subgroup c report
TRANSCRIPT
PIANC PTC II WORKING GROUP 28
Recommendationsfor the
Constructionof
Breakwaterswith
Vertical and Inclined Concrete Walls
Report of Sub-Group C
Final Report . . . July 1997Issued by the Sub-Group Chairman
to the Main Working Group Chairman, July 1997
Investigations into the implication of Construction aspects in Design
Performance of Concrete
Identification of “Hot Spots” in Design and Construction
00
MEMBERS OF THE SUB-GROUP
Mr. J. L. Diaz Rato, Ingenieros de Caminos, Canales y PuertosInfrastructure Department, Gijón Port Authority, Asturias, Spain.
Dr. O. Kiyomiya, PhD Civil EngineeringThe Port and Harbour Research Institute, Japan.
Mr. F. Ropert, ITPEService Technique Central des Ports Maritimes et des Voies Navigable, France.
Mr. B. N. Sharp, MSc, C.Eng, FICESir William Halcrow & Partners Ltd., London, UK
Professor Dr.-Ing. T. Stückrath, (Chairman)Technical University of Berlin, Germany.
Other Authorities consulted during preparation of the Recommendations
Mr. T. N. W. Akroyd, MSc Tech, LLB, C.Eng, FIStruct E, UKChairman, BS 8002 code drafting committee
Mr. P. Chubileau, CSPService Technique Central des Port Maritimes et des Voies Navigable, France
Dr. F. T. ChristensenApplied Science Associates, USA
Dr. T. A. Harrison, BSc, PhD, C.Eng, MICE, FICTBritish Ready Mixed Concrete Association, UK
Professor (Associate) D.Eng, P.A. HedarGothenburg, Sweden
Dr. B. Simpson, F.Eng, Eur Ing, MA, PhD, FICEOve Arup and Partners, Arup Geotechnics, UK
Mr. G. Leone, BSc, C.Eng, MICESir William Halcrow and Partners Ltd., UK
Mr. E. Longaygue, ITPEService Technique Central des Ports Maritimes et des Voies Navigable, France.
Dr.-Ing. J. SchwarzHamburg Ship Model Basin, Germany
Mr. D. Slater, BSc, C.Eng, MICE, MIStructESir William Halcrow and Partners Ltd., UK
Professor G. Somerville, PhD, C.EngBritish Cement Association, UK
Mr. D. Wimpenny, BSc, MPhilSir William Halcrow and Partners Ltd., UK
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Page
1 INTRODUCTION AND TERMS OFREFERENCE
1.1 Methodology for the Recommendations of Sub-Group C.... 1
1.2 General Headings .................................... 1
1.3 Format of Recommendations................... 1
2 DESIGN CRITERIA AND MATERIALS
2.1 Different Loadings not covered bySub-Group A
2.1.1 Earthquake ............................................... 22.1.2 Ice Pressure.............................................. 32.1.3 Ship Collision (deleted as insignificant).. 42.1.4 Earth Pressures ........................................ 42.1.5 Fill Pressures............................................ 82.1.6 Friction..................................................... 92.1.7 Handling and Float-Out Loads ................ 112.1.8 First Grounding........................................ 12
2.2 Resistance Analysis, Internal Analysis
2.2.1 Structural Analysis in Element Design.... 132.2.2 Scale Models............................................ 132.2.3 Limit State Design and Risk Analysis ..... 14
2.3 Durability of Concrete
2.3.1 Durability, Introduction ........................... 152.3.2 Design Working Life (or Service Life) ... 162.3.3 Processes of Deterioration ....................... 172.3.4 Exposure Classification ........................... 192.3.5 Influence of Cement Type ....................... 212.3.6 Influence of Cement Content................... 222.3.7 Cracking and the Influence of Cracks ..... 232.3.8 Influence of Curing.................................. 242.3.9 Monitoring and Maintenance................... 25
2.4 Materials
2.4.1 Rock and Rubble ..................................... 262.4.2 Filling and Backfilling............................. 262.4.3 Concrete Durability, General -
Design, Detailing and Workmanship....... 262.4.4 Unreinforced Concrete (Plain or Mass)... 262.4.5 Reinforced Concrete including
Selection of Cover to Reinforcement ...... 282.4.6 Prestressed Concrete................................ 302.4.7 Cement..................................................... 30
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Page
2.4.8 Aggregates ............................................... 312.4.9 Cracking and Crack Width ...................... 322.4.10 Reinforcing Steel ..................................... 322.4.11 Admixtures .............................................. 322.4.12 Additional Protective Measures:
Coatings, Coated Reinforcement,Cathodic Protection ............................. 32
2.4.13 Corrosion of Structural Steel ................... 33
3 CONSTRUCTION RELATEDCRITERIA AND CONSTRUCTIONMETHODS
3.1 Caissons
3.1.1 Float-Out Loading ................................... 343.1.2 First Grounding........................................ 343.1.3 Caisson Fill Methods and Pressures ........ 343.1.4 Sea Condition Data and Limits for
Construction Risk .................................... 343.1.5 Construction Joints .................................. 343.1.6 Settlement ................................................ 353.1.7 Early Thermal Cracking .......................... 373.1.8 Slipforming.............................................. 393.1.9 Curing ...................................................... 403.1.10 Developments in Caissons....................... 40
3.2 Blocks
3.2.1 Blocks from Concrete .............................. 423.2.2 Types of Concrete Blocks........................ 423.2.3 Common Problems .................................. 42
3.3 Rubble Mounds ...................................... 44
3.4 Curtain and Pile Type ........................... 44
4 SUMMARY
4.1 Different Loadings not covered by Sub-Group A............................................ 45
4.2 Resistance Analysis, Internal Analysis.... 464.3 Durability of Concrete ............................. 474.4 Materials .................................................. 494.5 Construction Related Criteria and
Methods - Caissons.................................. 514.6. Blocks ...................................................... 514.7 Curtain and Pile Type Breakwaters ......... 51
5 REFERENCES ...................................... 52
Contents
1.1 Methodology for the Recommend-ations of Sub-Group C
The task of Sub-Group C was drawn up at theMeeting of the Working Group in Hannover inFebruary 1993:-
“Investigations into the implication ofconstruction aspects in design. Performance ofconcrete. Identification of “hot spots” in designand construction”.
By consultation with the members of WorkingGroup 28 and, especially amongst the members ofSub-Group C, tasks and topics were assembledwhich Sub-Group C undertook to study. The Sub-Group decided that one of its tasks was to identifysignificant loading cases and items related to design,structural analysis and construction, which may notbe covered by the predominantly wave loading andanalysis considerations of the other Sub-Groups.The list was enlarged and changed several times.The main difficulty encountered in drawing up aframework for the topics was that a clear structurefor a consistent arrangement of all recommendationscould not be established. It is unavoidable thatcertain subjects have to be repeated, because anexclusive relation of one item to a single headingdoes not exist.
At the first meeting in Berlin in February 1994,the following basic methodology was agreed:-
• identify appropriate topics• study and prepare recommendations, and
discuss nationally with colleagues• compare guidance and codes nationally and
internationally• check if topics are already covered by exist-
ing documents• list the references to topics in a standard
alphabetical order form.At the three subsequent meetings of the
Sub-Group (held in Compiègne, France in July1994, Gijón, Spain in February 1995, and London,UK in July 1995) and by correspondence, the listsof topics were identified and initial reports wereprepared by each member on specific topics. It wasagreed which topics would be studied by whichmembers and, subsequently, these were developedinto several stages of draft text.
The final draft Sub-Group Report was completedin October 1995 and issued to the Sub-Groupmembers and, by the Chairman, to the MainWorking Group 28. After receipt of the furthercomments arising from the Sub-Group membersand comments of the main Working Group at themeetings held in Berlin, in March 1996, and in Delftin September 1996, the Sub-Group Report wasupdated for issue as a Final Sub-Group Report inJuly 1997.
The members of the Sub-Group and thoseexternal authorities who were consulted during thisperiod (either within the Working Group or outsidePIANC) are listed in the opening pages of thisSub-Report.
1.2 General HeadingsThe topics selected for consideration were as
follows:-DESIGN CRITERIA AND MATERIALS
• Different loadings not covered by Sub-GroupA
• Resistance analysis, internal analysis• Durability and maintenance• Materials.
CONSTRUCTION RELATED CRITERIAAND CONSTRUCTION METHODS
• Caissons• Blocks• Rubble mounds• Pile and Curtain type.
1.3 Format of RecommendationsThe recommendations of Sub-Group C, where
possible, consist of information mostly drawn fromthe experience and studies of its members, refer-ences, and brief summarised recommendations. Therecommendations are brief or given as a reference,except where it is considered that it may be morehelpful to give more detail. According to a decisionat the Meeting of the Main Working Group 28 onthe 26th of April 1995 in London, it will be left tothe Main Working Group to extract from thebroader formulated recommendations of the Sub-Groups, those statements which should be includedin the Recommendations of Working Group 28. Asummary of the main recommendations is given atthe end of the Report, as Section 4. In accordancewith the practice of PIANC, the Main Group Reportwill be circulated to members. Copies of Sub-Groupreports are available on application to PIANC.
1. INTRODUCTION AND TERMS OF REFERENCE
1
2.1 Different Loadings not covered bySub-Group A
2.1.1. EarthquakeThe closest understanding of the behaviour of
breakwater caissons during earthquakes can, as forstructures above ground, be obtained by dynamicresponse analysis. Due to advances in computertechniques, this can now be achieved by finiteelement methods. The caisson, rubble mound andsoil foundations and surrounding water should beconsidered in the modelling. This model has notbeen applied generally in design, except for veryimportant breakwater caissons located in a vigorousearthquake activity zone. The response of the modelis illustrated in Fig. 1.
Generally, however, the simple equivalent staticload method is appropriate for breakwater structuresbecause the seismic force has a relatively shortperiod of natural vibration, with heavy damping. Inaccordance with the Technical Standards for Portsand Harbour Facilities in Japan, 1991, the earth-quake load acting on a breakwater located in a seis-mic activity area should be calculated by the follow-ing formula:
• Earthquake load = (dead load + surcharge) x seismic coefficient.
The seismic coefficient is determined by takingthe regional probability of occurrence of an earth-quake, the condition of the foundation soil and theimportance of the structure into consideration.
As a general rule, it is considered that the earth-quake load acts horizontally at the centre of gravityof the structure. The vertical component of earth-quake load is not usually considered. The seismiccoefficient is obtained as follows:
Seismic coefficient =Regional seismic coefficient x factor for sub-soil condition x importance factor
In the Japanese Technical Standards, the factorsare as follows:
• regional seismic coefficient is 0.05, 0.10 or0.15
• factor for subsoil condition is 0.8, 1.0 or 1.2• importance factor is 0.5, 1.0, 1.2 or 1.5.These factors are classified in Tables 1, 2, & 3
on page 3.Usually, wave forces are larger than earthquake
forces and therefore earthquake forces can beneglected in the design, except in the case of verylarge breakwater caissons located in areas of vigor-ous seismic activity. The stability of the foundationsmust also be checked for earthquake loading. Evenin the major Hansin earthquake in Japan, in January1995, the breakwater caissons did not slide orcollapse. The most significant effect was settlementof the caissons, due to liquefaction of loose sand inthe foundation strata. Seismic forces should also betaken into consideration for piers, jetties, etc.
The vertical component is considered only for anearthquake with a narrow epicentre or for structureslocated near a fault. In this case, the verticalcomponent coefficient is usually adopted as half thehorizontal component.
2. DESIGN CRITERIA AND MATERIALS
2
m
b l o c k
armourstone
(a) Model
Fig. 1 Earthquake dynamic response model(ref The Port and Harbour Research Institute - Japan)
(b) Vibration Mode
Subsoil conditions are divided into three classifi-cations : 1st, 2nd and 3rd kind.
The classification of the subsoil condition andhence the subsoil condition factor depends on thethickness of the quaternary deposit and the types ofsubsoil condition, as given in Tables 1 & 2.
The factor relating to the importance of thestructure is classified in four categories: Specialclass, A class, B class and C class, as in Table 3.
2.1.2 Ice PressureAccording to Schwarz J, and Christensen F T,
there are still no formulae which can be recom-mended as being conclusive and reliable in thecalculation of ice pressures for all cases. Thefollowing guidelines should contribute a betterunderstanding of ice pressures for major works.
In countries with long frost periods, verticalbreakwaters should be designed to resist icepressure, which can arise from different causes:
a) pressure caused by the closed cover of ice ina harbour, when the cover expands withrising temperature
b) pressure caused by ice fields drifting parallelto the coast with tidal current or littoral drift
c) vertical pressures caused by piling upprocesses.
As a rule, ice pressure is not considered for slop-ing breakwaters in Europe (Hedar P A, 1995) but isrecommended for America (Christensen F T et al,1995, Bruun P, 1985).
For vertical breakwaters, assessment of icepressure is a matter of judgement depending on theice conditions and the possibilities of ice-structureinteraction. It is doubtful if the pressures given forconcrete dams for fresh water lakes, such as 300 kNper m2 and 50 to 200 kN per metre length accord-ing to Hedar P A, 1995, or typical pressures of 200to 400 kN per metre with extremes of 400 to 600 kNper metre according to Christensen F T, 1996, canbe applied to vertical breakwaters in sea water
Small scale experiments on ice forces on verticalstructures carried out at the Iowa Institute forHydraulic Research have resulted in the followingformula (ref Schwarz J, 1994).
σeff = 0.564 . d-0.5 . h0.1 . σc
whereσeff = effective indentation strength in MPad = structure width in mh = ice thickness in mσc = maximal compressive strength of an ice
prism at a strain rate of 0.003 persecond in MPa. (For a prism length of0.2 m the deformation speed must be0.0006 m/s). For instance for the BalticSea, a value of σc = 1.8 MPa is feasibleaccording to EAU 1990.
According to this formula the effective icepressure decreases with the square root of thestructure width. This finding is in accordance withSanderson’s investigations (ref. Sanderson T J O,1988) which however do not consider the effect ofthe ice thickness, as illustrated in Fig. 2 (overleaf).
The Iowa formula (Schwarz J, 1994), which stillhas to be proved for larger structures, has receivedsupport from full scale measurements in China andin the Baltic as well as recently by Canadianengineers (Fitzpatrick J, 1994 and Kennedy K, 1996).
3
Thickness of Gravel Sand or Clay Soft Groundquaternary deposit
Less than 5m 1st kind 1st kind 2nd kind
5-25m 1st kind 2nd kind 3rd kind
More than 25m 2nd kind 3rd kind 3rd kind
TABLE 1CLASSIFICATION OF SUBSOIL CONDITION
(ref Technical Standards for Ports and Harbour Facilities in Japan)
Classification Factor
1st kind 0.8
2nd kind 1.0
3rd kind 1.2
TABLE 2SUBSOIL CONDITION FACTOR
(ref Technical Standards for Port and Harbour Facilities in Japan)
Classification Characteristics of Structure Importanceof Structure Factor
Special Class
A Class
B Class
C Class
The structure has significant char-acteristics covered by items (1)-(3)in A Class.
(1) If the structure is damaged byan earthquake, a large number ofhuman lives and property will pos-sibly be lost.(2) The structure will perform animportant role in the reconstructionwork of the region after an earth-quake.(3) The structure handles haz-ardous or dangerous activities withrisk that the damage to the struc-ture will cause a great loss ofhuman life or property.(4) If the structure is damaged,economical and social activity ofthe region will suffer severely.(5) If the structure is damaged,repair work will be difficult.
The structure is other than Special,A or C Class.
The structure is small and easy torepair, excluding structures whichfall in the Special or A Classes.
1.5
1.2
1.0
0.5
TABLE 3CLASSIFICATION OF IMPORTANCE FOR EARTHQUAKE
(ref Technical Standards for Port and Harbour Facilities in Japan)
4
2.1.3 Ship Collision (deleted)The effects are not significant, and only relate to
small vessels.
2.1.4 Earth Pressures
This section concerns the principles of earthpressure in general and external to structures such ascaissons (see Fig. 5). Earth pressure in filling tocaissons is considered in 2.1.5. The general casemust be considered first, although a breakwater willhave fill placed against it only if it is protected by arubble mound on the exposed side, or it retainsreclamation on the lee side. The water pressures tobe used in conjunction with earth pressures will, inthe case of breakwaters, be a function of wave andtidal variation loading, although on the lee side andinternally they may only relate to tidal effect. Waveloading is not considered here. The general
principles applying to struc-tural design are discussed in2.2.
Traditional “workingstress” codes recommend"active" or "at-rest" pressurecoefficients to apply to thedry or submerged soil mass,as appropriate, as a working(or “characteristic”) load. The“fully active” coefficient isusually applied to blockworkwalls, where slight rotationalmovement can be tolerated,and can be applied for theoverall stability of caissonsas, for example, in Spain.However, the “at-rest” co-efficient is usually applied to
calculate the overall stability of rigid structures suchas caissons and to calculate the sizes of the structuralmembers. If the fill is to be compacted aggressively,it could be necessary to consider compactionpressures, i.e. pressures greater than "at rest". This“working load” is applied in conjunction withsuitable factors of safety for overall static(equilibrium) stability calculations, and to deriveworking structural material stresses. “Passive”pressures are calculated at the ultimate, and dividedby a relevant factor of safety. The Japanese code (refTechnical Standards for Port and Harbour Facilitiesin Japan, 1991) still continues this approach. TheGerman Waterfront Structures Code (ref EAU 1990)has now introduced the concepts of limit states inaccordance with draft Eurocode 7 while still retain-ing traditional methods as a permitted option. Dueto the difficulties of applying limit states tofoundation problems, traditional methods remain anoption in other new codes, such as BS 8002 (ref BS8002, 1994).
Most structural codes now adopt the limit statephilosophy, although the application of limit statesto earth pressure and variable water loading formaritime structures is not as straightforward as forstandard structural loading for buildings andbridges. The selection of partial factors to matchwith and correspond to the reliability and failureprobability of wave and water loading is the task ofSub-Groups A & D. See also 2.2.3.
One of the main difficulties in bringing togetherearth pressure theory and limit state design ofstructural members is the fact that earth pressures inretained materials reduce as movement increases. Inthis respect, earth pressures differ from almost allother types of loads considered in structural design.In addition, the failure mechanisms and pressuresare different for flexible steel sheet piling and rigidreinforced concrete or masonry walls. The problemis that, in geotechnical terms, the soil load at ultimate(failure) is often less than at working conditions,whereas for structural design of the reinforced
Fig. 2 Effective ice pressure versus pressure area (ref Sanderson T J O, 1988)
IND
EN
TATI
ON
PR
ES
SU
RE
MP
a
Then and NowBeaufort global and local ice loads
Load (tonnes) Pressure (MPs)
1 500 000
100 000
1980 1985 1990 1995
Local
Global
15
0
Fig. 3 Development of ice force predictions(ref Fitzpatrick J, 1994)
8
Load (tonnes) Pressure (MPa)
These last publications conclude that ice pres-sures for large works have been overestimated in thepast by a considerable factor, as illustrated in Fig. 3.According to Schwarz J, 1996, and Christensen F T,1996, a validation study is planned in an EUresearch project.
5
The serviceability limit state earth pressures areusually calculated directly from the characteristicdensity and strength properties and the appropriateearth pressure coefficient. The ultimate limit statepressures are usually calculated by modifyingseveral of the main parameters involved from theirexpected values, by the application of partial factors.Sometimes the partial factor is applied directly tothe serviceability limit state pressure, and some-times it is applied to a parameter, such as tan Ø'. Insome current codes, for example, such as BritishStandard 8110 : 1985, and the Spanish MaritimeWorks Recommendations ROM 0.2-90 and ROM0.5-94, the ultimate limit state pressures forstructural design are calculated by applying partialfactors to the serviceability limit state pressures.However, for many years the Danish and otherScandinavian codes adopted the less conservativeprinciple of applying a partial factor to tan Ø' thusderiving a smaller value of Ø' and hence a largervalue of the applied earth pressure, together withalternative factors on the water pressure.
This latter approach to the ultimate limit stateearth pressure for structural design can produce amuch smaller difference between the serviceabilityand ultimate limit states than the classical workingstress approach, or if the partial factors fromstructural codes are applied to the characteristic (i.e.working or serviceability) pressures. Note that thepartial factor for material strength will also beapplied, and that the total ULS using direct factorsfrom structural codes has been calibrated to give
similar results to the old working stress codes. SeeFig. 4 (ref Daniels R J and Sharp B N, 1979).However, some codes achieve a similar reductionby the use of much lower partial factors applied tothe serviceability limit state than are used in othercurrent structural codes. See below and Fig. 6(d)and 6(g).
British Standard 8002 : 1994, Code of Practicefor Earth Retaining Structures, incorporating thelimit state approach, has only recently been issued.BS 8002 adopts limit state philosophy but it doesnot, in fact, involve partial factors. It reduces tan Ø'by a “mobilisation” factor which operates in a similarway to a partial factor and can be imagined as apartial factor. However, BS 8002 does not deal withmaritime structures (ref Akroyd T N W, 1996 andBolton M D, 1996). Alternative limit state methodsare incorporated in Eurocode 7 : Part 1:1993.Eurocode 7 has two appropriate cases, B and C.Case B derives straight from structural design, andCase C comes from geotechnical stability analysis.
A comparison of various applications of partialfactor methods for the calculation of structuralmembers is given in Figs. 5 and 6. A breakwatercaisson with sand fill on the inside (land side) isillustrated in Fig. 5. Typical lateral pressurediagrams are given for the loading on the buriedface, i.e. the components of soil above water level,surcharge, submerged soil and water. To illustratethe comparison of methods in Fig. 6, the cases ofsubmerged soil and water loading with water levelat a MSL of +3m, only, are considered, all asunfavourable loads. Tidal variation of water level,wave loading and surcharge are not considered here
Fig. 4Caisson walls : Design wall
pressure distribution(Loading on one side only -i.e.
all loads have been calculated asunfavourable loads on the loaded
side and air on the other side.)
(Adapted from: Daniels R J andSharp B N, 1979)
+4.5m Dock deck level
+3.0m WL
Fill and water levelsto same vertical scale
Ultimate limit state using factors of 1.15 on tan∅1 and 1.2 on water
Serviceability limit state
Ultimate limit state using factors of 1.5 onsurcharge and 1.35 on soil and water asEurocode 7 Case B i.e. similar to working
stress results excluding the material factor.
+4
+2
0
-2
-4
-6
-8
-10
-12
-13.290 50 100 150 200 250 300 350 400
effective factor approx. 1.2effective factor approx. 1.35
Surcharge
Horizontal wall pressure: kN/m2
Red
uced
leve
l : m
concrete section the designer seeks to apply afactored higher loading for the ultimate case.
but the appropriate worst case situations of waterlevel will have to be considered in a specific design.
Case (a) in Fig. 6 shows the “at-rest” character-istic pressure (serviceability limit state) for sub-merged sand and water for a 20m depth of fill, forthe soil properties given in Fig. 5. Only granularmaterial is considered.
Cases (b), (c) and (d) show the ultimate limitstate pressures for structural design calculated bythe direct application of partial factors to case (a). InCase (b) to the British structural code BS 8110 :1985, the submerged sand pressure is multiplied by1.4 and water, as an adverse (i.e. not a reducing)load, is also multiplied by 1.4. The factor forvariable loads is 1.6. Case (c) shows the recommen-dations of the Spanish ROM 0.2-90 1990, whichuses a partial factor of 1.35. However, variablewater loads based on statistical data can be factoredby 1.0, and other variable loads by 1.5. Similar rulesapply to Eurocode 7, Case B. Case (d) shows asimilar USA application noted in the PIANC reporton floating breakwaters (ref PIANC, 1994), wherethe standard partial factor is 1.2. In this application,design wave loading can be used with a factor of1.3, or 1.6 on a yearly return period.
Cases (e) and (f) give examples of applying thepartial factor to tan Ø'. In (e), to earlier Scandinav-ian rules, the worst of two cases is taken, i.e. afactor on tan Ø' of 1.3 combined with water(adverse) at 1.0, or a factor on tan Ø' of 1.15 pluswater (adverse) at 1.2. In (f), the method of BS8002:1994 is shown with a “mobilisation” factor(nb not a “partial” factor) on tan Ø' of 1.2. However,BS 8002 does not appear to give adequate guidance
6
on appropriate factors for water pressure forstructural design. A similar result is given by CaseC of Eurocode 7, with a factor on tan Ø' of 1.25with, again, inadequate guidance on water pressuresto apply to structural elements in conjunction withthis soil loading. An answer to this serious ommis-sion has been suggested by Cole & Watt (ref Cole &Watt, 1994). They suggest that the earth pressurecalculated from the reduced tan Ø' should be treatedas a “worst credible load” to BS 8110 : Part 2, 1985,and therefore both this and the water pressure bemultiplied by a factor not less that 1.1 or 1.2 forstructural design of the section. This suggestion isrefuted by Akroyd (ref Akroyd T N W, 1996) and, incomparison with other “limit state” codes, it wouldindeed appear incorrect to apply a further factor tothe earth pressure component. However, the suggest-ion to apply a factor of 1.2 to the water componentat least produces a comparable philosophy to someother codes, and is illustrated in Case (f). Case (g)illustrates proposed Japanese factors (ref KiyomiyaO, 1994) which are similar to cases (b), (c) and (d)but with a lower general factor of 1.1, and a higherKo of 0.6. The factor for variable loads is 1.3.
It can be noted from the above that higher, moretraditional, ratios of ULS loads to serviceabilitylimit state loads for the calculation of member sizesare given by structural codes such as BS 8110,Eurocode 7 Case B, and ROM 0.2-90, althoughROM 0.2-90 has the facility for varying environ-mental load factors. The less-conservative loadingfrom the USA, earlier Scandinavian, BS 8002 andEurocode 7 Case C, and Japanese methods, are allof a similar magnitude.
Breakwater detailsas relevant
Sea andWave side(Front side)+ 3m msl
Land side
Reference water level for examples only.Different water level cases will be necessary
for design
Soilabovewaterlevel
Surcharge(Not
consideredinFig.6)
Submergedsoil.
Water
SOIL PROPERTIESCharacteristic ∅' above water 33°
∅' submerged 30°Density γ submerged 9.5kN/m3
γ sea water 10.2kN/m3
Characteristic Ko (At Rest) 1-sin∅'N.B. Soil pressures and water in soil only considered on one side,as unfavourable loads. Not wave loading.
Fig. 5 Breakwater caisson with sand fill on the inside (land-side)
- 17m
Fill Fill
7
SubmergedSandKo = 0.500Wall friction O°
95 kN/m2 204 kN/m2 299 kN/m2
(a) At-Rest characteristic pressures of submerged sand and water only. Serviceability limit state.
Water Soil and Water
SubmergedSand(a) x 1.4
133 kN/m2 286 kN/m2 419 kN/m2(b) Ultimate limit state to BS8110 structural factors. (Variable loads 1.6)
Water(Adverse)(a) x 1.4
(a) x 1.4
(a) x 1.35
128 kN/m2 275 kN/m2 403 kN/m2
(c) ULS to Spanish ROM. Permanent loads 1.35 (n.b. Variable loads 1.5) and Eurocode 7, Case B.(n.b. Variable environmental loads based on statistical data 1.0, other variable loads 1.5)
(a) x 1.35 (a) x 1.35
(a) x 1.2
114 kN/m2 245 kN/m2 359 kN/m2
(d) ULS to USA (Ref in text) 1.2 (DL & LL) + 1.3 Design Wave.
(a) x 1.2 (a) x 1.2
Case (i) Factor on tan ∅' 1.3Ko = 0.594 plus water at 1.0
Case (ii) Factor on tan ∅' 1.15Ko = 0.551 plus water at 1.2
105 113 kN/m2 204 245 kN/m2 317 350 kN/m2(e) ULS to Scandinavian factors. Case (ii) rules here
Case (i) water Total case (ii)
BS 8002.Mobilisation Factor on tan ∅' 1.2Ko = 0.566
Eurocode 7 Case C Factor on tan ∅' 1.25Ko = 0.581
108 110 kN/m2 245 kN/m2 353 355 kN/m2(f) ULS to BS 8002 and Eurocode 7 Case C.
Water to BS 8002??????
Water to Eurocode 7Case C ??????
Total Eurocode Case CSoil
+ waterat 1.2
Total case (i)
(a) x 1.1with Ko = 0.6
114 kN/m2 224 kN/m2 338 kN/m2(g) ULS. Proposed Japanese Factors. (Variable water pressure 1.3)
Fig. 6 Comparison of earth pressure calculations(Note: loads considered on one side only, all as unfavourable (unvariable) loads, to illustrate the principles. In practice there may be soil and/or water loads on
both sides with combinations as favourable or unfavourable and variable loads. See 2.2.3.)
Total
+ 3 m
- 1 7 m
Case (ii) water
Water at 1.2
Total BS 8002Soil +Water at 1.2
(a) x 1.1
2.1.5 Fill Pressures
The loading within caisson cells is generallyderived from silo theory. Practical verification ofthis approach is available from measurements incaisson cells in 1976/77 (ref Daniels R J and SharpB N, 1979) which demonstrated that the Janssen silodistribution of pressure with a wall friction of 20°compared closely with the measured pressures, Fig. 7.
The various methods for calculating internalpressures are illustrated in Figure 8. As in 2.1.4,submerged sand to 20m depth is used to illustratethe principles. Case (a) shows the characteristicat-rest unconfined pressure, for the materials givenin Fig. 5.
Case (b) shows calculations to DIN 1055 Part6:1987, ROM 0.2-90 and ACI 313:1983, all ofwhich are based on the Janssen method, applied inthe filling condition. Eurocode 7 refers , for silopressures, to Eurocode 1 Part 4, which is presum-ably based on the same method. ACI 313 adopts theactive pressure coefficient (0.333 in this case) andoverpressure factor. However, all of these methodsgive the same result with an at-rest coefficient of 0.5.
As shown in case (c), the Technical Standardsfor Port and Harbour Facilities in Japan, 1991, rec-ommend a fixed at-rest earth pressure coefficient of0.6, which is applied to a depth equal to the width ofthe cell. Below that 45° line, the pressure is constanton a silo basis. This calculation applies up to celldimensions of 5m x 5m, above which standard silopressures are applied. The results are simpler andnot greatly dissimilar from the others in case (b).
8
Fig. 7Caisson walls.
Comparison between measured and predictedwall pressures within a caisson cell
(ref Daniels R J andSharp B N, 1979)
+4
+2
0
-2
-4
-6
-8
-10
-12
0 20 40 60 80 100 120 140 1 6 0Horizontal wall pressure : kNm2
Caisson floor
EffectiveJanssenDistributionδ = 20°
Total wall pressure(measured)
Effectiveearth pressure(measured)
+ 1.5 m Fill level(partially filled)
+1.0 m WL
+4.5 m Final pavement level
Red
uced
leve
l : m
EffectiveRankineDistributionδ = 0°
Fill and water levels tosame vertical scale
Difference= water pressure
Note: Total wall pressures were measuredwith pressure cells. Effective earth pressure = total measuredpressure minus static water pressure
Water pressure is an important loading case tobe applied to external water pressure for floatingcaissons, and to variable water levels in the soil asaffected by wave and tidal effects. Different factorsneed to be applied to the water pressure, dependingupon whether the water pressure is considered to beassisting or resisting the earth pressure. As waterpressure is much larger than soil pressure in the caseof granular materials, there would seem to be littlepoint in over-refinement of the earth pressurefraction of the total load, unless much clearerguidance and rational partial factors are also givento variable water loading.
As a general recommendation, results such asgiven by the USA method with a general factor of1.2, the Japanese method, or by factors on tan Ø'such as the Scandinavian method, BS 8002 andEurocode 7 Case C, provided a factor of at least 1.2is applied to water pressure, appear to be similar andappropriate for the analysis of both overall stabilityand member strength. Some authorities, however,would recommend the higher, conventional, struct-ural factors to be used for member strength. TheHong Kong Geoguide 1 (ref Hong Kong Guide toretaining wall design) advocates this course. It givespartial factors similar to BS 8002 for soil equilibrium,but advises that structural members be calculatedusing unfactored soil properties in conjunction withthe relevant structural code. The Hong Kong Guidedoes not apply to maritime structures. See also 2.2.3.
Please refer to 2.1.5 for soil loading within cells,and to 2.2 and 2.4.2 for loading cases, materials andfilling methods.
9
The loading on the base plate of caissonsdepends on the form of wall pressure assumption.When a silo theory is used, the base plate will berelieved compared to that of a simple earth pressureat rest assumption, because the fill is hanging on thewalls. The soil reactions on the base plate can alsobe calculated to the Danish Standard DS 415. TheTechnical Standards for Port and Harbour Facilitiesin Japan 1990 give clear guidance on the design load-ings for walls and base plate, as does ROM 0.2-90.
There are two schools of thought in Spain. Theconservative approach adopts the pressure on thefoundation in service, after deducting the mass ofthe submerged base slab itself, and the thrust at thebottom of the fill, taking into account the silo effect.It is, however, considered more realistic to calculatethe pressure on the cell bottoms as equivalent to thatextended by the whole mass of caisson and filling,supposing that just one cell remains unfilled.
The above characteristic pressures can be usedfor the serviceability limit state or for traditionaldesign by “working stress” methods. The considera-tions affecting the choice of partial factor to applyto derive the ultimate limit state pressures, and thevarious alternatives available are the same as notedabove in 2.1.4, and in 2.2.3. Again a factor of theorder of 1.2 appears to be generally appropriate.
The density achieved for sand poured into water- filled caisson cells is low, much lower than forsand poured into open water. If it is required to limitsettlement of the fill within the caisson cells, or
+3
0
-3
-6
-13
-17
14
29
43
76
95 kN/m2
∅ ' Submerged 30°
Wall Friction 0°
γ Submerged 9.5 kN/m3
γ Sea water 10.2 kN/m3
Ko = 0.5
+3
- 17
(a) Submerged Sand. At-Rest Unconfined Pressure
+3
0
-3
-6
-13
-17
12
20
25
32
34
8 13
15 24
20 34
28 51
30 56
ACI 313 with Ka = 0.333
DIN 1055 Pt6, Spanish ROM, andACI 313 with K0 = 0.5. (Janssen)
ACI 313 ditto with Overpressure Factor
Wall Friction 17° Cells 4.8m square
(b) Submerged Sand. Silo Pressure (Janssen) to DIN, ROM 0.2-90 and ACI 313
4.8m
kN/m2
0
-3
-6
-13
-17
+3
-1.827
27(c) Submerged Sand Silo Pressure. Japanese Technical Standards
(applicable up to cell dimensions 5m x 5m)
Fig. 8 Comparison of Silo Fill Earth Pressure Calculations with External Earth Pressures
2.1.6 Friction
Although the concepts of friction are classicaland have been studied for so long, there is a surpris-ing divergence in the figures used in design, andlack of agreed experimental data. There is a widevariation in the coefficients of friction assumed, andthe factors of safety against sliding although thesedifferences may tend to cancel out in the resultingstability equation. In practice, the friction co-efficient and factor of safety must be related to thepermitted displacement.
Various coefficients of friction are compared inTable 4.
The Japanese values have been determined frommodel tests (see Table 5). The friction coefficientdepends on the size, configuration, strength andkind of stone, and the compacted condition, and isrelated to the coefficient of internal friction of themound. The coefficient is low at initial construction,but increases in time after compaction by stormsand self- weight consolidation. Weak rocks lead to alow value of µ.
In Japan, the friction resistance is increased byplacing a rubber mat or an asphalt mat under the baseof the caisson. This measure enables the size of thecaisson to be reduced. Other measures to increase thefriction coefficient include corrugations or dowelsin the base slab. The depth of the corrugations are
t
(a) Submerged sand. At-rest unconfined pressure
(b) Submerged sand. Silo pressure (Janssen) to DIN, ROM 0.2-90 and ACI 313
(c) Submerged sand. Silo pressure. Japanese Technical Standards(applicable up to cell dimensions 5m x 5m)
Fig. 8 Comparison of silo fill earth pressure calculations withexternal earth pressures
increase the density for reduction of liquefactionrisk in an earthquake, this can be achieved by vibra-tion (ref Daniels R J and Sharp B N, 1979/CochraneG H, Chetwin D J L and Hogbin W, 1979). Due tothe inverse relationship between density and Ø',densification causes no significant increase in wallpressure. Densification is usually only necessary ifthe fill is used to support surface works.
There is conflicting guidance as to whether thereduction of earth pressure by silo effect applies ifthe caisson is vibrated by earthquake shock or waveloading. According to Japanese experience, thefrequency of earthquake vibrations does not destroythe silo effect. The problem should be reduced byusing appropriate fill materials and achievingappropriate fill density to preclude earthquake lique-faction. The density, in place, can be verified byDutch cone or similar site investigation methods.
The load cases for the design of the cells and theinter-cell walls must reflect the worst cases whichcan apply in practice. For example, the maximumwater level case with water level at the surface mustbe considered. Also, the urgency with which thecells have to be filled when a caisson is placed in anexposed breakwater precludes the use of anyprocedure which involves special precautions, suchas restrictions on the level to which adjacent cellsare filled, or emptying the cells of water. Thecaisson must be designed to allow any sequence andproportion of filling.
-17t
some two-times the dimension of the stone size atthe surface of the rubble mound.
Experimental results for friction coefficients inJapan are given in Table 5. The coefficients dependupon the factors related to the rock type and dimen-sions noted above and to the amount of displace-ment, weight of caisson etc.
In comparison, a recent series of French tests(ref CÉTÉ - Laboratoire Régional Nord - Pas deCalais) are given in Table 6. The parameters inves-tigated included :
• two forms of bottom slab -smooth and corrugated
• two types of gravels -crushed gravel 0 - 80 mm andnatural sea gravel 20 - 80 mm
• varying vertical loadThese results were considered to be lower than
used in practice, and were accompanied by signifi-cant displacement, up to 180 mm, before maximummobilisation.
10
TABLE 5EXPERIMENTAL TEST RESULTS ON FRICTION COEFFICIENT (JAPAN)
Precast concrete against stone
No. Kind of Stone µ Average of Dimension Conditionµ of stone. mm of mound
1 crushed stone 0.460 - 0.801 - 30 screeded surface
rubble stone 0.564 - 0.679 0.624 120 not screeded
2 rubble stone 0.45 - 0.69 - 50 Surface “blinded” (i.e. smoothed)
with smaller stone
3 crushed stone 0.77 - 0.89 0.82 30 - 80 screeded
cobble stone 0.69 - 0.75 0.70 30 - 50 not screeded
4 crushed stone 0.607 - 0.790 0.725 20 - 30 not screeded
5 crushed stone 0.486 - 0.591 0.540 10 - 50 “
6 crushed stone 0.41 - 0.56 - 13 - 30 not uniform
µ = PmaxW
block
W P (Pull)
Stone
TABLE 6FRENCH RESULTS FOR FRICTION COEFFICIENT
Vertical Normalload. TON Stress T/m2
Horizontal Force. TON Friction Coefficient µ
Smooth Corrugated Smooth Corrugated
Natural Sea Gravel 20-80 mm24.1 10.5 12.6 13.7 0.53 0.5818.4 8 10.3 11.3 0.56 0.62
Crushed gravels 0-80 mm24.1 10.5 - 10.4 - 0.4318.4 8 - 8.6 - 0.47
P (Pull force)
δ (displacement)
Coefficient of Friction µ
UK and FranceSpain Germany (Fasci-(ROM (BS6349 cule0.5-94) BS8002 No .62,
EAU 90) titre V)
Condition
Precast concreteagainst bedrockor concretePrecast concreteagainst precastconcretePrecast concreteagainst rubblePrecast corrugat-ed or sloping baseagainst rubblePrecast concreteon a rubber mator asphalt matagainst rubbleIn-situ concrete against rubble
TABLE 4COMPARISON OF DESIGN VALUES FOR
COEFFICIENT OF FRICTION
0.5 -
- 0.7
0.6 0.7 δ = 2/3∅r tan∅'(often0.58)
- - δ = ∅r
0.7 -
- 1.0
Japan(TechnicalStandards
for Port andHarbour
Facilities)
2.1.7 Handling and Float-Out Loads
From the viewpoint of limit states, all handlingand float-out loads belong to transient load situa-tions. The partial factor for these loads can be keptlow; a figure of γF = 1.1 is suggested. The loads canarise from different reasons, for instance from:
a) movement of caissons or blocks after dis-mantling the formwork and during launching,
b) storage on supports during the curing periodor in stacks of blocks and before float-out,
c) loads from handling with cranes during thecuring period or tugs during placement.
Loads which can arise during construction mustbe included in the stress analysis. If the handlingprocess which will be used during construction bythe contractor cannot be anticipated by the designer,the stress analysis has to be rechecked by thedesigner as soon as the construction method isknown.
Float-out loads must never be considered asloads at the contractor’s own risk, for which thedesigner is not responsible. Damage arising fromfloat-out processes will not always become apparentduring the construction period.
The wave climate and the sea condition limitswhich can be tolerated during float-out activitieshave to be specified by the designer and have to beknown by the contractor. These waves could becalculated by using Recommendation 3.1.4.
If there are certain climatic periods during theyear for which it can be guaranteed that high wavesdo not occur, float-out activity could be limited tothese seasons. The relatively short-float-out processfor a breakwater can very often rule the design andthe work on site. Experience shows that in mostplaces the times for float-out can only be fixed fromday to day by the use of a local weather forecast.
In harbour construction a lot of damage has beencaused by wrong or incautious handling of concreteelements during the construction period (ref BruunP, 1985). Surveys have shown that one of thereasons for the destruction of many rubble moundbreakwaters is the fact that the concrete elements of
the outer cover for the breakwaters were alreadybroken during construction and placement. Ananalogous situation can arise for the elements ofvertical breakwaters. If damage arises, it will mainlyoccur under water and, even by good monitoringduring the construction period, this will not alwaysbecome obvious. There is a rule which has becomecompulsory for dam construction: all loads, waterlevel changes and other factors that influence thestability of a dam also have to be anticipated for theperiod of construction and not only from the time atwhich the structure has been handed over. This ruleshould also be applied to the design of harbourstructures.
The Technical Standards for Port and HarbourFacilities in Japan (1991) give the followingadditional recommendations for the calculation ofexternal loads during float out.
a) the water pressure should be calculated withan additional draft of 1 m,
b) the tractive force during towing should be cal-culated by using the formula (n.b. expressedhere in SI units):
T = ½ ρwCDv2A
whereT = tractive forceCD = drag coefficientv = towing speedA = submerged area of the leading wallρw = density of sea waterDynamic pressures and wave pressures are not
considered but a rise of the water level, δ in front ofthe caisson has to be considered which is shown onFig. 9. A head δ, of 1m has to be applied, accordingto the Japanese Technical Standards, 1991.
It must be ensured that a positive freeboardexists at all times. The distance between the centreof gravity and the metacentre, the metacentricheight, should be at least 5% of the draught of thecaisson.
11
Fig. 9Tractive force during towing and hydrostatic head
between different cells.(ref Technical Standards for Port and Harbour Facilities in
Japan)
Various values of the minimum Factor of Safetyagainst sliding for use with the friction coefficientsare given below :
Japan (ref Technical Standards for Port and Harbour Facilities) 1.2Spain (ref ROM 0.5- 94)
Permanent situation 1.5Momentary situation 1.3
UK (ref BS 6349 Pt. 2 - 1988) 1.75Germany (ref EAU 90):• using active earth pressure and wall friction
2/3 ∅' 1.5• using at-rest earth pressure 1.0.
2.1.8 First groundingSpecial loads will arise when the lowered ele-
ments or blocks first touch the ground. In mostcases the caissons will never again undergo a com-parable distribution of load. The loads are excep-tional, mainly for the following reasons:
a) It is not always known in advance which partof the floating element will touch the groundfirst
b) It is not always known on which supportsthe element will rest after the first placement
c) The impact load is a dynamic load whichcannot be determined exactly, even if thegrounding velocity and the elastic and plasticproperties of the ground are known.
The German EAU 1990 (Recommendations ofthe Committee for Waterfront Structures) states in10.5.2 that a first landing on one point in the middleor on two points on two edges has to be anticipatedas indicated in Fig. 10.
Fig. 10View under a caisson, floated out. The caisson could land on
point A or on the opposite corners B1 and B2.
To prevent high stresses arising from the firstgrounding, the caissons for the closing of theBrouwerhavensche Gat in Holland have beenprovided with “camel humps” (Fig. 11) and on othersites downstand legs have been provided. (Fig. 12).
The humps or legs have to protrude at least morethan double the tolerance of the ground levelling.
In Japan the rubble is sometimes not trimmed toa horizontal formation, but is provided with anexcess height in the middle (extra banking) toenforce a first contact between caisson and groundin the middle of the bottom slab.
The settling velocity of the caisson before thefirst contact is related to the motion of the seaaround the caisson. Even if massive float-out equip-ment is used (large cranes) a certain movement byswell during placing cannot be ruled out. Repeated
hitting of the ground must also be anticipated.During the sinking operation, sometimes only arelatively small volume of water has to be filled-inbut, as the inherent mass and volume of the elementis considerable, this leads to two effects:
• Due to high mass element of the impact force(Force = mass x acceleration) there is a highimpact force even though the acceleration islow.
• The large volume or surface area leads tolarge variations in the displacement of theelement for only small changes in watersurface level.
Both of these effects can lead to impact loadswhich are higher than any other load in the lifetimeof the caisson.
12
18.0 m
16.2
m
58.0 m
Fig. 11 "Camel Humps” of the caissons used for the
Brouwerhavensche Gat, Holland
Fig. 12Legs which provide a first landing on three points.
B1
A
B2
2.2 Resistance Analysis, Internal Analysis
2.2.1 Structural Analysis in Element Design
This section applies particularly to structuresbuilt up of several elements, such as caissons, forwhich extensive analysis is required for the structur-al design of individual elements of the structure.
Structural modelling of the actions on caissonscan be carried out in two ways.
(i) In the simpler and traditional approach, thestructure is split up into sets of beams andslabs and calculations carried out by tradition-al manual methods and two-dimensionalframe analysis, using computers whereapplicable. Guidance for the traditionaldesign of caissons is given in national codes,such as the Spanish ROM 0.2-90, TechnicalStandards for Ports and Harbour Facilities inJapan, EAU 90 etc. It must be stressed that,in such methods, calculations are carried outon each member separately and thereforerequire assumptions to be made about theconnections between members. Theseassumptions can lead to questionable approx-imations as to the boundary conditions intro-duced in the calculations.
(ii) The increase in computer power has made itfeasible to carry out a full three-dimensionalmodel analysis, using finite elementmethods. Nevertheless for simple geometriessuch as for caissons, it is not essential to usevolume elements. A Japanese example isillustrated in Figs 28 and 29. Generally, thickshell elements will suffice for engineeringpurposes, while volume elements may behelpful for specific local analyses, forinstance to map the local stresses and strainsaround a fastening point (e.g. in the loadingcase of towing).
For most engineering problems, static modelsare the most familiar and most widely employed.Even if the phenomenon is dynamic in essence(earthquake, waves, wind, etc.) the loads applied aregenerally equivalent static loads.
Two types of models can be differentiated.• first order models, based upon the initial
structure geometry• second order models, based upon the strained
structure geometry.First order models are satisfactory in most cases,
although it is necessary to use second order modelsfor slender members that are liable to second orderinstability (buckling, warping, etc.). In this respectEurocode 1: (ENV 1991-1994) specifies that theeffects of strains and displacements should beconsidered if they result in an increase of the effectsof actions (loads) by more than 10%.
As regards the behaviour of materials, theassumption of elastic behaviour is the most widelyused for simple structures, even though it is wellknown that the behaviour of concrete is not elastic,by far.
Going further into the analysis, it is necessary toassess the moments of inertia of the sections -depending upon the cross section area of the steelbars used for reinforcement, which are unknown atthe primary stage of design. Therefore the initiallyestimated dimensions of the concrete section aregenerally considered in the first place, yet iterativecalculations can still be made if in doubt.
In implementing structural finite elementmodels, the main problem may be the modelling ofsoil behaviour, i.e. the definition of a stress-strainrelationship. In this respect, Winckler’s springmethod is a traditional approach, consisting of a lin-ear relationship, for each point of the bottom slab,between the vertical pressure (effective stress)exerted upon the soil and and the vertical displace-ment. The ratio between these two values, is termedthe subgrade modulus K:
σ' = K ∆z if ∆z < 0σ' = 0 if ∆z ≥ 0
where σ' is the effective pressure exerted by the soilupon the structure.
The value of K commonly varies from approxi-mately 10 to 50 mN/m3 for a rubble foundation seton natural soil.
It should be noted that this simplistic methodtotally disregards shear stress resistance within thesoil. Furthermore, the value of K may vary acrossthe slab, and according to depth as well, in the caseof non-homogeneous soil layers, and so theappropriate mean value is not easy to determine.Therefore, a sensitivity analysis of the value of K,chiefly upon the stress values in the bottom slab,must be performed. The considerable influence ofthis parameter generally means that the assumptionsmade for the resistance of the soil are inadequate.
If the uncertainties about the soil behaviour arecritical to the design, more complex element models,such as finite element models, may also be used, butsuch finite element models require that there shouldbe adequate soil testing and adequate interpretationof the tests. So complex as the soil model may be, itseems sensible to test the design against a localreduction of K and/or against settlements in areasthat may suffer scour (e.g. in corners).
2.2.2 Scale Models
The assessment of external stability for verticalstructures is often assessed by means of “scalemodels”, which can also provide the more particularloading that can be generated on some members thatare particularly exposed.
13
Such members are isolated mechanically fromthe structure and supported by weighing gauges.This can be useful for some types of structures witha perforated front wall. The effects of impacts relat-ed to wave breaking can also be derived from thiskind of device.
Alternately, this can be carried out using atwo-stage numerical model:
• first one can implement a global model and aloose meshing, apply the loadings butexclude the impacts at this stage.
• then the specific members subject to theimpacts can be computed using a morerefined meshing that is fitted into the firstone; the limit displacments implemented inthese “zoom” models are those derived fromthe initial model.
2.2.3 Limit State Design and Risk Analysis
Structural analysis was traditionally carried outto working stress design limits, until limit statedesign methods were introduced in the early 1970’s,since when they have been used for caisson design.When limit state methods are used, the proceduresrequired are as follows:
• specification of the limit states to be consid-ered, regarding the requirements the structureis intended to fulfil
• selection of the parameters that are deemedrepresentative of events to be accounted forwith special care for variable loads
• setting up of load cases with application ofpartial coefficients
• implementation of structural models• assessment of relevant concrete sections,
using national regulations as regards thepartial coefficients to be applied to the resist-ance parameters of the materials (steel andconcrete).
It is not simply a case of adapting partial factorsfrom one source to another, because the principlesof reinforced concrete design may be different. Itmust also be noted that when the partial factorswere set up for structural purposes, the results wereintended to be similar to designs using traditionalworking stress codes. Even if the limit state conceptinvolves a semi-probabilistic approach, to date, itsimplementation has been carried out in a determin-istic way. However, the application of generalconcrete codes of practice to maritime structuressuffers from a lack of guidance, which leads to avariety of interpretations for the designer. Indeed, it
proves hard to specify the frequency of the events tobe considered and the representative parameters as afunction of both the lifetime of the structure and thenature of the limit state that is being considered.
For example, for some configurations, even withgeometries as simple as caissons, it may not beobvious whether to classify all the loads asfavourable or unfavourable, as this may depend onthe section, the member, or the other loads. Shouldthese loads be critical to the design of the member,it may even be advisable to perform calculationswith both favourable and unfavourable partial loadfactors. Otherwise, the partial factor may be chosenequal to 1.
It, however, remains the case that partial factorvalues, as recommended by general national regula-tions for the assessment of concrete or steel (etc.)structures, have been calibrated for the design ofland-based structures in relation to the failure proba-bility for the failure mode under consideration andfor some reference period - e.g. the design workinglife, (buildings, bridges, etc. Refer to Eurocode 1and Section 2.3.2 for the definition of design work-ing life) so they may not be appropriate to structuressubjected primarily to environmental actions (loads),such as waves. Similar problems of application existin relation to earth pressure loading and geotech-nical calculations, as referred to in 2.1.4 and 2.1.5.There are, for example, problems in applying theprinciples of BS 8002 and Eurocode 7. As noted in2.1.4, BS 8002 adopts limit state philosophy butdoes not involve partial factors. There is afundamental difference between the development ofearth pressure loading with respect to strain, and theforms of loading met in buildings and bridges.There is a strong body of opinion which claims thata partial factor approach is incompatible with soilmechanics.
At present there is no programme for issue of aPart 5 of Eurocode 2, for marine and maritimestructures. There are groups in Europe working onthis subject and the suitability of partial factors, andthe Japanese have introduced a set of factors whichthey have used for the design of prestressed caissons(ref Kiyomiya O, 1994 and also Kiyomiya O andYamada M, 1995)
PIANC’s safety approach, as can be seen fromWG12 on rubble mound breakwaters, seems to offera more appropriate design approach. Much workstill needs to be done to work out operationalmethods for the structural design of caissons. Themain problem is to determine damage functions thatwould be valid for the various limit states to be con-sidered, for they must address physical phenomenafar more complex than those relating to externalstability.
14
2.3 Durability of Concrete
2.3.1 Durability, IntroductionSpecification for durability for most materials is
derived from a “materials” point of view, i.e. fromthe properties of the material, the environment andthe expectations of protective treatments such aspainting or cathodic protection. The ranges ofsuccessful performance are subject to research andexperience, and can often be specified by referenceto compositional limits and by performance tests,fatigue limits etc.
Until the present time this “materials” approachhas also been applied to concrete and reinforcedconcrete. Prescriptive forms of specification (refsBeeby A W, 1992, Clifton J R, 1993) and empiricalrelationships between concrete mixes and laboratoryand field performances given in codes of practicehave been deemed to achieve a satisfactory perfor-mance in certain classes of exposure.
Although this approach has achieved relativesuccess in most land-based building applications,this has often proved to be far from the case in theseverity of maritime exposure, particularly to seawater or de-icing salts. The most dramatic failuremechanism is that of reinforcement corrosion,which can impose severe limitations in relation tothe design and economic feasibility of complex thinwalled structures such as caisson breakwaters orlight structural superstructures.
Great advances have been made in the computa-tion of wave loading, risk analysis and the selectionof the partial safety factors to apply to the level ofrisk in conjunction with the modern computationalpower for analytical design. In the same period, thedurability of reinforced concrete subject to severemarine exposure or de-icing salts has been adramatic failure (refs Aïtkin P C, 1993, Rostam S &Schiessl P, 1993). The present generation of nationalcodes of practice and even the latest joint EuropeanCommittee for Standardardisation (CEN) codes stillreflect the prescriptive approach, and are seriouslyout of date. A consensus view for appropriateguidance will not emerge in less than ten years.
The concept of durability is intrinsically con-nected with the concept of the service and designlife of the structure, which is also related to theanalysis of appropriate loading conditions and limitstate analyses. Owners need to define the servicelives required from their assets and plan a strategyfor maintenance. Current codes do not give arational basis for design of concrete to meet such alife.
The assessment of this further “Durability” limitstate (it is not actually a limit state but a means bywhich the other limit states are maintained over theoperational life) is a fundamental part of design andmust both predate and form part of all stages ofstructural design and detailing. A structured "dura-bility plan" is required to ensure that durability
issues are addressed at all stages and primarilyinclude:
• choice of a service and hence a design life inorder to achieve this service
• recognition of the severity of the specificenvironment(s) affecting the structure
• recognition of the consequences of theenvironment(s) on design, detailing andmaterials
• analysis for durability, by analytical modelwhere applicable
• quality assurance and quality control of bothdesign and construction
• monitoring and maintenance strategy.
These recommendations for a rational approachto design for durability can be matched with thefollowing sources:(i) The CEB Guide to Durable Concrete Struct-
ures (ref CEB Design Guide, 1992). Thismainly relates to buildings rather than seastructures, but gives excellent explanations ofthe mechanisms of deterioration and factorsinvolved. Major contributions to the CEBdesign guide were made by a number of mem-bers of the CEB (Comité Euro-Internationaldu Béton) General Task Group 20 : Durabilityand Service Life of Concrete Structures).
(ii) The work of RILEM committees (refs SchiesslP, 1993, Rostam S & Schiessl P, 1993), andCEN committees.
(iii) The work of the British committees such asthe Concrete Society Working Party onDurability Design and Performance BasedSpecification of Concrete, Report CS 109,1996, and “Durability by Intent”, a strategyfor the UK Dept of Environment programmeon durability of concrete and reinforcedconcrete. (UK Dept of EnvironmentProgramme on durability of concrete andreinforced concrete).
(iv) Important regular international conferencessuch as Bahrain (refs Bahrain Conferences)and Durability of Buildings and Components(refs Durability of Buildings and Components).
(v) The work of the Japanese Bureau of Ports andHarbours, Port and Harbour Research Instituteand the Overseas Coastal Area Institute ofJapan (ref Technical Standards for Ports andHarbour Facilities in Japan, 1991), and papersfrom the Institute and the Japanese ConcreteInstitute (refs Proceedings of Symposia -Japanese Concrete Institute, 1988 and 1989).
(vi) The work of the Australian CSIRO Divisionof Building, Construction and Engineering (refHo D W S & Cao H T, 1993).
(vii) Collaborative research projects such as theBrite-Euram Project 4062 on the ResidualService Life of Concrete Structures.
15
In the following sections 2.3.2 to 2.3.8, thefactors and mechanisms which control durability areoutlined. In the “materials” section 2.4, specificmeasures to achieve durability are advised.
2.3.2 Design Working Life (or Service Life)The definitions of service life, design life,
economic life, etc, must be considered with greatcare, as the meaning of these terms is not the same.The terms are used with different meanings indifferent papers and different contexts and the firsttask is to re-define these.
The recommended definition of the operational“service life” is that given in the draft Eurocode 1(ref Eurocode 1 ENV 1991-1:1994) as follows:
Design Working Life - “The assumed period forwhich a structure is to be used for its intendedpurpose with anticipated maintenance but with-out major repair being necessary”.In the Spanish Maritime Works Recommend-
ation (ref ROM 0.2-90, 1990) the term “MinimumDesign Life” is used for this period.
For maritime works subject to the probabilityand return periods of waves in addition to all otherprobabilities, this definition may require someadjustment, such as the following suggestion:
“The assumed period for which a structure is tobe used for its intended purpose with anticipatedmaintenance but without major repair beingnecessary within a probability appropriate to thefunction of the structure.” It is already stated inENV 1991-1: 1994 that a different level of reliabilitymay be generally adopted for structural safety andfor serviceability and that a different level ofreliability may depend upon the cause or mode offailure, amongst other factors.
With respect, now, to durability in relation todeterioration of construction materials, the “DesignWorking Life” as so defined is the period specifiedby the Owner and is related to operational strategy.Following on from this, the designer has to selectdesign methods and safety factors in order to ensure
a reasonable probability of achieving the specifiedlife. Thus “Design Life” in the British Standardguide to durability (ref BS 7543, 1992) is defined asthe period of use intended by the designer to supportengineering specification and analytical decisions.The “Design Life” as estimated will therefore be atleast equal to or exceed the specified “DesignWorking or Service Life” by a prudent margin,which includes factors of safety and ignorance.
Different “lives” may need to be considered foreconomic and feasibility considerations (ref PortDevelopment 1978, UNDP) and different strategiesfor monitoring deterioration and maintenance canapply.
The concepts of “Design Working Life” and“Design Life”, and the differences between thevarious definitions may not be immediately obviousto readers unfamiliar with papers on the subject and,from experience, can lead to argument. Thedifference between the various definitions may beclarified by Table 7.
A logical structure of design working (orservice) lives for maritime structures (although theretermed as “minimum design lives”) is given inTable 2.2.1.1 of the Spanish Maritime WorksRecommendations (ref ROM, 0.2-90, 1990) as setout in Table 8. The “design working lives” corres-pond with Classes 2, 3 and 4 of the draft Eurocode 1(ref Eurocode 1 ENV 1991-1:1994), but are usefullyexpanded to include general use and specificindustrial infrastructure, which can be especiallyapplicable to port and coastal works. This Table isrecommended for use.
It is important to note that, despite the recentintroduction of more clearly defined “service” lives,current codes of practice and design guides do notprovide guidance for adequate analysis or the meansto satisfy the stated lives. However, a framework formodifying partial factors commensurate withdesigning for different lives and probabilitites isavailable. (ref Eurocode 1, Part 1, Annex A andRilem Report 14, 1996, edited by Sarja andVesikari) and from the work of Sub-Group D ofPIANC WG 28.
16
Designated “Life”
1. Design Working Life (ref Eurocode 1) orService Life (ref BS 7543) or MinimumDesign Life (ref ROM 02-90, 1990).
2. Design Life (ref BS 7543, 1992) withrespect to durability.
3. Economic Life (ref Port DevelopmentUNDP, 1978 ).
Explanation
The utilisation period (or periods) specified by the Owner, with respect to structuralsafety, serviceability, or durability of structure and components (components likelyto have shorter periods).
A period at least equal to (1) or greater than (1) by a prudent factor, employed bythe designer in order to achieve (1). Note: the estimates are not a precise scienceand cannot be guaranteed, but they can be subject to rational analysis.
A period used for economic and financial studies, i.e. for comparison of alternativecapital and maintenance policies, using discounted cash values.
TABLE 7SUMMARY OF DEFINITIONS OF VARIOUS “DESIGN” LIVES
17
TYPE OF WORKOR INSTALLATION
GENERAL USEINFRASTRUCTURE
SPECIFIC INDUSTRIALINFRASTRUCTURE
REQUIRED SECURITY LEVEL
LEVEL 1 LEVEL 2 LEVEL 3
25 50 100
15 25 50
LEGEND
GENERAL USE INFRASTRUCTUREGeneral character works: not associated with the use of anindustrial installation or of a mineral deposit.
SPECIFIC INDUSTRIAL INFRASTRUCTUREWorks in the service of a particular industrial installation orassociated with the use of transitory natural deposits ofresources (e.g. industry service port, loading platform for amineral deposit, petroleum extraction platform, etc).
LEVEL 1Works and installations of local or auxiliary interest. Smallrisk of loss of human life or environmental damage in case offailure.(Defence and coastal regeneration works, works in minor portsor marinas, local outfalls, pavements, commercial installa-tions, buildings, etc).
NB: 1. The General Use period of 25 years corresponds withClass 2 of draft Eurocode 1.
LEVEL 2Works and installations of general interest.Moderate risk of loss of human life or environmental damagein case of failure. (Works in large ports, outfalls of large cities,etc).NB: 1. The General Use period of 50 years corresponds withClass 3 of draft Eurocode 1.
LEVEL 3Works and installations for protection against inundations orinternational interest. Elevated risk of human loss or environ-mental damage in case of failure.(Defence of urban or industrial centres, etc).NB: 1. The General Use period of 100 years correspondswith Class 4 of draft Eurocode 1.
*Defined as Design Working Life in draft Eurocode 1.
TABLE 8DESIGN WORKING-LIVES (SERVICE LIVES) DEFINED IN ROM 0.2-90 AS
"MINIMUM DESIGN LIVES"* FOR WORKS OR STRUCTURES OF DEFINITIVE CHARACTER(IN YEARS)
2.3.3 Processes of Deterioration
There are a number of well-known deteriorationprocesses for the concrete matrix in sea water, andtheir relative significance depends on the specificlocation and climate, but the most widespread andserious problem is that of chloride-induced corrosionof reinforcement or embedded metal generally.
A schedule of the deterioration mechanismsapplicable to maritime structures is given in Table 9.
Guidance and limits relating to these forms ofdeterioration are covered in National Standards andother Codes (refs EAU - German WaterfrontStructures Code, BRE Digest 363, 1991 - Sulfateand acid resistance of concrete in the ground,Concrete Society Report TR 30, 1995 - Minimisingthe risk of ASR, BRE Digest 330, 1988 - Alkaliaggregate reaction in concrete, NF P 15-010, 1985- Guide d’utilisation des ciments, NF P 18-011,1992 - Classification des environments agressifs)some of which are mandatory in their country oforigin. Care must be taken to clarify to owners andcontractual parties if and where departures are to be
made from National Standards on rational grounds,such as mainly apply to reinforcement corrosion.
The dominant factors involved in the durabilityof concrete are:
• the recognition that concrete is a porousmaterial and its behaviour depends on thepore structure achieved and, where applicable,cracks
• the transport mechanisms for water and dis-solved deleterious agents and gases withinthe pore structure and, where applicable,cracks
• the macro, meso and microclimate for thestructure and particular element.
An excellent explanation of the significance ofthese factors and the transport mechanisms is givenin the CEB Guide (ref CEB Design Guide, 1992).
The following recommendations mainly concen-trate on the case of reinforcement corrosion. Refer-ence to the other forms of deterioration generallycan be left to National Standards and referencessuch as given in Table 10. Some details from recentpublications are given in 2.4.3 and 2.4.4.
18
TABLE 10GUIDELINE DOCUMENTS FOR DETERIORATION MECHANISMS
Deterioration Mechanism
Sulfate attack
Salt weathering
Alkali aggregate reaction
Frost action
Abrasion
Early Thermal cracking
Plastic shrinkage and settlement
References
National CodesBRE Digest 363, 1991 (UK)CEB Guide to Durable Concrete Structures, 1992NF P 15-010 (France)NF P 18-011 (France)
Advice is not generally collated in Guides, but given in papers:-Fookes P G, 1993, Bijen J M, 1992 Al-Rabiyah A R, Rasheeduzzafar, Baggott R, 1989
Concrete Society TR 30, 1995 (UK)BRE Digest 330, 1987 (UK)CEN Technical Report, 1994NF P 15-010 (France)NF P 18-011 (France)
Japanese Papers (Koh Y and Kamada E, 1993)CEB Guide to Durable Concrete Structures, 1992
Advice is not widely availableCEB Guide to Durable Concrete Structures, 1992
BS 8007 1987CIRIA Report 91:1993 (UK)CEB Guide to Durable Concrete Structures, 1992 Appendix AACI 207.2R-90 Japan Concrete Institute. Manual of Massive Concrete, 1986(Japanese only)
CEB Guide to Durable Concrete StructuresConcrete Society Report 22, 1992 (UK)
Deterioration Mechanism
Reinforcement corrosion (due to chlo-rides)
Sulfate attack on concrete matrix
Salt weathering of concrete surface
Alkali-aggregate reaction
"Frost" (freeze-thaw) action
Abrasion
Early thermal cracking
Plastic shrinkage cracking (workman-ship)
Plastic settlement cracking (workman-ship)
Locations most likely to occur
Elements wetted but subject to drying -especially hot dry climates. See Figure 13.Corners subject to increased wetting andthen drying. Areas of low cover.
Delayed action in seawater.Colder waters may be more critical.
Elements subject to concentration of saltsby drying - intertidal zone. Paradoxically,cements which achieve the finer porestructure and resistance to steel corrosionmay be most susceptible.
Susceptible aggregates, pessimum re-action with mixed aggregates. Alkalisfrom sea water and marine aggregates.Rich mixes.
In cool with freezing zones withprolonged and repeated freezing.
Subject to abrasive bed movement,shingle, vessel impacts, ropes and moor-ings.
Thick sections and massive structuresbuilt in separate pours, causing restraintto shrinkage during cooling from heat ofhydration.
Arid climates, drying winds, low bleedmixes.
Deep sections, high bleed mixes.
Method of Avoidance
Analysis, design and detailing. Properlydesigned cover to reinforcement for spe-cific exposure conditions and tolerances.
Specification and tests.
Specification.Extensive water curing.
Specification and tests Petrography. Mixlimitations.
Specification and detailing. Air entrain-ment spacing factor.
Higher strength concrete, detailing,extensive curing, controlled permeabilityformwork, permanent steel protection.
Design and detailing, specification,pre-cooling of mix, cooling pipes inbuiltfor the hydration period.
Curing and protection at casting.
.Mix design, reduction of bleeding.Revibration.
TABLE 9DETERIORATION MECHANISMS FOR MARITIME CONCRETE
2.3.4 Exposure ClassificationThe policy of current European Code
Committees for concrete (for the revision of ENV206 : 1989 to pr EN 206) is to classify exposureconditions specific to the various deteriorationmechanisms, where appropriate. The transportmechanisms affecting the deterioration of either theconcrete matrix or embedded metal are largelydependent on the pore structure and the environ-mental conditions within and immediately exteriorto its surface. Deleterious substances are transportedby the medium of water (excluding CO2 and O2),and the moisture content of the concrete controlsboth the rate and effect of the transport of suchsubstances by air or water. The mechanism ofreinforcement corrosion is foremost in mind through-out the subsequent text but other mechanisms, i.e.frost, also apply in colder regions. In seawaterenvironments it is necessary to recognise three basicclimatic influences; that of the macro, meso andmicroclimates (ref Fookes P G, 1993) i.e. climateson the scale of the country, the site, and theparticular element of the structure respectively. Formost practical purposes the macro and meso envi-ronments can be considered together. Themicro-environment, i.e. the location of a specificmember in relation to sea level, and the degree andfrequency of inundation by seawater and drying out,is particularly critical.
The most important macro and meso climaticfactors are temperature and rainfall. Temperaturecontrols the rate of chemical reactions and thedegree of drying out of the cover concrete. Rainfall,humidity and the location of the member in relationto sea water level movement control the wetness ofthe concrete, which affects the mechanism for pene-tration of chlorides and controls the penetration ofoxygen to fuel the corrosion process. Generally,only the surface layer of concrete “dries out”, thedepth of the drying and wetting zone being muchgreater in arid climates than in temperate climatesand, consequently, extremely critical for corrosionof reinforcement. The wetting and drying depth in atemperate climate may not exceed 20mm, where itmay be at least 75mm to 100mm in arid conditions.(refs Bakker R F M & Roessink G, 1992, Bijen J M,1992).
There are four main sub-divisions of macro-climate:
• cool with freezing• temperate• hot wet• hot dry
and some seven micro-environmental cases ofexposure applicable to chloride-induced corrosionin maritime works in ascending order of severity:
(NB The classification below has been adaptedfrom cases XS1 to XS3 now being proposed byEuropean code committees for concrete exposed tochlorides).
XS1 Exposed to airborne salt but not in directcontact with sea water. Contrary to popularbelief, airborne chlorides alone, without thevehicle of water penetration, do not achieveenough concentration to cause reinforcementcorrosion in superstructures (ref Hussain S E,Paul I S & Bashenini M S, 1993)
XS2 Submerged (Also a subdivision, XS2A - back-filled)
XS3 The tidal, splash and spray zones. Dependingon the macro-climate and degree of wetness,this class may need to be broken down furtherinto four cases of ascending severity:XS3.1 Mid and lower tidal. The continuity of
saturation in this part of the tidal zoneis beneficial in restricting the flow ofoxygen, and conditions can be similarto XS2 (refs John D G, 1992, John DG, Leppard N W & Wyatt B S, 1993)
XS3.2 Upper tidal and capillary rise zonesXS3.3 Splash/spray zonesXS3.4 Mostly dry infrequently wetted: i.e.
concrete which is above the splashzone but subject to seasonal change insea level, storm events, testing of firehydrants and run off from mooringlines at capstans and bollards.
Again, contrary to popular belief, the worst caseis not necessarily the splash and intertidal zone. Thelatter is likely to be the worst case for the materialitself, i.e. bare and painted steel, timber, masonryand plain concrete, and for frost damage, but not forsteel embedded in concrete. In cool and temperateconditions there may not be much difference betweencases XS3.3 and 3.4, but the difference can be muchmore pronounced in hot wet and, especially, hot dryconditions.
Fookes (Fookes P G, 1993) has suggested an11-point scale for exposure risk of all concrete in ahot salty environment, both land-based and maritime.In a similar way, but restricted to coastal structures,severity ratings for concrete in a salt-water environ-ment have been expressed on a scale of 1 to 12, asillustrated in Table 11 and Fig 13 (ref Slater D andSharp B N, scheduled for publication late 1997). Itis important to note that the “very severe” and“extreme” exposure conditions of BS 8110 : 1985and BS 5328 1990/1991, and the most aggressiveClass 4 of DIN 1045:1988, only reach about 3 onthis 12 point scale, and it follows that such codesapply only to the cooler and temperate parts ofEurope and similar climates. ACI 318:1995 andEurocode 2 1992 classifications for seawater ex-posure range between 1 and 6 on this scale, depend-ing on location within the tidal and splash regimeand the ambient climate. Cooler European condi-tions may, of course, read only 4 or 5 on this scale.
Sea water contains both sulfates and chloridesand therefore the concrete has to provide both achemical resistance and protection for the reinforce-
19
ment. The latter requirement tends to control theconcrete quality. In concrete that is submerged, XS2or XS3.1, chlorides penetrate the concrete by diffu-sion, a slow process when compared with capillarysuction. In submerged sections, the lack of oxygenis likely to stifle corrosion to an extremely low rate.
Exposure class XS3 has the worst conditions.Chlorides will penetrate by a combination of ab-sorption/capillary suction and diffusion. In periodsof calm weather, wick action and evaporation leadsto a maximum concentration of chlorides near theexposed surface some distance above the water line.
Chloride is transported into concrete by themedium of water by a number of mechanisms.Chloride permeation is a complex phenomenon,including capillary suction and diffusion mecha-nisms, and hydration suction. The most rapid trans-port mechanism, capable of conveying the greaterquantity of ions into dry or partially saturated con-crete, is the capillary suction or absorption process.
The much slower process of diffusion takesplace in saturated concrete. The largest increase inchloride content of the cover zone is achieved bycapillary rise or suction followed by surface evapo-ration, or unbalanced cyclic wetting and drying, i.e.irregular inundation by salt water followed by aperiod in conditions which enable the concrete todry, due to seasonal water level changes or stormevents (refs Saetta A V, Scotta R V and Vitaliani RV, 1993, John D G, 1992, Sandberg P J P,Petterson K, Arup H and Tuutti K, 1996).
20
TABLE 11SUGGESTED SEVERITY RATINGS FOR MARINE ENVIRONMENTS
(ref Slater D and Sharp B N, scheduled for publication late 1997)The ratings are estimated factors for the relative rates of chloride induced corrosion for the same concrete element exposed to
different marine environments. The higher the rating the more severe durability risk7
CLIMATE ZONE
Cool with freezing Temperate Hot wet Hot dryLocation1
Mostly dry2 3 3 43 9-124
Splash/Spray zone 3 3 4 6
Extreme Upper tidal 3 3 4 6
Mid and Lower tidal 2 2 3 3
Underwater 1 1 2 2
Backfilled Faces 5,6 2 2 3 3
1 See Figure 13.2 Infrequent splashing by seawater but otherwise exposed to
weather e.g. copes.3 Concrete exposed to direct sunlight but protected from
rainfall in a hot wet macro-climate may experience amicro-climate of higher severity rating because of the dryingand increase of salt concentration caused by the absence ofwetting.
4 Rating increases with increase in ambient temperature andduration of dry periods e.g. Lower Arabian Gulf coastline 12.
5 For concrete above water level, if capillary rise and evapora-tion is not prevented this may cause increased salt concentra-tions in the fill and lead to a higher severity rating.
6 Sulfate attack due to ground conditions not taken intoaccount. Surface coating recommended if concrete mixrequired to resist chlorides not adequate to meet sulfate class.
7 Abrasion effects not taken into account. Additional covermay be needed over and above Table 15 values to providerequired design life allowing for abrasion loss.
Notes:-
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Fig. 13 Suggested severity ratings on a scale of 1-12(Refer to Table 11)
Hot-dry conditions are particularly hazardous forconcrete in maritime works. They apply pre-dominantly in the Arabian Gulf and Middle Eastgenerally, where aridity greatly exceeds thatexperienced in other regions, including the UnitedStates (ref Fookes P G, 1993). Absorption ofchlorides and penetration of oxygen is usuallyrestricted in wetter and cooler climates whichprevent the concrete drying out at the surface. Thiswetness, itself, is not the cause of corrosion but thereason for reduced corrosion. In hot arid climatesthe concrete dries to greater depths in long seasonsof calm weather or seasonal lower tides. Wetting ofthe dry surface which then occurs in occasionalstorms, seasonally higher tides, the testing of firehydrants or from mooring ropes, causes chloride-laden water to be sucked in deeply. Subsequentdrying causes the salt content to concentrate. In awet climate, rainfall would both reduce this concen-tration and the penetration of oxygen. Lack of waterin the pores permits the free flow of oxygen. (RefJohn D G, 1992, John D G, Leppard N W and WyattB S, 1993). Damage from this cause can particularlyapply to precast concrete (such as caissons) which isleft dry for a long period and then exposed to seaconditions. Torben Hansen, 1989, graphicallydescribes the self-destruct possibilities in theseconditions.
The chloride concentration is only one factor, asthe rate of development of corrosion depends upon anumber of factors, principally regulated by theaccess of oxygen. The presence of water reduces theingress of oxygen and the rate of reaction may beaffected by cracks, in which case anodic andcathodic areas can be relatively close to each other.Depending on the relative magnitude of the coverand degree of saturation, corrosion may be generalwith closely adjacent anodes and cathodes, or theanodic reduction process at a location may be drivenby cathodic activity caused by the ingress of oxygenat another location remote from the site ofcorrosion. Only free chlorides cause corrosion andconsequently the “threshold” value for the criticalchloride content depends upon the chloride bindingcapacity of the particular binder type, and thedegree of saturation.
For a given set of exposure conditions, thepropensity for and rate of corrosion is influencedmainly by cover to reinforcement, binder type,water-cement ratio and binder content, broadly inthat order.
If capillary suction is the main transport mecha-nism it is unlikely that, using most modern Portlandcements alone, concrete quality and cover willprovide an adequate design life in severe exposurecases, if only conventional cover thicknesses andplacing tolerances are employed (ref Bamforth P B,1993, Neville A M, 1995). However, concrete ofreadily achievable production quality used inconjunction with realistic reinforcement placingtolerances is more likely to achieve design lives
when Portland cement is blended with otherhydraulic or pozzolanic materials, such as blast-furnace slag, pulverised fuel ash or microsilica,provided that these materials are of appropriatequality.
For the past 20 years the most popular model foranalysing the ingress of chlorides has been that ofdiffusion, using Ficks second law, all arising fromthe pioneering work of Tuutti (ref Tuutti K, 1982).There have subsequently been developments withmore realistic models dealing with partially saturatedconcrete (ref Grace W R, 1991). A number ofEuropean workers have more recently developedpowerful suites of computer programs whichattempt to model the whole range of variables ofconcrete condition and permit calibration fromsimple laboratory measurements (ref Kiessl K R,1983, Roelfstra P E, 1989, Saetta A V, Scotta R Vand Vitaliani R V, 1993). The limitations of the Fickmodel, and the explanation why its use in interpreta-tion of chloride profile figures can lead to over-pessimistic estimates of the rate of ingress, aregiven in a recent paper by Danish workers (refJohansen V, Golterman P and Thaulow N, 1995).
The exposure cases specific to “frost” (actuallyfreeze-thaw) damage are given as follows:
XF1 Moderate water saturation without saltXF2 Moderate water saturation with saltXF3 High water saturation without saltXF4 High water saturation with salt.Recommendations in connection with freeze-
thaw resistance are given in 2.4.3.
2.3.5 Influence of Cement TypeThe principal choices for cement include plain
(previously known as “ordinary”) Portland cement,sulfate resisting Portland cement (i.e. a cement withspecific limits on the calcium tri-aluminate (C3A)content: the lower the C3A the more resistance tosulfate attack) or various types of blended cements.Blended cements include a combination of Portlandcement with blastfurnace slag (gbs), which is a latenthydraulic binder, or pozzolanic materials which canbe natural but are more likely to be pulverised-fuelash (pfa). The latest available material is microsilica,which can be mixed with either unblended orblended cements. Guidance on its use is given inConcrete Society Report 41 (ref Concrete SocietyReport 41, 1991).
It is most important to stress that any compari-son between materials to different NationalStandards must be made in full knowledge andcomparison of the different methods of test andspecification. This particularly applies to cementstrength class. Cements are produced to a number ofstrength classes and cement content for equivalentdurability may need to be increased when lowerstrength cement classes are used.
21
Sulfate resisting Portland cement is unlikely tobe necessary in maritime concrete, as the disruptiveeffect of sulfates is reduced in the presence ofchlorides, and more so in warmer waters. It isclaimed that, on the contrary, such sulfate-resistingcement is less resistant to reinforcement corrosion,but this may only be true of chlorides inbuilt into amix. Sulfate-resisting cement is a relatively “lowerheat” cement (relative to plain Portland cement) andits use is beneficial for reducing the generation ofheat of hydration.
Where only Portland cement is available, it isusual to limit the tri-calcium aluminate content inreinforced concrete to not less than 5% and notmore than 10%. This constitutes a compromise byspecifying, in effect, a moderate sulfate resistanceby virtue of the upper limit on C3A whilst avoidinga too low content C3A which may be less able toprotect the steel from corrosion. ASTM C-150Type II cement properties can cover this require-ment. In circumstances where reinforcementcorrosion is not the critical problem, or for whichadequate provision has been made, the long termeffects of the reaction between sulfate and thehydrates formed from C3A could possible dictatethe choice of low C3A Portland cement. See 2.4.4.However in these circumstances, blast furnacecements could combine the optimum solution.
High proportions of blastfurnace slag, of from50% to 70% and more, are beneficial from a numberof viewpoints, including resistance to chlorideingress, sulfate resistance, minimising the effect of“alkali aggregate reaction” and reducing the rate ofgeneration of heat of hydration. Alkali aggregatereactions are the chemical processes which can takeplace when the natural alkalinity of cement causedby the calcium hydroxide (pH about 12.5) isincreased to a pH of over 13 due to the oxides ofpotassium and sodium. Proportions of pulverisedfuel ash of the order of 30% to 40% have similar butlesser benefits. There appears to be benefit inadding microsilica in doses of about 5% in conjunc-tion with slag or pulverised fuel ash. Where slag orpulverised fuel ash is unavailable, microsilica maybe added in amounts up to 10% of the total cementi-tious content. It must be noted that the quality ofthese materials have to be appropriate. In some partsof the world slags and pulverised fuel ash containunsuitable constituents, and in some places plainPortland cement appears to be very effective.Blended cements may be either factory produced orblended from separate constituents at the mixer.
Blast furnace cements have been traditionallyemployed for maritime and other works in Germany,Holland and Spain, and pozzalanic cements usedelsewhere. Often blast furnace cement was usedbecause it was the principal material available.Recommendations are now made by manyauthorities, including the Japanese TechnicalStandards (ref Technical Standards for Port andHarbour Facilities in Japan, 1991), for the use of
high slag cements in maritime concrete. Althoughblended cements are almost a “must” to counteractcorrosion, their resistance to surface scaling andtolerance to poor curing is less than unblendedcements. Lower levels of slag, say 50%, may bemore appropriate in colder conditions when thegenerally slower action of slag cements may causeproblems in achieving early strenght.
The weakness of much prescriptive advice incurrent codes is that guidance on mixes and coverthickness is given largely independently of cementtype, coupled with imprecise description of expo-sure conditions. From the comparisons summarisedabove, the conclusion can only be that modernunblended Portland cement generally (there areexceptions) has the lowest resistance to chloridepenetration and, where severe chloride exposureconditions exist, is unlikely to guarantee a long ser-vice life, even in temperate climates, with cover toreinforcement as often recommended. For values ofcover in the accepted magnitude of 50 to 80mm,blended cements are a must.
For each cement type and blending ratio, onecan estimate minimum cement contents and maxi-mum water-cement ratios appropriate to the differ-ent exposure cases. Typical figures are given inSection 2.4.3 to 2.4.5. See Tables 14 & 15.
2.3.6 Influence of Cement ContentConcrete strength is directly related to water-
cement ratio. The pore structure of concrete dependson the water-cement ratio, the degree of hydrationand the cement type (ref CEB Guide).
Using superplasticisers, even under difficultconditions for controlling water demand withoutlosing workability (i.e. high temperature, pooraggregates, large pours, large handling distances), itis now possible to reduce the water-cement ratiobelow 0.40, when this is required, and at the sametime achieve both the desired fine pore structure andself healing properties due to the presence of unhy-drated cement (ref Aïtcin P C, 1993).
Another simple fact is that, ignoring the effect ofwater reducing admixtures, for a given workability,the amount of water required per cubic metre of agiven aggregate type and gradation is effectivelyindependent of the cement content (ref Barber P,1989 and Fig. 14). This is the basis of the ACImethod of proportioning concrete mixes (ref ACI211, 1991), in which the water content is first estab-lished, and the cement content then derived fromappropriate requirements for the water-cement ratioon the basis of durability.
Another important fact is that the water demandrequired to achieve a level of workability variesinversely to aggregate size. The larger the aggregatesize, the less water and cement paste is requiredand, consequently, less cement is required toachieve a given water-cement ratio. Illustration ofthe range of this effect is given in Table 12 (ref BS
22
5328, BS 8110 and derivation using Neville AM,1995, and also recent mix design software). It mustbe noted that the precise definition of aggregate sizecan differ between codes. This principle is coveredby French codes NF P 15-010 and 18-011, wherethe minimum cement content for concrete in seawater or prestressed work is given by the followingexpression, together with a water-cement ratio lessthan or equal to 0.50:
C = 7005√D
where: C = cement content in kg/m3
D = maximum aggregate size in mm.This expression leads to the following figures:
10 mm aggregate : 440 kg/m3
20 mm aggregate : 385 kg/m3
40 mm aggregate : 335 kg/m3
80 mm aggregate : 290 kg/m3.It is now possible to achieve both low water-
cement ratio and adequate workability using a rangeof normal to high performance water-reducingadmixtures. The type of cement, in addition, has itsown effect on water demand. Slags and pulverisedfuel ash generally require less water. The designmix must provide the constructor with adequateworkability to transport, place and compact theconcrete.
As the quality of the concrete in relation to bothstrength and durability is directly related to water-cement ratio and unit water-content, withinreasonable lower limits, the actual cement content istherefore a secondary consequence of the waterdemand for a given mix. Cement contents do notneed to be especially high for plain concrete or evenreinforced concrete providing requisite parametersare met. Cement contents must always be assessedin relation to aggregate size and grading, and a goodway of reducing cement content is to increaseaggregate size and/or reduce the fines content. This
device has traditionally been used in maritimeworks, where 40mm, 80mm and even 150mmaggregate has been used for plain concrete. This isimportant because cement content may be a control-ling factor from the point of view of minimisingalkali aggregate reaction (ref Concrete SocietyReport TR30, 1995) or controlling heat evolution inrelation to early thermal effects (refs CIRIA Report91, 1992, BS 8007, 1987, UK Dept of Transport BA24/87 and BD 28/87, 1987).
Mix specification limits given in specificationdocuments must be realistically drawn from a ratio-nal mix design using available materials, and notmerely numbers drawn from literature, as is oftenthe case.
The latest mix design methods available in themarket make use of computer simulation to achieveproportions based on packing theory. Analysis usingsuch software is given in brackets in Table 12 (refMixsim Issue 3, 1995). Computer simulationincludes the effect of grading, and could incorporatethe effect of mimimising the fines content, which isanother approach to limiting cement content.
2.3.7 Cracking and the Influence of Cracks
The causes and consequences of cracking inconcrete structures have often been misunderstood.In the past, most incidences of early thermal crack-ing were erroneously attributed to drying shrinkage.Most causes of cracking during the plastic state, i.e.
plastic shrinkage and plastic settlement cracking canbe resolved by attention to mix design, protectionfrom drying winds, and curing in arid conditions(ref Fookes P G, 1993, Al-Rabiyah A R, Rasheed-uzzafar, Baggott R, 1989). Early thermal cracking isan important case for design and construction oflarge masses of concrete, including both blocks andcaissons, and is explained in 3.1.7.
An excellent explanation of the processes andinfluence of cracking in concrete is given in theCEB Guide 1992, Concrete Society Report No. 44,1995, and BRE Digest 389, 1994.
23
TABLE 12APPROXIMATE ADJUSTMENTS TO MINIMUM
CEMENT CONTENTS AND WATER DEMANDS FORAGGREGATES OTHER THAN 20mm NOMINAL
MAXIMUM SIZE FOR A GIVEN CEMENTSTRENGTH, AGGREGATE, TYPE, WORKABILITY,
ETC.(Main figures from BS5328, BS8110 etc. Figures inbrackets from computer simulation Mixsim, 1995)
Nominal1 Adjustment2 to Difference in watermaximum minimum demand for a given
aggregate size mm cement content kg/m3 workability kg/m3
10 +40 (+30) +15 to + 20 (+15)20 0 040 -30 (-20) -15 to -20 (-15)80 -70 (-50) -35 to -40
(-20 to -30)
0 100 200 300 400 500
Mas
s of
mat
eria
l per
cub
ic m
etre Water
Stone
Sand
Cement
Fig . 14 Graphical illustration of variation of constituents forpractical range of cement content (ref Barber P, 1989)
over this range of normal mixes, thewater content is effectively constant
Cement content per cubic metre kg
1. Note that definitions of maximum aggregate size can differ betweendifferent codes.
2. According to the French expression overleaf, the adjustment is some50% greater, but of course depends on the choice of "minimum".
Although the significance of cracking may bemuch less than traditionally assumed in the case ofreinforced concrete, it may be significant if it resultsin the reduction of the mass of some unreinforcedconcrete structures (for example armour units) andcan cause unnecessary concern. Attention to detailin both design and construction, particularly inrelation to early thermal effects, can obviate theincidence of cracking. Indeed, a number ofEuropean tunnel structures are currently beingdesigned as “crack free”.
Most structural codes have flexural crack widthlimitations appropriate to specified environmentalconditions related to the serviceability limit state.The British method of design for crack control inrelation to early thermal stresses (ref BS 8007, 1987)has similar limits. However it is now generallyaccepted (ref CEB Guide, 1992, Schiessl P andRaupach M, 1997) that, once crack widths exceedsome 0.1mm, there is no significance in relation tothe ingress of deleterious substances to causereinforcement corrosion from wider cracks of up to0.5mm. Accordingly, the careful gradation of theeffect of cracks between 0.1mm and above is mean-ingless (ref CEB Guide, 1992). This statement refersto cracks perpendicular to main reinforcement andnot to cracking above and along the length of a barsuch as can occur in plastic settlement or if a bar(say a stirrup) acts as a crack inducer. Early thermaland workmanship related plastic cracking tends tocreate wider, uncontrolled, cracking coincident withreinforcement, which is therefore much moredetrimental than flexural cracks which intersect withthe reinforcement at right angles. The progress ofreinforcement corrosion is largely dependent on theproperties of the concrete itself, the moisture state,and the relative cover to reinforcement. However,cracks are likely to promote accelerated pitting cor-rosion. Appropriate detailing to distribute reinforce-ment spacing and limit crack widths remains impor-tant for various reasons.
Cracks do not significantly increase the effect offreeze-thaw damage, as the scaling or splitting ofconcrete by freeze-thaw action is due to the increaseof volume of completely water-filled pores, and is a
feature of the general surface of the concrete (CEBGuide, 1992).
The recommendation for increased cover givenin this Report is likely to be met by the objectionthat the width of cracks on the surface will beunacceptably wider than permitted by design codes.Many engineers will be worried that increased coverwill lead to wider flexural cracks and that reinforce-ment further distant from the surface will lead towider cracking of the cover zone. However,chloride ions and oxygen penetrate the concreteeverywhere and not just at cracks, and the width ofthe crack at the surface is not as critical as the widthat the reinforcement itself. Research has shown thatthe crack width at the steel bar is almost independ-ent of cover thickness, and that the width of the ‘V’shaped crack increases almost linearly with coverthickness (see Fig. 15). Therefore, the cover thick-ness should not be limited for crack width reasons.When checking crack width limits in accordancewith codes, one may increase the permissible crackwidth pro rata to the ratio of preferred cover to thetypical cover given in a code. This principle isalready included in some instances, for examplewhere permitted surface crack width is given as aproportion of the cover thickness (i.e. such as 0.004times cover thickness). (refs Merkblattsammlung,German Concrete Association, 1991, Dept ofTransport (UK) BA24/87 and BD 28/87, 1987)
Crack widths caused by early thermal effects canbe controlled by appropriate design (ref CIRIA 91,BS 8007). It is not necessary to sum the effects ofearly thermal and flexural cracking (ref Departmentof Transport (UK) BA24/87 and BD 28/87).
2.3.8 Influence of CuringThe need for curing continues to be hotly
debated and is the subject of contemporary studies(ref CIRIA Research project on the influence ofpractical on-site curing, in Progress 1994/95). Inthe past, with much less reactive cements, pro-longed water curing was necessary in order toachieve strength. This is no longer the case and,obviously, the influence of water curing on thesurface is unlikely to influence the centre of massivesections. On the contrary, indiscriminate applicationof cold water could cause thermal shock crackingand as many problems as it attempts to solve.
It is perhaps the case that, in the case of wettemperate climates, relaxation of curing is less dele-terious than is generally held. However, in hot dryarid climates, water loss from the surface can besignificant and result in incomplete hydration. As(ref CEB Guide, 1992) the relationship of water-cement ratio and degree of hydration dictates theresulting pore structure of the concrete, and as theduration of water curing is inversely proportional towater-cement ratio, one main advantage of using thelowest possible water-cement ratio is to reduce the
24
concrete covering thickness
crack width
Fig. 15Decreasing crack width, measured from the outside face to the
steel , for fixed stresses σs and different cover thickness c.(ref Merkblattsammlung, German Concrete Association, 1991)
Cov
er m
m
duration of water curing. From the graph/nomogramof the relationship between water permeability,volume of capillary pores and degree of hydration(CEB Guide, 1992), it follows that the lower thewater-cement ratio, the smaller is the duration ofcuring required for the preclusion of continuouscapillaries. This follows from the classic work ofPowers and Brownyards (ref Powers T C and Brown-yards T L, 1988) and as explained by Hansen T C,1989, Ho D W S and Lewis R K, 1983, and NevilleA M, 1995. A higher strength concrete obtained byadopting a lower water-cement ratio is therefore, toa degree, self-curing. Work at the University ofDundee (ref Dhir R K, Hewlett P C, Lota J S andDyer T B J, 1994) is in progress for the achievementof self-curing by adding water-soluble chemicals toreduce water evaporation in the set concrete.
The requirements for curing are also intimatelyrelated to the type of cement and bleeding character-istics of the mix. All of these properties tend to haveconflicting effects on the result. For example a highbleeding rate may increase the water-cement ratio ofthe cover concrete and lead to plastic settlement.However an unduly low bleeding rate (as can occurwith microsilica) may lead to dessication of thesurface and hence to plastic shrinkage cracking.Slag cements are generally held to require greatattention to curing to prevent surface breakdown.However it can also be the case that surface break-down by salt weathering of slag cements is due tothe fineness of the surface pore structure, whichleaves inadequate room for the accommodation ofsurface salt crystals which may be provided by aconcrete with a coarser pore structure.
2.3.9 Monitoring and MaintenanceStrategic planning for maintenance is directly
linked with the concepts of service life and -durability. In the past, very little guidance for themaintenance of concrete and other structures hasbeen available.
The subject is only now receiving due attention.The Technical Standards for Port and HarbourFacilities in Japan, 1991, places much more em-phasis on maintenance and durability than in itsprevious version and points out that maintenance islargely related to monitoring and managementstrategy. Maintenance needs depend on feedbackfrom regular monitoring and data-collection, forcomparison with “base-line” data, i.e. line, level etc,
of the situation “as constructed”. This emphasis isalso made in the CIRIA/CUR Manual on the use ofrock, 1992.
PIANC’s Permanent Committee for DevelopingCountries (PCDC) undertook, in 1978, to produce aseries of maintenance manuals for port infra-structure, (ref PIANC Bulletin 1985-No.50) but, sofar, manuals have only been produced for the moreobvious cases of mechanical equipment, roads andrailways. PIANC Working Group No.17 publishedtheir report on the Inspection, Maintenance andRepair of Maritime Structures exposed to MaterialDegradation caused by a salt water environment in1990 (ref PIANC, 1990). This Report gives helpfulguidance on inspection methods and monitoring butcovers such a broad range of materials includingtimber, stone, unreinforced concrete, reinforced andprestressed concrete and steel, that the specificexplanations and guidance for each material aresomewhat abbreviated.
Monitoring must be planned and adequaterecords taken, beginning with “base-line” measure-ments of line, level etc. immediately after com-pletion of construction. Computers can now greatlyfacilitate this kind of work.
Regular inspections of the structure (or beach orcoastal defence) should be carried out at least onceper year, most likely following the winter stormperiod, and after any major storm event. The princi-pal objects of the survey are to determine:
• the integrity of armour units and elements ofthe structure
• indication of movement and settlement• scour.Measurements must be taken at specific loca-
tions to provide adequate mapping of the structureat clearly located profiles or to a grid, and plotted onlarge-scale drawings. Computer and modern digitalmethods are now available. Aerial survey andunderwater video recording can be used.
In the case of reinforced concrete elements,base-line data includes records of “as-constructed”measurements of cover to reinforcement, and crackand damage mapping. Non-destructive methods ofmonitoring the performance of reinforced concreteare not as practicable or meaningful as often claimedand suffer problems of interpretation, but equipmentis available for monitoring corrosion potentials byhalf-cell and other methods. It is now possible tobuild in probes and take electrical measurements toassess the onset of any change.
25
2.4 Materials
2.4.1 Rock and rubbleFor guidance for, and the specification of, rock
and rubble construction it is recommended to usethe recently published “Manual of the use of rock incoastal and shoreline structures”, published jointlyin the UK and Holland (ref CIRIA SpecialPublication 83/CUR Report 154, 1992.)
2.4.2 Filling and BackfillingGranular materials are recommended for any
backfilling to walls and filling of caisson cells inorder that the earth pressure loading can be deter-mined with confidence, as outlined in Sections 2.1.4and 2.1.5. There are no strict requirements for thecharacteristics of the material used, i.e. quarry runrock, sand etc. It is sufficient that their saturateddensity is equal to or greater than the mass used inthe stability calculations and that their angle ofinternal friction is equal to or greater than the oneused in earth pressure calculations. Generally, if thematerial is suitable for pouring directly in positionunder water, it will be acceptable. Material pouredinto water (except for rockfill), will achieve a lowrelative density and measures may be necessary toincrease the density, as described in 2.1.5
2.4.3 Concrete Durability General -Design, Detailing and Workmanship
The requirements for achieving durability ofconcrete in maritime works will usually outweighthe requirements for achieving strength, or density.The factors influencing durability and the specificdeterioration mechanisms involved were sum-marised in Sections 2.3.1 to 2.3.8. This section and2.4.4 to 2.4.12 summarise the practical steps andspecifications appropriate to maritime concrete inbreakwater applications.
The measures to be taken against specificdeterioration mechanisms as listed in 2.3.3 andTable 9 are most concisely set out in the CEB Guideand as listed in Table 10. Many problems can bemitigated by good detailing and by good curing ofthe concrete surface. Risk of frost damage andreinforcement corrosion can be lessened byattention to drainage of water from near horizontalsurfaces, to prevent the ponding of water, and toprevent the run-down of water on vertical surfaces.Reliability of design and workmanship and thechance of avoidance of “gross errors” can beimproved by employing the discipline of QualityAssurance in both design and construction, and the
Quality Control and audit procedures included inthis process.
National Codes of Practice and Design Guides,and their associated test methods and limits canusually be relied upon for appropriate design forprotection against physical and mechanical action,freeze-thaw damage, chemical and sulfate attackand alkali-aggregate reactivity. However, at thepresent time, national codes and standards cannot berelied upon to assure durability against reinforce-ment corrosion in the long term, particularly inaggressive conditions and climates other than coolor temperate.
It is recommended that, in the maritime environ-ment, design for durability needs to be “explicit” inthe sense understood by European code committees,meaning a move away from prescriptive limitsgiven in codes to methods based on rational analysisand methods of test (ref Concrete Society Report109 on Durability Design and Performance-basedSpecification of Concrete 1996, Rostam & Schiessl,1993, Clifton J R, 1993).
The principal deterioration mechanisms in thecase of unreinforced concrete are those of saltweathering, surface scaling and freeze-thaw, andabrasion. The principal cause of deterioration ofreinforced and prestressed concrete is the corrosionof embedded metal but, obviously, the other deterio-ration mechanisms as for unreinforced concrete stillapply. The performance of different concrete mixesdepends upon the environmental conditions and thetype of cement.
The durability of unreinforced concrete incorpo-rating an appropriate water and cement content isgood in sea water conditions and, where seriouslyaggressive conditions exist, as defined in 2.3.4, con-sideration should be given to designs appropriate tounreinforced concrete as opposed to designs involv-ing embedded metal.
2.4.4 Unreinforced Concrete (Plain or Mass)The term “Unreinforced Concrete” will be used
to define concrete without steel reinforcement. Suchconcrete is usually termed “mass” in Europe andUK. In the USA and Japan, mass concrete meansconcrete of a size significant to heat generationwhich will require measures to be taken on accountof heat generation.
The surface of unreinforced concrete requires tobe designed and constructed with freeze-thaw, abra-sion, salt weathering and sulfate attack in mind.Although sulfate attack is mitigated in sea water bythe presence of chlorides and is less harmful inwarmer waters (refs BRE Digest 363, 1991, Bijen JM, 1984, Matta Z G, 1993, NF P (18-011)), it ispossible that the long-term effects of the reactionbetween sulfates and the hydrates formed from C3Amay be more significant in cases where the much
26
earlier and dramatic effects of reinforcementcorrosion are absent.
The choice of grade of concrete can be par-ticularly problematical in the case of unreinforcedconcrete in blockwork, large sections and armourunits. A number of contemporary code recommend-ations may lead the designer to concrete mixes withexcessive strength and cement content, and hence toproblems of early thermal cracking and brittleness.Due to the changing chemistry of cement it is notalways valid to compare specifications withsuccessful mixes used in the past. The UK MaritimeCode BS 6349 Part 1 : 1984 recommends a minimumcement content of 350 kg/m3 and maximumwater-cement ratio of 0.50, but this is meaninglesswithout reference to the options of increasedaggregate size or the use of admixtures, both ofwhich could enable the water content and hence thecement content to be reduced. The PIANC recom-mendations (ref PIANC Final Report of 3rd Inter-national Commission for the Study of Waves) forunreinforced concrete in 1980 are given in Table 13.
Depending on the exposure conditions, theanswer may lie between these figures and a suggest-ed range of mixes as given in Table 14 below. In allcases it is necessary to carefully consider anydeparture from local National Standards.
Abrasion resistance is obviously important inlocations where concrete may be abraded by shingle
or suspended sand. Abrasion resistance is usuallyachieved by using strong aggregates and higherstrength concrete, not less than C40/50 (cylinder/cube). Test panels in the UK have shown lossesbetween 2mm and 12mm per annum over a periodof 7 years. Exposure cases for abrasion have onlyrecently been proposed in the Eurocode committees,but their adoption is uncertain. Achievement ofabrasion resistance depends critically on the finish-ing operations and curing. There are a number ofabrasion tests.
The CEB Guide is very informative in relationto “frost” damage which, more correctly, should betermed “freeze-thaw” damage. Freeze-thaw resis-tance generally increases with reduced water-cement ratio, increased cement content, and highercontent of air pores. A higher proportion of blend-
27
TABLE 13PIANC RECOMMENDATIONS FORUNREINFORCED CONCRETE 1980
(ref PIANC 1980)
Nominal maximum Minimum cement28 day cube
Aggregate mm content kg/m3(size of cube not
defined)strength MPa
40 220 to 270 }20 to 3520 250 to 31010 360
TABLE 14SUGGESTED CONCRETE MIXES FOR UNREINFORCED CONCRETE FOR DIFFERENT MARINE CONDITIONS TO
AVOID SALT WEATHERING/SURFACE SCALING FOR 40mm MAXIMUM AGGREGATE SIZE8
(ref Slater D and Sharp B N scheduled for publication late 1997)
Exposureseverityrating
1-3
4,5
6-12
Maximumwater/
cementitiousratio2
0.55
0.50
0.45-0.40
Minimumcementitiouscontent1,3,4
kg/m3
300
325
3507
Suggested cement typeand blends 8, 9
Minimum Dimension of Pour
< 500mm > 500 mm5
pc, ASTM, I/II/IV/V 100% ASTM IV or Vor pfa blend 75% pc : 25% pfaor gbs blend 50% pc : 50% gbs
pc, ASTM I/II/IV/V 100% ASTM IVor pfa blend 75-70% pc : 25-30% pfaor gbs blend 30-50% pc : 70-50% gbs
pc, ASTM I/II/IV/V 100% ASTM IVor pfa blend 75-70% pc: 25-30% pfaor gbs blend 30-50% pc : 70-50% gbs
Typical ConcreteGrade6
(cyl/cube strength10)MPa
C30/37C25/30C25/30
C35/45C30/37C30/37
C45/55C40/50C40/50
Notes:1 Where appropriate, in large sections, advantage can be taken of the
lower cement paste and cement contents required for largeraggregate size. Minimum cement content for 80mm aggregate isapproximately 40 kg/m3 less. See Table 12.
2 Maximum water/cementitious ratio may need to be reduced to meetNational Standards. For any exposure rating, the water/cementi-tious ratio should be as low as practicable and economic.
3 Minimum cementitious content may need to be increased forabrasion resistance or to meet National Standards, with increasedrisk of early thermal cracking.
4 The figures for minimum cement content ranges applies to Portlandcement only. Larger figures may be required for blended cements.
5 Cement type or blend chosen to control heat of hydration. 40mmmaximum aggregate size recommended to allow cement contents tobe reduced by 20-40 kg/m3.
6 Approximate equivalent grade consistent with the minimumcement content and the maximum water/cement ratio. The accurateequivalent grade for controlling water/cement ratio and cementcontent should be established for the actual mix, if necessary bytrials.
7 Water reducing admixture recommended.8 Add air entrainment 4-6% for freezing conditions.9 Pfa and gbs blends require good curing conditions to avoid surface
defects.10 Values are characteristic compressive strengths tested at 28 days in
accordance with ENV 206, below which 5% of all possiblestrength test results may be expected to fall.Cylinder strength applies to 150mm diameter 300mm long cylin-ders and cube strength applies to 150mm cubes as defined by ENV206:1990 and tested in accordance with BS 1881: Part 116, 1983.
ing materials may unfavourably affect scaling.Exposure classes for freeze-thaw damage were givenin 2.3.4. Concrete with moderate water saturationwill not suffer from freeze-thaw damage. Salt causesa substantial drop in temperature at the concretesurface during thawing, and the different temperaturebetween the surface and the internal concrete causesinternal stress. One of the worst potential damagesituations is where there is a prolonged freezingcycle with a source of external sea water. The mostserious condition is reported to be when thesubstrate remains frozen but the surface thaws dueto solar gain. Water from the melted snow/ice entersthe concrete to form added ice, which re-freezesduring the night to cause “squeeze freezing”.Freeze-thaw conditions are likely to be much moreserious for roads as opposed to maritime structures.There are a number of standard freeze-thaw tests,such as ASTM and Scandinavian tests. Japanesetests are carried out to determine the change inlength of specimens.
Salt scaling is a prevalent phenomenon in hotarid conditions. It occurs in maritime conditionsand, due to sulfate, in low quality mass concrete insalt flat conditions. The subject is not covered inmore well-known guides and reference needs to bemade to Fookes (ref Fookes P G, 1993, Bijen J M,1992), and various papers on the Bahrain causewayby Rasheeduzzafar et al (ref Al-Rabiyah A R,Rasheeduzzafar, Baggott R, 1989) of the King FahadUniversity of Petroleum and Minerals, Dahran,Saudi Arabia. Salt scaling has affinity with bothfrost and sulfate attack. The quality of aggregate hasa marked influence and also the cement type.Scaling can be noticeably higher with slag cements,and it may be necessary to tolerate this in order toachieve requisite resistance to corrosion in the caseof reinforced concrete.
The mix proportions should be selected usingmix design methods commencing with the derivationof water demand, which can be restricted to achievea low water-cement ratio (i.e. less than 0.5) by usingplasticisers or super-plasticisers to ensure that ade-quate workability is available. Unduly restrictivelimits should not be imposed on workability wherethis property is required to ensure the integrity ofconcrete within the formwork.
2.4.5 Reinforced Concrete, including Selectionof Cover to Reinforcement
The durability of reinforced concrete is primari-ly dependent on countering the effects of chlorideinduced corrosion of steel reinforcement and ismore a design and detailing matter than a materialsmatter. As developed in 2.3.1 to 2.3.8, durability isa function of environmental loading and dependsprimarily on:
• exposure conditions• cover to reinforcement• cement type• pore structure/water-cement ratio.
For reinforced concrete, the principal designparameters include the exposure rating derived fromthe macro and micro environmental conditions, thecement type, the mix quality as determined by thewater-cement ratio, and the cover to reinforcement.
As the protective capacity of a given concrete isbroadly related to the square of the cover thickness,the provision of appropriate cover to reinforcementis the simplest and most direct way of reducingdamage from reinforcement corrosion. The dramaticeffect of cover on the service life is illustrated inFigs. 16 and 17. In this context cover must beproperly specified and detailed as summarisedbelow. The nominal cover for placing the reinforce-ment, to be used in the design and stated on thedrawings, can be derived by computing a minimumcharacteristic design value to meet durability (oraggregate size, fire or bond etc) requirements,together with a placing tolerance, which may varyfrom 5mm to 15mm, or more depending on thestandard of control.
An illustration of the range of nominal coverwhich is believed to be necessary for an estimated“design” life of 60 years (say to meet a DesignWorking Life of 50 years) is suggested in Table 15.The figures include 15mm tolerance but noallowance for abrasion or salt scaling. It will benoted that as the aggressivity increases, the recom-mended figures for cover are much higher than haveoften been used, especially for unblended cements.The table also demonstrates the importance ofextremely low water-cement ratios for Portlandcement, although even such provision is inadequate
28
Fig. 16 Example of the effect of the thickness of the concrete cover
(ref CEB Guide)
Fig. 17 A prediction by service life model for exposure to sea-water with
a chloride content of 10,000 ppm(ref Clear K C in Hognestad E, 1986)
Dep
th: c
m
Time: Years
without larger cover in the more aggressive condi-tions. Anyone surprised by this table should consultthe following supporting references:
• Standard specifications for Design andConstruction of Concrete Structures Parts 1& 2, Japan Society of Civil Engineers, 1986
• Miyagawa T, 1991• Bamforth PB and Price WF, 1993• Bamforth P B, 1993.The JSCE Standard Specification Part 1, 1986,
states that “in a ‘corrosive condition’ a cover notless than 75mm is advisable (not less than 60mm ifexamination and repair is easy), and not less than100mm in ‘severely corrosive condition’ (not lessthan 80mm if examination and repair is easy).Where the quality of concrete is hampered bydifficult construction conditions or the structurerequires a long life-time, concrete cover shall beincreased. Although cover of precast concrete maybe decreased by 20%, it is not advisable to decreasewhere sufficient corrosion resistance is necessary.”
The selection of appropriate cover for the expo-sure conditions and the fixing tolerance has a majorinfluence on the minimum wall thickness for a cais-son. Minimum wall thicknesses of 200mm and300mm are not likely to be appropriate for rein-forced concrete under extreme conditions of expo-sure. Thin walled structures are not appropriate forreinforced concrete in extreme exposure. In severeconditions it may be advisable to use designs avoid-ing reinforced concrete and, instead, to use a com-posite design. In the submerged and mid to lowertidal part, the sections may be adequate in thin wallcaisson design, but the exposed sections undersevere exposure may need to be thicker to enablethe cover to reinforcement to be increased or, iffeasible, constructed in plain concrete. The changein micro environment will often coincide with achange in construction conditions (i.e. from under-water to tidal or above-tide working) and so acomposite design may be chosen to improveconstructability while at the same time matchingdurability to different exposure conditions. In the
29
TABLE 15SUGGESTED NOMINAL COVER FOR REINFORCED CONCRETE (BEFORE ABRASION ALLOWANCE)
FOR DIFFERENT MARINE ENVIRONMENTS FOR 60 YEARS "DESIGN LIFE"(suggested as appropriate for 50 years "design working life")
(ref Slater D and Sharp B N, scheduled for publication late 1997)
Suggested Nominal Cover1,2,3mmExposureSeverityRating 75% pc : 25% pfa
50% pc : 50% gbs 4,5
70% pc : 30 pfa30% pc : 70% gbs90% pc : 10% ms6,7
100% pcw/c ratio
0.458
100% pcw/c ratio
0.409
Notes:
1 5010 5010 75 65
2 50 5010 95 85
3 65 50 120 105
4 80 60 14511 130
5 95 70 17011 15511
6 115 85 20011 18011
9-12 13512 10012 23011,12 20511,12
1 Includes an allowance of 15mm for workmanship tolerancesand reduction of cover during concreting.
2 Add an extra 10mm for prestressing strand to reducepercentage non-compliance of nominal cover to minimalvalue in recognition of risk of pitting corrosion.
3 A combination of the suggested nominal cover plus concretemix appropriate to higher exposure rating will provideextended service life.
4 Appropriate mix for exposure severity rating 2: GradeC35/45, minimum cementitious content 370 kg/m3, maxi-mum water/cement ratio 0.45, 20 mm aggregate. See note10, Table 14, for definition of Grade.
5 Assumed apparent diffusion coefficient at 20°C 3.0 x 10-13m2 sec-1.
6 Appropriate mix for exposure severity rating 5: MinimumGrade C40/50-55/65) minimum cementitious content 400kg/m3, maximum water/cement ratio 0.40. Appropriate mixfor exposure severity rating 6-12: Minimum Grade
C45/55-55/65, minimum cementitious content 425 kg/m3,maximum water/cement ratio. 0.34-0.38 20 mm aggregate.See note 10, Table 14, for definition of Grade.
7 Assumed apparent diffusion coefficient at 20°C 1.5 x10-13m2 sec-1.
8 Assumed apparent diffusion coefficient at 20°C 15.0 x10-13m2 sec-1.
9 Assumed apparent diffusion coefficient at 20°C 11.0 x10-13m2 sec-1.
10 Nominal cover of 50mm dictated by bond requirements with20mm maximum aggregate size and allowing for workman-ship tolerances.
11 Blended cementitious mix more suited to the exposureseverity recommended.
12 Note that this Severity Rating is for hot arid conditions andinfrequently wetted members. See Section 2.3.4. Extraprotection may be required by means of coatings orprovision for cathodic protection, depending uponapplication and estimated severity rating.
colder and wetter parts of Europe this problem isless critical, and a change from thin-walled caissonto unreinforced construction often occurs in anycase at about mid-tide level, and the problem isgreatly reduced.
At least half the problem of achieving adequatecover is created by failure to acknowledge the needfor adequate steel fixing tolerances. The answer tothis problem lies in the German codes and the newEurocodes, by considering a minimum characteristiccover to meet durability requirements, to whichmust be added a tolerance, ∆h, to derive the nominalplacing cover which is used in the design calcula-tions and the detailing. The resulting compoundtolerances of the concreting process can result, inparticular, in a wide variation in cover, as illustratedin Fig. 18.
Minimum Cover should be regarded as a“characteristic” value, and a margin of at least10mm and up to 15mm should be added to reducethe rate of failure to achieve the minimum cover towithin 5% of measurements (ref DIN 1045, 1988and Betonwerk and Fertigteil-Technik (ConcretePrecasting, Plant and Technology) Issue 51, 1992).This approach applies to measurement of cover inplace after concreting. Cover should be checked bycovermeter on trial panels at the commencement ofwork and “as constructed” records of cover should beprepared for all projects. The size of the tolerance∆h may depend on the class of work, but should beverified by control procedures. Possible specifica-tion methods and compliance limits are suggested inTable 16.
Spacers need to be made from appropriatematerials and adequately distributed. There areseveral references on spacers (refs Spacers forreinforcement, 1981 Cement & ConcreteAssociation UK, Spacers for reinforced concrete,Concrete Society 1989). There are a number ofingenious modern designs of plastic spacers, but itis possible for these to deform or achieve inadequatebond. For seriously aggressive conditions, spacersneed to made of materials equal to or better than theparent mix and treated to improve bond.
2.4.6 Prestressed ConcreteThe use of prestressed concrete in maritime
works is less common. Most engineers have beenafraid for chloride effects but there is a substantialhistory of success in specific applications (refGerwick B C, 1990). A recently constructed exam-ple in Japan is a new type of breakwater in Sakai,with a perforated caisson screen 16m diameter and10m high (ref PCI Journal, July/August 1994).However, the corrosion of prestressing steelpresents a much more dangerous situation than canoccur with non-prestressed concrete. The corrosionof pre-stressed steel proceeds at a faster rate thanthat of non-prestressed reinforcement under identicalconditions, and presents a higher risk of buildingfailure (ref Perl G C and Blades J T, 1993). There isalso the same problem as reinforced concrete, due tothe detailing of secondary reinforcement such asstirrups and reinforcement to the anchorage zones.One of the principal reasons for successful perform-ance has been the general employment of higherstrength, lower water-cement ratio concrete andincreased cover and, perhaps, location in temperateor cool conditions. Prestressed piles have performedwell in generally submerged and wet conditions inthe lower tidal zone in even hot-arid locations, andprestressed decks have performed well in their usuallocation above the splash zone where, of course,they are not usually subject to an aggressiveenvironment, even in the worst hot-arid conditions.
The concept of using prestressing to reduce theeffect of cracks is not the positive advantage it mayseem, as the pore structure of the cement paste andthe state of saturation is more important than thecracks. Concrete strength is usually required byCodes to be higher for prestressed work, but this isnot the case in all codes. Cover to tendons inmaritime work is usually, but not always, requiredto be greater than that for reinforced concrete. Somecodes limit the permitted chloride content of mixesto half that for reinforced concrete, but other codesdo not. A useful comparison of codes is given byPerl and Blades (ref Perl G C and Blades J T, 1993).
2.4.7 CementRefer to 2.3.5. and 2.3.6.
30
���� �� � �����������%)4')��
5������������� ������)$��
�����������������������&(��
�����������������������(-��
&*$
&&$
%,$
%($
%$$
*$
&$
%$�����&$������'$�����($������)$�����*$������+$������,$������-$����%$$������6��7����������������)��
5��.���������������������)����������������
Fig. 18Distribution of cover in practice. Analysis of 1600 cover-
meter readings for a 13m high retaining wall(ref Concrete Society Special Publication CS 109, 1996)
2.4.8 Aggregates
The quality of the aggregates, themselves, has alesser impact on the quality of the finished concretethan may be supposed, provided that appropriatemeasures are taken to achieve a dense concretematrix with strong cement paste of low permeability.However, the quality of aggregates has a significantinfluence on abrasion, freeze-thaw and salt scalingresistance.
As a bulk natural resource, it is usually necessaryto make the best use of locally available materialswherever possible (including, possibly, re-cycledmaterials) and it may be unrealistic to set qualitylimits too high except when standards are set by par-ticularly demanding circumstances. For that reasonit is important to refer to national standards and the
results of locally determined aggregate studies forthe limits to be applied to aggregate properties in theprevailing environmental conditions.
Most national standards set limits for the physi-cal properties of strength and absorption. Wherethese do not exist it is often necessary to refer toother national standards. Tests and limits particular-ly apply with regard to alkali aggregate reaction andchloride and sulfate content.
The properties of shape, texture and absorbencyof aggregates affect water-demand, which have asignificant effect on durability. There are a numberof very informative publications on the properties ofaggregate and comparisons between various nation-al standards (refs Geological Society (UK), SpecialPublication No.9, 1990, Pike D C, 1990, and BREpublications 243 and 244, 1993)
31
TABLE 16SPECIFICATION OF COVER TO REINFORCEMENT
Design Dimensions and Acceptance Criteria
* This acceptance criteria must be applied to a specific area in m2, or a specific member face of appropriate size. At least 35measurements should be taken using a covermeter calibrated by direct measurement, and checked by direct measurement.
Design and Detailing Dimensions
Nominal (Target) Cover
mm
Heavy Civils WorkTolerance 15mm
1008060
Normal In-SituTolerance 10mm(ref Eurocode 2.
5mm ≤∆h ≤10mm)60 5040
Precision WorkTolerance 5mm(ref Eurocode 2
0mm ≤∆h ≤5mm)Say504030
Note that this small tolerance canonly be used if the feasibility of
achieving it has been proven by testsand that results are verified by
production control.
No more than 5%of measurements
less than* mm
856545
504030
Precision work not usuallyapplicable to maritime work,
except for precast planks and submerged concrete.
453525
Note that the standard of control(i.e. as measured by the standarddeviation ÷ mean) exceeds near-
laboratory precision
No Single Readingless than* mm
705538
433425
Precision work not usuallyapplicable to maritime work,
except for precast planksand submerged concrete.
403020
Acceptance Criteria after Concreting - Measured Cover
1. When inspected in the forms, prior to concreting, the cover to reinforcement shall be the nominal (target) coverspecified within appropriate + and - tolerances, say 5mm, or 10mm.
2. When checked by covermeter after concreting, by a covermeter calibrated by direct measurement, the cover toreinforcement shall comply with a similar philosophy to that given below.
2.4.9 Cracking and Crack WidthRefer to 2.3.7 and 3.1.7.
2.4.10 Reinforcing SteelProvisions for reinforcing steel are covered by
National Standards. Steel used in coastal works isusually hot rolled high yield deformed.
Stainless steel may be necessary in some appli-cations, at a premium of some six times the cost ofblack steel. However it should be noted that thepremium on the overall concrete cost is relativelysmall, up to about 16%, say (ref McDonald D R,Sherman M R, Pfeifer D W and Virmani Y P, 1995).Suitability for seawater depends on the Chromium/Nickel proportions. Higher grade stainless steel toBS 6744, Type 316 S 33 has 11-14%Ni and 16.5-18.5%Cr. The lower grade, Type 304 S 31 has 8-11%Ni and 17-19%Cr. Unified European standardsare under preparation and some, including EN10088-1, which lists the steel grades, are alreadyissued. The standard European terminology issimilar to, but not precisely the same as, the DINsystem. The terms 316 and 304 are common to UK,USA and Japan. The higher grade to other standardsis: France Z6 CND 17.12, Germany 1.4436, ItalyX5 CrNiMo 17 13, Sweden 14 23 43. The lowergrade is: France Z6 CN 18.09, Germany 1.4301,Italy X5 CrNiMo 18 10, Sweden 14 23 32. On noaccount should lower grades of stainless steel (BSType 304) be used in chloride bearing water, the useof at least a hot rolled austenitic BS Type 316 S 33being necessary. Higher qualities of stainless steel ata further cost are available, but are produced formarine engine, piping and similar applications.
A CEB document (ref CEB, 1995, Coatingprotection for reinforcement) provides a currentstate-of-the-art report on three coating protectionsystems; hot dip galvanising, epoxy coating andPVC coating.
Fusion bonded epoxy coated reinforcement isavailable to some national standards (ref ASTMA775, 1990), on the principle of isolating thereinforcement from chloride ions. Both successfuland unsuccessful applications are reported andapplication must depend on the severity of theconditions.
Reinforcement is also available in the form ofsteel, polypropylene and other fibres, and is used inboth conventional and sprayed concrete applica-tions. There are standards for materials in USA (refASTM A820-90, ACI 544R82) and for design andconstruction in Japan (ref Japan Society of Engin-eers Recommendations, 1983). Steel fibres havebeen used for the reinforcement of armour units anddurability is claimed to be better than for barreinforcement. The use of stainless or “nearstainless” fibres is claimed to be promising.
2.4.11 AdmixturesIn some countries (UK in particular) the use of
admixtures was overtly discouraged and some traceof this antipathy remains in Specifications which puta barrier on admixtures without express approval.While there is no doubt that admixtures need to be“approved”, it is now often the case that admixturesare essential to provide workability with low watercontent or cohesion in underwater concrete andtherefore need “positive” encouragement, notdiscouragement. Types of admixtures are coveredby national standards or guidelines. Due to thecomplexities of selecting proprietary products, it isnormal to maintain faith in tried and provenmanufacturers.
2.4.12 Additional Protective Measures :Coatings, Coated Reinforcement, Cathodic Protection
In certain cases additional measures are requiredto achieve durability. Such measures can includecoatings to the exterior concrete surface or to thesteel reinforcement, or the recently developed appli-cation of cathodic protection to reinforced concrete.Claims are also made for rust inhibition by addingcalcium nitrite as an admixture, or for pore blockingadmixtures.
None of these solutions are as simple as theysound and usually introduce maintenance problemsof their own : for example, coatings to concreterequire special preparation of the surface such asgrit blasting, and require to be maintained; coatingsto reinforcement are susceptible to handling damage;cathodic protection systems have limits to life dueto cathodic and anodic reactions. However, eachmeasure may add a number of years to the servicelife and, usually, the better the initial concrete, thebetter the performance of the additional protectivemeasure.
Similar difficulties for the choice of admixtureapply to specifying and selecting coatings forconcrete. A UK Concrete Society working groupaims to issue guidance on this subject in 1997 (refGuide to Surface Treatment, 1997). Some countriestraditionally paint concrete. Coatings may benecessary in some cases to achieve extendedperformance, either to delay the onset of reinforce-ment corrosion or protect from chemical attack.Carbonation is unlikely to be a problem in marinestructures, and therefore coatings to protect againstcarbonation are of little advantage. Coatings areusually required to limit the passage of oxygen,carbon dioxide and liquid water, while enabling acertain transmission of water vapour which wouldotherwise reduce adhesion of the coating. Inaggressive conditions, only high-build, high qualityproducts are dependable. These usually compriseexpensive systems of acrylics, polyurethanes andepoxy resins built up to a substantial thickness.
32
Usually, the concrete surface needs to be preparedto a very high standard, using grit blasting, with allthe small air holes formed at the shutter surfacerevealed by this process filled with mortar by hand,because the coating will not be able to bridge theholes. These activities are labour intensive andexpensive.
Cathodic protection is very well established inapplications to steel maritime structures andpipelines. It is now adopted as a repair techniquewhen no other simpler alternative is feasible. It isless likely to be adopted in new construction, but insuch cases it is much easier to design and detail theelectrical continuity of reinforcement for cathodicprotection than it is in the case of retrofixing to anexisting structure.
Some Italian bridges have been designed withcathodic protection incorporated. It is claimed thatthe application of a protective current at an earlierstage in operation inhibits the migration of chlorideions, whilst the applied current is much less thanwould be needed in later years to combat corrosioncurrents made possible by the depassivating conse-quences of chloride ingress (ref Pedeferri P, 1992).
A significant benefit from the latest electro-chemical techniques can be the building in ofmonitoring circuits. A continuous record ofpotentials and/or corrosion currents can then beobtained, which can enable problems to beidentified and appropriate maintenance strategiesundertaken.
2.4.13 Corrosion of Structural SteelThe performance of structural steel in sea water
is much better known and understood than that of
reinforced concrete. Materials for sheet and bearingpiles and similar members met in coastal engineer-ing are covered by National Standards. Advice oncorrosion rates and anti-corrosion measures is givenin the British Maritime Structures Code (BS 6349),the German Water-front Structures Code, theJapanese Technical Standards and by sheet pilingmanufacturers. It is usual to either add extra thick-ness for a "corrosion allowance", or to use highergrades of steel with reduced stress levels.
The corrosion rate is usually greatest in thesplash and low water zones, less in the inter-tidalzone and least in the submerged and buried zone.High corrosion rates can be experienced at lowestastronomical tide level, where anaerobic corrosionand reduction by bacterial action can occur.
There appears to be no merit in the use of spe-cial alloy or copper bearing steels. Coal tar epoxyand similar protective treatments have providedexcellent protection to steel during recent decades.However, current legislation for environmental pro-tection and health and safety, as well as commercialpressure to reduce construction time and cost isleading to the replacement of traditional multi-coatmetallic and duplex organic systems which do notmeet these requirements. A range of new “compli-ant” coatings has been and is being developedwhich include water-borne coatings, high solid coat-ings and solvent-free coatings. There is inadequateguidance on these new systems which are, as yet,unproven in practice. This may especially be thecase for maritime work. A CIRIA research projectwas due to be completed in September 1996 (refCIRIA Project 523, 1996). Where appropriate,cathodic protection can be confidently designed forsubmerged areas and protective paint treatmentsminimise the amount of protective current requiredin the submerged areas.
33
3.1 Caissons
3.1.1 Float-Out LoadingRefer to 2.1.7
3.1.2 First GroundingRefer to 2.1.8
3.1.3 Caisson Fill Methods and PressuresRefer to 2.1.5
3.1.4 Sea Condition Data and Limits for Con-struction Risks
For each construction stage the designer has tospecify the waves which can be tolerated, in height,period and direction. If this wave condition issurpassed by more dangerous waves, the work has tobe interrupted. The design wave for different stagesof the construction can be different. Float out canoften be carried out under more severe conditionsthan grounding. The rubble on which the caissonwill be founded has to resist a certain wave actionwhich normally is more severe than the wave actionwhich has been foreseen during grounding. On theother hand, erosion of the rubble base can berepaired quickly at low cost, and the last levellingon most sites is, in any case, programmed to becarried out shortly before float-out.
To enable the designer and the contractor to esti-mate the wave height for the different constructionstages, they have to be provided with wave forecastdata. This data should not indicate the probability ofthe occurence of a certain wave height during theyear but it should give information about the timeduring which a certain wave height is surpassed,which is shown in Fig. 19. From Fig. 19 it can beseen which time of the year a certain wave height(either given as Hs or as Hmax) is surpassed. In theexample, a wave height of Hs = 1 m is surpassedduring 60% of the year, while a wave height of Hs =2 m is only surpassed during 23% of the time. Also,an indication of the length of calms, necessary forthe float-out processes, has to be given.
There will be only a few places in the worldwhere enough wave measurements have been madein the past to construct Fig. 19. On the other hand,the normal wave climate during the constructionperiod is much more significant to the progress on
site than the wave height which occurs during ashort storm once a year. Some pre-informationabout the changing weather conditions can becollected from local weather stations, airports andfrom weather maps. But for the planning andconstruction of a vertical breakwater, for which thework on water is sometimes restricted to very fewfloat-out dates during the construction period, abetter weather service is needed than that requiredfor the construction of a rubble mound breakwater,where the activity can be changed from day to dayaccording to the wave height.
There are not many references to the limits forconditions suitable for sinking caissons, or quantifi-cation of appropriate limits. Spanish experiencesuggests that, in the case of caissons of the size usedin vertical breakwaters, the sinking operationsshould be restricted to the following conditions:
T1/3 < 7 secondsHmax < 0.7m - tending to decrease.In periods of calm weather, the required cell fill-
ing operation must also be carefully considered. Inlarge caissons the cells have a considerable volume.Consequently the filling operation takes a long time,which must be estimated carefully, in order that theworks programme is compatible with the forecastweather conditions.
3.1.5 Construction JointsFrequent construction joints occur in nearly all
vertical breakwaters, because (with few exceptions)they are built in a discontinuous way by the use ofprefabricated concrete elements.
Where horizontal joints occur as, for instance, inblock type breakwaters, the blocks placed on top ofeach other can be easily linked together by a slotand key system or by wells and dowels, as shown inFig. 20. Another way to achieve a good inter-connection between small elements is to useinclined joints.
34
3. CONSTRUCTION RELATED CRITERIA ANDCONSTRUCTION METHODS
Fig. 19Probability for the exceedence of a certain wave height during
the year (ref Stückrath T, 1982 and Clutterbuck P G, 1977)
highest wave height Hmax
perc
enta
ge o
f ex
ceed
ence
Vertical joints, which are unavoidable for allvertical breakwaters constructed from largerelements, have numerous advantages but they canlead to serious difficulties during construction. Oneadvantage of a vertical joint is that adjoiningelements can undergo different settlement.Therefore these joints should not be filled withinflexible materials as long as the elements stillundergo differential movement. On the other hand,storms arising during the time in which the jointsare left open can lead to high water velocities in thejoints and to damage (Agerschou H et al, 1985).
Interconnection of adjoining elements is in mostcases necessary to distribute the local wave load onto more than one element. Therefore joints whichare permanently open are an exception.
A male and female slot and key system whichhas often been used, for instance in Brighton Marina(Agerschou H et al, 1985) or in Helsingborg 1981 isshown in principle in Fig. 21a. This detail has somedisadvantages. Free settlement of two adjoiningelements is restricted because differential tilting oftwo elements is prevented. Additionally, during thetime of placement of the elements (even, if they are
placed by crane) the sea must be absolutely calm.Even small wave heights lead to impact stresses inthe slot and key, due to the very large masses thatwill try to follow the orbital swell motion. Thereforea joint with a double female slot which is filledlater, as shown on Fig. 21b, leads to a lessvulnerable construction method. The vertical opengaps on both sides of the slots are usually sealedwith grout- filled tubes or "bolsters".
The materials that have been used to fill the gapin the joint in Fig. 21b have been coarse aggregates(in Helsinki 1981) or bitumen (ref Press H, 1962),but in most cases nearly inflexible tremie concretehas been used. The use of modern flexible materialsor a first filling with a soft material which is later
replaced by tremie concrete could be recommended,but experience with the use of these constructionmethods is rare. Surveys of breakwaters (refTanimoto, 1983) that describe the displacement ofcaissons after wave attack, lead to the conclusionthat most displacements of vertical breakwaters arejust horizontal slips. Therefore a differential move-ment of adjoining elements, caused by wave attack,can be prevented by a shear connection located onlyin the bottom slab, as proposed by Lundgren andJuhl, 1995 .
Double slot joints between rectangular caissonsas used in Spain are shown in Fig 22 and for thequay-wall caissons at Dubai Dry Docks in Fig 23.
3.1.6 SettlementVertical breakwaters show much greater settle-
ment than most other structures designed and builtby civil engineers. There are mainly three reasonsfor the high magnitude of settlement.
a) The sea bed on which the structures arefounded is, with few exceptions, loose andfine and cannot be precompacted.
b) The rubble mound which is used as a bearinglayer under the vertical elements and whichhas a considerable thickness, especially forvertically composite breakwaters, cannot
35
(a) Slot and key (b) Double slotFig. 21 Vertical joints between caissons with a circular
horizontal cross-section.
(a) Slot and key system (b) Wells and dowelsFig. 20 Vertical connections at horizontal joints between
concrete blocks placed on top of each other
a b
Fig. 22 Joints between caissons
Fig. 23 Typical joint details - Dubai Dry Docks(continued on page 36)
36
easily be precompacted and will therefore becompressed by the load of the caissons, andby wave loading.
c) Rubble mounds undergo considerable settle-ment during construction and time-dependent"creep" settlement for many years thereafter.These phenomena are explained below.
d) Sometimes the elements have to be placed onfoundations for which a trench has beendredged into the sea bed. In soft ground thismay involve the classical sand replacementmethod (ref Barberis MC, PIANC, 1935). Inthese cases the inflow of soft sea bed mudafter the last cleaning and before the place-ment of the elements cannot be fully prevent-ed, and the depth of fill will undergo bothcompression and consolidation settlement.
In most cases the vertical load of the elementson their foundations (weight minus bouyancy) isincreased from zero after first placement to the fullload after filling with sand and after concreting thecap. The greatest settlements have to be anticipatedduring the construction period. Many breakwatershave shown additional settlements during the firstyears of operation because the settlement is not anelastic movement and because it can be increased bythe shaking by waves. If possible, the last layer ofthe concrete cap which is visible to the eye, shouldtherefore be completed as long as possible after themajor construction period. Settlements can never beexcluded. They are, even if they are high, unavoid-able characteristics of vertical breakwaters. Visibledifferential settlements of adjoining elements shouldbe minimised as far as possible, and the uglyappearance these constructions can exhibit (becauseof the uneven surfaces), should be overcome byappropriate detailing features.
Although the size and the extent of the settle-ments can be measured easily and figures have beencollected for many harbours, not many publicationshave been made about settlement, and very littleguidance appears in the CIRIA/CUR Manual on theuse of rock, 1992. The magnitudes of settlementsgiven in available literature, are as follows:
• "settlements up to 1m are normal" (refLamberti A and Franco L, 1994)
• "settlements of 97cm have been measuredbut the influence of earthquake cannot beexcluded" (ref Ching T K, 1994, page 228)
• "Diagrammi dei cedimenti dei cassoni" showsettlements of 1m during the first threemonths and maximum settlements of 1.5mafter eight years (ref Ing Mantelli & Co.,Volti Harbour, Genova, 1994)
• In the Working Group 28 meeting in Londonon 26 April 1995, Sub-Group B reported thefollowing figures for the vertical breakwaterat Gela (Italy): Overall settlement 1m,differential settlement of two adjoiningelements 0.2m.
Fig. 23 (continued from page 35)Typical joint details - Dubai Dry Docks
In fact, considerable information on the timedependent post construction settlement of rockfill isavailable from dam construction. An immenseamount of investigation has been devoted to thissubject at the UK Building Research Establishment(BRE) and by American workers. Penman’s paperof 1971 (ref Penman ADM, 1971) reviewed thedevelopment of compacted rockfill as a constructionmaterial and illustrated its post-construction behav-iour. Other helpful papers include those by Sowers,Williams and Wallace of 1965 and Matheson et al,1986 and 1989. A series of papers on the latest stateof the art of rockfill structures was presented inLisbon in 1990 (ref Maranha das Neves E, 1990)including a paper by Charles which updates Penman(ref Charles J A, 1990). A UK Institution of CivilEngineers paper by Sharp (ref Sharp B N, 1996)reviews the available data and gives practical casehistories of the settlement of quay walls in ports androckfill used in quay wall construction. The signifi-cant factor is that the post-construction settlement ofa rockfill embankment is time-dependent, due to thecrushing of the highly stressed points of contactbetween the individual rocks. The pattern is similarto the secondary consolidation of clay, in that itreduces logarithmically with respect to time accord-ing to Penman’s expression :
δ = Η α (log t2 - log t1)100where δ = settlement in mm
H = height in mmα = creep coefficientt2 & t1 are any two times from the end
of construction for a settlement δ to occur.
The dimensions of time (months or years) areimmaterial as it is the difference that matters, andthe base of the logs is immaterial. The coefficient α,can vary between 0.2 and 1, and is often about 0.5.The rate of settlement of a 15 m high marineembankment could typically be 25 mm per month ormore at the end of construction and 5 mm per monthone year after completion of construction.Thereafter, the rate of settlement reduces slowlyaccording to the logarithmic law, which means thatsignificant settlement can continue for 20 years. Thelong term settlement due to this cause is unlikely tobe very large, in comparison with the settlement ofclay and seabed mud, but could be of the order of100 to 200 mm over many years. Although thisamount is unlikely to lead to “failure” of a verticalbreakwater, it may prove unsightly and lead tounnecessary worry and incorrect diagnosis of thecause of continuing settlement. Usually thissettlement is made up by additional fill.
The logarithmic decline in settlement rate isillustrated in Fig. 24. It will take nine years sincecompletion of fill to double the amount of settle-ment that occurred from year one to year three sincecompletion of the fill. It will take until year 27 for
the next increment to equal that which occurredfrom year one to year three. Fig. 24 illustrates thecase of a 20 m high embankment with a typicalcreep coefficient of 0.524. Substituting in Penman’sexpression:
H = 20,000 mmα = 0.524log10 27 = 1·434log10 l = 0
δ (from year 1 to year 27)
= 20,000 x 0.524 (1.434-0)100
= 150 mm.Obviously, the expression can’t deal with Year 0
(which has no log) but one can estimate from year 0to year 1 by using decimals of a year (or months).The slope of the log plot m equals αH , which in this case equals 104.8. The settlement100from years 1 to 27 can also be expressed as
δ = m(log t2 - t1), which equals 104.8 x 1.434 = 150 mm, as above.
3.1.7 Early Thermal CrackingControl of the temperature of concrete at the
placing stage and during hydration is virtuallyessential for the construction of massive concretesections. Thermal contraction from the heat generat-ed by hydrating cement results in severe crackingwherever the geometry of the section or thesequence of adjacent pours during constructionimposes a restraint to free contraction.
The subject was dealt with in the 1930’s in theUSA for low cement content mixes in massivedams, but ACI code recommendations do not extendto convenient crack calculation methods for rein-forcement design. Detailed recommendations,which are simple to apply, are given in UK publica-tions (ref CIRIA 91, 1990, BS 8007, 1987, Depart-ment of Transport BA 24/87 and BD 28/87, 1987)and the phenomena is the focus of interest in Japanfor the design of caissons, etc.
37
Fig. 24 Logarithmic decline of rockfill settlement(ref Sharp B N, 1996)
Time in years since completion of rockfillSe
ttlem
ent m
m
*Settlement from year 0 to year 1approximately 100 mm to 150 mm.
A major step forward is now possible due to thedevelopment of computer software which modelsthe transfer of heat and moisture in concrete. Theeffect of different mixes and sequences of construct-ion can thereby be compared and the propensity forcracking, stresses etc, determined (ref CEB DesignGuide, 1992, Appendix A).
Appropriate measures to control the heat ofhydration include:
• the use of the largest appropriate size ofaggregate and water reducing admixtures inorder to reduce the cement content demandedby low water-cement ratios
• the use of low heat cements, usually involv-ing blending with pozzolans or slag
• cooling by the cooling of materials, additionof flaked ice, the injection of liquid nitrogento negate the heat rise, or by cast-in incooling water pipes
• thermal curing and insulation to minimiseheat differences and gradients.
Early thermal stresses may occur in the con-struction of caissons due to the following reasons:
• restraint of adjacent pours• infilling of cells with mass concrete.Specifications frequently exacerbate thermal
effects by demanding unnecessary intervals betweenadjacent pours whereas, to minimise early thermalloading, the maximum freedom from restraintwould result from continuous casting, such asoccurs in slipforming. The old "alternate bay"method, with continuous reinforcement, causes themaximum restraint and is the least favourable.
The calculation of thermal crack widths andcrack control reinforcement by conventional meansare covered in references CIRIA 91, BS 8007 etc.
As the cement content has a significant effect onthe heat evolution during hydration and therefore thetemperature differences applied in the calculations,the temperature effects due to the likely maximumcement content should be used, and very carefulmonitoring of the actual cement content be made inrelation to the calculated reinforcement. The designermust bear in mind that the strength and cementcontent of the mix in practice may be considerablyhigher than a “minimum” requirement of a generalspecification. Ready-mix suppliers may be obligedto overshoot cement contents in order to complywith strength specifications and QA requirements,and this is often a factor which exacerbates earlythermal cracking.
The type of cracking caused by the two types ofrestraint is illustrated in Figs. 25 and 26. "Internal"restraint concerns the change in temperature acrossa thick section, such as a block, whereas "External"restraint concerns adjacent pours. Thermal cracksrecorded in caissons in Japan are illustrated in Fig.27. It is necessary to design the reinforcement to
38
Fig. 26 External restraint for various slab or wall pour sequences(ref CIRIA 91, 1992)
Fig. 25 Internal restraint to early thermal cracking(ref Concrete Society Digest No 2, 1984)
replace the formation of a few wide cracks by anumber of finer, controlled cracks. Bands of addi-tional reinforcement may need to be calculated forthe base of a restrained pour.
Methods for calculating reinforcement related tocrack width are given in ACI 207.2R and theJapanese Society of Civil Engineers StandardSpecification for Design and Construction ofConcrete Structures. These methods are based onproviding reinforcement to resist early thermaltensile stresses in excess of the early-age tensilestrength of the concrete. However, the UK method(BS 8007, CIRIA 91 etc), appears to be more con-venient, using the expression :
ρ = fct ∅ R α (Τ1 + Τ2) fb 2wmax
where ρ = percentage of steel areafct = concrete strength in tensionfb = average steel/concrete bond strength
(In fact fct is usually given direct asfb a quotient)∅ = bar diameter in mmwmax= maximum crack width in mmR = Restraint factorα = coefficient of thermal expansion of
concreteT1 = difference between centre line peak
temperature and mean ambient temp-erature
T2 = maximum temperature difference between adjoining sections.
The amount of reinforcement depends upon thejoint spacing (i.e. wide or close joint spacing) or forcontinuous (jointless) construction. For each joint
spacing criteria, the reinforcement percentage is, ineffect, inversely proportioned to bar size, and is easyto calculate. The result is not, of course, a rigorouslyexact figure, but a likely approximation.
3.1.8 SlipformingVertical slipforming lends itself to the casting of
caissons, silos, walls and towers. Because the plasticconcrete is placed in the forms which act as movingdies to shape the concrete by an “extrusion”process, the concrete is joint-free and is cast andhardens free of restraint from adjacent pours so thatearly thermal effects are minimised.
Explanations and guidelines for the formworkitself are given in ACI 347R, 1988 - Formwork.Tolerances are recommended in ACI 117 1990. Theforms are constructed with a slight taper such thatthe width between the forms is greater at the bottomthan the top. The true wall thickness is measured atthe elevation where hardened concrete is maintainedin the form. The allowable ACI tolerance for cross-sectional dimensions is 20mm.
Slipforming of caissons has been described inseveral references (Cochrane G H, Chetwin D J L,and Hogbin W, 1979, Philip Holzmann literaturefor Port of Damman, 1978). The slipforming of thetapered cylindrical legs of the Condeep platforms isdescribed by Moksnes J, 1975 and Condeep promo-tional literature. The slipforming of the Ekofisk arti-ficial island in the North Sea is described by MarionH and Mahfouz G, 1974. A technical discussion onslipforming (ref Fort G B and Davis P D, 1981)reports the procedures for casting the central plat-form of the Ninian oil field, both in dry dock andafloat, together with other advice and informationon mixes etc. A recent review was given by JonesM N and Horne R D, 1996. The method is lesssuccessful in dealing with discontinuities, such aswindows or slots in the walls. Otherwise dimensionalvariation should be small and the accuracy inplacing reinforcement has been found to be relativelygood, due to the location of vertical steel by guideswithin the forms.
The curing of massive vertical slipformedsurfaces presents logistical difficulties as recognisedin ACI 308:1981 (Revised 1986) “....structureserected using vertical slipforming methods shouldbe cured in accordance with the procedures used incuring other vertical surfaces recognising theparticular problem of slipform construction”. Theimmediately slipformed surface is, clearly, unprotect-ed by formwork in its early hours and is moresensitive to the application of curing activity and, asin the case of horizontally slipformed pavements,there may be little alternative to the earlyapplication of curing membranes. For subsequentapplication of curing activities see 2.3.8 and 3.1.9.
39
Cracks at side wallsFig. 27 Thermal cracking in caissons in Japan
Slit
3.1.9 CuringAs noted in 2.3.8, the purpose of and need for
wet curing is now questioned. The wet curing ofmassive vertical surfaces such as occur in caissonsintroduces severe logistical problems in fixingcuring materials, securing them against wind andweather, and the supply and drainage of water.
The need for wet curing of those parts of astructure which are subsequently to remain totallyimmersed may be particularly questionable.However, wet curing cannot be disregarded forthose parts of a reinforced concrete caisson structurewhich are manufactured in and/or going to beexposed to severely aggressive conditions in a hotarid environment or which are going to be subject toabrasion or frost damage. All of these circumstancesrequire a refined pore structure as influenced by wetcuring amongst other factors.
Indiscriminate wet curing with cold water,and/or the removal of formwork can lead to shockdue to temperature or moisture gradients. Forformed surfaces, it is generally advisable to leavethe forms in place as long as practically possible.The use of higher strength (i..e. low water-cementratio) concrete which is less susceptible to curingduration is recommended, as explained in 2.3.8.There appears to be advantage in the recently devel-oped permeable formwork liners, for the controlledremoval of or supply of water to the surface. If suchliners are left in place after removal of formwork,they provide a protective covering. Vacuum de-watering, to reduce the water-cement ratio of thesurface concrete, has been utilised since the thirties,but is more appropriate to horizontal surfaces.
The procedures for and timing of immersingcaissons in sea water require consideration.Provided the concrete surface is saturated and of aconcrete strength which can be cured in a shortduration, there is likely to be no disadvantage andeven positive advantage of early immersion The“Condeep” oil rig platforms for the North Sea andthe Ekofisk central reservoir were slipformed whenafloat (ref Marian H and Mahfouz G, 1974, andMoksnes J, 1975 and Condeep promotionalliterature). There is very great risk of disastrousabsorption of salts into concrete which has been leftto thoroughly dry at the surface in an arid climateand then suffer unbalanced periods of immersionwith long drying cycles. (ref Hansen T C, 1980).
A matrix of requirements must be consideredwhen deciding the plus and minus factors forcuring, including the appropriate surface porestructure in relation to abrasion or frost resistance orprotection of reinforcement, environmental con-ditions during and after construction, whetherfurther surface treatments are to be applied (inwhich circumstances curing membranes would beinappropriate) etc.
3.1.10 Developments in CaissonsRecent designs for composite caissons in Japan
include a caisson with a superstructure broken upinto wave breaking shapes and infilled with armourunits. The arrangement of this is shown in Fig 28and the stress plot in Fig 29. In the analytical modelthere are 3482 No. elements and 10,070 No. nodes.New developments in prestressed concrete doublewall cylindrical breakwaters, designed by limit statemethods, are described in the 1995 FIP Notes byKiyomia O and Yamada M. See Fig 30.
40
Fig. 28a Perspective of a composite caisson, with armour unitinfill
Fig. 28b Cross section
41
Fig. 30 Recent Japanese double cylindrical wall-type breakwater inprestressed concrete.
(ref Kiyomiya O and Yamada M, 1995)
Fig. 29 Stress analysis of caisson in Fig. 27
3.2 Blocks
3.2.1 Blocks from ConcreteBlockwork is inherently robust and durable, and
was, historically, used for vertical breakwaters inthe UK, France and Spain. The older forms, some-times using smaller blocks, often laid to a slopingbatter, are no longer used. Vertical blockworkbreakwaters in the Mediterranean were less success-ful and a list of famous failures includes theMustapha Jetty at Algiers, and at Catania andGenoa. Guidelines for blockwork breakwaters aregiven in the UK Maritime Code BS 6349 : Part 7,1991 which refers also to BS 6349 : Part 2, 1988 forblockwork quay walls, and examples are given inthe Japanese Technical Standards for port and har-bour facilities and the Spanish Diques de abrigo enEspaña (Breakwaters in Spain), 1988.
Blockwork quay walls are normally dry jointed.BS 6349 : Part 7, 1991 recommends that, where set-tlement is not significant, joints should be sealedand grouted to minimise air and water pressureeffects under wave action. However this adviceappears impractial and not applicable to separateblock construction. Sealing would only be effectiveif it is achieved by infill pours of in-situ concrete, asdescribed below.
A significant difference between blocks inbreakwaters as opposed to quay walls must lie in thesignificance of cracks. In dry bonded quaywallconstruction, some cracks are likely to occur duringplacing and preloading blocks in position. As thequaywall remains in compression due to earthretaining loads, a limited and random incidence ofcracking is insignificant. In the case of unreinforcedarmour units and breakwater blocks, such crackingreduces the mass of individual elements and mayrequire further consideration.
3.2.2 Types of Concrete BlocksModern blockwork breakwater walls can be
classified into four main types:-(i) With massive blocks of cyclopean dimen-
sions of mass up to 400 or 500 tonne. (SeeFig. 31).The block length is equal to the structurewidth. The blocks are usually stacked toform separate vertical columns which per-mits independent settlement of each column.When significant settlement is not expected,the blocks can be bonded (staggered) alongthe length of the wall.Connections between blocks are by mortisesin the horizontal direction as shown in Fig.31, and in the vertical direction sometimes bywells, which are infilled with tremie concrete,and can be armoured by steel dowels. SeeFig. 20(b). The wells can have a dualfunction as the holes for lifting tongs. (SeeFig. 32).
(ii) Built up of smaller blocks, of the order up to60 tonne, as shown in Fig. 33, or less asshown in Fig. 34.
(iii) Forms of blockwork construction whichincorporate large voids or discontinuities inthe sea-face, to absorb wave energy.
(iv) Composite walls of blockwork on a sub-merged rubble mound.
3.2.3 Common ProblemsA number of common “problems” are met in the
manufacture and placing of blocks.1. Planarity of horizontal surfaces
It is not as simple as it may seem to achievehorizontal faces to blocks. Lack of planaritymay result from the hand screeding of theopen top face of a block, or from the tendencyof block edges to curl. It must also be notedthat, to maintain verticality of a wall, it canbe necessary to shim between blocks. Suchlack of planarity can result in cracking ofsome courses of blocks.This problem has been overcome, in the caseof solid blocks, by casting the blocks on theirsides, such that the seating faces are formedvertically between the forms. This cannot bedone in the case of hollow blocks, moreusually used for quay walls.
2. Loss of friction on seating facesIt is dangerous to use felt or building paperto form the base of a pour. If this materialsticks to the block and is not removed, it cancause serious loss of friction. It is preferableto use proprietary "surfectant" (soap-type) ofshutter release products and not to use shutteroils or unsuitable oil products, to reduce therisk of loss of friction and damage toconcrete faces.Another solution is to cast the blocks on asteel base plate which has been treated with aformwork release liquid which is soluble inwater, which dissappears instantly on contactwith air.
3. Cracking of blocks during manufacture,handling and placingCracking during manufacture should beavoided by due attention to the early thermaldesign of the blocks, formwork and protect-ion during curing.Cracking during handling should be avoidedby appropriate lifting and handling methods,and avoidance of premature lifting. It may benecessary to add handling reinforcement tothe walls of hollow blocks, although thisshould be avoided if possible.Cracking during placing and particularlyduring any preloading may occur as a resultof 1. above.
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Cracking can be triggered by changes ofsection at mortice joints, and by the holesformed for interconnection wells and liftingpurposes. It is recommended that holes arenot made for lifting purposes. It is preferableto use steel lifting points, even though thismay necessitate an increased number of lift-ing points and a special lifting device whichbalances the lifting load between the extralifting points. Corrosion of such exposed lift-ing points will be insignificant under water,or to otherwise unreinforced concrete.
4. Quay wall blocks are usually bedded downby stacking blocks on the completedcolumns, either to form a preload surchargeto accelerate and therefore "take-up" settle-
ment, or to bed the blocks onto the rubblebed (i.e. a pressing down or "paper-weight"effect) This measure is unlikely to be
suitable for breakwater blocks,and it may be necessary to con-struct the in-situ concrete crownearly, in order to achieve this"paper-weight" effect. Earlyexecution will of course lose thebenefits to be gained by laterexecution, as recommended in3.1.6, which may minimiseunsightly movement and crack-ing. The joints in the crownblock can be formed on theslant, to permit unevensettlement. Joint spacing in thecrown block can be of the orderof 10 or 12m. The joints areusually sealed with bituminousmaterial.
43
Min 1m50
Fig .32 Lifting tongs recess and vertical well connection
Fig .33Bonded blockwork - Spain
Fig .34Breakwater Eng. Castor - Port of Algeciras, 1935
(ref Diques de abrigo en España, 1958)
Marina "Los Gigantes", 1973.Concrete blocks up to 15 Tonne.
(ref Diques de abrigo en España, 1958)
Seaward Side
Concrete Crown
H.W.L.
Foot Protection Concrete Block
Armour Stone
Harbour Side
Concrete Block
Foot Protection Concrete Block
Armour Store
Rubble
Fig .31 Concrete block breakwater - large blocks(ref Technical Standards for Port and Harbour Facilities in Japan)
Minimum1.5 m
or prestressed precast concrete. There are twosub-divisions of this type, namely single curtainwall and double curtain wall as illustrated in Fig. 35.Practical examples of the various types are shown inFig 36. In the double type of curtain wall, slits oropenings, are provided at the front curtain walls. Thewave energy is dissipated between the two curtainwalls and reflection wave height is attenuated.However, consideration must be given to the risk ofscour of the sea bed and the provision of protectionrock. When steel piles are adopted, corrosionprotection is necessary. The exposure conditions forconcrete elements and their fixings are likely to bevery severe. Cover to reinforcement in suchelements must be adequate to suit the exposureconditions and the materials as recommended in thisReport.
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3.3 Rubble MoundsFor construction aspects of any rubble mound
element of vertical breakwaters refer to the refer-ences of 2.4.1 and 3.1.6.
3.4 Curtain and Pile TypeCurtain and pile type breakwaters are effective
in locations where(1) the water depth is shallow and wave height
is small(2) the sea bed is soft (mud).This type of breakwater usually consists of both
piles and curtain walls, but sometimes the break-water is constructed with piles alone. Piles areusually of steel tube and curtain walls are reinforced
(b) Double curtain breakwater
Fig. 35. Types of Curtain and Pile Breakwaters (ref Technical Standards for Port and Harbour Facilities in Japan)
(a) At Osaka Port (b) At Kelhin Port
Fig. 36
Japanese examples of Curtainand Pile Breakwaters
(c) At Hakata Port
upper concrete
(a) Single curtain breakwater
4. SUMMARY
The topics examined by Sub-Group C compriseda wide range of subjects related to design, construc-tion and performance. They were mostly chosen tofill in the gaps of the considerations of the othersub-groups, whose tasks were to concentrate onloading and reliability criteria for stability underwave impact. The information presented is drawnfrom experience and practice worldwide and atten-tion is drawn to facts and references which may notbe obvious or available in any single comparabledocument. This being the case, it is neither possiblenor meaningful to abridge the considerations. ThisSummary, therefore, mainly provides a key to thedata and tables given in the text, together with sum-marised abstracts and recommendations whereappropriate.
The topics were examined under two overlap-ping headings:
(i) Design Criteria and Materials• different forms of loading other than
wave loading, i.e. earthquake, ice, soiletc.
• structural analysis and limit states forelement design
• durability and maintenance, particularlyof concrete structures
• materials.
(ii) Construction Related Criteria and Methodsrelating specifically to:- • caissons • blocks• rubble mounds• pile and curtain type.
The text is accompanied by an exhaustive refer-ence list.
4.1 Different Loadings not covered bySub-Group A
4.1.1 Earthqnake Although advances in computer techniques
enable dynamic response to be analysed by finiteelement methods, the simple equivalent static loadmethod is generally acceptable for breakwater struc-tures. In many countries, and for the obvious exam-ple of Japan, the horizontal earthquake load is stillcalculated by multiplying the vertical dead load andsurcharge by a seismic coefficient determined froma number of factors, as set out in Tables 1 to 3.
4.1.2 Ice PressureLoad from ice pressure on a vertical breakwater
seldom exceeds the wave load. The effective pres-sure from ice loading decreases with structure sizeand there are, at present, no conclusive formulaewhich can be applied to large works. Therefore, inthose countries where ice loading is a consideration,ice pressures are derived from local experience andjudgement. Some details are given in Section 2.1.2.
4.1.3 Deleted
4.1.4 Earth Pressures for Structural DesignEarth pressure is relevant to vertical breakwaters
with rubble or fill placed against them, and to theload from retained materials within caissons.
Traditional “working stress” codes recommend“active” or “at-rest” pressure coefficients to beapplied to the dry or submerged soil mass, appropri-ate to different forms of construction. Differentapproaches are taken in different countries. Due tothe problems of reconciling limit state methods forsoil mechanics analysis with structural analysis(because the fundamental relationship between loadand movement for soils differs from that for struc-tural materials) there is a lack of agreement in theformulation of limit state codes. Therefore, tradi-tional methods still remain as an option in mostcodes.
New structural analysis codes and geotechnicalcodes now adopt limit state philosophy. Structuralanalysis to limit state codes requires the applicationof partial factors for loading cases and materials forthe calculation of the ultimate and the serviceabilitylimit state conditions. However the application oflimit states and/or partial factors to earth pressureand variable water loading is not as straightforwardas for buildings and bridges. The selection of partialfactors to match with the reliability and probabilityof water and wave loading is the task of Sub-GroupsA and D. The subject is also discussed in 4.2.
There are two distinctive methods of applyinglimit state methods and partial factors to the struc-tural design of earth retaining structures. Onemethod derives directly from structural design: thepartial factors from Eurocode 2 and similar nationalcodes are applied to the characteristic or serviceabil-ity limit state loading. The other method derivesfrom geotechnical stability analysis: a partial factor(or, in the case or BS 8002, a “mobilisation” factor)is applied to a parameter, such as tan ∅´.
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However, the new geotechnical codes tend toconcentrate on equilibrium and stability, and do notgive adequate clarification of the loads to apply tothe structural design of members or to water loading.In maritime structures, the hydrostatic pressurecomponent greatly overshadows the load fromsubmerged earth and there seems to be no point inover-refinement of earth pressure loading unlessclearer consideration is given to water loadingwhich is also, of course, associated with variablewater loading due to waves. A comparison of sevenvarious national applications of partial factormethods for the calculation of structural members isgiven in 2.1.4, and is illustrated by an example,compared visually in Fig. 6. The example demon-strates the range of results for calculation of the loadon one side of a member in 20 m depth of fill ofsome 1.5 to 1. (Note that in practice, there will alsobe water loading on the other side, except in thecase of a lock or dry dock, when the water loadingis critical).
The range of factors lies between the applicationof the partial factors in the structural codes (i.e. 1.4or 1.35 on dead load and 1.6 or 1.5 on live load) tothe unfactored soil properties, and the less conserva-tive loading from new USA, Japanese and olderScandinavian codes and BS 8002 and the draftEurocode 7, depending upon interpretation (wherethe factor is of the order of 1.2). It must be notedthat BS 8002 is understood not to relate to maritimestructures, and the formulation of an interpretationof Eurocode 7 remains a matter of controversy.
4.1.5 Fill Pressures within CaissonsThe loading within caissons is generally derived
from silo theory. Field verification of this approachis illustrated in Fig. 7. An example of how fillpressures calculated to various national standardscompares with the “at-rest” unconfined pressure isillustrated in Fig. 8. The silo pressure of submergedsand is seen to range from 30% to 60% of theunconfined “at-rest” pressure.
4.1.6 FrictionThere is a surprising divergence in the various
national codes between the figures used in designfor friction and for a factor of safety against sliding.The coefficient of friction, compared in Table 4,varies between 0.5 and 1.0 (for different cases) andthe factor of safety between 1.0 to 1.75. In the latestgeotechnical approach to limit state codes, factors ofsafety against sliding or overturning have beenovertaken by the assessment of equilibrium atmodified soil strength parameters.
4.1.7 Handling and Float-Out LoadsLoads, which can arise during construction,
although transient, can be significant and must beconsidered carefully. From the viewpoint of ultimatelimit state design, a partial factor of γF = 1.1 is sug-gested. The forces arising from towing can be takenfrom Japanese standards, as illustrated in Fig. 9.
4.1.8 First GroundingSevere loading cases can arise when a lowered
caisson first makes contact with the preparedfoundation. In most cases the caissons will neveragain undergo a comparable distribution of load.These dynamic loads can not be predicted precisely,but the designer can influence and reduce the risk ofindeterminate load imposition by various means,including downstand legs which predetermine thelocation of first grounding.
4.2 Resistance Analysis, Internal Analysis
Structural analysis of caissons can be carried outby the traditional approach, in which the structure issplit into sets of beams and slabs, guidance onwhich is amply given in national codes. Computermethods are likely to be used for two-dimensionalframe analysis. For detailed final design it is morelikely that full three dimensional model analysiswill be used, using finite element analysis.
In implementing finite element models, the mainproblem may be the modelling of soil behaviour, i.e.the definition of stress-strain relationship. The sim-plest approach assumes a linear unconnected springrelationship, as per Winkler. This simplisticassumption disregards the inter-connection of thesoil elements, and these can either be modelled aswell or the simpler method used with sensitivitytests on the soil elasticity parameters. It is suggestedthat as a complex soil model is critically dependenton soil testing and interpretation, as well as its com-parison with the stiffness of the structure, it is sensi-ble to test the design against local reductions ofground support.
It is emphasised that caution must be exercisedin making the transition from traditional workingstress design methods to limit state methods whichare now general, worldwide. It is not simply a caseof adapting partial factors from one national code toanother, because the underlying principles of rein-forced concrete design may be different. The recom-mended partial factors in most national codes werederived for land-based building and bridges andrelate to broad probabilities of failure drawn fromhistorical precedent. These factors are not necessari-ly applicable to maritime structures in which themain loading cases are caused by environmental
46
loads which have to be derived from a probabilisticapproach. Similar problems relate to earth pressureloading and there is dispute over whether partialfactors are appropriate to limit state considerationsfor soils, as explained in 2.1.4. Also, it is not neces-sarily obvious whether loads are to be classified asfavourable or unfavourable.
The points raised in 2.1.4 and 2.2.3 are intended,also, to provide input to and guidance from the workof Sub-Groups A and D on the appropriate assess-ment of reliability for wave loading cases.
4.3 Durability of Concrete
4.3.1 General PrinciplesThe performance of concrete in seawater is a
subject for which knowledge and guidance remainsfragmented and ill-understood, despite the existenceof practical reports on the experience of Portlandcement concrete in the sea since 1850 andreinforced concrete from 1896. The subject issurrounded by myths and lack of understanding.The main reasons for this are that climatic andexposure conditions vary widely, different materialshave been used in various countries, and that theproperties of cement have changed during thecentury. A basic reason is, also, that deteriorationcan take a sufficiently long time such that it can bedifficult to connect cause with effect. The mecha-nisms for the deterioration of concrete structureshave not been adequately understood. It is believedthat recent work is beginning to resolve thesituation, but that a consensus view of appropriateguidance for practitioners will not be available for atleast ten years.
For this reason, the subject of concrete durabilityis the largest single clement of the Report. It is dis-cussed in more detail than other items, and containsmore reference to data which is either not generallyavailable or collated in a single document.
Whereas most forms of concrete deteriorationare now adequately covered by national codes, thisis not the case in relation to the most dramaticfailure mechanism, that is reinforcement corrosion.This can impose severe limitations in relation to thedesign of complex thin-walled structures such ascaissons or light superstructures and, in some casesthere may be little purpose in refinement of waveloading analysis if durability presents a significantrisk, albeit of a different nature. Durability is not initself a limit state but a means by which the principallimit states are maintained over the operational life.Current codes of practice deal with durability in aprescriptive manner, and do not provide a rationalbasis for design of concrete to meet a service life.Durability is not a matter of materials and choice ofmaterials, but a question of a holistic approach todesign.
4.3.2 Design Working Life (or Service Life)“Design working life” is the term and definition
from Eurocode 1, and has three main implicationsfor maritime structures:
• probability levels for wave return periods• probability levels for limit state design• time-dependent factors such as corrosion and
durability.A period of operating or service life (related to
operational and maintenance strategy) has to beconsidered by the owner of a structure and themeans of achieving this be addressed by the designer.The definitions of service life, design life and eco-nomic life require careful consideration, as there aremany different definitions in use. The maincategories of definition are compared in Table 7.For maritime structures, subject to the probabilityand return periods of environmental loading, thefollowing definition is recommended in which thedefinition of Eurocode 1 is supplemented by therider expressed in italics: “The assumed period forwhich a structure is to be used for its intendedpurpose with anticipated maintenance but withoutmajor repair being necessary within a probabilityappropriate to the function of the structure”.
Figures for design working lives specific tomaritime structures within the classification ofEurocode 1 are given in Table 8, drawn from theSpanish maritime recommendations.
It must be noted that a different level of reliabilitymay be adopted for different limit states and causesor modes of failure. Also that, despite the increasinguse of the concept of “service” life in respect ofstructural safety and durability, current codes do notgive adequate guidance for analysis to achieve suchlives. It is recommended that explicit analysis for a“design life” to satisfy the “design working life” isrequired for ensuring the durability of maritimestructures, and should be adopted in preference tothe current “prescriptive” guidance.
4.3.3 Processes of DeteriorationThe various deterioration mechanisms which
affect the durability of concrete maritime structures,the locations in which they are likely to occur, andmethods of avoidance are scheduled in Table 9. Themost widespread and critical problem is that ofchloride-induced corrosion of steel reinforcement,and the sections which follow concentrate on thisphenomenon. Adequate guidance on other forms ofdeterioration is usually given in national standards,as scheduled in Table 10.
The dominant factors involved in the durabilityof concrete, and particularly with regard to chlorideinduced corrosion are:
• recognition of the porous nature of concrete• understanding the transport mechanisms for
water and gases within the pore structure
47
• assessing the macro and micro-climaticexposure conditions for the whole structureand its individual elements.
4.3.4 Exposure classificationThe most important macro-climatic factors are
temperature and rainfall. Temperature controls therate of chemical reactions and the degree of dryingout of the cover concrete. Rainfall, humidity and thelocation of a member in relation to sea level move-ment control the wetness of concrete. The wetnessof the concrete determines the mechanism for thepenetration of chlorides and controls the penetrationof oxygen to fuel the corrosion process.
Contrary to the case of structural steel, timberand masonry, plain concrete and for freeze-thawdamage, reinforcement corrosion is less severe inthe regularly wetted tidal and splash zones. In cooland temperate climates, the concrete does not dryout to appreciable depth. However in the infrequentlywetted and mostly dry zone above the tidal zone butsubject to irregular inundation from seasonalchanges in sea level, storms etc., concrete dries outto greater depths. Especially in hot-arid areas suchas the Middle-East, and also where elements aresheltered from rain or in artificial climates such asin tunnels, the sporadic wetting of the dried-outconcrete enables chloride-laden water to be veryrapidly sucked in to greater depth by absorption.The processes of absorption, capillary suction andwick-action lead to much more rapid chlorideingress than the diffusion process which operates insaturated concrete. In a wet climate the chlorideconcentration at depth is reduced and the penetra-tion of oxygen is limited.
The proposed new Eurocode exposure classifica-tion system is explained and new suggestions forseverity ratings for concrete expressed on a scale of1 to 12 are set out in Table 11 and Fig. 13.
4.3.5 Influence of Cement TypeThe weakness of much prescriptive advice in
current codes is that guidance on mixes and -associated cover thickness to reinforcement is givenindependently of cement type. The behaviour of thevarious types of cement is compared and it isconcluded that:
• modern unblended Portland cement generallyhas the lowest resistance for chloride pene-tration and, where severe chloride exposureconditions exist, even in temperate climates,traditional thickness of cover may be inap-plicable. There are exceptions in somenational products and conditions
• blast furnace cements are highly recom-mended and have been traditionally used insome countries (originally on account of sul-fate resistance) and enable more traditional
thicknesses of cover to be used. Their toler-ance to surface scaling and poor curing is,however, less than for unblended Portlandcement
• other blending materials, such as fly ash andmicrosilica, have their benefits and limita-tions
• sulfate resistant Portland cements (i.e. withC3A less than 5%) are unlikely to be neces-sary in maritime concrete. A compromisesolution is often reached by controlling theC3A to between 5% and 10% for moderatesulfate resistance. In conditions wherereinforcement corrosion is not critical and,especially in colder waters, the long termeffects of sulfates may lead to a need for lowC3A Portland or slag cement.
4.3.6 Influence of cement contentAs is well known, the quality of a concrete mix
in relation to both strength and durability (as relatedto the pore structure) is controlled primarily by thewater-cement ratio and the unit water content. Thewater-cement ratio is therefore more important as aparameter to be specified than is cement content.The cement content is established, mainly, by divid-ing the water demand for a given mix by the water-cement ratio.
As it is desirable to use the lowest possiblewater-cement ratio to achieve durability, (generallythe requirement for durability may be more onerousthan for strength) and to reduce water movementand shrinkage effects, the cement content is con-trolled by the water content required to achieveappropriate workability.
Both water and, it follows, cement content canbe reduced by the judicious use of a range of waterreducing admixtures which still enable adequateworkability to be achieved at lower water contentwhile at the same time reducing the heat of hydra-tion consequences at higher cement content. Bothwater demand and cement content depend on thetype of cement and on aggregate size and grading.The effects of varying aggregate size from l0mm to80mm are scheduled in Table 12.
4.3.7 The Influence of CrackingThe causes and consequences of cracking have
often been misunderstood. Early thermal crackingcaused by restraint to shrinkage during cooling fromthe rise in temperature due to heat of hydration is amain cause of cracking which was previously, anderroneously, attributed to drying shrinkage.
48
Cracking caused in the plastic state can be pre-vented by good mix design, protection against dryingwinds and by good curing under arid conditions.
Most structural codes have crack width limita-tions, for flexural and early thermal cracking but thesignificance of these specific crack widths has beenover-estimated. Once crack width exceeds 0.1mmthere is no significance in relation to the ingress ofdeleterious substances. Therefore it is recommendedthat the cover to reinforcement should not bereduced for crack width reasons, despite theincrease in crack width at the surface. The reasonfor this is that, for flexural cracks, the crack width atthe reinforcement is independent of surface crackwidth, as the cracks are ‘V’ shaped, and the chlorideand oxygen ions penetrate everywhere through thepores, and not just at cracks.
Cracks do not significantly affect freeze-thawdamage as the scaling caused by freeze-thaw is,again, due to the effect of frost on water-filledpores. Cracking may, of course, be more significantin the case of unreinforced concrete if it reduces themass of armour units or blocks.
4.3.8 Influence of CuringIt appears that prolonged water curing in wet
and temperate climates may be of limited advantageand may even lead to adverse effects such asthermal shock. It may be essential in hot and aridclimates.
As the duration of curing is inversely propor-tional to water-cement ratio, adoption of a lowwater-cement ratio enables the curing period to bereduced.
4.3.9 Monitoring and MaintenanceInadequate guidance on this strategic topic is
available in the literature and national codes, but itappears that it is, at last, receiving more attention.Regular inspections should be carried out at leastonce per year, most likely following the winterstorm period. The principal objects of the survey areto determine:
• the integrity of armour units and elements ofthe structure
• indication of movement and settlement• scour.It is essential to record “base-line” measure-
ments of line and level immediately on completionof construction. This should include “as constructed”measurements of cover to reinforcement and crackand damage mapping. Computers, underwater videorecorders and corrosion measurement devices cannow be utilised.
4.4 Materials
4.4.1 Rock and RubbleReference is made to the CIRIA/CUR manual on
the use of rock in coastal and shoreline structures,1992.
4.4.2 Filling and BackfillingCurrent requirements are outlined in 2.4.2,
including the recommendation that measures may benecessary to increase the density of infill material.
4.4.3 Concrete - General. - Design, Detailingand Workmanship
The requirements for achieving durability ofconcrete in maritime works will usually outweighthe requirements for achieving strength and notemust be taken of the factors outlined above. Manyproblems can be “designed out” by good detailingand specification. Reliability of both design andworkmanship can be improved and “gross errors”avoided by employing the discipline of QualityAssurance and Quality Control audit procedures forboth design and workmanship.
4.4.4 Unreinforced ConcreteThe factors affecting unreinforced concrete
(more usually termed “mass concrete” in Europeand UK) primarily concern deterioration of theexposed surface and include freeze-thaw, abrasion,and sulfate attack. These forms of deterioration havesimilar and overlapping effects, and are described inmore detail in 2.4.4. Early-thermal design is impor-tant for crack avoidance.
Table 14 sets out suggestions for the choice ofwater-cement ratio, and hence minimum cement-itious contents and grades, for various cement typesand blends and minimum dimensions of pour, forthe range of exposure ratings on a scale of 1-12suggested in Table 11 and Fig. 13.
4.4.5 Reinforced Concrete, including theSelection of Cover to Reinforcement
The durability of reinforced concrete requiresconsideration of the same factors which affectunreinforced concrete, together with the majorphenomenon of chloride-induced corrosion. Themain conflict point in the design and productionprocess is the selection of and practical achievementof the appropriate cover to reinforcement. This ismore a design and detailing matter than a materialsmatter. As the protective power of a given concrete isbroadly related to the square of the cover thickness,the provision of appropriate cover is the simplest
49
and most positive way of reducing corrosion dam-age. The cover to be specified is influenced by theexposure severity rating, the cement type, the mixquality as determined by the water-cement ratio andthe placing tolerance which can be achieved.
The minimum cover considered necessary forcorrosion protection should be regarded as a“characteristic” value and a margin of at least 10mmto 15mm should be added to the figure in order toreduce the rate of failure to achieve the characteris-tic value to within 5%, by analogy with concretestrength compliance. Without this margin it isstatistically impossible to achieve the necessarycover as, in practice, the variation in position ofreinforcement about the mean position exceedscommon perception, as illustrated in Fig. 18. Sug-gested nominal cover for different severity ratings,cementitious materials and qualities is scheduled inTable 15. A nominal cover of 50mm is the lowestpracticable figure and is only suitable for the lowestseverity rating and using blended cements.
Nominal cover thicknesses between 75mm and100mm have to be considered as normal. For severeexposure combined with unblended Portlandcements, it may be necessary to double the cover.Methods of specifying cover in relation to a charac-teristic value, with separate acceptance criteriabefore concrete is placed and after concreting, aresuggested in Table 16.
The selection of appropriate cover and fixingtolerance has a major influence on the memberthickness for caissons and thin precast units. Min-imum member thicknesses of 200mm or 300mm aredifficult to achieve under aggresssive conditions.Under such conditions it may be necessary tochange the member type and section in thevulnerable upper tidal and splash zone, and either toadopt increased cover or, if feasible, change to plainconcrete in this zone.
4.4.6 Prestressed ConcretePrestressed concrete is less common for
maritime works but there is a substantial history ofsuccessful use in North Sea oil structures and inrecent caisson breakwaters. Reasons for successinclude the necessity for higher quality concrete andlarger cover, on structural grounds, and the highstate of saturation in cold and wet climates.Prestressing reduces the problem of cracking butdoes not provide any help in reducing the ingress ofchloride ions and oxygen through the body of theconcrete. The corrosion rate of prestressing steel isunderstood to be faster than that of lower stressedsteel and the detailing of secondary reinforcementpresents the same problems as ordinary reinforcedconcrete.
4.4.7 CementRefer to 4.3.5 and 4.3.6
4.4.8 AggregatesThe quality of aggregates has a lesser impact on
the strength and quality of concrete than may besupposed, but it has an influence on abrasion,freeze-thaw and salt scaling resistance. It is usuallynecessary to make the best use of locally availablematerials. National standards usually provide suffi-cient guidance.
4.4.9 Crack widthRefer to 4.3.7 and 4.5.6
4.4.10 Reinforcing SteelProvisions for reinforcing steel are adequately
covered by national standards. In addition to normalblack steels, various coating protection systems arealso available, but have not been greatly used inmaritime works and may be at a disadvantage inaggressive situations.
Stainless steels have not often been used butmay be appropriate provided the correct grades areused. The price penalty of six times that for blacksteel is less when viewed as a proportion of overallcost or the cost of premature failure. Steel fibre andnon-metallic fibber and strand reinforcement appearspromising.
4.4.11 AdmixturesAdmixtures are covered by national guidelines.
Their use should be positively encouraged in orderto provide adequate workability with low watercontent mixes.
4.4.12 Protective Measures such as Coatings,and Cathodic Protection
In aggressive conditions only high-build, highquality coating products are dependable, which areexpensive and require the concrete surface to beprepared to a high standard by grit blasting andother means. Cathodic protection is not regularlyused for the protection of new construction, althoughin some cases allowance for later implementationare made by ensuring continuity of reinforcementand facility for electrical connection. There isgrowing experience of its use as a repair techniquewhere simpler alternatives are not feasible.
4.4.13 Corrosion of Structural SteelThe corrosion performance of structural steel in
maritime conditions is much better known that thatof reinforcing steel embedded in concrete. Thecorrosion rate is usually higher in the splash zoneand at low astronomical tide levels, and very low indeep water. Either an extra thickness of metal as a“corrosion allowance”, high duty coating or cathodicprotection can be used. Successful traditional coat-ings, however, may no longer meet environmentaland health and safety regulations for application andthere is, as yet, inadequate experience with somenew water-based systems.
50
4.5 Construction Related Criteria andMethods - Caissons
4.5.1 Sea Condition Data and Limits forConstruction Risks
Wave forecast data must be available to enablethe designer and the contractor to estimate the seastate at each stage of construction, especially forfloat-out, grounding and filling of caissons. Thisdata should include information on the proportion oftime during a year in which certain wave heights arenot surpassed and the length of windows for calmweather.
4.5.2 Construction JointsConstruction joints are an important feature of
vertical breakwaters. Horizontal joints seldom giveproblems. Vertical joints are necessary to allowdifferential settlement to occur between adjoiningelements, but at the same time the interconnectionof elements is required in order to distribute the loadfrom local wave attack over more than one element.Typical examples of jointing methods are illustratedin Figs. 21, 22 and 23.
4.5.3 SettlementThe magnitude of settlement observed for
vertical breakwaters is higher than for most otherforms of construction, for reasons explained in3.1.6. Settlement is rarely critical but the range oflikely settlement and differential settlement shouldbe anticipated, and the visible effects of settlementshould be minimised by appropriate detailingfeatures.
Examples of measured settlement are quoted,including absolute settlement of up to 1.5m and adifferential of 0.2m. The contribution to settlementcaused by the time-dependent consolidation of rock-fill for many years after construction is notgenerally appreciated. Examples are given using thelogarithmic expression published by Penman andothers.
4.5.6 Early Thermal CrackingThermal contraction from the heat generated by
hydrating cement results in severe cracking whereverthe size or geometry of the section or the sequenceof adjacent pours during construction imposes arestraint to free contraction. Control of the tempera-ture and temperature gradient in large unreinforcedelements is essential to restrict cracking. The formsof restraint and expressions for calculating the rein-forcement required to restrict cracking to acceptable
orders of magnitude in monolithic reinforcedconstruction are given. Measures to control the heatof hydration rise are outlined. It is recommendedthat the early thermal design is checked against theactual cement contents used in the works, as thesemay exceed values assumed in design. Failure tomake this check often results in problems.
4.5.7 SlipformingSlipforming is a suitable method for casting
caissons. It is a method which lends itself to a high-rate of production. Because the concrete is castjoint-free and hardens free of restraint, problems ofconstruction joints and early thermal cracking areminimised. Practical guidance available in the liter-ature is given and problem areas identified.
4.5.8 CuringThe need for and relevance of wet curing in all
circumstances is discussed in 4.3.8 and 2.3.8 and, inthe case of massive caissons, introduces severelogistical problems. For caissons, the timing ofimmersion has a great influence on the rate ofproduction. If the concrete is saturated, and of asufficiently high grade which can be cured in a shortduration, early immersion may be beneficial.Concrete left to thoroughly dry in the surface layersin arid climates can lead to disastrous absorption ofsalts into the concrete upon immersion.
4.5.9 Development in CaissonsSome recent developments of composite and
prestressed concrete caissons in Japan are illustratedin Figs. 28, 29 and 30.
4.6 BlocksBlockwork is an inherently durable method,
especially for constructing quay walls, which hashistorically been used for vertical breakwaters.Guidance is given in Japanese, Spanish and UKdocuments. A degree of cracking is permitted inquay walls but could endanger breakwaters. Typicaltypes of blockwork for modern breakwaters areillustrated in Figs. 30, 31, 32, 33 and 34. Theproblems common to blockwork are outlined.
4.7 Curtain and Pile Type Breakwaters Curtain and pile type breakwaters are effective
in shallow depths of water with a low wave heightclimate, and for a soft sea bed. Various types ofmodern single and double wall breakwaters areillustrated in Figs. 35 and 36.
51
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