The behaviour of dynamically loaded pipes with circumferential flaws under internal pressure and external bending loads

Nuclear Engineering and Design 96 (1986) 99-113 North-Holland, Amsterdam

99

THE BEHAVIOUR OF DYNAMICALLY LOADED PIPES WITH CIRCUMFERENTIAL UNDER INTERNAL PRESSURE AND EXTERNAL BENDING LOADS

D. S T U R M , W. S T O P P L E R

FLAWS

a n d J. S C H I E D E R M A I E R

Staatliche Materialprfifungsansalt (MPA), University of Stuttgart, Pfaffenwaldring 32, 7000 Stuttgart 80, Fed. Rep. Germany Received 9 June 1986

For the determination of the load bearing capacity, the deformation, leak-before break- and fracture behaviour, bending tests on vessels out of ferritic and austenitic materials containing a circumferential defect were performed. The vessels were loaded over a temperature range extending from 20°C to 580°C by internal pressure and additionally by an externally applied quasi-static or dynamic bending moment. For the test a 4-point bending test rig providing a bending moment up to 3 MNm was used. Parallel to the tests the failure curves as well as the leak-before-break curves were calculated with the aid of engineering type calculational methods. No significant difference as far as load bearing capacity is concerned, can be established between quasi-static and dynamic loading. In most cases fracture occurred after considerable plastic deformation in the test cross section. The engineering type methods applied to the theoretical calculation of the leak-before-break curve on the corresponding defect curve proved to be conservative with respect to safety.

1. Introduction In the design, construction and calculation of light water reactor primary circuit components the current, dedicated Codes such as K T A 3201.2, T R D 302 and the A S M E Pressure Vessel Code Section III are to be applied. In this context the dimensioning of the load bearing cross-sections has to be such that the accommodation of the internal pressure and the external forces, whilst observing the limiting values laid down for the primary stresses, is assured and that neither unacceptable deformation nor fracture will occur. Since under upset conditions such as earthquakes, aircraft crash, water hammer, etc., high additional loading from primary bending can arise along with the internal pressure loading due to operation, the load bearing capacity reserves of piping must be known and all the more so if it contains faulty circumferential joints. For the experimental analysis of the mechanical behaviour of piping of important safety related significance under upset conditions, appropriate tests have been carried out for some years at M P A Stuttgart on pipes of both ferritic and austenitic materials within the scope of various research projects, [1,2]. These research

projects, aimed at demonstration of the integrity of piping and promoted partly by the Federal Ministry for Research and Technology and partly by industry and plant operators, may be divided into three individual areas with respect to their objectives as shown in table 1. Whilst in the impact tensile tests and component tests, the test pieces are subjected to a single, dynamically-applied loading, tests for the demonstration of strain tolerance are basically conducted under multiple i.e. cyclic, loading. The impact tensile and component tests provide information on the load bearing capacity and deformation behaviour whereas tests for the demonstration of cyclic strain tolerance for which bar and tubular specimens are likewise used give information on cyclic deformation behaviour. The present report concerns component tests consisting of bending tests on vessels which are loaded by internal pressure and additionally by an externally applied quasi-static or dynamic bending moment.

2. Load bearing capacity, leak-before-break criterion The load bearing capacity of a vessel containing a circumferential defect when subjected to loading by

0 0 2 9 - 5 4 9 3 / 8 6 / $ 0 3 . 5 0 © E l s e v i e r S c i e n c e P u b l i s h e r s B.V. ( N o r t h - H o l l a n d Physics P u b l i s h i n g D i v i s i o n )

4

I

~

Cyclic Strain Tolerance

Test for Demonstration of

- Component Tests

- Notch impact Bend Tests

Combination of the Results with the

Testing in Accordance with KTA

Replacement by Notch Impact Bend

Testing (External Systems) and its

Underpinninq of the Droppinq of NDT

Upset Conditions

lDuantifieation of the Safety Margin

I

Behaviour

Local Plastic Flow

Lowest Operating Temperature

- Loadinq Rate

Loading

Determination of Upset Related

from Circumferential 3oints

aqainst Catastrophic Failure Arisinq

Demonstration of Tolerance of

[

Conservative Assurance Against Brittle

~

Range

cance in Nuclear Power Plant under

Respect to Single Impact Type Loading Single Monotonic/Dynamic Loading

in the Region of Small Numbers of Cycles and Large Plastic Strain

the Load Bearinq Capacity Behaviour

of Pipinq of Safety-Related Signifi-

of Cracked Circumferential 3oints in

Pipes of Different' Materials with

Alternating Deformation Behaviour

I

II

Produetion and Improvement of Data on

Pipe Bend Tests

OF PIPING

PROOF OF INTEGRITY

Deformation and Load Bearinq Behaviour

Impact Tensile Tests

Table 1 Working areas for the proof of integrity of piping

E-

2-

r~

r~

101

D. Sturm et a L / Behauiour of loaded pipes with flaws

Table 2 Calculation of failure moment of a cylindrical vessel with a circumferential defect under internal pressure loading Plastic collapse

v--'-v uJ :~O

CATASTROPHIC-~ FAILURE t

I ! Mb

CALCULATION OF FAILURE M O M E N T OF V E S S E L S WITH CIRCUMFERENTIAL DEFECTS

~'

THEORY:

P L A S T I C COLLAPSE IN NET CROSS SECTION

DEFECT FORM! CIRCUMFERENTIAL N O T C H CRITERION FORMULA <~ Z

N~.~L E A K BEFORE X ~ J . FAILURE CURVE

e: ill

\\~,.~

Xl-"

LEAKAGE ~ . _

SENDING

t,/s=O,8 ,.,

CIRUMFERENTIAL SLIT

FLOW STRESS I

2qo~r2s(2sinl~ - f sin o.)

2q o~r2s{ 2cosl3- sin eL]- 2p, r3cos~3

MOMENT FACTORS

,,,

-f

r, - 2ors-p, r2

Oo

180 °

f_tk - 5

360 °

q=1.15

ANGLE 20~ of =~- ( R~.F~o.~

Fig. 1. Leak before break diagram.

q=0.98

internal pressure a n d an external b e n d i n g m o m e n t can be described clearly by means of the leak-before-break diagram, fig. 1. D e t e r m i n i n g factors are the strength a n d d e f o r m a t i o n characteristic properties of the pipe

RANGE OF VALIDITY

THIN WALLED VESSELS

tk

Q01i

3 MNm- Rohrbiegevorrichtung 3 MNm-PIPEBEN0tNGDEVICE

',~

Aufiager REAl:lIONBEARING

Prbfrohr TEST PIPE Oruckstempel

10 MNm

BENDING DEVICE

LOAOBEARING

Stutze SUPPOR1

Zylinder[3MN]

CYLINDER[3MN]

Grundrahmen

ST VESSEL~ /

Sm

BASEFRAME

Fig. 2. 3 MNm pipe bending rig.

Fig. 3.10 MNm pipe bending rig.

o,

Table 3 Calculation of the failure stress of a cylindrical vessel with defects in the circumferential direction on loading by internal pressure and external bending moment Classical bending theory

VESSELS WITH STRESS

CIRCUMFERENTIAL AXIAL

AT LOCATION

OF MAXIMUM

STRESS

+

(PRESSURE)

STRESS

Ormx

SURFACE

=

BENDING (AXIAL

A4A_WAp,

•

FAILURE

CRACKS STRESS

+

LOAD)

BENDING (OUTER

%().b)A,,p,

STRESS MOMENT)

+

~Mb

o~x =Rm

F L A W SIZE

INNER S U R F A C E

OUTER SURFACE

SLIT

LOCATION OF MAXIMUM STRESS

ELEMENTS CROSS SECTION AREA :

Ao = ~ 2 4 d,

AREA OF RING :

AQ = ~-(do 2- d. 2)

AREA OF FLAW:

A, = ~ orc(~(d, 2- d,2)

ACTING AREA :

A.= Ao* A,

LE~K~THOF LEVER ARM :

a =9* +

+ arc ot(d~ - d, 2)

-~- orc o~(d~2 -d, 2) A.= Ao 9 . .~z (cos(z)n ;,~}f=docos=~dl ;},u, docoS=, d,

~'. -~ cosCt

=

DISPLACEMENT OF CENTER OF GRAVITY :

MOIvENT OF INERTIA

x-AXIS: -~-- AXIS:

x=Q fufa.90° x = i f~Jr c z > 9 0 °

b =1_ sin~(d~3-d. 3) Ao*A,

b =0 1 sin~d~-d~) ~- Aq-Af

=71~2sin aidS" d,3l Aq-A,

1 sin~(d~-d, 3) ~ Aq-A,

Jx = 6-~E(d~-d,~) J, = J,- 92(A~r-A,)

~

-IA,. c~

•

J,- 92(AcA,)

~, ~,

J, - 92(~-A,)

- IA,..s~.~.(go-C~],~, 8

16

- [A,+

~Y

b

A.

~2=

,~2 ~lld, 2

e ,-:",', Is

103

D. Sturm et al. / Behaviour of loaded pipes with flaws

PRE CYLI

Fig. 4. Rig for quasistatic and dynamic bend tests. material and the defect geometry. The so-called "leakbefore-break curve" which may be related to the failure load of a through-wall defect (slit) divides the "leakage" region from that of "catastrophic failure". If, for example, a failure point lies below this limit curve then only snap through of the defect ligament in the wall thickness direction need to be considered. Should however the failure point lie above the curve then a catastrophic failure i.e. catastrophic disruption involving a guillotine break with high energy release is to be expected. The failure curves for pipes with surface defects, determined in bend tests and also the associated leak-before-break curve are however displaced towards smaller defect lengths with increasing internal pressure.

3. Methods of calculation

Not only the failure curves but also the leak-beforebreak curves may be calculated with the aid of engineering type calculational methods. These may be described as follows: - Proceeding from the assumption of plastic col-

lapse in the net cross-section then by the formulation of an equilibrium of forces the loading in the weakened cross-section which is considered as ideally fully plastifled, may be calculated, table 2. Failure is then assumed

I~,NOTCH

~I

()

I1~11

~

D

G

E

s

MEASUREMENIPOINTDESIGNATION • TEMPERATURE HT STRAINGAGE DEFLECTION ./-L CRACK OPENING • PRESSURETRANSDUCER -

Fig. 5. Instrumentation layout.

-N

104

D. Sturm et al. / Behaviour of loaded pipes with flaws

when the stress resulting from this loading attains a particular material characteristic value, usually the flow stress. Depending on the type of defect (slit or notch) an empirically determined correction factor q is applied to the flow stress [3]. - A further possible means of calculation is derived from classical bending theory, table 3. The expressions developed for the individual defect forms are based on the exact calculation of the moment of inertia in the weakened annular cross-section and the displacement of the neutral axis by the defect. Failure is then deemed to have occurred when the stress at the most highly stressed point reaches a material characteristic value which experience has shown for ductile materials to be the ultimate tensile stress [4,5]. Both methods have been examined for various pipe dimensions and defect geometries in both ferritic and

austenitic materials. Further possibilities of calculation, albeit at a considerably higher cost, are offered by Finite and Boundary Element methods. By these methods additional fluid dynamics processes which arise can also be handled.

4. T esti ng technique 4.1. Bending test rigs

At present there are four 4-point bending test rigs available at the MPA. Fig. 2 shows a 4-point bending test rig with a 3 MN pressure cylinder which provides bending moments up to 3 MNm. Pipes up to diameters of about 600 mm with wall thicknesses up to 25 mm may be tested in this rig.

measured acceleration distribution

I

I I--

--]

tl

01tt )~

mass o f beam m

l blfx,t)

Support

force:

1,1 F (t) B1(t) = ~2

-2 m • 11

1

b(x,t)

dx

0

B e n d i n g moment a t c e n t r e o f b e a m :

11

11

11

M (~-, t l = T

m

B 1 (tl + ~1

,~

b(x,t)

11

• (T

x) dx

o

Fig. 6. Determination of the inertial forces from the acceleration distribution in accordance with ref. [6].

D. Sturm et al. / Behaviour of loaded pipes with flaws

Since October 1985 a further bending test rig having a 10 M N pressure cylinder giving bending moments up to about 10 M N m , fig. 3, and taking pipes up to 800 mm diameter and 50 mm wall thickness is available. Tests on vessels of 800 mm outer diameter and 47 mm wall statically preloaded by an internal pressure of 150

bar are under way. The pressure medium for the pipes is air or pressurised water, with the test temperature lying in the range from ambient up to 300°C. The pressure cylinder for the generation of the bending moment can be operated with either liquid, gas or a propellent charge so that tests may be conducted with different

~

EOMSI.

105

q)=O

DMS3 m~ending Axis

grad =

%.120°~

Mean Longitudinal Strain: (Tensile Force/Internal Pressure)

ELZ -

Bending Strain over Pipe Circumference :

Eb

Total Strain over Pipe Circumference:

E

Maximum Bending Strain:

Angle of Bending Axis with respect to Axis I~0 : 0:

E l +E2 +E 3 3

= Ebmax sin(~0 +~B )

: Ebmax sin(~0 +kpB) + ELZ

E I -ELz Ebmax =sin(kPb +kpo) 3 (E I -ELZ ] ~0B = arctan(E I + 2 ~ 3E:LZ

Bending Stress:

O b : EE b

Bending Moment :

Mb =ObW b = E Wb E b

-Eo

Fig. 7. Relationships for calculation of longitudinal and bending stratus and the position of the bending axis in pipes under internal pressure loaded by bending.

106

D. Sturm et al. / Behauiour of loaded pipes with flaws

Crack Flank Displacement Fracture

'N~,.X

~endinq Deflection

/

/

_\

>1 20

30

40

~.~

"a

58

70

BO

TIME Cms] Fig. 8. Time dependence of bending deflection, bending strain, vessel internal pressure, ram pressure and crack flank displacement in dynamic test on Pipe 8 having partial circumferential internal notch. Table 4 Material properties Material

Temp. Structure l)

0.2% Proof Stress Rpo,2

°C

X10CrNiNb189

11NiMnCrMo55

Notch Impact Energy

Z

Av

%

J

417

622

18

48

50

428

605

22

68

200

BM

548

701

24

69

100

WM

428

635

--

37

120

BM

28O

608

48

66

170

697

.

RT

RT WM

X10CrMoV121

A5

Reduction of Area

BM

20MnMoNi55 17MnMoV64

Rm MPa

NiMoCr-Special Cast 300

Ultimate Elongation Tensile Strength at Fracture

.

.

.

108

BM

367

428

40

81

116

WM

374

46O

26

69

106

54

80

550

580

BM

RT

BM

I) Base Metal Specimens longitudinal

Creep Strength 13 min 310 MPa 658

710

29

D. Sturm et aL / Behaviour of loaded pipes with flaws

107

Table 5 Test parameters and results of pipe bending tests NO

MATERIAL

1

DIMENSIONS CRACKSIZE O. DIAMETER WALLTHICKN LENGTH DEPTH 2-t tk LENGTH ;

BOO

grd

mm

360

4O,B

CRACK POSITION

20 MnMoNi 55

3

NiMoCrSPEZIAL MELT

2500

17 MnMo V 64

)68 II 3600

Pi I Mb MPa I kNm

°C

BASEMATERIAL

2

TEMP LOADINGAT THEORETICAL CAPACITY INSTABILITY LOADING OF LBADING EXPERIMENT THEORY

23,7

Pi = 22,2 MPB

1,07

26,7

Pi :

23,1 MPa

1,16

16,4

Pi =

15,3 MPa

1,07

320

3B,9

47

OUTSIDE

270 41,B

REMARKS

CATASTROPHIC

310

FAILURE 4 • 5 6 135

X 10 CrNiNb 18 9

9

14,2

120

2500

I0

}60

ii

•

3,7

DEPOSIT MAT. CIRCUMF. SEAM OUTSIDE

615

642

. X 20 CrMo V 12 1 12

447 24 5000

13

17,0 (73)

14,2

DEPOSIT MAT.

12,1

CIRCUMFERENT

12,4

SEAM INSIDE

3,6

90

22,5

II0

20,8

20

17,0

OUTSIDE DEPOSIT/BASE MATERIAL OUTSIDE

-

-ii

NiMnCrMo 55

12,5

15

12,5

16

Mb = 3 4 , 5 kNm

1,10

CATAST.FAIL.

41

Mb =

3 0 , 3 kNm

1,35

LEAKAGE

36

Mb :

28,2 kNm

1,28

LIMITEDFAIL.

Mb :

6 7 , 2 kNm

1,19

CATASTR.FAIL.

kNm

1,08

550

7,1

391

Mb = 389

kNm

1,00

172

Mb = 102,5 kNm

1,68

Mb = 47,7 kNm

2,05

73,9

M b = 35,2 kNm

2,10

36,7

Mb =

2,17

LEAKAGE

97,9

20

IBO

16,9 kNm

4.2.2. Measurement of the bending moment The controlling loading parameter in bend tests is the bending moment acting on the test piece. Although the determination of the bending moment in the static case is still possible simply by measurement of the external forces and the geometry of the force application points, this is no longer admissible in the dynamic case without further considerations due to the mass inertial forces which arise. At high rates of load increase up to fracture the bending moment acting on the test piece is usually determined as follows: Measurement of the force acting on the pipe taking into account the inertial forces, by calculation from direct acceleration measurements made on the pipe, fig. 6 [6]. Measurement of the strains in a cross-section which remains elastic during the test and from this the calculation which is possible, of the actual bending moment acting in the most highly stressed cross-section, fig. 7. -

4.2.1. Instrumentation For the recording of the time dependent variation of bending moment and fracture opening area both the bending rig and the test pipe were instrumented as shown in fig. 5. In the dynamic tests data acquisition was achieved by means of a PCM unit with a total recording rate of about 560 Kword/s.

CEASED

38,2

Mb = 3 5 )

MATERIAL

4.2. Measurement technique

1,25 0,76

80

SLIT

levels of machine stiffness as well as long or short rise times for the bending moment application. The rig on which a number of the tests here described were carried out is shown in fig. 4. This rig is suitable for either quasi-static or dynamic load application. For dynamic loading the bending moment is applied by the detonation of a propellent charge in the combustion chamber of the pressure cylinder.

kNm

Mb = 95,5 kNm

381

BASE

120 llO0

M b : 515

1,23

7,3

• ithout crack 90

kNm

578

1)9 14

Mb : 502

8,0

without crack

7

8

4,3 50

-

108

D. Sturm et al. / Behaviour of loaded pipes with flaws 120

4.2.3. Determination of the fracture opening area For the determination of the jet forces after snap through of the ligament, the recording of fracture opening area a n d its development as a function of time are of decisive importance. To a good approximation the fracture opening area can be calculated from the instantaneous values of the crack flank displacement and crack length. For a sufficiently accurate m e a s u r e m e n t of these two quantities n o n - c o n t a c t inductive sensors developed by M P A Stuttgart, were employed, the signals of which remain unaffected by the outflowing medium.

[kNm] 100

I-'Z LLI

80

~

X 10CrNiNb 18 9 0.D.xs =133x 1/.,,2m m ~

/

V10dyn.(t,t=0.25s)

O (_'3 Z

5

I '~

50

Z LU

, ~ CATASTROPHICFRACTURE

03

5. Materials, test vessels and conduct of the tests

.d

The mechanical properties of the ferritic a n d austenitic piping materials investigated up to the present time are summarised in tabular form in table 4. The test pipes having outer diameters between 133

< Z n.'LU I-X LLI

L,0

\ ~ ¥ 8 star. (tk =0.B S) ~'~--\'(][~V9 dye It~ =0.89s)

L

y

o LEAKAOE

O0 LEAKAGE ~ "~tx., C % ~. CATASTROPHIFRACTURE LEAK-BEFORE-BREAK CURVE.j%%----~o.09"~

1000

[kNm]

i

i

0

17MnMoV 64 O,D-xs = 368xllmm

0 Pi = 8MP(a r

90

100

270

grd

360

NOTCH CIRCUMFERENTIALANGLE20. 100

200

300

mm

DEFECTLENGTH(EXTERNAL)

(00

8O0

£

Fig. 10. Load bearing behaviour of Pipes 8. 9 and 10 in comparison with the defect curves calculated in accordance with table 3.

I.--

Z LLI 0~" 600~___~Vl,

stat. (tk=O,36s)

0

z_

r--i z uJ m

800x/.,7

< 7 n~ LU

I-× w

MM /d-,Tx2/., I 139x12,5

DIMENSIONS QQxs IN

t~=O.30s+O,37s -- - -

~oo

~oo

I

LEAKAGE~ 120

133xl/.,,2

'

[%1 ~-f~l 60

136Bxll I

180

CATASTROPHIC FRACTURE I I 2&O

1

300 grd 350

NOTCH CIRCUMFERENTIALANGLE 20L 200

t,O0

600

0~)0 mm I000

'nS6

DEFECT LENGTH(EXTERNAL) Fig. 9. Load bearing behaviour of Pipes 4 and 5 in comparison with the defect curves calculated in accordance with table 3.

Fig. 11. Comparison between calculated (from table 3) and experimentally determined failure loads.

D. Sturm et al. / Behaoiour of loaded pipes with flaws

109

5O X 10 CrNigb 18 9 O.O.xs. =133xlL, mm2

[kNm 5O "5"

FRACTURE

/ Z Z

30

f ~

20

10

O

f

TATIC Test 8

5

10

15

[mm]

20

BENDINGDEFLECTION f Fig. 12. M o m e n t - d e f l e c t i o n curve from tests 8 and 9.

'J

w

.

t

~z

o

Fig. 13. Pipe section 8 after failure (X 10 CrNiNb 18 9).

2

3 ~:4

5

6

7 :

8

9 i,,s

Fig. 14. Pipe section 9 after failure (X 10 CrNiNb 18 9).

110

D. Sturm et aL / Behaviour of loaded pipes" withflaws

mm and 800 mm with wall thicknesses from 11 mm to 47 mm cover to a large extent, the range of piping of interest. The same may also be said for the types and positions of defects (through- and part-wall defects both internal and external, in base- and weld-metal respectively). The testing temperature was predominantly 20°C but did however extend up to 580°C. Tests were conducted partly without and partly with static internal pressure loading in which the level of loading was pitched at the particular operating conditions of interest. In three cases failure of the test pipe was achieved by quasi-static, monotonic increase of the internal pressure without supplementary loading by a bending moment but in all other cases it was achieved by the application of a moment. In order to determine the influence of a dynamic bending load on the load bearing capacity, appropriate spot-check tests were carried out. By way of example the time traces of some selected measurements are reproduced in fig. 8. It may be seen that the rise of the bending moment represented by the bending strain goes from 10% to its maximum value within 26 ms. From this a mean rise rate of bending moment of 1 M N m / s results. In certain tests values of up to 100 M N m / s were achieved. Fracture of the ligament is signified by a sudden fall in the bending strain, the onset of pressure loss in the pipe and the rapid increase in crack flank displacement. It is noteworthy that the ram pressure where the bending moment is generated by rapid combustion of a propellent, rises further in spite of the fall in the bending strain. From this it is clear that under dynamic loading conditions the determination of the bending moment from the ram force is not permissible. The bending deflection speed in the dynamic tests lay between 1 m / s and 10 m / s and the strain rate in the compression fibres laying opposite the defect between 1/s and lO/s.

SECTION A2/1 Fig. 15. Fracture surface of Pipe 4 (17 MnMoV 6 4).

sufficient accuracy by means of engineering type calculational methods (compare table 3). As the tests No 4 and 5 or 8 and 9 show, no significant difference as far as load beating capacity is concerned, can be established between quasi-static and dynamic loading.

6. Test results

6.2. Deformation behaviour

6.1. Load bearing capacity

Fig. 12 shows the form of the bending momentbending deflection curve for both the quasi-static test No. 8 and the dynamic test No. 9. Both curves follow an almost identical path, the maximum bending deflection at fracture in dynamic loading indeed being somewhat less than for quasi-static loading but this is however, attributable to the smaller failure moment. This again results from the slightly deeper notch (t k/S = 0.89) in test No. 9 compared with test No. 8 (tk/s = 0.85).

The results of the bending tests are given in table 5 and to the extent that quasi-static and dynamic loading was applied are also shown graphically in figs. 9 and 10. All of the quasi-static and dynamic tests exhibited a greater experimental failure load than the calculated value, fig. 11. In addition the type of failure, whether leakage or catastrophic failure could be determined with

D. Sturrn et al. / Behaviour of loaded pipes with flaws

111

A2./1 Fig. 16. Section A2/1 from the fracture surface of Pipe 4 (17 MnMoV 6 4).

6.3. Metallographic and fractographic investigations

The metallographic and fractographic investigations of the fracture region of the austenitic pipes Nos. 8 and 9 showed that the fractures ran mainly in the weld metal, where, in fact, they were initiated, figs. 13 and 14. In the quasi-static and dynamically loaded pipes at times there was considerable necking of the remaining ligament before fracture. In the ferritic pipes Nos. 4 and 5 the fractures likewise ran mainly in the weld metal where they were

initiated. Both fractures showed shear and flat fracture areas, the shear fracture fraction in the statically loaded pipe No. 4 being about 90% and about 60% in the dynamically loaded pipe No. 5, figs. 15 to 18. Whilst the crystalline portion of the fracture surface only occurred in the immediate locality of the notch extension in the statically tested pipe, fig. 16, in the dynamically tested pipe it was present over about half the pipe circumference lying largely symmetrically about the notch, fig. 18.

SECTION B1/1

'ig. 17. Fracture surface of Pipe 5 (dynamic test) 17 MnMoV 6 4).

"SECTION B2/1

Zig. 18. Sections B I / I and B2/1 from Pipe 5 (17 MnMoV 6 4).

SECTION 3 211

J

D. Sturm et al. / Behaviour of loaded pipes with flaws

113

7. Summary and conclusions

References

The c o m p a r a b l e b e n d i n g tests carried out at the M P A on vessels of different materials and dimensions, weakened by circumferential notches and subjected to loading by internal pressure and an externally applied b e n d i n g moment, showed no discernible difference in b e h a v i o u r as regards load bearing capacity irrespective of the manner, either static or dynamic, in which the single, monotonically rising b e n d i n g load was applied. T h e fracture surface of the dynamically tested ferritic pipes had a somewhat smaller fraction of ductile fracture than those of the quasi-statically tested ones. In all other cases fracture occurred after considerable plastic d e f o r m a t i o n in the test cross-section. In all cases the engineering type m e t h o d applied to the theoretical calculation of the leak-before-break curve on the corresponding defect curve proved to be conservative with respect to safety.

[1] K. Kussmaul, W. Stoppler, D. Sturm and P. Julisch, Ruling-out of fractures in pressure boundary piping, 1AEA-SM 269/7 (March 1983). [2] G. Bernecker, G. Gnirss, M.D. Schulze and P. Julisch, High temperature burst test on X20 CrMoV 12 1 components for proof of the exclusion of fracture for the piping of THTR 300, 9th MPA Seminar, October 13 and 14, 1983, Stuttgart. [3] M.F. Kanninen et al., Instability predictions on circumferentially cracked Type 304 stainless steel pipes under dynamic loading, NP-2347, Vol. 1, Research Project T 118-2 Final Report (April 1982) Appendix E. [4] D. Sturm et al., Research Programme "Phenomenological Vessel Burst Tests" 150 279, Final Report Phase l, July 1985, Staatliche Materialpriifungsanstalt, University of Stuttgart (1985). [5] P. Julisch, W. Stoppler and D. Sturm, Exclusion of rupture of safety-relevant piping systems by component tests and simple calculations, Transactions of the 8th International Conference on Structural Mechanics in Reactor Technology, Vol. G, 1985, Brussels, Belgium. [6] K. Brandes, E. Limberger and J. Herter, Kinetic limit load bearing capacity of impact loaded reinforced concrete components, BAM Research Report 90 (April 1983) Berlin.