Failure Analysis & Integrity Assessment of the Steam Drum ...

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Failure Analysis & Integrity Assessment of the Steam Drum

on the Incineration and Heat Recovery Plant at a Sewage Works

Report Number M16044

Date XXXXXXXXXXXXXXXXXX

Purchase Order Number XXXXXXXXXXXXXXXXXX

Customer XXXXXXXXXXXXX XXXXXXXXX XXXXXXXXXXX XXXXXXXXX XXXXXX

Customer Contact XXXXXXXXXXXXXX

Authors:

S L Bagnall J M Brear Metallurgical Manager Director R-Tech Consultants John Brear – Plant Integrity This report is Copyright © R-Tech Consultants Limited and John Brear – Plant Integrity Cyfyngedig and is technically and commercially confidential.

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Summary

The waste incinerator at the subject sewage works is equipped with a heat recovery system

for steam generation.

In 2014 a leak occurred on the steam drum, at the 5 o’clock position on the circumferential

weld between the cylindrical shell and the dished head at the north end. It was repaired

without detailed investigation. A further water leak was detected in December 2015. On

de-lagging a through-crack was found, also at the 5 o’clock position but on the

circumferential weld at the south end; further cracks, also through, were observed on the

feed-water sleeves. As work has progressed, several further cracks have been found at

various nozzle positions in the lower part of the drum.

R-Tech Consultants and John Brear – Plant Integrity (JB-PI) have been contracted to provide

a failure analysis, integrity assessment, and to advise on decision making with regard to

future management of the drum. The following conclusions and recommendations arise:

Conclusions:

1. The cracking experienced is due to a stress corrosion mechanism, quite possibly

caustic in nature, driven by the welding residual stresses and an unfavourable water

chemistry.

2. Drum material chemistry, mechanical properties and microstructure are acceptable,

but the welding procedures used in construction were not optimised and have given

rise to high weld residual stresses and poor microstructures.

3. Fracture mechanics analysis confirms that welding residual stresses provided the

driving force for cracking and gives predictions on defect stability that match

observation.

4. Whilst transverse cracks on the structural welds are constrained from propagation

and opening, a significant risk has been identified associated with cracking of nozzle

welds leading to nozzle severance.

5. Calculations show that an exclusion zone of 12 metres should be adequate if the

drum is returned to service after repair.

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Recommendations:

1. Advice be obtained on the procedures necessary to ensure correct operational

control of water level and water chemistry – for this drum if repaired, for its

replacement, and for other steam-raising plant within the company.

2. Should a repair option be followed, then a suitable weld procedure should be

employed. A temper-bead technique is recommended.

3. Following repair, hydrostatic testing at design pressure plus ten percent and at an

ambient temperature of no less than 10°C is considered appropriate.

4. Should the present drum be returned to service, then a carefully considered 12

metre exclusion zone is recommended.

5. Should a replacement option be followed, then a like-for-like approach should be

adopted, with an improved weld procedure imposed and PWHT considered.

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Contents 1. Section 1 – Failure Analysis……………………………………………………………………………………5 2. Section 2 – Integrity Assessment……………………………………………………………………………45

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Section 1 - Examination of a Welded Section from a Steam Drum 1. Introduction Following a leak of the steam drum at the XXXXXXXXXXXXX in December 2015, NDT

inspection was carried out on both the external and internal surfaces. This inspection

revealed a longitudinal crack across the south end circumferential weld. A circumferential

crack was also detected in the weld at the feed water nozzle, at the south end. The positions

of these cracks in the steam drum are shown in figure 1.1. The cracks, in-situ, are shown in

figures 1.2 and 1.3. R-Tech were informed that following a previous issue in 2014, at the

north end of the drum, all the structural welds and longitudinal welds were inspected using

NDT with no evidence of any indications. This suggests that the cracks present at the south

end had occurred within 1-2 years.

Figure 1.1 Diagram of steam drum, showing failure locations

R-Tech Consultants Ltd received a section incorporating the crack detected across the

circumferential weld with the request to determine the cause of failure. R-Tech were

informed that the operating pressure was 50 bar at a temperature of 260-270°C. The steam

drum had been shut down 4-5 times per year. The material was advised as being carbon

steel, BS1501 223 grade 490B. The wall thickness was 32 mm. The section, as-received, is

shown in figure 1.4.

Crack across circum weld

Crack at feed water nozzle

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Figure 1.2 External surface of steam drum, showing cracking across south end circumferential

weld (supplied by XXXXXXX)

Figure 1.3 External surface of steam drum, showing cracking in weld at feed water nozzle

(supplied by XXXXXX)

Shell Head

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Figure 1.4 Steam drum section, as-received

2. Visual Examination

The longitudinal crack evident across the circumferential weld in the steam drum section is

shown in figures 2.1 and 2.2. Prior to removal of the section, the crack tips had been drilled

to prevent propagation.

Figure 2.1 Steam drum section, inner surface, showing cracking across circumferential weld

Shell

Head

Weld

Shell

Head

Weld

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Figure 2.2 Steam drum section, outer surface, showing cracking across circumferential weld

A section incorporating the crack was removed and both surfaces were ground and polished

using a hand-held polisher. It was clear that the holes did not correspond to the location of

one of the crack tips on the inner surface and neither of the crack tips on the outer surface,

see figures 2.3 and 2.4. On the inner surface, two cracks were evident; the smaller fine crack

appeared to step down from the main crack. At the inner surface, the cracking was

approximately 45 mm in length, and on the outer surface was approximately 27 mm in

length.

Figure 2.3 Inner surface of cracking (after polishing)

Shell

Weld

Head

Crack tip

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Figure 2.4 Outer surface of cracking (after polishing)

The through-thickness (side) profile of the section incorporating the crack was ground and

polished in order to determine the postion of the crack in relation to the weld, see figure

2.5. The black marks evident on the section indicate the length of the crack at both the inner

and outer surfaces. It is clear that on the inner surface, some propagation had occurred into

the shell parent material (approximately 6 mm). At the outer surface, there was no evidence

of any crack propagation into the parent material. Examination of this section showed a

double V weld preparation, as expected. An additional weld was also evident at the inner

surface (indicated in figure 2.5). This weld overlapped what is thought to be the original

weld i.e. the weld had been produced after the original weld. It is therefore thought to be a

repair weld.

Crack tip

Crack tip

Large Crack

Small crack

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Figure 2.5 Side profile of section with crack, showing crack length in relation to the weld at

both the inner and outer surfaces

Sections were taken through the cracking (as shown in figures 2.6 and 2.7), so that the

resultant section fell apart revealing the fracture surface. The fracture surface, shown in

figure 2.8, was associated with a tightly adherent dark brown scale. This scale was analysed

by EDX, see later in section 5.

Figure 2.6 Inner surface of section incorporating the crack, showing sections taken

Shell Head

Outer

Inner

Repair Weld Original Weld

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Figure 2.7 Outer surface of section incorporating the crack, after taking sections shown in

figure 2.6

Figure 2.8 Section through cracking (sample B), showing fracture surface

One half of the fracture surface was cleaned using inhibited hydrochloric acid to remove the

corrosion prodcuct, see figure 2.9. The fracture surface was brittle in nature with no

evidence of any deformation or necking. Furthermore, the flow lines evident on the surface

indicated that the crack had initiated from the inner surface (highlighted in figure 2.9).

Inner surface

Outer surface

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Figure 2.9 Section through cracking (sample B), showing fracture surface (after cleaning)

3. Scanning Electron Microscopy

The clean fracture surface was examined using a Zeiss EVO 60 scanning electron microscope

(SEM) with Oxford INCA Energy Dispersive X-Ray (EDX) microanalysis. Multiple cracks were

evident across the fracture surface, see figures 3.1 and 3.2. At higher magnification, the

majority of the fracture surface exhibited micro-void coalescence which is indicative of a

ductile fracture mechanism, see figures 3.3 and 3.4. In some areas, a pearlite/ferrite

microstructure was evident on the fracture surface, see figures 3.5 and 3.6.

Inner surface

Outer surface

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Figure 3.1 Towards midpoint of fracture, showing cracking

Figure 3.2 Towards outer surface of fracture, showing cracking

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Figure 3.3 Towards inner surface of fracture, showing micro-void coalescence

Figure 3.4 Towards outer surface of fracture, showing micro-void coalescence

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Figure 3.5 Towards midpoint of fracture, showing pearlite/ferrite microstructure on the

surface

Figure 3.6 Towards midpoint of fracture, showing pearlite/ferrite microstructure on the

surface

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4. Metallographic Examination

4.1 Surface of Cracking

Examination of the polished sections shown in figures 2.3 and 2.4 at higher magnification

showed the cracking to be transgranular and branched in nature, see figures 4.1 to 4.4.

Figure 4.1 External surface of section, showing large crack (red arrow) and small crack (yellow

arrow)

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Figure 4.2 External surface of section, showing tip of large crack

Figure 4.3 External surface of section, showing small crack

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Figure 4.4 Internal surface of section, showing cracking

4.2 Section through cracking (section D)

Section D in figure 2.7 was ground and polished on a through-thickness face to a one micron

finish and etched using a 5% nital solution. A macro image of this section, shown in figure

4.5, shows the crack propagating through the parent material. At this section, the crack had

not penetrated through the entire wall thickness.

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Figure 4.5 Section D, showing crack along the majority of the wall thickness

The crack tip evident towards the outer surface was very blunt, see figure 4.6 and 4.7. The

remaining length of the crack exhibited some branching (see figures 4.8 to 4.11) and was

associated with corrosion product. In some areas, the corrosion product was associated

with elemental copper, see figure 4.12.

Weld

HAZ

Parent (head)

HAZ

Weld (repair)

Outer

Inner

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Figure 4.6 Section D, showing crack tip in weld at outer surface

Figure 4.7 Section D, showing crack tip in weld at outer surface

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Figure 4.8 Section D, showing cracking at inner surface

Figure 4.9 Section D, showing cracking at inner surface

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Figure 4.10 Section D, showing cracking in inner surface weld

Figure 4.11 Section D, showing cracking in parent material

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Figure 4.12 Section D, showing elemental copper associated with corrosion product in the

crack

The microstructure of the inner (repair) and outer welds was bainitic in nature with ferrite

evident at the prior austenite grain boundaries, see figures 4.13 to 4.16. These prior

austenite grains were very coarse for both the inner and outer welds. The heat affected

zones were also bainitic in nature (see figure 4.17) with the area close to the repair weld

and outer weld, also exhibiting very coarse grains. The parent material consisted of ferrite

and pearlite, see figure 4.18.

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Figure 4.13 Inner weld microstructure

Figure 4.14 Inner weld microstructure

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Figure 4.15 Outer weld microstructure

Figure 4.16 Outer weld microstructure

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Figure 4.17 HAZ microstructure, close to outer weld

Figure 4.18 Parent material (head) microstructure

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5. EDX Analysis

The corrosion product evident on the fracture surface and within the cracking was analysed

by EDX. The areas analysed are shown in figures 5.1 to 5.4 and the results are shown in table

5.1.

Figure 5.1 Uncleaned fracture surface, showing areas analysed

Figure 5.2 Corrosion product associated with cracking, showing areas analysed

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Table 5.1 EDX microanalysis of corrosion product

Sample ID Approximate Weight % Element (Note)

O Na Si P S Cl K Ca Ti Cr Mn Fe Cu Zn Mg

Fig 5.1-spec 1 19 0.7 0.6 0.5 0.3 1.9 72 5.7

Fig 5.1-spec 2 21 0.6 0.4 0.6 0.2 1.3 70 5 0.6

Fig 5.1-spec 3 5.7 0.5 0.4 0.4 2.3 85 4.8 0.6

Fig 5.1-spec 4 34 0.5 0.3 0.2 1.0 61 1.8 1.5

Fig 5.1-spec 5 32 0.5 0.4 0.3 1.6 63 2.6

Fig 5.1-spec 6 18 0.5 0.4 0.3 0.2 1.8 76 2.3

Fig 5.2-spec 1 32 0.3 67

Fig 5.2-spec 2 15 0.2 0.8 83 0.7

Fig 5.2-spec 3 54 22 0.5 1.8 1.5 2.8 6.4 0.5 8.2 1.1

Fig 5.2-spec 4 29 0.5 62

Fig 5.3-spec 1 37 0.6 0.2 1.0 61

Fig 5.3-spec 2 26 0.6 0.2 0.2 1.2 71

Fig 5.3-spec 3 30 0.5 0.5 1.0 68

Fig 5.3-spec 4 18 0.4 0.2 1.3 80

Fig 5.4-spec 1 18 0.6 0.6 0.8 78 1.3 0.5

Fig 5.4-spec 2 21 0.6 0.4 0.4 78 0.5

Fig 5.4-spec 3 24 0.8 0.3 0.3 2.2 0.8 67 1.1 2.9 0.4

Fig 5.4-spec 4 9 0.4 0.9 0.9 87 1.5 0.8

Note

The quantification procedure strictly applies to polished surfaces and therefore the results on rough surfaces such as particulates may only be considered semi-quantitative. The results indicate only the relative proportions of each element.

The corrosion product was found to consist predominantly of iron and oxygen. In some

areas, small amounts of chlorine, potassium, sulphur, calcium, sodium, phosphorus and

magnesium were detected. In one area, a significant amount of sodium was detected.

Copper and zinc were also evident in many of the analyses.

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6. Vickers Hardness Testing

Vickers hardness testing was conducted across the weld profile according to BS EN 6507-

1:2005 using a 10kg load. The hardness positions are shown in figure 6.1 and the results are

shown in table 6.1.

Figure 6.1 Section through weld, showing position of hardness measurements

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Table 6.1 Vickers hardness results

Position Hardness indent Hardness HV10

Parent (Head)

1 166

2 162

3 162

Repair weld HAZ – (head side)

4 202

5 195

6 204

Repair weld HAZ – (head side) near fusion line

7 255

8 251

9 231

Repair weld

10 263

11 236

12 233

Inner weld HAZ - (head side)

13 217

14 217

15 219

Inner weld

16 233

17 227

18 242

Inner weld HAZ – (shell side)

19 210

20 196

21 213

Parent (Shell)

22 180

23 175

24 160

Outer weld HAZ (shell side)

25 217

26 206

27 222

Outer Weld

28 215

29 240

30 233

Outer weld HAZ (head side) near fusion line

31 228

32 224

33 227

Outer weld HAZ - (head side)

34 203

35 197

36 203

The hardness levels in the parent materials were considered acceptable for a low carbon

steel. The hardness levels in the welds and heat affected zones (particularly near the fusion

line) were slightly higher than would be expected, though the levels were not considered

excessive.

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7. Charpy Impact Testing

Charpy impact testing was conducted in the inner and outer welds and in both parent

materials to BS EN ISO 148-1 at ambient temperature. The specimens taken from the welds

are identified as 9 and 12 in figure 7.1. These samples were notched so that the crack would

propagate through the thickness of the section, see figure 7.2. The specimens taken from

the head parent material are identified as 14 in figure 7.1 and as 8 and 13 for the shell

parent material. The position of the notch for these specimens is shown in figure 7.3. The

results are shown in table 7.1. For sample 14, sub-size specimens were used due to

insufficient material.

Figure 7.1 Steam drum section, showing sections taken

14

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Figure 7.2 Through thickness (side profile) of section 12, showing position of notch for weld

specimens

Figure 7.3 Section 13, showing position of notch for parent material specimens

Table 7.1 Charpy impact results at ambient temperature

Sample Identity Geometry of test specimen Impact Toughness J

Sample 8 (inner surface) Shell parent 10 x 10 x 55 2mm V notch 171

Sample 8 (mid section) Shell parent 10 x 10 x 55 2mm V notch 180

Sample 13 (outer surface) Shell parent 10 x 10 x 55 2mm V notch 186

Sample 14 (inner surface) Head parent 10 x 7.5 x 55 2mm V notch 133

Sample 14 (outer surface) Head parent 10 x 7.5 x 55 2mm V notch 140

Sample 9 (inner surface) Inner weld 10 x 10 x 55 2mm V notch 90

Sample 9 (outer surface) Outer weld 10 x 10 x 55 2mm V notch 102



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Charpy impact testing was also undertaken at 0°C across the weld of the section identified as 3 in figure 7.1 above, at both the inner and outer surfaces. These samples were also notched so that the crack would propagate through the thickness of the section. Furthermore, charpy impact testing was undertaken at 0°C in the head parent material on the section identified as 11 in figure 7.1 above. The results are shown in table 7.2.

Table 7.2 Charpy impact results at 0°C

Sample Identity Geometry of test specimen Impact Toughness J



Sample 11 (inner surface) Head parent 10 x 10 x 55 2mm V notch 169

Sample 11 (outer surface) Head parent 10 x 10 x 55 2mm V notch 112

The fracture surfaces of the impact samples are shown in figures 7.4 to 7.12.

Figure 7.4 Sample 8 (shell-inner surface), ambient test temperature

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Figure 7.5 Sample 8 (shell-mid section), ambient test temperature

Figure 7.6 Sample 13 (shell-outer section), ambient test temperature

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Figure 7.7 Left image – Sample 14 (head-inner surface), ambient test temperature

Right image – Sample 11 (head-inner surface), 0°C test temperature

Figure 7.8 Left image – Sample 14 (head-outer surface), ambient test temperature

Right image – Sample 11 (head-outer surface), 0°C test temperature

Figure 7.9 Sample 9 (Inner weld), ambient test temperature

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Figure 7.10 Sample 9 (Outer weld), ambient test temperature

Figure 7.11 Left image – Sample 12 (inner weld), ambient test temperature Right image – Sample 3 (inner weld), 0°C test temperature

Figure 7.12 Left image – Sample 12 (outer weld), ambient test temperature

Right image – Sample 3 (outer weld), 0°C test temperature

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The impact properties for the parent and weld materials were considered relatively high at both ambient temperature and 0°C. The fracture surfaces for those samples tested at ambient exhibited a ductile fracture.

Samples 9-inner, 9-outer, and 12-outer exhibited some crystalline areas, though these areas

were very small. For the samples tested at 0°C, the fractures were predominantly ductile,

with the exception of sample 11-outer surface, which exhibited a predominantly crystalline

(brittle) fracture. This difference in fracture surface corresponds to the fall in impact

toughness observed for this specimen.

8. Chemical Analysis Samples of the inner weld, repair weld (at the inner surface), outer weld, head parent and shell parent were analysed by ICP-OES and the results are shown in table 8.1.

Table 8.1 ICP-OES analysis results

Element Shell parent Inner weld Repair weld Head parent Outer weld Grade 490

Carbon 0.19* 0.12* 0.12* 0.19* 0.08* 0.20 max

Silicon 0.22 0.62 0.51 0.24 0.67 0.10-0.50

Manganese 1.14 1.96 1.74 1.17 2.04 0.90-1.60

Sulphur 0.003* 0.005* 0.005* 0.003* 0.007* 0.030 max

Phosphorus 0.012 0.023 0.019 0.012 0.023 0.030 max

Chromium <0.01 <0.01 <0.01 0.01 <0.01 0.25 max

Molybdenum 0.01 <0.01 0.01 0.01 <0.01 0.10 max

Nickel 0.05 0.02 0.03 0.04 0.02 0.75 max

Aluminium 0.018 0.012 0.012 0.018 0.011

Copper 0.02 0.14 0.08 0.03 0.12 0.30 max

Titanium <0.01 <0.01 <0.01 <0.01 <0.01

Niobium 0.01 0.01 0.01 0.01 <0.01

Vanadium 0.01 0.01 0.01 0.01 0.01

Nitrogen 0.007* 0.008* 0.008* 0.007* 0.007*

Tin 0.005 <0.003 0.003 0.005 <0.003

Antimony <0.003 <0.003 <0.003 <0.003 <0.003

Arsenic <0.003 <0.003 <0.003 <0.003 <0.003

*analysed by combustion

The parent materials complied with requirements specified for the 490 grade.

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9. Tensile Testing

Two specimens were machined from section 6 shown in figure 9.1 (i.e. the shell parent

material) for tensile testing at ambient temperature to BS EN ISO 6892-1:2012. The results

are shown in table 9.1.

Figure 9.1 Steam drum section, showing sections taken

Table 9.1 Tensile test results

Sample Identity Upper Yield

(MPa) UTS (MPa) Elongation (%)

Sample 6 (inner surface) Shell 355 544 29.0

Sample 6 (mid surface) Shell 375 531 28.6

BS1501 223 grade 490B 340 min 490-610 20 min

The tensile properties of the shell parent material complied with the limits specified for

BS1501 223 grade 490B.

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10. Discussion Failure of the steam drum had occurred due to the presence of a through-wall longitudinal

crack across one of the circumferential welds. An additional crack was also evident in the

weld of the feedwater nozzle, though this had not been received for examination. Both

failures had occurred at the south end of the steam drum, at approximately the 5 o’clock

position.

For the cracking examined, failure had initiated from the inner surface of the weld. The

fracture surface was associated with a tight adhering dark-brown scale. Corrosion product

was also evident inside the cracks, which in some areas was associated with particles of

elemental copper. This may have originated from copper based materials in the system.

Corrosion products from these materials dissolve in the feedwater. The metal ions react

with the steel and leave copper particles mixed with other deposits.

The corrosion product was found to consist predominantly of iron and oxygen. In some

areas, small amounts of chlorine, potassium, sulphur, calcium, sodium, phosphorus and

magnesium were detected. In one area, a significant amount of sodium was detected.

Copper and zinc were also evident in many of the areas analysed.

The fracture surface, on a macro scale, was brittle in nature with no evidence of any

deformation or necking. On a micro scale, the majority of the fracture surface exhibited

micro-void coalescence which is indicative of a ductile fracture mechanism (i.e. a tensile

overload). The brittle fracture observed, on a macro scale, is due to constraint created by

the thick section. With geometric constraint, plastic strain may be concentrated and fracture

can occur without visible macroscale deformation (1). In some areas, the ferrite pearlite

microstructure was evident on the surface of the fracture. This is sometimes observed when

corrosion is a contributing factor to crack propagation.

The cracking was transgranular and slightly branched in nature. The crack measured 45 mm

long at the outer surface and 27 mm long at the inner surface. This is considered to be a

very short crack length relative to the crack propagation through the entire wall thickness

(32 mm). Although some crack propagation was evident in the parent material, this was

confined to the weld area. The confinement of the crack to the weld area suggests that

residual stress in the weld was a major driving force for crack propagation and explains why

the crack had not extended significantly outside of the weld where the residual stresses

would be reduced.

The weld had been produced with a double V geometry, as expected. However, an

additional weld was evident at the inner surface which overlapped the original weld; this is

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thought to be a repair weld. The presence of this weld could be associated with an increase

in residual stress in the joint. Furthermore, the additional weld would cause grain growth in

areas of the original weld and heat affected zone. Examination of all three welds (outer,

repair and inner) revealed a bainitic microstructure with very coarse prior-austenite grains

highlighted by ferrite grains throughout the weld materials. A coarse grain structure was

also evident in the heat affected zones close to the weld materials. The presence of a coarse

grain structure in both the heat affected zone and weld materials indicates high

temperature input during welding.

The hardness levels in the weld and the HAZ close to the fusion line were slightly higher

than would normally expect, though these were not considered excessive. The impact

properties for the parent and weld materials were also considered relatively high at both

ambient temperature and 0°C. The fracture surfaces for those samples tested at ambient

exhibited a ductile fracture. For three of the specimens taken from the weld, some areas

were crystalline in nature, though these areas were very small. For the samples tested at

0°C, the fractures were predominantly ductile, with the exception of one of the samples

taken from the head parent material, which exhibited a predominantly crystalline (brittle)

fracture. This difference in fracture surface was also associated with a fall in impact

toughness. This may suggest that at this temperature the material is beginning to approach

the ductile to brittle transition temperature curve.

The parent materials complied with the chemical analysis requirements specified for the

490 grade. The tensile properties of the shell parent material complied with the limits

specified for BS1501 223 grade 490B.

The presence of corrosion product associated with the cracking and the branched nature of

the cracking suggests that fracture had occurred due to a stress corrosion mechanism. The

presence of some sodium, magnesium and potassium within the cracking may indicate

caustic cracking, though this not certain since caustic cracking is normally intergranular in

nature. On the other hand, there have been cases in the literature where caustic cracking

was reported as transgranular. Caustic cracking occurs when stressed carbon steel is

exposed to hot alkaline solutions. Caustic cracking can occur over a wide range of

temperatures, generally between 80-350 degrees C, see figure 10.1 (2). This mechanism can

also occur over a range of caustic concentrations; the lower limit is generally 3-5 wt% which

can be achieved by concentration of caustic species. The driving force for crack propagation

is a high tensile stress applied externally or residing in the steel as a result of

welding/fabrication. If a stress relieving heat treatment is not performed on the steel after

welding, the residual stresses remaining in the weld can be of the order of the tensile

strength of the steel. These residual stresses remain in the weld and the adjacent base

metal unless they are relaxed by a stress relief treatment or by cracking (2).

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Figure 10.1 Caustic cracking susceptibility diagram (2)

Although the failure at the feedwater nozzle was not received for examination, both failures

had occurred in a similar area, towards the bottom of the steam drum (at approximately the

5 o'clock position). The confinement of the failures to this position is more likely to be due

to the corrosion element of the failure, rather than mechanical. Concentration of caustic

species can occur at a waterline and generally, the waterline area is always most sensitive to

corrosion. It is possible that during shut down, the 5 o’clock position corresponded to the

stagnant water line which had then allowed caustic species to concentrate, though this

would have had to have occurred at elevated temperature (above at least 85°C). During

normal service, the water line would be volatile and therefore concentration of caustic

species would be less likely.

Although the corrosive species related to the stress corrosion mechanism are not certain,

the loading mechanism is probably the high residual stress in the joint. If the stress element

to the failure was eliminated by applying an appropriate post weld heat treatment, it is

highly likely that cracking would have been prevented.

11. Conclusions

11.1 Failure of the steam drum had occurred due to the presence of a through-wall

longitudinal crack across one of the circumferential welds. An additional crack was

also evident in the weld of the feedwater nozzle. Both failures had occurred at the

south end of the steam drum, at approximately the 5 o’clock position.

11.2 For the cracking examined, failure had initiated from the inner surface of the weld.

The fracture surface, on a macro scale, was brittle in nature. On a micro scale, the

majority of the fracture surface exhibited micro-void coalescence which is indicative

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of a ductile fracture mechanism (i.e. a tensile overload). The brittle fracture

observed, on a macro scale, is due to constraint created by the thick section.

11.3 The corrosion product associated with the fracture was found to consist

predominantly of iron and oxygen. In some areas, small amounts of chlorine,

potassium, sulphur, calcium, sodium, phosphorus and magnesium were detected. In

one area, a significant amount of sodium was detected. Copper and zinc were also

evident in many of the areas analysed.

11.4 The cracking was transgranular and slightly branched in nature. The crack length was

considered to be very short relative to the crack propagation through the entire wall

thickness (32 mm). Although some crack propagation was evident in the parent

material, this was confined to the weld area.

11.5 The confinement of the crack to the weld area suggests that residual stress in the

weld was a major driving force for crack propagation and explains why the crack had

not extended significantly outside of the weld where the residual stresses would be

reduced.

11.6 The weld had been produced with a double V geometry, as expected. However, a

repair weld was also evident at the inner surface which overlapped the original weld.

The presence of this repair weld could be associated with an increase in residual

stress in the joint.

11.7 All three welds and the heat affected zones close to the welds exhibited very coarse

grain structure which indicates high temperature input during welding.

11.8 The hardness levels in the weld and the HAZ close to the fusion line were slightly

higher than would normally expect, though these were not considered excessive.

The impact properties were also considered relatively high at ambient temperature

and 0°C.

11.9 The cause of failure is thought to be attributable to a stress corrosion mechanism.

The presence of some sodium, magnesium and potassium within the cracking may

indicate caustic cracking, though this not certain.

11.10 The confinement of the failures to the 5 o’clock position of the steam drum is more

likely to be due to the corrosion element of the failure, rather than mechanical.

Concentration of caustic species can occur at a waterline and generally, the

waterline area is always most sensitive to corrosion. It is possible that during shut

down, the 5 o’clock position corresponded to the stagnant water line which had then

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allowed caustic species to concentrate, though this would have had to have occurred

at elevated temperature (above at least 85°C).

11.11 Although the corrosive species related to the stress corrosion mechanism are not

certain, the loading mechanism is probably the high residual stress in the joint. If the

stress element to the failure was eliminated by applying an appropriate post weld

heat treatment, it is highly likely that cracking would have been prevented.

12. References

1) Fatigue and fracture: understanding the basics. F.C. Campbell. Chapter 3: Ductile

and brittle fractures. Page 89.

2) Stress Corrosion Cracking: Theory and Practice. V.S. Raja. SCC in low and medium

strength carbon steels. Page 170.

John Brear – Plant Integrity Cyfyngedig © 18-3-2016

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Section 2 - Integrity assessment of the steam drum on the incineration and

heat recovery plant at a Sewage Works

1. Introduction

The waste incinerator at the subject sewage works is equipped with a heat recovery system

for steam generation.

In 2014 a leak occurred on the steam drum, at the 5 o’clock position on the circumferential

weld between the cylindrical shell and the dished head at the north end. It was repaired

without detailed investigation. A further water leak was detected in December 2015. On

de-lagging a through-crack was found, also at the 5 o’clock position but on the

circumferential weld at the south end; further cracks, also through, were observed on the

feed-water sleeves. A portion of material containing the circumferential weld crack was

removed and sent to R-Tech Consultants Ltd for investigation. As work has progressed,

several further cracks have been found at various nozzle positions in the lower part of the

drum.

2. Drum design

The general arrangement drawing specifies the design, manufacture and NDT to be to BS

1113: 1992. Design conditions are given as:

Temperature, °C 265

Pressure, bar g 50

Hydrostatic test pressure, bar g 75*

* The manufacturer’s plate gives the hydrostatic test pressure as 82.5 bar g.

The general configuration of the drum is typical of designs for the service and operating

conditions (see Figure M.1.1). It is noted that the shell and heads are designed to 32mm

wall thickness, so as to avoid the requirement for post-weld stress relief heat treatment.

3. The drum material

The general arrangement drawing specifies the material for the drum plate and semi-

ellipsoidal dished ends as BS1501 Pt 1: 223-490B. Although BS 1501 was officially

withdrawn on 15th February 1993, it remains a steel specification that is regularly

requested. Thus the requirement for this material on the drum drawing dated 29-9-95 is

neither unusual nor problematic.


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BS1501 223-490 is a specification for fully killed, niobium treated, carbon manganese steel

plates, up to 150 mm thick, with a minimum tensile strength of 490 MPa. It may be supplied

as Type A, with specified minimum tensile properties at room temperature, or as Type B,

with additionally specified elevated temperature properties. In addition, low temperature

impact properties can be specified at either -30°C (LT30) or -50°C (LT50). For elevated

temperature service this material has largely been replaced by BS EN 10028 P355GH,

though that has a lower strength level of 355 MPa.

The drum drawing specifies Type B, though the required elevated temperature properties

are not stated, and makes no low temperature impact requirement.

3.1 Chemical requirements

Chemical composition (ladle analysis, %) of steel 223 Grade 490, is specified as:

C Si Mn P S Nb Cr Cu Mo Ni

min - 0.10 0.90 - - 0.01 - - - -

max 0.20 0.50 1.60 0.030 0.030 0.06 0.25* 0.30* 0.10* 0.75

* Cr + Cu + Mo = 0.50% max

3.2 Mechanical properties

3.2.1 Tensile requirements

Plate Thickness Tensile, Rm Yield, Re Elongation, A

mm MPa Min, MPa Min, %

3 - 16 490 - 610 355 21

16 - 40 490 - 610 345 21

40 - 63 490 - 610 340 20

63 - 100 490 - 610 * 20

100 - 150 490 - 610 * 20

* The value of yield strength for plates over 63 mm thick shall be the values specified for

plates of thickness between 40 mm and 63 mm reduced by 1% for each 5 mm or part

thereof increase in thickness over 63 mm.


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3.2.2 Impact requirements

No low temperature impact requirement is specified for the drum but, for information, the

LT30 specification gives minimum impact test values of:

Temperature, °C RT 0 -15 -30

Impact energy, J 61 55 41 27

These values apply to plates up to and including 80 mm thick and are based on the average

of three tests.

3.3 Welding

The drum comprises two rolled-plate cylindrical sections and two semi-ellipsoidal heads.

The two longitudinal and three circumferential structural welds are all indicated as double-V

on the drawing, but the weld procedure is not available. The north end circumferential

weld is noted as being the closure weld. The longitudinal welds are at the ± 45° positions,

measured from the top.

Nozzles are either set-through or set-on. The former appear to be welded with internal and

external fillets, rather than a fully penetrating weld.

As the vessel thickness is limited to 32mm, no post-weld heat treatment (PWHT) was

applied.

4. Investigation of the cracking

4.1 Incidence

The operators have supplied a list of cracks found to date; this is given in Table 1. A column

has been added to show the contact phase at each location.

It is immediately apparent that locations in contact with steam are not affected; neither are

set on nozzles. However, all but one (N3) of the set through nozzles in contact with water

have crack indications and the cracks on the circumferential welds are also in water-touched

regions (5 o’clock, as viewed from the north).

Photographs supplied show all cracks to be of similar appearance, largely confined to the

weldments and oriented in a manner consistent with residual stress fields (see Figures M.1.2

and M.1.3).

These observations point strongly to weld design and procedure and to water chemistry as

causative factors.


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4.2 The transverse crack in the south end circumferential weld

Detailed examination of the sample of material containing the crack has been performed

and reported by R-Tech. This has included metallography and fractography, chemical

analysis, tensile testing and impact testing.

4.2.1 Materials chemistry

Chemical analysis (Table M.8.1) shows the shell and head parent materials to conform to

the standard. The three weld metals are a little high in silicon and manganese, but might

well conform to the specified (but presently unknown) consumable. Chromium,

molybdenum, nickel and copper are all low.

Residual elements known to contribute to weldment cracking - phosphorus, sulphur,

arsenic, antimony and tin - are all low and accord with good steel-making practice at the

date of manufacture.

4.2.2 Welding

The weld is of double-V geometry, as specified. However, a single-bead repair run is evident

on the inner surface, head side (Figure M.2.5). This shows a noticeably different electrode

chemistry to the main weld. Whether it was applied to correct a poor weld geometry, or to

repair a post-weld crack, is not known.

Weld beads are large, indicating at least an 8mm electrode (Figure M.2.5). Large grains are

seen in the weld metal and heat affected zone (HAZ). There is little weld bead overlap,

resulting in long, continuous runs of coarse-grained material in the HAZ (Figure M.4.5). All

these features indicate a high heat input; this is poor practice as it results in high residual

stresses and microstructures susceptible to cracking. An improved weld procedure should

be specified for any replacement drum.

Parent materials show the expected ferrite-pearlite microstructure.

4.2.3 Mechanical properties

The elevated temperature mechanical properties specification implied by the ‘B’ suffix to

the materials standard is not known. Little change is to be expected over the temperature

range ambient to 265°C for materials of this general type, however.

Tensile testing of the shell parent material at ambient temperature (Table M.9.1) shows the

proof stress, UTS and ductility to conform to the standard.

Hardness testing (Table M.6.1) shows values in accord with the microstructures observed.

Those for parent material are normal for the strength grade; those for weld and HAZ are on

the high side, reflecting the high heat input and absence of PWHT. Following data given in


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various standards, it is possible to estimate tensile strength from hardness measurements.

Taking R-Tech’s mean data, the following results are obtained:

Location Hardness Estimated UTS, MPa Measured

HV10 HV30 HV10 HV30 UTS, MPa

Parent, head 163.3 165.7 517.6 525.4

Repair weld HAZ, head 200.3

640.9

Repair weld HAZ, head, near fusion line 245.7 792.0

Repair weld 244.0 786.5

Inner weld HAZ, head 217.7 698.7

Inner weld 234.0 753.1

Inner weld HAZ, shell 206.3 660.9

Parent, shell 171.7 163.0 545.4 516.5 531, 544

Outer weld HAZ, shell 215.0

689.8

Outer Weld 229.3 737.6

Outer weld HAZ, head, near fusion line 226.3 727.6

Outer weld HAZ, head 201.0 643.1

These indicate parent material strength values that conform to the standard but weld and

HAZ strengths that are generally high, reflecting the high heat input weld procedure used.

For guidance, the specified strength range for parent material, 490 – 610 MPa, corresponds

to a hardness range of 155 – 191 HV. For the shell parent material, the estimated and

measured UTS values are in a good agreement.

Impact properties of both parent materials and both main weld metals (Tables M.7.1,

M.7.2) clearly exceed the LT30 requirement, even though that is not part of the given

specification. Weld metal results show more scatter than those from parent material, as

expected. Fracture appearances are generally ductile (Figures M.7.4 – M.7.12). The test on

head parent material, nearer the outer surface, shows a somewhat reduced impact energy

and a more brittle fracture surface at 0°C, indicating that this temperature is probably

within the transition region.

4.2.4 Fracture mode and morphology

The crack is clearly confined to the width of the weld, propagating into the parent material

only within the confines of the double-V geometry. The fracture surface shows initiation in

the weld metal, at the inner surface (Figure M.2.9). In section it is seen to be transgranular

and shows some branching (Figures M.4.2, M.4.9). Whilst macroscopically brittle,

microscopically the fracture demonstrates the features of ductile tensile overload. This

combination indicates a degree of constraint on crack opening, in part due to the section

thickness but mainly a reflection of the local stress profile.


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On opening, the fracture surface was found to be covered in corrosion product

demonstrating a stress-corrosion mechanism. The crack characteristics point towards a

caustic corrosion mechanism, but not conclusively.

5. Integrity analysis

The general distribution of the cracking and the detail gained from examination of the

transverse crack in the south end circumferential weld point to a stress corrosion cracking

mechanism, driven by the action of the welding residual stresses under a poor water

chemistry environment. It is important to note that, after towards 20 years of operation,

welds have progressed from NDE clear to through cracking in around a year. As the welding

stresses have not changed during operation, this must reflect a recent and deleterious

change in water chemistry control – either absolutely or resulting from local concentration

due to running the drum at low water level.

5.1 Drum loadings

5.1.1 Pressure stress

With a design diameter of 1676 mm and wall thickness of 32 mm, the mean diameter hoop

stress at the design pressure of 50 bar g is 128.4 MPa; at the hydrostatic test pressures of 75

and 82.5 bar g it is 192.7 and 211. MPa respectively. All these figures are well below yield.

5.1.2 Effect of blocked support movement

The north saddle is indicated as fixed, the south as sliding; the saddle separation is 2,550

mm. It has been reported that the sliding saddle has, of late, not moved freely.

For a temperature range of 240 degC (ambient of 16°C to design of 256°C) and an expansion

coefficient of 13.3 x 10-6 mm/mm/degC, the expansion over the separation distance is 8.1

mm; the sliding support should account for that, plus any fit up errors. If this movement

were totally blocked, then a modulus of 195 GPa would result in an elastic stress of 622.4

MPa – comparable to the maximum UTS.

Whilst it is clear that such an extreme circumstance has not occurred, this calculation

demonstrates the need for care to ensure free support movement. Numerous boiler

failures have resulted from blocked supports, notably that on board the SS Norway in 2003.

A more likely scenario is that the support responds to the rising stress by moving in a series

of steps, rather than smoothly. This will transmit shock loads to the body of the drum, and

quite possibly further through the system, with a potential for damage or as a trigger to

crack initiation.


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5.1.3 Welding residual stresses

As noted previously, the crack orientations are consistent with expected weld residual

stresses. As a weld is deposited the temperature difference between the bead and the

substrate leads to thermal stresses – tensile in the weld metal and compressive in the

parent material, in internal equilibrium. These stresses can reach yield level, even exceed it

in situations of high constraint.

5.2 Fracture mechanics assessment

5.2.1 Estimation of fracture toughness

Figure 1 shows the measured impact data together with the LT30 specification. The head

parent material data obtained on sub-size specimens are shown as-measured and corrected

to 10 mm on a simple area basis. Corrected, they coincide with the shell parent material

data.

Fitting an inverse tangent function to the LT30 points results in a credible transition curve.

Likewise, a similar fit to the mean parent material data results in a curve with a transition

temperature of -8°C and an upper shelf energy of 200 J – both figures are credible when

compared with data on similar steels. Scaling this curve down to the mean weld impact

energy at 20°C results in a curve with an upper shelf energy of 125 J.

Using a Sailors and Corten (Reference 1) type correlation of the form:

Kic = A CvB

allows construction of fracture toughness (Kic) transition curves from the Charpy impact

energies (Cv), as shown in Figure 2. The estimated upper shelf toughnesses are 235 and 175

MPa√m for parent and weld metal respectively. In comparison, the LT30 impact

specification implies an upper shelf toughness of 120 MPa√m; this is consistent with the

conservative value of 100 MPa√m generally used for assessment purposes where no

measured values are available.

It is considered that this approach generates reasonable estimates of toughness for the

parent and weld metals in question.

5.2.2 Estimation of flow stress

For constrained geometries, the flow stress may be taken as the mean of the yield strength

and the ultimate tensile strength. For most regions of the weldment the UTS has been

estimated from the hardness data, as described above, and the yield to ultimate strength

ratio has been assumed the same as defined by the materials standard. The tensile data

obtained on the shell parent material confirm that this approach is reasonable.


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5.2.3 Defect assessment

The transition from stable crack growth to fast fracture is assessed against two criteria: the

tendency to brittle fracture and the tendency to plastic collapse. Susceptibility to brittle

fracture is determined by the ratio of the stress intensity at the crack tip – which depends

on the stress, the crack size and the component geometry – to the fracture toughness.

Susceptibility to plastic collapse is determined by the ratio of the stress in the ligament

ahead of the crack to the flow stress.

The interaction between these two processes is represented by the failure assessment

diagram, shown in Figure 3 (References 2, 3). The vertical axis represents the brittle

fracture ratio and the horizontal axis the collapse ratio. Two interaction curves are shown,

one for a purely elastic-plastic material and one for a material showing strain hardening.

The latter is considered more appropriate for this steel and constrained crack geometry. A

defect whose parameters lie below the assessment line is expected to be static, or to grow

in a stable manner. A defect whose parameters lie above the line is predicted to be

unstable.

It is not possible to determine from the fractography what the initial defect geometry was,

nor is the growth mechanism sufficiently well-defined to estimate growth rates. However,

the as-found geometry is clear. For a weld defect acted on by a flow-stress level residual

stress, the highest value possible, the position on the horizontal axis will be, by definition,

unity. The vertical component will be governed by this stress, the crack length and the weld

metal toughness. Taking the crack size observed and the properties estimated as described

in the previous section results in the plotted point – almost exactly on the assessment line.

This confirms the realism of the assumptions made.

In practical terms it shows that a weld metal defect, under the maximum residual stress, will

grow to reach the stability limit and that its further rapid growth will only be restricted by

the extent of the tensile residual stress field. Once it reaches the compressive stress zone,

growth will halt, as is confirmed by the rounded crack tip seen in Figure M.4.7. The ratio of

unity between the flow-stress and the residual stress that of necessity pertains in these

circumstances will lead to a ductile tearing mechanism at the micro scale, but the constraint

to crack opening consequent on the transition from tensile to compressive stresses at the

weld metal boundary will lead to apparently brittle behaviour at the macro scale. At these

high stress levels, it is likely that, once initiated, stable crack growth will be fairly rapid and

the observed one-year timescale is not unreasonable.

In contrast, a through crack of similar size in the parent material, under the pressure hoop

stress, will be well within the assessment line, as shown by the points plotted for parent

material under the operating and hydrostatic test stresses. Thus there is little, if any,

likelihood of a transverse weld metal crack on a circumferential weld propagating beyond


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the weldment.

The greatest risk is associated with the cracking seen around certain nozzles, as this could

lead to nozzle severance and a major water/steam emission.

6. Exclusion zone calculation

Should a nozzle weld fail, causing partial or even complete nozzle severance, water would

be ejected and this would flash off as steam at atmospheric pressure. As the water level in

the drum fell, complete boil-off of the water circuit contents would be possible, as the leak

could not be isolated. Given the drum geometry and nozzle configuration, a steam jet could

occur in any direction.

6.1 Type of release

As the failure situation addressed is that of a steam drum, it is necessary to consider

possible water, steam and two-phase releases. Whilst the potential ballistic trajectory of a

discharged water slug is considerable, it can readily be shown that instantaneous

vaporisation would occur at the rupture point. Similar considerations apply to a two-phase

release. It is therefore only necessary to examine a pure steam release in detail, as water

and mixed discharges would immediately converge to that situation.

6.2 Characteristics of a steam release

At any realistic boiler operating pressures it can be shown that a steam release would be

choked, i.e. sonic in nature (that is, the local velocity of the escaping steam has reached the

speed of sound at the local conditions). Calculations of sonic release rates have been

performed in a manner consistent with the methods given in recognised hazard assessment

procedures (References 4, 5). However, it is noted that the formulae adopted by these

procedures assume ideal gas behaviour; this is not appropriate for steam at these

conditions. In particular, such formulae implicitly underestimate the density of the steam,

and thus the mass discharge rate. Accordingly, more accurate models have been used here,

following standard texts on fluid dynamics (References 6, 7). In all cases, a full severance of

the nozzle has been assumed.

6.3 Mass and temperature profile of the release

Results are provided in Tables 2 and 3. Table 2 gives the release characteristics – mass

release rate, velocity – for the range of nozzle sizes present on the drum. For nozzle

diameters of 88.9 mm and above, the mass released in 3 minutes exceeds the 10,000 lbs

criterion of API 580 (Reference 4) and the releases should therefore be classed as

‘instantaneous’. Assessment is therefore based on the immediate effects of the release,

rather than on subsequent dispersion of the released material. It should be noted,

however, that the total fluid inventory in the heat recovery unit is large and therefore a


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considerable amount of material could be released after this three-minute period.

Table 2 also gives the corresponding heat released from the steam to the air and the

quantity of air required to cool the resultant mix to 50°C. Standard relationships have been

used to determine the distance over which this occurs (References 4, 5). At the design

pressure of 50 bar g, the exclusion zone size is calculated to be 25.2 metres for a jet release

from the largest nozzle.

This calculation, however, assumes that there is no obstacle or protection between the

ruptured nozzle and any personnel; the steam/air mixture thus forms a conical jet. The

cladding on the drum, while not gas-tight, is highly likely to muffle the release such that the

jet will be dispersed into a discharge approximating more to a hemispherical cloud. By

adjusting the shape factor in the distance equation to a limiting value of unity, more realistic

safe distances for this situation can be calculated. At the design pressure of 50 bar g, the

exclusion zone size is calculated to be 8.1 metres for a hemispherical cloud release from the

largest nozzle. Figure 4 shows the variation in safe distance to 50°C as a function of nozzle

size, for both jet and cloud releases.

Table 3 and Figure 5 give the variation in temperature of the steam/air mixture as a function

of distance from the release point, for a design pressure of 50 bar g and for both conical jet

and hemispherical cloud releases from the largest nozzle. Significant injury, even fatality,

might be expected on exposure to temperatures above 60°C. Brief exposure to a

temperature of 50°C may be tolerable (8.1 metres for a hemispherical cloud release) but for

anyone involved in work, and therefore possibly restricted from rapid escape, a limit of 40°C

(11.6 metres for a hemispherical cloud release) is considered more appropriate.

Exclusion zones should be planned recognizing that situations might arise where personnel

are required briefly to move towards the rupture point in effecting their escape.

7. Options

There are three potential options for the drum, immediate retirement, repair and operation

for around six months as a replacement is procured, repair and operation for the three

years required of the unit in general. There are also implications for the design and

operation of a new drum.

It has to be recognised that the present drum is badly damaged, as a consequence of stress

corrosion cracking caused by high weld residual stresses and poor water chemistry control.

If the drum is to be retained, then actions must be taken to alleviate the situation.

7.1 Drum operation

Firstly, for this drum, for its replacement, and potentially for other units within the

company, attention must be given to the operational procedures necessary to control water


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level and water chemistry. On units of this nature, with an open steam/water cycle, such

control is a fine art and it is strongly recommended that expert advice be sought.

7.2 Drum repair

If it is to be retained, the present drum must be repaired. This is physically achievable, but

may not be economic.

Prior to any repair work, free movement of the sliding saddle must be ensured.

7.2.1 Repair method

All damaged areas must be identified by NDE, then cut out, ensuring that all cracking and all

existing HAZ material is removed. As PWHT will be difficult, except on the circumferential

welds, a repair weld procedure that minimises residual stresses and aims to generate the

most favourable microstructures should be employed. Given the drum material and design,

a temper-bead technique is recommended. Here, an initial weld deposit is made using fine-

gauge electrodes, to minimise heat input into the parent material; subsequent layers use

progressively larger electrodes such that the heat input from each successive bead tempers

and stress relieves the previous layer. The usual NDE procedures should be followed.

There is good confidence that sound repairs could be achieved by this route, with six

months of operation possible, potentially longer with regular inspection, monitoring and

close operational control.

Future service with a repaired drum should be subject to application of an exclusion zone.

Provided the lagging and casing are secure, and subject to consideration of the need for

additional shielding, a zone of 12 metres should be adequate.

7.2.2 Hydrostatic testing

Following repair, there will be a requirement for hydrostatic testing. The drawing and

maker’s plate differ as to the original test pressure requirement.

In the circumstances, it is considered that a test pressure of design plus ten per cent would

be appropriate.

Given the measured impact properties, testing at an ambient temperature no less than 10°C

should be satisfactory.

7.3 Drum replacement

Should drum replacement be actioned, then essentially a like-for-like approach should be

adopted. An improved weld procedure should be imposed on the specification and, even at

this wall thickness, PWHT should be considered.


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8. Conclusions and recommendations

It is concluded that:

1. The cracking experienced is due to a stress corrosion mechanism, quite possibly

caustic in nature, driven by the welding residual stresses and an unfavourable water

chemistry.

2. Drum material chemistry, mechanical properties and microstructure are acceptable,

but the welding procedures used in construction were not optimised and have given

rise to high weld residual stresses and poor microstructures.

3. Fracture mechanics analysis confirms that welding residual stresses provided the

driving force for cracking and gives predictions on defect stability that match

observation.

4. Whilst transverse cracks on the structural welds are constrained from propagation

and opening, a significant risk has been identified associated with cracking of nozzle

welds leading to nozzle severance.

5. Calculations show that an exclusion zone of 12 metres should be adequate if the

drum is returned to service after repair.

It is recommended that:

1. Advice be obtained on the procedures necessary to ensure correct operational

control of water level and water chemistry – for this drum if repaired, for its

replacement, and for other steam-raising plant within the company.

2. Should a repair option be followed, then a suitable weld procedure should be

employed. A temper-bead technique is recommended.

3. Following repair, hydrostatic testing at design pressure plus ten percent and at an

ambient temperature of no less than 10°C is considered appropriate.

4. Should the present drum be returned to service, then a carefully considered 12

metre exclusion zone is recommended.

5. Should a replacement option be followed, then a like-for-like approach should be

adopted, with an improved weld procedure imposed and PWHT considered.


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9. References

1. Sailors, RH and Corten, HT

ASTM STP514, pp 174-191, 1973

2. American Petroleum Institute

‘Fitness for service’ API Recommended Practice 579

First Edition, 2000

3. British Standard BS7910

Guide on Methods for Assessing the Acceptability of Flaws in Structures

4. American Petroleum Institute

Base Resource Document on Risk-Based-Inspection

Supplement to API Recommended Practice RP580/581

5. Technica Ltd.

Techniques for Assessing Industrial Hazards

World Bank Technical Paper WTP 55

6. Walshaw, AC and Jobson, DA

Mechanics of Fluids, Third edition 1980

7. Rogers, GFC and Mayhew, YR

Engineering Thermodynamics: Work and Heat Transfer, fourth edition 1992


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Table 1: Defect indications

Location Nozzle size Design Phase Condition (as revealed by MPI)

2014

North circ weld, 5 o'clock Double-V Water Transverse crack - repaired

2015/16

North circ weld, 5 o'clock Repair Water No indications

South circ weld, 5 o'clock Double-V Water Transverse crack - repaired

N1 73 o/d Sch 80 Set through Water Minor indication toe weld on pipe

N2 168.3 o/d 21.95 wall Set through Water Crack in toe of weld on pipe

N3 168.3 o/d 21.95 wall Set through Water No indications

N4 48.3 o/d Sch 80 Set on Steam No indications



N7 60.3 o/d Sch 80 Set on Mixed No indications

N8 48.3 o/d Sch 80 Set through Water Crack in toe of weld on shell



N11 88.9 o/d Sch 160 Set through Water Crack in toe of weld on shell - repaired

N12 33.4 o/d Sch 80 Set through Water Crack in shell - repaired

N13 88.9 o/d Sch 160 Set through Water Crack in toe of weld on shell - repaired

N14 26.7 o/d Sch 80 Set on Steam No Indications



N17 114.3 o/d Sch 160 Set on Steam Inaccessible


N19 60.3 o/d 5.5 wall Set on Steam No Indications

N20 PSV outlet support Set on Steam No Indications

N21 141.3 o/d 15.9 wall Set through Water Crack in weld and shell

N22 48.3 o/d Sch 80 Set on Water No indications

N23 168.3 o/d 18.3 Set through Water Crack in weld and shell

N24 114.3 o/d Sch 160 Set through Water Crack in weld and shell

N25 48.3 o/d Sch 80 Set through Water Crack in weld, shell around bracket

N26 168.3 o/d Sch 160 Set through Water Crack indication weld

Information supplied by the operator


Table 2: Release characteristics as a function of nozzle size

Nozzle

33.4 o/d

Sch 80

48.3 o/d

Sch 80

73 o/d

Sch 80

88.9 o/d

Sch 160

114.3 o/d

Sch 160

141.3 o/d

15.9 wall

168.3 o/d

21.95

wall

168.3 o/d

18.3 wall

168.3 o/d

Sch 160

Pipe OD mm 33.40 48.30 73.00 88.90 114.30 141.30 168.30 168.30 168.30

Wall mm 4.55 5.08 7.01 11.13 8.56 15.90 21.95 18.30 18.26

Pipe ID mm 24.30 38.14 58.98 66.64 97.18 109.50 124.40 131.70 131.78

Area m2 4.64E-04 1.14E-03 2.73E-03 3.49E-03 7.42E-03 9.42E-03 1.22E-02 1.36E-02 1.36E-02

Pressure bar g 50.00 50.00 50.00 50.00 50.00 50.00 50.00 50.00 50.00

Saturation temperature °C 265.27 265.27 265.27 265.27 265.27 265.27 265.27 265.27 265.27

Release rate Kg/s 3.59 8.84 21.14 26.98 57.38 72.85 94.03 105.39 105.52

Release velocity m/s 298.94 298.94 298.94 298.94 298.94 298.94 298.94 298.94 298.94

3-minute release Kg 645.82 1,590.98 3,804.62 4,857.05 10,328.95 13,113.86 16,925.56 18,970.28 18,993.33

Release temperature °C 188.22 188.22 188.22 188.22 188.22 188.22 188.22 188.22 188.22

Total heat release MW 7.19 17.72 42.37 54.09 115.03 146.04 188.49 211.26 211.52

Air required Kg/s 216.01 532.13 1,272.53 1,624.54 3,454.72 4,386.19 5,661.09 6,344.99 6,352.70

Conical jet distance* m 4.65 7.29 11.28 12.75 18.59 20.94 23.79 25.19 25.20

Hemispherical cloud

distance* m

1.49 2.33 3.61 4.08 5.95 6.70 7.61 8.06 8.07

*To a safe temperature of 50°C


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Table 3. Temperature of the steam/air mixture as a function of distance from the release (168.3 OD Schedule 160 nozzle)

Temperature, °C 100 95 90 85 80 75 70 65 60 55 50 45 40 35 30 25 20 17

Conical jet, m 9.4 10.1 10.9 11.7 12.8 14.0 15.4 17.1 19.2 21.8 25.2 29.8 36.3 46.1 63.1 98.9 224.2 900.6

Hemispherical cloud, m 3.0 3.2 3.5 3.8 4.1 4.5 4.9 5.5 6.1 7.0 8.1 9.5 11.6 14.8 20.2 31.6 71.7 288.2


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Figure 1: Impact properties and fitted transition curves


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Figure 2: Fracture toughness estimated from the impact transition curves


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Figure 3: Failure assessment diagram for a through crack


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Figure 4: Distance to a safe temperature of 50°C as a function of nozzle area


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Figure 5: Temperature of the steam/air mixture as a function of distance from the release – 168.3mm OD schedule 160 nozzle

Failure Analysis & Integrity Assessment of the Steam Drum ...

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