failure analysis & integrity assessment of the steam drum
TRANSCRIPT
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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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Figure 5.3 Corrosion product associated with cracking, showing areas analysed
Figure 5.4 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
Sample 12 (inner surface) Inner weld 10 x 10 x 55 2mm V notch 150
Sample 12 (outer surface) Outer weld 10 x 10 x 55 2mm V notch 108
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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 3 (inner surface) Inner weld 10 x 10 x 55 2mm V notch 143
Sample 3 (outer surface) Outer weld 10 x 10 x 55 2mm V notch 148
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.
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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
N5 48.3 o/d Sch 80 Set on Steam No indications
N6 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
N9 48.3 o/d Sch 80 Set through Water Crack in toe of weld on shell
N10 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
N15 26.7 o/d Sch 80 Set on Steam No Indications
N16 168.3 o/d Sch 160 Set on Steam No Indications
N17 114.3 o/d Sch 160 Set on Steam Inaccessible
N18 168.3 o/d Sch 160 Set on Steam No Indications
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
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
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
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
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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
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
Page 61 of 65
Figure 1: Impact properties and fitted transition curves
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
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Figure 2: Fracture toughness estimated from the impact transition curves
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
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Figure 3: Failure assessment diagram for a through crack
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
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Figure 4: Distance to a safe temperature of 50°C as a function of nozzle area
John Brear – Plant Integrity Cyfyngedig © 18-3-2016
Page 65 of 65
Figure 5: Temperature of the steam/air mixture as a function of distance from the release – 168.3mm OD schedule 160 nozzle