hydrogen cracking of legs - health and safety executive · hydrogen cracking of legs and spudcans...

OTH 91 351

HYDROGEN CRACKING OF LEGSAND SPUDCANS ON JACK-UPDRILLING RIGS - A Summary of

Results of an Investigation

Authors

K AbernethyBritish Steel Technical Research Organisation

C M FowlerCAPCIS, University of ManchesterInstitute of Science and Technology

R JacobGlobal Corrosion Consultants Ltd

V S DaveyOffshore Safety Division

Health and Safety Executive

With the assistance of

Techword Services 153-155 London Road

Hemel HempsteadHerts HP3 9SQ

HSE BOOKS

Health and Safety Executive - Offshore Technology Report

© Crown copyright 1993Applications for reproduction should be made to HMSO

First published 1993ISBN 0 7176 0614 7

This report is published by the Health and Safety Executive aspart of a series of reports of work which has been supported byfunds formerly provided by the Department of Energy and latelyby the Executive. Neither the Executive, the Department nor thecontractors concerned assume any liability for the reports nor dothey necessarily reflect the views or policy of the Executive orthe Department.

Results, including detailed evaluation and, where relevant,recommendations stemming from their research projects arepublished in the OTH series of reports.

Background information and data arising from these researchprojects are published in the OTI series of reports.

FOREWORD

This study was initiated by the Department of Energy when the responsibility foroffshore safety was with its Petroleum Engineering Division. In April 1991 thisresponsibility was transferred to the Offshore Safety Division of the Health andSafety Executive, which is now publishing this report.

The Fourth Edition Guidance referred to in this report was issued by the Departmentof Energy in 1990 and is now the responsibility of HSE. Guidance on mattersdiscussed in this report will be included in Amendment No 2 to the Fourth Edition tobe published in 1993.

4

CONTENTS

85Appendix: Comments on interim guidance document

83References7.

7980

General conclusions6.1 Draft interim guidance

6.

737373707879

Cathodic protection at limited potentials5.1 Introduction5.2 The need for limited CP potentials5.3 Methods of limiting CP potentials5.4 Possible risks of limiting potentials5.5 Conclusions

5.

383839394146

The hydrogen susceptibility of Type A and Type B steels4.1 Introduction4.2 Test technique4.3 Test programme4.4 Results and discussion4.5 Conclusions and recommendations

4.

2121212526

Metallurgical characterisation of Type B steel3.1 Introduction3.2 Examinations3.3 Discussion3.4 Conclusions

3.

4459

10

Metallurgical characterisation of Type A steel2.1 Introduction2.2 Examinations2.3 Discussion2.4 Conclusions

2.

1133

Introduction1.1 History of the problem1.2 The research programme1.3 Layout of this report

1.

viSummary

6

SUMMARY

Hydrogen-assisted cracking is known to be a problem with some high strength steelsused in the fabrication of jack-up offshore drilling rigs, particularly in areas adjacentto welds. Following the discovery of such cracks in a number of rigs during routinesurveys in the late 1980s, the Department of Energy commissioned a series ofresearch investigations into different aspects of the problem. This report contains theresults of the investigations and makes recommendations to minimise the problem inthe future. The investigation covered:

� detailed characterisation of samples of two high strength steels known tohave suffered hydrogen-assisted cracking

� experiments to establish the limiting level of cathodic protection potentialbeyond which these steels are susceptible to hydrogen cracking (slow strainrate testing was used for this purpose)

� an assessment of different ways of ensuring that CP potentials in practice donot exceed the limiting level established by the experimental results.

Results of the characterisation tests showed the materials to be of variable strengthsand hardnesses. The slow strain rate tests established that hydrogen damage to thesteels was small provided CP potentials did not become more negative than -825 mV(with respect to a saturated calomel electrode). The only practical way of ensuringthis limit is not exceeded on site is by use of Schottky barrier diodes to separate theCP anodes from the structure being protected.

This report also includes a copy of interim guidance drafted by the Department ofEnergy on the subject of 'Hydrogen-assisted cracking in high strength steelsimmersed in seawater'. The document was discussed at the end of 1990 at a meetingattended by representatives of interested parties. Important points from thisdiscussion are included in this report.

vi

1. INTRODUCTION

1.1 HISTORY OF THE PROBLEM

It has been known for some time that hydrogen may attack high strength steels,particularly at areas adjacent to welds where increased hardness and residual stressmay exist. Some information on this phenomenon is contained in OffshoreInstallations: Guidance on Design, Construction and Certification(1) and in othercodes of practice associated with the construction and operation of equipment usinghigh strength steels.

Salient points in the history of the problem as it has affected jack-up drilling rigs are:

� The problem was first noticed in February 1988 during surveys of two rigsoperating on the UK continental shelf. Centrifugally cast steel had been usedfor their main leg chords. Cracking was found at the intersection betweenthe leg chords and the spud can top plate and at some internal connectionsbetween bulkhead members and the abutting leg chords. All cracking hadoccurred in the heat affected zone of the high strength material. Boatsamples were taken for metallurgical examination but the full implications ofwhat they showed were not immediately recognised.

� In March 1988 a rig which had been operating in Argentina was found tohave extensive cracking inside the spud cans. It required immediate repairon return to the UK but no metallurgical examination was carried out.

� In June 1988 a further rig was found to have serious cracking within thespud cans. This was associated with hydrogen sulphide (H2S) found to bepresent in the spud can residues. As the damage appeared to be confined tothe spud cans, this led to the assumption that hydrogen-assisted cracking wasoccurring due to the presence of H2S. Further examination by metallurgistsshowed that the cracking was of a form induced by hydrogen and that itcould also occur in the absence of H2S.

� In August 1988 a fifth rig was found to have severe cracking within the spudcans (all the rigs so far affected had been of the same type). The crackingextended to the external welds between the leg chords and the spud can topplates. At this point the Department of Energy asked the owners of allremaining units of this type to have them subjected to dry-dock survey asquickly as possible, but no later than 31 October 1988. In the case of thisfifth unit a dry-dock survey in early October showed that extensive crackingwithin the spud cans was complemented externally by severe cracking atbrace-to-chord connections up to the third horizontal level. This lattercracking had occurred mainly where the paint coating had been removed toallow underwater inspection. The internal spud can cracking, however, hadoccurred where the pain coating had deteriorated. All cracking was in theheat affected zone of welds in the high strength material. Boat samplesconfirmed that in each case hydrogen-induced stress corrosion cracking wasthe only form of damage present. No cracking due to fatigue was found.

� From mid October to early December 1988, three further units weresubjected to dry-dock survey. In each case the installations were found to becracked to some degree but, more seriously, one unit which had beenrepaired in April 1988 showed signs of further cracking within six months.

� The last four rigs surveyed were also subjected to an investigation by theDepartment of Energy with a view to understanding what influences may

1

have caused the cracking. This included measuring cathodic protection (CP)potentials within spud cans and an examination of the spud can contents forthe presence or absence of H2S. (On only one rig had CP potentials beenmeasured in service.) The tests showed that CP potentials were excessive forapplication to high strength steel. The absence of oxygen in the spud cancontents and the intermittent presence of H2S in small quantities could haveexacerbated the conditions already created by the CP. (The H2S wasnormally retained in the sediment on the bottom of spud cans and thereforewas not necessarily available to cause chemical attack.)

� Several other types of rigs coming up for survey were examined for this formof cracking. Two of one type showed hydrogen attack to a greater or lesserextent and a further type showed confirmed hydrogen attack at welds on theexterior of spud cans where the high strength steel chords were exposed.Two other rigs suffered from cracking which was confirmed on one as due tohydrogen attack.

� In the period to the end of February 1989, twelve rigs of five types had beenshown to be affected by spud can and leg cracking. Of the five typesinspected, all showed damage confirmed as being due to hydrogen attack ofthe heat affected zone in the high strength steel adjacent to welds. Somecracking due to faulty fabrication was also found.

� During the 1990 inspection period, one Certifying Authority reported nofurther problems on two rigs that had suffered from the problem initially.However, serious cracking was discovered on the inside of spud cansadjacent to high strength steel chords on a third rig that had been affectedearlier. Some minor cracking had also occurred at similar positions on theoutside of the spud cans.

� Another Certifying Authority reported that they had not had problemsrevealed in 1990 on any rigs other than those of the type that started theconcern in 1988. Four rigs of this type were subjected to full survey and thefindings could be summarised as:� Where CP anodes had been removed from spud cans and the spud

can water was treated with corrosion inhibitor and biocide, minordamage was observed but some of this may have been historic andresidual from previous repairs.

� Where CP systems had not been modified immediately after theprevious survey, recurrent damage had been found although someappeared to be due to the previous repairs causing excessive localresidual stresses.

� Some minor cracking had been observed on leg members outsidespud cans, but was not as serious as that found in 1988.

� In summary, there had been improvements overall in the performanceof the material in these rigs but not all problems had been overcome.Further efforts would be made to improve this situation using theCertifying Authority's own research.

� A third Certifying Authority reported that suspected hydrogen attack on tworigs had not been confirmed; it was a different form of cracking not related inany way to hydrogen or CP systems.

2

1.2 THE RESEARCH PROGRAMME

As soon as it became clear that there was a common problem here affecting a numberof jack-up drilling rigs, the Department of Energy convened a meeting of the ownersof the affected rigs, the relevant Certifying Authorities and the offshore concessionowners currently chartering the rigs. The meeting was held on 11 January 1989 andthe subject was discussed extensively.

Subsequently, the Department of Energy sponsored further research to investigate theproblem and possible solutions. This stage of the work was carried out in three parts:

� A detailed characterisation of two commercially available high strengthsteels used in jack-up chords. Carried out by British Steel TechnicalResearch Organisation. Reported in this document in Chapters 2 and 3.

� An experimental investigation of the susceptibility of these two high strengthsteels to hydrogen embrittlement at different applied CP potentials. Carriedout by Corrosion and Protection Centre Industrial Services at the Universityof Manchester Institute of Science and Technology. Reported in thisdocument in Chapter 4.

� A consultancy opinion on the feasibility of prevent corrosion of the steelswithout exceeding the apparent CP threshold for significant hydrogenembrittlement. Carried out by Global Corrosion Consultants Ltd. Reportedin this document in Chapter 5.

1.3 LAYOUT OF THIS REPORT

After this introduction (Chapter 1), the report gives a full account of the researchprogramme described in Section 1.2 above (Chapters 2-5). Chapter 6 drawsconclusions from the research and contains a draft version of interim guidance on thesubject to be published by the Health and Safety Executive (the successor to theDepartment of Energy on matters concerning the safety of offshore installations). Anappendix summarises a discussion that took place on 4 December 1990 when thedraft interim guidance was discussed by representatives of offshore owners,contractors, Certifying Authorities and the government bodies concerned.

3

2. METALLURGICAL CHARACTERISATION OF TYPE A STEEL

2.1 INTRODUCTION

An examination of material from six tubular sections has been carried out in order tocharacterise the material from which the tubulars were made. The tubulars werestated to be of high strength steel manufactured to the chemical and mechanicalproperty specifications shown in Tables 1 and 2. The steel is referred to in thisreport as 'Type A'. It is used in a number of jack-up rigs operating on the UnitedKingdom continental shelf, and the particular tubular supplied for detailedexamination had been damaged in a collision. After initial telephone discussions withthe client, a programme of work was agreed for the detailed examination of onetubular in order to characterise the microstructural, mechanical and chemicalproperties of the steel. The following tests were made on the major length of tubularmaterial, Sample 1:

� on each quadrant of the tubular (ie at Specimen positions 1-4), longitudinaltensile tests for yield stress, ultimate tensile stress, elongation and reductionin area (all in accordance with BS 18:1987 Tensile testing of metals.

� Charpy impact tests on each quadrant, in longitudinal orientation andundertaken at -40oC for sets of three 10 mm x 10 mm specimens, one settaken at each of the inner and outer surfaces and the mid-wall position

� hardness traverses across the tube wall for each quadrant� chemical composition check via spectrographic analysis for specimens from

both the inner and outer tube surfaces� microstructural characterisation across the wall thickness including

estimation of both the levels of porosity present and the extent of segregation� determination of through-thickness tensile properties on specimens produced

by welding stub pieces to both inner and outer surfaces (these test were onone quadrant only).

Having completed the detailed examination of material from this tubular, a morelimited examination was requested of material from five further tubulars. This wasnecessary because the first tubular examined apparently consisted of three layers,whilst material being used for sulphide stress corrosion cracking (SSCC) tests byother people was apparently mono-layer. Three of the further tubulars (Samples 2, 3and 4) came from the same source as the first, Sample 5 came from another rig andSample 6 was some of the material examined for SSCC. Samples 2-6 were subjectedto microscopical examination and hardness testing only.

Table 1Chemical specification of Type A steel

<0.0200.01-0.052.50-3.500.05-0.080.20-0.401.70-2.30<0.020<0.0120.10-0.400.05-0.10

NA1totalNiNbMoMnPSSiC

4

Table 2Mechanical specification of Type A steel

40 J at -40oC (average of three) 27 J (minimum individual)

14%780 N/mm2 min700 N/mm2 min

Impact (KCV)ElongationUTSYield stress

2.2 EXAMINATIONS

2.2.1 Visual examination

Sample 1 as supplied was a tubular section of approximately 0.8m diameter and2.8m long, with a wall thickness of 30 mm. One end of the section had been gas cutwhilst the other end showed some original, badly rusted, fracture surfaces and furthergas cut regions. There was evidence of accident damage close to this fracturedregion in that the tubular had been dented to a depth of approximately 160 mm. Thewall thickness at this point was 28 mm and the mastic around this area had spalledoff revealing a machined outer pipe surface. The wall of the pipe gave theappearance of having been stretched in the accident. Figures 1 and 2 arephotographs of the tubular.

2.2.2 Tensile properties

As requested, on each quadrant of Sample 1 (ie at Specimen positions 1-4), tensilespecimens were machined from longitudinal blanks cut from the outside third, centralthird and inside third of the wall thickness (referred to as 'top', 'middle' and 'bottom' inTables 3, 4 and 5). Duplicate specimens were taken at each position. The specimenswere tested in accordance with BS 18 and the results are given in Table 3. Thestress-strain curves did not exhibit a discontinuous yield, hence 0.2% proof stressvalues are reported. All of the values observed are in excess of the 700 N/mm2

minimum required in Table 2, being in the range 704-746 N/mm2. The lower valueswere associated with specimens taken from the central position (average 715 N/mm2)compared with an average of 734 N/mm2 for specimens from the outer third and731 N/mm2 for specimens from the inner third of the wall thickness.

Table 3Measured longitudinal tensile properties - Sample 1

56/6216796733Top 250/5518799728Top 13

(cont'd)

35/4515798733Bottom 247/5219797727Bottom 140/5014783719Middle 255/6018774708Middle 156/6219806737Top 255/6019799732Top 12

50/5517805737Bottom 245 (2) 17796733Bottom 130/4012764704Middle 245/5016783712Middle 152/5617793724Top 263(2)16793738Top 11

Reduction of area(%)

Elongation(%)

Ultimate tensile stress(N/mm2)

0.2% proof stress(N/mm2)

Tensile propertiesDepth position(1)Specimen

position(1)

5

-14780700Specification minimum

54/6016790722Bottom 255/6017788722Bottom 146/5515790721Middle 255/6015795734Middle 147/5513807738Top 254/5721809739Top 14

45/5017807746Bottom 250/5515792730Bottom 148/5215774711Middle 253/5816781710Middle 1

Reduction of area(%)

Elongation(%)


0.2% proof stress(N/mm2)

Tensile propertiesDepth position(1)Specimen

position(1)

1 See Section 2.2.2 of the text for positions of specimens2 All specimens oval except those that are marked (2)

2.2.3 Impact properties

Longitudinal Charpy impact specimens were prepared from material taken from theouter third, central third and inner third of the wall thickness from each specimenposition of Sample 1. Sets of three 10 mm x 10 mm specimens were taken at eachposition and testing performed at -40oC. The results obtained are shown in Table 4.The distributions of absorbed energy from the different positions through the wallthickness are shown in Figure 3, where it can be seen that the incidence of lowerenergy values increases from the outside to the inside of the tubular. The averagevalues are in excess of the 40 J requirement and individual values are all in excess ofthe 27 J individual minimum requirement.

Table 4Longitudinal Charpy results at -40oC - Sample 1

40 J min at -40oC (average of three)27 J (minimum individual)

Specification minimum

42424144Bottom56605453Middle 56565260Top4




Average321

Energy absorbed (J at -40oC)Depth positionSpecimen

position

6

2.2.4 Hardness results

Vickers hardness traverses were carried out using a 30 kg load at 1 mm intervalsacross the wall thickness on polished specimens from each quadrant of Sample 1, ieon Specimen positions 1-4. The results obtained are shown in Figure 4. It is clearfrom the figure that hardness variations exist with up to 30 HV difference in hardnessbeing recorded for any one quadrant. A further hardness traverse was carried out atSpecimen position 2, using a 5 kg load in order to determine the maximum hardnessat segregated positions more accurately. The results are shown in Figure 5, where itcan be seen that although individual hardness values are more erratic the generalshape of the hardness profile is similar to that obtained at the same position using a30 kg load. The maximum hardness recorded using the 5 kg load was 325 HV5,compared with 298 HV30 at the same position and 304 HV30 at any position.Arrows are used in both Figures 4 and 5 to mark the positions of the apparentinterfaces between the different layers. It can be seen that the hardness values arelower towards the inside of each layer.

2.2.5 Chemical composition

The chemical composition of the steel was determined by spectrographic analysis onspecimens from each specimen position (quadrant) of Sample 1. Analyses wereperformed on ground portions of the outer and inner surfaces. The results obtainedare shown in Table 5, where it can be seen that whilst the compositions satisfy thechemical specification in Table 1 the analyses performed on the outer surfaces areconsistently richer in most elements, particularly Mn and Ni.

Table 5Measured chemical composition - Sample 1

-0.05/0.08

<0.02-0.01/0.05

2.5/3.5

0.2/0.4

-<0.012

<0.021.7/2.3

0.10/0.40

0.05/0.10

Specification

0.6600.0580.0110.100.0392.800.240.040.0060.0082.040.310.071Bottom

0.7190.0850.0110.110.0443.090.260.040.0050.0112.250.330.071Top4

0.6520.0550.0110.100.0392.780.240.040.0060.0062.020.300.067Bottom

0.7390.0800.0140.110.0443.100.260.040.0100.0102.250.340.090Top3

0.6550.0580.0110.100.0392.810.240.040.0060.0072.040.310.065Bottom

0.7210.0800.0140.110.0453.060.260.040.0100.0092.220.340.080Top2

0.6710.0580.0110.100.0342.830.240.040.0060.0062.050.310.078Bottom

0.7260.0790.0150.110.0443.050.260.040.0100.0092.210.350.087Top1

CEVNbNCuA1totalNiMoCrSPMnSiCDepthposition

Specimenposition

CEV = C + Mn6 + Ni+Cu

15 + Cr+Mo+V5

2.2.6 Microstructure

Macrographs of transverse sections at all four specimen positions examined fromSample 1 are shown in Figure 6. It is apparent that Sample 1 consists of threelayers, each showing varying macrostructure, with more cored dendrites beingpresent towards the outside of each layer. Examples of typical microstructures atvarious depths from the outer surface for Specimen 2 are shown in Figure 7. The

7

general microstructure consists of a mixture of tempered martensite and bainite, withsome areas more enriched by segregation.

Porosity measurements were carried out at various depths on specimens from each ofthe four specimen positions around Sample 1. Results generated by automatic imageanalysis are shown for the mean area of individual pores (Figure 8) and the areapercentage of all pores (Figure 9) with 95% confidence limits. Some inclusions weredetected and included in the porosity data. Specimens from Specimen positions 1 and2 contained most porosity and it is considered that the inclusions would notsignificantly affect the results but at Specimen positions 3 and 4, with lower porositylevels, the results may be affected by the inclusions. However, it is clear that someporosity exists and that its extent is variable.

Compositional profiles were generated across a specimen from Specimen position 1using energy dispersive X-ray spectrometry in a scanning electron microscope.Analyses were carried out at 50 µm intervals along a line from OD to ID of thesample. At each analysis position the electron beam was scanned over an area ofabout 18 x 15 µm during spectrum accumulation to give an average composition ateach small area. The resultant profiles for Mn, Ni, Si, Mo, Cu and Nb are shown inFigure 10. For the first three elements, it is clear that the average concentrationincreases at positions corresponding to the more cored regions of the microstructure.The profiles for Mo, Cu and Nb are not as obvious in demonstrating segregationbecause of their lower concentrations (and hence larger relative errors in theanalyses) and because of the possibility in the Mo results of some additional scatterdue to small MnS inclusions (Mo and S energy peaks overlaps). However, the Mo,Cu and Nb results show highly significant correlations with the Mn and Ni results,indicating that segregation of these elements follows a similar pattern.

2.2.7 Through-thickness tensile properties

Through-thickness tensile specimens were produced in accordance with BS 6780: 1986 by friction welding stubs onto specimens cut from one specimen position ofSample 1. The outer and inner surfaces of the specimens were ground to remove rustprior to friction welding. The specimens were machined to give a gauge length whichonly included material from the tubular. Results for UTS and percentage reductionof area are shown in Table 6. No proof stress data were generated. The resultsobtained are in good agreement with those obtained on the longitudinal tensilespecimens for the same quadrant (Specimen position 1).

Table 6Through-thickness tensile properties - Sample 1

40/47*7854637953507632647871

Reduction of area (%)


Specimen

*Specimen oval

8

2.2.8 Through-thickness impact properties

Having observed that the Charpy properties of longitudinal specimens only just metspecification requirements, it was considered important to determine whether theapparent triple-layer nature of the casting was detrimental to impact resistance.Hence through-the-thickness Charpy specimens were produced by friction weldingstubs onto specimens cut from one quadrant (Specimen position 2) of Sample 1.After machine to 10 x 10 mm sections, the specimens were etched in nital. Notcheswere then located in three specimens from each of three positions - at the topapparent interface line and 2 mm above and 2 mm below the top apparent interfaceline.

The specimens were broken at -40oC and the results obtained are shown in Table 7.The results are all in excess of the average 40 J required at -40oC.

Table 7Through-thickness Charpy results - Sample 1

747676702 mm below top interfaceline

694882762 mm above top interface line

60576064Top interface line2

Average321

Absorbed energy (J at -40oC)Notch locationSpecimen

position

2.2.9 Examinations of material from Samples 2-6

Macrographs of sections prepared from Samples 2-6 are shown in Figure 11.

Vickers hardness profiles across the wall thickness of each of the samples are shownin Figure 12. They were generated using a 5 kg load at 1 mm intervals. As withSample 1, arrows are used to mark apparent interfaces between layers. Theseprofiles too are variable through the thickness, with hardness peaks corresponding tothe outside of the different layers. The samples showing the most uniformappearance macroscopically, ie Samples 4, 5 and 6, show the least fluctuation inhardness.

2.3 DISCUSSION

Examination has shown that the Sample 1 material possesses a structure consistentwith it having been centrifugally cast and built up in three main layers. The tubularappears to have been heat treated to produce a microstructure consisting of temperedmartensite and bainite. The structure provides evidence of significant segregationwith cored dendrites being present, particularly at the outside of each layer.Compositional and hardness variations exist across the wall thickness. Someporosity is also evident.

It is clear from examination of Samples 2, 3 and 4 material and comparisons withSample 1 results that the macrostructure of this material, which all came from thesame platform, is extremely variable. Sample 3 appears to contain four layers whilstSample 4 is mono-layer and Sample 2 may also contain more than one layer (itshows evidence of a chill zone at the outer tubular surface). Sample 5 (from another

9

platform) also appears to consist of more than one layer whilst Sample 6, althoughapparently mono-layer, shows much coarser grain size than any other section.

The mechanical properties of Sample 1 are variable across the wall thickness, withtensile results generally being poorest at the mid-thickness position, whilst Charpyabsorbed energies are generally lowest towards the inside of the tubular.

Whilst the chemical compositions measures at the outside and inside of the tubularSample 1 are within specified limits, large differences in composition are apparentand these are also revealed by the analysis profiles generated by X-ray spectroscopy.Generally, the higher compositions appear to be associated with the outer positions ofeach of the three layers. These compositional variations are responsible for thefluctuations observed in the hardness profiles and are the result of segregation.

The hardness profiles generated on all six samples show variation in hardness acrossthe wall thickness, this variation being most evident in the samples showing amulti-layer appearance.

Generally, the examinations have shown that the tubulars are of good quality, withonly slight apparent deficiencies in ultimate tensile strength at the centre. It shouldbe noted, however, that most the tensile tests in this examination were longitudinaland were taken at three positions through the wall thickness whilst the manufacturerwas required to make a tangential test on a ring removed from one end of the tubularafter quenching and tempering. Such testing would not necessarily give resultssimilar to those in the present examination but there is reason to believe that the steelwould meet the specification requirements.

2.4 CONCLUSIONS

Examination of material from a tubular section of Type A high strength steel hasshown:

� The material did not exhibit a discontinuous yield point, but its 0.2% proofstress exceed the specified minimum.

� The UTS results showed three values less than the required minimum of780 N/mm2. All of these were from specimens from the central third of thetubular wall thickness.

� Charpy properties were satisfactory, with no individual absorbed energyresult less than the specified minimum at 27 J at -40oC. The absorbedenergy values appeared to be lower towards the inner surface of the originaltubulars.

� Hardness testing showed variable results with the highest hardnesses beingobserved towards the outside of each of the microstructural bands.

� Chemical analysis and X-ray spectrometry showed variable compositionswith higher element levels being measured towards the outside of each band.Overall, the chemical compositions were within specified limits.

� Metallographic examination showed the general microstructure to betempered martensite and bainite. Macrosegregation in the form of coreddendrites and porosity was observed.

� Through-thickness tensile properties, UTS and percentage reduction of areawere similar to corresponding longitudinal properties.

10

3. METALLURGICAL CHARACTERISATIONOF TYPE B STEEL

3.1 INTRODUCTION

Following the examination of Samples 1-6 material as described in Chapter 2 here, asomewhat similar examination was carried out on another high strength steel, rolledplate, which has also been used in the jack-up chords of the rig from which Samples1-4 had been obtained. This steel is referred to in this report as 'Type B' steel. Asection was extracted from each of four samples (Samples 7-10) and prepared formetallographic examination. A hardness traverse through the wall thickness was alsomade. This examination was much more limited than that carried out on Samples1-6 to characterise the Type A material.

Later, after the initial results had been obtained on Samples 7-10, some tensile andCharpy tests were requested. Finally, some of the tests were repeated on specimenssubjected to further tempering.

3.2 EXAMINATIONS

3.2.1 Visual examination

Samples 7-10 each consisted of part of a tubular which had been gas cut from the legof the rig. Each sample had at least one tubular brace attached. The material was,therefore, a curved section approximately 770 mm wide in each case with the lengthsvarying between about 900 and 1070 mm. The metallographic samples were takenas far as possible remote from any gas-cut edges and away from welds associatedwith the braces.

3.2.2 Macroscopic examination

The metallographic specimens were extracted by sawing and were prepared forexamination. Figure 13 shows macrographs from each sample etched in 4% nitaland in ammonium persulphate. It is apparent that Samples 7 and 8, which are of31 mm wall thickness, show more segregation than Samples 9 and 10, which have 28mm wall thickness. The thicker plates also show slightly less working of the caststrucutre.

3.2.3 Microscopic examination

Micrographs were taken from each sample at specific positions through the wallthickness, as shown in Figures 14a-d. The microstructures are all similar, showing afine-grained tempered martensite. Some evidence of segregation is again apparent,particularly on Sample 8. Examination of the inclusion morphology shows themajority are globular and relatively dark in appearance. These were probably oxidesor oxysulphides. A few paler elongated inclusions were observed, particularly in thecentrally segregated regions. These were probably unmodified manganese sulphidesand suggest that no effective inclusion shape treatments were carried out duringsteelmaking.

21

3.2.4 Hardness tests

Vickers hardness tests were carried out at 1mm intervals through the thickness ofeach sample using a 5kg load. The results obtained are shown graphically inFigure 15. It is clear that Samples 7 and 8 show reduced hardness at the centre-line,whilst Samples 9 and 10 are slightly harder towards the quarter-depth andthree-quarter-depth positions, with lower hardness at the surface and centre-line.

Figure 16 shows the frequency distribution of the hardness values obtained andshows that the lowest values obtained on the thinner sections (Samples 9 and 10) arehigher than the highest values obtained on the thicker sections, Samples 7 and 8. Inview of these marked differences in hardness it was agreed that chemical analysisshould be carried out on all four samples and that Charpy and tensile propertiesshould be determined on one of the 31 mm thick samples and one of the 28 mm thicksamples.

3.2.5 Chemical analysis

The chemical composition of each of the four samples is shown in Table 8. It can beseen that the two 31 mm thick samples, Samples 7 and 8, are very similarcompositions. The two thinner samples, Samples 9 and 10, are also very similar incomposition to each other, but generally have a higher alloy content (Si, Mn, Cr, Niand Cu) than the thicker samples. This is reflected in the higher carbon equivalentvalue. The sulphur and phosphorous contents of all four samples are low, indicativeof good steelmaking practice.

Table 8Chemical analysis - Samples 7-10

<1.2<0.45<1.25<0.025<0.030.8/1.50.15/0.5<0.19Specificationrequirement

0.742<0.0050.004<0.0050.17<0.00051.000.331.190.0040.0061.140.400.1710

0.742<0.0050.004<0.0050.17<0.00051.000.331.190.0030.0061.140.400.179

0.703<0.0050.005<0.0050.13<0.00050.970.351.050.0010.0041.080.370.178

0.693<0.0050.005<0.0050.13<0.00050.960.351.050.0010.0041.080.370.167

VTiNbCuBNiMoCrSPMnSiC

CEV(%)

Sample

CEV = C + Mn6 + Ni+Cu

15 + Cr+Mo+V5

3.2.6 Tensile properties

Longitudinal tensile specimens were machined from the central portion of one sampleof each thickness (Samples 7 and 9). The specimens were of round cross-sectionwith a cross-sectional area of approximately 150mm2. Duplicate specimens weretested for each thickness and the results obtained are shown in Table 9. As iscommon on quenched and tempered steels the specimens did not exhibit a yield point,hence 0.2% proof stresses are quoted. It is apparent that Sample 7 meets the tensilerequirements of the Type B steel specification in all respects, whilst Sample 9 hashigher UTS and lower percentage elongation than permitted.

22

Table 9Longitudinal tensile properties - Samples 7 and 9

16 minimum785-930685 minimumSpecificationrequirement

64161010882631410369129701788378768198737787

Reduction of area

(%)

Elongation (%)

(5.65ªA)


0.2% proof stress (N/mm2)

Tensile properties

Sample

3.2.7 Charpy properties

Three 10 mm x 10 mm Charpy Vee-notch specimens were machined longitudinallyfrom the outer subsurface and central regions of Samples 7 and 9. As required bythe Type B material specification, the specimens were tested at -40oC. The impactenergies obtained are reported in Table 10. It can be seen that the specimens fromSample 7 gave good impact performance at both subsurface and central positionsalthough the central position results are lower, reflecting the segregation observed atthis position. The results from specimens from Sample 9 are unsatisfactory, withboth subsurface and central positions showing average energies lower than the 40 Jrequired at -40oC. The subsurface results are marginally worse than those from thecentre-line, indicating that, in the absence of segregation, the harder temperedmartensite observed has poorer toughness.

Table 10Longitudinal Charpy results - Samples 7 and 9

4028 individual Specification minimum

3239

2340

3637

3840

SubsurfaceCentre

9

177145

178120

180174

172140

SubsurfaceCentre

7Average321

Energy absorbed (J at -40oC)PositionSample

3.2.8 Additional testing

In view of the high hardness, high tensile strength and low Charpy result on Sample9, it was agreed that further tensile and Charpy testing, in addition to hardnesssurveys, should be carried out on specimens subjected to further tempering. Thespecification for the Type B steel allows for the quenched steel to be temperedbetween 550 and 650oC. Because of availability of specimen material, couponswere extracted from Sample 10. They were tempered for one hour at 550oC, 600oCand 650oC before tensile and Charpy specimens were cut and machined. Similarspecimens were also prepared from the as-received Sample 10 for comparison.

23

Metallographic samples were also prepared in each condition and these weresubjected to optical examination and hardness testing.

Micrographic examinations

Representative micrographs in each tempered condition are shown in Figure 17.There are no marked differences in microstructure.

Hardness tests

The results of through-thickness hardness testing for specimens in each temperedcondition are shown in Figure 18. It is apparent that an additional temper of 1 h at550oC has little effect on hardness, whilst tempering for 1 h at 600oC substantiallyreduces the hardness. This is further indicated in Figure 19 which shows thedistribution of hardness values obtained in each condition.

Tensile properties

The tensile results obtained are shown in Table 11. In the as-received condition, the0.2% proof stress values are similar to those obtained on Sample 9 material, whilstthe UTS values are slightly lower (in line with the slightly lower hardness values).As with Sample 9.

Table 11Longitudinal tensile properties of Sample 10 material, as-received and after

retempering for one hour

16 minimum785-930685 minimumSpecificationrequirement

7020789687*7120793688*6506819850753

6819859767600

6521994923

6317976893550

611410128916114963925As received

Reduction of area

(%)

Elongation (%)

(5.65ªA)


0.2% proof stress (N/mm2)

Tensile propertiesRetemperingtemperature

(oC)

*Yield point

the UTS results are still, however, higher than the 930 N/mm2 maximum required bythe specification whilst the percentage elongation values are below the minimumrequirement. Tempering for 1 h at 550oC appears to have improved the ductilityslightly (resulting in higher percentage elongation and reduction of area values) butthe effect on proof stress and UTS are negligible. Tempering at 600oC produces amarked reduction in proof stress and UTS as well as further improvement in ductilitywhilst tempering at 650oC causes further reduction in strength and improvement inductility. At this tempering temperature, the tensile specimens exhibited a yieldplateau, with some Lüders extension rather than the proof stress values shown by all

24

other specimens. The results at this temperature are close to the minimum requiredfor both yield and ultimate tensile stress.

Charpy properties

The Charpy results obtained on specimens prepared from Sample 10 in theas-received condition and after retempering at 550oC, 600oC and 650oC are shown inTable 12. In the as-received condition, the results are similar to those obtained onspecimens from Sample 9 and fail to meet the requirements of the specification forthe Type B steel. At the subsurface position, two of the three samples failed to meetthe individual requirement of 27 J and the average of the three results were less thanthe 40 J at -40oC required. The samples from the centre-line also failed in this latterrespect, though all three values met the individual requirement. In line with thechanges produced on hardness and tensile properties with retempering, the Charpyresults show a slight improvement after 1 h at 550oC, with subsurface resultsaveraging 55 J at -40oC being obtained. The tempering effect was slightly less at thecentre but the impact values averaged 40 J and just satisfied the specificationrequirement. Tempering at 600oC caused a marked improvement in absorbed energyvalues, giving results similar to those previously obtained on Sample 7 material, withthe subsurface results being higher than the centre-line though both easily satisfy thespecification requirements. At 650oC, the results are more uniform throughout thethickness. The effects of increased tempering are illustrated in Figure 20.

Table 12Longitudinal Charpy properties of Sample 10 material, as-received and after

retempering for one hour

4028 individuals Specification minimum

147141147153151156145153650Subsurface

Centre

126130115133184185181186600Subsurface

Centre

4039424055614560550Subsurface

Centre

3435363027243919As-receivedSubsurface

Centre

Average321Energy absorbed (J at -40oC)Retemperting temperature

(oC)Specimenposition

3.3 DISCUSSION

The examinations of Samples 7-10 clearly demonstrated differences in propertiesbetween the two thicker tubulars (Samples 7 and 8) and the two thinner tubulars(Samples 9 and 10). The two thicker ones are slightly leaner in composition and theirmeasured properties in terms of tensile strength and Charpy absorbed energy aresatisfactory when compared to the Type B steel specification requirements. As wellas leaner composition they also contain more macrosegregation. The extent ofsegregation is partly related to the casting conditions and may be worse in the loweralloyed steels due to different casting conditions.

25

The two thinner plates (Samples 9 and 10) have been shown to fail to meet thespecification tensile strength requirements, being too strong, and to have lower thanthe minimum required Charpy properties. It has been shown that retempering canimprove both tensile and Charpy properties to bring them in line with thespecification. Such changes in mechanical properties produced by temperingdifferences would not produce marked changes in microstructure. At thetemperatures used, the redistribution of the carbides would be relatively slow. It istherefore suggested that the abnormal properties displayed in the as-receivedcondition are a result of incorrect tempering. The chemical analyses and mechanicalproperties of the two thicker tubulars are similar enough to suggest that theyprobably came from the same case and heat treatment batch. Similarly, the analysesand properties of the thinner plates suggest that they also came from one case andheat treatment batch, but a different one from the thicker plates. Hence the availableevidence suggests that the product of one heat treatment batch has been incorrectlytempered.

3.4 CONCLUSIONS

Examinations of material from four samples of Type B high strength steel haveshown:

� Of the four tubulars examined, the two thicker ones appeared to be verysimilar in terms of chemistry and hardness and the mechanical properties ofthe only one tested were satisfactory.

� The two thinner tubulars also appeared to be very similar to each other interms of chemistry, hardness and mechanical properties, but they wereslightly more alloyed than the thicker tubulars.

� The mechanical properties of the thinner tubulars were unsatisfactory, beingexcessively strong and sub-standard in toughness. The incorrect propertiesprobably resulted from inadequate tempering.

26

4. THE HYDROGEN SUSCEPTIBILITY OF TYPE A ANDTYPE B STEELS

4.1 INTRODUCTION

This chapter describes an investigatory programme to characterise the hydrogensusceptibility of the Type A material described in Chapter 2 and the Type B materialdescribed in Chapter 3.

The brittle cracking behaviour observed in jack-up rigs was thought to have beencaused by exposure to hydrogen sulphide or by over-protection from the rig'scathodic protection system. Prior to the test programme described in this chapter, theon-site conditions were examined to ascertain the presence of hydrogen sulphide andthe actual cathodic protection levels on the rigs. This work has been reported (3) andthe data may be summarised as follows:

� H2S was detected but was considered to be locked up in a sludge at thebottom of the spud cans - the water phase in the spud cans contains little, ifany, free H2S

� the level of cathodic protection was measured as -981 mV to -1013 mVagainst an Ag/AgC1 reference electrode.

Materials were tested in two batches: the first in an initial test programme and thesecond in a 'focused' test programme to assess the effect of applying various levels ofcathodic protection:

� Batch 1 - initial test programmeTwo pieces were received from a chord of the same rig that Sample 5 hadcome from (see Chapter 2). Each piece measured approximately 300 mm x100 mm and they are referred to here as Sample 11.

The material was supplied to the specifications detailed in Tables 1 and 2 ofChapter 2. It was understood that one piece of the material was belowspecification. Later examination revealed casting defects in both pieces ofmaterial, with more in one than the other.

� Batch 2 - focused test programmeThe specimens for the focused test programme were taken from Samples12-16 (the correlations of these samples with earlier samples is explained inTable 13). Their metallographic characteristics are described in Chapters 2and 3. Materials exhibiting a multi-layer microstructure were specificallyselected so that the effect of any interfaces could be assessed during the testprogramme.

38

Table 13Material samples selected for 'focused' test programme

Homogeneous strucutre Type B916Two layersType A515Single layer containing weld Type A414Single layerType A113Single layer containing weldType A412

MicrostructureSteeltype

Corresponding Sample Nofrom Chapters 2 and 3

Sample No forChapter 4

These samples were obtained from different rigs from those referred to in Chapter 2

4.2 TEST TECHNIQUE

It was considered that slow strain rate testing was the most suitable method toproduce the required results within a short period. The technique comprises loadinga round tensile specimen slowly under conditions of constant strain rate, with thegauge area exposed to the test environment. There are four parameters to assessusing this test method:

� time to failure� reduction in cross-sectional area� sample elongation� fracture surface examination.

For this investigation, the time to failure and fracture surface examination wereconsidered to be the more important parameters. The reduction in area was not usedsince the variable porosity and large grain size of the material would have renderedthe data of little value.

The slow strain rate technique has the ability to differentiate between brittle andductile behaviour.

In all test programmes of this type an air control sample is run to provide referenceinformation. A typical load versus time plot is illustrated in Figure 21. For thematerials under consideration this represents the behaviour of a ductile material astypified by the fracture morphology shown in Figure 22. It a material behaves in abrittle manner, the plot illustrated in Figure 23 would be typical and the fracturemorphology shown in Figure 24 would be characteristic of such a failure.Essentially, brittle behaviour is characterised by a reduction in time to failure and achange in fracture morphology from ductile dimples to planar cleavage.

The equipment used to conduct slow strain rate tests is shown in Figure 25. A veryimportant consideration is the rate of strain; it must be slow enough for theenvironment to act. A strain rate of =1.6 x 10-7 S-1 was used as experience hasshown this rate provides satisfactory results.

4.3 TEST PROGRAMME

Initially, a series of electrochemical tests was conducted on Sample 11 material(Type A) in order to generate a set of polarisation curves. These data indicate thefree corrosion potential of the steel under different test conditions and allowpredictions of possible hydrogen evolution to be made. Potentiometric curves wereobtained on Sample 11 material for the following conditions:

39

� aerated seawater, pH 8.26� deaerated seawater, N2 purged, pH 8.8� collected spud can water, N2 blanket, pH 7.5.

In total, 48 slow strain rate tests were conducted. The first tests were conducted onSample 11 material in varying environments. The remaining 36 tests were conductedon material from Samples 12-16 in order to focus on the lowest cathodic protectionlevel which caused little or no hydrogen damage to the material under the imposedtest conditions. The full catalogue of test environmental conditions is listed in Tables14 and 15.

Table 14Test environmental conditions for Type A material - Sample 11

-945Spud can seawater under nitrogen blanket12-920Spud can seawater under nitrogen blanket11-945Spud can seawater under nitrogen blanket10-890Spud can seawater under nitrogen blanket9-800Spud can seawater under nitrogen blanket 8

No CPSeawater with dilute hydrogen sulphide (approximately 500 ppm)

7

No CPSeawater control (spud can seawater under nitrogen blanket)

6

No CPSeawater and hydrogen sulphide at saturation level(approximately 3000 ppm)

5-980Spud can seawater and nitrogen blanket 4

-1050Deaerated seawater 3-1050Aerated seawater 2

-Air control 1

Cathodic protection level (mV vs SCE)

Environmental conditionsTest No

Table 15Test environmental conditions for the 'focused' series of tests

Samples 12-16

-837Spud can seawater under nitrogen blanket28-837Spud can seawater under nitrogen blanket27-825Spud can seawater under nitrogen blanket26-825Spud can seawater under nitrogen blanket25-850Spud can seawater under nitrogen blanket24-850Spud can seawater under nitrogen blanket23

-Air control 22-Air control 21

Sample 13



Sample 12



40

Table 15 (Cont)

-850Spud can seawater under nitrogen blanket48-850Spud can seawater under nitrogen blanket47-980Spud can seawater under nitrogen blanket46-980Spud can seawater under nitrogen blanket45


Sample 16



Sample 15

-825Spud can seawater under nitrogen blanket34-825Spud can seawater under nitrogen blanket33-850Spud can seawater under nitrogen blanket32-850Spud can seawater under nitrogen blanket31


Sample 14



4.4 RESULTS AND DISCUSSION

4.4.1 Electrochemical potential tests on Sample 11 material

The three potential curves for different environmental conditions are illustrated inFigure 26. It can be seen that for all three conditions H2 or Habsorbed can be producedbelow -700 mV (versus a saturated calomel electrode (SCE), which is = 9 mV awayfrom Ag/AgC1). The effect of the oxygen level in both aerated and deaeratedseawater environments can be seen in the change in slope in the cathodic sweep in theregion -750 mV to -950 mV (v SCE) where a distinct limiting current density ofabout 100 µA/cm2 can be seen for the aerated seawater. This is not as pronouncedin the dearated environment and may even be due to oxide reduction from thepreviously imposed anodic polarisation.

At any potential more negative than -700 mV, some hydrogen will be produced aspart of the cathodic reaction sustaining Fe → Fe2+ + 2e-. However, at neutral pH, theoxygen reduction reaction will predominate until the potential is below the intersectof the oxygen line with the H2 line as shown in Figure 26. It should be noted that thecathodic reaction will be mixed hydrogen evolution and oxygen reduction but that, asthe potential falls (becomes more negative), the proportion of OH-/H2 in the overallcathodic reaction will change to produce more hydrogen.

The tests indicate that significant hydrogen could be produced at cathodic protectionlevels of -800 mV or at more negative values.

4.4.2 Slow strain rate tests on Sample 11 material

The load v time plots for Tests 1-12 are illustrated in Figure 27. The data aresummarised in Table 16. All specimens had a reduction in time to failure whencompared with the air control test, although it should be noted that some specimens

41

exhibited casting defects which would have contributed to premature failure(especially Tests 3, 8, 10 and 11).

Scanning electron microscopy was undertaken on the fracture surfaces and typicalmicrographs of salient features are included in Figure 28. Table 17 summarises thefeatures illustrated. As expected, the air control test exhibited typical ductilebehaviour.

Table 16Summary of slow strain rate testing - Sample 11

00102-945Spud can seawater under N2 blanket120093-920Spud can seawater under N2 blanket110087-945Spud can seawater under N2 blanket10--203-890Spud can seawater under N2 blanket9--140-800Spud can seawater under N2 blanket8

--66-Spud can seawater and dilute H2S (500 ppm)

7

--190-Spud can seawater under N2 blanket(solution control)

6--80-Seawater and H2S (3000 ppm) 5

0.1897.0108-980Spud can seawater under N2 blanket40.1485.5109-1050Deaerated seawater30.3513154-1050Aerated seawater2

-37260-Air control 1

Reduction in area

normalisedto air control

Reductionin area(%)

Time tofailure (hours)

Cathodic protection

level(mV vs SCE)


Table 17Failure modes of Sample 11 material subjected to slow strain rate testing

Ductile/brittle-945Spud can seawater under N2 blanket12-**-920Spud can seawater under N2 blanket11-**-945Spud can seawater under N2 blanket10Ductile -890Spud can seawater under N2 blanket9Ductile**-800Spud can seawater under N2 blanket8Brittle**-H2S level 500 ppm 7Ductile-Seawater control 6Brittle-H2S level 3000 ppm 5Brittle*-980Spud can seawater under N2 blanket4Brittle**-1050Deaerated seawater3Brittle*-1050Aerated seawater2Ductile -Air control 1

Failure modeCathodic protection level (mV vs SCE)


*Brittle areas evident around edge of specimen: ductile areas due to specimen overload caused by reduction inarea also observed.

**Casting defects observed which probably contributed to the failure, however the environment still causedsignificant embrittlement.

Two tests were conducted in H2S - containing environments at levels of 3000 ppm(the accepted saturation level) and 500 ppm (the maximum level attainable under theaction of sulphate-reducing bacteria). They showed that significant hydrogendamage would occur to material exposed to such an environment. However, the siteinspections of drilling rigs had reported very little, if any, presence of H2S in theregions of cracking and therefore more emphasis was placed on the cathodicprotection system investigation.

The environment in the spud cans was simulated for the remaining tests in thisprogramme (and in the focused programme - see Section 4.4.3) using collected spudcan under a nitrogen blanket. This produced a partially deaerated condition similarto that observed on site.

42

A control was run in seawater with no cathodic protection to verify that the seawateritself was not causing the problem. The results were acceptable and the remainingeight tests were all run under applied cathodic protection conditions ranging from-1050 mV to -800 mV vs SCE. All these tests conducted under cathodic protectionexhibited an overall loss of ductility, a reduction in time to failure and,fractographically, some degree of cleavage. The material under investigation didcontain casting pores and the data are subject to a degree of variability. Inparticular, the test conducted at -800 mV, although having a reduced time to failure,exhibited very little evidence of brittle cleavage.

4.4.3 Focused cathodic protection tests on Samples 12-16 material

Combining the cathodic protection levels measured on site and the slow strain ratetesting data from Tests 1-12, it was evident that a level of cathodic protection shouldbe focused upon where hydrogen damage was minimised and protection afforded.Therefore a further 36 tests (Tests 13-48) were conducted to ascertain:

� if other materials being used exhibited similar characteristics� the optimum cathodic protection level to use in practice.

Four Type A materials (Samples 12-15) were tested and one Type B material(Sample 16) was tested for comparison purposes. Each test at each chosen level ofcathodic protection was duplicated to limit experimental variations.

The load v time plots are illustrated in Figures 29-33 and summarised in Table 18.Scanning electron fractography was undertaken on typical specimens and the featuresobserved are illustrated in Figures 34-38. Table 19 summarises the fractographicexaminations.

Sample 12 material - Type A (Tests 13-20)

The two air control test samples showed normal ductile failure. The two -850 mVcathodic protection tests exhibited a significant reduction in time to failure and somecleavage on the fracture surface. Two tests were run at -825 mV, one produced aresult nearing the air control condition but the second exhibited a defective fracturesurface (casting porosity). Two further tests were conducted at -837 mV and againone sample had a defective area. The other exhibited signs of hydrogen damage butnot as severe as the -850 mV test.


Tests 21-28 followed the pattern of Tests 13-20. The air tests gave typical ductileresults and the -850 mV tests showed a reduction in time to failure but not assignificant as that of Sample 12. The remaining four tests, at -825 mV and -837 mV,also exhibited a small reduction in time to failure with some evidence of cleavage.


This material had the same parent metal as Sample 12 but contained acircumferential weld and the specimens were machined so that the weld portion was

43

in the centre of the gauge lengths. The two air tests produced ductile failures asexpected. However, the time to failure was lower than that for the parent metal andshowed a reduced elongation of the specimens. Four tests under cathodic protectionconditions were conducted, two each at -850 mV and -825 mV. The 850 mV and-825 mV. The -850 mV specimens had a reduced time to failure and showed brittlefracture characteristics but the -825 mV test specimens had the same time to failureas the air control tests and exhibited a ductile fracture mode.

Table 18Summary of slow strain rate testing of focused cathodic protection programme

Samples 12-16

0.47630240-850480.40325170-850470.2214215-980460.2918180-98045

-63255Air control44-62240Air control43

Sample 16

0.90920305-837420.83831265-837411.1628305-825400.59422250-825390.85314230-850380.59422265-85037

-24*235Air control36-37*290Air control35

Sample 15

0.4925280-825340.39517245-825330.35218210-850320.1868190-85031


Sample 14

0.6338290-837280.529265-837270.8551280-825260.56933265-825250.5533265-850240.4124240-85023


Sample 13

0.37516165*-837200.17510225-837190.2512175*-825180.17510315-825170.1567.5225-850160.074205-85015


Sample 12

Reduction in areanormalised to

air control

Reduction in area(%)

Time to failure(hours)


Test No

*High levels of porosity observed (casting defects)

Note: In all cases the final gauge diameter away from the fracture was reduced by approximately 3.5%, therefore

figures are comparative and not definitive.

Sample 15 material - Type A (Tests 35/42)

A high level of porosity was noted in this material.

The -850 mV tests exhibited some brittle behaviour but a comparison with the aircontrol specimens is difficult as the high level of porosity masked the embrittling

44

effect. Overall, some brittle behaviour was evident, the quantity reducing as thecathodic protection level approached -825 mV.

Sample 16 - Type B (Tests 43-48)

Six tests were conducted, two as air controls, two at -980 mV and two at -850 mVfor general comparison with the earlier work. The strength of the material wassignificantly higher than that of the Type A steel samples, 1120 MPa compared with850 MPa. The air controls exhibited typical ductile failure. The cathodic protectiontest specimens all exhibited a degree of brittle failure with associated reduction intime to failure.

Table 19Fractography summary from focused cathodic protection testing programme -

Samples 12-16

Brittle facets48Low magnification fracture48Brittle with secondary cracking47Low magnification fracture47Cleavage with secondary cracking46Low magnification failure46Brittle failure45Low magnification failure45Low magnification necking 44Ductile dimples43Low magnification fracture43

Sample 16

Cleavage facets41Casting porosity 41Ductile features 39Low magnification fracture39Brittle area37Casting porosity37Ductile dimples35Low magnification fracture35

Sample 15

Intergranular attack 34Ductile failure34Low magnification fracture34Ductile failure33Cleavage planes32Brittle features32Low magnification fracture32Casting porosity30Ductile dimples30Low magnification fracture30

Sample 14

Cleavage27Low magnification fracture27Brittle/cleavage area24Ductile areas24Low magnification fracture 24Ductile dimples21Low magnification fracture21

Sample 13

Secondary cracking19Ductile 19Casting porosity18Ductile dimples18Low magnification fracture18Secondary cracking16Low magnification fracture16Ductile dimples 13Ductile necking13

Sample 12 Feature Test No

45

4.4.4 Results summary

The Sample 11 material (Type A steel) did not tolerate H2S at levels of 500 ppm orabove when loaded under conditions approaching yield. The application of cathodicprotection of the order of -900 mV (v SCE) produced hydrogen damage. Thematerial neared its UTS level as measured in the air control test before the effect wasevident. Hydrogen severely reduced the capacity of the material for plastic flow andhence induced embrittlement.

Overall, the tests conducted at -850 mV on Samples 12-15 materials (Type A) allshowed a reduction in time to failure when compared with the air control specimens.The effect of testing at less negative cathodic protection levels was to reduce theeffect of hydrogen. In most cases, the material approached its UTS under the testconditions and the effect of the cathodic protection hydrogen was again to lower thecapacity of the material for plastic flow.

The Sample 16 material (Type B) performed similarly to the Type A material.Overall, even in the air tests, its plastic flow capacity was lower than that of the TypeA material, making the embrittling effect of cathodic protection less marked.

4.5 CONCLUSIONS AND RECOMMENDATIONS

4.5.1 Conclusions

Slow strain rate tests have shown that:

� All the materials under investigation suffered hydrogen-induced stresscorrosion cracking under the application of cathodic protection at loadinglevels approaching the measured UTS. The level of damage was reducedwith decreasing cathodic protection levels (ie less negative potentials).

� At a CP level of -825 mV, the hydrogen damage appeared to be small andthe performance of the materials in slow strain tests was generally lessaffected.

� Exposure of Type A material to wet H2S caused severe damage, failureoccurring well before the 0.2% proof stress.

4.5.2 Recommendations

The following recommendations can be made:

� Exposure of the materials examined to wet H2S should be avoided. It wouldbe useful to clarify the effects of very dilute H2S concentrations (ie less than500 ppm) to ascertain if the biocide packages commonly used in the industryto control sulphate-reducing bacteria activity are necessary.

� The applied cathodic protection level should be limited to -825 mV. If suchcontrol is not practicable, a coating system (together with regularinspections) should be considered.

� The load at which failure occurred was approaching the measured UTS.Any tensional loading should be limited to below the 0.2% proof stress.

46

5. CATHODIC PROTECTION AT LIMITED POTENTIALS

5.1 INTRODUCTION

It has been demonstrated in Chapter 4 that some high strength steels used in jack-uprigs suffer from hydrogen embrittlement in the potential range conventionally usedfor cathodic protection (CP). One option for overcoming this problem is to limit thepotential produced on steel by the CP system to a value more positive than that atwhich embrittlement has been shown to become significant, ie no more negative than-830 mV v Ag/Ag CI (for -825 mV v SCE). This chapter provides advice on thefeasibility of this approach and discusses the inherent problems associated with it.What corrosion risks could be inherent in limiting the potential to this limit? Is thereany practical experience of operating offshore CP systems at such potentials?

5.2 THE NEED FOR LIMITED CP POTENTIALS

Uncoated carbon steel in seawater is protected from corrosion in seawater atpotentials more negative then -800 mV (v Ag/Ag CI) and any properly designed CPsystem will meet this criterion. In systems designed to comply with publishedrecommendations and codes of practice (4-6) produce more negative potentials. Thus abare steel strucutre should achieve -800 mV at launch and the potential will driftmore negative as protective calcareous deposits are formed by the CP process.

For a system using sacrificial anodes, the most negative potential possible will bethat of the anodes (-1050 or -1100 mV dependent on anode type). This potentialcould well be achieved close to anodes on a bare steel strucutre. The situation isreversed for a well coated offshore strucutre or pipeline. Here, the initial potentialwill be close to the anode potential at all points, and will move less negative with timeas the coating deteriorates. The result is that for either coated or bare steel structuresprotected by sacrificial CP, potentials at some point during the operating life arelikely to be substantially more negative than the -830 mV limit apparently necessaryto avoid hydrogen embrittlement in high strength steels.

This situation will be exacerbated if impressed current is used for CP since there isthen no 'natural' negative limit and each anode will usually have a higher currentoutput than a sacrificial anode, giving more negative local potentials.

Clearly, a non-conventional approach to CP design is required if potentials are to belimited to no more negative than -830 mV. There are a variety of possibleapproaches to this problems.

5.3 METHODS OF LIMITING CP POTENTIALS

5.3.1 Control of anode distribution

A conventional CP design balances the current required to provide protection to thesteel against the current output (IA) available from the sacrificial anodes. The latteris calculated using Ohm's Law:

IA = VRA

73

where V is the anode driving voltage, ie the difference between the anodepotential and the local steel potential

RA is the anode resistance and is dependent on the anode size and shape.

For a bare steel strucutre, the current required to polarise the steel initially issubstantially higher than the maintenance current needed for protection oncepolarised. Sufficient anodes are provided in a conventional CP design to supply theinitial, higher, current at the target protected steel potential of -800 mV. For, say,zinc anodes this is at a driving voltage, V, of 250 mV. Since RA is essentiallyconstant, once the initial current requirement falls to the maintenance level, Ohm'sLaw requires that V also falls. Polarisation current densities are typically three timeshigher than maintenance levels, and hence V will fall from 250 mV to about 80 mV,giving a steel potential (the maintenance level) of -970 mV - much more negativethan the target level of -830 mV.

In order to guarantee that steel potentials never become more negative than -830 mVit would be necessary to limit the number of anodes to that capable of providing onlythe maintenance current density at (for zinc) a driving voltage of 220 mV. A similareffect could be achieved by limiting anode sizes and hence increasing the value of RA.There are major problems with attempting to control potential by this method:

� It cannot be applied to a coated structure, where the current demanded fromthe anodes increased with time.

� By definition, there will be insufficient anodes to provide the initialpolarisation current on a bare strucutre, ie the steel would never becomeprotected. This could be overcome by the additional use of a small numberof short life, high output, anodes to provide for polarisation. Such anodeswould need to be in the form of ribbon. This technique has been used byAramco to reduce the weight requirements on small platforms in the ArabianGulf. There, magnesium ribbon with a life of 1-2 months was used for initialpolarisation, with a limited number of very large aluminium anodes cateringfor maintenance current needs. Magnesium anodes could not be used withthe high strength steels under consideration here since their very negativepotentials could exacerbate the embrittlement that the design seeks to avoid.However, a design using aluminium alloy ribbon anodes could, in principle,be developed.

� It would be necessary to know accurately the maintenance currentrequirement at the design stage. This is a more serious obstacle. Jack-uprigs are used in a variety of locations and maintenance current densitydepends, inter-alia, on seawater temperature, oxygen content and flow rate,as well as the previous polarisation history of the steel. It is highly unlikelythat a 'correct' value could be defined.

It is the last point in this list, rather than any practical design problems, that makesthis approach unviable.

5.3.2 Use of dielectric shields

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Dielectric shields are used in impressed current CP systems to limit steel potentialsclose to high current output anodes. In principle they could be used for the samepurpose with sacrificial anodes but the size of shield required is, to a firstapproximation, inversely proportional to the difference between the generalprotection level required and the most negative acceptable potential. For the problemunder consideration, this difference (30 mV) is much smaller than that for a typicalimpressed current installation (500 mV) and very large sizes (several metres across)would be needed. This approach can, therefore, also be ruled out.

5.3.3 Low driving voltage anodes

In principle, the CP potential limit of -830 mV could be met by using a sacrificialanode material with an operating potential of about -850 mV. There appears to beno basic theoretical impediment to the production of such an alloy but the authorsknow of no such material having been developed commercially - despite its attractionfor the protection of steels susceptible to hydrogen embrittlement.

5.3.4 Sacrificial coatings

Flame sprayed aluminium (FSA) coatings have been used for the protection of steeloffshore over a number of years. They have been in use on a steel tower off theDurhan coast since 1958 and on steel pilling in both Norway and the USA(7). Bothsealed and unsealed aluminium coatings have been subjected to considerable testingand are reported as giving good protection in seawater for periods up to 18 years.

Probably the best reported case of FSA coatings offshore is their use on the Conoco'Hutton' TLP. They have been used to protect tension leg components after ConocoUSA had conducted a series of investigations on the metallurgical effects of the flamespraying process on the substrate(7) and its effect on corrosion fatigue. Theelectrochemical properties have also been investigated(8) and a recent paper(9) reviewsthe performance of these coatings on the TLP after two years service. It isunderstood that tension leg elements removed for examination in 1990, after 6 yearsexposure, were found to be in excellent condition with no significant change inappearance since the two year study(10). FSA has also been used on the 'Hutton',which operate at 80oC.

FSA coatings are useful for the protection of high tensile steels (eg the 'Hutton'tension legs) because they maintain potentials in the range -850 mV to -900 mV overlong periods. They can also be applied at sufficient thickness to ensure a 20 yearoperating life(8). Although these potentials are somewhat more negative than the -830mV target of this investigation, FSA coatings may be worth further study for thejack-up leg application.

5.3.5 Voltage-limiting devices

Schottky barrier diodes have been successfully used by Boeing to preventoverprotection of ferritic stainless steel struts and hydrofoils on marine hydrofoilvessels with aluminium hulls(11). The stainless steel components are electricallyisolated from the hull but connected to the aluminium sacrificial anode system on thevessels using low-voltage-drop diodes. The diodes only allow the passage of currentwhen the potential difference between the aluminium anodes and the stainless steel is

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greater than 300 mV. Thus the steel cannot reach potentials more negative than -800mV when using AI-Zn-In alloy anodes, or -750 mV when using zinc.

This arrangement has recently been used for the protection of duplex stainless steelflowlines installed as part of the Marathon Central Brae development in the NorthSea. AI-Zn-In bracelet anodes have been installed over the neoprene pipe coatingthus isolating them from the pipeline steel. Connection was then made to the latterthrough Schottky barrier diodes. See Figure 39. Data obtained from the fixedmonitoring system on the flowlines show clearly that the system is working well,giving protection potentials in the range -840 to -790 mV. The overprotection atPMU 411 shown in Figure 40 has been attributed to faulty electronics. The

fluctuations in potentials between -840 and -790 mV can presumably be attributed tothe diodes 'switching'.

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It would appear that this approach is capable of achieving the -830 mV limit and isworthy of further study. Detailed design attention would need to be paid to theengineering of anode mountings and electrical connections, but the installed systemhas the advantage of being passive in the sense that it would require no regularmaintenance attention after installation. It represents by far the best option for theprotection of high strength steels in jack-up rigs.

5.3.6 Potentiostatic impressed current

It is possible to control an impressed current system potentiostatically. However,with such a narrow range of control potentials required, a large number of relativelylow output anodes would be needed. Although possible, this would lead toengineering and control difficulties and is not considered worth pursuing.

5.4 POSSIBLE RISKS OF LIMITING POTENTIALS

The free corrosion potential of steel in seawater is approximately -640 mV.Application of CP current moves the potential in a negative direction, the relationshipbetween the (anodic) current applied and potential being given by the Tafelequation(12).

(E − Eo) = a + b log I

where E is the potential reachedEo is the free corrosion potentialI is the currenta and b are the constants.

It is not necessary to pursue the theory of corrosion and CP further here, but itfollows from the application of the theory that:

� If the potential of carbon steel is lowered, by the application of a cathodiccurrent, to a value more negative than a certain value of E then corrosionwill cease. Experience shows that the relevant potential value is -800 mV forcarbon steel in seawater.

� Lowering the potential of carbon steel to a value between the free corrosionpotential (-640 mV) and -800 mV produces a reduction in corrosion ratewithout preventing corrosion completely. This requires less current than fullprotection.

� Due to the logarithmic nature of the current potential relationship, the firstincrement of potential shift produces a greater effect than later increments.

� Lowering the potential below -800 mV does not confer any additionalcorrosion protection and is wasteful in terms of additional current.

It follows from the foregoing that protection of carbon steel in the range - 830 to 800mV, as required to present hydrogen embrittlement, represents the most efficientmode of cathodic protection and has no adverse protection consequences. If thetechnique used to limit the CP potential allows potentials less negative than -800 mV,say down to -760 mV, then the effect will not be disastrous due to the logarithmic

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relationship discussed above. At potentials around -760 mV, the corrosion rate ofsteel will be some two orders of magnitude less than the free corrosion rate - not aserious consequence.

5.5 CONCLUSIONS

From the investigation reported here it can be concluded that:

� The most appropriate technique for limiting CP potentials to -830 mVappears to be the use of Schottky barrier as voltage-limiting devices. Thisshould be capable of providing potentials in the range -830 to -770 mV ifproperly engineered.

� Flame sprayed aluminium coatings will also restrict the potential, but only tothe range -900 to -850 mV.

� No other method is available at present for achieving the -830 mV limit.

� There appear to be no significant corrosion risks in operating CP in a limitedpotential range.

6. GENERAL CONCLUSIONS

Some high strength steel types used for the construction of jack-up drilling rigs aresusceptible to hydrogen embrittlement, causing cracking in highly stressed areas.This has been shown to take place in both aerobic and anaerobic conditions, but ismost likely to occur where hydrogen is generated by the cathodic protection systemand where there is insufficient oxygen available at the site to remove the hydrogen.

To avoid these conditions:

� CP levels should be arranged so that generation of hydrogen is avoided or atleast minimised

� particular care should be taken where anaerobic conditions exist (ie eitherwithin spud cans or where spud cans and lower leg sections are buried inmud) to ensure that CP potentials are within the safe range for the steel beingused

� care should be taken during repair welding, by using agreed procedures, tominimise residual stresses in the areas adjacent to the weld

� the weld procedures set up for new building at the design and constructionstages should ensure that the combined residual stress plus applied stressesmet in service do not exceed the yield stress or the 0.2% proof stress. Weldand fabrication procedures (including the sequence of fabrication) should bearranged to minimise residual stresses, for example by avoiding major weldsin high strength steel which is highly restrained.

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6.1 DRAFT INTERIM GUIDANCE

As a result of the research described in earlier chapters of this report and discussionsheld with interested parties, draft interim guidance has been prepared to control andprevent hydrogen-assisted cracking in high strength steels. The guidance isreproduced here, and comments made on it at a discussion meeting held on 4December 1990 are summarised in the appendix to this report.

Interim Guidance Document

HYDROGEN ASSISTED CRACKING OF HIGH STRENGTHSTEELS IMMERSED IN SEAWATER

INTRODUCTION

Since early 1988, the Department of Energy has been investigating a widespreadproblem of hydrogen assisted cracking of high strength steels, particularly those usedin the leg chords and spud cans of many self-elevating offshore installations.

The findings of the investigations and of several meetings with owners, operators,Certifying Authorities and relevant industry associations are set out below asconclusions and interim guidance so that all may be aware of the problem and of howit may be dealt with.

This information should be read in conjunction with Sections 4, 12, 21 and 33 of theDepartment's Guidance on design, construction and certification 4th Edition 1990.It will be included in an amendment to the 4th Edition, issued by the Health andSafety Executive.

CONCLUSIONS

1. Some high strength steels can be susceptible to embrittlement by hydrogenwhich can arise from the CP system, particularly in the absence of oxygen, or fromthe generation of hydrogen sulphide in anaerobic conditions (eg in spud cans).

2. The susceptibility of the steel to hydrogen assisted cracking is influenced byresidual stresses which can be high in the case of high strength steels with nopost-weld heat treatment.

3. The susceptibility of steels to hydrogen embrittlement can be established bytests.

4. Commonly used CP systems are likely to give excessive negative voltageswhich will render high strength steels susceptible to hydrogen embrittlement.However, CP systems are available which can offer controlled voltages with apredetermined limit to the negative value.

5. The potential susceptibility of high strength steels to hydrogen embrittlementshould be taken into account at the design stage as part of the material selectionprocess, as should the design of the CP system. Internal volumes which can besealed, such as the interior of spud cans, may be protected by a corrosion inhibitorwith an approved biocide.

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6. Surveys need to take account of the possibility of the occurrence of hydrogenassisted cracking.

7. The finding and characterisation of hydrogen assisted cracks requires highquality NDT. Where paint coatings are removed (eg in dry dock) to facilitateinspection of areas which would normally be submerged and thus protected by CP,the coatings must be fully reinstated. If underwater inspections are necessary, paintcoatings should not be removed, but NDT methods which can give adequate levels ofdetection through paint coatings should be used.

8. Consideration should be given to special surveys of legs which have beensubmerged in mud or slit.

FUTURE ACTIONS

9. Jack-ups known to be vulnerable.

a) A further review of in-service experience and trend analysis is required oncompletion of the latest round of structural surveys.

b) Requirements for future surveys must be defined. It is suggested thatdetailed NDT surveys of vulnerable areas on legs and within spud cans beundertaken at intervals not exceeding 2 years.

c) CP systems should be modified as soon as possible to maintain closecontrol of the CP voltage at all times, within the limits determined by thesusceptibility of the steels in the strucutre.

d) The above should apply also when the jack-up unit is working alongsideor over a fixed installation to which it may be connected electrically,either directly or fortuitously.

e) CP surveys should be undertaken on a routine basis (at least annually) tocheck the functioning of the CP systems.

f) There should be consideration of special surveys to check on weld repairswhich involve high strength material.

g) Consideration should be given to special surveys of legs which have beensubmerged in mud or silt. To assist with these surveys the owner shouldmaintain records which include the depth of penetration, the type of soil,the duration, the state of the CP system and any modifications made to itto accommodate the changed circumstances.

10. Other existing jack-ups

a) Surveys should include a methodical search for evidence of hydrogenassisted cracking. Survey schedules should be amended accordingly andSurveyors properly trained in what to look for.

b) All cracks in high strength material found during survey should beinvestigated. Cracks should not be dismissed as arising due to previouslyundetected fabrication defects.

c) Tests should be undertaken to determine the susceptibility, of thematerials used, to hydrogen embrittlement/cracking at the existing CPlevels. (The simplest method available is slow strain rate testing).

d) CP measurements should be taken routinely to ensure that CP levels arenot such as to give rise to hydrogen embrittlement/cracking.

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e) Consideration should be given to special surveys of legs which have beensubmerged in mud or silt. To assist with these surveys the owner shouldmaintain records which include the depth of penetration, the type of soil,the duration, the state of the CP system and any modifications made to itto accommodate the changed circumstances.

11. New construction

a) Materials should be selected having regard to their susceptibility tohydrogen embrittlement.

b) Proposed materials should be tested to determine their susceptibility atappropriate CP levels (refer to 9c above).

c) CP system design should be compatible with 10c.d) Welding procedures, pre- and post-weld heat treatment and NDT should

take due account of potential delayed cracking problems. Final NDTshould take place a sufficient time after welding to identify delayedcracking (typically 48 hours or more for high strength steels).

e) Assembly methods should be arranged so that the residual stress andhardness levels in the heat affected zones of welds in high strength steelare kept to values shown to be acceptable in the tests described in 10cabove.

f) Quality assurance programmes should take account of all of the aboveand include schedules for ensuring that specification details (eg theprocessing and heat treatment of high strength steels) are met.

A R McIntosh30/9/91

REFERENCES

Department of Energy (now Health and Safety Executive). Offshore installations: Guidance on design, construction and certification. Fourth edition, 1990, [ISBN 011 412961 4] and amendments [ISBN 0 11 886389 4], HMSO London.

Marine Technology Directorate Ltd. Design and operational guidance on cathodicprotection of offshore structures, subsea installations and pipelines. 1990, MTDPublication 90/102, London.

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7. REFERENCES

1 DEPARTMENT OF ENERGYOffshore Installations: Guidance on Design, Construction andCertification Department of Energy, fourth edition, 1990, available from HMSO.

2 BRITISH STANDARDS INSTITUTIONTensile Testing of Metals (Including Aerospace Materials)BSI, BS 18: 1987.

3 DAVEY, V S Hydrogen Assisted Cracking of High Strength Steels in the Legs of Jack-UpRigsInternational Jack-Up Conference, City University, London,September 1991.

4 DET NORSKE VERITASCathodic Protection DesignDnV Recommended Practice RP B401, March 1986.

5 NATIONAL ASSOCIATION OF CORROSION ENGINEERSCorrosion Control of Steel Fixed Offshore Platforms Associated withPetroleum ProductionNACE Standard RP 0176-83.

6 BRITISH STANDARDS INSTITUTIONCode of Practice for Cathodic ProtectionBSI Code of Practice CP 1021: 1973.

7 COOPER, M T, THOMASON, W H AND VARDON, J D C Flame Sprayed Aluminium Coatings for Corrosion Control for the HuttonTension Leg ComponentsProceedings of Conference 'UK Corrosion 83', I Corr ST, Birmingham, 1983pp 23-17.

8 THOMASON, W H Materials Performance, Vol 24 No 3, 1984, pp 20-28.

9 ROSEBROOK, T, THOMASON, W H AND BYRD, J DA Review of the Performance of Flame Sprayed Aluminium Coatings Usedon Subsea ComponentsProceedings of Conference 'Corrosion 89', NACE, Paper No 624, Houston,1989.

10 ROSBROOK, TPrivate Communication

11 DEES, D DCombined Cathodic Protection of 5000 Series Aluminium Alloys and 15-5PM Stainless Steel in SeawaterProceedings of Conference 'Corrosion 87' NACE, Paper No 73, Houston,1987.

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12 TAFEL, JPhysik. Chem. 50A 641, 1905.

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APPENDIX: COMMENTS ON INTERIM GUIDANCEDOCUMENT

The draft interim guidance document Hydrogen Assisted Cracking of High StrengthSteels Immersed in Seawater (see Section 6.1) was presented for discussion at ameeting attended by owners, contractors, Certifying Authorities and governmentbodies on 4 December 1990. The chairman of the meeting stated that it was intendedto circulate the document to the industry in the near future, subject to modification bycomments from the meeting. He also referred to the guidance already provided bythe Department of Energy in Part 12 of the 4th edition of Offshore Installations: Guidance on Design, Construction and Certification. In addition, the publicationDesign and Operational Guidance on Cathodic Protection of Offshore Structures,Subsea Installations and Pipelines published by the Marine Technology DirectorateLimited dealt in some detail with the protection of high strength steels.

During the discussion the following points were raised:

� What was meant by high quality NDT? The NDT method used should bereliable in identifying cases of very fine cracking typical of that associatedwith hydrogen embrittlement. Although magnetic particle inspection hadbeen used exclusively to date, other methods based on the detection ofelectromagnetic fields were able to detect and characterise defects throughpaint coatings without the need to clean down to bare metal. The reliabilityof these techniques for use underwater had not been fully evaluated yet.Operators of this equipment needed to be specially trained and briefed whenexamining for hydrogen-related cracking.

� A request was made to establish a standard format for tests, which should becarried out at the design stage or as a result of inspections and surveys, todetermine the quality and resistance of materials to hydrogen attack.

� One Certifying Authority had prepared a note on the operation of one type ofjack-up rig where the legs may penetrate into mud so that the cathodicprotection system can be modified to cope with anaerobic conditions.Surveys (for hydrogen attack and other reasons) would need to include bracelevels 0-4 immediately above the spud cans.

� A strenuous effort must be made by all concerned to reduce the activity ofCP systems to acceptable limits, ie to ensure the voltage applied to the steeldoes not go more negative than -850 mV of -825 mV should tests indicatethat this is necessary. Owners were reminded that it is their obligation toensure that adequate CP voltage readings are undertaken for surveypurposes. This may mean that the owners would have to commission studieson how to measure the CP voltages (including inside spud cans), how toreduce them to adequate levels and how to avoid interference with the rig CPsystem from an adjacent fixed platform when the rig was performing a tenderoperation alongside another installation. Owners and Certifying Authoritieswere reminded that CP voltage measurements should be carried out on allrigs, not just on those believed to be susceptible to hydrogen cracking.

� Serious consideration should be given to replacing conventional anodes withdiode-controlled CP systems at least at the bottom of the legs and withinspud cans, ie in those areas where anaerobic conditions may exist for more

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than short periods. Parameters indicating correct operation of these systemsor their failure are quite definite and can easily be measured by conventionaltechniques, eg:

� open circuit at anode: -700 mV� correct operation: -825 mV� short circuit of diode or contact with another system: 950 mV or

more negative.

� Special surveys may be required from time to time and attention should begiven during these surveys to previous weld repairs in order to detectincipient cracking caused by the process of repair. Surveys need goodreporting methods to record and identify cracking. Owners and CertifyingAuthorities were reminded that proper inspections must be carried out toidentify this form of damage. Special surveys may not be necessary forexisting jack-ups which are not believed to be susceptible to this problem.

� Where new designs of jack-up rig were intended for construction it was theduty of Certifying Authorities and owners to ensure that the steels to be usedand the weld-induced heat affected zones were examined for strength,hardness and toughness levels as indicators of their susceptibility tohydrogen attack. Previously this had been carried out using hydrogensulphide testing but this is very difficult to control. Testing using a slowstrain rate with a range of CP voltages is more easily controlled and largenumbers of tests can be carried out in a relatively short time. TheDepartment of Energy believed that all steels in the high strength range (ieactual yield stress greater than 650 MPa) should be examined for thepossibility of hydrogen damage in service, both for the parent material andweldments.

� Where spud cans were protected by biocide and corrosion inhibitor, annualinspection programmes should include tests to ensure adequate performanceof the protection system.

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hydrogen cracking of legs - health and safety executive · hydrogen cracking of legs and spudcans...

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