Scielo RSS <![CDATA[Journal of the South African Institution of Civil Engineering]]> vol. 54 num. 1 lang. es <![CDATA[SciELO Logo]]> <![CDATA[<b>The creation and application of a national freight flow model for South Africa</b>]]> South Africa suffered from a historical lack of freight-flow information that was detrimental to infrastructure planning, optimal network development and market structuring. This paper proposes a methodology that can fill this gap, be repeated annually and is, by its nature, not prone to the errors of market surveys. The methodology develops a comprehensive description of South Africa's surface freight flow market space based on the definition of four definitive freight flow market segments. The results from the annual South African National Roads Agency (SANRAL) traffic counts are allocated to these segments to develop national road freight flows. For rail freight flows, the rail database is reclassified on a station-to-station (i.e. origin-destination) basis to match the freight flow market segments developed. Consequently, modal flows, market share and total flows for all freight flow market segments and the geographical groupings with the segments can be analysed and reported each year. The results confirm the deteriorating role of South Africa's rail system amidst growing freight demand, as well as the concomitant over-cropping of the road network, and therefore enable the development of specific national freight transport policy recommendations. <![CDATA[<b>Challenges confronting road freight transport and the use of vehicle-pavement interaction analysis in addressing these challenges</b>]]> Traditional arguments for maintaining riding quality of pavement are expanded in this paper to examine the effects of deteriorating riding quality on vehicle operating costs, freight damage and logistics. The objectives of this paper are to analyse the effects of different levels of riding quality on a truck and its freight, and to discuss potential applications of the analysis in terms of effectiveness of the freight transport system. The paper discusses needs and drivers influencing freight transport costs, vehicle-pavement interaction concepts, and the potential physical effects and costs from roads with deteriorating riding quality. A case study is presented analysing vehicle-pavement interaction for selected roadways in California. It is concluded that investments in pavement and freight transport industry improvements can be investigated by applying vehicle-pavement interaction analysis to evaluate damage to pavement, vehicle and freight that would result from alternative levels of pavement riding quality. The paper recommends that existing concepts, tools and resources such as dedicated truck lanes and vehicle-pavement interaction analysis can help to improve the freight transport system. A framework is proposed to better understand the scale of potential impacts of riding quality from localised effects to larger-scale influences, including costs to customers and global competitiveness. <![CDATA[<b>Air void characterisation of HMA gyratory laboratory-moulded samples and field cores using X-ray computed tomography (X-ray CT)</b>]]> The research work presented in this paper deals with the characterisation of the internal structure of hot-mix asphalt (HMA), incorporating both gyratory compacted samples produced in the laboratory and field cores. The primary objective was to determine the optimum trim depth on either end of laboratory-moulded HMA cylindrical samples that would optimise the air void (AV) uniformity in the test specimens. The analysis was based on the X-ray Computed Tomography (X-ray CT) scanning tests and subsequent image analyses. Two Texas HMA mixes, namely a coarse-graded (Type B) and a fine-graded (Type D) mix, with gyratory samples compacted in the laboratory to two different heights (110 and 164 mm) were evaluated for their internal structure in terms of the distribution of both the AV content and AV size. Analysis of the results indicated that the coarse-graded HMA mix (Type B) and the taller (164 mm in height) gyratory-moulded samples would be more likely associated with a more heterogeneous distribution of the AV content and AV size, respectively. Supplemented with field cores, the X-ray CT results indicated significantly poor AV content distribution (i.e. higher AV content and weakest area) at the ends, particularly in the top and bottom 20 mm zone of the samples. Thus, for 150 mm diameter samples of height equal to or greater than 110 mm, trimming a minimum of 20 mm on either side of the gyratory compacted samples should be given due consideration without compromising the specimen aspect ratio and NMAS coverage requirements (NMAS - nominal maximum aggregate size). In general, test specimens should always be cut from the middle zone of the SGC moulded samples where the AV is less heterogeneously distributed. <![CDATA[<b>Weak interlayers in flexible and semi-flexible road pavements</b>: <b>Part 1</b>]]> Weak layers, interlayers, laminations and/or interfaces in the upper structural layers of road pavements are specifically prohibited in most road-building specifications. However, such layers are extremely common and often lead to premature pavement distress. In Part 1 of this two-part set of papers, it is shown that from experience with heavy vehicle simulator (HVS) and dynamic cone penetrometer (DCP) testing, the presence of such layers and/or conditions at any depth in the structural layers of a flexible or semi-flexible pavement is far more deleterious than is commonly appreciated. In Part 2 the effects of these weak layers are further modelled and discussed using various examples based an HVS testing and mechanistic pavement analyses. In particular, a weak upper base course of a cemented pavement under a thin bituminous surfacing may lead to severe surfacing (and upper base) failure within a matter of weeks to months after opening to traffic, not excluding failure even during construction. In this paper (Part 1), the causes of weak layers, interlayers, laminations and/or interfaces, together with simple methods for their detection during construction and analyses of their effects on the structural capacity of flexible and semi-rigid (cemented) road pavements, are briefly discussed. <![CDATA[<b>Mechanistic modelling of weak interlayers in flexible and semi-flexible road pavements</b>: <b>Part 2</b>]]> This paper (Part 2 of a two-part set of papers) discusses models and illustrates the adverse effects of weak layers, interlayers, laminations and/or weak interfaces in flexible and semi-flexible pavements, also incorporating lightly cemented layers. The modelling is based on mechanistic analyses for pavement design and evaluation. In Part 1, the effects of these relatively weak layers, interlayers, laminations and/or weak interfaces were discussed. It was shown that methodologies are available to detect and investigate the existence of these weak layers in cemented pavement layers. In Part 2, several cases of the above conditions for different road pavement types are discussed, with field examples. Mechanistic analyses were done on a typical hot mix asphalt (HMA), several cases of a cemented base pavement and a granular base pavement, with and without these weak layers and interface conditions to demonstrate their adverse effects. The analyses focus on the strain energy of distortion (SED) as a pavement response parameter to indicate the potential for structural damage expected within the pavement structure or layer. Generally, the higher the SED, the higher the potential damage in the pavement layer. SED shows some potential for quantifying the relative effects of these weak layers, interlayers, laminations and/or weak interfaces within flexible and semi-flexible pavements. <![CDATA[<b>The effect of parameters on the end buffer impact force history of the crane</b>]]> An overarching investigation was conducted to provide engineers with guidelines for designing crane supporting structures. The focus of this study was to determine whether the identified parameters had an effect on the end buffer impact force history when the electric overhead travelling crane collides with the end stops of the supporting structure. Seven design codes which were reviewed do not consider the crane and its supporting structure as a coupled system. This simplification ignores some of the parameters which have a significant influence on the impact force, which could lead to the codified estimates being sometimes unconservative. During the experimental tests it was discovered that some of the parameters could not be accurately controlled and/or monitored. This led to the development of a finite element (FE) model of the full-scale experimental configuration which was used to conduct advanced simulations. The FE model considered the crane and the supporting structure as a coupled system, in which the parameters were individually varied to obtain its effect on the impact force history. The results showed that some of the individual parameters do have a significant effect on the impact force history. <![CDATA[<b>Estimation of the maximum end buffer impact force for a given level of reliability</b>]]> The first paper in this set of two, titled The effect of parameters on the end buffer impact force history of the crane (see page 55), examined the effect of a change in the magnitude of the parameter on the end buffer impact force history. This paper investigates to what degree a change in the magnitude of the parameter alters the impact force history. This was accomplished through a sensitivity analysis performed by individually varying the magnitude of the parameter in the FE model. For each case individual maximum impact forces were obtained. The maximum impact force could not simply be selected by choosing the greatest value from the sensitivity study. A constraint optimisation technique for a given level of reliability ((3) using the FE simulation data was used to determine the maximum impact force. A comparison between the constraint optimisation and codified results showed that SABS 0160-1989 underestimates the impact force by 18%, while SANS 10160-2010 substantially overestimates the impact force by 64% for a level of reliability of β = 3. If the relevant clauses of SANS 10160-6 that pertain to end stop design are used in their present form, this will result in a conservative design, whereas SABS 0160 has a probability of 2.3% of being exceeded. <![CDATA[<b>Assessment of the behaviour factor for the seismic design of reinforced concrete structural walls according to SANS 10160 - Part 4</b>]]> Reinforced concrete structures, designed according to proper capacity design guidelines, can deform inelastically without loss of strength. Therefore, such structures need not be designed for full elastic seismic demand, but could be designed for a reduced demand. In codified design procedures this reduced demand is obtained by dividing the full elastic seismic demand by a code-defined behaviour factor. There is, however, no consensus in the international community regarding the appropriate value to be assigned to the behaviour factor. The purpose of this study is to assess the value of the behaviour factor currently prescribed by SANS 10160-4 (2011) for the design of reinforced concrete structural walls. This is done by comparing displacement demand to displacement capacity for a series of structural walls. The first step in seismic force-based design is the estimation of the fundamental period of the structure. The influence of this first crucial step is investigated in this study by considering two period calculation methods. It was found that, regardless of the period calculation method, the current behaviour factor value prescribed in SANS 10160-4 (2011) is adequate to ensure that inter-storey drift of structural walls would not exceed code-defined drift limits. <![CDATA[<b>A rational approach to predicting the buckling length of compression chords in prefabricated timber truss roof structures braced by means of diagonal bracing</b>]]> In South Africa, timber-trussed roofs supporting concrete tiles have for many years often been braced solely by means of diagonal braces. Failures have shown that the diagonal brace was inadequate for larger span roofs, and the use of diagonal bracing has subsequently been limited to spans of less or equal to 10 m. When designing the compression chords of a timber truss in a braced roof, SANS 10163:1 (2003) recommends a minimum effective length for out-of-plane buckling of not less than 15 x b, which is 540 mm for a 36 mm wide member. This effective or out-of-plane buckling length of the top chord was later assumed to be equal to the spacing of the trusses. With the availability of PC-based packages that are able to perform three-dimensional buckling analyses, it is perhaps useful to investigate the validity of using the effective length equal to the truss spacing, and then also the 10 m limit on span for roofs braced by means of diagonal braces. A common error made when analysing three-dimensional buckling problems is to assume connectivity on the centreline of the members, thereby neglecting eccentricity between the centreline of the bracing and the centreline of the member being braced (see Figure 1). In timber-trussed roofs, the diagonal brace is nailed to the underside of the top chord of a number of adjacent trusses. The brace runs at more or less 45° and triangulation appears to be complete when viewed on plan, as the battens form the other elements of the bracing system triangulation. Trusses some distance from the trusses that are connected to the diagonal brace can, however, only obtain lateral support via the battens that are connected to the top of the compression chords. The authors feel that a more correct way of analysing a timber-trussed roof, braced by means of a diagonal brace, requires that the eccentricity between the centreline of the battens on top of the compression chords and the centreline of the braced points on the underside of the compression chords be taken into account. Furthermore, the connections between the battens and the top chord are not infinitely stiff and this stiffness, together with the low torsional rigidity of the timber members, should be taken into account in the buckling analysis. The analysis can be further improved by taking the out-of-plane bending stiffness of the web members into account. All these factors will influence the buckling length of the compression chords to some degree. In this paper, the authors show how incorrect assumptions may mislead the designer into believing that the buckling length is equal to or less than the spacing of the trusses. They also show that, even though the bracing members have been placed on the correct sides of the top chord in the analysis, incorrect assumptions about the torsional stiffness of the top chords can lead to buckling lengths that are slightly less than when a more realistic torsional stiffness is used.