Subgroup C Report

PIANC PTC II WORKING GROUP 28

Recommendations for the

Construction of

Breakwaters with

Vertical and Inclined Concrete Walls

Report of Sub-Group C Final Report . . . July 1997 Issued 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 Puertos Infrastructure Department, Gijón Port Authority, Asturias, Spain. Dr. O. Kiyomiya, PhD Civil Engineering The Port and Harbour Research Institute, Japan. Mr. F. Ropert, ITPE Service Technique Central des Ports Maritimes et des Voies Navigable, France. Mr. B. N. Sharp, MSc, C.Eng, FICE Sir 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, UK Chairman, BS 8002 code drafting committee Mr. P. Chubileau, CSP Service Technique Central des Port Maritimes et des Voies Navigable, France Dr. F. T. Christensen Applied Science Associates, USA Dr. T. A. Harrison, BSc, PhD, C.Eng, MICE, FICT British Ready Mixed Concrete Association, UK Professor (Associate) D.Eng, P.A. Hedar Gothenburg, Sweden Dr. B. Simpson, F.Eng, Eur Ing, MA, PhD, FICE Ove Arup and Partners, Arup Geotechnics, UK Mr. G. Leone, BSc, C.Eng, MICE Sir William Halcrow and Partners Ltd., UK Mr. E. Longaygue, ITPE Service Technique Central des Ports Maritimes et des Voies Navigable, France. Dr.-Ing. J. Schwarz Hamburg Ship Model Basin, Germany Mr. D. Slater, BSc, C.Eng, MICE, MIStructE Sir William Halcrow and Partners Ltd., UK Professor G. Somerville, PhD, C.Eng British Cement Association, UK Mr. D. Wimpenny, BSc, MPhil Sir William Halcrow and Partners Ltd., UK (i)

Contents Page 1

INTRODUCTION AND TERMS OF REFERENCE

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 by Sub-Group A

2.1.1 2.1.2 2.1.3 2.1.4 2.1.5 2.1.6 2.1.7 2.1.8

Earthquake ............................................... Ice Pressure.............................................. Ship Collision (deleted as insignificant).. Earth Pressures ........................................ Fill Pressures............................................ Friction..................................................... Handling and Float-Out Loads ................ First Grounding........................................

2.2

Resistance Analysis, Internal Analysis

2.2.1 2.2.2 2.2.3

Structural Analysis in Element Design.... 13 Scale Models............................................ 13 Limit State Design and Risk Analysis ..... 14

2.3

Durability of Concrete

2.3.1 2.3.2 2.3.3 2.3.4 2.3.5 2.3.6 2.3.7 2.3.8 2.3.9

Durability, Introduction ........................... Design Working Life (or Service Life) ... Processes of Deterioration ....................... Exposure Classification ........................... Influence of Cement Type ....................... Influence of Cement Content................... Cracking and the Influence of Cracks ..... Influence of Curing.................................. Monitoring and Maintenance...................

2.4

Materials

2.4.1 2.4.2 2.4.3

Rock and Rubble ..................................... Filling and Backfilling............................. Concrete Durability, General Design, Detailing and Workmanship....... Unreinforced Concrete (Plain or Mass)... Reinforced Concrete including Selection of Cover to Reinforcement ...... Prestressed Concrete................................ Cement.....................................................

2.4.4 2.4.5 2.4.6 2.4.7

Page 2.4.8 2.4.9 2.4.10 2.4.11 2.4.12

Aggregates ............................................... Cracking and Crack Width ...................... Reinforcing Steel ..................................... Admixtures .............................................. Additional Protective Measures: Coatings, Coated Reinforcement, Cathodic Protection ............................. 2.4.13 Corrosion of Structural Steel ...................

2 3 4 4 8 9 11 12

3

CONSTRUCTION RELATED CRITERIA AND CONSTRUCTION METHODS

3.1

Caissons

3.1.1 3.1.2 3.1.3 3.1.4

Float-Out Loading ................................... First Grounding........................................ Caisson Fill Methods and Pressures ........ Sea Condition Data and Limits for Construction Risk .................................... 3.1.5 Construction Joints .................................. 3.1.6 Settlement ................................................ 3.1.7 Early Thermal Cracking .......................... 3.1.8 Slipforming .............................................. 3.1.9 Curing ...................................................... 3.1.10 Developments in Caissons.......................

15 16 17 19 21 22 23 24 25

26 26 28 30 30 (ii)

32 33

34 34 34 34 34 35 37 39 40 40

3.2

Blocks

3.2.1 3.2.2 3.2.3

Blocks from Concrete .............................. 42 Types of Concrete Blocks........................ 42 Common Problems .................................. 42

3.3

Rubble Mounds...................................... 44

3.4

Curtain and Pile Type ........................... 44

4

SUMMARY

4.1

4.6. 4.7

Different Loadings not covered by Sub-Group A............................................ Resistance Analysis, Internal Analysis.... Durability of Concrete ............................. Materials .................................................. Construction Related Criteria and Methods - Caissons.................................. Blocks ...................................................... Curtain and Pile Type Breakwaters .........

5

REFERENCES ...................................... 52

4.2 4.3 4.4 4.5

26 26

31 32 32 32

45 46 47 49 51 51 51

1.

INTRODUCTION AND TERMS OF REFERENCE

1.1

Methodology for the Recommendations of Sub-Group C

The task of Sub-Group C was drawn up at the Meeting of the Working Group in Hannover in February 1993:“Investigations into the implication of construction aspects in design. Performance of concrete. Identification of “hot spots” in design and construction”. By consultation with the members of Working Group 28 and, especially amongst the members of Sub-Group C, tasks and topics were assembled which Sub-Group C undertook to study. The SubGroup decided that one of its tasks was to identify significant loading cases and items related to design, structural analysis and construction, which may not be covered by the predominantly wave loading and analysis considerations of the other Sub-Groups. The list was enlarged and changed several times. The main difficulty encountered in drawing up a framework for the topics was that a clear structure for a consistent arrangement of all recommendations could not be established. It is unavoidable that certain subjects have to be repeated, because an exclusive relation of one item to a single heading does 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 existing 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 July 1994, Gijón, Spain in February 1995, and London, UK in July 1995) and by correspondence, the lists of topics were identified and initial reports were prepared by each member on specific topics. It was agreed which topics would be studied by which members and, subsequently, these were developed into several stages of draft text. The final draft Sub-Group Report was completed in October 1995 and issued to the Sub-Group members and, by the Chairman, to the Main Working Group 28. After receipt of the further comments arising from the Sub-Group members and comments of the main Working Group at the meetings held in Berlin, in March 1996, and in Delft in September 1996, the Sub-Group Report was updated for issue as a Final Sub-Group Report in July 1997.

The members of the Sub-Group and those external authorities who were consulted during this period (either within the Working Group or outside PIANC) are listed in the opening pages of this Sub-Report.

1.2

General Headings

The topics selected for consideration were as follows:-

DESIGN CRITERIA AND MATERIALS

• • • •

Different loadings not covered by Sub-Group A Resistance analysis, internal analysis Durability and maintenance Materials.

CONSTRUCTION RELATED CRITERIA AND CONSTRUCTION METHODS

• • • •

1.3

Caissons Blocks Rubble mounds Pile and Curtain type.

Format of Recommendations

The recommendations of Sub-Group C, where possible, consist of information mostly drawn from the experience and studies of its members, references, and brief summarised recommendations. The recommendations are brief or given as a reference, except where it is considered that it may be more helpful to give more detail. According to a decision at the Meeting of the Main Working Group 28 on the 26th of April 1995 in London, it will be left to the Main Working Group to extract from the broader formulated recommendations of the SubGroups, those statements which should be included in the Recommendations of Working Group 28. A summary of the main recommendations is given at the end of the Report, as Section 4. In accordance with the practice of PIANC, the Main Group Report will be circulated to members. Copies of Sub-Group reports are available on application to PIANC. 1

2.

DESIGN CRITERIA AND MATERIALS

2.1

Different Loadings not covered by Sub-Group A

As a general rule, it is considered that the earthquake load acts horizontally at the centre of gravity of the structure. The vertical component of earthquake load is not usually considered. The seismic coefficient is obtained as follows: Seismic coefficient = Regional seismic coefficient x factor for subsoil condition x importance factor In the Japanese Technical Standards, the factors are as follows: • regional seismic coefficient is 0.05, 0.10 or 0.15 • 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

2.1.1. Earthquake The closest understanding of the behaviour of breakwater caissons during earthquakes can, as for structures above ground, be obtained by dynamic response analysis. Due to advances in computer techniques, this can now be achieved by finite element methods. The caisson, rubble mound and soil foundations and surrounding water should be considered in the modelling. This model has not been applied generally in design, except for very important breakwater caissons located in a vigorous earthquake activity zone. The response of the model is illustrated in Fig. 1. m

armour stone

b l o c k

(a) Model

(b) Vibration Mode

Fig. 1 Earthquake dynamic response model (ref The Port and Harbour Research Institute - Japan)

Generally, however, the simple equivalent static load method is appropriate for breakwater structures because the seismic force has a relatively short period of natural vibration, with heavy damping. In accordance with the Technical Standards for Ports and Harbour Facilities in Japan, 1991, the earthquake load acting on a breakwater located in a seismic activity area should be calculated by the following formula:

•

Earthquake load = (dead load + surcharge) x seismic coefficient.

The seismic coefficient is determined by taking the regional probability of occurrence of an earthquake, the condition of the foundation soil and the importance of the structure into consideration. 2

forces and therefore earthquake forces can be neglected in the design, except in the case of very large breakwater caissons located in areas of vigorous seismic activity. The stability of the foundations must also be checked for earthquake loading. Even in the major Hansin earthquake in Japan, in January 1995, the breakwater caissons did not slide or collapse. The most significant effect was settlement of the caissons, due to liquefaction of loose sand in the foundation strata. Seismic forces should also be taken into consideration for piers, jetties, etc. The vertical component is considered only for an earthquake with a narrow epicentre or for structures located near a fault. In this case, the vertical component coefficient is usually adopted as half the horizontal component.

Subsoil conditions are divided into three classifications : 1st, 2nd and 3rd kind. The classification of the subsoil condition and hence the subsoil condition factor depends on the thickness of the quaternary deposit and the types of subsoil condition, as given in Tables 1 & 2. The factor relating to the importance of the structure is classified in four categories: Special class, A class, B class and C class, as in Table 3. TABLE 1 CLASSIFICATION OF SUBSOIL CONDITION (ref Technical Standards for Ports and Harbour Facilities in Japan) Thickness of quaternary deposit

Gravel

Sand or Clay

Soft Ground

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 2 SUBSOIL CONDITION FACTOR (ref Technical Standards for Port and Harbour Facilities in Japan)

Classification

Factor

1st kind

0.8

2nd kind

1.0

3rd kind

1.2

TABLE 3 CLASSIFICATION OF IMPORTANCE FOR EARTHQUAKE (ref Technical Standards for Port and Harbour Facilities in Japan) Classification of Structure

Characteristics of Structure

Importance Factor

Special Class The structure has significant characteristics covered by items (1)-(3) in A Class.

1.5

A Class

(1) If the structure is damaged by an earthquake, a large number of human lives and property will possibly be lost. (2) The structure will perform an important role in the reconstruction work of the region after an earthquake. (3) The structure handles hazardous or dangerous activities with risk that the damage to the structure will cause a great loss of human life or property. (4) If the structure is damaged, economical and social activity of the region will suffer severely. (5) If the structure is damaged, repair work will be difficult.

1.2

B Class

The structure is other than Special, A or C Class.

1.0

C Class

The structure is small and easy to repair, excluding structures which fall in the Special or A Classes.

0.5

2.1.2 Ice Pressure According to Schwarz J, and Christensen F T, there are still no formulae which can be recommended as being conclusive and reliable in the calculation of ice pressures for all cases. The following guidelines should contribute a better understanding of ice pressures for major works. In countries with long frost periods, vertical breakwaters should be designed to resist ice pressure, which can arise from different causes: a) pressure caused by the closed cover of ice in a harbour, when the cover expands with rising temperature b) pressure caused by ice fields drifting parallel to the coast with tidal current or littoral drift c) vertical pressures caused by piling up processes. As a rule, ice pressure is not considered for sloping breakwaters in Europe (Hedar P A, 1995) but is recommended for America (Christensen F T et al, 1995, Bruun P, 1985). For vertical breakwaters, assessment of ice pressure is a matter of judgement depending on the ice conditions and the possibilities of ice-structure interaction. It is doubtful if the pressures given for concrete dams for fresh water lakes, such as 300 kN per m2 and 50 to 200 kN per metre length according to Hedar P A, 1995, or typical pressures of 200 to 400 kN per metre with extremes of 400 to 600 kN per metre according to Christensen F T, 1996, can be applied to vertical breakwaters in sea water Small scale experiments on ice forces on vertical structures carried out at the Iowa Institute for Hydraulic Research have resulted in the following formula (ref Schwarz J, 1994). σeff = 0.564 . d-0.5 . h0.1 . σc where σeff = effective indentation strength in MPa d = structure width in m h = ice thickness in m σc = maximal compressive strength of an ice prism at a strain rate of 0.003 per second in MPa. (For a prism length of 0.2 m the deformation speed must be 0.0006 m/s). For instance for the Baltic Sea, a value of σc = 1.8 MPa is feasible according to EAU 1990. According to this formula the effective ice pressure decreases with the square root of the structure width. This finding is in accordance with Sanderson’s investigations (ref. Sanderson T J O, 1988) which however do not consider the effect of the ice thickness, as illustrated in Fig. 2 (overleaf). The Iowa formula (Schwarz J, 1994), which still has to be proved for larger structures, has received support from full scale measurements in China and in the Baltic as well as recently by Canadian engineers (Fitzpatrick J, 1994 and Kennedy K, 1996). 3

INDENTATION PRESSURE MPa

principles applying to structural design are discussed in 2.2. Traditional “working stress” codes recommend "active" or "at-rest" pressure coefficients to apply to the dry or submerged soil mass, as appropriate, as a working (or “characteristic”) load. The “fully active” coefficient is usually applied to blockwork walls, where slight rotational movement can be tolerated, and can be applied for the overall stability of caissons Fig. 2 Effective ice pressure versus pressure area as, for example, in Spain. (ref Sanderson T J O, 1988) However, the “at-rest” coefficient is usually applied to These last publications conclude that ice prescalculate the overall stability of rigid structures such sures for large works have been overestimated in the as caissons and to calculate the sizes of the structural past by a considerable factor, as illustrated in Fig. 3. members. If the fill is to be compacted aggressively, According to Schwarz J, 1996, and Christensen F T, it could be necessary to consider compaction 1996, a validation study is planned in an EU pressures, i.e. pressures greater than "at rest". This research project. “working load” is applied in conjunction with suitable factors of safety for overall static (equilibrium) stability calculations, and to derive Then and Now working structural material stresses. “Passive” Beaufort global and local ice loads pressures are calculated at the ultimate, and divided Load(tonnes) (tonnes) Pressure (MPa) Load Pressure (MPs) by a relevant factor of safety. The Japanese code (ref 1 500 000 Technical Standards for Port and Harbour Facilities 15 in Japan, 1991) still continues this approach. The Local German Waterfront Structures Code (ref EAU 1990) has now introduced the concepts of limit states in 80 accordance with draft Eurocode 7 while still retaining traditional methods as a permitted option. Due Global to the difficulties of applying limit states to 100 000 foundation problems, traditional methods remain an 1980 1985 1990 1995 option in other new codes, such as BS 8002 (ref BS Fig. 3 Development of ice force predictions 8002, 1994). (ref Fitzpatrick J, 1994) Most structural codes now adopt the limit state philosophy, although the application of limit states 2.1.3 Ship Collision (deleted) to earth pressure and variable water loading for maritime structures is not as straightforward as for The effects are not significant, and only relate to standard structural loading for buildings and small vessels. bridges. The selection of partial factors to match with and correspond to the reliability and failure 2.1.4 Earth Pressures probability of wave and water loading is the task of Sub-Groups A & D. See also 2.2.3. This section concerns the principles of earth One of the main difficulties in bringing together pressure in general and external to structures such as earth pressure theory and limit state design of caissons (see Fig. 5). Earth pressure in filling to structural members is the fact that earth pressures in caissons is considered in 2.1.5. The general case retained materials reduce as movement increases. In must be considered first, although a breakwater will this respect, earth pressures differ from almost all have fill placed against it only if it is protected by a other types of loads considered in structural design. rubble mound on the exposed side, or it retains In addition, the failure mechanisms and pressures reclamation on the lee side. The water pressures to are different for flexible steel sheet piling and rigid be used in conjunction with earth pressures will, in reinforced concrete or masonry walls. The problem the case of breakwaters, be a function of wave and is that, in geotechnical terms, the soil load at ultimate tidal variation loading, although on the lee side and (failure) is often less than at working conditions, internally they may only relate to tidal effect. Wave whereas for structural design of the reinforced loading is not considered here. The general 4

similar results to the old working stress codes. See Fig. 4 (ref Daniels R J and Sharp B N, 1979). However, some codes achieve a similar reduction by the use of much lower partial factors applied to the serviceability limit state than are used in other current structural codes. See below and Fig. 6(d) and 6(g). British Standard 8002 : 1994, Code of Practice for Earth Retaining Structures, incorporating the limit state approach, has only recently been issued. BS 8002 adopts limit state philosophy but it does not, in fact, involve partial factors. It reduces tan Ø' by a “mobilisation” factor which operates in a similar way to a partial factor and can be imagined as a partial factor. However, BS 8002 does not deal with maritime structures (ref Akroyd T N W, 1996 and Bolton M D, 1996). Alternative limit state methods are incorporated in Eurocode 7 : Part 1:1993. Eurocode 7 has two appropriate cases, B and C. Case B derives straight from structural design, and Case C comes from geotechnical stability analysis. A comparison of various applications of partial factor methods for the calculation of structural members is given in Figs. 5 and 6. A breakwater caisson with sand fill on the inside (land side) is illustrated in Fig. 5. Typical lateral pressure diagrams are given for the loading on the buried face, i.e. the components of soil above water level, surcharge, submerged soil and water. To illustrate the comparison of methods in Fig. 6, the cases of submerged soil and water loading with water level at a MSL of +3m, only, are considered, all as unfavourable loads. Tidal variation of water level, wave loading and surcharge are not considered here

concrete section the designer seeks to apply a factored higher loading for the ultimate case. The serviceability limit state earth pressures are usually calculated directly from the characteristic density and strength properties and the appropriate earth pressure coefficient. The ultimate limit state pressures are usually calculated by modifying several of the main parameters involved from their expected values, by the application of partial factors. Sometimes the partial factor is applied directly to the serviceability limit state pressure, and sometimes it is applied to a parameter, such as tan Ø'. In some current codes, for example, such as British Standard 8110 : 1985, and the Spanish Maritime Works Recommendations ROM 0.2-90 and ROM 0.5-94, the ultimate limit state pressures for structural design are calculated by applying partial factors to the serviceability limit state pressures. However, for many years the Danish and other Scandinavian codes adopted the less conservative principle of applying a partial factor to tan Ø' thus deriving a smaller value of Ø' and hence a larger value of the applied earth pressure, together with alternative factors on the water pressure. This latter approach to the ultimate limit state earth pressure for structural design can produce a much smaller difference between the serviceability and ultimate limit states than the classical working stress approach, or if the partial factors from structural codes are applied to the characteristic (i.e. working or serviceability) pressures. Note that the partial factor for material strength will also be applied, and that the total ULS using direct factors from structural codes has been calibrated to give Surcharge

+4.5m Dock deck level

+4 +3.0m WL

+2 Fill and water levels to same vertical scale

0

Reduced level : m

-2 Ultimate limit state using factors of 1.15 on tan∅1 and 1.2 on water Ultimate limit state using factors of 1.5 on surcharge and 1.35 on soil and water as Eurocode 7 Case B i.e. similar to working stress results excluding the material factor.

-4

-6

Fig. 4 Caisson walls : Design wall pressure distribution (Loading on one side only -i.e. all loads have been calculated as unfavourable loads on the loaded side and air on the other side.)

-8 Serviceability limit state

-10

(Adapted from: Daniels R J and Sharp B N, 1979)

-12 -13.29 0

50

100

150

200

250

Horizontal wall pressure: kN/m2

300

350

400 effective factor approx. 1.2 effective factor approx. 1.35

5

on appropriate factors for water pressure for structural design. A similar result is given by Case C of Eurocode 7, with a factor on tan Ø' of 1.25 with, again, inadequate guidance on water pressures to apply to structural elements in conjunction with this soil loading. An answer to this serious ommission has been suggested by Cole & Watt (ref Cole & Watt, 1994). They suggest that the earth pressure calculated from the reduced tan Ø' should be treated as a “worst credible load” to BS 8110 : Part 2, 1985, and therefore both this and the water pressure be multiplied by a factor not less that 1.1 or 1.2 for structural design of the section. This suggestion is refuted by Akroyd (ref Akroyd T N W, 1996) and, in comparison with other “limit state” codes, it would indeed appear incorrect to apply a further factor to the earth pressure component. However, the suggestion to apply a factor of 1.2 to the water component at least produces a comparable philosophy to some other codes, and is illustrated in Case (f). Case (g) illustrates proposed Japanese factors (ref Kiyomiya O, 1994) which are similar to cases (b), (c) and (d) but with a lower general factor of 1.1, and a higher Ko of 0.6. The factor for variable loads is 1.3. It can be noted from the above that higher, more traditional, ratios of ULS loads to serviceability limit state loads for the calculation of member sizes are given by structural codes such as BS 8110, Eurocode 7 Case B, and ROM 0.2-90, although ROM 0.2-90 has the facility for varying environmental load factors. The less-conservative loading from the USA, earlier Scandinavian, BS 8002 and Eurocode 7 Case C, and Japanese methods, are all of a similar magnitude.

but the appropriate worst case situations of water level will have to be considered in a specific design. Case (a) in Fig. 6 shows the “at-rest” characteristic pressure (serviceability limit state) for submerged sand and water for a 20m depth of fill, for the soil properties given in Fig. 5. Only granular material is considered. Cases (b), (c) and (d) show the ultimate limit state pressures for structural design calculated by the direct application of partial factors to case (a). In Case (b) to the British structural code BS 8110 : 1985, the submerged sand pressure is multiplied by 1.4 and water, as an adverse (i.e. not a reducing) load, is also multiplied by 1.4. The factor for variable loads is 1.6. Case (c) shows the recommendations of the Spanish ROM 0.2-90 1990, which uses a partial factor of 1.35. However, variable water loads based on statistical data can be factored by 1.0, and other variable loads by 1.5. Similar rules apply to Eurocode 7, Case B. Case (d) shows a similar USA application noted in the PIANC report on floating breakwaters (ref PIANC, 1994), where the standard partial factor is 1.2. In this application, design wave loading can be used with a factor of 1.3, or 1.6 on a yearly return period. Cases (e) and (f) give examples of applying the partial factor to tan Ø'. In (e), to earlier Scandinavian rules, the worst of two cases is taken, i.e. a factor on tan Ø' of 1.3 combined with water (adverse) at 1.0, or a factor on tan Ø' of 1.15 plus water (adverse) at 1.2. In (f), the method of BS 8002:1994 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

Sea and Wave side (Front side) + 3m msl

Breakwater details as relevant

Fill

Reference water level for examples only. Different water level cases will be necessary for design Land side

Fill

- 17m

Characteristic ∅' ∅' Density γ γ

SOIL PROPERTIES above water 33° submerged 30° submerged 9.5kN/m3 sea water 10.2kN/m3

Soil above water level

Surcharge (Not considered inFig.6)

Submerged soil.

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)

6

Water

+ 3 m Submerged Sand Ko = 0.500 Wall friction O°

Soil and Water

Water

-17m 95 kN/m2

204 kN/m2 (a) At-Rest characteristic pressures of submerged sand and water only. Serviceability limit state.

Submerged Sand (a) x 1.4

133 kN/m2

299 kN/m2

Water (Adverse) (a) x 1.4

(a) x 1.4

286 kN/m2 (b) Ultimate limit state to BS8110 structural factors. (Variable loads 1.6)

(a) x 1.35

419 kN/m2

(a) x 1.35

(a) x 1.35

128 kN/m2 275 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 1.2

114 kN/m2

(a) x 1.2

(a) x 1.2

245 kN/m2 (d) ULS to USA (Ref in text) 1.2 (DL & LL) + 1.3 Design Wave.

Case (i) Factor on tan ∅' 1.3 Ko = 0.594 plus water at 1.0

Case (i) water

Case (ii) Factor on tan ∅' 1.15 Ko = 0.551 plus water at 1.2 105

403 kN/m2

113 kN/m2

Total case (i)

204 245 kN/m2 (e) ULS to Scandinavian factors. Case (ii) rules here

Eurocode 7 Case C Factor on tan ∅' 1.25 Ko = 0.581 108 110

(a) x 1.1 with Ko = 0.6

Total case (ii)

Case (ii) water

317

Water to BS 8002 ??????

BS 8002. Mobilisation Factor on tan ∅' 1.2 Ko = 0.566

kN/m2

359 kN/m2

Water to Eurocode 7 Case C ?????? Water at 1.2

Total BS 8002 Soil + Water at 1.2

Total Eurocode Case C Soil + water at 1.2

245 kN/m2 (f) ULS to BS 8002 and Eurocode 7 Case C. (a) x 1.1

350 kN/m2

353

355 kN/m2

Total

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.)

7

Water pressure is an important loading case to be applied to external water pressure for floating caissons, and to variable water levels in the soil as affected by wave and tidal effects. Different factors need to be applied to the water pressure, depending upon whether the water pressure is considered to be assisting or resisting the earth pressure. As water pressure is much larger than soil pressure in the case of granular materials, there would seem to be little point in over-refinement of the earth pressure fraction of the total load, unless much clearer guidance and rational partial factors are also given to variable water loading. As a general recommendation, results such as given by the USA method with a general factor of 1.2, the Japanese method, or by factors on tan Ø' such as the Scandinavian method, BS 8002 and Eurocode 7 Case C, provided a factor of at least 1.2 is applied to water pressure, appear to be similar and appropriate for the analysis of both overall stability and member strength. Some authorities, however, would recommend the higher, conventional, structural factors to be used for member strength. The Hong Kong Geoguide 1 (ref Hong Kong Guide to retaining wall design) advocates this course. It gives partial factors similar to BS 8002 for soil equilibrium, but advises that structural members be calculated using unfactored soil properties in conjunction with the relevant structural code. The Hong Kong Guide does 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 and filling methods.

2.1.5

Fill Pressures

The loading within caisson cells is generally derived from silo theory. Practical verification of this approach is available from measurements in caisson cells in 1976/77 (ref Daniels R J and Sharp B N, 1979) which demonstrated that the Janssen silo distribution of pressure with a wall friction of 20° compared closely with the measured pressures, Fig. 7. The various methods for calculating internal pressures are illustrated in Figure 8. As in 2.1.4, submerged sand to 20m depth is used to illustrate the principles. Case (a) shows the characteristic at-rest unconfined pressure, for the materials given in Fig. 5. Case (b) shows calculations to DIN 1055 Part 6:1987, ROM 0.2-90 and ACI 313:1983, all of which are based on the Janssen method, applied in the filling condition. Eurocode 7 refers , for silo pressures, to Eurocode 1 Part 4, which is presumably based on the same method. ACI 313 adopts the active pressure coefficient (0.333 in this case) and overpressure factor. However, all of these methods give the same result with an at-rest coefficient of 0.5. As shown in case (c), the Technical Standards for Port and Harbour Facilities in Japan, 1991, recommend a fixed at-rest earth pressure coefficient of 0.6, which is applied to a depth equal to the width of the cell. Below that 45° line, the pressure is constant on a silo basis. This calculation applies up to cell dimensions of 5m x 5m, above which standard silo pressures are applied. The results are simpler and not greatly dissimilar from the others in case (b).

+4.5 m Final pavement level

+4 + 1.5 m Fill level (partially filled)

+2

+1.0 m WL

0

Fill and water levels to same vertical scale

Note: Total wall pressures were measured with pressure cells.

Reduced level : m

-2

Effective earth pressure = total measured pressure minus static water pressure

Effective earth pressure (measured)

-4

Total wall pressure (measured)

-6

-8

Fig. 7 Caisson walls. Comparison between measured and predicted wall pressures within a caisson cell

Difference = water pressure

-10

Effective Janssen Distribution δ = 20°

Effective Rankine Distribution δ = 0°

(ref Daniels R J and Sharp B N, 1979)

-12 Caisson floor

0

20

40

60

80

100

Horizontal wall pressure : kNm2

8

120

140

160

The loading on the base plate of caissons depends on the form of wall pressure assumption. When a silo theory is used, the base plate will be relieved compared to that of a simple earth pressure at rest assumption, because the fill is hanging on the walls. The soil reactions on the base plate can also be calculated to the Danish Standard DS 415. The Technical Standards for Port and Harbour Facilities in Japan 1990 give clear guidance on the design loadings for walls and base plate, as does ROM 0.2-90. There are two schools of thought in Spain. The conservative approach adopts the pressure on the foundation in service, after deducting the mass of the submerged base slab itself, and the thrust at the bottom of the fill, taking into account the silo effect. It is, however, considered more realistic to calculate the pressure on the cell bottoms as equivalent to that extended by the whole mass of caisson and filling, supposing that just one cell remains unfilled. The above characteristic pressures can be used for the serviceability limit state or for traditional design by “working stress” methods. The considerations affecting the choice of partial factor to apply to derive the ultimate limit state pressures, and the various alternatives available are the same as noted above in 2.1.4, and in 2.2.3. Again a factor of the order of 1.2 appears to be generally appropriate. The density achieved for sand poured into water - filled caisson cells is low, much lower than for sand poured into open water. If it is required to limit settlement of the fill within the caisson cells, or +3

+3

t

∅' Submerged 30°

14

0

Wall Friction 0°

29

-3

γ Submerged 9.5 kN/m3

43

-6

γ Sea water 10.2 kN/m3 Ko = 0.5

76

-13

95

-17

kN/m2

- -17 17 t

(a) At-Rest Unconfined Pressure (a) Submerged SubmergedSand. sand. At-rest unconfined pressure +3 0 -3 -6

8

12

13

ACI 313 with Ka = 0.333

15 20 24 20

ACI 313 ditto with Overpressure Factor

25 34 DIN 1055 Pt6, Spanish ROM, and ACI 313 with K0 = 0.5. (Janssen) 28

51 32 Wall Friction 17° Cells 4.8m square 30 56 -17 34 (b) (b) Submerged SubmergedSand. sand.Silo SiloPressure pressure(Janssen) (Janssen)totoDIN, DIN,ROM ROM0.2-90 0.2-90and and ACI ACI313 313 -13

+3 -1.8

4.8m

0 -3

27

-6

-13 -17

kN/m2 27 (c) (c) Submerged sand. SiloSilo pressure. Japanese Technical Standards Submerged Sand Pressure. Japanese Technical Standards (applicable up to dimensions 5m 5m x 5m) (applicable upcell to cell dimensions x 5m)

Fig. of of Silo Fill Fig.8 8Comparison Comparison silo fillEarth earthPressure pressureCalculations calculationswith with External Earth Pressures external earth pressures

increase the density for reduction of liquefaction risk in an earthquake, this can be achieved by vibration (ref Daniels R J and Sharp B N, 1979/Cochrane G H, Chetwin D J L and Hogbin W, 1979). Due to the inverse relationship between density and Ø', densification causes no significant increase in wall pressure. Densification is usually only necessary if the fill is used to support surface works. There is conflicting guidance as to whether the reduction of earth pressure by silo effect applies if the caisson is vibrated by earthquake shock or wave loading. According to Japanese experience, the frequency of earthquake vibrations does not destroy the silo effect. The problem should be reduced by using appropriate fill materials and achieving appropriate fill density to preclude earthquake liquefaction. The density, in place, can be verified by Dutch cone or similar site investigation methods. The load cases for the design of the cells and the inter-cell walls must reflect the worst cases which can apply in practice. For example, the maximum water level case with water level at the surface must be considered. Also, the urgency with which the cells have to be filled when a caisson is placed in an exposed breakwater precludes the use of any procedure which involves special precautions, such as restrictions on the level to which adjacent cells are filled, or emptying the cells of water. The caisson must be designed to allow any sequence and proportion of filling. 2.1.6

Friction

Although the concepts of friction are classical and have been studied for so long, there is a surprising divergence in the figures used in design, and lack of agreed experimental data. There is a wide variation in the coefficients of friction assumed, and the factors of safety against sliding although these differences may tend to cancel out in the resulting stability equation. In practice, the friction coefficient and factor of safety must be related to the permitted displacement. Various coefficients of friction are compared in Table 4. The Japanese values have been determined from model tests (see Table 5). The friction coefficient depends on the size, configuration, strength and kind of stone, and the compacted condition, and is related to the coefficient of internal friction of the mound. The coefficient is low at initial construction, but increases in time after compaction by storms and self- weight consolidation. Weak rocks lead to a low value of µ. In Japan, the friction resistance is increased by placing a rubber mat or an asphalt mat under the base of the caisson. This measure enables the size of the caisson to be reduced. Other measures to increase the friction coefficient include corrugations or dowels in the base slab. The depth of the corrugations are 9

some two-times the dimension of the stone size at the surface of the rubble mound. Experimental results for friction coefficients in Japan are given in Table 5. The coefficients depend upon the factors related to the rock type and dimensions noted above and to the amount of displacement, weight of caisson etc. In comparison, a recent series of French tests (ref CÉTÉ - Laboratoire Régional Nord - Pas de Calais) are given in Table 6. The parameters investigated included : • two forms of bottom slab smooth and corrugated • two types of gravels crushed gravel 0 - 80 mm and natural sea gravel 20 - 80 mm • varying vertical load These results were considered to be lower than used in practice, and were accompanied by significant displacement, up to 180 mm, before maximum mobilisation.

TABLE 4 COMPARISON OF DESIGN VALUES FOR COEFFICIENT OF FRICTION Coefficient of Friction µ Condition

Japan (Technical Spain Standards (ROM for Port and 0.5-94) Harbour Facilities)

Precast concrete against bedrock or concrete Precast concrete against precast concrete Precast concrete against rubble Precast corrugated or sloping base against rubble Precast concrete on a rubber mat or asphalt mat against rubble In-situ concrete against rubble

UK and Germany (BS6349 BS8002 EAU 90)

France (Fascicule No .62, titre V)

tan∅' (often 0.58)

0.5

-

-

0.7

0.6

0.7

δ = 2/3∅r

-

-

δ = ∅r

0.7

-

-

1.0

TABLE 5 EXPERIMENTAL TEST RESULTS ON FRICTION COEFFICIENT (JAPAN) Precast concrete against stone No.

Kind of Stone

µ

Average of µ

Dimension of stone. mm

Condition 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= block W

Pmax W

P (Pull) Stone

P (Pull force)

δ (displacement) TABLE 6 FRENCH RESULTS FOR FRICTION COEFFICIENT Vertical load. TON

Normal Stress T/m2

24.1 18.4

10.5 8

24.1 18.4

10.5 8

Friction Coefficient µ

Horizontal Force. TON Smooth

Corrugated

Natural Sea Gravel 20-80 mm 12.6 13.7 10.3 11.3

Smooth

Corrugated

0.53 0.56

0.58 0.62

-

0.43 0.47

Crushed gravels 0-80 mm

10

-

10.4 8.6

Various values of the minimum Factor of Safety against sliding for use with the friction coefficients are given below : Japan (ref Technical Standards for Port and Harbour Facilities) 1.2 Spain (ref ROM 0.5- 94) Permanent situation 1.5 Momentary situation 1.3 UK (ref BS 6349 Pt. 2 - 1988) 1.75 Germany (ref EAU 90): • using active earth pressure and wall friction 2/ ∅' 1.5 3 • using at-rest earth pressure 1.0. 2.1.7

Handling and Float-Out Loads

From the viewpoint of limit states, all handling and float-out loads belong to transient load situations. The partial factor for these loads can be kept low; a figure of γF = 1.1 is suggested. The loads can arise from different reasons, for instance from: a) movement of caissons or blocks after dismantling the formwork and during launching, b) storage on supports during the curing period or in stacks of blocks and before float-out, c) loads from handling with cranes during the curing period or tugs during placement. Loads which can arise during construction must be included in the stress analysis. If the handling process which will be used during construction by the contractor cannot be anticipated by the designer, the stress analysis has to be rechecked by the designer as soon as the construction method is known. Float-out loads must never be considered as loads at the contractor’s own risk, for which the designer is not responsible. Damage arising from float-out processes will not always become apparent during the construction period. The wave climate and the sea condition limits which can be tolerated during float-out activities have to be specified by the designer and have to be known by the contractor. These waves could be calculated by using Recommendation 3.1.4. If there are certain climatic periods during the year for which it can be guaranteed that high waves do not occur, float-out activity could be limited to these seasons. The relatively short-float-out process for a breakwater can very often rule the design and the work on site. Experience shows that in most places the times for float-out can only be fixed from day to day by the use of a local weather forecast. In harbour construction a lot of damage has been caused by wrong or incautious handling of concrete elements during the construction period (ref Bruun P, 1985). Surveys have shown that one of the reasons for the destruction of many rubble mound breakwaters is the fact that the concrete elements of

the outer cover for the breakwaters were already broken during construction and placement. An analogous situation can arise for the elements of vertical breakwaters. If damage arises, it will mainly occur under water and, even by good monitoring during the construction period, this will not always become obvious. There is a rule which has become compulsory for dam construction: all loads, water level changes and other factors that influence the stability of a dam also have to be anticipated for the period of construction and not only from the time at which the structure has been handed over. This rule should also be applied to the design of harbour structures. The Technical Standards for Port and Harbour Facilities in Japan (1991) give the following additional recommendations for the calculation of external loads during float out. a) the water pressure should be calculated with an additional draft of 1 m, b) the tractive force during towing should be calculated by using the formula (n.b. expressed here in SI units): T = ½ ρwCDv2A where T = tractive force CD = drag coefficient v = towing speed A = submerged area of the leading wall ρw = density of sea water Dynamic pressures and wave pressures are not considered but a rise of the water level, δ in front of the caisson has to be considered which is shown on Fig. 9. A head δ, of 1m has to be applied, according to the Japanese Technical Standards, 1991. It must be ensured that a positive freeboard exists at all times. The distance between the centre of gravity and the metacentre, the metacentric height, should be at least 5% of the draught of the caisson.

Fig. 9 Tractive force during towing and hydrostatic head between different cells. (ref Technical Standards for Port and Harbour Facilities in Japan)

11

B1

m 58.0 m

Fig. 11 "Camel Humps” of the caissons used for the Brouwerhavensche Gat, Holland

The settling velocity of the caisson before the first contact is related to the motion of the sea around the caisson. Even if massive float-out equipment is used (large cranes) a certain movement by swell during placing cannot be ruled out. Repeated

Fig. 12 Legs which provide a first landing on three points.

A B2 Fig. 10 View 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 first grounding, the caissons for the closing of the Brouwerhavensche Gat in Holland have been provided with “camel humps” (Fig. 11) and on other sites downstand legs have been provided. (Fig. 12). The humps or legs have to protrude at least more than double the tolerance of the ground levelling. In Japan the rubble is sometimes not trimmed to a horizontal formation, but is provided with an excess height in the middle (extra banking) to enforce a first contact between caisson and ground in the middle of the bottom slab.

12

18.0

16.2 m

2.1.8 First grounding Special loads will arise when the lowered elements or blocks first touch the ground. In most cases the caissons will never again undergo a comparable distribution of load. The loads are exceptional, mainly for the following reasons: a) It is not always known in advance which part of the floating element will touch the ground first b) It is not always known on which supports the element will rest after the first placement c) The impact load is a dynamic load which cannot be determined exactly, even if the grounding velocity and the elastic and plastic properties of the ground are known. The German EAU 1990 (Recommendations of the Committee for Waterfront Structures) states in 10.5.2 that a first landing on one point in the middle or on two points on two edges has to be anticipated as indicated in Fig. 10.

hitting of the ground must also be anticipated. During the sinking operation, sometimes only a relatively small volume of water has to be filled-in but, as the inherent mass and volume of the element is considerable, this leads to two effects: • Due to high mass element of the impact force (Force = mass x acceleration) there is a high impact force even though the acceleration is low. • The large volume or surface area leads to large variations in the displacement of the element for only small changes in water surface level. Both of these effects can lead to impact loads which are higher than any other load in the lifetime of the caisson.

2.2

Resistance Analysis, Internal Analysis

2.2.1

Structural Analysis in Element Design

This section applies particularly to structures built up of several elements, such as caissons, for which extensive analysis is required for the structural design of individual elements of the structure. Structural modelling of the actions on caissons can be carried out in two ways. (i) In the simpler and traditional approach, the structure is split up into sets of beams and slabs and calculations carried out by traditional manual methods and two-dimensional frame analysis, using computers where applicable. Guidance for the traditional design of caissons is given in national codes, such as the Spanish ROM 0.2-90, Technical Standards for Ports and Harbour Facilities in Japan, EAU 90 etc. It must be stressed that, in such methods, calculations are carried out on each member separately and therefore require assumptions to be made about the connections between members. These assumptions can lead to questionable approximations as to the boundary conditions introduced in the calculations. (ii) The increase in computer power has made it feasible to carry out a full three-dimensional model analysis, using finite element methods. Nevertheless for simple geometries such as for caissons, it is not essential to use volume elements. A Japanese example is illustrated in Figs 28 and 29. Generally, thick shell elements will suffice for engineering purposes, while volume elements may be helpful for specific local analyses, for instance to map the local stresses and strains around a fastening point (e.g. in the loading case of towing). For most engineering problems, static models are the most familiar and most widely employed. Even if the phenomenon is dynamic in essence (earthquake, waves, wind, etc.) the loads applied are generally 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 models for slender members that are liable to second order instability (buckling, warping, etc.). In this respect Eurocode 1: (ENV 1991-1994) specifies that the effects of strains and displacements should be considered if they result in an increase of the effects of actions (loads) by more than 10%.

As regards the behaviour of materials, the assumption of elastic behaviour is the most widely used for simple structures, even though it is well known that the behaviour of concrete is not elastic, by far. Going further into the analysis, it is necessary to assess the moments of inertia of the sections depending upon the cross section area of the steel bars used for reinforcement, which are unknown at the primary stage of design. Therefore the initially estimated dimensions of the concrete section are generally considered in the first place, yet iterative calculations can still be made if in doubt. In implementing structural finite element models, the main problem may be the modelling of soil behaviour, i.e. the definition of a stress-strain relationship. In this respect, Winckler’s spring method is a traditional approach, consisting of a linear relationship, for each point of the bottom slab, between the vertical pressure (effective stress) exerted upon the soil and and the vertical displacement. The ratio between these two values, is termed the subgrade modulus K: σ' = K ∆z if ∆z < 0 σ' = 0 if ∆z ≥ 0 where σ' is the effective pressure exerted by the soil upon the structure. The value of K commonly varies from approximately 10 to 50 mN/m3 for a rubble foundation set on natural soil. It should be noted that this simplistic method totally disregards shear stress resistance within the soil. Furthermore, the value of K may vary across the slab, and according to depth as well, in the case of non-homogeneous soil layers, and so the appropriate 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 of this parameter generally means that the assumptions made for the resistance of the soil are inadequate. If the uncertainties about the soil behaviour are critical to the design, more complex element models, such as finite element models, may also be used, but such finite element models require that there should be adequate soil testing and adequate interpretation of the tests. So complex as the soil model may be, it seems sensible to test the design against a local reduction of K and/or against settlements in areas that may suffer scour (e.g. in corners).

2.2.2

Scale Models

The assessment of external stability for vertical structures is often assessed by means of “scale models”, which can also provide the more particular loading that can be generated on some members that are particularly exposed. 13

Such members are isolated mechanically from the structure and supported by weighing gauges. This can be useful for some types of structures with a perforated front wall. The effects of impacts related to wave breaking can also be derived from this kind of device. Alternately, this can be carried out using a two-stage numerical model: • first one can implement a global model and a loose meshing, apply the loadings but exclude the impacts at this stage. • then the specific members subject to the impacts can be computed using a more refined meshing that is fitted into the first one; the limit displacments implemented in these “zoom” models are those derived from the initial model.

2.2.3

Limit State Design and Risk Analysis

Structural analysis was traditionally carried out to working stress design limits, until limit state design methods were introduced in the early 1970’s, since when they have been used for caisson design. When limit state methods are used, the procedures required are as follows: • specification of the limit states to be considered, regarding the requirements the structure is intended to fulfil • selection of the parameters that are deemed representative of events to be accounted for with special care for variable loads • setting up of load cases with application of partial coefficients • implementation of structural models • assessment of relevant concrete sections, using national regulations as regards the partial coefficients to be applied to the resistance parameters of the materials (steel and concrete). It is not simply a case of adapting partial factors from one source to another, because the principles of reinforced concrete design may be different. It must also be noted that when the partial factors were set up for structural purposes, the results were intended to be similar to designs using traditional working stress codes. Even if the limit state concept involves a semi-probabilistic approach, to date, its implementation has been carried out in a deterministic way. However, the application of general concrete codes of practice to maritime structures suffers from a lack of guidance, which leads to a variety of interpretations for the designer. Indeed, it

14

proves hard to specify the frequency of the events to be considered and the representative parameters as a function of both the lifetime of the structure and the nature of the limit state that is being considered. For example, for some configurations, even with geometries as simple as caissons, it may not be obvious whether to classify all the loads as favourable or unfavourable, as this may depend on the section, the member, or the other loads. Should these loads be critical to the design of the member, it may even be advisable to perform calculations with both favourable and unfavourable partial load factors. Otherwise, the partial factor may be chosen equal to 1. It, however, remains the case that partial factor values, as recommended by general national regulations for the assessment of concrete or steel (etc.) structures, have been calibrated for the design of land-based structures in relation to the failure probability for the failure mode under consideration and for some reference period - e.g. the design working life, (buildings, bridges, etc. Refer to Eurocode 1 and Section 2.3.2 for the definition of design working life) so they may not be appropriate to structures subjected primarily to environmental actions (loads), such as waves. Similar problems of application exist in relation to earth pressure loading and geotechnical calculations, as referred to in 2.1.4 and 2.1.5. There are, for example, problems in applying the principles of BS 8002 and Eurocode 7. As noted in 2.1.4, BS 8002 adopts limit state philosophy but does not involve partial factors. There is a fundamental difference between the development of earth pressure loading with respect to strain, and the forms of loading met in buildings and bridges. There is a strong body of opinion which claims that a partial factor approach is incompatible with soil mechanics. At present there is no programme for issue of a Part 5 of Eurocode 2, for marine and maritime structures. There are groups in Europe working on this subject and the suitability of partial factors, and the Japanese have introduced a set of factors which they have used for the design of prestressed caissons (ref Kiyomiya O, 1994 and also Kiyomiya O and Yamada M, 1995) PIANC’s safety approach, as can be seen from WG12 on rubble mound breakwaters, seems to offer a more appropriate design approach. Much work still needs to be done to work out operational methods for the structural design of caissons. The main problem is to determine damage functions that would be valid for the various limit states to be considered, for they must address physical phenomena far more complex than those relating to external stability.

2.3

Durability of Concrete

2.3.1 Durability, Introduction Specification for durability for most materials is derived from a “materials” point of view, i.e. from the properties of the material, the environment and the expectations of protective treatments such as painting or cathodic protection. The ranges of successful performance are subject to research and experience, and can often be specified by reference to compositional limits and by performance tests, fatigue limits etc. Until the present time this “materials” approach has also been applied to concrete and reinforced concrete. Prescriptive forms of specification (refs Beeby A W, 1992, Clifton J R, 1993) and empirical relationships between concrete mixes and laboratory and field performances given in codes of practice have been deemed to achieve a satisfactory performance in certain classes of exposure. Although this approach has achieved relative success in most land-based building applications, this has often proved to be far from the case in the severity of maritime exposure, particularly to sea water or de-icing salts. The most dramatic failure mechanism is that of reinforcement corrosion, which can impose severe limitations in relation to the design and economic feasibility of complex thin walled structures such as caisson breakwaters or light structural superstructures. Great advances have been made in the computation of wave loading, risk analysis and the selection of the partial safety factors to apply to the level of risk in conjunction with the modern computational power for analytical design. In the same period, the durability of reinforced concrete subject to severe marine exposure or de-icing salts has been a dramatic failure (refs Aïtkin P C, 1993, Rostam S & Schiessl P, 1993). The present generation of national codes of practice and even the latest joint European Committee for Standardardisation (CEN) codes still reflect the prescriptive approach, and are seriously out of date. A consensus view for appropriate guidance will not emerge in less than ten years. The concept of durability is intrinsically connected with the concept of the service and design life of the structure, which is also related to the analysis of appropriate loading conditions and limit state analyses. Owners need to define the service lives required from their assets and plan a strategy for maintenance. Current codes do not give a rational basis for design of concrete to meet such a life. The assessment of this further “Durability” limit state (it is not actually a limit state but a means by which the other limit states are maintained over the operational life) is a fundamental part of design and must both predate and form part of all stages of structural design and detailing. A structured "durability plan" is required to ensure that durability

issues are addressed at all stages and primarily include: • choice of a service and hence a design life in order to achieve this service • recognition of the severity of the specific environment(s) affecting the structure • recognition of the consequences of the environment(s) on design, detailing and materials • analysis for durability, by analytical model where applicable • quality assurance and quality control of both design and construction • monitoring and maintenance strategy. These recommendations for a rational approach to design for durability can be matched with the following sources: (i) The CEB Guide to Durable Concrete Structures (ref CEB Design Guide, 1992). This mainly relates to buildings rather than sea structures, but gives excellent explanations of the mechanisms of deterioration and factors involved. Major contributions to the CEB design guide were made by a number of members of the CEB (Comité Euro-International du Béton) General Task Group 20 : Durability and Service Life of Concrete Structures). (ii) The work of RILEM committees (refs Schiessl P, 1993, Rostam S & Schiessl P, 1993), and CEN committees. (iii) The work of the British committees such as the Concrete Society Working Party on Durability Design and Performance Based Specification of Concrete, Report CS 109, 1996, and “Durability by Intent”, a strategy for the UK Dept of Environment programme on durability of concrete and reinforced concrete. (UK Dept of Environment Programme on durability of concrete and reinforced concrete). (iv) Important regular international conferences such 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 and Harbours, Port and Harbour Research Institute and the Overseas Coastal Area Institute of Japan (ref Technical Standards for Ports and Harbour Facilities in Japan, 1991), and papers from the Institute and the Japanese Concrete Institute (refs Proceedings of Symposia Japanese Concrete Institute, 1988 and 1989). (vi) The work of the Australian CSIRO Division of Building, Construction and Engineering (ref Ho D W S & Cao H T, 1993). (vii) Collaborative research projects such as the Brite-Euram Project 4062 on the Residual Service Life of Concrete Structures. 15

In the following sections 2.3.2 to 2.3.8, the factors and mechanisms which control durability are outlined. In the “materials” section 2.4, specific measures 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 great care, as the meaning of these terms is not the same. The terms are used with different meanings in different papers and different contexts and the first task 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 for which a structure is to be used for its intended purpose with anticipated maintenance but without major repair being necessary”. In the Spanish Maritime Works Recommendation (ref ROM 0.2-90, 1990) the term “Minimum Design Life” is used for this period. For maritime works subject to the probability and return periods of waves in addition to all other probabilities, this definition may require some adjustment, such as the following suggestion: “The assumed period for which a structure is to be used for its intended purpose with anticipated maintenance but without major repair being necessary within a probability appropriate to the function of the structure.” It is already stated in ENV 1991-1: 1994 that a different level of reliability may be generally adopted for structural safety and for serviceability and that a different level of reliability may depend upon the cause or mode of failure, amongst other factors. With respect, now, to durability in relation to deterioration of construction materials, the “Design Working Life” as so defined is the period specified by the Owner and is related to operational strategy. Following on from this, the designer has to select design methods and safety factors in order to ensure

a reasonable probability of achieving the specified life. Thus “Design Life” in the British Standard guide to durability (ref BS 7543, 1992) is defined as the period of use intended by the designer to support engineering specification and analytical decisions. The “Design Life” as estimated will therefore be at least equal to or exceed the specified “Design Working or Service Life” by a prudent margin, which includes factors of safety and ignorance. Different “lives” may need to be considered for economic and feasibility considerations (ref Port Development 1978, UNDP) and different strategies for monitoring deterioration and maintenance can apply. The concepts of “Design Working Life” and “Design Life”, and the differences between the various definitions may not be immediately obvious to readers unfamiliar with papers on the subject and, from experience, can lead to argument. The difference between the various definitions may be clarified by Table 7. A logical structure of design working (or service) lives for maritime structures (although there termed as “minimum design lives”) is given in Table 2.2.1.1 of the Spanish Maritime Works Recommendations (ref ROM, 0.2-90, 1990) as set out in Table 8. The “design working lives” correspond with Classes 2, 3 and 4 of the draft Eurocode 1 (ref Eurocode 1 ENV 1991-1:1994), but are usefully expanded to include general use and specific industrial infrastructure, which can be especially applicable to port and coastal works. This Table is recommended for use. It is important to note that, despite the recent introduction of more clearly defined “service” lives, current codes of practice and design guides do not provide guidance for adequate analysis or the means to satisfy the stated lives. However, a framework for modifying partial factors commensurate with designing for different lives and probabilitites is available. (ref Eurocode 1, Part 1, Annex A and Rilem Report 14, 1996, edited by Sarja and Vesikari) and from the work of Sub-Group D of PIANC WG 28.

TABLE 7 SUMMARY OF DEFINITIONS OF VARIOUS “DESIGN” LIVES Designated “Life”

Explanation

1. Design Working Life (ref Eurocode 1) or Service Life (ref BS 7543) or Minimum Design Life (ref ROM 02-90, 1990).

The utilisation period (or periods) specified by the Owner, with respect to structural safety, serviceability, or durability of structure and components (components likely to have shorter periods).

2. Design Life (ref BS 7543, 1992) with respect to durability.

A period at least equal to (1) or greater than (1) by a prudent factor, employed by the designer in order to achieve (1). Note: the estimates are not a precise science and cannot be guaranteed, but they can be subject to rational analysis.

3. Economic Life (ref Port Development UNDP, 1978 ).

A period used for economic and financial studies, i.e. for comparison of alternative capital and maintenance policies, using discounted cash values.

16

TABLE 8 DESIGN WORKING-LIVES (SERVICE LIVES) DEFINED IN ROM 0.2-90 AS "MINIMUM DESIGN LIVES"* FOR WORKS OR STRUCTURES OF DEFINITIVE CHARACTER (IN YEARS) REQUIRED SECURITY LEVEL

TYPE OF WORK OR INSTALLATION

LEVEL 1

LEVEL 2

LEVEL 3

GENERAL USE INFRASTRUCTURE

25

50

100

SPECIFIC INDUSTRIAL INFRASTRUCTURE

15

25

50

LEGEND GENERAL USE INFRASTRUCTURE General character works: not associated with the use of an industrial installation or of a mineral deposit. SPECIFIC INDUSTRIAL INFRASTRUCTURE Works in the service of a particular industrial installation or associated with the use of transitory natural deposits of resources (e.g. industry service port, loading platform for a mineral deposit, petroleum extraction platform, etc). LEVEL 1 Works and installations of local or auxiliary interest. Small risk of loss of human life or environmental damage in case of failure. (Defence and coastal regeneration works, works in minor ports or marinas, local outfalls, pavements, commercial installations, buildings, etc).

2.3.3

Processes of Deterioration

There are a number of well-known deterioration processes for the concrete matrix in sea water, and their relative significance depends on the specific location and climate, but the most widespread and serious problem is that of chloride-induced corrosion of reinforcement or embedded metal generally. A schedule of the deterioration mechanisms applicable to maritime structures is given in Table 9. Guidance and limits relating to these forms of deterioration are covered in National Standards and other Codes (refs EAU - German Waterfront Structures Code, BRE Digest 363, 1991 - Sulfate and acid resistance of concrete in the ground, Concrete Society Report TR 30, 1995 - Minimising the risk of ASR, BRE Digest 330, 1988 - Alkali aggregate 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 of origin. Care must be taken to clarify to owners and contractual parties if and where departures are to be

NB: 1. The General Use period of 25 years corresponds with Class 2 of draft Eurocode 1. LEVEL 2 Works and installations of general interest. Moderate risk of loss of human life or environmental damage in case of failure. (Works in large ports, outfalls of large cities, etc). NB: 1. The General Use period of 50 years corresponds with Class 3 of draft Eurocode 1. LEVEL 3 Works and installations for protection against inundations or international interest. Elevated risk of human loss or environmental damage in case of failure. (Defence of urban or industrial centres, etc). NB: 1. The General Use period of 100 years corresponds with Class 4 of draft Eurocode 1. *Defined as Design Working Life in draft Eurocode 1.

made from National Standards on rational grounds, such as mainly apply to reinforcement corrosion. The dominant factors involved in the durability of concrete are: • the recognition that concrete is a porous material and its behaviour depends on the pore structure achieved and, where applicable, cracks • the transport mechanisms for water and dissolved deleterious agents and gases within the pore structure and, where applicable, cracks • the macro, meso and microclimate for the structure and particular element. An excellent explanation of the significance of these factors and the transport mechanisms is given in the CEB Guide (ref CEB Design Guide, 1992). The following recommendations mainly concentrate on the case of reinforcement corrosion. Reference to the other forms of deterioration generally can be left to National Standards and references such as given in Table 10. Some details from recent publications are given in 2.4.3 and 2.4.4.

17

TABLE 9 DETERIORATION MECHANISMS FOR MARITIME CONCRETE Deterioration Mechanism

Locations most likely to occur

Method of Avoidance

Reinforcement corrosion (due to chlorides)

Elements wetted but subject to drying especially hot dry climates. See Figure 13. Corners subject to increased wetting and then drying. Areas of low cover.

Analysis, design and detailing. Properly designed cover to reinforcement for specific exposure conditions and tolerances.

Sulfate attack on concrete matrix

Delayed action in seawater. Colder waters may be more critical.

Specification and tests.

Salt weathering of concrete surface

Elements subject to concentration of salts by drying - intertidal zone. Paradoxically, cements which achieve the finer pore structure and resistance to steel corrosion may be most susceptible.

Specification. Extensive water curing.

Alkali-aggregate reaction

Susceptible aggregates, pessimum reaction with mixed aggregates. Alkalis from sea water and marine aggregates. Rich mixes.

Specification and tests Petrography. Mix limitations.

"Frost" (freeze-thaw) action

In cool with freezing zones with prolonged and repeated freezing.

Specification and detailing. Air entrainment spacing factor.

Abrasion

Subject to abrasive bed movement, shingle, vessel impacts, ropes and moorings.

Higher strength concrete, detailing, extensive curing, controlled permeability formwork, permanent steel protection.

Early thermal cracking

Thick sections and massive structures built in separate pours, causing restraint to shrinkage during cooling from heat of hydration.

Design and detailing, specification, pre-cooling of mix, cooling pipes inbuilt for the hydration period.

Plastic shrinkage cracking (workmanship)

Arid climates, drying winds, low bleed mixes.

Curing and protection at casting.

Plastic settlement cracking (workmanship)

Deep sections, high bleed mixes.

. Mix design, reduction of bleeding. Revibration.

TABLE 10 GUIDELINE DOCUMENTS FOR DETERIORATION MECHANISMS Deterioration Mechanism

References

Sulfate attack

National Codes BRE Digest 363, 1991 (UK) CEB Guide to Durable Concrete Structures, 1992 NF P 15-010 (France) NF P 18-011 (France)

Salt weathering

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

Alkali aggregate reaction

Concrete Society TR 30, 1995 (UK) BRE Digest 330, 1987 (UK) CEN Technical Report, 1994 NF P 15-010 (France) NF P 18-011 (France)

Frost action

Japanese Papers (Koh Y and Kamada E, 1993) CEB Guide to Durable Concrete Structures, 1992

Abrasion

Advice is not widely available CEB Guide to Durable Concrete Structures, 1992

Early Thermal cracking

BS 8007 1987 CIRIA Report 91:1993 (UK) CEB Guide to Durable Concrete Structures, 1992 Appendix A ACI 207.2R-90 Japan Concrete Institute. Manual of Massive Concrete, 1986 (Japanese only)

Plastic shrinkage and settlement

CEB Guide to Durable Concrete Structures Concrete Society Report 22, 1992 (UK)

18

2.3.4 Exposure Classification The policy of current European Code Committees for concrete (for the revision of ENV 206 : 1989 to pr EN 206) is to classify exposure conditions specific to the various deterioration mechanisms, where appropriate. The transport mechanisms affecting the deterioration of either the concrete matrix or embedded metal are largely dependent on the pore structure and the environmental conditions within and immediately exterior to its surface. Deleterious substances are transported by the medium of water (excluding CO2 and O2), and the moisture content of the concrete controls both the rate and effect of the transport of such substances by air or water. The mechanism of reinforcement corrosion is foremost in mind throughout the subsequent text but other mechanisms, i.e. frost, also apply in colder regions. In seawater environments it is necessary to recognise three basic climatic influences; that of the macro, meso and microclimates (ref Fookes P G, 1993) i.e. climates on the scale of the country, the site, and the particular element of the structure respectively. For most practical purposes the macro and meso environments can be considered together. The micro-environment, i.e. the location of a specific member in relation to sea level, and the degree and frequency of inundation by seawater and drying out, is particularly critical. The most important macro and meso climatic factors are temperature and rainfall. Temperature controls the rate of chemical reactions and the degree of drying out of the cover concrete. Rainfall, humidity and the location of the member in relation to sea water level movement control the wetness of the concrete, which affects the mechanism for penetration of chlorides and controls the penetration of oxygen to fuel the corrosion process. Generally, only the surface layer of concrete “dries out”, the depth of the drying and wetting zone being much greater in arid climates than in temperate climates and, consequently, extremely critical for corrosion of reinforcement. The wetting and drying depth in a temperate climate may not exceed 20mm, where it may 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 macroclimate: • cool with freezing • temperate • hot wet • hot dry and some seven micro-environmental cases of exposure applicable to chloride-induced corrosion in maritime works in ascending order of severity: (NB The classification below has been adapted from cases XS1 to XS3 now being proposed by European code committees for concrete exposed to chlorides).

XS1 Exposed to airborne salt but not in direct contact with sea water. Contrary to popular belief, airborne chlorides alone, without the vehicle of water penetration, do not achieve enough concentration to cause reinforcement corrosion in superstructures (ref Hussain S E, Paul I S & Bashenini M S, 1993) XS2 Submerged (Also a subdivision, XS2A - backfilled) XS3 The tidal, splash and spray zones. Depending on the macro-climate and degree of wetness, this class may need to be broken down further into four cases of ascending severity: XS3.1 Mid and lower tidal. The continuity of saturation in this part of the tidal zone is beneficial in restricting the flow of oxygen, and conditions can be similar to XS2 (refs John D G, 1992, John D G, Leppard N W & Wyatt B S, 1993) XS3.2 Upper tidal and capillary rise zones XS3.3 Splash/spray zones XS3.4 Mostly dry infrequently wetted: i.e. concrete which is above the splash zone but subject to seasonal change in sea level, storm events, testing of fire hydrants and run off from mooring lines at capstans and bollards. Again, contrary to popular belief, the worst case is not necessarily the splash and intertidal zone. The latter is likely to be the worst case for the material itself, i.e. bare and painted steel, timber, masonry and plain concrete, and for frost damage, but not for steel embedded in concrete. In cool and temperate conditions there may not be much difference between cases XS3.3 and 3.4, but the difference can be much more pronounced in hot wet and, especially, hot dry conditions. Fookes (Fookes P G, 1993) has suggested an 11-point scale for exposure risk of all concrete in a hot salty environment, both land-based and maritime. In a similar way, but restricted to coastal structures, severity ratings for concrete in a salt-water environment have been expressed on a scale of 1 to 12, as illustrated in Table 11 and Fig 13 (ref Slater D and Sharp B N, scheduled for publication late 1997). It is important to note that the “very severe” and “extreme” exposure conditions of BS 8110 : 1985 and BS 5328 1990/1991, and the most aggressive Class 4 of DIN 1045:1988, only reach about 3 on this 12 point scale, and it follows that such codes apply only to the cooler and temperate parts of Europe and similar climates. ACI 318:1995 and Eurocode 2 1992 classifications for seawater exposure range between 1 and 6 on this scale, depending on location within the tidal and splash regime and the ambient climate. Cooler European conditions may, of course, read only 4 or 5 on this scale. Sea water contains both sulfates and chlorides and therefore the concrete has to provide both a chemical resistance and protection for the reinforce19

ment. The latter requirement tends to control the concrete quality. In concrete that is submerged, XS2 or XS3.1, chlorides penetrate the concrete by diffusion, a slow process when compared with capillary suction. In submerged sections, the lack of oxygen is likely to stifle corrosion to an extremely low rate. Exposure class XS3 has the worst conditions. Chlorides will penetrate by a combination of absorption/capillary suction and diffusion. In periods of calm weather, wick action and evaporation leads to a maximum concentration of chlorides near the exposed surface some distance above the water line. Chloride is transported into concrete by the medium of water by a number of mechanisms. Chloride permeation is a complex phenomenon, including capillary suction and diffusion mechanisms, and hydration suction. The most rapid transport mechanism, capable of conveying the greater quantity of ions into dry or partially saturated concrete, is the capillary suction or absorption process. The much slower process of diffusion takes place in saturated concrete. The largest increase in chloride content of the cover zone is achieved by capillary rise or suction followed by surface evaporation, or unbalanced cyclic wetting and drying, i.e. irregular inundation by salt water followed by a period in conditions which enable the concrete to dry, due to seasonal water level changes or storm events (refs Saetta A V, Scotta R V and Vitaliani R V, 1993, John D G, 1992, Sandberg P J P, Petterson K, Arup H and Tuutti K, 1996).

Infrequently wetted, wetted or overtopped in seasonal conditions, run-off from ropes, hydrants

Cold/ temperate

Hot Wet H D

3 4

9

splash wave crest

Splash/Spray Zone

6

Upper Tidal 0 1 2 3 45 6 7 8 910 Severity R Scale Mid and Lower Tidal

2

3

Cold/ Temperate

Hot Wet and Hot Dry

wave trough Hot W and H Dry

Submerged or Backfilled Cold/ Temperate Ditto Backfilled Fig. 13

Ditto Backfill 1 2 3

Suggested severity ratings on a scale of 1-12 (Refer to Table 11)

TABLE 11 SUGGESTED 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 Location1

Cool with freezing

Temperate

Hot wet

Hot dry

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

Notes:1 2 3

4

20

See Figure 13. Infrequent splashing by seawater but otherwise exposed to weather e.g. copes. Concrete exposed to direct sunlight but protected from rainfall in a hot wet macro-climate may experience a micro-climate of higher severity rating because of the drying and increase of salt concentration caused by the absence of wetting. Rating increases with increase in ambient temperature and duration of dry periods e.g. Lower Arabian Gulf coastline 12.

5

6

7

For concrete above water level, if capillary rise and evaporation is not prevented this may cause increased salt concentrations in the fill and lead to a higher severity rating. Sulfate attack due to ground conditions not taken into account. Surface coating recommended if concrete mix required to resist chlorides not adequate to meet sulfate class. Abrasion effects not taken into account. Additional cover may be needed over and above Table 15 values to provide required design life allowing for abrasion loss.

Hot-dry conditions are particularly hazardous for concrete in maritime works. They apply predominantly in the Arabian Gulf and Middle East generally, where aridity greatly exceeds that experienced in other regions, including the United States (ref Fookes P G, 1993). Absorption of chlorides and penetration of oxygen is usually restricted in wetter and cooler climates which prevent the concrete drying out at the surface. This wetness, itself, is not the cause of corrosion but the reason for reduced corrosion. In hot arid climates the concrete dries to greater depths in long seasons of calm weather or seasonal lower tides. Wetting of the dry surface which then occurs in occasional storms, seasonally higher tides, the testing of fire hydrants or from mooring ropes, causes chlorideladen water to be sucked in deeply. Subsequent drying causes the salt content to concentrate. In a wet climate, rainfall would both reduce this concentration and the penetration of oxygen. Lack of water in the pores permits the free flow of oxygen. (Ref John D G, 1992, John D G, Leppard N W and Wyatt B S, 1993). Damage from this cause can particularly apply to precast concrete (such as caissons) which is left dry for a long period and then exposed to sea conditions. Torben Hansen, 1989, graphically describes the self-destruct possibilities in these conditions. The chloride concentration is only one factor, as the rate of development of corrosion depends upon a number of factors, principally regulated by the access of oxygen. The presence of water reduces the ingress of oxygen and the rate of reaction may be affected by cracks, in which case anodic and cathodic areas can be relatively close to each other. Depending on the relative magnitude of the cover and degree of saturation, corrosion may be general with closely adjacent anodes and cathodes, or the anodic reduction process at a location may be driven by cathodic activity caused by the ingress of oxygen at another location remote from the site of corrosion. Only free chlorides cause corrosion and consequently the “threshold” value for the critical chloride content depends upon the chloride binding capacity of the particular binder type, and the degree of saturation. For a given set of exposure conditions, the propensity for and rate of corrosion is influenced mainly by cover to reinforcement, binder type, water-cement ratio and binder content, broadly in that order. If capillary suction is the main transport mechanism it is unlikely that, using most modern Portland cements alone, concrete quality and cover will provide an adequate design life in severe exposure cases, if only conventional cover thicknesses and placing tolerances are employed (ref Bamforth P B, 1993, Neville A M, 1995). However, concrete of readily achievable production quality used in conjunction with realistic reinforcement placing tolerances is more likely to achieve design lives

when Portland cement is blended with other hydraulic or pozzolanic materials, such as blastfurnace slag, pulverised fuel ash or microsilica, provided that these materials are of appropriate quality. For the past 20 years the most popular model for analysing the ingress of chlorides has been that of diffusion, using Ficks second law, all arising from the pioneering work of Tuutti (ref Tuutti K, 1982). There have subsequently been developments with more realistic models dealing with partially saturated concrete (ref Grace W R, 1991). A number of European workers have more recently developed powerful suites of computer programs which attempt to model the whole range of variables of concrete condition and permit calibration from simple laboratory measurements (ref Kiessl K R, 1983, Roelfstra P E, 1989, Saetta A V, Scotta R V and Vitaliani R V, 1993). The limitations of the Fick model, and the explanation why its use in interpretation of chloride profile figures can lead to overpessimistic estimates of the rate of ingress, are given in a recent paper by Danish workers (ref Johansen V, Golterman P and Thaulow N, 1995). The exposure cases specific to “frost” (actually freeze-thaw) damage are given as follows: XF1 Moderate water saturation without salt XF2 Moderate water saturation with salt XF3 High water saturation without salt XF4 High water saturation with salt. Recommendations in connection with freezethaw resistance are given in 2.4.3.

2.3.5 Influence of Cement Type The principal choices for cement include plain (previously known as “ordinary”) Portland cement, sulfate resisting Portland cement (i.e. a cement with specific limits on the calcium tri-aluminate (C3A) content: the lower the C3A the more resistance to sulfate attack) or various types of blended cements. Blended cements include a combination of Portland cement with blastfurnace slag (gbs), which is a latent hydraulic binder, or pozzolanic materials which can be natural but are more likely to be pulverised-fuel ash (pfa). The latest available material is microsilica, which can be mixed with either unblended or blended cements. Guidance on its use is given in Concrete Society Report 41 (ref Concrete Society Report 41, 1991). It is most important to stress that any comparison between materials to different National Standards must be made in full knowledge and comparison of the different methods of test and specification. This particularly applies to cement strength class. Cements are produced to a number of strength classes and cement content for equivalent durability may need to be increased when lower strength cement classes are used. 21

Sulfate resisting Portland cement is unlikely to be necessary in maritime concrete, as the disruptive effect of sulfates is reduced in the presence of chlorides, and more so in warmer waters. It is claimed that, on the contrary, such sulfate-resisting cement is less resistant to reinforcement corrosion, but this may only be true of chlorides inbuilt into a mix. Sulfate-resisting cement is a relatively “lower heat” cement (relative to plain Portland cement) and its use is beneficial for reducing the generation of heat of hydration. Where only Portland cement is available, it is usual to limit the tri-calcium aluminate content in reinforced concrete to not less than 5% and not more than 10%. This constitutes a compromise by specifying, in effect, a moderate sulfate resistance by virtue of the upper limit on C3A whilst avoiding a too low content C3A which may be less able to protect the steel from corrosion. ASTM C-150 Type II cement properties can cover this requirement. In circumstances where reinforcement corrosion is not the critical problem, or for which adequate provision has been made, the long term effects of the reaction between sulfate and the hydrates formed from C3A could possible dictate the choice of low C3A Portland cement. See 2.4.4. However in these circumstances, blast furnace cements could combine the optimum solution. High proportions of blastfurnace slag, of from 50% to 70% and more, are beneficial from a number of viewpoints, including resistance to chloride ingress, sulfate resistance, minimising the effect of “alkali aggregate reaction” and reducing the rate of generation of heat of hydration. Alkali aggregate reactions are the chemical processes which can take place when the natural alkalinity of cement caused by the calcium hydroxide (pH about 12.5) is increased to a pH of over 13 due to the oxides of potassium and sodium. Proportions of pulverised fuel ash of the order of 30% to 40% have similar but lesser benefits. There appears to be benefit in adding microsilica in doses of about 5% in conjunction with slag or pulverised fuel ash. Where slag or pulverised fuel ash is unavailable, microsilica may be added in amounts up to 10% of the total cementitious content. It must be noted that the quality of these materials have to be appropriate. In some parts of the world slags and pulverised fuel ash contain unsuitable constituents, and in some places plain Portland cement appears to be very effective. Blended cements may be either factory produced or blended from separate constituents at the mixer. Blast furnace cements have been traditionally employed for maritime and other works in Germany, Holland and Spain, and pozzalanic cements used elsewhere. Often blast furnace cement was used because it was the principal material available. Recommendations are now made by many authorities, including the Japanese Technical Standards (ref Technical Standards for Port and Harbour Facilities in Japan, 1991), for the use of 22

high slag cements in maritime concrete. Although blended cements are almost a “must” to counteract corrosion, their resistance to surface scaling and tolerance to poor curing is less than unblended cements. Lower levels of slag, say 50%, may be more appropriate in colder conditions when the generally slower action of slag cements may cause problems in achieving early strenght. The weakness of much prescriptive advice in current codes is that guidance on mixes and cover thickness is given largely independently of cement type, coupled with imprecise description of exposure conditions. From the comparisons summarised above, the conclusion can only be that modern unblended Portland cement generally (there are exceptions) has the lowest resistance to chloride penetration and, where severe chloride exposure conditions exist, is unlikely to guarantee a long service life, even in temperate climates, with cover to reinforcement as often recommended. For values of cover in the accepted magnitude of 50 to 80mm, blended cements are a must. For each cement type and blending ratio, one can estimate minimum cement contents and maximum water-cement ratios appropriate to the different exposure cases. Typical figures are given in Section 2.4.3 to 2.4.5. See Tables 14 & 15. 2.3.6 Influence of Cement Content Concrete strength is directly related to watercement ratio. The pore structure of concrete depends on the water-cement ratio, the degree of hydration and the cement type (ref CEB Guide). Using superplasticisers, even under difficult conditions for controlling water demand without losing workability (i.e. high temperature, poor aggregates, large pours, large handling distances), it is now possible to reduce the water-cement ratio below 0.40, when this is required, and at the same time achieve both the desired fine pore structure and self healing properties due to the presence of unhydrated cement (ref Aïtcin P C, 1993). Another simple fact is that, ignoring the effect of water reducing admixtures, for a given workability, the amount of water required per cubic metre of a given aggregate type and gradation is effectively independent of the cement content (ref Barber P, 1989 and Fig. 14). This is the basis of the ACI method of proportioning concrete mixes (ref ACI 211, 1991), in which the water content is first established, and the cement content then derived from appropriate requirements for the water-cement ratio on the basis of durability. Another important fact is that the water demand required to achieve a level of workability varies inversely to aggregate size. The larger the aggregate size, the less water and cement paste is required and, consequently, less cement is required to achieve a given water-cement ratio. Illustration of the range of this effect is given in Table 12 (ref BS

TABLE 12 APPROXIMATE ADJUSTMENTS TO MINIMUM CEMENT CONTENTS AND WATER DEMANDS FOR AGGREGATES OTHER THAN 20mm NOMINAL MAXIMUM SIZE FOR A GIVEN CEMENT STRENGTH, AGGREGATE, TYPE, WORKABILITY, ETC. (Main figures from BS5328, BS8110 etc. Figures in brackets from computer simulation Mixsim, 1995) Nominal1 Adjustment2 to Difference in water maximum minimum demand for a given aggregate size mm cement content kg/m3 workability kg/m3 10 20 40 80

+40 (+30) 0 -30 (-20) -70 (-50)

+15 to + 20 (+15) 0 -15 to -20 (-15) -35 to -40 (-20 to -30)

1. Note that definitions of maximum aggregate size can differ between different codes. 2. According to the French expression overleaf, the adjustment is some 50% greater, but of course depends on the choice of "minimum".

device has traditionally been used in maritime works, where 40mm, 80mm and even 150mm aggregate has been used for plain concrete. This is important because cement content may be a controlling factor from the point of view of minimising alkali aggregate reaction (ref Concrete Society Report TR30, 1995) or controlling heat evolution in relation to early thermal effects (refs CIRIA Report 91, 1992, BS 8007, 1987, UK Dept of Transport BA 24/87 and BD 28/87, 1987). Mix specification limits given in specification documents must be realistically drawn from a rational mix design using available materials, and not merely numbers drawn from literature, as is often the case. The latest mix design methods available in the market make use of computer simulation to achieve proportions based on packing theory. Analysis using such software is given in brackets in Table 12 (ref Mixsim Issue 3, 1995). Computer simulation includes the effect of grading, and could incorporate the effect of mimimising the fines content, which is another approach to limiting cement content. 2.3.7

Cracking and the Influence of Cracks

The causes and consequences of cracking in concrete structures have often been misunderstood. In the past, most incidences of early thermal cracking were erroneously attributed to drying shrinkage. Most causes of cracking during the plastic state, i.e. 0

Cement content per cubic metre kg 100 200 300 400

500 Water

Mass of material per cubic metre

5328, BS 8110 and derivation using Neville AM, 1995, and also recent mix design software). It must be noted that the precise definition of aggregate size can differ between codes. This principle is covered by French codes NF P 15-010 and 18-011, where the minimum cement content for concrete in sea water or prestressed work is given by the following expression, together with a water-cement ratio less than or equal to 0.50: C = 700 5√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 watercement ratio and adequate workability using a range of normal to high performance water-reducing admixtures. The type of cement, in addition, has its own effect on water demand. Slags and pulverised fuel ash generally require less water. The design mix must provide the constructor with adequate workability to transport, place and compact the concrete. As the quality of the concrete in relation to both strength and durability is directly related to watercement ratio and unit water-content, within reasonable lower limits, the actual cement content is therefore a secondary consequence of the water demand for a given mix. Cement contents do not need to be especially high for plain concrete or even reinforced concrete providing requisite parameters are met. Cement contents must always be assessed in relation to aggregate size and grading, and a good way of reducing cement content is to increase aggregate size and/or reduce the fines content. This

Stone

Sand Cement over this range of normal mixes, the water content is effectively constant

Fig . 14 Graphical illustration of variation of constituents for practical range of cement content (ref Barber P, 1989)

plastic shrinkage and plastic settlement cracking can be resolved by attention to mix design, protection from drying winds, and curing in arid conditions (ref Fookes P G, 1993, Al-Rabiyah A R, Rasheeduzzafar, Baggott R, 1989). Early thermal cracking is an important case for design and construction of large masses of concrete, including both blocks and caissons, and is explained in 3.1.7. An excellent explanation of the processes and influence of cracking in concrete is given in the CEB Guide 1992, Concrete Society Report No. 44, 1995, and BRE Digest 389, 1994. 23

Although the significance of cracking may be much less than traditionally assumed in the case of reinforced concrete, it may be significant if it results in the reduction of the mass of some unreinforced concrete structures (for example armour units) and can cause unnecessary concern. Attention to detail in both design and construction, particularly in relation to early thermal effects, can obviate the incidence of cracking. Indeed, a number of European tunnel structures are currently being designed as “crack free”. Most structural codes have flexural crack width limitations appropriate to specified environmental conditions related to the serviceability limit state. The British method of design for crack control in relation to early thermal stresses (ref BS 8007, 1987) has similar limits. However it is now generally accepted (ref CEB Guide, 1992, Schiessl P and Raupach M, 1997) that, once crack widths exceed some 0.1mm, there is no significance in relation to the ingress of deleterious substances to cause reinforcement corrosion from wider cracks of up to 0.5mm. Accordingly, the careful gradation of the effect of cracks between 0.1mm and above is meaningless (ref CEB Guide, 1992). This statement refers to cracks perpendicular to main reinforcement and not to cracking above and along the length of a bar such as can occur in plastic settlement or if a bar (say a stirrup) acts as a crack inducer. Early thermal and workmanship related plastic cracking tends to create wider, uncontrolled, cracking coincident with reinforcement, which is therefore much more detrimental than flexural cracks which intersect with the reinforcement at right angles. The progress of reinforcement corrosion is largely dependent on the properties of the concrete itself, the moisture state, and the relative cover to reinforcement. However, cracks are likely to promote accelerated pitting corrosion. Appropriate detailing to distribute reinforcement spacing and limit crack widths remains important for various reasons. Cracks do not significantly increase the effect of freeze-thaw damage, as the scaling or splitting of concrete by freeze-thaw action is due to the increase of volume of completely water-filled pores, and is a

Cover mm

concrete covering thickness

crack width

Fig. 15 Decreasing 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)

24

feature of the general surface of the concrete (CEB Guide, 1992). The recommendation for increased cover given in this Report is likely to be met by the objection that the width of cracks on the surface will be unacceptably wider than permitted by design codes. Many engineers will be worried that increased cover will lead to wider flexural cracks and that reinforcement further distant from the surface will lead to wider cracking of the cover zone. However, chloride ions and oxygen penetrate the concrete everywhere and not just at cracks, and the width of the crack at the surface is not as critical as the width at the reinforcement itself. Research has shown that the crack width at the steel bar is almost independent of cover thickness, and that the width of the ‘V’ shaped crack increases almost linearly with cover thickness (see Fig. 15). Therefore, the cover thickness should not be limited for crack width reasons. When checking crack width limits in accordance with codes, one may increase the permissible crack width pro rata to the ratio of preferred cover to the typical cover given in a code. This principle is already included in some instances, for example where permitted surface crack width is given as a proportion of the cover thickness (i.e. such as 0.004 times cover thickness). (refs Merkblattsammlung, German Concrete Association, 1991, Dept of Transport (UK) BA24/87 and BD 28/87, 1987) Crack widths caused by early thermal effects can be controlled by appropriate design (ref CIRIA 91, BS 8007). It is not necessary to sum the effects of early thermal and flexural cracking (ref Department of Transport (UK) BA24/87 and BD 28/87).

2.3.8 Influence of Curing The need for curing continues to be hotly debated and is the subject of contemporary studies (ref CIRIA Research project on the influence of practical on-site curing, in Progress 1994/95). In the past, with much less reactive cements, prolonged water curing was necessary in order to achieve strength. This is no longer the case and, obviously, the influence of water curing on the surface is unlikely to influence the centre of massive sections. On the contrary, indiscriminate application of cold water could cause thermal shock cracking and as many problems as it attempts to solve. It is perhaps the case that, in the case of wet temperate climates, relaxation of curing is less deleterious than is generally held. However, in hot dry arid climates, water loss from the surface can be significant and result in incomplete hydration. As (ref CEB Guide, 1992) the relationship of watercement ratio and degree of hydration dictates the resulting pore structure of the concrete, and as the duration of water curing is inversely proportional to water-cement ratio, one main advantage of using the lowest possible water-cement ratio is to reduce the

duration of water curing. From the graph/nomogram of the relationship between water permeability, volume of capillary pores and degree of hydration (CEB Guide, 1992), it follows that the lower the water-cement ratio, the smaller is the duration of curing required for the preclusion of continuous capillaries. This follows from the classic work of Powers and Brownyards (ref Powers T C and Brownyards T L, 1988) and as explained by Hansen T C, 1989, Ho D W S and Lewis R K, 1983, and Neville A M, 1995. A higher strength concrete obtained by adopting a lower water-cement ratio is therefore, to a degree, self-curing. Work at the University of Dundee (ref Dhir R K, Hewlett P C, Lota J S and Dyer T B J, 1994) is in progress for the achievement of self-curing by adding water-soluble chemicals to reduce water evaporation in the set concrete. The requirements for curing are also intimately related to the type of cement and bleeding characteristics of the mix. All of these properties tend to have conflicting effects on the result. For example a high bleeding rate may increase the water-cement ratio of the cover concrete and lead to plastic settlement. However an unduly low bleeding rate (as can occur with microsilica) may lead to dessication of the surface and hence to plastic shrinkage cracking. Slag cements are generally held to require great attention to curing to prevent surface breakdown. However it can also be the case that surface breakdown by salt weathering of slag cements is due to the fineness of the surface pore structure, which leaves inadequate room for the accommodation of surface salt crystals which may be provided by a concrete with a coarser pore structure.

2.3.9 Monitoring and Maintenance Strategic planning for maintenance is directly linked with the concepts of service life and durability. In the past, very little guidance for the maintenance of concrete and other structures has been available. The subject is only now receiving due attention. The Technical Standards for Port and Harbour Facilities in Japan, 1991, places much more emphasis on maintenance and durability than in its previous version and points out that maintenance is largely related to monitoring and management strategy. Maintenance needs depend on feedback from regular monitoring and data-collection, for comparison with “base-line” data, i.e. line, level etc,

of the situation “as constructed”. This emphasis is also made in the CIRIA/CUR Manual on the use of rock, 1992. PIANC’s Permanent Committee for Developing Countries (PCDC) undertook, in 1978, to produce a series of maintenance manuals for port infrastructure, (ref PIANC Bulletin 1985-No.50) but, so far, manuals have only been produced for the more obvious cases of mechanical equipment, roads and railways. PIANC Working Group No.17 published their report on the Inspection, Maintenance and Repair of Maritime Structures exposed to Material Degradation caused by a salt water environment in 1990 (ref PIANC, 1990). This Report gives helpful guidance on inspection methods and monitoring but covers such a broad range of materials including timber, stone, unreinforced concrete, reinforced and prestressed concrete and steel, that the specific explanations and guidance for each material are somewhat abbreviated. Monitoring must be planned and adequate records taken, beginning with “base-line” measurements of line, level etc. immediately after completion of construction. Computers can now greatly facilitate this kind of work. Regular inspections of the structure (or beach or coastal defence) should be carried out at least once per year, most likely following the winter storm period, and after any major storm event. The principal objects of the survey are to determine: • the integrity of armour units and elements of the structure • indication of movement and settlement • scour. Measurements must be taken at specific locations to provide adequate mapping of the structure at clearly located profiles or to a grid, and plotted on large-scale drawings. Computer and modern digital methods are now available. Aerial survey and underwater 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 crack and damage mapping. Non-destructive methods of monitoring the performance of reinforced concrete are not as practicable or meaningful as often claimed and suffer problems of interpretation, but equipment is available for monitoring corrosion potentials by half-cell and other methods. It is now possible to build in probes and take electrical measurements to assess the onset of any change.

25

2.4

Materials

2.4.1 Rock and rubble For guidance for, and the specification of, rock and rubble construction it is recommended to use the recently published “Manual of the use of rock in coastal and shoreline structures”, published jointly in the UK and Holland (ref CIRIA Special Publication 83/CUR Report 154, 1992.)

2.4.2 Filling and Backfilling Granular materials are recommended for any backfilling to walls and filling of caisson cells in order that the earth pressure loading can be determined with confidence, as outlined in Sections 2.1.4 and 2.1.5. There are no strict requirements for the characteristics of the material used, i.e. quarry run rock, sand etc. It is sufficient that their saturated density is equal to or greater than the mass used in the stability calculations and that their angle of internal friction is equal to or greater than the one used in earth pressure calculations. Generally, if the material is suitable for pouring directly in position under water, it will be acceptable. Material poured into water (except for rockfill), will achieve a low relative density and measures may be necessary to increase the density, as described in 2.1.5

2.4.3

Concrete Durability General Design, Detailing and Workmanship The requirements for achieving durability of concrete in maritime works will usually outweigh the requirements for achieving strength, or density. The factors influencing durability and the specific deterioration mechanisms involved were summarised in Sections 2.3.1 to 2.3.8. This section and 2.4.4 to 2.4.12 summarise the practical steps and specifications appropriate to maritime concrete in breakwater applications. The measures to be taken against specific deterioration mechanisms as listed in 2.3.3 and Table 9 are most concisely set out in the CEB Guide and as listed in Table 10. Many problems can be mitigated by good detailing and by good curing of the concrete surface. Risk of frost damage and reinforcement corrosion can be lessened by attention to drainage of water from near horizontal surfaces, to prevent the ponding of water, and to prevent the run-down of water on vertical surfaces. Reliability of design and workmanship and the chance of avoidance of “gross errors” can be improved by employing the discipline of Quality Assurance in both design and construction, and the 26

Quality Control and audit procedures included in this process. National Codes of Practice and Design Guides, and their associated test methods and limits can usually be relied upon for appropriate design for protection against physical and mechanical action, freeze-thaw damage, chemical and sulfate attack and alkali-aggregate reactivity. However, at the present time, national codes and standards cannot be relied upon to assure durability against reinforcement corrosion in the long term, particularly in aggressive conditions and climates other than cool or temperate. It is recommended that, in the maritime environment, design for durability needs to be “explicit” in the sense understood by European code committees, meaning a move away from prescriptive limits given in codes to methods based on rational analysis and methods of test (ref Concrete Society Report 109 on Durability Design and Performance-based Specification of Concrete 1996, Rostam & Schiessl, 1993, Clifton J R, 1993). The principal deterioration mechanisms in the case of unreinforced concrete are those of salt weathering, surface scaling and freeze-thaw, and abrasion. The principal cause of deterioration of reinforced and prestressed concrete is the corrosion of embedded metal but, obviously, the other deterioration mechanisms as for unreinforced concrete still apply. The performance of different concrete mixes depends upon the environmental conditions and the type of cement. The durability of unreinforced concrete incorporating an appropriate water and cement content is good in sea water conditions and, where seriously aggressive conditions exist, as defined in 2.3.4, consideration should be given to designs appropriate to unreinforced concrete as opposed to designs involving embedded metal.

2.4.4 Unreinforced Concrete (Plain or Mass) The term “Unreinforced Concrete” will be used to define concrete without steel reinforcement. Such concrete is usually termed “mass” in Europe and UK. In the USA and Japan, mass concrete means concrete of a size significant to heat generation which will require measures to be taken on account of heat generation. The surface of unreinforced concrete requires to be designed and constructed with freeze-thaw, abrasion, salt weathering and sulfate attack in mind. Although sulfate attack is mitigated in sea water by the presence of chlorides and is less harmful in warmer waters (refs BRE Digest 363, 1991, Bijen J M, 1984, Matta Z G, 1993, NF P (18-011)), it is possible that the long-term effects of the reaction between sulfates and the hydrates formed from C3A may be more significant in cases where the much

earlier and dramatic effects of reinforcement corrosion are absent. The choice of grade of concrete can be particularly problematical in the case of unreinforced concrete in blockwork, large sections and armour units. A number of contemporary code recommendations may lead the designer to concrete mixes with excessive strength and cement content, and hence to problems of early thermal cracking and brittleness. Due to the changing chemistry of cement it is not always valid to compare specifications with successful mixes used in the past. The UK Maritime Code BS 6349 Part 1 : 1984 recommends a minimum cement content of 350 kg/m 3 and maximum water-cement ratio of 0.50, but this is meaningless without reference to the options of increased aggregate size or the use of admixtures, both of which could enable the water content and hence the cement content to be reduced. The PIANC recommendations (ref PIANC Final Report of 3rd International Commission for the Study of Waves) for unreinforced concrete in 1980 are given in Table 13. Depending on the exposure conditions, the answer may lie between these figures and a suggested range of mixes as given in Table 14 below. In all cases it is necessary to carefully consider any departure from local National Standards. Abrasion resistance is obviously important in locations where concrete may be abraded by shingle

TABLE 13 PIANC RECOMMENDATIONS FOR UNREINFORCED CONCRETE 1980 (ref PIANC 1980) Nominal maximum Aggregate mm

Minimum cement content kg/m3

40 20 10

220 to 270 250 to 310 360

28 day cube (size of cube not defined) strength MPa

}

20 to 35

or suspended sand. Abrasion resistance is usually achieved by using strong aggregates and higher strength concrete, not less than C40/50 (cylinder/ cube). Test panels in the UK have shown losses between 2mm and 12mm per annum over a period of 7 years. Exposure cases for abrasion have only recently been proposed in the Eurocode committees, but their adoption is uncertain. Achievement of abrasion resistance depends critically on the finishing operations and curing. There are a number of abrasion tests. The CEB Guide is very informative in relation to “frost” damage which, more correctly, should be termed “freeze-thaw” damage. Freeze-thaw resistance generally increases with reduced watercement ratio, increased cement content, and higher content of air pores. A higher proportion of blend-

TABLE 14 SUGGESTED 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) Maximum water/ cementitious ratio2

Minimum cementitious content1,3,4 kg/m3

1-3

0.55

300

pc, ASTM, I/II/IV/V or pfa blend or gbs blend

100% ASTM IV or V 75% pc : 25% pfa 50% pc : 50% gbs

C30/37 C25/30 C25/30

4,5

0.50

325

pc, ASTM I/II/IV/V or pfa blend or gbs blend

100% ASTM IV 75-70% pc : 25-30% pfa 30-50% pc : 70-50% gbs

C35/45 C30/37 C30/37

6-12

0.45-0.40

3507

pc, ASTM I/II/IV/V or pfa blend or gbs blend

100% ASTM IV 75-70% pc: 25-30% pfa 30-50% pc : 70-50% gbs

C45/55 C40/50 C40/50

Exposure severity rating

Suggested cement type and blends 8, 9 Minimum Dimension of Pour > 500 mm5

< 500mm

Typical Concrete Grade6 (cyl/cube strength10) MPa

Notes: 1

2

3

4 5

Where appropriate, in large sections, advantage can be taken of the lower cement paste and cement contents required for larger aggregate size. Minimum cement content for 80mm aggregate is approximately 40 kg/m3 less. See Table 12. Maximum water/cementitious ratio may need to be reduced to meet National Standards. For any exposure rating, the water/cementitious ratio should be as low as practicable and economic. Minimum cementitious content may need to be increased for abrasion resistance or to meet National Standards, with increased risk of early thermal cracking. The figures for minimum cement content ranges applies to Portland cement only. Larger figures may be required for blended cements. Cement type or blend chosen to control heat of hydration. 40mm maximum aggregate size recommended to allow cement contents to be reduced by 20-40 kg/m3.

6

7 8 9 10

Approximate equivalent grade consistent with the minimum cement content and the maximum water/cement ratio. The accurate equivalent grade for controlling water/cement ratio and cement content should be established for the actual mix, if necessary by trials. Water reducing admixture recommended. Add air entrainment 4-6% for freezing conditions. Pfa and gbs blends require good curing conditions to avoid surface defects. Values are characteristic compressive strengths tested at 28 days in accordance with ENV 206, below which 5% of all possible strength test results may be expected to fall. Cylinder strength applies to 150mm diameter 300mm long cylinders and cube strength applies to 150mm cubes as defined by ENV 206:1990 and tested in accordance with BS 1881: Part 116, 1983.

27

2.4.5

Reinforced Concrete, including Selection of Cover to Reinforcement The durability of reinforced concrete is primarily dependent on countering the effects of chloride induced corrosion of steel reinforcement and is more a design and detailing matter than a materials matter. As developed in 2.3.1 to 2.3.8, durability is a function of environmental loading and depends primarily on: • exposure conditions • cover to reinforcement • cement type • pore structure/water-cement ratio. For reinforced concrete, the principal design parameters include the exposure rating derived from the macro and micro environmental conditions, the cement type, the mix quality as determined by the water-cement ratio, and the cover to reinforcement. As the protective capacity of a given concrete is broadly related to the square of the cover thickness, the provision of appropriate cover to reinforcement is the simplest and most direct way of reducing damage from reinforcement corrosion. The dramatic effect of cover on the service life is illustrated in Figs. 16 and 17. In this context cover must be properly specified and detailed as summarised below. The nominal cover for placing the reinforcement, to be used in the design and stated on the drawings, can be derived by computing a minimum characteristic design value to meet durability (or aggregate size, fire or bond etc) requirements, together with a placing tolerance, which may vary from 5mm to 15mm, or more depending on the standard of control. An illustration of the range of nominal cover which is believed to be necessary for an estimated “design” life of 60 years (say to meet a Design Working Life of 50 years) is suggested in Table 15. The figures include 15mm tolerance but no allowance for abrasion or salt scaling. It will be noted that as the aggressivity increases, the recommended figures for cover are much higher than have often been used, especially for unblended cements. The table also demonstrates the importance of extremely low water-cement ratios for Portland cement, although even such provision is inadequate

Depth: cm

ing materials may unfavourably affect scaling. Exposure classes for freeze-thaw damage were given in 2.3.4. Concrete with moderate water saturation will not suffer from freeze-thaw damage. Salt causes a substantial drop in temperature at the concrete surface during thawing, and the different temperature between the surface and the internal concrete causes internal stress. One of the worst potential damage situations is where there is a prolonged freezing cycle with a source of external sea water. The most serious condition is reported to be when the substrate remains frozen but the surface thaws due to solar gain. Water from the melted snow/ice enters the concrete to form added ice, which re-freezes during the night to cause “squeeze freezing”. Freeze-thaw conditions are likely to be much more serious for roads as opposed to maritime structures. There are a number of standard freeze-thaw tests, such as ASTM and Scandinavian tests. Japanese tests are carried out to determine the change in length of specimens. Salt scaling is a prevalent phenomenon in hot arid conditions. It occurs in maritime conditions and, due to sulfate, in low quality mass concrete in salt flat conditions. The subject is not covered in more well-known guides and reference needs to be made to Fookes (ref Fookes P G, 1993, Bijen J M, 1992), and various papers on the Bahrain causeway by Rasheeduzzafar et al (ref Al-Rabiyah A R, Rasheeduzzafar, Baggott R, 1989) of the King Fahad University of Petroleum and Minerals, Dahran, Saudi Arabia. Salt scaling has affinity with both frost and sulfate attack. The quality of aggregate has a 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 to achieve requisite resistance to corrosion in the case of reinforced concrete. The mix proportions should be selected using mix design methods commencing with the derivation of water demand, which can be restricted to achieve a low water-cement ratio (i.e. less than 0.5) by using plasticisers or super-plasticisers to ensure that adequate workability is available. Unduly restrictive limits should not be imposed on workability where this property is required to ensure the integrity of concrete within the formwork.

Time: Years

Fig. 16 Example of the effect of the thickness of the concrete cover (ref CEB Guide)

28

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)

without larger cover in the more aggressive conditions. Anyone surprised by this table should consult the following supporting references: • Standard specifications for Design and Construction 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 not less than 75mm is advisable (not less than 60mm if examination and repair is easy), and not less than 100mm in ‘severely corrosive condition’ (not less than 80mm if examination and repair is easy). Where the quality of concrete is hampered by difficult construction conditions or the structure requires a long life-time, concrete cover shall be increased. Although cover of precast concrete may be decreased by 20%, it is not advisable to decrease where sufficient corrosion resistance is necessary.”

The selection of appropriate cover for the exposure conditions and the fixing tolerance has a major influence on the minimum wall thickness for a caisson. Minimum wall thicknesses of 200mm and 300mm are not likely to be appropriate for reinforced concrete under extreme conditions of exposure. Thin walled structures are not appropriate for reinforced concrete in extreme exposure. In severe conditions it may be advisable to use designs avoiding reinforced concrete and, instead, to use a composite design. In the submerged and mid to lower tidal part, the sections may be adequate in thin wall caisson design, but the exposed sections under severe exposure may need to be thicker to enable the cover to reinforcement to be increased or, if feasible, constructed in plain concrete. The change in micro environment will often coincide with a change in construction conditions (i.e. from underwater to tidal or above-tide working) and so a composite design may be chosen to improve constructability while at the same time matching durability to different exposure conditions. In the

TABLE 15 SUGGESTED 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) Exposure Severity Rating

1 2

3

4

5 6

Suggested Nominal Cover1,2,3mm 75% pc 50% pc

: 25% pfa : 50% gbs 4,5

70% pc 30% pc 90% pc

: : :

30 pfa 70% gbs 10% ms6,7

100% pc w/c ratio 0.458

100% pc w/c ratio 0.409

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

Includes an allowance of 15mm for workmanship tolerances and reduction of cover during concreting. Add an extra 10mm for prestressing strand to reduce percentage non-compliance of nominal cover to minimal value in recognition of risk of pitting corrosion. A combination of the suggested nominal cover plus concrete mix appropriate to higher exposure rating will provide extended service life. Appropriate mix for exposure severity rating 2: Grade C35/45, minimum cementitious content 370 kg/m3, maximum water/cement ratio 0.45, 20 mm aggregate. See note 10, Table 14, for definition of Grade. Assumed apparent diffusion coefficient at 20°C 3.0 x 10-13m2 sec-1. Appropriate mix for exposure severity rating 5: Minimum Grade C40/50-55/65) minimum cementitious content 400 kg/m3, maximum water/cement ratio 0.40. Appropriate mix for exposure severity rating 6-12: Minimum Grade

Notes:

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 x 10-13m2 sec-1. 8 Assumed apparent diffusion coefficient at 20°C 15.0 x 10-13m2 sec-1. 9 Assumed apparent diffusion coefficient at 20°C 11.0 x 10-13m2 sec-1. 10 Nominal cover of 50mm dictated by bond requirements with 20mm maximum aggregate size and allowing for workmanship tolerances. 11 Blended cementitious mix more suited to the exposure severity recommended. 12 Note that this Severity Rating is for hot arid conditions and infrequently wetted members. See Section 2.3.4. Extra protection may be required by means of coatings or provision for cathodic protection, depending upon application and estimated severity rating.

29

260

Number of results in each 5mm cover interval

220

Characteristic cover is 24mm

Standard deviation is 15.35mm

180 Nominal specified cover 50mm

140

100 Statistical mean of results 49mm

60

20 10

20

30 40 50 60 70 80 Cover (mm) in intervals of 5mm

90 100

Fig. 18 Distribution of cover in practice. Analysis of 1600 covermeter readings for a 13m high retaining wall (ref Concrete Society Special Publication CS 109, 1996)

colder and wetter parts of Europe this problem is less critical, and a change from thin-walled caisson to unreinforced construction often occurs in any case at about mid-tide level, and the problem is greatly reduced. At least half the problem of achieving adequate cover is created by failure to acknowledge the need for adequate steel fixing tolerances. The answer to this problem lies in the German codes and the new Eurocodes, by considering a minimum characteristic cover to meet durability requirements, to which must be added a tolerance, ∆h, to derive the nominal placing cover which is used in the design calculations and the detailing. The resulting compound tolerances of the concreting process can result, in particular, in a wide variation in cover, as illustrated in Fig. 18. Minimum Cover should be regarded as a “characteristic” value, and a margin of at least 10mm and up to 15mm should be added to reduce the rate of failure to achieve the minimum cover to within 5% of measurements (ref DIN 1045, 1988 and Betonwerk and Fertigteil-Technik (Concrete Precasting, Plant and Technology) Issue 51, 1992). This approach applies to measurement of cover in place after concreting. Cover should be checked by covermeter on trial panels at the commencement of work and “as constructed” records of cover should be prepared for all projects. The size of the tolerance ∆h may depend on the class of work, but should be verified by control procedures. Possible specification methods and compliance limits are suggested in Table 16. 30

Spacers need to be made from appropriate materials and adequately distributed. There are several references on spacers (refs Spacers for reinforcement, 1981 Cement & Concrete Association UK, Spacers for reinforced concrete, Concrete Society 1989). There are a number of ingenious modern designs of plastic spacers, but it is possible for these to deform or achieve inadequate bond. For seriously aggressive conditions, spacers need to made of materials equal to or better than the parent mix and treated to improve bond.

2.4.6 Prestressed Concrete The use of prestressed concrete in maritime works is less common. Most engineers have been afraid for chloride effects but there is a substantial history of success in specific applications (ref Gerwick B C, 1990). A recently constructed example in Japan is a new type of breakwater in Sakai, with a perforated caisson screen 16m diameter and 10m high (ref PCI Journal, July/August 1994). However, the corrosion of prestressing steel presents a much more dangerous situation than can occur with non-prestressed concrete. The corrosion of pre-stressed steel proceeds at a faster rate than that of non-prestressed reinforcement under identical conditions, and presents a higher risk of building failure (ref Perl G C and Blades J T, 1993). There is also the same problem as reinforced concrete, due to the detailing of secondary reinforcement such as stirrups and reinforcement to the anchorage zones. One of the principal reasons for successful performance has been the general employment of higher strength, lower water-cement ratio concrete and increased cover and, perhaps, location in temperate or cool conditions. Prestressed piles have performed well in generally submerged and wet conditions in the lower tidal zone in even hot-arid locations, and prestressed decks have performed well in their usual location above the splash zone where, of course, they are not usually subject to an aggressive environment, even in the worst hot-arid conditions. The concept of using prestressing to reduce the effect of cracks is not the positive advantage it may seem, as the pore structure of the cement paste and the state of saturation is more important than the cracks. Concrete strength is usually required by Codes to be higher for prestressed work, but this is not the case in all codes. Cover to tendons in maritime work is usually, but not always, required to be greater than that for reinforced concrete. Some codes limit the permitted chloride content of mixes to half that for reinforced concrete, but other codes do not. A useful comparison of codes is given by Perl and Blades (ref Perl G C and Blades J T, 1993). 2.4.7 Cement Refer to 2.3.5. and 2.3.6.

TABLE 16 SPECIFICATION OF COVER TO REINFORCEMENT Design Dimensions and Acceptance Criteria 1.

When inspected in the forms, prior to concreting, the cover to reinforcement shall be the nominal (target) cover specified within appropriate + and - tolerances, say 5mm, or 10mm.

2.

When checked by covermeter after concreting, by a covermeter calibrated by direct measurement, the cover to reinforcement shall comply with a similar philosophy to that given below.

Design and Detailing Dimensions Nominal (Target) Cover

No more than 5% of measurements less than* mm

No Single Reading less than* mm

Heavy Civils Work Tolerance 15mm 100 80 60

85 65 45

70 55 38

Normal In-Situ Tolerance 10mm (ref Eurocode 2. 5mm ≤∆h ≤10mm) 60 50 40

50 40 30

43 34 25

Precision work not usually applicable to maritime work, except for precast planks and submerged concrete.

Precision work not usually applicable to maritime work, except for precast planks and submerged concrete.

45 35 25

40 30 20

mm

Precision Work Tolerance 5mm (ref Eurocode 2 0mm ≤∆h ≤5mm) Say 50 40 30 Note that this small tolerance can only be used if the feasibility of achieving it has been proven by tests and that results are verified by production control. *

Acceptance Criteria after Concreting - Measured Cover

Note that the standard of control (i.e. as measured by the standard deviation ÷ mean) exceeds nearlaboratory precision

This acceptance criteria must be applied to a specific area in m2, or a specific member face of appropriate size. At least 35 measurements should be taken using a covermeter calibrated by direct measurement, and checked by direct measurement.

2.4.8

Aggregates

The quality of the aggregates, themselves, has a lesser impact on the quality of the finished concrete than may be supposed, provided that appropriate measures are taken to achieve a dense concrete matrix with strong cement paste of low permeability. However, the quality of aggregates has a significant influence on abrasion, freeze-thaw and salt scaling resistance. As a bulk natural resource, it is usually necessary to make the best use of locally available materials wherever possible (including, possibly, re-cycled materials) and it may be unrealistic to set quality limits too high except when standards are set by particularly demanding circumstances. For that reason it is important to refer to national standards and the

results of locally determined aggregate studies for the limits to be applied to aggregate properties in the prevailing environmental conditions. Most national standards set limits for the physical properties of strength and absorption. Where these do not exist it is often necessary to refer to other national standards. Tests and limits particularly apply with regard to alkali aggregate reaction and chloride and sulfate content. The properties of shape, texture and absorbency of aggregates affect water-demand, which have a significant effect on durability. There are a number of very informative publications on the properties of aggregate and comparisons between various national standards (refs Geological Society (UK), Special Publication No.9, 1990, Pike D C, 1990, and BRE publications 243 and 244, 1993) 31

2.4.9 Cracking and Crack Width Refer to 2.3.7 and 3.1.7.

2.4.10 Reinforcing Steel Provisions for reinforcing steel are covered by National Standards. Steel used in coastal works is usually hot rolled high yield deformed. Stainless steel may be necessary in some applications, at a premium of some six times the cost of black steel. However it should be noted that the premium on the overall concrete cost is relatively small, 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 to BS 6744, Type 316 S 33 has 11-14%Ni and 16.518.5%Cr. The lower grade, Type 304 S 31 has 811%Ni and 17-19%Cr. Unified European standards are under preparation and some, including EN 10088-1, which lists the steel grades, are already issued. The standard European terminology is similar to, but not precisely the same as, the DIN system. The terms 316 and 304 are common to UK, USA and Japan. The higher grade to other standards is: France Z6 CND 17.12, Germany 1.4436, Italy X5 CrNiMo 17 13, Sweden 14 23 43. The lower grade is: France Z6 CN 18.09, Germany 1.4301, Italy X5 CrNiMo 18 10, Sweden 14 23 32. On no account should lower grades of stainless steel (BS Type 304) be used in chloride bearing water, the use of at least a hot rolled austenitic BS Type 316 S 33 being necessary. Higher qualities of stainless steel at a further cost are available, but are produced for marine engine, piping and similar applications. A CEB document (ref CEB, 1995, Coating protection for reinforcement) provides a current state-of-the-art report on three coating protection systems; hot dip galvanising, epoxy coating and PVC coating. Fusion bonded epoxy coated reinforcement is available to some national standards (ref ASTM A775, 1990), on the principle of isolating the reinforcement from chloride ions. Both successful and unsuccessful applications are reported and application must depend on the severity of the conditions. Reinforcement is also available in the form of steel, polypropylene and other fibres, and is used in both conventional and sprayed concrete applications. There are standards for materials in USA (ref ASTM A820-90, ACI 544R82) and for design and construction in Japan (ref Japan Society of Engineers Recommendations, 1983). Steel fibres have been used for the reinforcement of armour units and durability is claimed to be better than for bar reinforcement. The use of stainless or “near stainless” fibres is claimed to be promising. 32

2.4.11 Admixtures In some countries (UK in particular) the use of admixtures was overtly discouraged and some trace of this antipathy remains in Specifications which put a barrier on admixtures without express approval. While there is no doubt that admixtures need to be “approved”, it is now often the case that admixtures are essential to provide workability with low water content or cohesion in underwater concrete and therefore need “positive” encouragement, not discouragement. Types of admixtures are covered by national standards or guidelines. Due to the complexities of selecting proprietary products, it is normal to maintain faith in tried and proven manufacturers.

2.4.12 Additional Protective Measures : Coatings, Coated Reinforcement, Cathodic Protection In certain cases additional measures are required to achieve durability. Such measures can include coatings to the exterior concrete surface or to the steel reinforcement, or the recently developed application of cathodic protection to reinforced concrete. Claims are also made for rust inhibition by adding calcium nitrite as an admixture, or for pore blocking admixtures. None of these solutions are as simple as they sound and usually introduce maintenance problems of their own : for example, coatings to concrete require special preparation of the surface such as grit blasting, and require to be maintained; coatings to reinforcement are susceptible to handling damage; cathodic protection systems have limits to life due to cathodic and anodic reactions. However, each measure may add a number of years to the service life and, usually, the better the initial concrete, the better the performance of the additional protective measure. Similar difficulties for the choice of admixture apply to specifying and selecting coatings for concrete. A UK Concrete Society working group aims to issue guidance on this subject in 1997 (ref Guide to Surface Treatment, 1997). Some countries traditionally paint concrete. Coatings may be necessary in some cases to achieve extended performance, either to delay the onset of reinforcement corrosion or protect from chemical attack. Carbonation is unlikely to be a problem in marine structures, and therefore coatings to protect against carbonation are of little advantage. Coatings are usually required to limit the passage of oxygen, carbon dioxide and liquid water, while enabling a certain transmission of water vapour which would otherwise reduce adhesion of the coating. In aggressive conditions, only high-build, high quality products are dependable. These usually comprise expensive systems of acrylics, polyurethanes and epoxy resins built up to a substantial thickness.

Usually, the concrete surface needs to be prepared to a very high standard, using grit blasting, with all the small air holes formed at the shutter surface revealed by this process filled with mortar by hand, because the coating will not be able to bridge the holes. These activities are labour intensive and expensive. Cathodic protection is very well established in applications to steel maritime structures and pipelines. It is now adopted as a repair technique when no other simpler alternative is feasible. It is less likely to be adopted in new construction, but in such cases it is much easier to design and detail the electrical continuity of reinforcement for cathodic protection than it is in the case of retrofixing to an existing structure. Some Italian bridges have been designed with cathodic protection incorporated. It is claimed that the application of a protective current at an earlier stage in operation inhibits the migration of chloride ions, whilst the applied current is much less than would be needed in later years to combat corrosion currents made possible by the depassivating consequences of chloride ingress (ref Pedeferri P, 1992). A significant benefit from the latest electrochemical techniques can be the building in of monitoring circuits. A continuous record of potentials and/or corrosion currents can then be obtained, which can enable problems to be identified and appropriate maintenance strategies undertaken.

2.4.13 Corrosion of Structural Steel The performance of structural steel in sea water is much better known and understood than that of

reinforced concrete. Materials for sheet and bearing piles and similar members met in coastal engineering are covered by National Standards. Advice on corrosion rates and anti-corrosion measures is given in the British Maritime Structures Code (BS 6349), the German Water-front Structures Code, the Japanese Technical Standards and by sheet piling manufacturers. It is usual to either add extra thickness for a "corrosion allowance", or to use higher grades of steel with reduced stress levels. The corrosion rate is usually greatest in the splash and low water zones, less in the inter-tidal zone and least in the submerged and buried zone. High corrosion rates can be experienced at lowest astronomical tide level, where anaerobic corrosion and reduction by bacterial action can occur. There appears to be no merit in the use of special alloy or copper bearing steels. Coal tar epoxy and similar protective treatments have provided excellent protection to steel during recent decades. However, current legislation for environmental protection and health and safety, as well as commercial pressure to reduce construction time and cost is leading to the replacement of traditional multi-coat metallic and duplex organic systems which do not meet these requirements. A range of new “compliant” coatings has been and is being developed which include water-borne coatings, high solid coatings and solvent-free coatings. There is inadequate guidance on these new systems which are, as yet, unproven in practice. This may especially be the case for maritime work. A CIRIA research project was due to be completed in September 1996 (ref CIRIA Project 523, 1996). Where appropriate, cathodic protection can be confidently designed for submerged areas and protective paint treatments minimise the amount of protective current required in the submerged areas.

33

3.

CONSTRUCTION RELATED CRITERIA AND CONSTRUCTION METHODS

3.1

Caissons

3.1.1

Float-Out Loading Refer to 2.1.7

3.1.2

First Grounding Refer to 2.1.8

3.1.3

Caisson Fill Methods and Pressures Refer to 2.1.5

Sea Condition Data and Limits for Construction Risks For each construction stage the designer has to specify the waves which can be tolerated, in height, period and direction. If this wave condition is surpassed by more dangerous waves, the work has to be interrupted. The design wave for different stages of the construction can be different. Float out can often be carried out under more severe conditions than grounding. The rubble on which the caisson will be founded has to resist a certain wave action which normally is more severe than the wave action which has been foreseen during grounding. On the other hand, erosion of the rubble base can be repaired quickly at low cost, and the last levelling on most sites is, in any case, programmed to be carried out shortly before float-out. To enable the designer and the contractor to estimate the wave height for the different construction stages, they have to be provided with wave forecast data. This data should not indicate the probability of the occurence of a certain wave height during the year but it should give information about the time during which a certain wave height is surpassed, which is shown in Fig. 19. From Fig. 19 it can be seen which time of the year a certain wave height (either given as Hs or as Hmax) is surpassed. In the example, a wave height of Hs = 1 m is surpassed during 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 for the float-out processes, has to be given. There will be only a few places in the world where enough wave measurements have been made in the past to construct Fig. 19. On the other hand, the normal wave climate during the construction period is much more significant to the progress on 34

percentage of exceedence

3.1.4

site than the wave height which occurs during a short storm once a year. Some pre-information about the changing weather conditions can be collected from local weather stations, airports and from weather maps. But for the planning and construction of a vertical breakwater, for which the work on water is sometimes restricted to very few float-out dates during the construction period, a better weather service is needed than that required for the construction of a rubble mound breakwater, where the activity can be changed from day to day according to the wave height. There are not many references to the limits for conditions suitable for sinking caissons, or quantification of appropriate limits. Spanish experience suggests that, in the case of caissons of the size used in vertical breakwaters, the sinking operations should be restricted to the following conditions: T1/3 < 7 seconds Hmax < 0.7m - tending to decrease. In periods of calm weather, the required cell filling operation must also be carefully considered. In large caissons the cells have a considerable volume. Consequently the filling operation takes a long time, which must be estimated carefully, in order that the works programme is compatible with the forecast weather conditions. highest wave height Hmax

Fig. 19 Probability for the exceedence of a certain wave height during the year (ref Stückrath T, 1982 and Clutterbuck P G, 1977)

3.1.5 Construction Joints Frequent construction joints occur in nearly all vertical breakwaters, because (with few exceptions) they are built in a discontinuous way by the use of prefabricated concrete elements. Where horizontal joints occur as, for instance, in block type breakwaters, the blocks placed on top of each other can be easily linked together by a slot and key system or by wells and dowels, as shown in Fig. 20. Another way to achieve a good interconnection between small elements is to use inclined joints.

Vertical joints, which are unavoidable for all vertical breakwaters constructed from larger elements, have numerous advantages but they can lead to serious difficulties during construction. One advantage of a vertical joint is that adjoining elements can undergo different settlement. Therefore these joints should not be filled with inflexible materials as long as the elements still undergo differential movement. On the other hand, storms arising during the time in which the joints are left open can lead to high water velocities in the joints and to damage (Agerschou H et al, 1985). Interconnection of adjoining elements is in most cases necessary to distribute the local wave load on to more than one element. Therefore joints which are permanently open are an exception. A male and female slot and key system which has often been used, for instance in Brighton Marina (Agerschou H et al, 1985) or in Helsingborg 1981 is shown in principle in Fig. 21a. This detail has some disadvantages. Free settlement of two adjoining elements is restricted because differential tilting of two elements is prevented. Additionally, during the time of placement of the elements (even, if they are

(a) Slot and key system

(b) Wells and dowels

Fig. 20 Vertical connections at horizontal joints between concrete blocks placed on top of each other

a

replaced by tremie concrete could be recommended, but experience with the use of these construction methods is rare. Surveys of breakwaters (ref Tanimoto, 1983) that describe the displacement of caissons after wave attack, lead to the conclusion that most displacements of vertical breakwaters are just horizontal slips. Therefore a differential movement of adjoining elements, caused by wave attack, can be prevented by a shear connection located only in the bottom slab, as proposed by Lundgren and Juhl, 1995 . Double slot joints between rectangular caissons as used in Spain are shown in Fig 22 and for the quay-wall caissons at Dubai Dry Docks in Fig 23.

Fig. 22 Joints between caissons

3.1.6 Settlement Vertical breakwaters show much greater settlement than most other structures designed and built by civil engineers. There are mainly three reasons for the high magnitude of settlement. a) The sea bed on which the structures are founded is, with few exceptions, loose and fine and cannot be precompacted. b) The rubble mound which is used as a bearing layer under the vertical elements and which has a considerable thickness, especially for vertically composite breakwaters, cannot

b

(a) Slot and key (b) Double slot Fig. 21 Vertical joints between caissons with a circular horizontal cross-section.

placed by crane) the sea must be absolutely calm. Even small wave heights lead to impact stresses in the slot and key, due to the very large masses that will try to follow the orbital swell motion. Therefore a joint with a double female slot which is filled later, as shown on Fig. 21b, leads to a less vulnerable construction method. The vertical open gaps on both sides of the slots are usually sealed with grout- filled tubes or "bolsters". The materials that have been used to fill the gap in 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 concrete has been used. The use of modern flexible materials or a first filling with a soft material which is later

Fig. 23 Typical joint details - Dubai Dry Docks (continued on page 36)

35

Fig. 23 (continued from page 35) Typical joint details - Dubai Dry Docks

36

easily be precompacted and will therefore be compressed by the load of the caissons, and by wave loading. c) Rubble mounds undergo considerable settlement during construction and time-dependent "creep" settlement for many years thereafter. These phenomena are explained below. d) Sometimes the elements have to be placed on foundations for which a trench has been dredged into the sea bed. In soft ground this may involve the classical sand replacement method (ref Barberis MC, PIANC, 1935). In these cases the inflow of soft sea bed mud after the last cleaning and before the placement of the elements cannot be fully prevented, and the depth of fill will undergo both compression and consolidation settlement. In most cases the vertical load of the elements on their foundations (weight minus bouyancy) is increased from zero after first placement to the full load after filling with sand and after concreting the cap. The greatest settlements have to be anticipated during the construction period. Many breakwaters have shown additional settlements during the first years of operation because the settlement is not an elastic movement and because it can be increased by the shaking by waves. If possible, the last layer of the concrete cap which is visible to the eye, should therefore be completed as long as possible after the major construction period. Settlements can never be excluded. They are, even if they are high, unavoidable characteristics of vertical breakwaters. Visible differential settlements of adjoining elements should be minimised as far as possible, and the ugly appearance these constructions can exhibit (because of the uneven surfaces), should be overcome by appropriate detailing features. Although the size and the extent of the settlements can be measured easily and figures have been collected for many harbours, not many publications have been made about settlement, and very little guidance appears in the CIRIA/CUR Manual on the use of rock, 1992. The magnitudes of settlements given in available literature, are as follows: • "settlements up to 1m are normal" (ref Lamberti A and Franco L, 1994) • "settlements of 97cm have been measured but the influence of earthquake cannot be excluded" (ref Ching T K, 1994, page 228) • "Diagrammi dei cedimenti dei cassoni" show settlements of 1m during the first three months and maximum settlements of 1.5m after eight years (ref Ing Mantelli & Co., Volti Harbour, Genova, 1994) • In the Working Group 28 meeting in London on 26 April 1995, Sub-Group B reported the following figures for the vertical breakwater at Gela (Italy): Overall settlement 1m, differential settlement of two adjoining elements 0.2m.

the next increment to equal that which occurred from year one to year three. Fig. 24 illustrates the case of a 20 m high embankment with a typical creep coefficient of 0.524. Substituting in Penman’s expression: H = 20,000 mm α = 0.524 log10 27 = 1·434 = 0 log10 l 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 0 to 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 settlement 100 from 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. Time in years since completion of rockfill

Settlement mm

In fact, considerable information on the time dependent post construction settlement of rockfill is available from dam construction. An immense amount of investigation has been devoted to this subject at the UK Building Research Establishment (BRE) and by American workers. Penman’s paper of 1971 (ref Penman ADM, 1971) reviewed the development of compacted rockfill as a construction material and illustrated its post-construction behaviour. 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 state of the art of rockfill structures was presented in Lisbon 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 Civil Engineers paper by Sharp (ref Sharp B N, 1996) reviews the available data and gives practical case histories of the settlement of quay walls in ports and rockfill used in quay wall construction. The significant factor is that the post-construction settlement of a rockfill embankment is time-dependent, due to the crushing of the highly stressed points of contact between the individual rocks. The pattern is similar to the secondary consolidation of clay, in that it reduces logarithmically with respect to time according to Penman’s expression : δ = Η α (log t2 - log t1) 100 where δ = settlement in mm H = height in mm α = creep coefficient t2 & t1 are any two times from the end of construction for a settlement δ to occur. The dimensions of time (months or years) are immaterial as it is the difference that matters, and the 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 marine embankment could typically be 25 mm per month or more at the end of construction and 5 mm per month one year after completion of construction. Thereafter, the rate of settlement reduces slowly according to the logarithmic law, which means that significant settlement can continue for 20 years. The long term settlement due to this cause is unlikely to be very large, in comparison with the settlement of clay and seabed mud, but could be of the order of 100 to 200 mm over many years. Although this amount is unlikely to lead to “failure” of a vertical breakwater, it may prove unsightly and lead to unnecessary worry and incorrect diagnosis of the cause of continuing settlement. Usually this settlement is made up by additional fill. The logarithmic decline in settlement rate is illustrated in Fig. 24. It will take nine years since completion of fill to double the amount of settlement that occurred from year one to year three since completion of the fill. It will take until year 27 for

*Settlement from year 0 to year 1 approximately 100 mm to 150 mm.

Fig. 24 Logarithmic decline of rockfill settlement (ref Sharp B N, 1996)

3.1.7

Early Thermal Cracking

Control of the temperature of concrete at the placing stage and during hydration is virtually essential for the construction of massive concrete sections. Thermal contraction from the heat generated by hydrating cement results in severe cracking wherever the geometry of the section or the sequence of adjacent pours during construction imposes a restraint to free contraction. The subject was dealt with in the 1930’s in the USA for low cement content mixes in massive dams, but ACI code recommendations do not extend to convenient crack calculation methods for reinforcement design. Detailed recommendations, which are simple to apply, are given in UK publications (ref CIRIA 91, 1990, BS 8007, 1987, Department of Transport BA 24/87 and BD 28/87, 1987) and the phenomena is the focus of interest in Japan for the design of caissons, etc. 37

A major step forward is now possible due to the development of computer software which models the transfer of heat and moisture in concrete. The effect of different mixes and sequences of construction can thereby be compared and the propensity for cracking, stresses etc, determined (ref CEB Design Guide, 1992, Appendix A). Appropriate measures to control the heat of hydration include: • the use of the largest appropriate size of aggregate and water reducing admixtures in order to reduce the cement content demanded by low water-cement ratios • the use of low heat cements, usually involving blending with pozzolans or slag • cooling by the cooling of materials, addition of flaked ice, the injection of liquid nitrogen to negate the heat rise, or by cast-in in cooling water pipes • thermal curing and insulation to minimise heat differences and gradients. Early thermal stresses may occur in the construction 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 between adjacent pours whereas, to minimise early thermal loading, the maximum freedom from restraint would result from continuous casting, such as occurs in slipforming. The old "alternate bay" method, with continuous reinforcement, causes the maximum restraint and is the least favourable. The calculation of thermal crack widths and crack control reinforcement by conventional means are covered in references CIRIA 91, BS 8007 etc. As the cement content has a significant effect on the heat evolution during hydration and therefore the temperature differences applied in the calculations, the temperature effects due to the likely maximum cement content should be used, and very careful monitoring of the actual cement content be made in relation to the calculated reinforcement. The designer must bear in mind that the strength and cement content of the mix in practice may be considerably higher than a “minimum” requirement of a general specification. Ready-mix suppliers may be obliged to overshoot cement contents in order to comply with strength specifications and QA requirements, and this is often a factor which exacerbates early thermal cracking. The type of cracking caused by the two types of restraint is illustrated in Figs. 25 and 26. "Internal" restraint concerns the change in temperature across a thick section, such as a block, whereas "External" restraint concerns adjacent pours. Thermal cracks recorded in caissons in Japan are illustrated in Fig. 27. It is necessary to design the reinforcement to 38

Fig. 25 Internal restraint to early thermal cracking (ref Concrete Society Digest No 2, 1984)

Fig. 26 External restraint for various slab or wall pour sequences (ref CIRIA 91, 1992)

Slit

Cracks at side walls Fig. 27 Thermal cracking in caissons in Japan

replace the formation of a few wide cracks by a number of finer, controlled cracks. Bands of additional reinforcement may need to be calculated for the base of a restrained pour. Methods for calculating reinforcement related to crack width are given in ACI 207.2R and the Japanese Society of Civil Engineers Standard Specification for Design and Construction of Concrete Structures. These methods are based on providing reinforcement to resist early thermal tensile stresses in excess of the early-age tensile strength of the concrete. However, the UK method (BS 8007, CIRIA 91 etc), appears to be more convenient, using the expression : f ρ = ct ∅ R α (Τ1 + Τ2) fb 2wmax where ρ = percentage of steel area fct = concrete strength in tension fb = average steel/concrete bond strength f (In fact ct is usually given direct as fb a quotient) ∅ = wmax= R = α =

bar diameter in mm maximum crack width in mm Restraint factor coefficient of thermal expansion of concrete T1 = difference between centre line peak temperature and mean ambient temperature T2 = maximum temperature difference between adjoining sections. The amount of reinforcement depends upon the joint spacing (i.e. wide or close joint spacing) or for continuous (jointless) construction. For each joint

spacing criteria, the reinforcement percentage is, in effect, inversely proportioned to bar size, and is easy to calculate. The result is not, of course, a rigorously exact figure, but a likely approximation.

3.1.8 Slipforming Vertical slipforming lends itself to the casting of caissons, silos, walls and towers. Because the plastic concrete is placed in the forms which act as moving dies to shape the concrete by an “extrusion” process, the concrete is joint-free and is cast and hardens free of restraint from adjacent pours so that early thermal effects are minimised. Explanations and guidelines for the formwork itself are given in ACI 347R, 1988 - Formwork. Tolerances are recommended in ACI 117 1990. The forms are constructed with a slight taper such that the width between the forms is greater at the bottom than the top. The true wall thickness is measured at the elevation where hardened concrete is maintained in the form. The allowable ACI tolerance for crosssectional dimensions is 20mm. Slipforming of caissons has been described in several references (Cochrane G H, Chetwin D J L, and Hogbin W, 1979, Philip Holzmann literature for Port of Damman, 1978). The slipforming of the tapered cylindrical legs of the Condeep platforms is described by Moksnes J, 1975 and Condeep promotional literature. The slipforming of the Ekofisk artificial island in the North Sea is described by Marion H and Mahfouz G, 1974. A technical discussion on slipforming (ref Fort G B and Davis P D, 1981) reports the procedures for casting the central platform of the Ninian oil field, both in dry dock and afloat, together with other advice and information on mixes etc. A recent review was given by Jones M N and Horne R D, 1996. The method is less successful in dealing with discontinuities, such as windows or slots in the walls. Otherwise dimensional variation should be small and the accuracy in placing reinforcement has been found to be relatively good, due to the location of vertical steel by guides within the forms. The curing of massive vertical slipformed surfaces presents logistical difficulties as recognised in ACI 308:1981 (Revised 1986) “....structures erected using vertical slipforming methods should be cured in accordance with the procedures used in curing other vertical surfaces recognising the particular problem of slipform construction”. The immediately slipformed surface is, clearly, unprotected by formwork in its early hours and is more sensitive to the application of curing activity and, as in the case of horizontally slipformed pavements, there may be little alternative to the early application of curing membranes. For subsequent application of curing activities see 2.3.8 and 3.1.9. 39

3.1.9 Curing As noted in 2.3.8, the purpose of and need for wet curing is now questioned. The wet curing of massive vertical surfaces such as occur in caissons introduces severe logistical problems in fixing curing materials, securing them against wind and weather, and the supply and drainage of water. The need for wet curing of those parts of a structure which are subsequently to remain totally immersed may be particularly questionable. However, wet curing cannot be disregarded for those parts of a reinforced concrete caisson structure which are manufactured in and/or going to be exposed to severely aggressive conditions in a hot arid environment or which are going to be subject to abrasion or frost damage. All of these circumstances require a refined pore structure as influenced by wet curing amongst other factors. Indiscriminate wet curing with cold water, and/or the removal of formwork can lead to shock due to temperature or moisture gradients. For formed surfaces, it is generally advisable to leave the forms in place as long as practically possible. The use of higher strength (i..e. low water-cement ratio) concrete which is less susceptible to curing duration is recommended, as explained in 2.3.8. There appears to be advantage in the recently developed permeable formwork liners, for the controlled removal of or supply of water to the surface. If such liners are left in place after removal of formwork, they provide a protective covering. Vacuum dewatering, to reduce the water-cement ratio of the surface concrete, has been utilised since the thirties, but is more appropriate to horizontal surfaces. The procedures for and timing of immersing caissons in sea water require consideration. Provided the concrete surface is saturated and of a concrete strength which can be cured in a short duration, there is likely to be no disadvantage and even positive advantage of early immersion The “Condeep” oil rig platforms for the North Sea and the Ekofisk central reservoir were slipformed when afloat (ref Marian H and Mahfouz G, 1974, and Moksnes J, 1975 and Condeep promotional literature). There is very great risk of disastrous absorption of salts into concrete which has been left to thoroughly dry at the surface in an arid climate and then suffer unbalanced periods of immersion with long drying cycles. (ref Hansen T C, 1980). A matrix of requirements must be considered when deciding the plus and minus factors for curing, including the appropriate surface pore structure in relation to abrasion or frost resistance or protection of reinforcement, environmental conditions during and after construction, whether further surface treatments are to be applied (in which circumstances curing membranes would be inappropriate) etc.

40

3.1.10 Developments in Caissons Recent designs for composite caissons in Japan include a caisson with a superstructure broken up into wave breaking shapes and infilled with armour units. The arrangement of this is shown in Fig 28 and the stress plot in Fig 29. In the analytical model there are 3482 No. elements and 10,070 No. nodes. New developments in prestressed concrete double wall cylindrical breakwaters, designed by limit state methods, are described in the 1995 FIP Notes by Kiyomia O and Yamada M. See Fig 30.

Fig. 28a Perspective of a composite caisson, with armour unit infill

Fig. 28b Cross section

Fig. 29 Stress analysis of caisson in Fig. 27

Fig. 30 Recent Japanese double cylindrical wall-type breakwater in prestressed concrete. (ref Kiyomiya O and Yamada M, 1995)

41

3.2

Blocks

3.2.1 Blocks from Concrete Blockwork is inherently robust and durable, and was, historically, used for vertical breakwaters in the UK, France and Spain. The older forms, sometimes using smaller blocks, often laid to a sloping batter, are no longer used. Vertical blockwork breakwaters in the Mediterranean were less successful and a list of famous failures includes the Mustapha Jetty at Algiers, and at Catania and Genoa. Guidelines for blockwork breakwaters are given in the UK Maritime Code BS 6349 : Part 7, 1991 which refers also to BS 6349 : Part 2, 1988 for blockwork quay walls, and examples are given in the Japanese Technical Standards for port and harbour facilities and the Spanish Diques de abrigo en España (Breakwaters in Spain), 1988. Blockwork quay walls are normally dry jointed. BS 6349 : Part 7, 1991 recommends that, where settlement is not significant, joints should be sealed and grouted to minimise air and water pressure effects under wave action. However this advice appears impractial and not applicable to separate block construction. Sealing would only be effective if it is achieved by infill pours of in-situ concrete, as described below. A significant difference between blocks in breakwaters as opposed to quay walls must lie in the significance of cracks. In dry bonded quaywall construction, some cracks are likely to occur during placing and preloading blocks in position. As the quaywall remains in compression due to earth retaining loads, a limited and random incidence of cracking is insignificant. In the case of unreinforced armour units and breakwater blocks, such cracking reduces the mass of individual elements and may require further consideration. 3.2.2 Types of Concrete Blocks Modern blockwork breakwater walls can be classified into four main types:(i) With massive blocks of cyclopean dimensions of mass up to 400 or 500 tonne. (See Fig. 31). The block length is equal to the structure width. The blocks are usually stacked to form separate vertical columns which permits independent settlement of each column. When significant settlement is not expected, the blocks can be bonded (staggered) along the length of the wall. Connections between blocks are by mortises in the horizontal direction as shown in Fig. 31, and in the vertical direction sometimes by wells, which are infilled with tremie concrete, and can be armoured by steel dowels. See Fig. 20(b). The wells can have a dual function as the holes for lifting tongs. (See Fig. 32). 42

(ii) Built up of smaller blocks, of the order up to 60 tonne, as shown in Fig. 33, or less as shown in Fig. 34. (iii) Forms of blockwork construction which incorporate large voids or discontinuities in the sea-face, to absorb wave energy. (iv) Composite walls of blockwork on a submerged rubble mound. 3.2.3 Common Problems A 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 achieve horizontal faces to blocks. Lack of planarity may result from the hand screeding of the open top face of a block, or from the tendency of block edges to curl. It must also be noted that, to maintain verticality of a wall, it can be necessary to shim between blocks. Such lack of planarity can result in cracking of some courses of blocks. This problem has been overcome, in the case of solid blocks, by casting the blocks on their sides, such that the seating faces are formed vertically between the forms. This cannot be done in the case of hollow blocks, more usually used for quay walls. 2. Loss of friction on seating faces It is dangerous to use felt or building paper to form the base of a pour. If this material sticks to the block and is not removed, it can cause serious loss of friction. It is preferable to use proprietary "surfectant" (soap-type) of shutter release products and not to use shutter oils or unsuitable oil products, to reduce the risk of loss of friction and damage to concrete faces. Another solution is to cast the blocks on a steel base plate which has been treated with a formwork release liquid which is soluble in water, which dissappears instantly on contact with air. 3. Cracking of blocks during manufacture, handling and placing Cracking during manufacture should be avoided by due attention to the early thermal design of the blocks, formwork and protection during curing. Cracking during handling should be avoided by appropriate lifting and handling methods, and avoidance of premature lifting. It may be necessary to add handling reinforcement to the walls of hollow blocks, although this should be avoided if possible. Cracking during placing and particularly during any preloading may occur as a result of 1. above.

Cracking can be triggered by changes of section at mortice joints, and by the holes formed for interconnection wells and lifting purposes. It is recommended that holes are not made for lifting purposes. It is preferable to use steel lifting points, even though this may necessitate an increased number of lifting points and a special lifting device which balances the lifting load between the extra lifting points. Corrosion of such exposed lifting points will be insignificant under water, or to otherwise unreinforced concrete. 4. Quay wall blocks are usually bedded down by stacking blocks on the completed columns, either to form a preload surcharge to accelerate and therefore "take-up" settle-

Minimum Min 1.51m50 m

Fig .32 Lifting tongs recess and vertical well connection

ment, or to bed the blocks onto the rubble bed (i.e. a pressing down or "paper-weight" effect) This measure is unlikely to be suitable for breakwater blocks, Seaward Side Harbour Side and it may be necessary to construct the in-situ concrete crown Concrete Crown early, in order to achieve this H.W.L. "paper-weight" effect. Early execution will of course lose the Concrete Block benefits to be gained by later execution, as recommended in Foot Protection Concrete Block Foot Protection Concrete Block 3.1.6, which may minimise Armour Stone Armour Store unsightly movement and cracking. The joints in the crown block can be formed on the Rubble slant, to permit uneven Fig .31 Concrete block breakwater - large blocks settlement. Joint spacing in the (ref Technical Standards for Port and Harbour Facilities in Japan) crown block can be of the order of 10 or 12m. The joints are usually sealed with bituminous material.

Fig .33 Bonded blockwork - Spain Marina "Los Gigantes", 1973. Concrete blocks up to 15 Tonne. (ref Diques de abrigo en España, 1958)

Fig .34 Breakwater Eng. Castor - Port of Algeciras, 1935 (ref Diques de abrigo en España, 1958)

43

3.3

Rubble Mounds

For construction aspects of any rubble mound element of vertical breakwaters refer to the references of 2.4.1 and 3.1.6.

3.4

Curtain and Pile Type

Curtain 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 breakwater is constructed with piles alone. Piles are usually of steel tube and curtain walls are reinforced

(a) Single curtain breakwater

or prestressed precast concrete. There are two sub-divisions of this type, namely single curtain wall and double curtain wall as illustrated in Fig. 35. Practical examples of the various types are shown in Fig 36. In the double type of curtain wall, slits or openings, are provided at the front curtain walls. The wave energy is dissipated between the two curtain walls and reflection wave height is attenuated. However, consideration must be given to the risk of scour of the sea bed and the provision of protection rock. When steel piles are adopted, corrosion protection is necessary. The exposure conditions for concrete elements and their fixings are likely to be very severe. Cover to reinforcement in such elements must be adequate to suit the exposure conditions and the materials as recommended in this Report.

(b) Double curtain breakwater

Fig. 35. Types of Curtain and Pile Breakwaters (ref Technical Standards for Port and Harbour Facilities in Japan) upper concrete

(a) At Osaka Port

(b) At Kelhin Port

Fig. 36 Japanese examples of Curtain and Pile Breakwaters

(c) At Hakata Port 44

4.

SUMMARY

The topics examined by Sub-Group C comprised a wide range of subjects related to design, construction and performance. They were mostly chosen to fill in the gaps of the considerations of the other sub-groups, whose tasks were to concentrate on loading and reliability criteria for stability under wave impact. The information presented is drawn from experience and practice worldwide and attention is drawn to facts and references which may not be obvious or available in any single comparable document. This being the case, it is neither possible nor meaningful to abridge the considerations. This Summary, therefore, mainly provides a key to the data and tables given in the text, together with summarised abstracts and recommendations where appropriate. The topics were examined under two overlapping headings: (i) Design Criteria and Materials • different forms of loading other than wave loading, i.e. earthquake, ice, soil etc. • structural analysis and limit states for element design • durability and maintenance, particularly of concrete structures • materials. (ii) Construction Related Criteria and Methods relating specifically to:• caissons • blocks • rubble mounds • pile and curtain type. The text is accompanied by an exhaustive reference list.

4.1

Different Loadings not covered by Sub-Group A

4.1.1 Earthqnake Although advances in computer techniques enable dynamic response to be analysed by finite element methods, the simple equivalent static load method is generally acceptable for breakwater structures. In many countries, and for the obvious example of Japan, the horizontal earthquake load is still calculated by multiplying the vertical dead load and surcharge by a seismic coefficient determined from a number of factors, as set out in Tables 1 to 3.

4.1.2 Ice Pressure Load from ice pressure on a vertical breakwater seldom exceeds the wave load. The effective pressure from ice loading decreases with structure size and there are, at present, no conclusive formulae which can be applied to large works. Therefore, in those countries where ice loading is a consideration, ice pressures are derived from local experience and judgement. Some details are given in Section 2.1.2.

4.1.3 Deleted

4.1.4 Earth Pressures for Structural Design Earth pressure is relevant to vertical breakwaters with rubble or fill placed against them, and to the load from retained materials within caissons. Traditional “working stress” codes recommend “active” or “at-rest” pressure coefficients to be applied to the dry or submerged soil mass, appropriate to different forms of construction. Different approaches are taken in different countries. Due to the problems of reconciling limit state methods for soil mechanics analysis with structural analysis (because the fundamental relationship between load and movement for soils differs from that for structural materials) there is a lack of agreement in the formulation of limit state codes. Therefore, traditional methods still remain as an option in most codes. New structural analysis codes and geotechnical codes now adopt limit state philosophy. Structural analysis to limit state codes requires the application of partial factors for loading cases and materials for the calculation of the ultimate and the serviceability limit state conditions. However the application of limit states and/or partial factors to earth pressure and variable water loading is not as straightforward as for buildings and bridges. The selection of partial factors to match with the reliability and probability of water and wave loading is the task of Sub-Groups A and D. The subject is also discussed in 4.2. There are two distinctive methods of applying limit state methods and partial factors to the structural design of earth retaining structures. One method derives directly from structural design: the partial factors from Eurocode 2 and similar national codes are applied to the characteristic or serviceability limit state loading. The other method derives from geotechnical stability analysis: a partial factor (or, in the case or BS 8002, a “mobilisation” factor) is applied to a parameter, such as tan ∅´. 45

However, the new geotechnical codes tend to concentrate on equilibrium and stability, and do not give adequate clarification of the loads to apply to the structural design of members or to water loading. In maritime structures, the hydrostatic pressure component greatly overshadows the load from submerged earth and there seems to be no point in over-refinement of earth pressure loading unless clearer consideration is given to water loading which is also, of course, associated with variable water loading due to waves. A comparison of seven various national applications of partial factor methods for the calculation of structural members is given in 2.1.4, and is illustrated by an example, compared visually in Fig. 6. The example demonstrates the range of results for calculation of the load on one side of a member in 20 m depth of fill of some 1.5 to 1. (Note that in practice, there will also be water loading on the other side, except in the case of a lock or dry dock, when the water loading is critical). The range of factors lies between the application of the partial factors in the structural codes (i.e. 1.4 or 1.35 on dead load and 1.6 or 1.5 on live load) to the unfactored soil properties, and the less conservative loading from new USA, Japanese and older Scandinavian codes and BS 8002 and the draft Eurocode 7, depending upon interpretation (where the factor is of the order of 1.2). It must be noted that BS 8002 is understood not to relate to maritime structures, and the formulation of an interpretation of Eurocode 7 remains a matter of controversy.

4.1.5 Fill Pressures within Caissons The loading within caissons is generally derived from silo theory. Field verification of this approach is illustrated in Fig. 7. An example of how fill pressures calculated to various national standards compares with the “at-rest” unconfined pressure is illustrated in Fig. 8. The silo pressure of submerged sand is seen to range from 30% to 60% of the unconfined “at-rest” pressure.

4.1.6 Friction There is a surprising divergence in the various national codes between the figures used in design for 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) and the factor of safety between 1.0 to 1.75. In the latest geotechnical approach to limit state codes, factors of safety against sliding or overturning have been overtaken by the assessment of equilibrium at modified soil strength parameters. 46

4.1.7 Handling and Float-Out Loads Loads, which can arise during construction, although transient, can be significant and must be considered carefully. From the viewpoint of ultimate limit state design, a partial factor of γF = 1.1 is suggested. The forces arising from towing can be taken from Japanese standards, as illustrated in Fig. 9.

4.1.8 First Grounding Severe loading cases can arise when a lowered caisson first makes contact with the prepared foundation. In most cases the caissons will never again undergo a comparable distribution of load. These dynamic loads can not be predicted precisely, but the designer can influence and reduce the risk of indeterminate load imposition by various means, including downstand legs which predetermine the location of first grounding.

4.2

Resistance Analysis, Internal Analysis

Structural analysis of caissons can be carried out by the traditional approach, in which the structure is split into sets of beams and slabs, guidance on which is amply given in national codes. Computer methods are likely to be used for two-dimensional frame analysis. For detailed final design it is more likely that full three dimensional model analysis will be used, using finite element analysis. In implementing finite element models, the main problem may be the modelling of soil behaviour, i.e. the definition of stress-strain relationship. The simplest approach assumes a linear unconnected spring relationship, as per Winkler. This simplistic assumption disregards the inter-connection of the soil elements, and these can either be modelled as well or the simpler method used with sensitivity tests on the soil elasticity parameters. It is suggested that as a complex soil model is critically dependent on soil testing and interpretation, as well as its comparison with the stiffness of the structure, it is sensible to test the design against local reductions of ground support. It is emphasised that caution must be exercised in making the transition from traditional working stress design methods to limit state methods which are now general, worldwide. It is not simply a case of adapting partial factors from one national code to another, because the underlying principles of reinforced concrete design may be different. The recommended partial factors in most national codes were derived for land-based building and bridges and relate to broad probabilities of failure drawn from historical precedent. These factors are not necessarily applicable to maritime structures in which the main loading cases are caused by environmental

loads which have to be derived from a probabilistic approach. Similar problems relate to earth pressure loading and there is dispute over whether partial factors are appropriate to limit state considerations for soils, as explained in 2.1.4. Also, it is not necessarily obvious whether loads are to be classified as favourable or unfavourable. The points raised in 2.1.4 and 2.2.3 are intended, also, to provide input to and guidance from the work of Sub-Groups A and D on the appropriate assessment of reliability for wave loading cases.

4.3

Durability of Concrete

4.3.1 General Principles The performance of concrete in seawater is a subject for which knowledge and guidance remains fragmented and ill-understood, despite the existence of practical reports on the experience of Portland cement concrete in the sea since 1850 and reinforced concrete from 1896. The subject is surrounded by myths and lack of understanding. The main reasons for this are that climatic and exposure conditions vary widely, different materials have been used in various countries, and that the properties of cement have changed during the century. A basic reason is, also, that deterioration can take a sufficiently long time such that it can be difficult to connect cause with effect. The mechanisms for the deterioration of concrete structures have not been adequately understood. It is believed that recent work is beginning to resolve the situation, but that a consensus view of appropriate guidance for practitioners will not be available for at least ten years. For this reason, the subject of concrete durability is the largest single clement of the Report. It is discussed in more detail than other items, and contains more reference to data which is either not generally available or collated in a single document. Whereas most forms of concrete deterioration are now adequately covered by national codes, this is not the case in relation to the most dramatic failure mechanism, that is reinforcement corrosion. This can impose severe limitations in relation to the design of complex thin-walled structures such as caissons or light superstructures and, in some cases there may be little purpose in refinement of wave loading analysis if durability presents a significant risk, albeit of a different nature. Durability is not in itself a limit state but a means by which the principal limit states are maintained over the operational life. Current codes of practice deal with durability in a prescriptive manner, and do not provide a rational basis for design of concrete to meet a service life. Durability is not a matter of materials and choice of materials, but a question of a holistic approach to design.

4.3.2 Design Working Life (or Service Life) “Design working life” is the term and definition from Eurocode 1, and has three main implications for 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 be considered by the owner of a structure and the means of achieving this be addressed by the designer. The definitions of service life, design life and economic life require careful consideration, as there are many different definitions in use. The main categories of definition are compared in Table 7. For maritime structures, subject to the probability and return periods of environmental loading, the following definition is recommended in which the definition of Eurocode 1 is supplemented by the rider expressed in italics: “The assumed period for which a structure is to be used for its intended purpose with anticipated maintenance but without major repair being necessary within a probability appropriate to the function of the structure”. Figures for design working lives specific to maritime structures within the classification of Eurocode 1 are given in Table 8, drawn from the Spanish maritime recommendations. It must be noted that a different level of reliability may be adopted for different limit states and causes or modes of failure. Also that, despite the increasing use of the concept of “service” life in respect of structural safety and durability, current codes do not give adequate guidance for analysis to achieve such lives. It is recommended that explicit analysis for a “design life” to satisfy the “design working life” is required for ensuring the durability of maritime structures, and should be adopted in preference to the current “prescriptive” guidance. 4.3.3 Processes of Deterioration The various deterioration mechanisms which affect the durability of concrete maritime structures, the locations in which they are likely to occur, and methods of avoidance are scheduled in Table 9. The most widespread and critical problem is that of chloride-induced corrosion of steel reinforcement, and the sections which follow concentrate on this phenomenon. Adequate guidance on other forms of deterioration is usually given in national standards, as scheduled in Table 10. The dominant factors involved in the durability of concrete, and particularly with regard to chloride induced 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-climatic exposure conditions for the whole structure and its individual elements.

4.3.4 Exposure classification The most important macro-climatic factors are temperature and rainfall. Temperature controls the rate of chemical reactions and the degree of drying out of the cover concrete. Rainfall, humidity and the location of a member in relation to sea level movement control the wetness of concrete. The wetness of the concrete determines the mechanism for the penetration of chlorides and controls the penetration of oxygen to fuel the corrosion process. Contrary to the case of structural steel, timber and masonry, plain concrete and for freeze-thaw damage, reinforcement corrosion is less severe in the regularly wetted tidal and splash zones. In cool and temperate climates, the concrete does not dry out to appreciable depth. However in the infrequently wetted and mostly dry zone above the tidal zone but subject to irregular inundation from seasonal changes in sea level, storms etc., concrete dries out to greater depths. Especially in hot-arid areas such as the Middle-East, and also where elements are sheltered from rain or in artificial climates such as in tunnels, the sporadic wetting of the dried-out concrete enables chloride-laden water to be very rapidly sucked in to greater depth by absorption. The processes of absorption, capillary suction and wick-action lead to much more rapid chloride ingress than the diffusion process which operates in saturated concrete. In a wet climate the chloride concentration at depth is reduced and the penetration of oxygen is limited. The proposed new Eurocode exposure classification system is explained and new suggestions for severity ratings for concrete expressed on a scale of 1 to 12 are set out in Table 11 and Fig. 13. 4.3.5 Influence of Cement Type The weakness of much prescriptive advice in current codes is that guidance on mixes and associated cover thickness to reinforcement is given independently of cement type. The behaviour of the various types of cement is compared and it is concluded that: • modern unblended Portland cement generally has the lowest resistance for chloride penetration and, where severe chloride exposure conditions exist, even in temperate climates, traditional thickness of cover may be inapplicable. There are exceptions in some national products and conditions • blast furnace cements are highly recommended and have been traditionally used in some countries (originally on account of sulfate resistance) and enable more traditional 48

• •

thicknesses of cover to be used. Their tolerance to surface scaling and poor curing is, however, less than for unblended Portland cement other blending materials, such as fly ash and microsilica, have their benefits and limitations sulfate resistant Portland cements (i.e. with C3A less than 5%) are unlikely to be necessary in maritime concrete. A compromise solution is often reached by controlling the C3A to between 5% and 10% for moderate sulfate resistance. In conditions where reinforcement corrosion is not critical and, especially in colder waters, the long term effects of sulfates may lead to a need for low C3A Portland or slag cement.

4.3.6 Influence of cement content As is well known, the quality of a concrete mix in relation to both strength and durability (as related to the pore structure) is controlled primarily by the water-cement ratio and the unit water content. The water-cement ratio is therefore more important as a parameter to be specified than is cement content. The cement content is established, mainly, by dividing the water demand for a given mix by the watercement ratio. As it is desirable to use the lowest possible water-cement ratio to achieve durability, (generally the requirement for durability may be more onerous than for strength) and to reduce water movement and shrinkage effects, the cement content is controlled by the water content required to achieve appropriate workability. Both water and, it follows, cement content can be reduced by the judicious use of a range of water reducing admixtures which still enable adequate workability to be achieved at lower water content while at the same time reducing the heat of hydration consequences at higher cement content. Both water demand and cement content depend on the type of cement and on aggregate size and grading. The effects of varying aggregate size from l0mm to 80mm are scheduled in Table 12.

4.3.7 The Influence of Cracking The causes and consequences of cracking have often been misunderstood. Early thermal cracking caused by restraint to shrinkage during cooling from the rise in temperature due to heat of hydration is a main cause of cracking which was previously, and erroneously, attributed to drying shrinkage.

Cracking caused in the plastic state can be prevented by good mix design, protection against drying winds and by good curing under arid conditions. Most structural codes have crack width limitations, for flexural and early thermal cracking but the significance of these specific crack widths has been over-estimated. Once crack width exceeds 0.1mm there is no significance in relation to the ingress of deleterious substances. Therefore it is recommended that the cover to reinforcement should not be reduced for crack width reasons, despite the increase in crack width at the surface. The reason for this is that, for flexural cracks, the crack width at the reinforcement is independent of surface crack width, as the cracks are ‘V’ shaped, and the chloride and oxygen ions penetrate everywhere through the pores, and not just at cracks. Cracks do not significantly affect freeze-thaw damage as the scaling caused by freeze-thaw is, again, due to the effect of frost on water-filled pores. Cracking may, of course, be more significant in the case of unreinforced concrete if it reduces the mass of armour units or blocks.

4.3.8 Influence of Curing It appears that prolonged water curing in wet and temperate climates may be of limited advantage and may even lead to adverse effects such as thermal shock. It may be essential in hot and arid climates. As the duration of curing is inversely proportional to water-cement ratio, adoption of a low water-cement ratio enables the curing period to be reduced.

4.3.9 Monitoring and Maintenance Inadequate guidance on this strategic topic is available in the literature and national codes, but it appears that it is, at last, receiving more attention. Regular inspections should be carried out at least once per year, most likely following the winter storm period. The principal objects of the survey are to determine: • the integrity of armour units and elements of the structure • indication of movement and settlement • scour. It is essential to record “base-line” measurements of line and level immediately on completion of construction. This should include “as constructed” measurements of cover to reinforcement and crack and damage mapping. Computers, underwater video recorders and corrosion measurement devices can now be utilised.

4.4

Materials

4.4.1 Rock and Rubble Reference is made to the CIRIA/CUR manual on the use of rock in coastal and shoreline structures, 1992. 4.4.2 Filling and Backfilling Current requirements are outlined in 2.4.2, including the recommendation that measures may be necessary to increase the density of infill material. 4.4.3

Concrete - General. - Design, Detailing and Workmanship The requirements for achieving durability of concrete in maritime works will usually outweigh the requirements for achieving strength and note must be taken of the factors outlined above. Many problems can be “designed out” by good detailing and specification. Reliability of both design and workmanship can be improved and “gross errors” avoided by employing the discipline of Quality Assurance and Quality Control audit procedures for both design and workmanship. 4.4.4 Unreinforced Concrete The factors affecting unreinforced concrete (more usually termed “mass concrete” in Europe and UK) primarily concern deterioration of the exposed surface and include freeze-thaw, abrasion, and sulfate attack. These forms of deterioration have similar and overlapping effects, and are described in more detail in 2.4.4. Early-thermal design is important for crack avoidance. Table 14 sets out suggestions for the choice of water-cement ratio, and hence minimum cementitious contents and grades, for various cement types and blends and minimum dimensions of pour, for the range of exposure ratings on a scale of 1-12 suggested in Table 11 and Fig. 13. 4.4.5

Reinforced Concrete, including the Selection of Cover to Reinforcement The durability of reinforced concrete requires consideration of the same factors which affect unreinforced concrete, together with the major phenomenon of chloride-induced corrosion. The main conflict point in the design and production process is the selection of and practical achievement of the appropriate cover to reinforcement. This is more a design and detailing matter than a materials matter. As the protective power of a given concrete is broadly related to the square of the cover thickness, the provision of appropriate cover is the simplest 49

and most positive way of reducing corrosion damage. The cover to be specified is influenced by the exposure severity rating, the cement type, the mix quality as determined by the water-cement ratio and the placing tolerance which can be achieved. The minimum cover considered necessary for corrosion protection should be regarded as a “characteristic” value and a margin of at least 10mm to 15mm should be added to the figure in order to reduce the rate of failure to achieve the characteristic value to within 5%, by analogy with concrete strength compliance. Without this margin it is statistically impossible to achieve the necessary cover as, in practice, the variation in position of reinforcement about the mean position exceeds common perception, as illustrated in Fig. 18. Suggested nominal cover for different severity ratings, cementitious materials and qualities is scheduled in Table 15. A nominal cover of 50mm is the lowest practicable figure and is only suitable for the lowest severity rating and using blended cements. Nominal cover thicknesses between 75mm and 100mm have to be considered as normal. For severe exposure combined with unblended Portland cements, it may be necessary to double the cover. Methods of specifying cover in relation to a characteristic value, with separate acceptance criteria before concrete is placed and after concreting, are suggested in Table 16. The selection of appropriate cover and fixing tolerance has a major influence on the member thickness for caissons and thin precast units. Minimum member thicknesses of 200mm or 300mm are difficult to achieve under aggresssive conditions. Under such conditions it may be necessary to change the member type and section in the vulnerable upper tidal and splash zone, and either to adopt increased cover or, if feasible, change to plain concrete in this zone. 4.4.6 Prestressed Concrete Prestressed concrete is less common for maritime works but there is a substantial history of successful use in North Sea oil structures and in recent caisson breakwaters. Reasons for success include the necessity for higher quality concrete and larger cover, on structural grounds, and the high state of saturation in cold and wet climates. Prestressing reduces the problem of cracking but does not provide any help in reducing the ingress of chloride ions and oxygen through the body of the concrete. The corrosion rate of prestressing steel is understood to be faster than that of lower stressed steel and the detailing of secondary reinforcement presents the same problems as ordinary reinforced concrete. 4.4.7

50

Cement Refer to 4.3.5 and 4.3.6

4.4.8 Aggregates The quality of aggregates has a lesser impact on the strength and quality of concrete than may be supposed, but it has an influence on abrasion, freeze-thaw and salt scaling resistance. It is usually necessary to make the best use of locally available materials. National standards usually provide sufficient guidance. 4.4.9 Crack width Refer to 4.3.7 and 4.5.6 4.4.10 Reinforcing Steel Provisions for reinforcing steel are adequately covered by national standards. In addition to normal black steels, various coating protection systems are also available, but have not been greatly used in maritime works and may be at a disadvantage in aggressive situations. Stainless steels have not often been used but may be appropriate provided the correct grades are used. The price penalty of six times that for black steel is less when viewed as a proportion of overall cost or the cost of premature failure. Steel fibre and non-metallic fibber and strand reinforcement appears promising. 4.4.11 Admixtures Admixtures are covered by national guidelines. Their use should be positively encouraged in order to provide adequate workability with low water content mixes. 4.4.12 Protective Measures such as Coatings, and Cathodic Protection In aggressive conditions only high-build, high quality coating products are dependable, which are expensive and require the concrete surface to be prepared to a high standard by grit blasting and other means. Cathodic protection is not regularly used for the protection of new construction, although in some cases allowance for later implementation are made by ensuring continuity of reinforcement and facility for electrical connection. There is growing experience of its use as a repair technique where simpler alternatives are not feasible. 4.4.13 Corrosion of Structural Steel The corrosion performance of structural steel in maritime conditions is much better known that that of reinforcing steel embedded in concrete. The corrosion rate is usually higher in the splash zone and at low astronomical tide levels, and very low in deep water. Either an extra thickness of metal as a “corrosion allowance”, high duty coating or cathodic protection can be used. Successful traditional coatings, however, may no longer meet environmental and health and safety regulations for application and there is, as yet, inadequate experience with some new water-based systems.

4.5

Construction Related Criteria and Methods - Caissons

4.5.1

Sea Condition Data and Limits for Construction Risks Wave forecast data must be available to enable the designer and the contractor to estimate the sea state at each stage of construction, especially for float-out, grounding and filling of caissons. This data should include information on the proportion of time during a year in which certain wave heights are not surpassed and the length of windows for calm weather.

4.5.2 Construction Joints Construction joints are an important feature of vertical breakwaters. Horizontal joints seldom give problems. Vertical joints are necessary to allow differential settlement to occur between adjoining elements, but at the same time the interconnection of elements is required in order to distribute the load from local wave attack over more than one element. Typical examples of jointing methods are illustrated in Figs. 21, 22 and 23.

4.5.3 Settlement The magnitude of settlement observed for vertical breakwaters is higher than for most other forms of construction, for reasons explained in 3.1.6. Settlement is rarely critical but the range of likely settlement and differential settlement should be anticipated, and the visible effects of settlement should be minimised by appropriate detailing features. Examples of measured settlement are quoted, including absolute settlement of up to 1.5m and a differential of 0.2m. The contribution to settlement caused by the time-dependent consolidation of rockfill for many years after construction is not generally appreciated. Examples are given using the logarithmic expression published by Penman and others.

4.5.6 Early Thermal Cracking Thermal contraction from the heat generated by hydrating cement results in severe cracking wherever the size or geometry of the section or the sequence of adjacent pours during construction imposes a restraint to free contraction. Control of the temperature and temperature gradient in large unreinforced elements is essential to restrict cracking. The forms of restraint and expressions for calculating the reinforcement required to restrict cracking to acceptable

orders of magnitude in monolithic reinforced construction are given. Measures to control the heat of hydration rise are outlined. It is recommended that the early thermal design is checked against the actual cement contents used in the works, as these may exceed values assumed in design. Failure to make this check often results in problems.

4.5.7 Slipforming Slipforming is a suitable method for casting caissons. It is a method which lends itself to a highrate of production. Because the concrete is cast joint-free and hardens free of restraint, problems of construction joints and early thermal cracking are minimised. Practical guidance available in the literature is given and problem areas identified.

4.5.8 Curing The need for and relevance of wet curing in all circumstances is discussed in 4.3.8 and 2.3.8 and, in the case of massive caissons, introduces severe logistical problems. For caissons, the timing of immersion has a great influence on the rate of production. If the concrete is saturated, and of a sufficiently high grade which can be cured in a short duration, early immersion may be beneficial. Concrete left to thoroughly dry in the surface layers in arid climates can lead to disastrous absorption of salts into the concrete upon immersion.

4.5.9 Development in Caissons Some recent developments of composite and prestressed concrete caissons in Japan are illustrated in Figs. 28, 29 and 30.

4.6

Blocks

Blockwork is an inherently durable method, especially for constructing quay walls, which has historically been used for vertical breakwaters. Guidance is given in Japanese, Spanish and UK documents. A degree of cracking is permitted in quay walls but could endanger breakwaters. Typical types of blockwork for modern breakwaters are illustrated in Figs. 30, 31, 32, 33 and 34. The problems 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 height climate, and for a soft sea bed. Various types of modern single and double wall breakwaters are illustrated in Figs. 35 and 36. 51

5.

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