Effect of microstructure on the tensile flow stress of ferritic compacted graphite cast iron in intermediate temperature ranges

Materials" Science and Engineering, A l l l (1989) 27-34

27

Effect of Microstructure on the Tensile Flow Stress of Ferritic Compacted Graphite Cast Iron in Intermediate Temperature Ranges C. G. CHAO, T. S. LUI and M. H. HON

Department of Materials Engineering, National Cheng Kung Universit3: Tainan (Taiwan) (Received March 28, 1988; in revised form August 22, 1988)

Abstract

The effects of varying carbon content and ferrite grain size on the tensile flow stress of ferritic compacted graphite cast iron from room temperature to 773 K are investigated in this study. The tensile flow stress of ferritic compacted graphite cast iron is affected by the triaxial stress field developed in the matrix between graphite particles and ferrite grain size. The flow stress of this material is described by the equation o= ( 1 - k 1V~) (oo+k:d: -'/:) where Vg is the graphite volume fraction and d/ the ferrite grain size. 1. Introduction Compacted graphite cast irons have recently gained significant engineering interest. The strength, ductility and toughness of compacted graphite cast iron is superior to gray iron and approaches that of ductile iron. Its thermal conductivity, machinability and damping capacity are also better than ductile iron. For these reasons, compacted graphite cast iron is being considered for a number of applications such as ingot moulds, brake drums, cylinder heads, valve bodies, and vehicular engine blocks [1-3]. Therefore, understanding the characteristics of compacted graphite cast irons at elevated temperatures becomes essential. According to Lui and Yanagisawa's investigations [4, 5], the tensile flow stress of ferritic spheroidal graphite cast iron is influenced by carbon content and ferrite grain size in the intermediate temperature range. In addition, Yanagisawa and Lui's investigation [4] also showed that dynamic strain aging in ferritic spheroidal graphite cast iron occurred between 473 and 673 K using a nominal tensile strain rate of 2.8 x 10 -4 s - i . However, the influence of microstructure and dynamic strain aging on the flow stress of ferritic compacted graphite 0921-51)93/89/$3.51)

cast irons at intermediate temperature ranges has received very little attention. In this study, several specimens with different structures (different ferrite grain size and graphite size) were strained in the temperature range between room temperature and 773 K, to investigate the influence of microstructure associated with triaxial stresses and precipitation associated with dynamic strain aging on the flow stress of ferritic compacted graphite cast irons. 2. Experimental procedure The materials used were melted in a basic high-frequency induction furnace from a charge of pig iron with the following composition: 4.0-4.1%C, 0.1-0.3%Si, 0.02-0.04%Mn, 0.03-0.05%P and 0.01-0.03%S. The melt was spheroidized with Fe-45wt.%Si-8wt.%Mg alloy, inoculated with Fe-75wt.%Si alloy and cast into sand moulds. Three kinds of Y-block shape castings were produced as shown in Fig. 1 and Table 1. The chemical compositions as analyzed by emission spectrometry are listed in Table 2. The castings were annealed to a ferritic matrix by holding at 1173 Kfor 2 h, 1073 Kfor 1.5 h, 1013

\

/

Fig. 1. Y-block casting. © Elsevier Sequoia/Printed in The Netherlands

28

K for 5 h, and furnace cooling as shown in Fig. 2. Specimens with dimensions of 6 mm x 30 mm were machined from the parallel parts of the castings. The tensile tests were carried out in an MTS machine from room temperature to 773 + 3 K at strain rate of 6.6 x 10 -4 s - 1 . Each specimen was held at the test temperature for 10 min prior to the test to ensure a uniform temperature in the specimen. The graphite volume fraction Vg and average ferrite grain size df was measured with a "Buehler Omnimet Image Analyzer".

TABLE 1

3. Results

3.1. Microstructure of specimens Figure 3 shows the microstructures of specimens (section size A) with various carbon contents. The graphite particle number increases and the ferrite grain size decreases slightly with increasing carbon content. Figure 4 shows the microstructures of specimens (2.75% C) with dif-

Dimensions ofcastingsize

Casting

dimension (mm)

A B C

TABLE 2

a

b

c

65 30 15

90 90 90

280 150 65

Chemicalcomposition ofspecimensin wt.%

Specimen

C

Si

Mn

P

S

Mg

1A 1B 1C 2A 2B 2C 3A 3B 4A 5A 6A 7A

2.75 2.75 2.75 3.28 3.28 3.35 3.86 3.86 1.65 2.57 3.08 3.55

2.53 2.53 2.53 2.42 2.42 2.64 2.41 2.41 2.81 2.82 2.79 2.92

0.086 0.086 0.086 0.028 0.028 0.028 0.034 0.034 0.117 0.153 0.111 0.079

0.024 0.024 0.024 0.026 0.026 0.026 0.026 0.026 0.025 0.026 0.027 0.030

0.012 0.012 0.012 0.014 0.014 0.014 0.014 0.014 0.026 0.011 0.014 0.014

0.017 0.017 0.012 0.020 0.016 0.017 0.018 0.017 0.011 0.012 0.014 0.013

1173 K-2h Ld Pr"

1073 K-l.Sh F'C---~"1013 K - 5h

OC

Ld 0_. Ld

TIME Fig. 2. Condition of heat treatment.

Fig. 3. Microstructures of specimens (section size A) with (a) 2.75% C, (b) 3.28% C and (c) 3.86% C.

29 i

•

Size A

LU N

80. 7 F'~

© 80 Iii I,FIF

L~ LL

A

40 L

2.5

2.0

5.'0

I

3:s

4.o

CARBON CONTENT (wt~) Fig. 5. Influence of casting size and carbon content on ferrite grain size.

Z

15

• • *

Size A B C

•

C12 IJ klJ I 0 ___J 0 > 5

-v Ck

< r-F

V@~0.038x (~C)

0

0

I

I

I

1

3

4

CARBON CONTENT (wt %) Fig. 6. Relation between graphite volume fraction V~ and carbon content. Fig. 4. Microstructure of specimens (2.75% C) with different section size: (a) 65 mm thick; (b) 30 mm thick; (c) 15 mm thick.

ferent casting section size. T h e graphite particle size decreases and also ferrite grain size decreases with decreasing casting section size. Figure 5 shows the influence of casting section size and carbon content on ferrite grain size. T h e ferrite grain size becomes finer as the casting section size is decreased and carbon contents increased. Figure 6 shows that the graphite

volume fraction Vg and carbon content are linearly related by Vg=0.038 ×(percentage C) irrespective of section size. 3.2. Flow stress f r o m room temperature to 773 K Figure 7 shows several nominal stress strain curves for specimens with 3 . 2 8 % C deformed between r o o m temperature and 773 K. At r o o m temperature (curve 1) and 373 K (curve 2), no serrations were observed, while periodic serrations appeared at 473 K (curve 3). With further

30

[22Mpa

0,4:4

o

1 573 K

W

/

/

/

/

/

/

,,,~cc:~

STRAIN

STRAIN

Fig. 8. Influence of carbon content on dynamic strain aging.

Fig. 7. Stress-strain curves of 2A specimens (3.28%C) at temperatures between room temperature and 773 K.

increase in temperature up to 573 K (curve 4) the number of serrations increased. In curve 5 and 6 in Fig. 7, it was found that between 673 K and 773 K the serrations disappeared. These results are similar to the findings of other researchers [6-8], and are due to the interaction of dislocations and interstitial solute atoms during deformation. Baird [9] gives an extensive discussion of the effect of inhomogeneous flow and interstitial content on ductility. Serrations occur because of a shortage of mobile dislocations due to the locking by interstitial atoms. The stress increases until the correct mobile density is restored and then drops because deformation can process at a lower stress [10, 11 ]. At lower temperatures, the deficiency is not too serious, arising only because of a decrease in the rate of re-mobilization. The serrations are, therefore, of small amplitude and low frequency such as curve 3. At higher temperatures, the shortage is much more significant because of the pinning of the mobile dislocations themselves. Under these conditions the serrations are characterized by a large amplitude and high frequency such as curve 4. The stress-strain curves of Fig. 8 illustrate the effect of carbon contents on dynamic strain aging. The temperature range of dynamic strain aging between 473 and 573 K in higher carbon alloys (2.57%C, 3.08%C) shifts to the temperature range between 573 and 673 K in lower carbon alloys (1.65% C). Serrations appear on the flow curve when the net carbon on the dislocation exceeds the critical amount of carbon needed to form sufficient atmospheres. When the net carbon on the dislocation line is less than the critical amount, even though atmospheres exist, the amount of net carbon is not sufficient to

produce serrated flow [12]. When the carbon content decreases, there will be less source in the matrix available for carbon to form Cottrell atmospheres around moving dislocations. With an increase in source spacing, the serrations occur at higher temperatures at which carbon atoms can be mobile enough to pin dislocations stopped at obstacles. Thus, lower carbon alloys need higher temperatures for similar serrations as the higher carbon alloys. Figure 9 illustrates the influence of temperature on yield stress with various microstructures. Considering the graphite in compacted graphite cast irons as voids [4, 5], the effective sectional area decreases when the carbon content increases. The yield stress should therefore increase with decreasing carbon content. In addition, the yield strength should also increase with decreasing casting section since the microstructure becomes finer. However, the morphology of compacted graphite has an irregular shape, so that the stress concentration condition is different and the effective sectional area of the vertical z axis (applied stress direction) is different. Therefore, the fluctuation of yield stress increases. The yield stress does not fully increase with decreasing carbon content and casting section as shown in Fig. 9. It is worth noting a yield stress plateau formed from 473 K. Garde et al. [13-15], suggested that the athermal region in the yield stress vs. temperature diagram corresponds to stress increment due to strain aging added to a curve of monotonically decreasing strength with increasing temperature. Figure 10 shows the influence of temperature on flow stress (e = 0.03) with various microstructures. The flow stresses have the same tendency as the yield stresses. A flow stress plateau formed from 473 to 6 7 3 K is due to dynamic strain aging.

31

('C) 100

200

'

I.

(°C)

300

'

400

500

'2.76~C',

100

35O

200

i

300

400

!

I

Si~e'

500

!

2.75 %C,

!

Size • A • B

300

cO

250

0

(a)

LL 2OO 300

15ol

b

(a)

300

,

400

,

500

300 -m'~x

,

600

I

I

I

600

700

800

Size

,

700

800

3so[

Size

\

5 0

4OO

'; c

3.28c" 3.35%C

cO u3

C

•

I_.d Of

LLI I-- 250 u')

:

I\

250

C] J LIJ

"

~

©

>-

2OO (b)

*

I

I

J

I

I

300

400

500

600

700

3.86 ~C ,

Size •

'~ ~ x ~ ~

3OO

I"

A

i 400 3

800

250

U') Of CO LLI

:b) _J LL 200 ' 300

t 500 .

•8

i 600

•

J 700

i 800

: %C

CO cO 25O I11 rF

• B 2OO

•

d~200

.

0

_2 D_

i

I

300

)-

(c) 150 a 300

,

400

400

I

I

500

600

I

700

I

800

TEMPERATURE ( K) 500

600

TEMPERATURE

700

800

(K)

Fig. 9. Influence of temperature on yield stress between room temperature and 773 K: (a) 2.75% C; (b) 3.28% C; (c) 3.86% C. (A, B, C: casting size).

4. Discussion

4.1. Effect of microstructure on flow stress Ever since Hall [16] and Petch [17] introduced their well-known relationship between the yield strength of low carbon steel and grain size, a great deal of effort has been devoted to explaining the relationship from a fundamental point of

Fig. 10. Influence of temperature on flow stress (e=0.03) between room temperature and 773 K: (a) 2.75%C; (b) 3.28% C; (c) 3.86% C. (A, B, C: casting size.)

view and applying the relationship to different metals and alloy systems. It is now generally recognized that a uniaxial tensile stress aeq (equivalent stress) in a single phase polycrystalline aggregate can be represented by the Hall-Petch relationship Ocq = O0 +

kdf-1/2

(1)

where a 0 is the friction stress, k is the slope and df is the ferrite grain size.

32

Equation (1) can be expected to apply to the ferrite matrix of compacted graphite cast irons as well. According to Yanagisawa and Lui [5], the flow stress of spheroidal graphite cast irons can be expressed as a function of Vg and de- 1/2 by

k2dr -1/2)

0=(1 - k 1Vg)oeq=(a - k, Vg)(Oo +

at room temperature ao.o3 =(1 - 2 . 1 8 Vg)(232.1 + 3 8 . 8 1 dr- I/2) MPa

(4) at 373 K ao.o3 = (1 - 2.61 Vg)(235.7 + 43.51 dr-1/2) M P a

(2)

(5)

where (7o, k I and k2 can be determined by multiple regression analysis. The values of o 0, k I and k 2 of compacted graphite cast irons at various temperatures are presented in eqns. (3)-(8).

at 473 K ao.o3 = ( 1 - 1.54 Vg)( 176.6 + 43.12 dr-1/2) M P a

(6) at 573 K

o0.03 = (1 - 2.63 Vg)(91.43 + 87.61 df-1/2)MPa

'(3) 400

RY.

/

573 "K

O'ooa=(1- 2.63 V g ) ( 9 1 . 4 3 + S7.61 df ½ )

z

E

ve 0 0.079 0.097

Z w

.-, oo:,o,%

w rr" 350

•

,

: o°:I 0.154

Fo

300 Ld

>rn

• ,,~"

I

Er --1.64

v Vv

300

/:

>-

R =0.88

,

. . . .

, , , 350

300

R =0.96

,(;)

/ 25O

400

(b) I

I

I

I

250

(7 BY FITTED EQUATION(HPa) 350

Er =0.43

133 b

'/~,

I

I

i

;

I

300

I

673 K

(7 BY FITTED EQUATION(MPa)

/

773K 0"°03 =0-

Z

A

I,I 300 r'r"

V/

•

2'67 Vg)(78"3

v

/

V

IA

0

mm

0 0

V

V

/

>- 200 m

25O

l /

eee ~

Ld re"

Er =1.37 R -'0.74

*

/

÷ 61:5 df~-)

*

CO < LtJ /

/

•

FZ Ld 250

Z Ill

I

350

/

7

CO < LLJ Z >133

0"o a =(1-1.

bJ rY

CO < LId

b

350

/

}-Z b_J

2

~

V~)(153.3+40.77df-ln)MPa (7)

ao.o3=(1 - 1.1

Er--2.19-R =0.75

b

(c) I

I

250

I

I

I

I

I

I

I

I

/

I

300

O" BY FITTED EQUATION(HP~:)

350

ee ,

I

200

(d) i

i

i

i

~

i

i

i

250

O' BY FITTED EQUATION(MPa)

Fig. 11. Comparison between flow stress by measurement and fitted equation: (a) room temperature; (b) 573 K; (c) 673 K; (d) 773 K. (Er, standard error deviation; R, multiple correlation coefficient.)

33

at 6 7 3 K and o0.,,3=(1 - 2 . 6 7 Vg)(78.3+61.15 d,-'/2)MPa (8) at 773 K. After obtaining these equations, the precision of fit was checked and the results are shown in Fig. 11. If the compacted graphite particles are considered to be voids with no strength [4, 5], the flow stress of ferritic compacted graphite cast irons can be presented by eqn. (9). It shows a simple decrease of flow stress a z along the z axis as effective sectional area decreases.

o_=(1- ~)O{q

(9)

This equation means that the kl coefficient of in eqn. (2) is unity. However, the experimental results in eqns. (3)-(8) show that k I is larger than unity. From the results, it is considered that the homogeneous stress state of the ferrite matrix is changed due to the presence of graphite in the matrix, i.e. a stress concentration of triaxial stress occurs around the compacted graphite in the matrix. These results are similar to those found in spheroidal graphite cast irons [5]. The parameter k~, a sort of triaxiality parameter, is affected by temperature to the extent that it changes from 2.63 at room temperature to 1.1 at 673 K and back to 2.67 at 773 K. The reason for k~ decreasing at 573 and 673 K is that the yield stress of the matrix decreases and plastic deformation of the matrix easily occurs to release stress concentration while the temperature increases to 573 and 673 K. However, while the temperature reaches 773 K, the specimen will neck at e = 0.03. This is shown in Fig. 7. k~ returns to 2.67 at 773 K because triaxial stress occurs by necking besides forming a stress concentration around graphite.

irrespective of the internal stress field. This tendency results in a homogeneous distribution of dislocation [15, 18, 19]. Therefore, the friction stress reverts to low value at 773 K. Figure 12 shows the values of o0 at room temperature and 773 K to be smaller than those at other temperatures. The k, coefficient of df 17_ in eqn. (2) was taken as the extent to which dislocations pile up at grain boundaries. The friction stress dominates the flow stress from 373 K which is affected by dynamic strain aging. The influence of grain boundaries on flow stress is not obvious, i.e. the values of k~ are small. On the other hand, the

100

('C)

200

300

400

500

250

d? DA 12:: 15C Z O F-Q) Q£ L

:

I

I

50

X

I00 '

48 0

5; 0

600 '

700 '

TEMPERATURE (K) Fig. 12. Friction stress o, of the ferritic compacted graphite cast iron in eqn. (2). D.S.A.: dynamic strain aging.

i001

('C) 100

300

200

1

400

500

i

i

I

h" 4. 2. Effect o/friction stress and grain size on flow stress The friction stress o0 is thought to be a measure of the intrinsic resistance of matrix to the motion of a dislocation. A high frictional stress in the temperature range between 373 and 673 K is due to dynamic strain aging promoting the deceleration of dislocations, i.e. there is a clustering of dislocation at high internal stress fields. This means that the strain hardening rate increases due to dynamic strain aging. Above the temperature range of dynamic strain aging, 773 K, dislocation tends to move at the average velocity

1.~ < c12

5o

°f o

D,S.A.

I

3OO

t

1

I

L

I

I

400

500

600

700

800

TEMPERATURE (K)

Fig. 13. k, coefficient of dr-i/2 in eqn. (2) for the ferritic compacted graphite cast irons.

34

influence of grain boundaries on flow stress is dominant at room temperature a n d 773 K because the friction stresses are small at these temperatures. Therefore, Fig. 13 shows the values of k 2 at room temperature and 773 K to be larger than those at other temperatures. 5. Conclusions (1) A tensile flow stress plateau of ferritic compacted graphite cast irons occurs owing to dynamic strain aging from 473 to 673 K. (2) The flow stress o f ferritic compacted graphite cast irons is affected by the triaxial stress field developed in the ferrite matrix between graphite particles. (3) The flow stress of ferritic compacted graphite cast irons can be expressed by the following equation as the function of graphite volume fraction Vg and ferrite grain size df-1/2 with the constants kl, k2 and a 0.

o=(1

- k 1 Vg)(O'o + k2df -1/2)

Acknowledgments The authors thank Professor O. Yanagisawa for comment on the manuscript. We are also grateful for the financial support provided by the National Science Council of Taiwan under Grant NSC 75-0405-E006-16.

References 1 E. Borghigiani and C. Marinari, Trans. Am. Foundrymen'sSoc., 90(1982) 529-549. 2 P. A. Green and A. J. Thomas, Trans. Am. Foundrymen's Soc,, 87(1979) 569-572. 3 C. M. Dunks and K. B. Truner, Trans. Am. Foundrymen's Soc, 89(1981) 575-586. 4 0 . Yanagisawa and T. S. Lui, Trans. Jpn. Inst. Met., 24, No. 12 (1983) 858-867. 5 0 . Yanagisawa and T. S. Lui, Metall. Trans. A, 16 (1985) 667-671. 6 S. C. Park, L. P. Beckerman and R. E. Reed-Hill, Metall Trans. A, 14 (1983) 463-469. 7 I. S. Kim and M. C. Chaturvedi, Trans. Jpn. Inst. Met., 28, No. 3 (1986) 205-212. 8 A.K. Sachdev, Metall. Trans. A, 13(1982) 1793-1797. 9 J. D. Baird, The Inhomogeneity of Plastic Deformation, 191-220. 10 Y. Bergstrom and W. Roberts, Acta Metall., 19 (1971) 815-823. 11 S. C. Park, L. P. Beckerman and R. E. Reed-Hill, Metall. Trans. A, 14 (1983) 463-469. 12 R. W. Hayes and W. C. Hayes, Acta Metall., 32 (1984) 259-267. 13 A. M. Garde, A. T. Santhanam and R. E. Reed-Hill, Acta Metall., 20(1972) 215. 14 S. I. Hong, Mater. Sci. Eng., 76 (1985) 77-81. 15 S.I. Hong, Mater. Sci. Eng., 79(1986) 1-7. 16 E. O. Hall, Proc. Roy. Soc. London, Set. B, 64 (1951) 474. 17 N.J. Petch, J. lronSteellnst., 174(1953) 25. 18 D. J. Lloyd, D. W. Chang and M. C. Chaturvedi, Acta Metall., 23 (1975) 93-100. 19 L. P. Kubin and Y. Estrin, Acta Metall., 33 (1985) 397-407.