Abutment of Bridges: Functions, Types, and Design

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September 4, 2022

End supports and intermediate supports are two separate categories of bridge substructures. The intermediate supports of multi-span bridges are referred to as “piers,” while the end supports are typically referred to as “abutments.” Abutments and piers of bridges are typically built from in-situ concrete. As a part of the bridge, the abutment connects the bridge to the approach roadway, gives the bridge superstructure vertical support at the bridge ends, and retains the roadway earth materials from the bridge spans.

Typically, bridges are built as part of a railway or road highway project. Although the cost of the bridges may only make up a small portion of the overall contract, the construction of the bridge substructures can significantly affect the overall contract schedule because it invariably falls on the critical path for construction and typically takes place concurrently with earthmoving and drainage operations. More than half of the costs of a bridge is frequently spent on the foundation that supports the bridge deck.

Types of Abutment

The selection of appropriate abutments for a bridge should be made at the same stage as the choice of the deck superstructure. There are many types of abutment in use all over the world. Abutments can be categorised into the following;

Solid or full height abutments
Skeletal or open abutments
Mass concrete bankseats
Integral abutments
Semi-integral abutments
Reinforced earth abutments

The criteria for the bridge’s design must be taken into account while choosing an abutment type. Bridge geometry, needs for the road and riverbanks, geotechnical conditions, right-of-way limits, requirements for the architect, and other factors might be among them. The ability to compare the benefits and drawbacks of the various types of abutments will help the bridge designer make the best choice for the bridge construction from the outset of the design process.

open spill through abutment — **Figure 2:** Spill-through or skeleton abutment

Cantilever Abutment Walls

For bridges without integral abutments, the T-section reinforced concrete cantilevered wall has remained the most popular method of construction for the solid wall type of bridge abutment. To meet various needs, the core concept has been modified in a number of ways.

For right bridge decks with spans under 12 metres, sloping abutments for aesthetic or clearance reasons, and counterfort walls for heights of 10 metres and above, propped cantilever walls are frequently employed. The overall height of a solid wall abutment is automatically in the range of 7-9 m because the minimum headroom for new highway bridges is often higher than 5.1 m. Because mass concrete retaining walls are not cost effective at this height, reinforced cantilever abutment walls are now used extensively.

The simplicity of this form of construction and the similarity with cantilever retaining walls also accounts for its economic success and popularity.

Free Cantilever Abutment Walls

For heights of 6 to 9 metres, plane cantilever abutment walls are the most popular type of construction, and despite their size, the main concrete wall is frequently poured in a single lift. The wall stem typically measures between 0.9 and 1.2 metres in width, making it possible for someone to enter the reinforcing cage while it is being constructed. The base will often be 0.4–0.6 times wider than it is tall, and the toe may extend 1.0–2.0 m in front of the wall.

The soil foundation conditions and available sliding resistance will, nevertheless, affect the base’s physical proportions and dimensions. Figure 6 shows a typical illustration of a cantilever abutment wall with horizontally cantilevered wing walls.

FREE CANTILEVER ABUTMENT — **Figure 3:** Side view of a cantilever abutment

Active earth pressure conditions are typically used for overturning, sliding, and bearing pressure calculations where an abutment wall can rotate around its base or slide horizontally. The lateral earth pressure behind the abutment wall has typically been assumed to be in the at-rest state for walls that are tightly supported, such as on a mix of vertical and raking piles.

To account for high pressures during the backfill material’s compaction, the wall stem design is typically based on at-rest conditions in all circumstances. On the assumption that the abutment wall solely functions as a vertical cantilever, the design forces are frequently estimated using a metre-wide strip. There is a compelling justification for taking the structure’s three-dimensional behaviour into account if wing walls are joined to the back of the abutment.

The main abutment wall’s need for vertical reinforcement can be decreased to a nominal proportion of the cross-sectional area when the weight of the wing wall and significant corner splays are combined. Figure 4 depicts an idealised system of forces acting vertically and horizontally on a straightforward cantilever wall. It is foolish to rely on passive pressure at the front of the wall since excavations for highway services may be introduced along the foundation’s toe, entirely removing the soil.

forces acting on abutment — **Figure 4:** Typical forces acting on a bridge abutment

Counterfort Abutment Walls

For heights more than 10 m, where the percentage reinforcement in a free cantilever becomes quite large, counterfort abutment walls become economically viable. To increase flexural stiffness and resist the lateral earth pressures created by the depth of backfill material, triangular counterforts are added to the back of the abutment wall slab.

The reinforcement and formwork surrounding the counterforts make the building more challenging, and it is more difficult to physically compact the backfill. The counterforts are vertical cantilevers that are separated at about half the height of the wall. Although the wall slab naturally spans the shorter horizontal distance between the counterforts, it can be treated as a slab clamped on three sides, allowing the wall thickness to be decreased. The heel of the base slab also spans between the counterforts.

The primary tensile reinforcement’s anchorage length at the back of the counterforts, however, is a limiting element, therefore there is typically minimal room for thickness reduction.

Propped Cantilever Abutments

Bridge decks up to a 12 m span have minimal longitudinal movement, making it possible to employ the deck as a strut for square or skew-free bridges. Although the abutments can be planned as a supported cantilever, at-rest earth forces are typically assumed for the design of reinforcement at the back of the wall and footing, as well as both stability and bearing pressure calculations, due to the inflexible character of the structure.

Since complete fixity of the foundation is improbable, it is common practise to estimate the front face reinforcement in an abutment by assuming that the wall is pinned at both the deck and base levels. The rear curtain walls at the top of the abutments are made to withstand the propping force and are typically used to separate the deck from the top of the abutments.

To prevent rotation and horizontal displacement of the abutments, it is frequently important to specify that the initial backfilling before the deck is built should be kept to a maximum of 50% of the abutment height. The deck’s completion is then postponed until the backfill behind the abutment walls is finished.

Open Abutments

The type of end supports required to extend a bridge’s central span and produce neighbouring “open side spans” are known as “open abutments,” and they are frequently used in construction terminology. In this case, abutments come in two different basic varieties; a subterranean reinforced concrete or piled skeleton or “spill-through” abutment formed at or below previously existing ground level beneath an embankment slope, or a mass concrete bankseat located at the top of the slope and includes a side span.

A three-span deck with intermediate piers and end abutment supports is an alternative to a single-span deck with solid cantilever abutments. Therefore, the prices of two intermediate piers, two end abutments, and two additional deck spans may be contrasted with the costs of two massive cantilever abutments, related wing walls, and chosen granular backfill. The choice of a three-span open structure must also take into account aesthetics, sightlines, flood relief, and pedestrian safety.

Bankseats

When the foundation level is near to the existing ground level, simple mass concrete or minimally reinforced sections may be used for abutment supports at the top of cuttings. This kind of structure is typically “stepped out” in sections to lessen foundation strain and keep the force that results on the bankseat inside the middle part of its base. To limit the immediate region of backfill behind the wall, little wing walls that hang easily from the back of the bankseat can be used.

Bankseats can also be utilised on embankments, where they can either be supported directly on pile foundations or allowed to settle with the fill. In the latter scenario, pile downdrag due to embankment settlement might lower the payload of the pile group unless isolating sleeves are utilised. Driving raking piles for a bankseat can be difficult at an embankment’s edge and is not advised if the embankment is anticipated to settle.

A solid abutment wall with substantial wing walls is usually more expensive than using a bankseat, intermediate pier, and additional deck for the side span. This is especially true for small bridges, but for wide constructions, the closed abutment is typically more cost-effective because the cost of the wing walls remains constant and decreases as the width increases.

Spill-through abutments

This type of abutment (shown below) is composed of two or more buried columns supported on a single foundation slab and topped by a cill beam to support the deck structure. To minimise long-term settlement, the backfill must be carefully compacted around the columns since it overflows between the legs.

It is frequently employed in embankment situations where a suitable foundation can be located at the original ground level. In this situation, it might be a more affordable option than a bankseat that is supported by piles pushed through the embankment fill. Figure 9 depicts a typical spill-through abutment bridge, however the completed embankment makes it impossible to see the abutment’s legs.

Since very few field tests have been conducted to ascertain the long-term movements and ground pressures on the subterranean structure, design assumptions for this sort of abutment vary substantially. Assuming full active earth pressure across the whole width of the abutment, regardless of the soil that pours between the columns, is one conservative, straightforward design strategy. While the fill between the columns may arch or experience “drag effects,” the columns and cill beams are typically thought of as being loaded by active earth pressure.

Integral Abutments

Conventional bridges typically include expansion joints and bearings inserted between the superstructure and the abutments to account for relative movement and prevent temperature-induced strains from building up between the superstructure and abutments. These expansion joints and bearings, however, may result in significant maintenance issues.

The concept of physically and structurally joining the superstructure and abutments to produce what is known as an integrated bridge has gained popularity as a result of the issues with conventional bridges that feature joints and bearings. All of the aforementioned issues with joints and bearings are prevented by integral bridges.

The abutments are however compelled to move away from the soil they hold onto when the temperature drops and the deck contracts (for example, in the winter), and towards the soil when the temperature rises and the deck expands because of the integral connection between the superstructure and the abutments (e.g. in the summer). Because of this, the soil behind the abutment experiences temperature-induced cyclic loading from the abutment, which might result in substantially higher earth pressures than originally intended.

Reinforced Earth Abutments

Modular facing panels, often made of pre-cast concrete, earth fill, and soil reinforcement make up a reinforced earth wall. The wall is constructed by repeatedly performing a series of tasks at various levels, including installing face panels, putting earth fill in place and compacting it, laying reinforcements (geotextiles), and putting more earth fill in place and compacting it. Until the necessary height is attained, the processes are repeated.

The facing panels shape the surface, enabling the construction of nearly vertical walls, and the finished wall is able to resist lateral pressure through friction along the reinforcing. When a bankseat is built on top, reinforced earth can be used as part of the abutment. To minimise any local loading effects that could result in local deformations on the face of the wall in this situation, the bankseat is often positioned back from the top of the wall.

Reinforced earth walls are widely used in conjunction with other types of abutment structures to create affordable retaining walls around bridge approaches.

Abutment Foundations

Most abutments are generally supported by either spread or piled foundations. Three issues need to be considered in the choice of foundation:

the available bearing capacity of the undisturbed natural soil at the site;
the settlement that the foundation will undergo (and impose on the superstructure); and
the tolerance of the abutments, deck, etc., to the expected differential settlements.

Most of the time, when the soil’s bearing capacity is sufficient to sustain a spread footing with minimal settlements, this will be the most cost-effective laying alternative. If rock is present at the founding level, this will necessarily result in a spread footing solution. Most dense sands, granular soils, or stiff clays will give appropriate bearing capacities for spread footings.

Abutment foundation — Types of abutment foundation

When soft compressible soils are present or the abutment is situated at the top of a steep embankment, piles are typically used. If a top-down strategy is used for building bridges in cutting, the use of piling can streamline the process.

In many cases, it might be more cost-effective to remove any soft material and replace it with well-compacted granular material or mass concrete rather than using pilings if a suitable soil or rock is present within a reasonable depth of the founding level.

Although differential settlement is difficult to predict with any degree of certainty, past experience indicates that it might be as much as two-thirds of the maximum total projected settlement. While spread foundations may be adequate for some sites based on their bearing capacity, their size frequently results in comparatively substantial overall settlements when compared to the settlements that a well-designed pile foundation is likely to encounter (typically less than 10mm at working load). The ability of the deck structure to contain the anticipated differential settlements may therefore dictate the choice of foundation method.

Abutment Approach Slab

One can anticipate that the embankment fill next to the deck will settle significantly (and perhaps a few per cent of the fill height). Without specific precautions, the intended vertical alignment of the highway pavement would be disrupted, which will result in a bad ride quality for vehicles on the approaches to bridge decks. There are two methods that could be used: a granular wedge next to the abutment or structural run-on (approach) slabs.

An approach slab will provide a smooth transition between the relatively flexible approach pavement and the nonflexible bridge superstructure by bridging across the settling fill immediately behind the abutment. However, for integral abutments backfilled by granular materials, the backfill will become gradually compacted under horizontal cyclic loading from the abutment and a void will form under the run-on slab, which may cause damage to the slab under traffic load if it has not been designed for this case.

Design of Abutments

The primary function of an abutment wall is to transmit all vertical and horizontal forces from a bridge deck to the ground, without causing overstress or displacements in the surrounding soil mass. The abutment wall also serves as an interface between the approach embankments and the bridge structure, so it must also function as a retaining wall.

The type of the bearing supports, if any, determines how much a bridge deck and an abutment wall interact with one another. Integral abutments are becoming more popular since they eliminate the need for additional bearings and the associated maintenance costs. It is simple to idealise the impact of bearing type, end fixity, or free supports throughout the design phase. It is more difficult to predict the impact of ground movements brought on by settlement, mining subsidence, or earth tremors, and these impacts must be taken into account specifically for a given structure.

The primary vertical loading acting on an abutment is due to the dead load and live load reactions from the bridge deck. Additional loading arises from the self-weight of the abutment, earth pressure, wall friction between the backfill and the abutment, and live loading immediately behind the abutment.

Traction and braking forces due to live loads on the deck are carried at the fixed bearings and may represent a substantial overturning moment on a tall abutment. Although these forces are applied to localised areas of the deck, they can usually be treated as a uniform load across the width of the abutment.

Loadings on Abutments

A number of actions are possible on abutments which must be thoroughly accounted for in the designs. Some of the actions are as follows;

Soil Loading

The earthfill retained at the back of abutments exert earth pressure as typical for retaining walls. This may be accompanied by water hydrostatic pressure if adequate drainage is not provided at the back of the wall. In a situation where a wall can move by tilting or sliding and the backfill is a free draining granular material, active pressures are assumed.

Vehicle loading (surcharge)

In the simplest case, for example a distributed load (q kN/m²) at the ground surface, such as an HA loading, an additional stress equal to K_aq can be added to the earth pressure assumed on the back of the abutment. For the design of highway bridges, a live load surcharge of 10 kN/m² for HA loading and 20 kN/m² for 45 units of HB is often used. For rail loading either a UDL of 150 kN/m along each track, applied over a 2.5m width, or an RU and RL surcharge of 50 kN/m² and 30 kN/m² respectively is taken over the track area.

Compaction Pressure

The application of compaction plant, such as heavy vibrating rollers, to abutment backfill in layers leads to temporary but quite large increases in both vertical and horizontal stress within the fill. Some of these stresses remain locked into the fill, and can give considerable additional lateral loading on a cantilever abutment, particularly over the depth just below the top of the wall.

Swelling Pressure

Compaction of cohesive fill produces even greater increases in lateral earth pressures than in granular fill, of the order 0.2–0.4 times the undrained shear strength. But for such clays the more significant issue is likely to be lateral swelling pressures. For clays placed relatively dry, a relaxation in lateral stress has been observed immediately after compaction.

However, as rainwater enters the fill, swelling starts to occur. In situ determinations of the average lateral stresses within a 6m high abutment backfill of London clay showed that horizontal total stresses rose up to 180 kPa near to the centre of the embankment, and up to 70 kPa close to the wing walls. Another pilot-scale experiment observed average lateral pressures on a 3m high wall of the order of 100 kPa. Given that these figures are of the order of many times higher than the commonly assumed equivalent fluid pressure, it is suggested that cohesive backfill should not be used behind abutments.

Effects of seasonal deck expansion and contraction

Longitudinal movements in the bridge deck due to creep, shrinkage and temperature changes cause forces at bearing level on non-integral abutments. The magnitude of these forces depends upon the shear characteristics or frictional resistance of the bearings. The coefficient of friction of most bearings lies in the range f_i = 0.03–0.06. The frictional force is derived from the nominal dead load and the superimposed dead loads on the deck.

Integral abutments do not have bearings, and therefore the backfill they support is subjected to seasonal increases and decreases in horizontal strain. The deck is stiff relative to the backfill and the soil provides insufficient restrain to prevent movement.

Load Combinations for Abutments

BD 30.87 – UK Standard

Case 1:
Backfill + Construction surcharge
Wall backfilled up to bearing shelf level only.

Case 2:
Backfill + HA surcharge + Deck dead load + Deck contraction

Case 3:
Backfill + HA surcharge + Braking behind abutment + Deck dead load

Case 4:
Backfill + HB surcharge + Deck dead load

Case 5:
Backfill + HA surcharge + Deck dead load + HB on deck

Case 6:
Backfill + HA surcharge + Deck dead load + HA on deck + Braking on deck
(Not applied to free abutment if sliding bearings are provided)

Load Combinations (European Standards)

Case 1:
Backfill + Construction surcharge

Case 2:
Backfill + Normal Traffic Surcharge + Deck Permanent load + Deck contraction/shrinkage

Case 3:
Backfill + Normal Traffic Surcharge + Deck Permanent load + gr1a on deck

Case 4:
Backfill + SV/100 and SV/196 Surcharge + Deck Permanent load + gr1a (frequent value) on deck

Case 5:
Backfill + Normal Traffic Surcharge (frequent value) + Deck Permanent load + gr5 on deck

Case 6:
Backfill + Normal Traffic Surcharge (frequent value) + Deck Permanent load + gr2 (ψ1LM1 with braking on deck)
(Braking not applied to free abutment if sliding bearings are provided)

Case 7:
Backfill + Deck Permanent load + gr6 (LM3 with braking on deck)
(Braking not applied to free abutment if sliding bearings are provided)

Stability of Abutments

The stability of an abutment should be checked for three basic modes of failure:

sliding;
overturning;
overall instability.

Sliding

When passive resistance in front of the toe can be relied upon, the minimum factor of safety taken in design is normally 2.0. If the passive pressure contribution is neglected, then a minimum factor of safety against sliding is usually 1.5. A shear key is sometimes provided in the base slab to mobilise greater soil resistance when otherwise the resistance to sliding is inadequate.

Overturning

Overturning is checked by taking moments about the toe when the most adverse load combination is acting on the structure. A minimum factor of safety of 2.0 is normally adopted providing the resultant reaction lies within the middle third. If there is ‘tension’ in the bearing pressure at the heel, then a higher factor of safety may be used as a further precaution against failure.

Overall instability

A slip circle analysis is essential for a bankseat form of construction and may be necessary for other types of abutment when the soil strata well below the structure is weaker than the soil layers at foundation level. Where soil strengths are based on tests, then a minimum factor of safety would be 1.5. Particular care is needed during construction if an intermediate pier foundation is being excavated at the toe of a cutting slope, when there is a bankseat positioned at the top.

Structural Design of Abutments

The structural design of abutments involves the selection of the proper thickness of the wall (stem) and base, and selection of the proper size and spacing of reinforcements to prevent ultimate and serviceability limit state failure.

Base Design

A base slab’s toe is made to withstand the highest ground forces pressing on the base, while some relief can be obtained from the toe’s self-weight and any added fill. The heel must be made to withstand upward ground pressure as well, however in this instance, fill, live load surcharge, and self-weight can generate high loading circumstances that can reverse the shears and moments that follow. The foundation slab may be supported by piles, in which case the predicted loads in each pile would take the place of the bearing pressures.

Wall Design

The stem of an abutment wall is designed to withstand the shears and bending moments caused by horizontal forces, as well as the bending imposed by the deck in the case of integral bridges. Since direct stresses from vertical loads are typically relatively minimal, they can be disregarded when designing walls. At the root of torsion blocks on horizontally cantilevered wing walls, significant in-plane strains can develop.

abutment wall construction — Construction of abutment walls

While for integral bridges the top of the wall and connection with the deck can also be crucial, the key section for moments and shear forces occurs at the root of the wall in the case of simple vertical cantilever walls. Due to traction and braking effects, concentrated horizontal stresses may exist at bearing level as well as at the back of the curtain wall. Calculating the bending moments in the wall typically involves distributing these loads vertically.

How to Design Pile Foundation using SPT and CPT Data

By

Ubani Obinna

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August 21, 2022

Table of Contents

A lot of correlations have been proposed by researchers for relating the results from Standard Penetration Test (SPT) and Cone Penetration Test (CPT) to pile load capacity. In this article, we are going to review how to design pile foundation using SPT and CPT data. It very important that these correlations used in this article and other similar ones should be used with caution, since they are statistical relationships that may not have taken all parameters into consideration. Experience is required to successful apply them to practice.

It is common to use the results of in-situ tests to determine the ultimate bearing capacity of piles in coarse soils because it is difficult to get undisturbed samples of these soils from boreholes. Two fundamental approaches will be discussed in this article, which are SPT and CPT.

118D99DE 1A45 43A7 A34A E233C66CF3BC — **Figure 1:** CPT process in the field

In all cases, the allowable load is calculated by dividing the ultimate bearing capacity by a safety factor which usually varies between 2 and 3. The diameter of the pile and the soil’s compressibility both have an impact on the allowable settlement at the working load. Experience has shown that a safety factor of 2.5 will guarantee that an isolated pile driven into coarse soil with a shaft diameter of no more than 600 mm won’t settle by more than 15 mm.

As an alternative, partial safety factors can be used in accordance with the steps outlined in EC7, and the serviceability limit state can be verified through calculation or experience, and if necessary, by performing loading tests. If the “global” safety factor of 2.5 is utilised, these tests are still required unless experience offers a more reliable indicator of settlement behaviour.

Standard Penetration Test (SPT)

For driven piles in coarse-grained soils (sands), Meyerhof (1956) proposed the following relationship between the skin friction (f_s) of the pile and the standard penetration number;

Skin friction

The skin friction in piles can be obtained from SPT data using the relationships in Equations (1-3);

Q_f = f_s × perimeter × length ——– (1)
For Displacement piles: f_s = 1.9N₆₀; f_s ≤ 100 kPa ——– (2)
For Non-displacement piles: f_s = 0.95N₆₀; f_s ≤ 50 kPa ——– (3)

End bearing

The end bearing in piles can be obtained from SPT data using the relationships in Equations (4-6);
Q_b= f_bA_b ——– (4)
f_b = CN₆₀ (kPa) ——– (5)
C = 38(L_s/D) ≤ 380 ——– (6)

where L_s is the length of pile driven in sand, and D is the diameter of the pile.

Design of Pile Foundations using SPT and CPT Data — **Figure 2**: Relationship between standard penetration number and angle of shearing resistance

Alternatively, the standard penetration number can be used to estimate the angle of internal friction of the soil (using Figure 2), which is then used to compute the allowable load on the pile.

Cone Penetration Test (CPT)

The cone penetrometer was originally developed to estimate the end bearing capacity of piles. The cone resistance, q_c, is a measure of the end bearing capacity and the sleeve resistance, q_s1, is a measure of the skin or shaft friction.

44119BCC 12F8 4BA5 A7FA BF9196161061 — **Figure 3:** Cone penetration test

End Bearing

The ultimate end bearing capacity of a single pile (Xu and Lehane, 2005) is estimated from;

Q_b = C_bq_c-avA_b ——– (7)

where C_b= 0.6 for closed-ended driven pipe piles in sand and C_b = 0.9 for jacked piles in sand; q_c–av is the average cone tip resistance over a distance 1.5 times the pile diameter above and the same distance below the pile base, and A_b is the area of the pile base. For open-ended pipe piles in siliceous sand (Lehane and Randolph, 2002);

C_b = 0.15[1 + 3(D*/D)²] ——– (8)

and D is the external diameter, D* is the effective diameter.

The maximum expected settlement from the equation above is about 10% of the pile diameter.

Several other empirical equations are used in practice. For example, Fleming and Thorburn (1983) suggested that for Equation (7), C_b = 1 and q_c–avis the average cone value over an influence zone of 8 pile diameters above the pile base and 2 pile diameters below the pile base, calculated as follows:

q_c-av = (q_c1 + q_c2 + 2q_c3)/4 ——– (9)

where q_c1 is the arithmetic average of cone resistance values over 2 pile diameters below the pile base, q_c2 is the minimum cone resistance value over 2 pile diameters below the pile base, and q_c3is the arithmetic average of minimum cone resistance values below q_c2 over 8 pile diameters above the pile base.

Shaft Resistance

The shaft or skin friction is calculated where the sleeve resistance is determined using either an arithmetic or geometric mean value of cone resistance over the buried pile depth. Cone sleeve friction is influenced by soil compressibility and relative density, whereas skin friction on a pile is influenced by pile geometry, roughness, relative density, installation technique, soil compressibility, and pile geometry.

In the case of fine-grained soils, soil consolidation around the sleeve has a significant impact on the cone sleeve friction value. As a result, utilising the results from the cone penetrometer, significant discrepancies between the short-term and long-term load capacity can be anticipated.

As evidenced by the results of the pile tests, the cone penetrometer estimates for the short-term load capacity can be as low as 20% of the long-term load capacity. One of many equations suggested in the literature can be used to predict the cone sleeve resistance if it is not measured. Some of these are as follows:

For both open-ended and closed-ended driven pipe piles, the skin frictional stress (Lehane et al., 2005) is given as;

f_s = C_sq_cA_rs × [max(2, h/D)]^-0.5 × tan (δ_cv) ——– (10)

where C_s is a constant (0.03 for compression piles and 0.0225 for tension piles), h is the distance of the pile section under consideration above the pile base, δ_cv is the soil–pile interface friction angle correlated to the mean particle size (≈23° for D₅₀ = 1 mm, increasing to 28.88 for D₅₀= 2 mm for sand on steel) and A_rs is the effective area ratio of the pile shaft given as;

Closed-ended pipe piles: A_rs = 1
Open ended piles: A_rs= 1 – min{1, (D_i/1.5)^0.2} (D_i/D)²

Other relationships that have been developed by several other authors are as follows;

Eslami and Fellenius (1997):
fs = C_sq_cs ——– (11)

where q_cs is the cone resistance after adjustments for porewater pressure measured at the cone shoulder, and C_s is a coefficient that depends on soil type, as shown in the Table below;

Type of soil	C_s
Soft, sensitive soils	0.08
Clay	0.05
Stiff clay and mixture of clay and silt	0.025
Mixture of silt and sand	0.01
Sand	0.004

Vesic (1977) —coarse-grained soils:
f_s = 0.11 exp(-3 tan ϕ’_cs)q_c ——– (12)

Jardine et al. (1998) —coarse-grained soils:
f_s = σ’_rctan ϕ_i ——– (13)

where σ’_rcis the radial effective stress on the shaft and is empirically related to the cone resistance.

Tumay and Fakhroo (1984) —stiff clays:
f_s = 0.5q_c ——– (14)

The ultimate skin friction is;
Q_f = f_s × perimeter × length

You have to exercise caution in using these empirical equations, as they were derived from pile load tests and cone penetrometer data in particular soil types and locations.

Worked Example: Pile Load Capacity Using SPT Data

A 350 x 350 mm closed-ended square pile is driven into a sand profile to a depth of 10 m. The SPT results are shown in the table below. Estimate the allowable load capacity for a factor of safety of 3.

Depth (m)	1	3	5	6	8	10	11	13
N₆₀ (blows/300 mm)	22	18	25	20	30	36	39	45

Use Meyerhof (1956) equations for displacement piles since the pipe pile is closed-ended.

Solution

Step 1: Determine the skin friction
N₆₀ = average N₆₀ = (22 + 18 + 25 + 20 + 30)/5 = 23
Displacement pile: f_s = 1.9N₆₀ = 1.9 × 23 = 44 kPa < 100 kPa; use fs = 44 kPa
q_f= f_s × perimeter × length = 44 × 4(0.35) × 10 = 616 kN

Step 2: Determine the end bearing and allowable load capacity.
f_b = CN₆₀ (kPa);
C = 38(L_s/D) = 38(10/0.35) = 1085 > 380
Use C = 380
N₆₀ = 36 (this is the N value at the base)
Q_b = f_bA_b= 380 × 36 × 0.35 × 0.35 = 1675.8 kN

Q_ult = Q_f+ Q_b = 616 + 1675.8 = 2291.8 kN
Q_a = Q_ult/FS = 2291.8/3 = 763.93kN

Lateral-Torsional Buckling Strength of Corrugated Web Girders

By

Charity Mmuo

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August 20, 2022

Table of Contents

Corrugated web girders are being utilized more commonly in structural engineering due to their many beneficial qualities, such as lower dead load, increased shear buckling strength, and a special structural behavior known as “accordion effect.” Because of this, designing girders is simpler than it would be for those with flat web.

For the design of these girders, the stability of the local and global members is essential. The effect of local stability on cross-sectional resistances, specifically the bending moment resistance, shear buckling resistance, and resistance against transverse force, has been well studied in the past. These resistances can now be calculated using precise design models.

According to Jager et al (2022), several research works have shown that girders with corrugated web have a higher elastic critical moment than those with flat web. This implies that the lateral-torsional buckling strength of beams increases when the web is corrugated.

However, the increment has been viewed in several ways by being associated with the cross-sectional characteristics. The majority of studies concur that the web contribution should not be taken into account when calculating inertias about the strong and weak axis, although there are discrepancies when taking the enhanced elastic critical moment into account.

corrugated web girder — **Figure 1:** Corrugated web girders in application

The elastic critical moment of a conventional flat web girder subjected to uniform bending moment can be calculated by Eq. (1) according to EN1993-1-1 as;

M_cr = π/kL √EI_z [(π/k_wL)² EI_w + GI_t] ——– (1)

Where L is the span of the girder, E is the elastic modulus, G is the shear modulus, I_z is the moment of inertia about the minor axis, I_tis the torsional constant, I_w is the warping constant, k is the effective length factor about the weak axis rotation and k_wis the effective length factor with respect to warping.

To account for the critical elastic moment of corrugated web girders, previous research work by Lindner (1990) suggested an additional term with a correction factor c_w in the warping constant given by Eqs. 2–4 for double symmetric sections in order to consider the greater performance (c₁ = 8, c₂ = 25); where I_w,flatis the warping constant of flat web girder.

I_w = I_w,flat+ c_w (L²/Eπ²) ——– (2)

c_w = a₃² (h_w + t_f)²/[c₁⋅u_x⋅(a₁ + a₄) ——– (3)

u_x = (h_w + t_f)/2⋅G⋅a₁⋅t_w + (h_w+ t_f)²⋅(a₁ + a₄)³/c₂⋅a₁²⋅E⋅b_f ⋅t_f³ ——– (4)

Larsson and Persson (2013) performed a notable FE study and found that the proposal of Lindner (1990) gives the best approximation to the numerical results. However, this additional term depends on the girder length which is not possible for a sectional constant. Therefore, they substituted the Lindner’s proposal into Eq. (1) in order to rearrange the additional term to the torsional constant according to Eq. (5).

I_t= I_t,flat + c_w/G ——– (5)

The appropriateness of Eq. (5) was confirmed by Lopes et al. (2017) and they proposed a slightly modified correction factor cw for trapezoidally and sinusoidally corrugated web girders (c₁ = 22, c₂ = 300 in Eqs. 3–4) based on FE analysis.

According to EN1993-1-1 [2] the reduction factor (χ_LT) for the lateral-torsional buckling strength for rolled sections or equivalent welded sections with flat web may be calculated by Eqs. 6–8, where α_LT is the imperfection factor (for different buckling curves: a – 0.21, b – 0.34, c – 0.49, d – 0.76), λ_LT is the relative slenderness, β is the multiplication factor and λ_LT,0 is the relative slenderness limit. The standard suggests to use buckling curve d if the depth-to-width ratio of the section is greater than 2; otherwise the buckling curve c shall be used. In Eq. (6) M_y is the cross-sectional bending moment resistance considering the local flange buckling.

λ_LT = √[M_y/M_cr] ——– (6)

λ_LT,0 = 0.4, and β = 0.75
ϕ_LT = 0.5[1+ α_LT (λ_LT – λ_LT,0) + βλ_LT²] ——– (7)

χ_LT = 1/[ϕ_LT + √(ϕ_LT² – βλ_LT²)] but χ ≤ 1.0 ——– (8)

In the past, nonlinear FE (finite element) analysis was mostly used to explore the ultimate lateral-torsional buckling strength; hence, there aren’t many test data available. Only a small number of experimental tests and some non-linear FE analysis performed by various researchers have been used to study the ultimate lateral-torsional buckling strength of corrugated web girders.

Researchers (Jager et al, 2022) from the Department of Structural Engineering at Budapest University of Technology and Economics, Hungary, recently worked on an experimental study using 11 large-scale test specimens to examine the lateral torsional buckling resistance of corrugated web girders.

The study was published in the journal, Structures (Elsevier). In the research work, the prior design recommendations for trapezoidal corrugated web girders’ lateral-torsional buckling strength were compared and evaluated based on the test results, and a preliminary design buckling curve was provided.

Test Setup and Experimental Study

11 large-scale test specimens were examined under four-point-bending condition as part of the research program by Jager et al (2022). Investigations were conducted on six distinct girder geometries with the same trapezoidal corrugation profile but varied flange sizes. Five more specimens were employed in test replications to check the accuracy of the findings. The widths of the parallel and inclined web folds are the same for all specimens (a₁ = a₂ = 98 mm), as they are often employed in bridges with a corrugation angle of 45°.

notation 1 — **Figure 2**: Notations used on the study (Jager et al, 2022)

Notation 2 — **Figure 3:** Cross-sectional layout of specimen’s geometry (Jager et al, 2022)

The six specimen types with varying flange sizes were depicted in scaled drawings as shown above, and the girder numbers were assigned in accordance with the rising flange diameters. The nominal flange thickness (t_f) of specimen types #1, #2, and #3 is 14 mm, whereas specimen types #4, #5, and #6 have a 16 mm flange thickness. There are five different nominal flange widths (b_f): 140, 160, 180, 220, 250, and 300 mm.

With a nominal web depth (h_w) of 520 mm, all specimens have a web thickness (t_w) of 6 mm. Vertical stiffeners with a nominal thickness (t_s) of 10 to 16 mm were installed at the load introduction and support positions. To analyze the lateral-torsional buckling behavior of corrugated web girders, a wide range of varied LTB slenderness ratios were covered by the specimens.

S355 steel grade was used for the flanges, while S235 steel grade was used for the web, with nominal yield strengths of 355 MPa and 235 MPa, respectively. Coupon tests in accordance with ISO6892-1 were used to assess the material’s characteristics.

test set up — **Figure 4:** Test setup and location of measuring devices (Jager et al, 2022)

Prior to testing, the rigid frame measuring system’s spatial coordinates were calculated using geodetic triangulation from three places using a total station theodolite to achieve high precision. Each specimen was subjected to a static load of around 0.1 kN/s]. To assess the elastic response of the structure and the stiffness of the investigated girders in the elastic domain, loops were run five times by equal loading steps during the loading and unloading process until achieving about 60% of the projected ultimate load bearing capacity.

The hydraulic jacks that were being used were coordinated and attached to the same hydraulic system. In order to assess the results and compare them using FE models, the test specimens’ observed load-displacement curves, buckling forms, and ultimate failure modes were documented on photos. Displacement control was used during the experiments, and the post-buckling ranges were also investigated. Following the testing, portions of steel plate were removed from the intact flange and web parts of each specimen for material testing.

Overall Results and Failure Modes

According to the findings of the study, girders with compact flanges experienced pure lateral-torsional buckling failure, whereas girders with slender flanges experienced flange buckling failure in conjunction with lateral-torsional buckling failure. Furthermore, because of the rigid cross-sectional movement failure modes, no distortional web failure developed.

Moreover, it was observed that after unloading, slender specimens can regain proportionately bigger displacements owing to elastic buckling, while heavier flanged samples can experience inelastic buckling. The study also showed that initial imperfections have significant resistance reduction effect on the lateral-torsional buckling strength.

Comparison with Design Proposals

The authors compared the lateral-torsional buckling curves of EN 1993-1-1 (β =0.75) to various reduction factors estimated from the test findings (χ_LT,test). By the calculation of the relative slenderness three different proposals for the elastic critical moment are used:

(i) without any modification in the torsional or warping constants,
(ii) with an additional term in the torsional constant proposed by Larsson and Persson (2013) using the correction factor of Lindner (1990) presented by Eq. (5), and
(iii) with the slightly modified correction factor of Lindner according to Lopes et al. (2017).

The web’s contribution to the moment of inertias surrounding the strong and weak axes was ignored in every case. Both effective length factors (k_w and k) were arbitrarily set to 0.5 in order to optimize the elastic critical moment and offer a safe side solution for the buckling curve because the test specimens were almost completely constrained against warping and rotation about the weak axis. It should be noted that all test results were above the previous proposal’s buckling curve band.

Comparison of test results — **Figure 5:** Comparison of the test results and previous proposals to EN1993-1-1 buckling curves (Jager et al, 2022)

The calculation methods considered in the study are in harmony with the “accordion effect” of corrugated web girders. It is to be noted all test results are above the buckling curve b and the proposal of Lopes et al. (2017) gives the larger elastic critical moment and smaller relative slenderness then Larsson and Persson’s (2013) solution.

The calculated resistances are determined according to the equation below;

M_b,R = χ_LT⋅M_y ——– (9)

Where the reduction factors (χ_LT) are obtained from Eqs. 6–8 using the above mentioned three different proposals for the elastic critical moment. It can be seen that the best fit and safe side solution is provided by the proposal of Lopes et al. (2017) using buckling curve b.

By comparing the test and computed lateral-torsional buckling resistances, The authors pointed out that the Lopes et al (2017) proposal’s suggestion offered the greatest match and safest approach to take. However, it should be highlighted that a more extensive statistical evaluation should be carried out in accordance with EN 1990 AnnexD, expanded by advanced FE analysis, in order to determine an applicable buckling curve that satisfies the safety criterion of the Eurocode.

Article Credit:
B. Jager, L. Dunai, B. Kovesdi (2022): Lateral-torsional buckling strength of corrugated web girders – Experimental study. Structures 43 (2022) 1275–1290 https://doi.org/10.1016/j.istruc.2022.07.053

The contents of the cited original article published by Cement and Concrete Research (Elsevier) is open access, under the CC BY license (http://creativecommons.org/licenses/by/4.0/) which allows you to share and adapt (remix) the article provided the appropriate credit is given, and the link to this license provided.

References

[1] Larsson M, Persson J, Lateral-torsional buckling of steel girders with trapezoidally corrugated webs, MSc thesis, Gothenburg, Sweden, 57, 2013.
[2] Lindner J. Lateral torsional buckling of beams with trapezoidally corrugated webs, Proceedings of the 4th International Colloquium on Stability of Steel Structures, Budapest, Hungary, 79-82, 1990.
[3] Lopes GC, Couto C, Real PV, Lopes N. Elastic critical moment of beams with sinusoidally corrugated webs. J Constr Steel Res 2017;129:185–94.

The T-shaped Civil Engineer

By

Anuoluwapo Kolade

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August 20, 2022

T-SHAPED ENGINEER — T-shaped people are those with a depth of knowledge in at least one discipline and a breadth of knowledge about innovation and entrepreneurship that allows them to work effectively with professionals on other disciplines to bring their ideas to life — Tina Seelig

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Perhaps you think you have stumbled on an article you would read in a few minutes; then, I suppose you might be wrong. I expect you to devote about 10 – 15 minutes to reading and assimilating this treasure coated in words. It will be worth your time.

Shall we? Yeah.

Recently, a friend and I had a conversation where he blatantly told me that Nigeria has not been producing good graduates in recent years. Well, you can make a case for or against his opinion, but I think I agree with him. I agree with him because the decay of our education system is obvious, and you may be disappointed after engaging students or graduates in a conversation.

This malady then begs the question: Who is to be blamed? Me? You? The government or schools? Well, let me tell you about my blame-sharing formula quickly. When I was in school, I ditched 10% of the blames to the government and another 10% to my school while I took 80%.

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Do you know why I took the largest share of the blames? I did so because I believed that whatever I become or becomes of me, even within the unfavourable education system, totally depends on my choices. It depends on what I make out of my life. It also depends on how I use my time during breaks and strikes. Moreover, I do not have control over the system, government, or school. Therefore, I had no choice but to become a T-shaped student who is well prepared to take the industry by storm.

images — *Building a set of T-shaped knowledge and skills is one of the most valuable things you can do for your future career and personal development*

The T-shaped Concept

Generally, the T-shaped concept is a metaphor for an individual’s depth and breadth in their skills. The vertical bar on the “T” represents the depth of related skills and expertise in a single field. In contrast, the horizontal bar represents a breadth of skills and the ability to collaborate across disciplines.

For students, the T-shaped concept not only means possessing deep, technical skills but also having broader characteristics or qualities. I want to emphasize that developing into a T-shaped civil engineering student gives you a competitive edge and could make all the difference in attaining a stable and successful career after graduation.

In this article, I will share my model for a T-shaped student. This model is not theoretical, but one that worked for me during my undergraduate days and shaped me into who I am today.

The T-shaped Model

The T-shaped model is a tripartite model with the vertical bar on the “T” representing professional development and the horizontal bar representing academic excellence and skill acquisition/entrepreneurship/leadership on either side.

Fitting into this model and finding appropriate balance is essential for staying competitive and becoming well-rounded professionals after graduation. Below, I will tell you everything you need to know about the T-shaped model and why you must become a T-shaped student.

Academic Excellence

Perhaps you believe that school is a scam; it is not. Nigeria’s tertiary education system presently finds itself in several challenges, including underfunding of institutions, infrastructural decay and neglect, academic corruption and fraud, wastage of resources, and distasteful conditions of learning and service. I was in the system for quite a long time, maybe 6-7 years, and I know what it means to be a student in such a system. However, school is and has never been a scam, and neither is academic excellence.

By the grace of God, I was privileged to graduate with a good grade, and I can tell you that a good grade grants you access to opportunities, boosts your confidence, and helps you earn respect. Indeed, grades are not a measure of a person, nor are they even the sole measure of academic accomplishment.

However, people care about grades. Many employers also care because your grades tell them if you can successfully handle tasks and produce results with less training or close supervision. I have two stories or experiences to share about academic excellence below.

First, I got a full-time job after the first month of leaving school. How did it happen? A Quantity Surveyor who contributed to training me during my 6 months of industrial training and knew about my professional capacity recommended me for the job. Honestly, I did not meet up with the job requirements. However, the company’s principal decided to take a chance on me because of my academic grade and punctuality.

I also passed the written test on structural design and showed a willingness to learn while working for the company. It was more like I was employed straight away, though on 3 months probation. Eventually, I left the job for NYSC camp before the end of the 3 months, and I could not return to the job even though the company was willing to allow me to keep my job.

Second, twelve other corps members and I were posted to an establishment for our NYSC year, and an interview was scheduled. One of the interviewers asked for my grade, which was the end of the interview. At the end of the interview, 7 of us were admitted, and I was the first person on the list. The academic grade was the difference.

images 1 1 — Education’s purpose is to replace an empty mind with an open one **— Malcolm Forbes**

Dear reader, you cannot afford to dismiss grades as unimportant, even if you have reservations about them, as many of us do. Furthermore, without good grades, though, gaining entrance into a master’s or PhD program or winning scholarships becomes much more difficult. Therefore, I can tell you sincerely that having a good grade will result in a life with many advantages.

Skill Acquisition/Entrepreneurship/Leadership

I have often heard people ask: what if I did not or cannot graduate with a good grade? It is no big deal because it does not mean you are doomed. You can always leverage acquired technical, soft, business and leadership skills. Being successful in school is more than just earning good grades. Going to school allows you to focus on developing new skills, making new friends, and crafting a path for your future.

Talking about leadership, taking up leadership position(s) as a student is a no-brainer because becoming a student leader will help you to develop innumerable leadership skills, including conflict resolution, cultural intelligence and professional advancement. Furthermore, becoming a student leader allows you to find how to build a team, how you work with others, and where your areas of improvement might be. Lastly, you will gain valuable soft and management skills, build networks and connections, and improve your CV or resume.

Having mentioned the importance of leadership skills and experience, I would like to remind you that losing focus on your academic and career goals is easier if you get financially stressed or frustrated. Do I also need to remind you that having access to money when you need it is enough to keep you going? Therefore, you must learn a skill or venture into a business that can earn money because trying to find your feet in the AEC industry can be very frustrating and overwhelming.

Furthermore, at the start of your journey, you may have to do many unpaid jobs or internships to get the experience and exposure you desire. Therefore, having a skill or business that pulls in money for you can make your journey much easier.

For example, I am not the type that is good at business (at least for now) and am not a fan of tech, but I ensured I left school with three skills — research, writing and teaching skills. These skills are what I have been leveraging from my penultimate year in school until now. Through these skills, I have been able to keep financial stress and frustration away with the earnings.

Screenshot 20220608 191454 — **Photo credit:** LinkedIn

Furthermore, I stumbled on a post on LinkedIn some time ago highlighting that it is wrong not to have a plan B, at least. In this post, the person in question lamented how he has been searching for jobs for almost two years without getting one. He even went as far as tearing his certificate. He must have been frustrated. Thankfully, he was able to get multiple job connections through the post.

Dear reader, with income-generating, leadership, and strong entrepreneurial skills and experience, not only will you have the option of bringing your projects and dreams into existence, but you are also likely to find yourself in high demand among potential employers who look for a mix of entrepreneurial and other professional skills.

Professional Development

How can you do better in a field when you don’t know better? It is vital to be true to yourself at this point. Do I have to remind you that great engineers are not born? Truly, they are made through professional and personal development. Therefore, evolving into a better you is the only way you will reach new heights professionally and personally.

Perhaps you had the chance to ask me what topped my priority list when I was in school; then, I would tell you it was striving to be a well-versed professional with high esthetic standards and a passion for excellence. I ensured I made use of every opportunity to advance the achievement of this goal because I never joked about it.

I did talk earlier about the importance of getting a good grade. However, your certificate holds little water compared to your experience. Therefore, you should never joke about internships and strike periods as undergraduates because they are great opportunities to gather relevant work experiences.

Similarly, if you are a recent graduate, I would advise that you pursue internship and apprenticeship opportunities that would give you relevant work experiences, even with little or no pay, rather than going for jobs with little pay and offering no relevant work experiences.

For example, from my experiences, I have discovered that anything that is meant to teach and instruct you will require patience, flexibility, growth and stretching. Engineering is not that easy because most of the concepts we adopt are partially abstract and partially concrete. Therefore, most of these things require time and careful direction or mentoring to understand and apply, and this is where your patience and willingness to learn through internships come into play.

Personally, internship experiences have changed and shaped my life and charted me on a path of career realization and helping people. It seemed like I was just being used or doing too much at first, but the benefits have been enormous. Funnily, I did not even get a penny until after my first 6 months at the organization, but I ensured that I put into practice all the skills I acquired, which has built my expertise. As a result, even while I was still an undergraduate, I was able to build my expertise and empower others alongside.

Dear reader, professional and personal development do not happen overnight but over time, that is, slowly and deliberately. It is a daily practice and lifestyle — a lifelong process!

Final Words

20160704114059 shutterstock 390101425 — *You cannot afford to be ordinary; come up higher*

As I round up this article, I want to tell you that self-confidence is the most beautiful thing you can possess in this field or career path. Unfortunately, many a time, young and aspiring engineers struggle with confidence issues because of their inability to measure up with the T-shaped model. Consequently, this results in fears that affect one’s performance and productivity.

However, the great news is: that you can do many things to measure up and help boost your confidence levels. Now is the best time to work on your confidence level and the lagging areas. Sooner, you will see yourself reach the heights you desire, after hours, days, weeks, and years of constant work and dedication.

To the young and aspiring just starting in their careers, please, invest in yourself, your learning and your career; and be unapologetic about it. The process is not always fun or easy but always beneficial. You can restart from the basics and no matter how challenging it gets, never forget to show up every day. Don’t give up. It will be worth it.

Lastly, my dear colleague, dream big, develop yourself, unleash your potential, collaborate with others, play to your strengths, work on your weaknesses, enjoy the process, share your unique gifts with the world, and grow your greatness by testing yourself, expanding yourself, learning and improving.

What is the Failure Load of Pile Foundation?

By

Ubani Obinna

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August 20, 2022

In our previous articles, we defined the failure load as the load that ultimately causes a pile to fail, or the load at which the soil’s bearing capacity is fully mobilised. However, in an engineering sense, failure might have occurred long before the structure was subjected to the maximum load because of the structure’s excessive settlement.

Engineers generally agree with Terzaghi’s assertion that, for practical purposes, the ultimate load can be defined as the load that results in a settlement of one-tenth of the pile’s diameter or width. However, the settlement at the working load may be excessive if this criterion is applied to piles with a large diameter and a nominal safety factor of 2.

The allowable load is almost always determined purely by tolerable settlement at the working load in situations when piles are serving as structural foundations. For every given type and size of pile in any soil or rock conditions, the engineer should be able to predict the load—settlement relationship up to the point of failure when calculating the allowable loads on piles.

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Failure Load and Settlement

In most cases a simple procedure is to calculate the ultimate bearing capacity of the isolated pile and to divide this value by a safety factor which experience has shown will limit the settlement at the working load to a value which is tolerable to the structural designer. But where settlements are critical it is necessary to evaluate separately the proportions of the applied load carried in shalt friction and end-bearing and then to calculate the settlement of the pile head from the interaction of the elastic compression of the pile shaft with the elasto-plastic deformation of the soil around the shaft and the compression of the soil beneath the pile base.

In all cases where piles are supported wholly by soil and are arranged in groups, the steps in calculating allowable pile loads are as follows;

Determine the base level of the piles which is required to avoid excessive settlement of the pile group. The practicability of attaining this level with the available methods of installing the piles must be kept in mind.
Calculate the required diameter or width of the piles such that settlement of the individual pile at the predetermined working load will not result in excessive settlement of the pile group.
Examine the economics of varying the numbers and diameters of the piles in the group to support the total load on the group.

The overall goal should be to adopt the highest working load on each individual pile while keeping the number of piles in each group as small as possible. As a result, pile caps will be smaller and less expensive, and the group settlement will be at a minimum. However, excessive settlement that causes intolerable differential settlements between neighbouring piles or pile groups may occur if the safety factor on the individual pile is too low.

The diameter and length of the piles in the case of isolated piles or piles arranged in very small groups will be determined only by taking into account the settlement of the isolated pile at the working load. Installation methods significantly impact the carrying capacity of piles. The interaction between the pile and the soil is influenced by a number of variables, including whether a pile is driven or cast in situ in a bored hole, whether it is straight-sided or tapered, and whether it is made of steel, concrete, or timber.

Engineers shouldn’t have very high expectations for formulas used to determine the carrying capacity of piles and shouldn’t be upset if the calculations show failure loads that are off by plus or minus 60% of the failure load determined by test loading. It should be kept in mind that a full-scale foundation is being evaluated when a pile is subjected to test loading.

It is not surprising that there could be relatively large differences in failure loads on any given site given the typical variability in ground conditions and the influence of installation techniques on ultimate resistance. If full-scale pad or strip foundations were loaded to failure, engineers would not be surprised to see such huge differences.

The alternative is to calculate allowable loads or design bearing capacities by dynamic formulae. These will give even wider variations than soil mechanics’ methods and, in any case, these dynamic formulae are largely discredited by experienced foundation engineers, unless they are used in conjunction with dynamic testing and analysis using standard equipment.

Digital Fabrication with Concrete and Sustainable Designs

By

Charity Mmuo

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February 27, 2024

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Over the past few years, the subject of digital fabrication with concrete has advanced significantly, with numerous alternative techniques having been created and numerous large-scale models having been built.

A recent assessment indicates that 3D concrete printing, the most extensively researched and commercially available of these technologies, is at a technology readiness level (TRL) of 6-7, comparable to polymer fused deposition modelling technology, putting it on the verge of becoming widely used. However, according to Flatt and Wangler (2022), the viability of such processes is still under discussion, which frequently results in divisive and pointless conversations.

Pioneers created these procedures with the goal of resolving productivity challenges in the building industry. However, in recent years, the need to expand architects’ creative areas and make it more cost-effective to build increasingly complex buildings made feasible by computer-aided design has been a major driving force behind digital fabrication in construction. With this capability, digital fabrication is being pushed more and more as a way to increase efficiency while also lowering the environmental impact of the building industry.

The fact that digitally fabricated structures would only use material where it was necessary, allowing for significant material savings, is a major defense of this claim. This reasoning encounters difficulties with concrete, too, because digitally created concrete frequently has a substantially larger environmental footprint per unit volume than regular concrete.

Additionally, the printing process itself may have some additional negative effects on the environment due to the manufacturing of the printing cell or the energy used to run it. It has been demonstrated that printing factors like printhead velocity and resolution have a significant impact on these process-related effects.

These well-known and currently investigated issues, as well as the fact that previous digital fabrication demonstrations have frequently focused more on “production prowess” than material savings, might result in dry and fruitless discussions of the technology’s sustainability.

Flatt and Wangler (2022) of the Institute for Building Materials, ETH Zurich, Zurich, Switzerland, recently published a paper in the journal, Cement and Concrete Research to highlight the real opportunities and challenges with regard to sustainability in digital fabrication with concrete, hopefully sparking fruitful discussions on the topic as the technology becomes more widely used.

Their article outlined a straightforward equation that incorporates the primary issues with regard to a structure’s environmental footprint. Three things come into play: Shape efficiency (or material utilized), Material footprint, and Service life (durability). The material itself was then further discussed in the context of concrete extrusion (3D printing), the technology that is most commonly used in many industries like cars, home improvement, computers, and PCB manufacturing. It was emphasized that in many cases, printed concrete is overdesigned and that well-chosen accelerators can address that issue quite effectively.

Main Factors of Concern

According to Flatt and Wangler (2022), three factors primarily determine a concrete structure’s environmental impact:

The total amount of materials used
The material’s embodied carbon dioxide, and
The durability.

According to the equation, it would be the product of the volume of material used and the environmental impact of that substance per unit volume, divided by the service life. It was highlighted that this first-order estimation does not account for variations in this and is only appropriate for comparing structures with equivalent load-bearing capacity. Additionally, the environmental effects of related changes in concrete production, such as formwork use or the aforementioned operating energy, are not expressly taken into account in this relationship.

2 6 — Schematic illustration of the main factors affecting the environmental impact of a structural element with a given load-bearing capacity per year of service life (Flatt and Wangler, 2022).

The primary benefit of digital fabrication is that it can require less material. Although the effect on durability is still being researched, this typically accompanies an increase in the material’s environmental imprint and a potential reduction in service life.

Such conflicts mean that the results of environmental balancing will typically not be trivial and will undoubtedly depend on the circumstances. According to Flatt and Wangler (2022), this necessitates a more thorough consideration of the issue, taking into account the true potential for material savings while keeping in mind the limitations of material composition and durability.

Shape Efficiency

The cost of constructing buildings that use less material to provide a certain load-bearing capacity is one of the key defenses for digital fabrication. Thus, it could facilitate structural design methods that are very successful but are all too frequently overlooked. In this context, it is worthwhile to reflect on Pier Luigi Nervi’s ribbed floor designs.

Although labour was inexpensive at the time these constructions were made, concrete was expensive. Today, the situation is the opposite, making it less expensive to create consistent floor slabs by utilizing a lot more concrete than is actually necessary. This is a particularly instructive example because it is simple to visualize the savings since floors are a significant consumer of concrete in structures.

2 7 — Ribbed floor system by Per Luigi Nervi at the Gatti Wool Factor

Material Footprint

Despite the potential for material savings, digital concrete frequently has a larger environmental impact than regular concrete, at least for the most popular kind of concrete extrusion. This can be reduced by using recycled components in place of new ones, looking into using different, lower-CO₂binders, or cutting the amount of paste in the cement (increasing aggregate content).

However, none of these approaches are specifically applicable to digital concrete and are instead investigated for concrete in general. The greater level of processing, particularly pumping, which normally increases paste volume, is principally responsible for digital concrete’s higher environmental impact.

In fact, given that mix designs often do not include coarse (>4 mm) aggregate, digital “concrete” is more appropriately referred to as digital “mortar” even though coarse aggregates are beginning to emerge in both academic and industrial settings. Whatever the case, the lower maximum aggregate size restricts the maximum packing fraction of aggregates, increasing the paste volume and, consequently, cement contents and carbon footprints that are larger.

Also keep in mind that while digital concretes have water-to-binder ratios that are more in line with infrastructure and high-performance concrete, their primary use has been in non-load bearing capacities, such as replacing concrete masonry or serving as a lost formwork for cast reinforced concrete. As a result, given their current utilization, digital concrete mixes frequently have twice as much cement content than is really required.

Durability

The durability of digital concrete is a crucial final aspect to take into account when talking about its sustainability. Concerningly, the prevalent technology of extrusion printing in this context can result in cold joints between the layers. However, in general, it depends on the material qualities, state of the extruded material and the previously deposited layer.

The formation of cold joints in 3D-printed concrete is still an active area of research. Therefore, it is influenced by the time interval (contour length and printing speed), and it also seems to be significantly influenced by the substrate’s surface drying throughout the layer time interval. If created, cold joints can weaken the bond between layers, but more significantly for durability, they open up channels for faster ingression of water and/or CO₂.

Investigations are now being done to determine how digital concrete reacts to freezing and thawing cycles, although early results show that it performs poorly when compared to conventional concretes. Digital concrete faces the added challenge of curing in full exposure, thus losing the formwork as a “skin,” which increases the likelihood of shrinkage cracking and opens channels for aggressive chemicals.

Only recently have shrinkage issues specific to digital concrete been studied, but future research on these issues will undoubtedly need to be expanded. Noting that lowering the paste content is an easy technique to reduce shrinkage, raising the maximum aggregate size is beneficial for lowering the material footprint as well as boosting durability.

Similar to regular concrete, durability describes a material’s performance under particular exposure conditions and for a particular use. In this situation, it is important to distinguish between structural applications—where reinforcement is required—and other situations because the majority of—but not all—concerns regarding the influx of aggressive species are caused by the presence of reinforcement.

Another noteworthy achievement is the Pantheon, a non-reinforced concrete building whose performance is dependent on sound structural planning. Digital manufacturing can therefore help with important design-related issues of durability.

Determining the appropriate applications for digital manufacturing technologies and whether steel reinforcing is necessary is then crucial. One choice is to largely abandon structural concrete in favour of competing with masonry, or to simply use printed concrete as a substitute for lost formwork.

This eliminates the challenging task of strengthening digital concrete. The problem of reinforcing, on the other hand, must be addressed if structural applications are the focus, and this involves a significant amount of continuing research. Despite this, significant advancements are still required before the majority of reinforcement schemes can be widely accepted and approved in practice.

Conclusions

The fundamental argument made in the study by Flatt and Wangler (2022) is that, in addition to being situation-specific, the subject of the environmental impact of digital fabrication is complex and has a difficult answer. Examples that highlight potential significant material savings were given, however they are now more expensive to build than bulkier pieces with simpler designs.

In fact, the potential for material efficiency is what sets digital concrete construction apart from conventional construction in terms of sustainability, a distinction that should not be lost on those who aim to promote the technology to this end.

This is especially important because adoption of the technology based solely on cost-related factors would entail accepting a higher carbon footprint in exchange for lower labour costs. The current incentives for the implementation of digital concrete processes appear to be primarily cost-driven, related to formwork and masonry labour. These technologies may still have other social advantages (or problems), but those are outside the focus of this research, which is just looking at the environmental aspect of these technologies.

For a variety of reasons, the footprint of digital concrete is bigger than that of traditional concrete. One has to do with using a stronger paste volume, and is a problem that can be solved by increasing the maximum aggregate size through material advancements, which may also have additional advantages like reduced shrinking and greater incentive to use local materials.

The use of high clinker cements and overly strong final designs appear to be another factor contributing to the high carbon footprint. Both are the outcome of ineffectively attempting to meet the needs for gaining strength. Instead, utilizing accelerators based on aluminum can increase strength when it is needed, preventing overdesign of final strength and allowing the use of carbon-lean cements. Therefore, such compounds could be taken into account for quick vertical building rates of thin (and more shape-efficient) structures as well as a way to perhaps reduce the environmental impact of digital concrete.

However, extra caution must be used while employing such compounds because doing so could result in the formation of cold joints, which could reduce durability. In fact, the effect of digital fabrication techniques on the durability of concrete must be further examined in light of durability’s relationship to environmental impact.

References:
Flatt R. J., Wangler T. (2022): On sustainability and digital fabrication with concrete. Cement and Concrete Research, Volume 158, 2022, 106837 https://doi.org/10.1016/j.cemconres.2022.106837

The contents of the cited original article published by Cement and Concrete Research (Elsevier) is open access, under the CC BY license (http://creativecommons.org/licenses/by/4.0/) which allows you to share and adapt (remix) the article provided the appropriate credit is given, and the link to this license provided.

Design of Rectangular Roadside Drains | Drainage Sewers and Channels

By

Anuoluwapo Kolade

-

August 6, 2022

Table of Contents

Roadside drains or channels are structures that are used for conveying storm water away from roads or streets. The complete design of rectangular roadside drains involves hydraulic design, geotechnical design, and structural design.

The hydraulic design involves the proper sizing of the drain to ensure that the design flood is properly discharged, while the geotechnical design involves the verification of the capacity of the supporting soil to carry the weight of the channel and the water. It also involves the verification of the soil-structure interaction since drains are buried structures. The structural design of drains involves the selection of the proper material, thickness, and reinforcement to withstand the pressures and forces exerted by the soil and water.

In previous articles, we extensively discussed how to determine the best hydraulic cross-section of roadside drains and the construction and cost comparison of rectangular and trapezoidal drains. In this article, you will discover everything you need to know about the geotechnical and structural design of rectangular roadside drains.

Similar to the design of retaining walls, roadside drains are also subjected to active and earth pressures. In the example treated below, active and passive earth pressures, surcharge loads and water pressures are considered.

Worked Example on the Design of Rectangular Roadside Drains

The rectangular drain shown below is backfilled with a typical cohesionless granular material, having a unit weight (γ) of 18 kN/m³, zero cohesion (C), and internal angle of friction (ϕ) of 30°. The allowable bearing pressure of the soil is 150 kN/m², the coefficient of friction (μ) is 0.5, the unit weight of reinforced concrete is 24 kN/m³, and surcharge loads of 15 and 5 kN/m² on both sides of the drain. The drain has been designed to cater to a flow of 400mm depth and the unit weight of water (γ_w) should be taken as 9.8 kN/m³.

Given the information above, design the drain wall and base reinforcements assuming f_cu = 20 N/mm², f_y = 460 N/mm², cover to reinforcement = 40 mm, diameter of reinforcements = 10 mm, and thickness of walls and base = 150 mm.

Geotechnical Design

Wall pressure calculations

K_a = (1 – sinϕ) / (1 + sinϕ)
K_a = (1 – sin30°) / (1 + sin30°) = 0.333

Wall 1

active 093917 — **Active pressure on drain wall**

Active pressure at the top of the drain wall = qK_a = 15 × 0.33 = 4.95 kN/m²
Active pressure at the base of the drain wall = qK_a + K_aγZ = 4.95 + (0.33 × 18 × 0.85) = 4.95 + 5.049 = 9.999 kN/m²

passive 093919 — **Passive pressure on drain wall**

Passive pressure at the top of the drain wall = 0
Passive pressure at the base of the wall = γ_wZ = (9.8 × 0.55) = 5.39 kN/m²
Net pressure at the base of the wall = 9.999 – 5.39 = 4.609 kN/m²

Wall 2
Active pressure at the top of the drain wall = qK_a = 5 × 0.33 = 1.65 kN/m²
Active pressure at the base of the drain wall = qK_a + K_aγZ = 1.65 + (0.33 × 18 × 0.85) = 1.65 + 5.049 = 6.699 kN/m²

Passive pressure at the top of the drain wall = 0
Passive pressure at the base of the wall = γ_wZ = (9.8 × 0.55) = 5.39 kN/m²
Net pressure at the base of the wall = 6.699 – 5.39 = 1.309 kN/m²

Total vertical load (N)
Walls (W_ws) = 2(0.15 x 0.7 x 24) = 5.04 kN/m
Base (W_b) = 1.1 x 0.15 x 24 = 3.96 kN/m
Water (W_w) = 0.4 x 0.8 x 9.8 = 3.136 kN/m
Total vertical load W_ws + W_b + W_w (N) = 5.04 + 3.96 + 3.136 = 12.136 kN/m

Horizontal forces on drain walls due to surcharge load and backfill

combination 093919 — ***Resultant pressure on drain wall***

Wall 1 = qK_aZ + (0.5 × K_aγZ × Z) – (0.5 × γ_wZ × Z) = (15 × 0.333 × 0.85) + (0.5 × 5.049 × 0.85) – (0.5 × 5.39 × 0.85) = 4.246 + 2.146 + 2.291 = 4.101 kN/m

Wall 2 = qK_aZ + (0.5 × K_aγZ × Z) – (0.5 × γ_wZ × Z) = (5 × 0.33 × 0.85) + (0.5 × 5.049 × 0.85) – (0.5 × 5.39 × 0.85) = 1.403 + 2.146 – 2.291 = 1.258 kN/m

Net horizontal force (P_A) = 4.101 – 1.258 = 2.843 kN/m

Resistance to sliding

Frictional Force (F_f) = μN = 0.5 × 12.136 = 6.068 kN/m
F.O.S = F_f / P_A = 6.068/2.843 = 2.134
The factor of safety 2.134 > 1.5. Therefore, the drain is very safe from sliding.

Resistance to overturning

Taking moment about wall 1;

Sum of overturning moments (M_o) = (4.101 – 1.258) × (0.85/3) = 0.806 kNm per m
Sum of restoring moments (M_R) = (W₁ × 0.075m) + (W_w × 0.55m) + (W₂ × 1.025) + (W_b × 0.55) = (2.52 × 0.075) + (3.136 × 0.55) +(2.52 × 1.025) + (3.96 × 0.55) = 0.189 + 1.725 + 2.583 + 2.178 = 6.675 kNm/m

F.O.S = M_R / M_O = 6.675/0.806 = 8.281
The factor of safety 8.281 > 2. Therefore, the drain is very safe from overturning.

Bearing capacity check

Bending moment about the centerline of the base;

M = (W₂ × 0.475m) + (4.101 × 0.85/3) – (W₁ × 0.475m) – (1.258 × 0.85/3) = (2.52 × 0.475m) + (4.101 × 0.85/3) – (2.52 × 0.475m) – (1.258 × 0.85/3) = 1.197 + 1.162 – 1.197 – 0.356 = 0.806 kNm per m

Total vertical load (N) = 12.136 kN/m
Eccentricity (e) = M/N = 0.806/12.136 = 0.066m

Check: D/6 = 1.1/6 = 0.183m
Since e < D/6, there is no tension in the drain base.

Maximum pressure in the drain base (q_max) = P/B (1 + 6e/B) = 12.136/1.1 [1 + (6 × 0.066)/1.1] = 15.005 kN/m²
Minimum pressure in the drain base (q_min) = P/B (1 – 6e/B) = 12.136/1.1 [1 – (6 × 0.066)/1.1] = 7.061 kN/m²

Since q_min and q_max are lower than the allowable bearing pressure of the soil (150 kN/m²), bearing capacity check is satisfied.

Structural Design

Design of the Walls

Since the horizontal force due to surcharge load and backfill on Wall 1 > Wall 2, we adopt Wall 1 parameters for design. Using the centroid formula of a parallelogram for the pressure diagram of wall 1 to determine the distance (x) from the centroid to the base of the wall and distance (y) from the centroid to the top of the wall;

x = 0.85 [((4.609 + (2 x 4.95)) / (3(4.609 + 4.95))] = 0.43m

Thus, y = 0.85 – 0.43 = 0.42m
Taking moment at the top of the drain wall due to the active force;
M = 4.101 x 0.42 = 1.722 kNm per m

Taking moment at the base of the drain wall due to the active force;
M = 4.101 x 0.43 = 1.763 kNm per m

Since the moment at the base of the drain wall is greater than that at the top, we adopt the moment at the base for design.

At ultimate limit state;
M = 1.4 × 1.763 = 2.468 kNm per m

Flexural Design (Bending)

Given: Thickness of wall (h) = 150mm, Cover = 40mm, f_cu = 20 N/mm², f_y = 460N/mm², Rebars = 10mm

Effective depth (d) = 150 – 40 – (10/2) = 105 mm

K = M/(f_cubd²) = (2.468 x 10⁶) / (20 x 1000 x 105²) = 0.0112 (K < 0.156)
l_a = 0.5 + (0.25 – k/0.9)^0.5 = 0.5 + (0.25 – 0.0112/0.9)^0.5 = 0.987
Since 0.987 > 0.95, l_a = 0.95

A_s,req = M/(0.95f_y.la.d) = (2.468 x 10⁶) / (0.95 × 460 x 0.95 x 105) = 56.62 mm²/m
A_Smin = (0.13bh)/100 = (0.13 x 1000 x 150) / 100 = 195 mm²

Provide Y10 @ 300mm c/c (A_Sprov = 260 mm²/m)

Steel ratio check

4.0 > (100A_Sprov / bh) > 0.13
4.0 > (100 x 260) / (1000 x 150) > 0.13
4.0 > 0.17 > 0.13 (Steel ratio is satisfied)

Shear check

Ultimate design shear force on drain wall (V) = (1.4 × 4.101) = 5.741 kN/m

Shear stress (v) = V/bd = (5.741 × 1000) / (1000 × 105) = 0.055 N/mm²

Shear strength (v_c) = 0.632 × (100A_s/bd)^1/3 × (400/d)^1/4 × (f_cu/25)^1/3
v_c = 0.632 × [(100 × 260)/(1000 × 105)]^1/3 × (400/302)^1/4 × (20/25)^1/3 = 0.632 × 1.3529 × 1.3971 × 0.9283 = 1.109 N/mm²
Since v < v_c, no shear reinforcement required.

Design of the base

The pressure distribution diagram on the base at serviceability limit state is shown below;
q_min = 7.061 kN/m²
q_max = 15.005 kN/m²

bearing capacity 082251 — **Pressure distribution on the drain base**

At the ultimate limit state;

q_min = 7.061 x 1.4 = 9.885 kN/m²
q_max = 15.005 x 1.4 = 21.007 kN/m²

On investigating the maximum design moment at point A;

Water = 1.4 × [9.8 × 0.4 × 0.8 × (0.8/2 + 0.15) = 2.415 kNm/m
Base = 1.4 × [24 × 0.15 × 0.8 × (0.8/2 + 0.15) = 2.218 kNm/m
Earth pressure = [9.885 × 1.1 × (1.1/2)] + [(21.007 – 9.885) × 1.1 x 0.5 × (1.1/3)] = 8.223 kNm/m

Net moment = 8.223 – 2.415 – 2.218 = 3.59 kNm/m

On investigating the maximum design moment at point B;

Water = 2.415 kNm/m
Base = 2.218 kNm/m
Earth pressure = [9.885 × 1.1 × (1.1/2)] + [(21.007 – 9.885) × 1.1 × 0.5 × (2 × 1.1/3)] = 10.466 kNm/m

Net moment = 10.466 – 2.415 – 2.218 = 5.833 kNm per m

Since net moment at B > moment at A, we adopt 5.8833 kNm for design.

Flexural Design (Bending)

Given: Thickness of base(h) = 150 mm, Cover = 40 mm, f_cu = 20 N/mm², f_y = 460 N/mm², Size of rebars = 10mm

Effective depth (d) = 150 – 40 – (10/2) = 105mm

K = M/(F_cubd²) = (5.833 × 10⁶) / (20 × 1000 × 105²) = 0.0265 (K < 0.156)
la = 0.5 + (0.25 – k/0.9)^0.5 = 0.5 + (0.25 – 0.0265/0.9)^0.5= 0.97

Since 0.97 > 0.95, La = 0.95

A_Sreq = M/(0.95F_y.La.d ) = (5.833 × 10⁶) / (0.95 × 460 × 0.95 × 105) = 133.82 mm²/m
A_Smin = (0.13bh)/100 = (0.13× 1000 × 150) / 100 = 195 mm²

Provide Y10 @ 300mm c/c (A_Sprov = 260 mm²/m)

Shear Check

Calculating the maximum shear force at any section of the drain base;

Water = 1.4 × (9.8 × 0.4 × 0.8) = 4.39 kN/m
Base = 1.4 × (24 × 0.15 × 0.8) = 4.032 kN/m
Earth pressure = 0.5 × (21.007 + 9.885) × 0.8 = 12.356 kN/m

Net shear force = 12.356 – 4.39 – 4.032 = 3.934 kN/m

Shear stress (v) = V/bd = (3.934 × 1000) / (1000 × 105) = 0.037 N/mm²

Shear strength (V_c) = 0.632 × (100A_s/bd)^1/3 × (400/d)^1/4 × (f_cu/25)^1/3 = 0.632 × (100 × 260)/(1000 × 105)]^1/3 × (400/302)^1/4 × (20/25)^1/3 = 0.632 × 1.3529 × 1.3971 × 0.9283 = 1.109 N/mm²

Since v < V_c, no shear reinforcement required.

Detailing

Conclusion

This article has discussed the geotechnical and structural design of rectangular roadside drains. However, readers must note that only one load case has been treated. Therefore, a designer must consider other load cases or load combinations to ascertain the accuracy of the design. For example, it would be appropriate to rerun the design with the drain filled and when the drain is empty to determine the most critical load case or combination.

Causes of Deterioration of Used Concrete Sewer Pipes

By

Charity Mmuo

-

July 28, 2022

Concrete is the building material that is most usually used for sewer systems because of its favorable structural qualities, capacity for prefabrication, and freedom from form restrictions. For a variety of reasons, such as the effects of (bio)chemical deterioration, ageing, and the loss of soil support, the structural integrity of concrete sewer pipes degrades with time.

The design life of a sewer system is several decades. Due to the capital-intensive nature of maintaining a sewer system as well as the severe societal and financial consequences of catastrophic failure, accurate condition evaluation has become increasingly important over time.

Concrete sewer pipe — **Figure 1:** Installation of concrete sewer pipes

The two most frequent sources of data used to determine whether to repair or replace sewers are Closed-Circuit Television (CCTV) inspection and age. The difficulty of revealing deterioration on the outside of the sewer pipe wall, the low accuracy and reliability of visual inspection data, and the weak correlation between visual inspection data and material properties are just a few drawbacks that have recently been discovered with regard to these inspection methods.

Additionally, the majority of nations lack a database that contains precise information on the state of the subsurface infrastructure. Thus, it is obvious that clear knowledge about the real structural state of sewer systems is required in order to enhance current inspection techniques and enable adequate condition assessments.

The structural state of sewer networks has been the subject of numerous study investigations during the past few decades. Much emphasis has been paid to the biogenic sulphuric acid-induced degradation process that typically occurs in concrete sewer pipes.

Investigations into the sulphuric acid-producing bacteria and the chemical deterioration mechanisms that result at the inner surface of sewer pipes have revealed that the concrete’s calcium hydroxide and calcium silicate hydrate react with the acid to create gypsum and/or ettringite, which frequently causes an increase in porosity and, as a result, a decrease in strength and stiffness.

deterorating sewer — **Figure 2:** Deteriorated concrete sewer pipes

Additionally, naturally occurring carbondioxide in soil may migrate to the outside of concrete sewer pipes, where it may interact with hydrated cement in the presence of moisture. The strength, porosity, and pore size distribution of the cement paste may change as a result of this carbonation process. The assessment of the structural failure behavior of sewer pipes was the focus of additional experimental researches.

Despite the fact that the aforementioned investigations have focused on important issues and facts, a complete understanding of how (bio)chemical attack affects the mechanical performance of in-situ concrete sewer pipes is still lacking. However, this information is necessary to boost the suitable recommendations that municipalities and other stakeholders can use when making decisions about the upkeep and replacement of concrete sewer pipe systems.

Evaluation of Deterioration of old Sewer Pipes

Recently, researchers (Luimes et al, 2022) from the Department of the Built Environment, Eindhoven University of Technology, Eindhoven, The Netherlands, and Department of Hydraulic Engineering, Deltares, MH Delft, The Netherlands, studied the effects of biochemical attack on the mechanical performance of used concrete sewer pipes which they published in the journal, Construction and building materials (Elsevier).

Thirty-five used, unreinforced concrete sewer pipes provided by The Netherlands’ Municipalities of The Hague and Arnhem were used for the experimental program. The tested pipes, which range in age, size, and geometry, have been in use as combined sewer systems (i.e., the tested pipes installed in the 1920s and 1950s) or as improved separated systems (i.e., the tested pipes installed in the 1990s) up until July/August 2019 (pipes from Arnhem) and January 2020 (pipes from The Hague), respectively, without undergoing rehabilitation or protection treatments.

The test program involved concrete sewer pipes that had been in use for between 22 and 95 years. During that time, exposure to particular in-situ circumstances had resulted in some chemical attack, which may have been accompanied by mechanical damage from traffic loads and excavation. This might have reduced the concrete’s mechanical qualities, which might have affected the pipe’s ability to support structural loads. The inner and outer surfaces of the sewer pipes were first given a rigorous visual inspection during the study, and were then classified using various surface condition classes in order to explore these issues in more detail.

samples — **Figure 3:** Four examples of unreinforced concrete sewer pipe specimens of different age that were considered in the experimental program (Luimes et al, 2022)

The cross-sections of the sewer pipes were also tested for residual alkalinity using a phenolphthalein method, where a pH indicator was provided by a 1 percent solution of phenolphthalein. The solution turns from pink to magenta in moderately alkaline conditions with a pH in the range of 9.2 to pH 10. The solution turns magenta in extremely alkaline surroundings with a pH > 10, but it stays colorless in somewhat alkaline and acidic conditions with a pH 9.2.

The dissolution of solid calcium-hydroxide in the pore solution causes the pH of the concrete to rise to a level above 12.5 to 13, which is shown by the magenta zones. On the other hand, the colorless zones indicate the presence of a chemical attack in the past, which may have been brought on by carbonation (8.3 < pH < 9) and biogenic suphuric acid corrosion (1< pH < 3).

X-ray diffraction (XRD) data were used by the researchers to further analyze the type of chemical attack. A total of 6 different surface condition classes were identified, and a particular place may be described by many surface conditions. The surface conditions were classified as follows in accordance with nomenclature generally used by the sewer asset management community:

Smooth – A surface that is virtually intact and aesthetically comparable to the surface of a brand-new sewer pipe
Exposed granulates – Exposed granulates and porous mortar between granulates referred to chemically harmed surfaces where the granulates are now visible and/or have fallen off due to the loss of the thin outer mortar layer. In the latter class, the mortar that once held the exposed granulates together has degraded into a porous, loose material that is simple to spall off.
The deposits class — As a result of variations in the wastewater level, dark colored bands indicate obvious color changes and adhering deposits along the inner pipe surface.
Rough – A surface with minor chemical assault symptoms, wherein a portion of the thin outer mortar layer is still covering the granulates.
Excavation damage – This refers to relatively big scraped or broken off pieces brought about by bulldozers and excavators during the mechanical removal of the sewer pipes from the soil.

different levels of pipe deterioration — Figure 4: Representations of the 6 surface condition classes: (1) Smooth, (2) Exposed granulates, (3) Porous mortar between granulates, (4) Deposits — Dark coloured bands,
(5) Rough, and (6) Excavation damage, as observed on the inner and outer surfaces of the tested concrete sewer pipes (Luimes et al, 2022)

The types and degree of biochemical attack were respectively assessed by performing XRD analyses and phenolphthalein tests. The following conclusions were obtained from the findings.

The process of biogenic sulphide corrosion can be blamed for a major portion of the structural deterioration of old sewer lines. This process often results in a porous mortar layer between the granulates and a weak, corroded layer that looks like exposed granulates at the inside of the pipe.
Carbonation may have an impact on the sewer pipe’s outside, but it seems to have a very small impact on the pipe’s surface condition and is thought to be less detrimental to the structural integrity.
In contrast to the relatively new pipes from the 1990s, the old pipes from the 1920s and 1950s typically exhibit quite significant levels of chemical attack. Despite the fact that the level of chemical attack tends to increase with pipe age, an explicit relationship between pipe age and the degree of chemical attack cannot be determined from measurement data because the pipe’s specific material composition and the surrounding environment also have a significant impact on the level of chemical degradation.

According to the researchers (Luimes et al, 2022), the study’s findings can be used to develop and improve inspection and condition assessment standards. Further research has shown that biogenic sulphide corrosion can significantly contribute to the mechanical deterioration of sewer pipes, making it prudent to keep an eye on how this corrosion process is progressing inside in-situ sewer pipes.

Source:
Luimes R. A., Scheperboer I.C., Suiker A.S.J., Bosco E., and Clemens F.H.L.R. (2022): Effect of biochemical attack on the mechanical performance of used concrete sewer pipes. Construction and Building Materials 346 (2022) 128390 https://doi.org/10.1016/j.conbuildmat.2022.128390

The contents of the cited original article published by Construction and Building Materials (Elsevier) is open access, under the CC BY license (http://creativecommons.org/licenses/by/4.0/) which allows you to share and adapt (remix) the article provided the appropriate credit is given, and the link to this license provided.

Corrosion of Buried Mild Steel Corrugated Sheets

By

Charity Mmuo

-

July 26, 2022

A high moisture content, good aeration, a high level of acidity, and a considerable number of soluble salts, can make a soil become corrosive. Metal alloys may undergo a dealloying process in corrosive soils as a result of the hostile surroundings there. The combination of soil corrosivity and dealloying corrosion is to blame for the spread of corrosion in buried steel structures.

Corrugated metal pipes (CMPs) and corrugated metal culverts (CMCs) are subterranean steel structures that have been utilized in traffic networks and water supply systems in North America and Europe since the 1850s . Corrugated profiles are being used in pipelines and culverts in order to help these structures interlock with the surrounding backfill soils and increase confinement properties as well as overall capacity. In cold coastal regions, salt used to melt the snow causes significant chloride deposits in the soil, which exposes buried corrugated steel structures to, resulting in the formation of thick rust layers.

Corrosion, which results from exposure to hostile conditions where chlorides attack metals with or without protective coatings, is the main reason why buried steel constructions deteriorate. Such deterioration is characterized by thickness loss and a decline in the axial and flexural stiffness of the steel. To avoid structural failure, it is necessary to reevaluate the structural capability and estimated service life. Around the years, soil corrosion has had an impact on several underground steel constructions all over the world (Ezzeldin et al., 2022).

Temperature variations, humidity, airborne sea salt, salts dissolved during snow thawing, and other chemical elements can all contribute to corrosive situations that cause steel to deteriorate over time. The risk of corrosion in buried steel structures is brought on by the very variable soil conditions in the area.

The corrosion process is accelerated by repeated daily and seasonal exposure to salt and water, particularly when there is an increase in temperature fluctuation in cold regions due to the effects of global warming. To monitor potential structural degradation and damage, regular in-service inspection of culvert performance is therefore essential.

The US Pipelines and Hazardous Material Safety Administration states that exterior corrosion is typically to blame for pipeline system ruptures brought on by corrosion. In addition to causing environmental issues, corrosion is a major element in the aging of networks and facilities, needing maintenance and rehabilitation that can put a large financial strain on the nation’s budget.

Recent Research Study on Corrosion

Recent research from the Department of Civil and Resource Engineering and the Department of Mechanical Engineering at Dalhousie University in Nova Scotia, Canada, studied the accelerated laboratory corrosion test on corrugated mild steel structures buried in cohesionless soils. The study utilized repeated wet/dry cycles to simulate the effects of chloride deposits on the buried steel structures. The findings of the study were published in the journal, Case Studies in Construction Materials.

corrugated steel pipe — Corrugated steel pipe

For the purpose of simulating the effects of chloride deposits on corrugated mild steel structures buried in cohesionless soil, the researchers (Ezzeldin et al., 2022) devised an accelerated laboratory corrosion test using multiple wet/dry cycles. The test was initiated by applying a 3.5% NaCl electrolyte solution to cohesionless soil above buried corrugated steel coupons, which were then subjected to repeated wet/dry cycles.

The study also examined the structural profile geometry’s loss of thickness and the degree to which the steel’s tensile strength, ductility, and hardness had been degraded. This investigation focused on the deterioration of mild steel coupons as a result of corrosion. Both micrometer gauge measurements and the weight loss method were used to calculate the corrosion damage.

Corrugated mild carbon steel (CS) type B coupons were used for the tests. The coupons had a corrugation depth of 13 mm, a wavelength of 68 mm, and a thickness of 1.5 mm. Each coupon’s total projection dimensions were 110 mm by 110 mm in order to accommodate multiple waves, one crest and two valleys at the surface facing the dirt. The interface geometry of buried CMPs and CMCs was simulated by the corrugated specimens.

Each coupon was buried beneath a layer of finely graded, well-compacted, cohesionless dirt. The system included two timers to regulate the wetting (spraying) and drying stages, a tank of distilled deionized water, a pump to transfer water from the tank, stainless steel pipes and fittings to carry the water from the tank to an oven, a convection oven to distribute heat evenly to the coupons during the drying stages, and other components.

Two programmed timers were used to regulate the timing for each stage while the wet/dry cycles were repeated. The process of soaking (spraying) took 4 seconds. The soil above each coupon was sprayed with the distilled water as it was transferred by the pump from the tank. To provide dissolved oxygen and keep the water at normal temperature, the tank containing the distilled water was left open to the atmosphere. The heat from the oven was then used to finish the drying process. The soil temperature rose gradually during the drying stage, reaching a nearly dry state (i.e., a recorded soil temperature of about 90 ℃) in about 60 minutes.

The Experimental Setup of the Corrosion test. — **The Experimental Setup**

Each full wet/dry cycle took about 60 minutes to complete because the spraying was done right away at the start of the cycle. When the soil was sprayed during the wetting stage, the temperature abruptly dropped from around 90℃ to about 60 ℃, a drop of about 30 degrees. Prior to the start of the following wetting stage, the temperature was raised once again during the ensuing drying stage in order to evaporate the majority of the remaining water.

To keep the salt content in the soil at the same level (i.e., 3.5 percent) during subsequent wetting procedures, which were carried out using just distilled water, the electrolyte solution was added just once, using the same quantity of distilled water as utilized for each wetting stage.

Gravity caused the salts to settle onto the steel coupons as a result of the repeated wetting process used to completely saturate the soil throughout each cycle. In order to provide aeration and break up salt crusts that had collected, a spatula was used to mix only the top layer of soil and then compact it on top of each coupon after every 20 wet/dry cycles.

In order to reduce any potential impact on corrosion propagation, the soil adhering to the metal surface’s interface was kept undisturbed by the researchers. To assess the spread of corrosion caused by each set of cycles, five coupons were evaluated with varying totals of wet/dry cycles (50, 100, 200, 400, and 800 cycles).

Findings from the Study

At the end of the experiment by Ezzeldin et al (2022), the following conclusions were made;

Mild steel corrosion was accelerated by repeated wet/dry cycles in the absence of a protective layer (such as zinc coating).
In the steel coupons, where more induced stresses were created during the production of the corrugated steel sheets, the degree of corrosion was greater at the corrugation crests and valleys.
Rust layers of a similar nature and morphology developed on all of the test specimens, imitating the effect of acidic environments on buried steel structures in cold climates. This effect could be clearly observed in both valleys of the coupon treated to 800 wet/dry cycles.
While the rate of corrosion steadily decreased, the level of corrosion damage increased when the number of wet/dry cycles was increased. Mixed corrosion modes, such as deep pitting that produced cavities, were a part of the corrosion that eventually evolved.
The structural geometry, which lost thickness, and the mechanical qualities, such as tensile strength, ductility performance, and hardness, all degraded as a result of the steel coupons’ deterioration. Due to the reduced axial and flexural rigidity, subterranean steel structures like CMPs and CMCs would no longer be able to function to their full potential.

A mathematical model, requiring the measurement of four physicochemical parameters at the interface between the soil and the mild steel surface, was used by the researchers to provide an approximate prediction of the depth of corrosion damage in buried steel structures. The present study suggests employing this mathematical model to make approximate predictions of corrosion damage over time, based on the following Eqs. (1,2):

ν_p = C₀exp[-(q₁pH + q₂ρ + q₃E_Redox+ q₄E_s-p)] ——– (1)

z(t) = ν_pt + [(υ₀ – ν_p)/q₀] [1 – exp(- q₀t)] ——– (2)

Where:
ν_p = the average long-term corrosion rate;
υ₀ = the initial corrosion rate = 0.6743
C₀ = constant 1, = 12.2652,
q₀= constant 2, = 1.7326,
q₁ = pH constant = 0.6623,
q₂ = resistivity constant, = 0.0069 Ωm,
q₃ = redox potential constant, i.e., 0.0027 mV/SHE,
q₄ = soil-structure electric potential constant, i.e., 0.981 V/Cu/CuSO₄,
z(t) = the maximum depth of corrosion damage at time (t).

The corrosion damage and reduction in nominal thickness (%) related to number of cycles from the accelerated wet/dry test and number of years from the mathematical model is shown in the Table below;

Reference(s)
Ezzeldin I., El Naggar H., Newhook J. and Jarjoura G. (2022): Accelerated wet/dry corrosion test for buried corrugated mild steel. Case Studies in Construction Materials 17 (2022) e01152. https://doi.org/10.1016/j.cscm.2022.e01152

The contents of the cited original article published by Case Studies in Construction Materials (Elsevier) is open access, under the CC BY license (http://creativecommons.org/licenses/by/4.0/) which allows you to share and adapt (remix) the article provided the appropriate credit is given, and the link to this license provided.

Number and Depth of Borings for Soil Investigation

By

Ubani Obinna

-

June 12, 2024

Table of Contents

The entire area of a project site cannot be fully explored for site investigation due to logistical and financial constraints. To provide enough information for the design and construction of the foundation of a building or highway, geotechnical engineering consultants must make good decisions about the location, number, and depth of borings for soil investigation. The zone of soil that will be affected by the structural loads should be covered by the number and depth of borings. There are no fixed guidelines to adhere to.

More often, the number and depths of borings are governed by experience based on the geological nature of the ground, the importance of the structure, the structural loads, and the availability of equipment. The minimum number and depth of borings may be specified by building regulations and regulatory authorities in the local area.

Whenever possible, boreholes should always be dug close to the intended foundation location. Where the bearing stratum’s depth is uneven, this is crucial. The boreholes should be precisely positioned in relation to the proposed structures, both in terms of level and location.

A grid of holes that are evenly spaced serves as an appropriate design of boreholes when the layout of the structures has not been established at the time the soil investigation is being conducted. It is feasible to use a grid of boreholes with in-situ probes of some kind, such as dynamic or static cone penetration tests, spaced more closely apart within the borehole grid for large areas. EC 7 recommends, for category 2 investigations, that the exploration points forming the grid should normally be at a mutual spacing of 20 — 40 m.

A challenging issue that is intimately related to the relative costs of the soil investigation and the project for which it is done is the necessary number of boreholes that must be sunk at any specific place. Normally, as more boreholes are drilled, more information about the soil conditions becomes available, allowing for more efficiency in the foundation design. Additionally, the likelihood of encountering unforeseen or challenging soil conditions, which would significantly raise the cost of the foundation work, decreases over time.

An economic limit, however, is reached when the cost of borings outweighs any savings in foundation cost and merely drives up the project’s overall cost. In order to determine the true dip of the strata, it is recommended that at least two and ideally three boreholes be drilled for all but the smallest structures. However, inaccurate assumptions about stratification can still be made.

However, it is very important that the number of boreholes be sufficient to detect any variances in the soil of the site. If the loads placements (such as column footing positions) on the structure’s footprint are known (which is frequently not the case), you should think about drilling at least one borehole where the heaviest load is.

The depth to which boreholes should be sunk is governed by the depth of soil affected by foundation-bearing pressures. The vertical stress on the soil at a depth of one and a half times the width of the loaded area is still one-fifth of the applied vertical stress at the foundation level, and the shear stress at this depth is still appreciable. Thus, borings in soil should always be taken to a depth of at least one to three times the width of the loaded area.

The borings are relatively shallow for narrow, widely spaced strip or pad foundations, but for big raft foundations, the borings must be deep unless rock is present within the required depth. When strip or pad footings are placed closely together, the pressure zones overlap, and the entire loaded region effectively becomes a raft foundation with correspondingly deep borings. To cover the zones of soil affected by loading transmitted through the piles in the case of piled foundations, the ground should be studied below the pile-point level.

EC 7 recommends a depth of five shaft diameters below the expected toe level. It is usual to assume that a large piled area in uniform soil behaves as a raft foundation with the equivalent raft at a depth of two-thirds of the length of the piles.

As a guide, a minimum of three boreholes should be drilled for a building area of about 250 m² (2500 ft²) and about five for a building area of about 1000 m² (10,000 ft²). Some guidelines on the minimum number of boreholes for buildings and for due diligence in subdivisions are given in Table 1.

Area (m²)	Numbers of boreholes (minimum)
< 100	2
250	3
500	4
1000	5
2000	6
5000	7
6000	8
8000	9
10000	10

Table 1: Guidelines for the Minimum Number of Boreholes for Buildings

Some general guidance on the depth of boreholes is provided in the following:

In a compressible soil such as clays, the borings should penetrate to at least between 1 and 3 times the width of the proposed foundation below the depth of embedment or until the stress increment due to the heaviest foundation load is less than 10%, whichever is greater.
In very stiff clays and dense, coarse-grained soils, borings should penetrate 5 m to 6 m to prove that the thickness of the stratum is adequate.
Borings should penetrate at least 3 m into the rock.
Borings must penetrate below any fills or very soft deposits below the proposed structure.
The minimum depth of boreholes should be 6 m unless bedrock or very dense material is encountered.

Guidelines for the Minimum Number and Depth of Borings for Common Geostructures

For foundation construction on compressible soils (clay and similar materials) with sufficient strength to initially support the structure, it is important to ensure that borings penetrate these compressible layers. Alternatively, borings should reach a depth where the additional stress placed on deeper strata is minimal, ensuring negligible consolidation that wouldn’t significantly impact the proposed structure’s settlement.

Exceptions exist for exceptionally heavy loads or situations where seepage or other factors are paramount. In such cases, borings may be terminated upon encountering bedrock or penetrating a stratum of exceptional bearing capacity and rigidity for a short distance.

However, this is only advisable if prior explorations in the vicinity or regional stratigraphic knowledge confirm that these strata possess adequate thickness or are underlain by even stronger formations. If these confirmations are lacking, a subset of the borings must be extended further to verify the thickness of the strong strata, regardless of the underlying material’s characteristics.

The recommended guidelines for the number and depth of borings for common civil engineering structures are provided below;

Shallow Foundation for Buildings

Minimum number of boreholes
1, but generally boreholes are placed at node points along grids of sizes varying from 15 x 15m to 40 x 40 m.

Minimum depth
The minimum depth of soil exploration for foundations should be 5 m or 1B to 3B, where B is the width of the foundation. Additionally, the depth of exploration should extend to a depth where the increment in stress is equal to or less than 10% of the maximum foundation pressure.

Deep (Pile) Foundation for Buildings

Minimum Number of Boreholes
1 boring, but generally boreholes are placed at node points along grids of sizes varying from 15 x 15m to 40 x 40 m

Minimum Depth of Boring
25m to 30m;
If bedrock is encountered, drill 3m into it

Bridge

Minimum number of boreholes
Abutments – 2
Piers – 2

Minimum Depth of Boring
25m to 30m;
If bedrock is encountered, drill 3m into it

Retaining Walls

Minimum Number of Boreholes
Length < 30 m: 1
Length > 30 m: 1 every 30 m, or 1 to 2 times the height of the wall

Minimum Depth of Boring
1 to 2 times the height of the wall. For walls located on bedrock, drill 3m into the bedrock

Cut Slopes

Minimum Number of Boreholes
Along the length of slope: 1 every 60 m;
if the soil does not vary significantly, 1 every 120 m
On slope: 3

Minimum Depth of Boring
6m below the bottom of the cut slope

Embankments, Including Highways

Minimum Number of Borings
1 every 60 m;
if the soil does not vary significantly, 1 every 120 m

Minimum Depth of Boring
The greater of 2 x height or 6 m

Types of Abutment

Cantilever Abutment Walls

Free Cantilever Abutment Walls

Counterfort Abutment Walls

Propped Cantilever Abutments

Open Abutments

Bankseats

Spill-through abutments

Integral Abutments

Reinforced Earth Abutments

Abutment Foundations

Abutment Approach Slab

Design of Abutments

Loadings on Abutments

Soil Loading

Vehicle loading (surcharge)

Compaction Pressure

Swelling Pressure

Effects of seasonal deck expansion and contraction

Load Combinations for Abutments

Stability of Abutments

Sliding

Overturning

Overall instability

Structural Design of Abutments

Base Design

Wall Design

Standard Penetration Test (SPT)

Skin friction

End bearing

Cone Penetration Test (CPT)

End Bearing

Shaft Resistance

Worked Example: Pile Load Capacity Using SPT Data

Test Setup and Experimental Study

Overall Results and Failure Modes

Comparison with Design Proposals

References

The T-shaped Concept

The T-shaped Model

Academic Excellence

Skill Acquisition/Entrepreneurship/Leadership

Professional Development

Final Words

Failure Load and Settlement

Main Factors of Concern

Shape Efficiency

Material Footprint

Durability

Conclusions

Worked Example on the Design of Rectangular Roadside Drains

Geotechnical Design

Wall pressure calculations

Resistance to sliding

Resistance to overturning

Bearing capacity check

Structural Design

Design of the Walls

Flexural Design (Bending)

Steel ratio check

Shear check

Design of the base

Flexural Design (Bending)

Shear Check

Detailing

Conclusion

Evaluation of Deterioration of old Sewer Pipes

Recent Research Study on Corrosion

Findings from the Study

Guidelines for the Minimum Number and Depth of Borings for Common Geostructures

Shallow Foundation for Buildings

Deep (Pile) Foundation for Buildings

Bridge

Retaining Walls

Cut Slopes

Embankments, Including Highways

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