and the difference is quite important in practice. Different fracture toughness at equal austenite grain size or the nearly constant fracture toughness of HSS M2 hardened from different austenitizing temperatures, f.i. from 1050 to 1230 C, and double tempered at 510 C, is in accordance vvith the hypothesis that the austenite grain size is not the dominant parameter effecting the fracture toughness of HSS M2.
The result of the present investigation is useful for the optimisation of vacuum heat treatment for different HSS tools submitted to tensile impact stress dur-ing use vvhere an optimal combination of hardness and fracture toughness are decisive.
4.3 Tool life tests
Long production runs have underlined the impor-tance of an improved fine blanking tool life, (Fig. 9).
On the basis of the experimental results shovvn in Fig. 8. it vvas found that the optimum vacuum heat treatment of fine blanking HSS M2 tools needs at least tvvo preheating stages (650 and 850°C respec-tively), a variable hardening temperatures (usually 1050 to 1150l C) and double tempering at the same temperature (510 C/1 hour).
Figure 9: Fine blanking tool for a ratchet vvheel Slika 9: Orodje za precizno štancanje zobnika varnostnega pasu
The life of a fine blanking tool varies considerably depending on the size and design of the blank, the type of blanking steel, and care and maintenance. To establish tool life, vve selected three tools for a fine blanking ratchet vvheel. The punches and dies vvere from HSS AISI M2. Blanks vvith a material gauge of 4 mm vvere from AISI C 1045 blanking steel in a spheroidized-annealed condition. The punches and dies vvere stress-relieved at 650 C 4 hours and vacuum heat-treated in a single chamber vacuum furnace vvith uniform high pressure gas quenching. Depending on the alloying and the condition of austenitization (1100°C/2min), the final hardness of 61.5 HRc and the final fracture toughness of 17.28 MNm 3'2 vvas reached after double tempering at 510 C for 1 hour (Fig. 8). The vacuum heat-treated punches and dies vvere tested on a 250t triple-action hydraulic press and compared vvith the fine blanking tools for fine blanking ratchet vvheels conventionally heat-treated in a salt bath as follovvs: stress relieved at 650°C/4hour, preheated at 450 C, 650 C and 850 C respectively, austeni-tized at 1100 C/2min and control-quenched to 550°C in 5 min, held isothermally at 550°C for 10 min and cooled to 80JC vvith air, follovved by double tempering at 600°C/1 hour. The final hardness of the tools achieved after double tempering at 600°C/1 hour, vvas 58 to 59 HRc, depending on the alloying.
The basic trial parameters (such as the cutting force, strikes per tirne unit, temperature, greasing etc.) vvere constant during the experiments, and did not affect the final results. The differences observed in fine blanking tools could only originate in the punches and dies themselves. 15.000 ratchet vvheels vvere made vvith each tool and edges exam-ined in a binocular microscope to estimate the dam-age. Minor defects vvere observed on the vacuum heat-treated punches and dies and those conven-tionally heat-treated in a salt bath.
Figure 10: Defects on the punch cutting edge Slika 10: Poškodbe na rezilnem robu pestiča
The wear of the punches and dies increases vvith the operation tirne. SEM observations show that the starting vvear of the cutting edges of the punches and dies could not be easily determined. The material shovved not only surface fatigue, but also adhesion and abrasion, (Fig. 10).
The experiments shovved that higher vvorking hardness (61.5 HRc) and improved fracture toughness of vacuum heat-treated punches and dies - particularly those double-tempered at 510° C - had significant ef-fect on the defects on the cutting edges. Compared to conventionally heat treated tools tempered at 600UC vvas the vacuum heat-treated tool life by 15 to 20%, greater. During the tool tests, no effects vvere found that could be related to the in service dimen-sional instability due to the later-transformed re-tained austenite. Namely, X-ray structural analyses shovved that the content of retained austenite did not exceed 1 vol% in ali the specimens.
5. Conclusions
The exact significance of fracture toughness as it affects HSS properties and service behavior is not thoroughly understood, and there are differences in behavior betvveen grades and process routes. Hovvever, the modified method for the establishment of fracture toughness improved by IMT, appeared to be a successful method for establishing the fracture toughness of HSS.
The measurements of vvear in the cutting edges of punches and dies shovv that double tempering at 510°C/1 h after vacuum heat treatment prolongs the life of fine blanking tools for ratchet vvheels by 15 to 20%, compared to conventionally heat-treated tools vvhich vvere hardened at the same austenitizing temperature and tempered tvvice at 600°C. It seams that it is not the type of process vvhich substantially af-
fects the tool life, but first of ali, the proper choice of hardening and tempering temperatures.
The vvear resistance of punches and dies cannot only be described as a material property, but as a property of a complex tribological system. Yet, it is proven that the vvear resistance depends, above ali, on the microstructure of the tool material and on its physical and chemical properties.
The presented results, obtained by the evaluation of daily confirmed data, justify the use of modem vacuum heat-treatment technology and the use of the nevvest high-performance HSS steels to achieve great improvements in tool lives and overall econo-my.
6. References
1	K. E. Thelning: Steel and its Heat Treatment. Bofors Handbook, 1975, 313 -319
2	R. VVilson: Metallurgy and Heat Treatment of Tool Steels, McGravv-Hill Book Company UK, 1975, 69
3	H. Czichos: Tribology, Elsevier Scientific Publishishing Company Amsterdam, 1978
4	G. E. Dieter,: Mechanical Metallurgy, McGravv-Hill. 1986, 358
5	B. Ule, D. Kmetic, A. Rodič: Rudarsko-Metalurški zbornik, 1989, 3, 509-519
6	V. Leskovšek, B. Ule, A. Rodič, D. Lazar: Vacuum. 43. 1992, 5-7, 713-716
7	Shen Wei et.al.: Engineering Fracture Mechanics. 16, 1982, 1, 69-92
8	L. Calliari, A. Molinari, E. Ramous, G. Torbol. G. Wolf: Vacuum and Conventional Treatment of Tool Steels, Instituto per la Ricerca Scientifica e Technologica 38050 Povo (Trento) ltaly, 49-57
9	V. Leskovšek, B. Ule, A. Rodič: Metals Ailoys Technologies, 27, 1993, 1-2, 195-204
10G. Hoyle, High Speed Steels, Buttervvorth & Co. Ltd. 1988, 143-146