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Fatigue crack growth rates and fracture toughness in electroslag refined En 52 steel
Fatigue crack growth rates and fracture toughness in electroslag refined En 52 steel M o h a n G. Hebsur, K. P. A b r a h a m and Y. V. R. K. Prasad
En 52 steel has been electroslag refined and the resultant effects of refining on its mechanical properties have been assessed. It was found that refining caused a decrease in fatigue crack growth rates and increases in fatigue strength, fracture toughness, Charpy fracture energy and tensile ductility. Fatigue crack growth rates in region I and in region III were found to be considerably lower in the electroslag refined steel: they were unaffected in region II. The fracture toughness values for the electroslag refined steel are nearly twice those estimated for the unrefined steel. Measurements on heat-treated samples have shown that the electroslag refined steel has a better response to heat.treatment. The improvement in the mechanical properties is explained in terms of the removal of nonmetallic inclusions and a reduction in the sulphur content of ~he steel.
Electroslag refining (ESR) is one of the secondary remelting processes by which high quality steels for sophisticated applications can be produced. The ESR technique is a corn. paritively simple and low cost process capable of producing steels with improved cleanliness, reduced sulphur content and better chemical homogeneity) In this process, the steel in the form of a consumable electrode is remelted under a cover of molten synthetic slag of high reactivity. In addition to the removal/redistribution of nonmetallic inclusions, efficient desulphurization occurs due to high slag basicity and the high temperature prevailing at the droplet/slag interface. All these factors influence the mechanical properties of the steel. The aim of the present investigation is to study the influence of electroslag refining on the mechanical properties, particularly on the fatigue crack propagation rates and the fracture toughness of En 52 steel: both these parameters are known to be improved by the removal of nonmetallic inclusions. En 52 steel was chosen for this investigation in view of its extensive applications as heavy duty inlet and exhaust valves in internal combustion engines and jet engines; in addition, data on the fracture toughness and the fatigue crack propagation rates for this steel are lacking. In this investigation, the approach has been to compare the data on ESR steel with unrefined steel. The studies have been made on the heat-treated material so that they will have practical utility.
EXPERIMENTAL Forged rods of commercial En 52 steel (referred to as 'unrefined steel'hereafter)were electroslagrefinedusing
a synthetic slagmixture containing CaF 2 :Al205 (80 :20) in a laboratory scaleelectrosiagrefiningunit. The description of the unit and the refiningprocedure have been published elsewhere.2 A current of 800 A and a voltage of 20 V were found to give satisfactoryingots.The chemical analysisof the steelbefore and afterrefiningis given in Table I. The microstructuraldetailssuch as grain sizeand inclusioncontent were recorded using standard metallographic techniques. Mechanical testingwas carriedout on forged and heat-treatedspecimens. O n the basisof details availablein the literature,3 the following heat treatment was given to the specimens: austenitizingat I050°C for one hour followed by oil quench and tempering at various temperatures in the range 600 °-850°C for two hours. The room temperature tensile,impact (Charpy) and fatigue strengthswere measured using standard A S T M specimens. Fatigue testingwas performed on smooth specimens under rotation bending load and the S/N curves were recorded. The fracturetoughness was measured on a 25.4 m m thick compact tension specimen following the procedure recommended 4 in A S T M - 3 9 9 . Fatigue precracking was carried out on a 50 tonne M T S machine using a haversine load cycle (1000 kg/t00 kg) at 30 Hz. The precracked specimens were pulled in tension at a cross-head speed of 1.0 ram/rain and the displacement was measured using a clip-ongauge. The load versus displacement curves were recorded on an x-y recorder. Fatigue crack growth rateswere measured on 25.4 m m thickcompact tension specimens following the recommendation A S T M E6 47-78. 5 The specimens were precracked using the haversineload cycle (I000 kg/ I00 kg) at a frequency of 40 Hz. Further crack growth
Table 1. Chemical analysis (in weight) of En 52 steel before and after ESR
Unrefined ESR ingot (top)
ESR ingot (middle) ESR ingot (bottom)
0142-1123/80/040147-06
C
Si
Mn
P
S
Cr
Ni
V
0A2
3 A0
0.50
0,030
0.014
8.15
0.47
0.03
0A3
3.39
0.49
0,029
0.006
8.13
0A7
0.03
0.42
3.38
0.49
0,030
0.005
8.14
0.46
0.03
0A4
3.40
0.48
0,030
0.006
8.12
0.47
0.03
$ 0 2 . 0 0 © IPC Business Press L i m i t e d 1 9 8 0
INT. J. F A T I G U E O c t o b e r 1 9 8 0
147
in the precracked specimenwas measured by fatigue loading with load control under sinusoidal tension at a frequency of 10 Hz. Two different values of stress intensity ratio (R = Krnm/Kmax,where Kmax and Kmin are the maximum and minimum stress intensities during each cycle) equal to 0.1 and 0.5 were used. The fatigue crack length was continuously monitored using a travelling microscope. The printing of Moir4 grid lines on the specimen surface and illuminating it with a stroboscope facilitated measurement of crack length. The average crack extension between two consecutive measurements was approximately 0.025 ram. All the fracture surfaces were examined with a scanning electron ndcroscope to determine the mode of fracture. R E S U L TS
From the chemical analysis data given in Table 1, it can be seen that the sulphur content of the steel is reduced considerably by electroslag refining while the contents of other alloying elements remain well within the limits. Further, the compositional variations from top to the bottom of the ingot were only marginal. The distribution of inclusions before and after ESR is shown in Fig. 1 in the form of number of inclusions per mm 2 versus inclusion size and the inclusion rating carried out according to ASTM E - 4 5 comparison method 6 is shown in Table 2. The ESR steel is free from large inclusions (>6/J.m) of all types. The microstructure of the as-quenched specimens consisted of martensite and specimens tempered at different temperatures had a tempered martensite structure. All these structures are essentially the same for both unrefined and ESR steel. The room temperature tensile properties of the refined and unrefined steel in different heat treatment conditions are given in Table 3. While the yield strength (YS) and ultimate tensile strength (UTS) are unaffected by ESR, ioa
En 52 steel
Table 2. Inclusion rating o f En 52 steel according t o A S T E 4 E45 comparison method Inclusion t y p e
Unrefined
Sulphides Alumina Silicates Oxides Volume fraction
ESR
Thin
Heavy
Thin
Heavy
25 20 25 25
15 10 15 15
05 05 05 05
-
(%)
Average inclusion spacing (/lm)
018
0.025
10
12
Table 3. Tensile data on En 52 steel before and after electroslag refi ning Treatment
1050°C OQ 600°C T 700°C T 800°C T
0~% YS (MPa)
UTS (MPa)
El (%)
UR
ESR
UR
860 810 716
870 815 724
1205 1220 3 1010 1020 16 960 960 26 920 930 28
ESR
UR
UR
ESR
6 23 36 48
6 28 45 50
12 37 62 70
the ductility is enhanced to some extent. In Fig. 2, the fatigue data (S/N curves) obtained on smooth specimens failed under rotating bend tests are given. These specimens were quenched in oil from 1050°C and tempered at 700°C for 2 h. The fatigue limit has increased by about 58% in ESR material compared with similarly treated unrefined steel. The fatigue crack growth rates (da/dN) are obtained from the crack length a versus number of cycles N data using a seven point polynomial method. 5 Since the cyclic load limits were kept constant, the applied z ~ (Kmax Kmin) increases as the crack length increases. In general,
En 52 steel ~
"~1
Refined
•
l0 = 3 L-- I I I I I I
f~
i
"6
I
ESR
Note: OQ -- oil quench; YS - yield stress; El -- elongation T -temper; UR -- unrefined; ESR - refined; RA -- reduction in area
"-i
"s
RA (%)
I I L.__ I I
10--
! 40( /
--
0
Unrefined
~ O
C
30(
I I--.. 1 L--- I I__.
ip--ESR
1 I
I 0
4
i
i
16 Inclusion size (/=rn) 12
L '--i--?-
20
24
4 x IO= 28
Fig. 1 Plot of number of inclusions per mm = versus inclusion size for En 52 steel before and after electroslag refining
I N T J. F A T I G U E O c t o b e r 1 9 8 0
IO4
I Io~ Number of cycles
Ioe
F i g 2 S/N curve for En 52 steel obtained on smooth rotating bending fatigue specimens oil quenched from 1050°C and tempered at 600°C
the fatigue crack growth data a r e represented in the form of da/dN versus AK as suggested by Paris. 7 The data obtained on unrefined and tel'reed En 52 steel in heat treated condition are shown in Figs 3 and 4. These curves exhibit three distinct regions. 8 At low AK (region I), the curve asymptotically approaches the so.ailed threshold ~ K value. The curve then changes to a linear region (region II) before turning asymptetic again as AK approaches the fracture toughness (KIc) of the material (region III). It is generally recoguised that region I and region III are sensitive to microstructure and load ratio effects; while the slope in stage II is insensitive to these variables. In the ESR steel, the (near) threshold stress intensity has increased and crack growth rates in region III have decreased whereas the slope in region II is unaffected in comparison with the unrefined steel. The fracture toughness ( / ~ ) and the Charpy V notch fracture energy obtained on the unrefined and the ESR steel are shown as a function of tempering temperature in Figs 5 and 6 respectively. These figures indicate that the ESR steel exhibits superior (nearly double) fracture toughness and impact properties in all conditions of heat treatment.
En 52 steel
A
16~ _
o•
.o
Ol •
~
•
• A A
•
,?
&
16~ _ ~
DISCUSSION
•
o O• 0
•
O& • O• • 0 • 0A o • •
O&
I050°C
oil
quench
800=C temperature
•
•
0 & o•
0
O•
•
&
Oa
• ~. •
•
•
O•
o"
i~ s _
oo
O&
A
~.
• •
•o A
•
•
•0
16 4 -
• t.
•
• •0
•
E
•
R=0.5
R=0.1
• •
Unrefined
o
Refined
•
• • •
A
D u r i n g e l e c t r o s l a g r e f i n i n g o f steels, t h e f o l l o w i n g c h a n g e s
Occur: (i)
The large sized non-metallic inclusions are removed and the finer inclusions are redistributed more uniformly. (ii) Sulphur content of the steel is reduced appreciably.
(iii)
I0
I
I
I
I
20
40
60
80
I
I
100 120
Alternating stress intensity, Z~K (MPoq-~) Fatigue crack growth rates for En 52 steel as a function of Fig. 4 A K at R = 0.1 and R = 0.5
Chemical homogeneit7 is achieved. En 52 steel
En 52 steel Austenifizir~ ~ a t u r e o
iO-~ __ &
•
O
A
O
•
: 1050°C
• •
•
O & &
oO & • • O O• 0 • 0 • •
&O O&
E
AO& Z
&O•
•
• •
•
:S
0 • 0 &A •
{
2 =-
%•e
& 0
ao • • Ao• • &o• •
'
I0"
R=0.1 1050eC oil quench Tempering tomper~uN 650=C
~0
,, o • • o "O • •
._~ -(
Io
&O & • O
Unrefined
800°C 0
Refined
Refined
•
I
•
I0--
E)
•
Umef~
t 20
I 40
I 60
I 80
I R)o
I 120
AlterrN]ting stress intensity, LIK (MPoq-~) Fig. 3 Fatigue crack growth rates for En 52 steel as a function o f AK at R = 0.1
o As Q
I 6O0
i
70|0
Tempering t e m ~
I ("C)
°1
Fig. 5 Fracture toughness o f En 52 steel as a function of tempering temperature
I N T . J. F A T I G U E
October 1980
149
of static fractures, including cleavage, intergranular and fibrous fracture. 14 The SEM fractographs taken on specimens of En 52 steel before and after refining, treated at 1050°C, oil quench and 800°C temper, are shown in Fig. 7. The unrefined specimen has failed by quasi-cleavage and an intergranular mode whereas refined steel has failed by a fibrous quasi-cleavage mode. The fatigue crack growth rate in region II is reported to be insensitive to microstructure and load ratio effects. Paris 7 has shown that the slope in this region follows the equation:
En 52 steel
3C
-D
zc ~t
(aa/dN) = C (~)m
(1)
where C is a m a t e r i a l c o n s t a n t a n d m is the slope o f the
Z > 0
IC
t
,
600 0
i
Refined
•
Unrefined
o
I
700
I
I
800
1
850
Temperin~ temperature (°C)
growth rate curve. The value of m remained unchanged in the unrefined and refined steel and is also not affected by heat treatment. The value of m for En 52 steel lies in the range 2.4-3.0. The crack growth in region II is controlled by the amount of crack opening in each cycle and gives rise to striations. The effect of nonmetallic inclusions on tensile ductility and fracture toughness is discussed by Schwalbe 9 in terms of well known void formation mechanisms. Due to the difference in the deformability of matrix and the inclusion, decohesion occurs at the interface leading to void
Fig. 6 Charpy fracture energy of En 52 steel as a function of tempering temperature
The electroslag refining has completely eliminated inclusions with sizes larger than 6 pm and reduced the number of those with sizes between 4-6/n-n. Furthermore, the lowering of sulphur content to very low values (0.005%) reduces the number and size of sulphide inclusions. The mechanical properties, particularly fatigue strength, fracture toughness and tensile ductility, are very sensitive to the presence of nonmetallic inclusions. 9 The fatigue strength depends to a large extent on the crack initiation at the debonded inclusion/matrix interface. The nonmetallic inclusions act as stress raisers and initiate cracks. It has been shown 1° that in a given material there is a critical inclusion size (larger than 20/~'n) which will be able to nucleate a fatigue crack. Brooksband and Andrews n have estimated the residual stress distribution surrounding an inclusion and have suggested that such tessellated stresses could be responsible for matrix yielding and crack initiation during fatigue. Lindley and Richards 12 have shown that there is no significant effect of particles on the fatigue crack propagation rates by alteration of flow parameters below general yield. Shih and Araki 13 have observed that fatigue growth rates associated with a dimple mechanism decreased with increasing cleanliness of the steel. They have also noticed that inclusions are more dangerous to fatigue crack growth at lower tempering temperatures. It is generally recognised that regions I and III of the fatigue crack growth rate versus z~K curve a;'e sensitive to microstructure and load ratio effects. 14 For similarly heat-treated steel, the increase in the threshold stress intensity factor and the decrease in fatigue crack growth rate in region III that occurred in the ESR steel (Figs 3 and 4) can be attributed to the removal of nonmetallic inclusions. In region I, the crack growth rates are slow and the threshold values could not be estimated accurately. In region III where Kmax approaches KIc, the growth rate is rapid and failure occurs by modes that are characteristic
150
I N T . J. F A T I G U E O c t o b e r 1 9 8 0
~ii~ .........
Fig. 7 (a) SEM fractograph obtained on the fracture surface of electroslag refined En 52 steel in region III of fatigue crack growth showing mixed mode fracture consisting of ductile and quasi-cleavage features; and (b) SEM fractograph of unrefined En 52 steel in region III of fatigue crack growth showing mixed mode fracture consisting of intergranular and cleavage features
formation. The void formation can also occur due to the fracture of inclusions ahead of the crack due to large plastic strains occurring in the plastic zone. The ductile fracture proceeds by a process of void coalescence. Thus the ductility depends on the volume fraction, kind, size and distribution of inclusions. The sulphide inclusions are most deleterious as far as the fracture toughness is concerned. Birkle et a115 have shown that Kic increases with decrease in sulphur content in the steel. Further, the Mn/S ratio decides the shape of the MnS inclusions and at higher ratios the inclusions prefer a globular shape to a plate.like shape. 16 Gladman 17 has shown that the plate-shaped inclusions are more dangerous than equiaxed ones as far as fracture toughness is concerned. The ductility is also affected by the size of the inclusions, since the plastic strain necessary to fracture an inclusion increases with decreasing inclusion size. Schwalbe 9 has shown that the particle size reduction causes an increase in fracture toughness. Whereas the particle size determines the onset of voids, the distance between the particles controls the void growth and crack propagation. The electroslag refining has eliminated the large size inclusions (>6 pro) at which void formation would readily occur. Also the volume fraction of inclusions is reduced by electroslag refining. Moreover, the reduction in sulphur content has increased the Mn/S ratio which helps in the formation of globular MnS inclusions. All these factors contribute towards improving the tensile ductility, fracture toughness and Charpy fracture energy in the ESR steel. The critical defect size can be calculated from the Kic values obtained on the refined and unrefined steel following the procedure suggested by Kiessling and Nordberg. l° This gives the maximum defect size that the material can tolerate with the stress intensity factor not exceeding the Kic value. For a disc shaped crack, the critical defect size is given by A = 1.57 k 2 (Kic/Oy) 2
(2)
and for an eMpsoidal surface crack, the critical defect size ( d e p t h ) is B = 0.S k 2 (Kic/Oy)2
(S)
where k = factor of safety, generally taken as i. The calculated values of the diameter of the inner spherical defect A and the depth of the ellipsoidal surface crack B at o = ay (Oy = yield stress) are presented in Table 4. Electroslag refining has increased the critical defect size in all conditions of heat-treatment. With a view to comparing the fracture toughnesses of the unrefined and the ESR steel, the KQ values are plotted against their individual UTS values in Fig. 8. For a given heat treatment the fracture toughness values in the ESR steel are the highest. The scanning electron micrographs obtained with the unrefined and refined steel are shown in Fig. 9. The refined
En 52 steel
10C
\
:'2.
9£
8C
v
o
! °° 40 3C
2C
I
19 00
I
900
I
1000
II00
UTS (MPo) Fig. 8
Fracture toughness versus tensile strength in En 52 steel
steel exhibits fibrous and quasi~leavage t y p e o f fracture whereas the unrefined steel has failed by intergranular and quasi-cleavage fracture. The large ~zed inclusions in the unrefined steel provide sites f o r early cleavage and the high sulphur content results in segregation o f sulphur at the grain boundaries 18 thus causing the l o w values o f fracture
toughness.
CONCL USIONS Electroslag refining of En 52 steel decreases the fatigue crack growth rates and increases the fatigue strength, fracture toughness, Charpy fracture energy and tensile ductility. The improvement in these mechanical properties can be attributed to the removal of nonmetallic inclusions and the reduction in sulphur content. Furthermore, the fatigue crack growth rates in region I and region III decrease in ESR steel.
REFERENCES 1.
Holzgruber, W. 'Possibilities and limitations to influence the structure of ESR ingots and properties of ESR products', Proc Fifth Int Syrnp on Electros/ag and Other Special Melting Technology (Pittsburgh, USA 1975) pp 7 9 - 8 0
Table 4. Critical defect size at o = ey for the En 5 2 steel
Refined
Unrefined Treatment 1050°C OQ 600 ° C T 700° C T 800 ° C T
YS (MPa)
KQ
(MPa~/m)
A* (ram)
B** (mm)
YS (MPa)
KQ (MPa~/m)
A (ram)
B (ram)
860 810 716
32 42 56
2.1 4.2 9.6
0.40 0.80 1.80
870
67
9.3
1.8
815
90
19.0
3.6
720
97
28.0
5.4
*A -- Diameter of inner spherical defect **B -- Depth of ellipsoid surface crack
INT. J. F A T I G U E October 1980
151
5.
6.
7.
8.
9.
10.
11.
12.
13.
14. i
15.
16.
17. Fig. 9 (a) SEM fractograph of the fracture toughness specimen of electroslag refined En 52 steel exhibiting fibrous and quasi-cleavage fracture; and (b) SEM fractograph of the fracture toughness specimen of unrefined En 52 steel showing intergranular and cleavage fracture 2.
3.
4.
152
Hebsur, M. G., Abraham, K. P. and Prasad, Y. V. R. K. 'Hot working characteristics of electroslag refined En 52 steel: a hot torsion study', Metals Techno/ (to be published) Woolman, J. and Mottram, R.A. The Mechanica/ and Physical Properties o f the British Standard En Steels, Vol 3 (Pergamon Press, England 1968) pp 180-185 'Standard Test Method for Plane Strain Fracture Toughness of Metallic Materials', Annual Book o f ASTM Standards, Part 10, Designation £ 3 9 9 - 7 4 (ASTM, 1978)
I N T . J. F A T I G U E O c t o b e r 1980
18.
"Tentative Method for Constant-Load-Amplitude Fatigue Crack Growth Rates Above 10 -s rnm/cycle', Annual Book o f ASTM Standards, Part 10, Designation E647--78T (ASTM, 1978) 'Recommended Practice for Determining the Inclusion Content of Steel', Annual Book o f ASTM Standards, Part 1 I, Designation E 4 5 - 7 6 (ASTM, 1976) Paris, P. C. "The fracture mechanics approach to fatigue', Proc lOth Sagamore Army Materials Research Conference (Syracuse University Press, USA, 1964) pp 107-132 Imhof, E. J. and Barsom, J. M. 'Fatigue and corrosion fatigue crack growth of 4340 steel at various yield strengths',ASTM STP 536 (1973) pp 183--205 Schwalbe, K. H. 'On the influence of microstructure on crack propagation mechanisms and fracture toughness of metallic materials', Engng Fracture Mech 9 (1977) pp 795--832 Keisling, R. and Nordberg, H. 'Influence of inclusions on mechanical properties of steel', Int Conf on Production and Application o f Clean Steels (BalatonfiJred, Hungary, 1971) pp 179--185 Brooksbank, D. and Andrews, K.W. 'Stress fields around inclusions and thei r relation to mechanical properties', Int Conf on Production and Application o f Clean Steels (Be latonfUred, Hungary, 1971) pp 186--195 Lindley,T.C.and Richards, C.E. "lnfluenceof microstructure on fatigue crack propagation in steels', in Effects o f Second Phase Particles on the Mechanical Properties o f Steel (Iron and Steel Institute, London, 1971) pp 119--123 Shih,T. Y.and Araki,T. 'The effect of nonmetallic inclusions and microstructure on the fatigue crack initiation and propagation in high strength carbon steels', Trans ISI Japan 13 (1973) pp 11--19 Parker, E. R. 'Inter-relation of composition, transformation kinetics, morphology and mechanical properties of alloy steels', Metal Trans 84 (1977)pp 1025--1042 Birkle, A. J., Wei, R. P. and Pellissier, G. E. 'Analysis of plane strain fracture in 0.45C - Ni--Cr--Mo steel', Trans ASM 59 (1966) pp 9 8 1 - 9 9 0 McTegart, W. J. and Gittins, A. 'The role of sulphides in the hot workability of steels', in Sulphide Inclusions In Steels, ASM 798--211 (1974) Gladman, T., Holmes, B. and Mclvor, I. D., 'Effects of second phase particles on strength, toughness and ductility', in Effects o f Second Phase Particles on the Mechanical Properties o f Steel (Iron and Steel Institute, London, 1971) pp 68--78 Hebsur, M. G., Abraham K. P. and Prasad, Y. V. R. K. 'Effect of electroslag refining on the fracture toughness and fatigue crack propagation rates in heat-treated AlSl 4340 steel', Engng Fracture Mech (to be published)
AUTHORS The authors are all with the Department of Metallurgy, Indian Institute of Science, Bangalore 560 012, India. Enquiries should be addressed to Dr Prasad in the first instance.
En 52 steel has been electroslag refined and the resultant effects of refining on its mechanical properties have been assessed. It was found that refining caused a decrease in fatigue crack growth rates and increases in fatigue strength, fracture toughness, Charpy fracture energy and tensile ductility. Fatigue crack growth rates in region I and in region III were found to be considerably lower in the electroslag refined steel: they were unaffected in region II. The fracture toughness values for the electroslag refined steel are nearly twice those estimated for the unrefined steel. Measurements on heat-treated samples have shown that the electroslag refined steel has a better response to heat.treatment. The improvement in the mechanical properties is explained in terms of the removal of nonmetallic inclusions and a reduction in the sulphur content of ~he steel.
Electroslag refining (ESR) is one of the secondary remelting processes by which high quality steels for sophisticated applications can be produced. The ESR technique is a corn. paritively simple and low cost process capable of producing steels with improved cleanliness, reduced sulphur content and better chemical homogeneity) In this process, the steel in the form of a consumable electrode is remelted under a cover of molten synthetic slag of high reactivity. In addition to the removal/redistribution of nonmetallic inclusions, efficient desulphurization occurs due to high slag basicity and the high temperature prevailing at the droplet/slag interface. All these factors influence the mechanical properties of the steel. The aim of the present investigation is to study the influence of electroslag refining on the mechanical properties, particularly on the fatigue crack propagation rates and the fracture toughness of En 52 steel: both these parameters are known to be improved by the removal of nonmetallic inclusions. En 52 steel was chosen for this investigation in view of its extensive applications as heavy duty inlet and exhaust valves in internal combustion engines and jet engines; in addition, data on the fracture toughness and the fatigue crack propagation rates for this steel are lacking. In this investigation, the approach has been to compare the data on ESR steel with unrefined steel. The studies have been made on the heat-treated material so that they will have practical utility.
EXPERIMENTAL Forged rods of commercial En 52 steel (referred to as 'unrefined steel'hereafter)were electroslagrefinedusing
a synthetic slagmixture containing CaF 2 :Al205 (80 :20) in a laboratory scaleelectrosiagrefiningunit. The description of the unit and the refiningprocedure have been published elsewhere.2 A current of 800 A and a voltage of 20 V were found to give satisfactoryingots.The chemical analysisof the steelbefore and afterrefiningis given in Table I. The microstructuraldetailssuch as grain sizeand inclusioncontent were recorded using standard metallographic techniques. Mechanical testingwas carriedout on forged and heat-treatedspecimens. O n the basisof details availablein the literature,3 the following heat treatment was given to the specimens: austenitizingat I050°C for one hour followed by oil quench and tempering at various temperatures in the range 600 °-850°C for two hours. The room temperature tensile,impact (Charpy) and fatigue strengthswere measured using standard A S T M specimens. Fatigue testingwas performed on smooth specimens under rotation bending load and the S/N curves were recorded. The fracturetoughness was measured on a 25.4 m m thick compact tension specimen following the procedure recommended 4 in A S T M - 3 9 9 . Fatigue precracking was carried out on a 50 tonne M T S machine using a haversine load cycle (1000 kg/t00 kg) at 30 Hz. The precracked specimens were pulled in tension at a cross-head speed of 1.0 ram/rain and the displacement was measured using a clip-ongauge. The load versus displacement curves were recorded on an x-y recorder. Fatigue crack growth rateswere measured on 25.4 m m thickcompact tension specimens following the recommendation A S T M E6 47-78. 5 The specimens were precracked using the haversineload cycle (I000 kg/ I00 kg) at a frequency of 40 Hz. Further crack growth
Table 1. Chemical analysis (in weight) of En 52 steel before and after ESR
Unrefined ESR ingot (top)
ESR ingot (middle) ESR ingot (bottom)
0142-1123/80/040147-06
C
Si
Mn
P
S
Cr
Ni
V
0A2
3 A0
0.50
0,030
0.014
8.15
0.47
0.03
0A3
3.39
0.49
0,029
0.006
8.13
0A7
0.03
0.42
3.38
0.49
0,030
0.005
8.14
0.46
0.03
0A4
3.40
0.48
0,030
0.006
8.12
0.47
0.03
$ 0 2 . 0 0 © IPC Business Press L i m i t e d 1 9 8 0
INT. J. F A T I G U E O c t o b e r 1 9 8 0
147
in the precracked specimenwas measured by fatigue loading with load control under sinusoidal tension at a frequency of 10 Hz. Two different values of stress intensity ratio (R = Krnm/Kmax,where Kmax and Kmin are the maximum and minimum stress intensities during each cycle) equal to 0.1 and 0.5 were used. The fatigue crack length was continuously monitored using a travelling microscope. The printing of Moir4 grid lines on the specimen surface and illuminating it with a stroboscope facilitated measurement of crack length. The average crack extension between two consecutive measurements was approximately 0.025 ram. All the fracture surfaces were examined with a scanning electron ndcroscope to determine the mode of fracture. R E S U L TS
From the chemical analysis data given in Table 1, it can be seen that the sulphur content of the steel is reduced considerably by electroslag refining while the contents of other alloying elements remain well within the limits. Further, the compositional variations from top to the bottom of the ingot were only marginal. The distribution of inclusions before and after ESR is shown in Fig. 1 in the form of number of inclusions per mm 2 versus inclusion size and the inclusion rating carried out according to ASTM E - 4 5 comparison method 6 is shown in Table 2. The ESR steel is free from large inclusions (>6/J.m) of all types. The microstructure of the as-quenched specimens consisted of martensite and specimens tempered at different temperatures had a tempered martensite structure. All these structures are essentially the same for both unrefined and ESR steel. The room temperature tensile properties of the refined and unrefined steel in different heat treatment conditions are given in Table 3. While the yield strength (YS) and ultimate tensile strength (UTS) are unaffected by ESR, ioa
En 52 steel
Table 2. Inclusion rating o f En 52 steel according t o A S T E 4 E45 comparison method Inclusion t y p e
Unrefined
Sulphides Alumina Silicates Oxides Volume fraction
ESR
Thin
Heavy
Thin
Heavy
25 20 25 25
15 10 15 15
05 05 05 05
-
(%)
Average inclusion spacing (/lm)
018
0.025
10
12
Table 3. Tensile data on En 52 steel before and after electroslag refi ning Treatment
1050°C OQ 600°C T 700°C T 800°C T
0~% YS (MPa)
UTS (MPa)
El (%)
UR
ESR
UR
860 810 716
870 815 724
1205 1220 3 1010 1020 16 960 960 26 920 930 28
ESR
UR
UR
ESR
6 23 36 48
6 28 45 50
12 37 62 70
the ductility is enhanced to some extent. In Fig. 2, the fatigue data (S/N curves) obtained on smooth specimens failed under rotating bend tests are given. These specimens were quenched in oil from 1050°C and tempered at 700°C for 2 h. The fatigue limit has increased by about 58% in ESR material compared with similarly treated unrefined steel. The fatigue crack growth rates (da/dN) are obtained from the crack length a versus number of cycles N data using a seven point polynomial method. 5 Since the cyclic load limits were kept constant, the applied z ~ (Kmax Kmin) increases as the crack length increases. In general,
En 52 steel ~
"~1
Refined
•
l0 = 3 L-- I I I I I I
f~
i
"6
I
ESR
Note: OQ -- oil quench; YS - yield stress; El -- elongation T -temper; UR -- unrefined; ESR - refined; RA -- reduction in area
"-i
"s
RA (%)
I I L.__ I I
10--
! 40( /
--
0
Unrefined
~ O
C
30(
I I--.. 1 L--- I I__.
ip--ESR
1 I
I 0
4
i
i
16 Inclusion size (/=rn) 12
L '--i--?-
20
24
4 x IO= 28
Fig. 1 Plot of number of inclusions per mm = versus inclusion size for En 52 steel before and after electroslag refining
I N T J. F A T I G U E O c t o b e r 1 9 8 0
IO4
I Io~ Number of cycles
Ioe
F i g 2 S/N curve for En 52 steel obtained on smooth rotating bending fatigue specimens oil quenched from 1050°C and tempered at 600°C
the fatigue crack growth data a r e represented in the form of da/dN versus AK as suggested by Paris. 7 The data obtained on unrefined and tel'reed En 52 steel in heat treated condition are shown in Figs 3 and 4. These curves exhibit three distinct regions. 8 At low AK (region I), the curve asymptotically approaches the so.ailed threshold ~ K value. The curve then changes to a linear region (region II) before turning asymptetic again as AK approaches the fracture toughness (KIc) of the material (region III). It is generally recoguised that region I and region III are sensitive to microstructure and load ratio effects; while the slope in stage II is insensitive to these variables. In the ESR steel, the (near) threshold stress intensity has increased and crack growth rates in region III have decreased whereas the slope in region II is unaffected in comparison with the unrefined steel. The fracture toughness ( / ~ ) and the Charpy V notch fracture energy obtained on the unrefined and the ESR steel are shown as a function of tempering temperature in Figs 5 and 6 respectively. These figures indicate that the ESR steel exhibits superior (nearly double) fracture toughness and impact properties in all conditions of heat treatment.
En 52 steel
A
16~ _
o•
.o
Ol •
~
•
• A A
•
,?
&
16~ _ ~
DISCUSSION
•
o O• 0
•
O& • O• • 0 • 0A o • •
O&
I050°C
oil
quench
800=C temperature
•
•
0 & o•
0
O•
•
&
Oa
• ~. •
•
•
O•
o"
i~ s _
oo
O&
A
~.
• •
•o A
•
•
•0
16 4 -
• t.
•
• •0
•
E
•
R=0.5
R=0.1
• •
Unrefined
o
Refined
•
• • •
A
D u r i n g e l e c t r o s l a g r e f i n i n g o f steels, t h e f o l l o w i n g c h a n g e s
Occur: (i)
The large sized non-metallic inclusions are removed and the finer inclusions are redistributed more uniformly. (ii) Sulphur content of the steel is reduced appreciably.
(iii)
I0
I
I
I
I
20
40
60
80
I
I
100 120
Alternating stress intensity, Z~K (MPoq-~) Fatigue crack growth rates for En 52 steel as a function of Fig. 4 A K at R = 0.1 and R = 0.5
Chemical homogeneit7 is achieved. En 52 steel
En 52 steel Austenifizir~ ~ a t u r e o
iO-~ __ &
•
O
A
O
•
: 1050°C
• •
•
O & &
oO & • • O O• 0 • 0 • •
&O O&
E
AO& Z
&O•
•
• •
•
:S
0 • 0 &A •
{
2 =-
%•e
& 0
ao • • Ao• • &o• •
'
I0"
R=0.1 1050eC oil quench Tempering tomper~uN 650=C
~0
,, o • • o "O • •
._~ -(
Io
&O & • O
Unrefined
800°C 0
Refined
Refined
•
I
•
I0--
E)
•
Umef~
t 20
I 40
I 60
I 80
I R)o
I 120
AlterrN]ting stress intensity, LIK (MPoq-~) Fig. 3 Fatigue crack growth rates for En 52 steel as a function o f AK at R = 0.1
o As Q
I 6O0
i
70|0
Tempering t e m ~
I ("C)
°1
Fig. 5 Fracture toughness o f En 52 steel as a function of tempering temperature
I N T . J. F A T I G U E
October 1980
149
of static fractures, including cleavage, intergranular and fibrous fracture. 14 The SEM fractographs taken on specimens of En 52 steel before and after refining, treated at 1050°C, oil quench and 800°C temper, are shown in Fig. 7. The unrefined specimen has failed by quasi-cleavage and an intergranular mode whereas refined steel has failed by a fibrous quasi-cleavage mode. The fatigue crack growth rate in region II is reported to be insensitive to microstructure and load ratio effects. Paris 7 has shown that the slope in this region follows the equation:
En 52 steel
3C
-D
zc ~t
(aa/dN) = C (~)m
(1)
where C is a m a t e r i a l c o n s t a n t a n d m is the slope o f the
Z > 0
IC
t
,
600 0
i
Refined
•
Unrefined
o
I
700
I
I
800
1
850
Temperin~ temperature (°C)
growth rate curve. The value of m remained unchanged in the unrefined and refined steel and is also not affected by heat treatment. The value of m for En 52 steel lies in the range 2.4-3.0. The crack growth in region II is controlled by the amount of crack opening in each cycle and gives rise to striations. The effect of nonmetallic inclusions on tensile ductility and fracture toughness is discussed by Schwalbe 9 in terms of well known void formation mechanisms. Due to the difference in the deformability of matrix and the inclusion, decohesion occurs at the interface leading to void
Fig. 6 Charpy fracture energy of En 52 steel as a function of tempering temperature
The electroslag refining has completely eliminated inclusions with sizes larger than 6 pm and reduced the number of those with sizes between 4-6/n-n. Furthermore, the lowering of sulphur content to very low values (0.005%) reduces the number and size of sulphide inclusions. The mechanical properties, particularly fatigue strength, fracture toughness and tensile ductility, are very sensitive to the presence of nonmetallic inclusions. 9 The fatigue strength depends to a large extent on the crack initiation at the debonded inclusion/matrix interface. The nonmetallic inclusions act as stress raisers and initiate cracks. It has been shown 1° that in a given material there is a critical inclusion size (larger than 20/~'n) which will be able to nucleate a fatigue crack. Brooksband and Andrews n have estimated the residual stress distribution surrounding an inclusion and have suggested that such tessellated stresses could be responsible for matrix yielding and crack initiation during fatigue. Lindley and Richards 12 have shown that there is no significant effect of particles on the fatigue crack propagation rates by alteration of flow parameters below general yield. Shih and Araki 13 have observed that fatigue growth rates associated with a dimple mechanism decreased with increasing cleanliness of the steel. They have also noticed that inclusions are more dangerous to fatigue crack growth at lower tempering temperatures. It is generally recognised that regions I and III of the fatigue crack growth rate versus z~K curve a;'e sensitive to microstructure and load ratio effects. 14 For similarly heat-treated steel, the increase in the threshold stress intensity factor and the decrease in fatigue crack growth rate in region III that occurred in the ESR steel (Figs 3 and 4) can be attributed to the removal of nonmetallic inclusions. In region I, the crack growth rates are slow and the threshold values could not be estimated accurately. In region III where Kmax approaches KIc, the growth rate is rapid and failure occurs by modes that are characteristic
150
I N T . J. F A T I G U E O c t o b e r 1 9 8 0
~ii~ .........
Fig. 7 (a) SEM fractograph obtained on the fracture surface of electroslag refined En 52 steel in region III of fatigue crack growth showing mixed mode fracture consisting of ductile and quasi-cleavage features; and (b) SEM fractograph of unrefined En 52 steel in region III of fatigue crack growth showing mixed mode fracture consisting of intergranular and cleavage features
formation. The void formation can also occur due to the fracture of inclusions ahead of the crack due to large plastic strains occurring in the plastic zone. The ductile fracture proceeds by a process of void coalescence. Thus the ductility depends on the volume fraction, kind, size and distribution of inclusions. The sulphide inclusions are most deleterious as far as the fracture toughness is concerned. Birkle et a115 have shown that Kic increases with decrease in sulphur content in the steel. Further, the Mn/S ratio decides the shape of the MnS inclusions and at higher ratios the inclusions prefer a globular shape to a plate.like shape. 16 Gladman 17 has shown that the plate-shaped inclusions are more dangerous than equiaxed ones as far as fracture toughness is concerned. The ductility is also affected by the size of the inclusions, since the plastic strain necessary to fracture an inclusion increases with decreasing inclusion size. Schwalbe 9 has shown that the particle size reduction causes an increase in fracture toughness. Whereas the particle size determines the onset of voids, the distance between the particles controls the void growth and crack propagation. The electroslag refining has eliminated the large size inclusions (>6 pro) at which void formation would readily occur. Also the volume fraction of inclusions is reduced by electroslag refining. Moreover, the reduction in sulphur content has increased the Mn/S ratio which helps in the formation of globular MnS inclusions. All these factors contribute towards improving the tensile ductility, fracture toughness and Charpy fracture energy in the ESR steel. The critical defect size can be calculated from the Kic values obtained on the refined and unrefined steel following the procedure suggested by Kiessling and Nordberg. l° This gives the maximum defect size that the material can tolerate with the stress intensity factor not exceeding the Kic value. For a disc shaped crack, the critical defect size is given by A = 1.57 k 2 (Kic/Oy) 2
(2)
and for an eMpsoidal surface crack, the critical defect size ( d e p t h ) is B = 0.S k 2 (Kic/Oy)2
(S)
where k = factor of safety, generally taken as i. The calculated values of the diameter of the inner spherical defect A and the depth of the ellipsoidal surface crack B at o = ay (Oy = yield stress) are presented in Table 4. Electroslag refining has increased the critical defect size in all conditions of heat-treatment. With a view to comparing the fracture toughnesses of the unrefined and the ESR steel, the KQ values are plotted against their individual UTS values in Fig. 8. For a given heat treatment the fracture toughness values in the ESR steel are the highest. The scanning electron micrographs obtained with the unrefined and refined steel are shown in Fig. 9. The refined
En 52 steel
10C
\
:'2.
9£
8C
v
o
! °° 40 3C
2C
I
19 00
I
900
I
1000
II00
UTS (MPo) Fig. 8
Fracture toughness versus tensile strength in En 52 steel
steel exhibits fibrous and quasi~leavage t y p e o f fracture whereas the unrefined steel has failed by intergranular and quasi-cleavage fracture. The large ~zed inclusions in the unrefined steel provide sites f o r early cleavage and the high sulphur content results in segregation o f sulphur at the grain boundaries 18 thus causing the l o w values o f fracture
toughness.
CONCL USIONS Electroslag refining of En 52 steel decreases the fatigue crack growth rates and increases the fatigue strength, fracture toughness, Charpy fracture energy and tensile ductility. The improvement in these mechanical properties can be attributed to the removal of nonmetallic inclusions and the reduction in sulphur content. Furthermore, the fatigue crack growth rates in region I and region III decrease in ESR steel.
REFERENCES 1.
Holzgruber, W. 'Possibilities and limitations to influence the structure of ESR ingots and properties of ESR products', Proc Fifth Int Syrnp on Electros/ag and Other Special Melting Technology (Pittsburgh, USA 1975) pp 7 9 - 8 0
Table 4. Critical defect size at o = ey for the En 5 2 steel
Refined
Unrefined Treatment 1050°C OQ 600 ° C T 700° C T 800 ° C T
YS (MPa)
KQ
(MPa~/m)
A* (ram)
B** (mm)
YS (MPa)
KQ (MPa~/m)
A (ram)
B (ram)
860 810 716
32 42 56
2.1 4.2 9.6
0.40 0.80 1.80
870
67
9.3
1.8
815
90
19.0
3.6
720
97
28.0
5.4
*A -- Diameter of inner spherical defect **B -- Depth of ellipsoid surface crack
INT. J. F A T I G U E October 1980
151
5.
6.
7.
8.
9.
10.
11.
12.
13.
14. i
15.
16.
17. Fig. 9 (a) SEM fractograph of the fracture toughness specimen of electroslag refined En 52 steel exhibiting fibrous and quasi-cleavage fracture; and (b) SEM fractograph of the fracture toughness specimen of unrefined En 52 steel showing intergranular and cleavage fracture 2.
3.
4.
152
Hebsur, M. G., Abraham, K. P. and Prasad, Y. V. R. K. 'Hot working characteristics of electroslag refined En 52 steel: a hot torsion study', Metals Techno/ (to be published) Woolman, J. and Mottram, R.A. The Mechanica/ and Physical Properties o f the British Standard En Steels, Vol 3 (Pergamon Press, England 1968) pp 180-185 'Standard Test Method for Plane Strain Fracture Toughness of Metallic Materials', Annual Book o f ASTM Standards, Part 10, Designation £ 3 9 9 - 7 4 (ASTM, 1978)
I N T . J. F A T I G U E O c t o b e r 1980
18.
"Tentative Method for Constant-Load-Amplitude Fatigue Crack Growth Rates Above 10 -s rnm/cycle', Annual Book o f ASTM Standards, Part 10, Designation E647--78T (ASTM, 1978) 'Recommended Practice for Determining the Inclusion Content of Steel', Annual Book o f ASTM Standards, Part 1 I, Designation E 4 5 - 7 6 (ASTM, 1976) Paris, P. C. "The fracture mechanics approach to fatigue', Proc lOth Sagamore Army Materials Research Conference (Syracuse University Press, USA, 1964) pp 107-132 Imhof, E. J. and Barsom, J. M. 'Fatigue and corrosion fatigue crack growth of 4340 steel at various yield strengths',ASTM STP 536 (1973) pp 183--205 Schwalbe, K. H. 'On the influence of microstructure on crack propagation mechanisms and fracture toughness of metallic materials', Engng Fracture Mech 9 (1977) pp 795--832 Keisling, R. and Nordberg, H. 'Influence of inclusions on mechanical properties of steel', Int Conf on Production and Application o f Clean Steels (BalatonfiJred, Hungary, 1971) pp 179--185 Brooksbank, D. and Andrews, K.W. 'Stress fields around inclusions and thei r relation to mechanical properties', Int Conf on Production and Application o f Clean Steels (Be latonfUred, Hungary, 1971) pp 186--195 Lindley,T.C.and Richards, C.E. "lnfluenceof microstructure on fatigue crack propagation in steels', in Effects o f Second Phase Particles on the Mechanical Properties o f Steel (Iron and Steel Institute, London, 1971) pp 119--123 Shih,T. Y.and Araki,T. 'The effect of nonmetallic inclusions and microstructure on the fatigue crack initiation and propagation in high strength carbon steels', Trans ISI Japan 13 (1973) pp 11--19 Parker, E. R. 'Inter-relation of composition, transformation kinetics, morphology and mechanical properties of alloy steels', Metal Trans 84 (1977)pp 1025--1042 Birkle, A. J., Wei, R. P. and Pellissier, G. E. 'Analysis of plane strain fracture in 0.45C - Ni--Cr--Mo steel', Trans ASM 59 (1966) pp 9 8 1 - 9 9 0 McTegart, W. J. and Gittins, A. 'The role of sulphides in the hot workability of steels', in Sulphide Inclusions In Steels, ASM 798--211 (1974) Gladman, T., Holmes, B. and Mclvor, I. D., 'Effects of second phase particles on strength, toughness and ductility', in Effects o f Second Phase Particles on the Mechanical Properties o f Steel (Iron and Steel Institute, London, 1971) pp 68--78 Hebsur, M. G., Abraham K. P. and Prasad, Y. V. R. K. 'Effect of electroslag refining on the fracture toughness and fatigue crack propagation rates in heat-treated AlSl 4340 steel', Engng Fracture Mech (to be published)
AUTHORS The authors are all with the Department of Metallurgy, Indian Institute of Science, Bangalore 560 012, India. Enquiries should be addressed to Dr Prasad in the first instance.