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Fracture toughness of alloy steels used in turbo-generator components
.Eng;ne~r;ng
Frocturc
Mechanics,
1970, Vol. 2, pp. 125- 143.
Pergamon Press.
Printed in Great Britdln
FRACTURE TOUGHNESS OF ALLOY STEELS TURBO-GENERATOR COMPONENTS Central Metallurgical Laboratories,
USED
IN
D. V. THORNTON English Electric Co, Ltd., Whetstone, Leicester, England
Abstract-The plane strain fracture toughness of various alloy steels used as forgings in turbo-generators has been determined. It has been shown that equivalent values of fracture toughness are obtained for these materials using both linear elastic and general yield fracture mechanics. The values obtained have been used to assess the rapid fracture susceptibility of various components from single defects. The effect of metallurgical factors on the fracture mode and toughness is illustrated.
NOTATION K/c FATT 6 eU T a
fl’y E G ” K8
plane strain fracture toughness determined by linear elastic fracture mechanics, ksi 6. fracture appearance transition temperature, “C crack opening displacement (COD), in X 10es yield strain specimen thickness, in. notch depth, in. yield stress, t.s.i. elastic modulus, t.s.i. strain energy release rate, in. lb/in’ Poisson’s ratio. stress intensity determined by general yield fracture mechanics. based on measured ksi 6 as K8 but based on corrected values, ksi 6 applied stress, t.s.i. distance from crack tip, in. fractional distance from crack tip compared to notch depth relative displacement stress intensification Charpy value at test temperature, ft-lb nominal fibre stress, t.s.i. crack depth, in. thickness, in.
values,
INTRODUCTION THE SIGNIFICANCE
of defects and the elimination of the hazard of rapid fracture, which in large rotating machinery causes catastrophic failure[ 11, have been a problem to the turbine builder for a considerable time. The present method of assessing the susceptibility of large forgings to fast, unstable fracture is based largely on routine quality control tests, rules of thumb and experience. It is obviously highly desirable to develop a quantitative approach for both acceptance and design of large turbogenerator components. The recent expansion of the aerospace industry has involved the use of high strength metallic materials and has emphasised the need for a quantitative assessment of rapid fracture. This has led to a renewed interest in linear elastic fracture mechanics originally developed by Griffith in 1920 to explain the fracture of glass. Recent advances in both the theoretical and practical aspects have now brought fracture mechanics to a stage at which some of the problems in the assessment of large forgings in medium strength steels used in turbo-generators can now be dealt with in a more scientific manner. 125
126
D. V. THORNTON
The major advantage is that quantitative information is provided for design, material assessment and inspection acceptance purposes, which can be used where service experience with new materials and new designs is not available. By consideration of the design stress and configuration, the anticipated defect size and the fracture toughness, the safety of the component can now be judged more precisely. The components considered include discs and rotors, alternator end rings, bolts and L P. blades. Turbo-generators usually operate at 3000 revlmin and the major stressing of rotating parts is by centrifugal forces. From the rapid fracture aspect, the low pressure cylinder and alternator are the most critical regions since the material strength and component size are high and the operating temperature relatively low. These factors increase the susceptibility to rapid fracture which may initiate at such stress concentrations as inherent material faults, design geometry and cracks which have grown in service by fatigue or stress corrosion. During the past ten years the size of single machine 3000 rev/min turbo-generator units has increased from 100 to 500 MW. Figure 1, which shows a 500 MW unit being assembled, illustrates the large size of these units. Typical materials and component sizes used in the 100 MW and 500 MW machines are shown in Table 1. This indicates that the design of increased output units has led to the use of higher strength materials in considerably larger section sizes, both of these factors increasing the likelihood of rapid fracture. Material assessment has been based on standard quality control testing. In spite of the progression to higher strength levels, developments in steel composition and treatment have enabled the tensile ductility to be maintained and the impact transition properties to be enhanced. Increasing use of modern non-destructive testing methods, particularly ultrasonic detection, has enabled the quality of large forgings to be determined and a high acceptance standard to be established[21. With the cooperation of the steel makers and forgemasters, materials and techniques have been developed whereby components of a high standard are consistently produced. The unfortunate occurrences experienced in the United States in 1953 and 1954, where several catastrophic failures of turbo-generators occurred [ll, have so far been avoided in the United Kingdom. However, as service conditions become more severe with the development of larger, higher efficiency machines, the need for quantitative assessment of materials and designs becomes more acute. TEST TECHNIQUES With the large forgings used in turbo-machinery the thickness is such that the fracture mode would be predominantly plane strain (e.g. large discs are 22 in. thick). Thus the objective is to measure the plane strain fracture toughness (K,,), which is the lowest energy fracture mode at the lowest operating temperature. At the last stage of an L.P. turbine, this is almost ambient temperature. Other components, while normally operating at higher temperatures, may be stressed under start-up conditions at a temperature near to ambient. One of the test techniques employed in this work to measure K,c is based on linear elastic fracture mechanics principles developed by Irwin [3,4]. The data have been determined using the testing parameters current in the ASTM literature, bend testing being employed for all tests. The notch configuration comprised a parallel sided milled slot terminating in a spark machined slot, O-1 in. deep and OXKK3 in wide which was subsequently lengthened by a fatigue crack. Testing of the medium strength, high
Fig. 1. 500 MW steam turbine assembly.
Fig. 5. Fracture
Tested
path following
the plastic
hinge in a
cold
worked
austenitic
at :+2o*c
-2ooc
-60°C Fig. 6. 34% Ni
Cr
. MO V fracture
surfaces.
steel.
(b)
(cl
Fig. 7. Typical fracture modes (a) Transgranular quasi-cleavage fracture Intergranular fracture X 150; (c) Mixed trans- and inter-granular fracture
480; (b) 150.
X X
Ib)
Fig. 8. Primary
carbides
in
12%Cr
Mo .V Nb steels. (a) Specimen
200; (c) Specimen
A 35 X 200.
A 40 X 200; (b) Specimen
A 4
IX
L.P. disc. L.P. shaft. I.P. rotor. Alternator rotor. End bell. L.P. blade.
Component
13 thick 25 dia. 37 dia. 3 5 dia. 3 thick 27 Aem foil
Size (in.) 39 22 30 19 47,5 28
Carbon 3%Cr,Mo 8% Mn 8% Ni Mn.Ni
~~inimum 0.2% PS (t.s.i.)
3%Cr.Mo Carbon
Steel
100 MW unit
464 55 diq. dia. 34 thick 36 Aerofoil
22 thick 29 dia.
Size (in.)
2%Ni.Cr.Mo.V l%Cr.Mo*V 18%Mn .Cr 12%Cr.Mo.V
3+%Ni.Cr.Mo.V 2t%Ni.Cr*Mo*V
Steel
500 MW unit
Table 1. Materials used for various components of a 100 MW and 500 MW unit
;: 57 50
50 42
Minimum 0.2% PS (t.s.i.).
128
D. V. THORNTON
toughness forgings used as discs, rotors and large L.P. blades has necessitated the use of large specimens in order to obtain plane strain failure, where sections up to 6 in. square have been used. The majority of specimens exhibit a maximum load failure. Destructive testing of forgings to measure typical fracture toughness values is expensive, both in obtaining representative material and in manufacture and testing of the specimens. In order to establish the range of toughness for a particular material and to carry out quality control tests based on fracture mechanics principles or correlated with fracture mechanics results, it would be necessary to use samples taken from production forgings. In this case the material available for testing would be severely limited and, therefore, a small scale test is required for this purpose. In view of this, general yield fracture mechanics [5[ has been investigated. The displacement at a crack tip is assumed to describe the local stress and strain conditions and it is assumed that fracture will occur at a critical displacement dependent on material, temperature, strain rate and through thickness stress. The crack opening displacement (COD) has been measured on the large specimens used for Klc determination and also on 0.4 in square bend specimens machined from these large specimens. The displacement was measured using a strain gauged, thin strip of copper-beryllium inserted to the bottom of the 0.008 in wide spark machined slot. A torque was applied to the gauge prior to testing and the change in signal due to decrease in torque as the crack opens on loading was autographically recorded. Each gauge was calibrated, using a micrometer arrangement, to determine the displacement corresponding to a particular change in signal. On the K,(. specimens the autographic trace served as a record of the specimen compliance and a measure of the COD. The same technique was used to measure COD on the 0.4 in. specimens. Initial tests on O-4 in. square specimens of a 3% Cr + MO . V disc steel indicated that the use of 10 per cent side grooves aided in the detection of first cracking i.e. the critical COD value. A limited number of tests were carried out with specimens containing slots of different width. Specimens were prepared with 0.014 in., 0.009 in. and 0.006 in. wide slots and also an 0.009in. slot terminating in a fatigue crack. Similar COD values were obtained on all except the widest slot, Fig. 2, where the COD value measured was larger and showed considerable scatter. x Tested n Tested
0
I
2
3
4
5
Notch
6
7
width
6
9
at at
IO
2O’C 60°C
II
12
13
I4
, in. x 10e3
Fig. 2. The effect of notch width on COD of a 3% Cr. MO . V steel.
15
Alloy steels in turbo-generator
components
129
In subsequent tests a specimen containing 10 per cent side grooves and having a notch depth of 20 per cent beam depth, with the slot terminating in a fatigue crack, was used. Summing up, the object of the work was to establish typical fracture toughness values for various materials and to examine the COD tests as a means of quality control testing. Testing variables have not been systematically investigated, the testing conditions being standardised where possible. Study of metallurgical variables was confined to a few special cases. RESULTS Linear elastic fracture
AND DISCUSSION
mechanics
The chemical composition, source of material and room temperature mechanical properties of the materials tested are shown in Table 2. The results of the large bend tests are given in Table 3. In cases where post-yield fractures have been obtained and the compliance and load deflection curve deviated from linearity, two values are quoted in the Klc column. These are a minimum value calculated at a failure stress equivalent to uniaxial yield stress, and the value calculated at the actual failure stress (shown in brackets). The recent tentative recommended testing practice for fracture toughness of the ASTM E-24 sub-committee I has proposed that for valid KIC measurement both the specimen thickness and crack length should be greater than 2.5 (K,c/cr,)2. It is felt that this requirement is conservative and that there may be a relaxation of this as more data becomes available. In the present tests a square section with a notch 25 per cent of the beam depth has been used and although the crack length is generally too short, the thickness in most cases is sufficient to meet this requirement. It is considered, therefore, that the data is sufficiently accurate for the present purpose of estimating the tolerance of a material to defects. General yieldfracture
mechanics
Results obtained to date using the small COD specimens are given in Fig. 3. Although the small COD specimens failed with a flat fracture and rapid crack propagation, the majority also failed after general yield. The conditions of constraint were not similar to those of the large bend specimens and to the actual components where plane strain elastic failure is anticipated. These COD results, therefore, are not expected to agree with Klc values determined at the same temperature. There is a large amount of evidence in the literature [6-81 that K,c of low alloy steels steels shows a rapid increase over a certain temperature range. Wells [51 and Burdekin [9] have shown that the COD of certain steels show a transition with increase in temperature similar to the well known impact transition. It is thought that a quality control test for medium strength low alloy steel forgings may possibly be established by correlation of the COD transition on standard specimens with the KIC transition. A correlation of Klc and Charpy FATT has been shown for materials used as turbine forgings and has been incorporated in a fracture analysis diagram[8]. The COD vs. temperature curves for the O-4 in. square specimens in the present work show a transitional behaviour, but as can be seen in Fig. 3. the temperature range over which this occurs is not similar to the respective Charpy FATT. The COD value at the upper shelf agrees
18
17
16
14b) 14c) 15
14a)
13
12
9 10 11
8b
la lb 2 3 4 5 6a 6b 7 8a
Material No.
tested-chemical
Source
properties
80.3
78.6
82.2
87.9
70.8
69.3
60.1 62.9
86.6 78.2 72.4
80.5 77.9 73.9 70.3
70.7
91.5
63.0
55.0 82.0
57 50 17 31
60.2 59.3 54.3 92.4
51.6 50.8 48.2 82.6
11
15
60
58
61 66 68
40
53
50 14 53 62 59 18 48 50 15
59.3 67.9 62.0 63.1 50.8 50.8 49.6 47.8 60.0
49.9 57.4 54.5 54.5 41.4 39.9 41.8 34.0 50.8
0.2% PS (t.s.i.)
and mechanical
Disc 3.5Cr,0.5Mo, O.lV. Disc 3.5Cr, 0.5Mo,O.lV. 2.4Ni, OGSCr, 0.5Mo,O.lV Disc 3.3Ni, 1.6Cr. 0~6Mo,O~l5V Disc Rotor l.OCr, 0.7Mo,O.32V, 0.7Ni Rotor 1.2Cr, l.lMo, 0.37V, 0.6Ni l.OCr, 0.7Mo,O.39V, 0.7Ni, 0.006B Rotor l.OCr, 0.7Mo,O.39V, 0.7Ni, 0.006B Rotor 11 .lCr, 0.6Mo,O.27V, 0.21Nb, 0.07N Disc 104 in. billet 9.9Cr,0.7Mo, 0.3OV. 0.22Nb outside O.l5C, 9.9Cr, 0.7Mo,O.3OV, 0.22Nb centre End bell 0.6C, 7.0Mn, 9.ONi, 4,OCr End bell 0.5C, 18.8Mn. 4.4Cr, O.lNi O.l4C, ll.lCr,0.8Ni,0.6Mo, Bar (5 in. dia.) 0.3V, 0.4Nb O.O4C, 14.3Cr, 5.4Ni, 1.7M0, I .7Cu Bar 0.8Nb (3 X 2: in.) Bar Ti, 2.3A1, 11.2Sn. 4.6Mo (2 X 2 in) Bar O.O5C, 14.2Cr. 5.3Ni, 1.6Mo,1.6Cu, O+Ib (3 X 24 in.) (3 X 2: in.) O.OSC, l4.2Cr, 5,3Ni, 1.6M0, 1,6Cu, 0.4Nb (3 X 2tin.) O.O5C, 14.2Cr, .5,3Ni, 1.6M0, 1.6Cu,0.4Nb O.llC, 11.4Cr,2.6Ni, 1.3Mo,O.l3V, Bar (6s x 26 in.) 0.2Nb, 0.06N O.l2C, 12.5Cr, 2,5Ni, 1.8Mo,O.3V, Bar O.OSN (66 X 2t in.) O.O4C, 12.6Cr, 4,4Ni, 0.3Ti, O.INb Plate (3 X 1 in.) 14.2Co,O.O7Al Plate O.O3C, 12,5Ni, 4,9Cr, 2.9Mo,0.2Ti, (36 X 1 in.) 0.5Al
0.35C, 0.35C, 0.28C. 0,35C, 0.3C, 0.3C, 0.2C, 0*2C, O.l6C, O.lSC,
Composition (Wt%)
Table 2. Materials
+6 -
-23
+60 -5 -25
(+ 140)
+ 120
+73 +72 _ _
+ 140
+41 -1 +69 + 123 +s5 _
+58 _
50% FATT (“C)
Alloy steels in turbo-generator
, OC
Temperature
k;[~~“;T,T~
-00
,
-40
0
40
Temperature
Temperature
131
components
:
120
80 ,OC
,
160
, OC
Fig. 3. The effect of test temperature on COD.
with that calculated according to Wells[5] for the transition from plane strain to plane stress behaviour, 6 = 27r e,T.
(1)
Use of the expression given by Burdekin[9],
6
=
ha -log
77
(
5+1
>
,
(2)
predicts the COD value at which plane strain conditions will not be dominant. This value is found to be in the transition range. These calculated values are shown in Fig. 3., together with the COD measured in large specimens. From this it would appear that plane strain COD values can be obtained at low temperature on small specimens which are similar to those obtained on larger bars. (Subsequently these will be shown to agree with Klc results.) However plane stress fracture begins to dominate as the transition range is approached. Further tests are in progress on other materials and these may show a more definite pattern of behaviour.
2 2
3 3
3
4
5 5 5
6a
6b
7
8a 8b
A21 A22
A24 A28
A29
A23
A33 A38 A39
A25
A34
A35
A40 A41
9 9
10
A30 A31
2281
End hell materials
+20 + 100
2 2
Al5 Al7
+20
+20 +20
+20 +20
+20
+20
+35
+20 +20 +20
+20
-30
+20 -60
+40 -75
+20 -65 -20 -30 +20
Materialt
Test temp. (“C)
Disc and rotor materials A8 la Al3 la Al4 la A27 la A26 lb
Specimen. ident.
82.6
48.2 48.2
51.6 50.8
50.8
34.0
41.8
39.9 39.9 39.9
41.4
54-5
54.5 58.8
52.8 58.5
54.5 51.0
49.9 57.0 52.2 50.3 57.4
0.2% PS (a,) t.s.i.
-
18 18
9
9
6 6 6
15
18
48 13
43 8
34 77
20 10 13 12
c,.+ ftlb
K,C
0.87
0.88 o-95
I.16 0.81
0.70
1.28
0.50
0.64 1.11 l-20
0.74
‘a
>
116
79 84
92 (107) 71
56
81 (104)
57
65 63 71
69
> 133 (177)
> 163 (233) 83
I .43 0.56 1.33
> 127 (169) 58
90 > 140 (215)
> 100 (145) 76 109 > 108 (126) 60
ksifi
1.33 0.48
0.60 1.53
1.44 0.65 1.04 1.17 0.49
rr,/Ou§
(0.6)
1.5
I.1 0.9
I.1 (0.9) 1.9
2.3
1.0 (0.6)
2.8
2.4 1.6 0.6
1.5
O-8 (0.4)
0.6 (0.3) 2.2
0.6 (0.4) 3.7
I .9 0.7 (0.3)
3.4
:I;
I.6
0.7 (0.4)
($
Table 3. Large bend test results
2.5
5.0 4.5
3.6 (2.7) 5.9
8.3
3,l (2.2)
10.6
10.0 3.5 2.8
5.9
3.5 (2.0)
2.8 ( 1.4) 10.1
2.6 (1.5) 12.7
8.3 3.3 t1.4)
3.3 (1.6) 14 3.0 2.9 (2.1) 12.1
$$
(in
6
IO-:‘)
IO.0
9.7 8.2
6.4 4.2
1.6
6.5
2.6 1.8 4-o
2.3
8.1
23 .O
1.9
10.3
1.1 -
6.1 2.0 3.0 5.3 2.4
X
247
186 171
160 130
77
127
77 73 108
84
180
305
200 91
67
150 92 116 140 101
K6 ksifi
-
2.13
2.35 2,04
1.54
1.58
1.89
1.20 1.29
1.48 1.83
1.07
1.14 1.38
1.22
0.79 1.09 1.16
1.04
1.22 1.18 1.21 1.53
0.93
1.13 -
0.72
1.08
0.96 -
0.81 0.99 0.95 1.08
1.05
K&K,>
1.02
1.31 -
1.18 1.56
0.75 -
1.11 1.68
1.06
I .03 1.21
KslK,,l
w N
t20
13
14a 14b 14c
15
16
17 18
A16
Al8 At9 A20
A58
A37
B94 895
tMaterials referenced as in Table 2. tC,.- Charpy value at test temperature. §o!--nominal fibre stress at fracture. “KeIK,c correction- see Appendix.
+20 +20
+20
+20
+20 +20 +20
+20
materials 12
Biading Al2
+20 +20 +20
11 II 11
material
Bolting B91 B92 B90
78.6 80.3
62.9
60.1
77.9 73.9 70‘3
71.0
82.0
55*0 55.0 55.0
24 56
28
61
26 60 87
-
5
9 9 9
0.34 1.27
0.83
I.55
0.66 1.33 163
0.29
0.28
0.65 0.56 0.84 .-
52 )z 162 (206)
152
> 173 (257)
107 r 155 (207) > 143 (234)
36
47
53 45 56
IO.8 1.4 (0.9)
1.4
0.7 (0.7)
(0.5) O-8 (0.3)
::f;
9-6
2.8
2.4 3.8 1.1
12.3 1.3 (0.8)
2.2
1.6 (1.5)
6.7 2.9 (1.6) 3.0 (1.1)
40.2
38.3
8.1 11.3 4.8
1.8 8.5
8.8
21.5
10.0 12.8 15.4
2.1
2.8
2.5 1.8 1.7
102 225
202
300
240 265 283
71
131
100 85 83
I.09
1.98
1.33
1.16
2.24 1.28 I.21
1.97
2.79
1.89 1.89 1.48
0.90 0.80
0.98
1.12
1.37 099 1.01
0‘74
1.13
1so2 0.89 0.95
E
% 3 a $
;
3 =: !: z 8 t; E’
D. V. THORNTON
134
Relation of COD and Krc values
Where both COD and Klc values were obtained in the same large bend test the theoretical relationship proposed by Wells has been used to calculate a stress intensity from COD Su,=G
(3) (4)
This is shown as KS in Table 3 and the values of Ks/KIc in Fig. 4(a). The figures obtained for the initial tests on 3% Cr. MO . V disc steel (material l(a) in Table 3) showed good agreement but the histogram, Fig. 4(a), shows that the COD generally over-estimates Klc. The instrumentation is considered satisfactory even for accurate measurement of the small COD values found in these tests. Therefore, other factors have been examined, namely the stress value to be used in the equation and errors due 9 0 7
I
6
(a)
;
3
l--l
4 3 2
G
5lr
0.4
0.6
0.6
I.0
16
12
l-h 2.2
16
2.4
2-6
2-8
3.0
KS ‘KIC
Il. IO . 9. 0-
( b)
7. ; g
6.
z
543-
n
2I .
04
06
06
10
!
! ! !--j
14
I.6
16
20
, , , , , 2.2
24
26
26
30
hdK,c
Fig. 4. Comparison of fracture mechanics (KS ics (K,,.). (a) based on rected
stress intensity derived from general yield and KS,(.) and linear elastic fracture mechanmeasured values; (b) based on values corby Marcal-King analysis.
Alloy steels in turbo-generator
components
135
to position of displacement measurement. The yield value used to calculate K8 was the uniaxial yield strength whereas the value across the yield zone required modification for stress intensification under the severe biaxial stress conditions at the notch tip. The experimental technique employed measures the COD in the mid-thickness of the specimen at the base of the machined notch. This point has been up to 0.4 in. from the tip of the fatigue crack, resulting in an erroneous result. The results of an analysis due to Marcal and King [ lo] have been used to correct the results for measurement position and stress intensification. The method used is shown in Appendix I. The resultant ordering of the results is shown in the histogram, Fig. 4(b). The three results which have corrected KS Ic/K,c ratios greater than 1.4 are cold worked austenitic steels. These fractures were significantly different from all the others in that the fracture path appears to follow the plastic hinge, Fig. 5, and not to occur normal to the specimen. The reason for this may be metallurgical in that heavy carbide networks were present at the austenite grain boundaries and the material is susceptible to stress corrosion. This may result in slow crack growth during the test and, therefore, errors in the application of the correction factors. These results show that excellent agreement between general yield and linear elastic fracture mechanics can be obtained for plane strain fracture provided that the correct COD value and stress are used in (3). A significant difference between fracture toughness values calculated by the two approaches has previously been reported by Knott[l l] for an L.P. disc steel, and possible reasons for this discrepancy have been suggested[ 11, 121. A sample of the 3% Cr. MO * V bar used by Knott in his investigation was obtained and heat treated in a similar manner to the 61.5 t.s.i. 0.2 per cent proof strength material used in[ll]. Fatigue cracked specimens were tested in three point bending and showed an elastic behaviour. Klc values calculated were 63 and 72 ksi X&. The actual COD was measured as 0.49 and O-53 x lo-” in. respectively from which calculated Ksrc values of 70 and 74 ksi fi were obtained using the Marcal-King corrections. Again very good agreement was obtained. In these tests the COD was measured during the test and therefore includes both the elastic and local plastic strain. Knott used double notched specimens and measured the displacement after the test and therefore any elastic contribution would be neglected. This is likely to lead to inaccuracies and low K*,C as in [ 111, since the total displacement is only 0.1 X 10e3 in. at the crack tip. Failure mode
Examination of the fracture faces by visual, optical and scanning electron microscopy has yielded information on the fracture characteristics of certain alloys. In general the toughness of the material is reflected in the degree of roughness of the fracture surface. Figure 6 shows 34% Ni . Cr . MO . V specimens tested at three temperatures. The room temperature specimen has a very rough surface showing the direction of propagation, whereas at low temperature where the material has low toughness the fracture is flat and featureless. However, all three specimens appeared to have fractured by a similar mode, being transcrystalline quasi-cleavage. Scanning electron microscopy has indicated three modes of fast fracture, transgranular, intergranular and a combination of these modes. Typical examples are shown in Fig. 7. The majority of fractures have been found to be transgranular quasi-cleavage, Fig. 7(a), the relative amounts of cleavage and ductile tearing depending on material and test temperature. In some specimens slow crack growth could be seen as ductile E.F.M. Vol. 2 No. 2-D
136
D. V. THORNTON
tears. Figure 7(b) shows the type of failure which occurred in all 3% Cr. MO - V specimens. This intergranular mode had been shown[ 131 to be associated with temper embrittlement of 3% Ni . Cr ’ MO - V disc steel, which may be induced in large forgings where the cooling rate can be very slow. Samples taken from the 3% Cr - MO . V did not show any response to standard temper embrittlement etches, even after ageing for 1000 hr at 470°C. This would indicate that temper embrittlement, if present in this material, was not at an advanced stage. The other disc and rotor materials did not show any tendency to intergranular failures, but the cold-worked austenitic materials did. Material 17 showed a coarse intergranular fracture, similar to Fig. 7(b), whereas material 18, although partially intergranular, showed a large amount of ductile tearing, Fig. 7(c). Optical metallography revealed almost complete carbide networks at the grain boundaries in material 17, which may account for the preferred fracture path. This difference in fracture mode appears to be a reflection of the relative toughness of the two materials, the high toughness material showing large areas of ductile tearing.
12% Cr * MO. V Nb steels Fracture toughness values measured on different 12% Cr. MO - V * Nb steels, all having the same minimum O-2 per cent proof strength, have shown considerable variation. Air melted material (7 and 11 in Table 2) was found to have a lower toughness than the vacuum remelted material 8, from which two specimens were taken. This vacuum remelted material again showed considerable variation, the specimen notched at a position corresponding to the centre of the billet (A41) having a lower value than A40, notched at the outside. Metallography and mechanical tests have been carried out on material from these two specimens, and also that taken from material 7 (A39 in order to investigate the variation in toughness. The strength of the three samples was similar, although the tensile ductility of A35 was much lower and the impact properties were also inferior. The only difference in mechanical properties between A40 and A41 was slightly lower tensile ductility at the inside position; chemical analysis did not show any significant variation. Metallographic examination showed the three specimens to have similar tempered martensitic structures, but the air melted sample had an inclusion count twice that of the vacuum remelted material. The major difference between the three samples was in the amount and form of the primary carbide which was present at the prior austenite grain boundaries. The specimens taken from the outside of the vacuum remelted billet, which had the highest fracture toughness, contained isolated particles of primary carbide, Fig. 8(a), whereas material at the centre of the billet contained incomplete grain boundary networks of primary carbide, Fig. 8(b). The air melted sample showed agglomerations of primary carbides at the grain boundaries in association with inclusions, Fig. 8(c). Examination of a tensile fracture of this material showed that in this case also the fracture path ran through grain boundaries containing primary carbides. These observations would indicate that the solution treatment was not sufficient to take these large primary carbides into solution. In the lO$in. square billet which had had a 1.8 : I reduction, more primary carbide was present at the centre, as would be anticipated from the original ingot segregation. Since the primary carbide distribution was the only difference noted between the two vacuum remelted samples, this may be responsible for the disparity in the toughness values.
137
Alloy steels in turbo-generatorcomponents
PRACTICAL SIGNIFICANCE AND INTERPRETATION OF RESULTS The object of the testing has been to establish the plane strain fracture toughness of various materials used in turbo-alternators. The specimens were taken from material representative of that used for component manufacture and the results have been assumed to be typical values. The critical defect size for fast fracture in certain components has been estimated using these results, assuming defects to lie normal to the direction of principal stress. Discs and rotors
The most critical region of disc and rotor forgings is at the central bore. The hoop stress is at a maximum due to the stress concentration effect of the bore and also this region is the position most likely to contain defects, which could be inclusions, unsealed porosity or hair-line cracks. The most serious defect is one which lies in an axial radial plane, reaches the bore surface and is infinitely long (length to depth ratio greater than 10). Table 4 shows the critical defect size calculated from the experimentally determined fracture toughness values for an applied stress equivalent to the uniaxial yield stress and an infinitely long surface crack. Conversion to an equivalent elliptical or circular defect, either at the surface or embedded, can be made by reference to Fig. 9 [ 141. It is to be noted that a typical design ratio of hoop stress to yield stress for discs and rotors at normal running speed is less than 0.5, which could increase to 0.72 on a 20 per cent overspeed. It is only under runaway conditions where stresses of yield point magnitude would occur. Table 4. Defect size at yield stress for large forgings at 20°C
Material No.
Material type
(zi.)
K,c
c (in.) 0.44 > 0.21 0.06
(ksi%)
la la lb
3 CMV
49.9
3 CMV
57.4
145 > 100 60
2
2f NCMV
54.5
90
0.14
3
3 NCMV
54.5
233 > 163
0.96 > 0.47
4 5
1 CMV 1 CMV
41.4 39.9
69 65
0.15 0.14
6a 6b 6b
1 CMVB 1 CMVB
41.8 34.0
57 104 1 81
0.10 0.49 > 0.30
7
12Cr
50.8
56
0.06
8a
12Cr (Vacuum remelted)
51.6
107
0.22
50.8
92 71
> 0.16 0.11
48.2
80
0.14
82.6
116
0.10
8a 8b 9 10
>
8Mn8Ni4Cr lSMn4Cr
Analysis for infinitely long surface defect (C = K,,‘/1.21 TR?).
138
D. V. THORNTON
b-6.44
t-“-v
w Surface
defects
Y
0 5.8
/
Internal
/
defects
Fig. 9. Flaw geometries for equivalent crack size effect, (a) Surface defects; (b) Internal defects.
In considering the likelihood of rapid fracture, an important aspect is the quality of forging which is accepted. The present inspection procedure includes a complete ultrasonic examination, magnetic inspection where possible, and a close visual examination of the bore. The present acceptance standards take into account the minimum possible operating temperature, the Charpy FATT, metallurgical cleanliness and the peak tangential stress at maximum overspeed. Cracks and porosity are not accepted. In the case of shafts and rotors the inclusions usually found are cigar shaped. Acceptance of these depends on number and dispersion, but the largest acceptable would be 0.25 in. long with a radial depth of approx 0.030 in. The maximum acceptable defect size on an L.P. disc is 0.25 in. dia near the bore surface (equivalent to 0.04 in. deep infinitely long surface defect). From Table 4 it can be seen that even at yield stress the critical defect size for rapid fracture is considerably in excess of these values. Before any advantage can be taken of this apparent safety, (e.g. relaxation of acceptance standards), it is necessary to obtain further information on such aspects, such as:(a) The effect of clouds of inclusions and interaction of defects. (b) The integrity of non-destructive testing techniques. (c) The equivalence of various types of defect and ‘flat bottomed holes’ (the term used to describe the size of defect from an ultrasonic defect echo). (cl) The rate of stable, slow crack growth of initial defects during service by fatigue or stress corrosion. The materials used as large forgings in the L.P. turbine and alternator, which have been chosen qualitatively and developed to give good impact transition properties,
Alloy steels in turbo-generator
components
139
show high values of fracture toughness. The test on the laboratory treated sample of 3% Cr - MO . V, where the proof strength was increased to 57 t.s.i. showed a marked loss in toughness. This indicates the danger which may arise in up-grading the strength of a material, even though it has a high toughness, without a quantitative assessment. The higher temperature service conditions dictate that the primary consideration for H.P. and I.P. rotors is creep resistance, and a 1% Cr * MO - V steel has been developed for these conditions. The toughness in various grades of this material is seen to be lower than for L.P. materials although the critical defect size at room temperature is still well in excess of the maximum defect acceptable by non-destructive testing standards. Intergranular fracture, which may be a result of temper embrittlement, occurred in one disc material even though it possessed a satisfactory fracture toughness. This could result from the slow rate used in cooling the disc after tempering in order to minimize internal stresses. Water quenching large forgings from the tempering temperature improves the impact transition properties, and therefore an improvement in fracture toughness may be possible. However, the effect of such a treatment on the residual stress pattern would have to be shown to be acceptable.
Alternator end rings The end ring has approximate dimensions of 2 ft long, 3 ft dia. and 3 in. wall thickness. The material is necessarily non-magnetic in order to minimize electrical losses. Cold expanded austenitic steels are used with a proof strength in the range 50-60 t.s.i. Various methods of expansion are currently being studied, including hydraulic and explosive methods, in order to increase the proof strength to 70 t.s.i. Two materials, 9 and 10, have been tested and the results used to estimate the critical defect size for rapid crack propagation at an applied stress equal to yield stress, Table 4. The higher strength material has a high degree of toughness and could tolerate a similar defect size to the lower strength material at a far higher design stress. These values are considerably larger than figures previously quoted for a martensitic stainless steel end-ring [ 151.
Bolts Turbine casing bolts are subjected to high temperature stress relaxation during service and are made from high proof strength, creep resistant materials. These materials generally have low impact properties and rapid fracture may occur, particularly on starting the turbine from cold when high transient stresses can be imposed while the bolt is relatively cold. The fracture toughness of a typical bolting material (material 11) has been determined and the value applied to the configuration of a 5f in. dia. bolt. The analysis used is for circumferentially notched round bar in tension, where the thread form has been assumed to be a sharp notch. The result suggests that failure would occur at a stress of 40 t.s.i. In practice the root of the thread is relatively blunt, which effectively increases the toughness of the component. However, during service creep damage may occur resulting in cracking at the first unsupported thread root, causing local sharpening and increasing the notch depth. In addition, deterioration in service may cause embrittlement and reduction in the fracture toughness of such components.
D. V. THORNTON
140
L.P. blades
The final stage L.P. turbine blades operate at a temperature of approximately 30°C in a condensing steam atmosphere. A stainless material is used to minimize corrosion, and the size of these forgings is approximately 40 X 6 X 24 in. The blades are supported by lacing wires, approximately +in. dia. which pass through the blade at an angle of about 20”, making an inclined elliptical hole in the blade where the blade thickness is about fin. An anticipated failure mechanism is the growth of a fatigue crack to critical size from the edge of a lacing hole. Tests have been carried out on 12% Cr stainless martensitic steels at present used as blades with a 36 in. aerofoil section, having a proof strength of 50 t.s.i. Shear fracture after general yield occurred in these specimens which simulated a fatigue cracked lacing hole stressed in tension. Future designs, blades 42 in. long or greater, will require higher strength stainless materials and the blade thickness may also be increased. In this case plane strain rather than plane stress failure may dominate. The plane strain fracture toughness of a number of potential materials has been determined and the values used to estimate the critical crack length for rapid fracture. The analysis used is for a centre cracked finite plate at a design stress of half yield strength. These results are shown in Table 5. The effective initial crack length due to the inclined elliptical lacing wire hole is of the order of I in., dependent on blade thickness, and several materials are clearly not satisfactory for such an application. With materials of high toughness, an additional advantage may be an effective further increase in toughness by domination of plane stress conditions. It is interesting to note that the fracture toughness of material 14 is reduced by a factor of approximately two as the proof strength is increased by about 10 per cent. The leading edge of large L.P. blades requires protection from erosion by water droplets contained in the condensing steam. This is accomplished at present by brazing Table 5. Defect
size at half yield stress for blading materials at 20°C WY
C
K/c
(t.s.i.)
12
82
47
0.17
13
71
36
0.13
14a 14b 14b 14c 14c
78 74
107 207 > 155 234 > 143
0.90 2.14 > I.64 2.38 > I.60
15
60
> 173
> 2-20
16
63
152
1.88
17
79
52
0.22
18
80
206 1 162
2.01 > 1.58
70
(ksiG)
(in.)
Material No.
Analysis for a centre cracked plate K,(.2 = d na/W where a = C/2 and W = 3 in.
W tan
Alloy steels in turbo-generator
components
141
on an erosion resistant shield, but possible alternative methods of protection include welding on an erosion resistant tip or local hardening of the blade material. This latter process introduces further difficulties, however, as martensitic stainless steels at the high strength level necessary to impart sufficient erosion resistance are susceptible to stress corrosion by hydrogen embrittlement in condensing steam. The limiting stress intensity below which stress corrosion cracking will not occur has been determined for two steels and has been found to be only 25 per cent of the plane strain fracture toughness. This information can be used to assess blade design and minimize the likelihood of slow crack growth by stress corrosion. SUMMARY The results of work completed to date have enabled the critical defect size for rapid fracture due to a single load application to be estimated for certain materials used as large forgings in turbo-machinery. In addition, quantitative assessment of potential materials for new designs has been made. The values determined have been assumed to be typical and further testing is necessary to establish lower bound values upon which design and acceptance procedures could be based. In the case of large forgings where initial defects may cause rapid fracture, it is essential that non-destructive testing methods are capable of detecting and accurately estimating the size and location of such defects. The present work covers the effect of single defects and further work is required to assess the interaction of defects and the effectiveness of clouds of small inclusions in initiating rapid fracture. The equivalence between linear elastic and general yield fracture mechanics has been shown by this work where excellent agreement between results using both approaches has been obtained for the plane strain fracture in a variety of alloy steels at various strength levels. There are many areas where a quantitative assessment of materials and design can aid the development of larger, high efficiency turbo-generators. This work begins to indicate the margin of safety from the hazard of fast fracture initiating from defects for virgin materials. Deleterious effects which may occur in service, such as slow crack growth by fatigue or stress corrosion and material embrittlement, are problems presently being investigated. Acknowledgements-The author would like to thank the Director of Research, English Electric Co, Ltd., for permission to publish this paper. The guidance of Professor A. A. Wells, particularly in the use of the Marcal-King analysis, is gratefully acknowledged.
REFERENCES [l] [2] [3] [4] [5] [6] [7] [8] [9]
E. E. Thum, Recent accidents with large forgings. Mef. Prog. 69.49-57 (1956). F. Buckley, Metallurgical aspects of turbine constructionII. fron & Stee/39.324-328 (1966). G. R. Irwin, Sfrucfural Mechanics, pp. 557-594. Pergamon Press, Oxford (1960). J. E. Srawley and W. F. Brown, Fracture toughness testing methods. ASTMSpec. Tech. Publ.No. 38 I ( 1965). A. A. Wells, Notched bar tests, fracture mechanics and brittle strengths of welded structures. Br. Weld. J. 12,2- I3 (I 965). E. T. Wessel, State of the art of the WOL specimen for K,c fracture toughness testing. Westinghouse Res. Rep. No. 67-l D6-BTLFR-RI (1967). D. F. Mowbray, A. J. Brothers and S. Yukawa, Fracture toughness determinations of A302B and Ni . MO. V steels with various size specimens. Trans. ASME, J. bus. Engng, Paper No. 66-Met. 1. A. J. Brothers, D. L. Newhouse and B. H. Wundt, Results of bursting tests of alloy steel discs and their application to design against brittle fracture. Gen. Elect. Res. Rep. Ger-22 18, presented at ASTM68rhA. Mrg(1965). F. H. Burdekin, Initiation of brittle fracture in structural steels. B.W.R.A. RepNo. E/l l/66 (1966).
D. V. THORNTON
142
[toI
P. V. Marcal and I. P. King, Elastic-plastic analysis of two-dimensional stress systems by the finite element method. Inr. J. mech. Sci. 9, I43- I55 ( 1967). L.P. discs. [111J. F. Knott, Application of fracture mechanics to an alloy steel used in turbo-generator J. Iron Steel Insf. 204. IO1 4- IO2 I ( 1966). on[ 1 l],J. Iron Steel Insr. 205,304 ( 1967). [I21 Correspondence in high purity 3.5 Ni-1.75 Cr-0.20C steel. Paper presented to r131 G. C. Gould, Temper embrittlement ASTM, October (I 967). of fracture information in specifications. ASTM Spec. Tech. Publ. No. 3X 1 [I41 W. F. Payne, Incorporation (1965). applications of Brittle fracture research. The Engineer 1151 M. G. Gemmill and K. Smith, Engineering 233,632-634(1967). (Received20
May
1968)
APPENDIX
K6 correction using the Marcal-King analysis [ lo] The Marcal-King method for the incremental elastic-plastic analysis of a two dimensional, plane strain stress system is based upon the finite element method of stress analysis and uses uniformly stressed triangular elements. The results derived from an analysis of a notched tension specimen, of notch depth to half width ratio of l-2 and total width 6 in., under plane strain conditions have been used. COD results have been converted to a stress intensity using (4), where
A more accurate value can be obtained if corrections are of measurement with respect to the crack tip, and to oy using the Marcal-King analysis. Table A 1 gives the displacement in the above specimens the crack tip (D) and values of applied to yield stress (a/~,). Table A
made to 6 for the position for stress intensification, at various distances from These values have been
I. Kelative displacement
D, ulu,
0
0.10
0.20
0.40
0.60
0.80
0.357 0.644 0.929 1.218 I.504
0.074 0.222 0.442 0.705 0.997
0.170 0.395 0.689 1.074 1.586
0.266 0.568 0.936 1.443 2.175
0,408 0.841 I.320 I .930 2.876
0.514
0601 I.219 I .874 2.616 3.634
D = distance
from crack
Table A2. Factored D, ulrr, 0,357 0.644 0.929 1.218 I.504
I.049 1.623 2.296 3.284
tip (in.). relative
displacement(f)
0
I /30
2130
4/30
6/30
8130
I .oo
2.30 1.78 1.56 I .52 I .59
3.59 2.56 2.12 2.04 2.18
5.51 3.78 2.98 2.73 2.88
6.95 4.72 3.68 3.25 3.30
8.12 5.48 4.23 3.71 3.64
1 .oo I .oo 1.00
Ia
D = fractional
distance
from crack tip.
Alloy steels in turbo-generator
143
components
Table A3. Stress intensification
a/u,
H
0.357 0644 0.929 1.218 1.504
1.385
I ,766 I.903 2.118 2.235
H _ or oil Table D, u/o, 0.357 0644 0.929 1.218
I.504
_ Maximum stress Yield stress
= Stress intensification.
A4. w
0
l/30
2/30
0.85 0.75 0.73 0.69 0.67
1.29 1.01 0.90 0.85 0.84
1.61 1.20 1.05 1.13 1.13
4130
6130
8/30
1.99
2.24 164
2.42 I .76 1.49 1.32 1.28
1.46 I .25
1.24 1.22
1.34 1.32 1.28
converted to a relative displacement (F) at fractional distances, with respect to notch depth, from the crack tip and are shown in Table A2. The stress intensification (H) at various applied to yield stress ratios are shown in Table A3. Hence
(5) and combining (4) and (5)
Values of m are shown in Table A4. These figures have been used to correct the K6 values calculated from (4) and are given in Table 3 of the paper as ratios of Klc determined by linear elastic techniques. R&meLa resistance a la rupture par tension plane de divers aciers allies utilises dans la fabrication de pieces forgees pour les turbo-generateurs a CtC determinee. On a montre que des valeurs Cquivalentes de resistance a la rupture sont obtenues a partir d’essais mecaniques de limite de resistance a la rupture generale et par allongement lineaire. Les valeurs obtenues ont servi a tvaluer la tendance rapide a la rupture des differents composants a partis de defauts simples. L’effet des facteurs metallurgiques sur le mode de rupture et la resistance a la rupture est illustre. ZusammenfassungDie Bruchzlhigkeit im ebenen Dehnungszustand von verschiedenen legierten Stahlen, die als Schmiedestticke in Turbogeneratoren verwendet wurden, wurde bestimmt. Es konnte gezeigt werden, dass Pquivalente Werte fur die Bruchzahigkeit dieser Werkstoffe erhalten werden bei Verwendung linearelastischer sowie vollplastischer Bruchmechanik. Die erhaltenen Werte wurden zur Beurteilung der Anfalligkeit fiir Rapidbruch, ausgehend von einzelnen Fehlern der verschiedenen Bestandteile, verwendet. Der Eintluss metallurgischer Faktoren auf die Bruchart und-z%higkeit wird veranschaulicht.
Frocturc
Mechanics,
1970, Vol. 2, pp. 125- 143.
Pergamon Press.
Printed in Great Britdln
FRACTURE TOUGHNESS OF ALLOY STEELS TURBO-GENERATOR COMPONENTS Central Metallurgical Laboratories,
USED
IN
D. V. THORNTON English Electric Co, Ltd., Whetstone, Leicester, England
Abstract-The plane strain fracture toughness of various alloy steels used as forgings in turbo-generators has been determined. It has been shown that equivalent values of fracture toughness are obtained for these materials using both linear elastic and general yield fracture mechanics. The values obtained have been used to assess the rapid fracture susceptibility of various components from single defects. The effect of metallurgical factors on the fracture mode and toughness is illustrated.
NOTATION K/c FATT 6 eU T a
fl’y E G ” K8
plane strain fracture toughness determined by linear elastic fracture mechanics, ksi 6. fracture appearance transition temperature, “C crack opening displacement (COD), in X 10es yield strain specimen thickness, in. notch depth, in. yield stress, t.s.i. elastic modulus, t.s.i. strain energy release rate, in. lb/in’ Poisson’s ratio. stress intensity determined by general yield fracture mechanics. based on measured ksi 6 as K8 but based on corrected values, ksi 6 applied stress, t.s.i. distance from crack tip, in. fractional distance from crack tip compared to notch depth relative displacement stress intensification Charpy value at test temperature, ft-lb nominal fibre stress, t.s.i. crack depth, in. thickness, in.
values,
INTRODUCTION THE SIGNIFICANCE
of defects and the elimination of the hazard of rapid fracture, which in large rotating machinery causes catastrophic failure[ 11, have been a problem to the turbine builder for a considerable time. The present method of assessing the susceptibility of large forgings to fast, unstable fracture is based largely on routine quality control tests, rules of thumb and experience. It is obviously highly desirable to develop a quantitative approach for both acceptance and design of large turbogenerator components. The recent expansion of the aerospace industry has involved the use of high strength metallic materials and has emphasised the need for a quantitative assessment of rapid fracture. This has led to a renewed interest in linear elastic fracture mechanics originally developed by Griffith in 1920 to explain the fracture of glass. Recent advances in both the theoretical and practical aspects have now brought fracture mechanics to a stage at which some of the problems in the assessment of large forgings in medium strength steels used in turbo-generators can now be dealt with in a more scientific manner. 125
126
D. V. THORNTON
The major advantage is that quantitative information is provided for design, material assessment and inspection acceptance purposes, which can be used where service experience with new materials and new designs is not available. By consideration of the design stress and configuration, the anticipated defect size and the fracture toughness, the safety of the component can now be judged more precisely. The components considered include discs and rotors, alternator end rings, bolts and L P. blades. Turbo-generators usually operate at 3000 revlmin and the major stressing of rotating parts is by centrifugal forces. From the rapid fracture aspect, the low pressure cylinder and alternator are the most critical regions since the material strength and component size are high and the operating temperature relatively low. These factors increase the susceptibility to rapid fracture which may initiate at such stress concentrations as inherent material faults, design geometry and cracks which have grown in service by fatigue or stress corrosion. During the past ten years the size of single machine 3000 rev/min turbo-generator units has increased from 100 to 500 MW. Figure 1, which shows a 500 MW unit being assembled, illustrates the large size of these units. Typical materials and component sizes used in the 100 MW and 500 MW machines are shown in Table 1. This indicates that the design of increased output units has led to the use of higher strength materials in considerably larger section sizes, both of these factors increasing the likelihood of rapid fracture. Material assessment has been based on standard quality control testing. In spite of the progression to higher strength levels, developments in steel composition and treatment have enabled the tensile ductility to be maintained and the impact transition properties to be enhanced. Increasing use of modern non-destructive testing methods, particularly ultrasonic detection, has enabled the quality of large forgings to be determined and a high acceptance standard to be established[21. With the cooperation of the steel makers and forgemasters, materials and techniques have been developed whereby components of a high standard are consistently produced. The unfortunate occurrences experienced in the United States in 1953 and 1954, where several catastrophic failures of turbo-generators occurred [ll, have so far been avoided in the United Kingdom. However, as service conditions become more severe with the development of larger, higher efficiency machines, the need for quantitative assessment of materials and designs becomes more acute. TEST TECHNIQUES With the large forgings used in turbo-machinery the thickness is such that the fracture mode would be predominantly plane strain (e.g. large discs are 22 in. thick). Thus the objective is to measure the plane strain fracture toughness (K,,), which is the lowest energy fracture mode at the lowest operating temperature. At the last stage of an L.P. turbine, this is almost ambient temperature. Other components, while normally operating at higher temperatures, may be stressed under start-up conditions at a temperature near to ambient. One of the test techniques employed in this work to measure K,c is based on linear elastic fracture mechanics principles developed by Irwin [3,4]. The data have been determined using the testing parameters current in the ASTM literature, bend testing being employed for all tests. The notch configuration comprised a parallel sided milled slot terminating in a spark machined slot, O-1 in. deep and OXKK3 in wide which was subsequently lengthened by a fatigue crack. Testing of the medium strength, high
Fig. 1. 500 MW steam turbine assembly.
Fig. 5. Fracture
Tested
path following
the plastic
hinge in a
cold
worked
austenitic
at :+2o*c
-2ooc
-60°C Fig. 6. 34% Ni
Cr
. MO V fracture
surfaces.
steel.
(b)
(cl
Fig. 7. Typical fracture modes (a) Transgranular quasi-cleavage fracture Intergranular fracture X 150; (c) Mixed trans- and inter-granular fracture
480; (b) 150.
X X
Ib)
Fig. 8. Primary
carbides
in
12%Cr
Mo .V Nb steels. (a) Specimen
200; (c) Specimen
A 35 X 200.
A 40 X 200; (b) Specimen
A 4
IX
L.P. disc. L.P. shaft. I.P. rotor. Alternator rotor. End bell. L.P. blade.
Component
13 thick 25 dia. 37 dia. 3 5 dia. 3 thick 27 Aem foil
Size (in.) 39 22 30 19 47,5 28
Carbon 3%Cr,Mo 8% Mn 8% Ni Mn.Ni
~~inimum 0.2% PS (t.s.i.)
3%Cr.Mo Carbon
Steel
100 MW unit
464 55 diq. dia. 34 thick 36 Aerofoil
22 thick 29 dia.
Size (in.)
2%Ni.Cr.Mo.V l%Cr.Mo*V 18%Mn .Cr 12%Cr.Mo.V
3+%Ni.Cr.Mo.V 2t%Ni.Cr*Mo*V
Steel
500 MW unit
Table 1. Materials used for various components of a 100 MW and 500 MW unit
;: 57 50
50 42
Minimum 0.2% PS (t.s.i.).
128
D. V. THORNTON
toughness forgings used as discs, rotors and large L.P. blades has necessitated the use of large specimens in order to obtain plane strain failure, where sections up to 6 in. square have been used. The majority of specimens exhibit a maximum load failure. Destructive testing of forgings to measure typical fracture toughness values is expensive, both in obtaining representative material and in manufacture and testing of the specimens. In order to establish the range of toughness for a particular material and to carry out quality control tests based on fracture mechanics principles or correlated with fracture mechanics results, it would be necessary to use samples taken from production forgings. In this case the material available for testing would be severely limited and, therefore, a small scale test is required for this purpose. In view of this, general yield fracture mechanics [5[ has been investigated. The displacement at a crack tip is assumed to describe the local stress and strain conditions and it is assumed that fracture will occur at a critical displacement dependent on material, temperature, strain rate and through thickness stress. The crack opening displacement (COD) has been measured on the large specimens used for Klc determination and also on 0.4 in square bend specimens machined from these large specimens. The displacement was measured using a strain gauged, thin strip of copper-beryllium inserted to the bottom of the 0.008 in wide spark machined slot. A torque was applied to the gauge prior to testing and the change in signal due to decrease in torque as the crack opens on loading was autographically recorded. Each gauge was calibrated, using a micrometer arrangement, to determine the displacement corresponding to a particular change in signal. On the K,(. specimens the autographic trace served as a record of the specimen compliance and a measure of the COD. The same technique was used to measure COD on the 0.4 in. specimens. Initial tests on O-4 in. square specimens of a 3% Cr + MO . V disc steel indicated that the use of 10 per cent side grooves aided in the detection of first cracking i.e. the critical COD value. A limited number of tests were carried out with specimens containing slots of different width. Specimens were prepared with 0.014 in., 0.009 in. and 0.006 in. wide slots and also an 0.009in. slot terminating in a fatigue crack. Similar COD values were obtained on all except the widest slot, Fig. 2, where the COD value measured was larger and showed considerable scatter. x Tested n Tested
0
I
2
3
4
5
Notch
6
7
width
6
9
at at
IO
2O’C 60°C
II
12
13
I4
, in. x 10e3
Fig. 2. The effect of notch width on COD of a 3% Cr. MO . V steel.
15
Alloy steels in turbo-generator
components
129
In subsequent tests a specimen containing 10 per cent side grooves and having a notch depth of 20 per cent beam depth, with the slot terminating in a fatigue crack, was used. Summing up, the object of the work was to establish typical fracture toughness values for various materials and to examine the COD tests as a means of quality control testing. Testing variables have not been systematically investigated, the testing conditions being standardised where possible. Study of metallurgical variables was confined to a few special cases. RESULTS Linear elastic fracture
AND DISCUSSION
mechanics
The chemical composition, source of material and room temperature mechanical properties of the materials tested are shown in Table 2. The results of the large bend tests are given in Table 3. In cases where post-yield fractures have been obtained and the compliance and load deflection curve deviated from linearity, two values are quoted in the Klc column. These are a minimum value calculated at a failure stress equivalent to uniaxial yield stress, and the value calculated at the actual failure stress (shown in brackets). The recent tentative recommended testing practice for fracture toughness of the ASTM E-24 sub-committee I has proposed that for valid KIC measurement both the specimen thickness and crack length should be greater than 2.5 (K,c/cr,)2. It is felt that this requirement is conservative and that there may be a relaxation of this as more data becomes available. In the present tests a square section with a notch 25 per cent of the beam depth has been used and although the crack length is generally too short, the thickness in most cases is sufficient to meet this requirement. It is considered, therefore, that the data is sufficiently accurate for the present purpose of estimating the tolerance of a material to defects. General yieldfracture
mechanics
Results obtained to date using the small COD specimens are given in Fig. 3. Although the small COD specimens failed with a flat fracture and rapid crack propagation, the majority also failed after general yield. The conditions of constraint were not similar to those of the large bend specimens and to the actual components where plane strain elastic failure is anticipated. These COD results, therefore, are not expected to agree with Klc values determined at the same temperature. There is a large amount of evidence in the literature [6-81 that K,c of low alloy steels steels shows a rapid increase over a certain temperature range. Wells [51 and Burdekin [9] have shown that the COD of certain steels show a transition with increase in temperature similar to the well known impact transition. It is thought that a quality control test for medium strength low alloy steel forgings may possibly be established by correlation of the COD transition on standard specimens with the KIC transition. A correlation of Klc and Charpy FATT has been shown for materials used as turbine forgings and has been incorporated in a fracture analysis diagram[8]. The COD vs. temperature curves for the O-4 in. square specimens in the present work show a transitional behaviour, but as can be seen in Fig. 3. the temperature range over which this occurs is not similar to the respective Charpy FATT. The COD value at the upper shelf agrees
18
17
16
14b) 14c) 15
14a)
13
12
9 10 11
8b
la lb 2 3 4 5 6a 6b 7 8a
Material No.
tested-chemical
Source
properties
80.3
78.6
82.2
87.9
70.8
69.3
60.1 62.9
86.6 78.2 72.4
80.5 77.9 73.9 70.3
70.7
91.5
63.0
55.0 82.0
57 50 17 31
60.2 59.3 54.3 92.4
51.6 50.8 48.2 82.6
11
15
60
58
61 66 68
40
53
50 14 53 62 59 18 48 50 15
59.3 67.9 62.0 63.1 50.8 50.8 49.6 47.8 60.0
49.9 57.4 54.5 54.5 41.4 39.9 41.8 34.0 50.8
0.2% PS (t.s.i.)
and mechanical
Disc 3.5Cr,0.5Mo, O.lV. Disc 3.5Cr, 0.5Mo,O.lV. 2.4Ni, OGSCr, 0.5Mo,O.lV Disc 3.3Ni, 1.6Cr. 0~6Mo,O~l5V Disc Rotor l.OCr, 0.7Mo,O.32V, 0.7Ni Rotor 1.2Cr, l.lMo, 0.37V, 0.6Ni l.OCr, 0.7Mo,O.39V, 0.7Ni, 0.006B Rotor l.OCr, 0.7Mo,O.39V, 0.7Ni, 0.006B Rotor 11 .lCr, 0.6Mo,O.27V, 0.21Nb, 0.07N Disc 104 in. billet 9.9Cr,0.7Mo, 0.3OV. 0.22Nb outside O.l5C, 9.9Cr, 0.7Mo,O.3OV, 0.22Nb centre End bell 0.6C, 7.0Mn, 9.ONi, 4,OCr End bell 0.5C, 18.8Mn. 4.4Cr, O.lNi O.l4C, ll.lCr,0.8Ni,0.6Mo, Bar (5 in. dia.) 0.3V, 0.4Nb O.O4C, 14.3Cr, 5.4Ni, 1.7M0, I .7Cu Bar 0.8Nb (3 X 2: in.) Bar Ti, 2.3A1, 11.2Sn. 4.6Mo (2 X 2 in) Bar O.O5C, 14.2Cr. 5.3Ni, 1.6Mo,1.6Cu, O+Ib (3 X 24 in.) (3 X 2: in.) O.OSC, l4.2Cr, 5,3Ni, 1.6M0, 1,6Cu, 0.4Nb (3 X 2tin.) O.O5C, 14.2Cr, .5,3Ni, 1.6M0, 1.6Cu,0.4Nb O.llC, 11.4Cr,2.6Ni, 1.3Mo,O.l3V, Bar (6s x 26 in.) 0.2Nb, 0.06N O.l2C, 12.5Cr, 2,5Ni, 1.8Mo,O.3V, Bar O.OSN (66 X 2t in.) O.O4C, 12.6Cr, 4,4Ni, 0.3Ti, O.INb Plate (3 X 1 in.) 14.2Co,O.O7Al Plate O.O3C, 12,5Ni, 4,9Cr, 2.9Mo,0.2Ti, (36 X 1 in.) 0.5Al
0.35C, 0.35C, 0.28C. 0,35C, 0.3C, 0.3C, 0.2C, 0*2C, O.l6C, O.lSC,
Composition (Wt%)
Table 2. Materials
+6 -
-23
+60 -5 -25
(+ 140)
+ 120
+73 +72 _ _
+ 140
+41 -1 +69 + 123 +s5 _
+58 _
50% FATT (“C)
Alloy steels in turbo-generator
, OC
Temperature
k;[~~“;T,T~
-00
,
-40
0
40
Temperature
Temperature
131
components
:
120
80 ,OC
,
160
, OC
Fig. 3. The effect of test temperature on COD.
with that calculated according to Wells[5] for the transition from plane strain to plane stress behaviour, 6 = 27r e,T.
(1)
Use of the expression given by Burdekin[9],
6
=
ha -log
77
(
5+1
>
,
(2)
predicts the COD value at which plane strain conditions will not be dominant. This value is found to be in the transition range. These calculated values are shown in Fig. 3., together with the COD measured in large specimens. From this it would appear that plane strain COD values can be obtained at low temperature on small specimens which are similar to those obtained on larger bars. (Subsequently these will be shown to agree with Klc results.) However plane stress fracture begins to dominate as the transition range is approached. Further tests are in progress on other materials and these may show a more definite pattern of behaviour.
2 2
3 3
3
4
5 5 5
6a
6b
7
8a 8b
A21 A22
A24 A28
A29
A23
A33 A38 A39
A25
A34
A35
A40 A41
9 9
10
A30 A31
2281
End hell materials
+20 + 100
2 2
Al5 Al7
+20
+20 +20
+20 +20
+20
+20
+35
+20 +20 +20
+20
-30
+20 -60
+40 -75
+20 -65 -20 -30 +20
Materialt
Test temp. (“C)
Disc and rotor materials A8 la Al3 la Al4 la A27 la A26 lb
Specimen. ident.
82.6
48.2 48.2
51.6 50.8
50.8
34.0
41.8
39.9 39.9 39.9
41.4
54-5
54.5 58.8
52.8 58.5
54.5 51.0
49.9 57.0 52.2 50.3 57.4
0.2% PS (a,) t.s.i.
-
18 18
9
9
6 6 6
15
18
48 13
43 8
34 77
20 10 13 12
c,.+ ftlb
K,C
0.87
0.88 o-95
I.16 0.81
0.70
1.28
0.50
0.64 1.11 l-20
0.74
‘a
>
116
79 84
92 (107) 71
56
81 (104)
57
65 63 71
69
> 133 (177)
> 163 (233) 83
I .43 0.56 1.33
> 127 (169) 58
90 > 140 (215)
> 100 (145) 76 109 > 108 (126) 60
ksifi
1.33 0.48
0.60 1.53
1.44 0.65 1.04 1.17 0.49
rr,/Ou§
(0.6)
1.5
I.1 0.9
I.1 (0.9) 1.9
2.3
1.0 (0.6)
2.8
2.4 1.6 0.6
1.5
O-8 (0.4)
0.6 (0.3) 2.2
0.6 (0.4) 3.7
I .9 0.7 (0.3)
3.4
:I;
I.6
0.7 (0.4)
($
Table 3. Large bend test results
2.5
5.0 4.5
3.6 (2.7) 5.9
8.3
3,l (2.2)
10.6
10.0 3.5 2.8
5.9
3.5 (2.0)
2.8 ( 1.4) 10.1
2.6 (1.5) 12.7
8.3 3.3 t1.4)
3.3 (1.6) 14 3.0 2.9 (2.1) 12.1
$$
(in
6
IO-:‘)
IO.0
9.7 8.2
6.4 4.2
1.6
6.5
2.6 1.8 4-o
2.3
8.1
23 .O
1.9
10.3
1.1 -
6.1 2.0 3.0 5.3 2.4
X
247
186 171
160 130
77
127
77 73 108
84
180
305
200 91
67
150 92 116 140 101
K6 ksifi
-
2.13
2.35 2,04
1.54
1.58
1.89
1.20 1.29
1.48 1.83
1.07
1.14 1.38
1.22
0.79 1.09 1.16
1.04
1.22 1.18 1.21 1.53
0.93
1.13 -
0.72
1.08
0.96 -
0.81 0.99 0.95 1.08
1.05
K&K,>
1.02
1.31 -
1.18 1.56
0.75 -
1.11 1.68
1.06
I .03 1.21
KslK,,l
w N
t20
13
14a 14b 14c
15
16
17 18
A16
Al8 At9 A20
A58
A37
B94 895
tMaterials referenced as in Table 2. tC,.- Charpy value at test temperature. §o!--nominal fibre stress at fracture. “KeIK,c correction- see Appendix.
+20 +20
+20
+20
+20 +20 +20
+20
materials 12
Biading Al2
+20 +20 +20
11 II 11
material
Bolting B91 B92 B90
78.6 80.3
62.9
60.1
77.9 73.9 70‘3
71.0
82.0
55*0 55.0 55.0
24 56
28
61
26 60 87
-
5
9 9 9
0.34 1.27
0.83
I.55
0.66 1.33 163
0.29
0.28
0.65 0.56 0.84 .-
52 )z 162 (206)
152
> 173 (257)
107 r 155 (207) > 143 (234)
36
47
53 45 56
IO.8 1.4 (0.9)
1.4
0.7 (0.7)
(0.5) O-8 (0.3)
::f;
9-6
2.8
2.4 3.8 1.1
12.3 1.3 (0.8)
2.2
1.6 (1.5)
6.7 2.9 (1.6) 3.0 (1.1)
40.2
38.3
8.1 11.3 4.8
1.8 8.5
8.8
21.5
10.0 12.8 15.4
2.1
2.8
2.5 1.8 1.7
102 225
202
300
240 265 283
71
131
100 85 83
I.09
1.98
1.33
1.16
2.24 1.28 I.21
1.97
2.79
1.89 1.89 1.48
0.90 0.80
0.98
1.12
1.37 099 1.01
0‘74
1.13
1so2 0.89 0.95
E
% 3 a $
;
3 =: !: z 8 t; E’
D. V. THORNTON
134
Relation of COD and Krc values
Where both COD and Klc values were obtained in the same large bend test the theoretical relationship proposed by Wells has been used to calculate a stress intensity from COD Su,=G
(3) (4)
This is shown as KS in Table 3 and the values of Ks/KIc in Fig. 4(a). The figures obtained for the initial tests on 3% Cr. MO . V disc steel (material l(a) in Table 3) showed good agreement but the histogram, Fig. 4(a), shows that the COD generally over-estimates Klc. The instrumentation is considered satisfactory even for accurate measurement of the small COD values found in these tests. Therefore, other factors have been examined, namely the stress value to be used in the equation and errors due 9 0 7
I
6
(a)
;
3
l--l
4 3 2
G
5lr
0.4
0.6
0.6
I.0
16
12
l-h 2.2
16
2.4
2-6
2-8
3.0
KS ‘KIC
Il. IO . 9. 0-
( b)
7. ; g
6.
z
543-
n
2I .
04
06
06
10
!
! ! !--j
14
I.6
16
20
, , , , , 2.2
24
26
26
30
hdK,c
Fig. 4. Comparison of fracture mechanics (KS ics (K,,.). (a) based on rected
stress intensity derived from general yield and KS,(.) and linear elastic fracture mechanmeasured values; (b) based on values corby Marcal-King analysis.
Alloy steels in turbo-generator
components
135
to position of displacement measurement. The yield value used to calculate K8 was the uniaxial yield strength whereas the value across the yield zone required modification for stress intensification under the severe biaxial stress conditions at the notch tip. The experimental technique employed measures the COD in the mid-thickness of the specimen at the base of the machined notch. This point has been up to 0.4 in. from the tip of the fatigue crack, resulting in an erroneous result. The results of an analysis due to Marcal and King [ lo] have been used to correct the results for measurement position and stress intensification. The method used is shown in Appendix I. The resultant ordering of the results is shown in the histogram, Fig. 4(b). The three results which have corrected KS Ic/K,c ratios greater than 1.4 are cold worked austenitic steels. These fractures were significantly different from all the others in that the fracture path appears to follow the plastic hinge, Fig. 5, and not to occur normal to the specimen. The reason for this may be metallurgical in that heavy carbide networks were present at the austenite grain boundaries and the material is susceptible to stress corrosion. This may result in slow crack growth during the test and, therefore, errors in the application of the correction factors. These results show that excellent agreement between general yield and linear elastic fracture mechanics can be obtained for plane strain fracture provided that the correct COD value and stress are used in (3). A significant difference between fracture toughness values calculated by the two approaches has previously been reported by Knott[l l] for an L.P. disc steel, and possible reasons for this discrepancy have been suggested[ 11, 121. A sample of the 3% Cr. MO * V bar used by Knott in his investigation was obtained and heat treated in a similar manner to the 61.5 t.s.i. 0.2 per cent proof strength material used in[ll]. Fatigue cracked specimens were tested in three point bending and showed an elastic behaviour. Klc values calculated were 63 and 72 ksi X&. The actual COD was measured as 0.49 and O-53 x lo-” in. respectively from which calculated Ksrc values of 70 and 74 ksi fi were obtained using the Marcal-King corrections. Again very good agreement was obtained. In these tests the COD was measured during the test and therefore includes both the elastic and local plastic strain. Knott used double notched specimens and measured the displacement after the test and therefore any elastic contribution would be neglected. This is likely to lead to inaccuracies and low K*,C as in [ 111, since the total displacement is only 0.1 X 10e3 in. at the crack tip. Failure mode
Examination of the fracture faces by visual, optical and scanning electron microscopy has yielded information on the fracture characteristics of certain alloys. In general the toughness of the material is reflected in the degree of roughness of the fracture surface. Figure 6 shows 34% Ni . Cr . MO . V specimens tested at three temperatures. The room temperature specimen has a very rough surface showing the direction of propagation, whereas at low temperature where the material has low toughness the fracture is flat and featureless. However, all three specimens appeared to have fractured by a similar mode, being transcrystalline quasi-cleavage. Scanning electron microscopy has indicated three modes of fast fracture, transgranular, intergranular and a combination of these modes. Typical examples are shown in Fig. 7. The majority of fractures have been found to be transgranular quasi-cleavage, Fig. 7(a), the relative amounts of cleavage and ductile tearing depending on material and test temperature. In some specimens slow crack growth could be seen as ductile E.F.M. Vol. 2 No. 2-D
136
D. V. THORNTON
tears. Figure 7(b) shows the type of failure which occurred in all 3% Cr. MO - V specimens. This intergranular mode had been shown[ 131 to be associated with temper embrittlement of 3% Ni . Cr ’ MO - V disc steel, which may be induced in large forgings where the cooling rate can be very slow. Samples taken from the 3% Cr - MO . V did not show any response to standard temper embrittlement etches, even after ageing for 1000 hr at 470°C. This would indicate that temper embrittlement, if present in this material, was not at an advanced stage. The other disc and rotor materials did not show any tendency to intergranular failures, but the cold-worked austenitic materials did. Material 17 showed a coarse intergranular fracture, similar to Fig. 7(b), whereas material 18, although partially intergranular, showed a large amount of ductile tearing, Fig. 7(c). Optical metallography revealed almost complete carbide networks at the grain boundaries in material 17, which may account for the preferred fracture path. This difference in fracture mode appears to be a reflection of the relative toughness of the two materials, the high toughness material showing large areas of ductile tearing.
12% Cr * MO. V Nb steels Fracture toughness values measured on different 12% Cr. MO - V * Nb steels, all having the same minimum O-2 per cent proof strength, have shown considerable variation. Air melted material (7 and 11 in Table 2) was found to have a lower toughness than the vacuum remelted material 8, from which two specimens were taken. This vacuum remelted material again showed considerable variation, the specimen notched at a position corresponding to the centre of the billet (A41) having a lower value than A40, notched at the outside. Metallography and mechanical tests have been carried out on material from these two specimens, and also that taken from material 7 (A39 in order to investigate the variation in toughness. The strength of the three samples was similar, although the tensile ductility of A35 was much lower and the impact properties were also inferior. The only difference in mechanical properties between A40 and A41 was slightly lower tensile ductility at the inside position; chemical analysis did not show any significant variation. Metallographic examination showed the three specimens to have similar tempered martensitic structures, but the air melted sample had an inclusion count twice that of the vacuum remelted material. The major difference between the three samples was in the amount and form of the primary carbide which was present at the prior austenite grain boundaries. The specimens taken from the outside of the vacuum remelted billet, which had the highest fracture toughness, contained isolated particles of primary carbide, Fig. 8(a), whereas material at the centre of the billet contained incomplete grain boundary networks of primary carbide, Fig. 8(b). The air melted sample showed agglomerations of primary carbides at the grain boundaries in association with inclusions, Fig. 8(c). Examination of a tensile fracture of this material showed that in this case also the fracture path ran through grain boundaries containing primary carbides. These observations would indicate that the solution treatment was not sufficient to take these large primary carbides into solution. In the lO$in. square billet which had had a 1.8 : I reduction, more primary carbide was present at the centre, as would be anticipated from the original ingot segregation. Since the primary carbide distribution was the only difference noted between the two vacuum remelted samples, this may be responsible for the disparity in the toughness values.
137
Alloy steels in turbo-generatorcomponents
PRACTICAL SIGNIFICANCE AND INTERPRETATION OF RESULTS The object of the testing has been to establish the plane strain fracture toughness of various materials used in turbo-alternators. The specimens were taken from material representative of that used for component manufacture and the results have been assumed to be typical values. The critical defect size for fast fracture in certain components has been estimated using these results, assuming defects to lie normal to the direction of principal stress. Discs and rotors
The most critical region of disc and rotor forgings is at the central bore. The hoop stress is at a maximum due to the stress concentration effect of the bore and also this region is the position most likely to contain defects, which could be inclusions, unsealed porosity or hair-line cracks. The most serious defect is one which lies in an axial radial plane, reaches the bore surface and is infinitely long (length to depth ratio greater than 10). Table 4 shows the critical defect size calculated from the experimentally determined fracture toughness values for an applied stress equivalent to the uniaxial yield stress and an infinitely long surface crack. Conversion to an equivalent elliptical or circular defect, either at the surface or embedded, can be made by reference to Fig. 9 [ 141. It is to be noted that a typical design ratio of hoop stress to yield stress for discs and rotors at normal running speed is less than 0.5, which could increase to 0.72 on a 20 per cent overspeed. It is only under runaway conditions where stresses of yield point magnitude would occur. Table 4. Defect size at yield stress for large forgings at 20°C
Material No.
Material type
(zi.)
K,c
c (in.) 0.44 > 0.21 0.06
(ksi%)
la la lb
3 CMV
49.9
3 CMV
57.4
145 > 100 60
2
2f NCMV
54.5
90
0.14
3
3 NCMV
54.5
233 > 163
0.96 > 0.47
4 5
1 CMV 1 CMV
41.4 39.9
69 65
0.15 0.14
6a 6b 6b
1 CMVB 1 CMVB
41.8 34.0
57 104 1 81
0.10 0.49 > 0.30
7
12Cr
50.8
56
0.06
8a
12Cr (Vacuum remelted)
51.6
107
0.22
50.8
92 71
> 0.16 0.11
48.2
80
0.14
82.6
116
0.10
8a 8b 9 10
>
8Mn8Ni4Cr lSMn4Cr
Analysis for infinitely long surface defect (C = K,,‘/1.21 TR?).
138
D. V. THORNTON
b-6.44
t-“-v
w Surface
defects
Y
0 5.8
/
Internal
/
defects
Fig. 9. Flaw geometries for equivalent crack size effect, (a) Surface defects; (b) Internal defects.
In considering the likelihood of rapid fracture, an important aspect is the quality of forging which is accepted. The present inspection procedure includes a complete ultrasonic examination, magnetic inspection where possible, and a close visual examination of the bore. The present acceptance standards take into account the minimum possible operating temperature, the Charpy FATT, metallurgical cleanliness and the peak tangential stress at maximum overspeed. Cracks and porosity are not accepted. In the case of shafts and rotors the inclusions usually found are cigar shaped. Acceptance of these depends on number and dispersion, but the largest acceptable would be 0.25 in. long with a radial depth of approx 0.030 in. The maximum acceptable defect size on an L.P. disc is 0.25 in. dia near the bore surface (equivalent to 0.04 in. deep infinitely long surface defect). From Table 4 it can be seen that even at yield stress the critical defect size for rapid fracture is considerably in excess of these values. Before any advantage can be taken of this apparent safety, (e.g. relaxation of acceptance standards), it is necessary to obtain further information on such aspects, such as:(a) The effect of clouds of inclusions and interaction of defects. (b) The integrity of non-destructive testing techniques. (c) The equivalence of various types of defect and ‘flat bottomed holes’ (the term used to describe the size of defect from an ultrasonic defect echo). (cl) The rate of stable, slow crack growth of initial defects during service by fatigue or stress corrosion. The materials used as large forgings in the L.P. turbine and alternator, which have been chosen qualitatively and developed to give good impact transition properties,
Alloy steels in turbo-generator
components
139
show high values of fracture toughness. The test on the laboratory treated sample of 3% Cr - MO . V, where the proof strength was increased to 57 t.s.i. showed a marked loss in toughness. This indicates the danger which may arise in up-grading the strength of a material, even though it has a high toughness, without a quantitative assessment. The higher temperature service conditions dictate that the primary consideration for H.P. and I.P. rotors is creep resistance, and a 1% Cr * MO - V steel has been developed for these conditions. The toughness in various grades of this material is seen to be lower than for L.P. materials although the critical defect size at room temperature is still well in excess of the maximum defect acceptable by non-destructive testing standards. Intergranular fracture, which may be a result of temper embrittlement, occurred in one disc material even though it possessed a satisfactory fracture toughness. This could result from the slow rate used in cooling the disc after tempering in order to minimize internal stresses. Water quenching large forgings from the tempering temperature improves the impact transition properties, and therefore an improvement in fracture toughness may be possible. However, the effect of such a treatment on the residual stress pattern would have to be shown to be acceptable.
Alternator end rings The end ring has approximate dimensions of 2 ft long, 3 ft dia. and 3 in. wall thickness. The material is necessarily non-magnetic in order to minimize electrical losses. Cold expanded austenitic steels are used with a proof strength in the range 50-60 t.s.i. Various methods of expansion are currently being studied, including hydraulic and explosive methods, in order to increase the proof strength to 70 t.s.i. Two materials, 9 and 10, have been tested and the results used to estimate the critical defect size for rapid crack propagation at an applied stress equal to yield stress, Table 4. The higher strength material has a high degree of toughness and could tolerate a similar defect size to the lower strength material at a far higher design stress. These values are considerably larger than figures previously quoted for a martensitic stainless steel end-ring [ 151.
Bolts Turbine casing bolts are subjected to high temperature stress relaxation during service and are made from high proof strength, creep resistant materials. These materials generally have low impact properties and rapid fracture may occur, particularly on starting the turbine from cold when high transient stresses can be imposed while the bolt is relatively cold. The fracture toughness of a typical bolting material (material 11) has been determined and the value applied to the configuration of a 5f in. dia. bolt. The analysis used is for circumferentially notched round bar in tension, where the thread form has been assumed to be a sharp notch. The result suggests that failure would occur at a stress of 40 t.s.i. In practice the root of the thread is relatively blunt, which effectively increases the toughness of the component. However, during service creep damage may occur resulting in cracking at the first unsupported thread root, causing local sharpening and increasing the notch depth. In addition, deterioration in service may cause embrittlement and reduction in the fracture toughness of such components.
D. V. THORNTON
140
L.P. blades
The final stage L.P. turbine blades operate at a temperature of approximately 30°C in a condensing steam atmosphere. A stainless material is used to minimize corrosion, and the size of these forgings is approximately 40 X 6 X 24 in. The blades are supported by lacing wires, approximately +in. dia. which pass through the blade at an angle of about 20”, making an inclined elliptical hole in the blade where the blade thickness is about fin. An anticipated failure mechanism is the growth of a fatigue crack to critical size from the edge of a lacing hole. Tests have been carried out on 12% Cr stainless martensitic steels at present used as blades with a 36 in. aerofoil section, having a proof strength of 50 t.s.i. Shear fracture after general yield occurred in these specimens which simulated a fatigue cracked lacing hole stressed in tension. Future designs, blades 42 in. long or greater, will require higher strength stainless materials and the blade thickness may also be increased. In this case plane strain rather than plane stress failure may dominate. The plane strain fracture toughness of a number of potential materials has been determined and the values used to estimate the critical crack length for rapid fracture. The analysis used is for a centre cracked finite plate at a design stress of half yield strength. These results are shown in Table 5. The effective initial crack length due to the inclined elliptical lacing wire hole is of the order of I in., dependent on blade thickness, and several materials are clearly not satisfactory for such an application. With materials of high toughness, an additional advantage may be an effective further increase in toughness by domination of plane stress conditions. It is interesting to note that the fracture toughness of material 14 is reduced by a factor of approximately two as the proof strength is increased by about 10 per cent. The leading edge of large L.P. blades requires protection from erosion by water droplets contained in the condensing steam. This is accomplished at present by brazing Table 5. Defect
size at half yield stress for blading materials at 20°C WY
C
K/c
(t.s.i.)
12
82
47
0.17
13
71
36
0.13
14a 14b 14b 14c 14c
78 74
107 207 > 155 234 > 143
0.90 2.14 > I.64 2.38 > I.60
15
60
> 173
> 2-20
16
63
152
1.88
17
79
52
0.22
18
80
206 1 162
2.01 > 1.58
70
(ksiG)
(in.)
Material No.
Analysis for a centre cracked plate K,(.2 = d na/W where a = C/2 and W = 3 in.
W tan
Alloy steels in turbo-generator
components
141
on an erosion resistant shield, but possible alternative methods of protection include welding on an erosion resistant tip or local hardening of the blade material. This latter process introduces further difficulties, however, as martensitic stainless steels at the high strength level necessary to impart sufficient erosion resistance are susceptible to stress corrosion by hydrogen embrittlement in condensing steam. The limiting stress intensity below which stress corrosion cracking will not occur has been determined for two steels and has been found to be only 25 per cent of the plane strain fracture toughness. This information can be used to assess blade design and minimize the likelihood of slow crack growth by stress corrosion. SUMMARY The results of work completed to date have enabled the critical defect size for rapid fracture due to a single load application to be estimated for certain materials used as large forgings in turbo-machinery. In addition, quantitative assessment of potential materials for new designs has been made. The values determined have been assumed to be typical and further testing is necessary to establish lower bound values upon which design and acceptance procedures could be based. In the case of large forgings where initial defects may cause rapid fracture, it is essential that non-destructive testing methods are capable of detecting and accurately estimating the size and location of such defects. The present work covers the effect of single defects and further work is required to assess the interaction of defects and the effectiveness of clouds of small inclusions in initiating rapid fracture. The equivalence between linear elastic and general yield fracture mechanics has been shown by this work where excellent agreement between results using both approaches has been obtained for the plane strain fracture in a variety of alloy steels at various strength levels. There are many areas where a quantitative assessment of materials and design can aid the development of larger, high efficiency turbo-generators. This work begins to indicate the margin of safety from the hazard of fast fracture initiating from defects for virgin materials. Deleterious effects which may occur in service, such as slow crack growth by fatigue or stress corrosion and material embrittlement, are problems presently being investigated. Acknowledgements-The author would like to thank the Director of Research, English Electric Co, Ltd., for permission to publish this paper. The guidance of Professor A. A. Wells, particularly in the use of the Marcal-King analysis, is gratefully acknowledged.
REFERENCES [l] [2] [3] [4] [5] [6] [7] [8] [9]
E. E. Thum, Recent accidents with large forgings. Mef. Prog. 69.49-57 (1956). F. Buckley, Metallurgical aspects of turbine constructionII. fron & Stee/39.324-328 (1966). G. R. Irwin, Sfrucfural Mechanics, pp. 557-594. Pergamon Press, Oxford (1960). J. E. Srawley and W. F. Brown, Fracture toughness testing methods. ASTMSpec. Tech. Publ.No. 38 I ( 1965). A. A. Wells, Notched bar tests, fracture mechanics and brittle strengths of welded structures. Br. Weld. J. 12,2- I3 (I 965). E. T. Wessel, State of the art of the WOL specimen for K,c fracture toughness testing. Westinghouse Res. Rep. No. 67-l D6-BTLFR-RI (1967). D. F. Mowbray, A. J. Brothers and S. Yukawa, Fracture toughness determinations of A302B and Ni . MO. V steels with various size specimens. Trans. ASME, J. bus. Engng, Paper No. 66-Met. 1. A. J. Brothers, D. L. Newhouse and B. H. Wundt, Results of bursting tests of alloy steel discs and their application to design against brittle fracture. Gen. Elect. Res. Rep. Ger-22 18, presented at ASTM68rhA. Mrg(1965). F. H. Burdekin, Initiation of brittle fracture in structural steels. B.W.R.A. RepNo. E/l l/66 (1966).
D. V. THORNTON
142
[toI
P. V. Marcal and I. P. King, Elastic-plastic analysis of two-dimensional stress systems by the finite element method. Inr. J. mech. Sci. 9, I43- I55 ( 1967). L.P. discs. [111J. F. Knott, Application of fracture mechanics to an alloy steel used in turbo-generator J. Iron Steel Insf. 204. IO1 4- IO2 I ( 1966). on[ 1 l],J. Iron Steel Insr. 205,304 ( 1967). [I21 Correspondence in high purity 3.5 Ni-1.75 Cr-0.20C steel. Paper presented to r131 G. C. Gould, Temper embrittlement ASTM, October (I 967). of fracture information in specifications. ASTM Spec. Tech. Publ. No. 3X 1 [I41 W. F. Payne, Incorporation (1965). applications of Brittle fracture research. The Engineer 1151 M. G. Gemmill and K. Smith, Engineering 233,632-634(1967). (Received20
May
1968)
APPENDIX
K6 correction using the Marcal-King analysis [ lo] The Marcal-King method for the incremental elastic-plastic analysis of a two dimensional, plane strain stress system is based upon the finite element method of stress analysis and uses uniformly stressed triangular elements. The results derived from an analysis of a notched tension specimen, of notch depth to half width ratio of l-2 and total width 6 in., under plane strain conditions have been used. COD results have been converted to a stress intensity using (4), where
A more accurate value can be obtained if corrections are of measurement with respect to the crack tip, and to oy using the Marcal-King analysis. Table A 1 gives the displacement in the above specimens the crack tip (D) and values of applied to yield stress (a/~,). Table A
made to 6 for the position for stress intensification, at various distances from These values have been
I. Kelative displacement
D, ulu,
0
0.10
0.20
0.40
0.60
0.80
0.357 0.644 0.929 1.218 I.504
0.074 0.222 0.442 0.705 0.997
0.170 0.395 0.689 1.074 1.586
0.266 0.568 0.936 1.443 2.175
0,408 0.841 I.320 I .930 2.876
0.514
0601 I.219 I .874 2.616 3.634
D = distance
from crack
Table A2. Factored D, ulrr, 0,357 0.644 0.929 1.218 I.504
I.049 1.623 2.296 3.284
tip (in.). relative
displacement(f)
0
I /30
2130
4/30
6/30
8130
I .oo
2.30 1.78 1.56 I .52 I .59
3.59 2.56 2.12 2.04 2.18
5.51 3.78 2.98 2.73 2.88
6.95 4.72 3.68 3.25 3.30
8.12 5.48 4.23 3.71 3.64
1 .oo I .oo 1.00
Ia
D = fractional
distance
from crack tip.
Alloy steels in turbo-generator
143
components
Table A3. Stress intensification
a/u,
H
0.357 0644 0.929 1.218 1.504
1.385
I ,766 I.903 2.118 2.235
H _ or oil Table D, u/o, 0.357 0644 0.929 1.218
I.504
_ Maximum stress Yield stress
= Stress intensification.
A4. w
0
l/30
2/30
0.85 0.75 0.73 0.69 0.67
1.29 1.01 0.90 0.85 0.84
1.61 1.20 1.05 1.13 1.13
4130
6130
8/30
1.99
2.24 164
2.42 I .76 1.49 1.32 1.28
1.46 I .25
1.24 1.22
1.34 1.32 1.28
converted to a relative displacement (F) at fractional distances, with respect to notch depth, from the crack tip and are shown in Table A2. The stress intensification (H) at various applied to yield stress ratios are shown in Table A3. Hence
(5) and combining (4) and (5)
Values of m are shown in Table A4. These figures have been used to correct the K6 values calculated from (4) and are given in Table 3 of the paper as ratios of Klc determined by linear elastic techniques. R&meLa resistance a la rupture par tension plane de divers aciers allies utilises dans la fabrication de pieces forgees pour les turbo-generateurs a CtC determinee. On a montre que des valeurs Cquivalentes de resistance a la rupture sont obtenues a partir d’essais mecaniques de limite de resistance a la rupture generale et par allongement lineaire. Les valeurs obtenues ont servi a tvaluer la tendance rapide a la rupture des differents composants a partis de defauts simples. L’effet des facteurs metallurgiques sur le mode de rupture et la resistance a la rupture est illustre. ZusammenfassungDie Bruchzlhigkeit im ebenen Dehnungszustand von verschiedenen legierten Stahlen, die als Schmiedestticke in Turbogeneratoren verwendet wurden, wurde bestimmt. Es konnte gezeigt werden, dass Pquivalente Werte fur die Bruchzahigkeit dieser Werkstoffe erhalten werden bei Verwendung linearelastischer sowie vollplastischer Bruchmechanik. Die erhaltenen Werte wurden zur Beurteilung der Anfalligkeit fiir Rapidbruch, ausgehend von einzelnen Fehlern der verschiedenen Bestandteile, verwendet. Der Eintluss metallurgischer Faktoren auf die Bruchart und-z%higkeit wird veranschaulicht.