Shaking table study of a brick veneer house subjected to blast vibrations

Abstract

This article reports on a unique shaking table test examining the performance of a single-room house to simulated blast vibrations. The unreinforced masonry veneer specimen that is constructed to represent typical Australian residential construction is subjected to a total of 564 blast vibrations increasing in peak component velocity from 1 to 383 mm/s. Damage thresholds in the masonry veneer are examined along with the damage–drift relationship with reference to standardised damage categories provided in Australian Standard 2870. Importantly, drifts recorded in the veneer at recommended vibration limits are found to be well below the lowest drifts needed to cause damage. The lowest threshold of damage in the masonry veneer occurred at a drift equal to 0.10% (1/960) at a peak component velocity of 148 mm/s, nearly 30 times the recommended environmental limit, demonstrating that current environmental limits are conservative with respect to damage potential for residential structures in good condition.

Keywords

brick crack damage drift threshold veneer vibration

Introduction

Throughout Australia and the world, houses are situated in close proximity to blasting operations from mines and quarries where blasting is employed for the efficient extraction of rock. Although blast vibrations are limited to low levels, they are typically perceptible and generate annoyance in nearby residents with concern that their home is being damaged by the blasting activity. Past studies have examined the onset of damage in existing structures with the US Bureau of Mines (USBM) RI8507 (Siskind et al., 1980) being the largest and most comprehensive body of work in this field. However, in-field subjects had unquantified levels of residual stress from ordinary occupant and environmental loads. Such studies have focused on generating vibration limits based upon soil particle velocity to prevent damage to structures. A racking response caused by blast vibrations has the potential to cause similar damage patterns to heave/subsidence-related damage which is prevalent in many areas subjected to blasting. This often leads to confusion between the causes of cracking in homes. This project examines the correlation between damage and drift in brick veneer as a step towards generating more rational criteria for limiting vibrations. A previous experimental study has investigated the damage–drift relationship for brick veneer through static testing (Heath et al., 2008). The most reliable method of identifying damage thresholds is to decouple residual stresses from the stresses induced from blast vibrations by constructing and testing a structure in pristine condition then progressively subjecting it to increasing levels of vibration to monitor the initiation and propagation of damage. This article describes a project unprecedented in nature involving a shaking table test of a brick veneer structure, the most common construction type in Australia, subjected to gradually increasing levels of blast vibrations. Levels of drift and vibration for threshold and increasing damage states relative to damage categories stipulated by Australian Standard (AS) 2870 (1996) are reported.

Literature review

With the exception of a single fatigue study reported by Stagg et al. (1984), no shaking table studies have been conducted to investigate damage thresholds in brick veneer residential structures subjected to vibrations from surface blasts in mines and quarries. Masonry behaviour is dependent on loading rate and pattern, characterised by greater stiffness and strength under dynamic loading (Abrams, 1996; Magenes and Calvi, 1994; Tomazevic, 1994; Tomazevic et al., 1996). A loading regime, including a limited number of high-intensity events, is appropriate for the examination of performance criteria while progressive increases in vibration are appropriate for investigating damage thresholds. However, this approach may be conservative since the structure is likely to be in a weakened state at the onset of damage (Gulkan et al., 1990). Numerous dynamic tests have been performed to examine in-plane and out-of-plane behaviour of masonry structures (Borchelt and Klingner, 2009; Clough et al., 1990; ElGawady et al., 2003; Gulkan et al., 1990; Klopp and Griffith, 1998; Magenes and Calvi, 1994; Mengi and McNiven, 1989; Paquette and Bruneau, 2000; Paulson and Abrams, 1990; Toranzo et al., 2009). Clough et al. (1990) reported on five full-scale tests of masonry structures subjected to seismic loading, and it was found that vertical table motions did not adversely affect results relative to uniaxial or biaxial horizontal shaking. No significant difference in performance was observed between structures excited in a uniaxial or biaxial manner (Gulkan et al., 1990). Tomazevic (1994) noted that the vertical component of three-dimensional ground motions may be disregarded in experimental studies. These findings demonstrate that shaking table studies investigating structural response to blast vibrations could neglect the vertical component of vibrations without adversely affecting the outcome. Similarly, the horizontal components of blast vibrations could be independently simulated in two orthogonal directions on a shaking table without significantly influencing the performance of the structure.

Blasting is an efficient method for the extraction of rock overlying valuable minerals. Between 20% and 30% of explosive energy is unavoidably lost from surface blasting which becomes responsible for disturbance to neighbours that are unavoidably close to mines and quarries in many regions throughout Australia and the world. Rock overburden is drilled, filled with explosive charges and then packed with stemming (crushed rock) to contain the blast (Figure 1). Controlled blasting techniques require the detonation of charges having carefully timed delays to maximise the efficiency of blasting, with surface mines typically detonating between 90 and 3200 kg of explosives per delay with more than 100 tonnes of explosives in total (Sharma, 2008; Singh and Singh, 2005). The undesirable outcome of this process is the generation of airblast and ground vibrations that may propagate many kilometres at perceptible levels (Crum et al., 1992). The geology of the propagating medium and timing (delay) between charges in the detonation sequence have the greatest impact on the ground vibration waveform.

Figure 1.

Example bench blast design used to remove rock overburden showing millisecond delays between detonation of charges.

Airblast (air overpressure) generated from the blasting travels in a spherical manner, therefore rapidly reducing in magnitude with distance and possessing less damage potential than ground vibrations. Walls and windows may experience a flexural response due to the face loading caused by airblast which is largely responsible for the audible structural response (secondary rattling). Due to the speed of transmission of airblast being limited by the speed of sound, this response occurs after the faster travelling passage of ground vibrations has excited the structure. Consequently, airblast is not important when investigating damage thresholds in brick veneer structures.

The absorption of energy is a function of the material’s deformational properties, thus the decay in amplitude becomes a function of energy loss per cycle (Dowding, 1996). Ground vibrations comprise compression (P) waves, surface (S) waves and Rayleigh (R) waves. Higher frequency P and S waves decay much more rapidly than R waves which are substantially lower in frequency. At large distances from the blast, R waves dominate ground particle motion since they have experienced fewer deformational cycles compared to the higher frequency P and S waves. Furthermore, a separation develops between wave types at large distances since P and S waves travel more rapidly than R waves through a given media (Dowding, 1996). Vibrations from blasting in coal mines typically have a trailing large-amplitude and low-frequency wave compared to vibrations generated from quarry blasting (Siskind et al., 1980). Vibrations from coal mine blasts have lower principal frequencies possessing greater damage potential due to their similarity with natural frequency of residential structures (Konon and Schuring, 1983). The above studies demonstrate that blast vibrations generating from large surface mining blasts that arrive in residential areas have a greater potential to excite residential structures relative to blast vibrations generated from smaller quarry blasts.

The typical brick veneer Australian dwelling includes a load-resisting timber frame lined internally with plasterboard, a brick veneer that may contain articulations to accommodate limited ground movement, a roof structure comprising timber trusses covered with tiles and a reinforced concrete stiffened raft footing system (Figure 2).

Figure 2.

Typical structure of an Australian brick veneer dwelling including a brick veneer, load-bearing timber frame and stiffened reinforced concrete raft footing.

Airblast is transmitted via walls and roof while ground vibrations are transmitted through the footing system (Eltschlager, 2001). Ground vibrations are predominantly responsible for generating a racking response, whereas airblast generates a flexural response (Figure 3). Studies of structural response to airblast and ground vibrations from nearby surface mines and quarries have demonstrated that low-rise houses are far more likely to sustain damage attributed to racking-related shear and tensile wall strains rather than bending strains generated from mid-wall response (Nicholls et al., 1971; Singh and Roy, 2010; Siskind et al., 1980). While these studies observed racking-related cracking, glass fracture was not. Such studies have demonstrated that walls are much more resilient to face loading of airblast with glass panels being most susceptible to damage, as has been verified by sonic boom studies (Dowding, 1996).

Figure 3.

Structural response to blast vibrations: (a) mid-wall, (b) racking and (c) torsional (Dowding, 1996).

Structural response to transient vibrations typically 2 to 4 s in length is more sensitive to changes in the structure’s natural frequency rather than damping (Siskind et al., 1980). The natural frequency of most residential structures is between 4 and 12 Hz, the percentage of critical damping in the range of 2–10% and racking amplification (γ_p), defined as the ratio of peak eave level displacement divided by peak table displacement, is generally between 1.0 and 5.0 (Aimone-Martin et al., 2003; Dowding and Murray, 1981; Medearis, 1977; Singh and Roy, 2010; Siskind et al., 1980). The above-reported range in performance provides the guidelines for what may be considered representative structural response when designing a suitable specimen for a shaking table study.

Threshold damage is widely accepted as superficial ‘hair-sized’ new cracks or lengthening of old cracks (Dowding, 1996; Northwood et al., 1963; Rainer, 1982; Singh and Roy, 2010; Siskind et al., 1980; Stagg et al., 1984). There have been numerous attempts to identify the threshold of blast- related damage, although no damage has been observed below a ground vibration level of 12.7 mm/s (Aimone-Martin et al., 2003; Gad et al., 2005; Northwood et al., 1963; Singh and Roy, 2010; Stagg et al., 1984). Complicating this matter is the presence of residual stresses generated by settlement, poor maintenance, weather cycles, and prior repair and renovation (Konon and Schuring, 1983). Hence, there may be no absolute minimum vibration-related damage threshold where blasting or environmental or occupant-related vibrations could precipitate a crack (Siskind et al., 1980). None of the reported studies involved the testing of a structure in pristine condition on a shaking table in order to investigate damage thresholds.

Investigations into vibration limits to avoid damage to structures have been investigated since the 1930s, and up until the 1980s soil particle velocity was the preferred indicator of damage potential with 50.8 mm/s being the common vibration limit (Duvall and Fogelson, 1962; Konon and Schuring, 1983; Nicholls et al., 1971). However, blast operators frequently had to reduce this due to damage-related complaints and fear of legal action (Medearis, 1977). A review of blasting guidelines and a large survey of houses damaged by blasting by the USBM culminated in the Report of Investigations 8507 by Siskind et al. (1980) which generated frequency-dependent velocity criteria. To date, there are no universally accepted vibration limits for blasting with some standards being viewed as a ‘guide’ with the velocity limit adopted at the discretion of the blasting operator (AS2187.2, 2006; Dowding, 1996; Siskind et al., 1980). The most conservative vibration limits are based upon human annoyance rather than anticipated threshold damage levels. The vibration limit specified by AS2187.2 (2006) is based upon limits derived by the Australian and New Zealand Environmental Council (ANZEC, 1990) and has been set at 5 mm/s to avoid human annoyance (with provision for 10% exceeding this value, but no greater than 10 mm/s) which recommends a further reduction to 2 mm/s over the long term. However, where there is no risk of complaint, AS2187.2 (2006) relaxes vibration limits to those in British Standard (BS) 7385 (BS 7385-2, 1993) and USBM RI8057 (Siskind et al., 1980). It should be noted that velocity is based on component rather than vector sum for BS 7385 (BS 7385-2, 1993) and AS2187.2 (2006). The least conservative limits relative to the frequency band of surface blasting belong to the USSR standard, the only standard other than AS2187.2 (2006) offering frequency independent limits. Figure 4 summarises Australian and International standard specifications for vibration limits applicable to residential structures. Although vibration standards have evolved, there is a clear need to develop a more rational method to limit structural response to avoid damage rather than relying on prescriptive techniques that seek to avoid damage by limiting soil particle velocity. An integral component of this development is the identification of the damage–drift relationship that may be determined via a shaking table study.

Figure 4.

Vibration limits applicable to modern residential structures in close proximity to ongoing blasting with USSR and Indian standards sourced from Singh and Roy (2010).

Test setup

Construction of test house

A single-room house of typical domestic form was constructed on a shaking table and subjected to simulated ground vibrations recorded from mine and quarry blasts (Figure 5). The biaxial shaking table facility consists of a 2.0 m × 2.0 m and 45-mm-thick steel table with a stiff extension frame bolted to it to increase plan dimensions to 2.57 m × 2.77 m. Two orthogonal 100-kN servo-controlled actuators including accumulators drive the table in the north–south (N-S) and east–west (E-W) directions with a ±100-mm stroke. Hydraulic power is supplied by a three-phase pump delivering 63 L/min. The table was driven in displacement mode from a PC via a micro-console unit.

Figure 5.

Specimen instrumented and setup for testing in the N-S direction.

The frame of the test house was constructed of seasoned MGP12 radiata (AS1684.1, 2006) with nailed joints and bolted to the shaker table. The north and south walls included a 1195-mm wide × 1385-mm high window, while the east and west walls had doorways 900-mm wide × 2070-mm high making the structure symmetrical. The internal walls and ceiling were clad with plasterboard.

The introduction of a lateral bracing system was necessary to preserve the natural frequency of the structure throughout testing to that of a dwelling in good condition. This permitted measurement of the damage–drift relationship with increasing vibration levels that would not have been possible had the lateral stiffness been achieved through lining of the timber frame (i.e. via plasterboard and plywood) alone that may have been compromised at lower levels of vibration.

The variable linear stiffness bracing system incorporated disc springs and was fixed to the frame to maintain a constant lateral stiffness. When the springs were stacked with the same orientation, the stiffness was increased. Conversely, stacking the discs with opposing orientations reduces net stiffness. The stacking sequence of the 36 individual discs in a spring system largely determined the net stiffness of the structure. The stiffness of the bracing system was adjusted once the natural frequency of the structure under racking motion had been identified. A structural frame with degrading stiffness was undesirable as it offers limited opportunity to examine the propagation of cracking damage with drift for a structure in good condition (without degraded stiffness). Two pairs of springs were installed for use in the direction aligned with shaking while flat bar bracing was used on walls orthogonal to the direction of shaking to resist small racking loads. Figure 6 depicts the disc spring setup assembled in the bracing system.

Figure 6.

Spring assembly mounted in bracing system.

The roof consisted of a 1560-kg reinforced concrete slab rigidly connected to the north and south timber frames which represented typical gravity loading on a 2.4-m length of an 8-m-wide house. The shaking table had the capacity to accurately simulate ground motion at very high levels for the full-scale specimen. For the purpose of investigating racking performance, the test specimen correctly represents typical Australian brick veneer construction with a representative roof mass at a 1:1 scale, and hence scaling was not necessary. Due to limitations related to the size of the shaking table it was not possible to test a larger specimen.

The east and west masonry walls were 2270 mm in length and included a central door penetration 2065-mm high by 850-mm wide (Figure 7). The north and south masonry walls were 2390 mm in length and contained a central window penetration 1385-mm high by 1210-mm wide. A professional bricklayer was employed to construct the masonry using extruded clay units of dimensions 230 mm (L) × 76 mm (H) × 110 mm (W), and a 1:1:6 (C:L:S) mortar by volume. The bond strength of the mortar was equal to 0.66 MPa with a standard deviation equal to 0.14 MPa, based upon bond wrench tests performed on 31 specimens carefully extracted from the veneer walls once testing had been completed. General purpose veneer ties spanning a 40-mm cavity connected the veneer to the frame at a maximum spacing equal to 600 mm vertically and horizontally and via a nailed connection. While offering good axial load transfer, offered minimal resistance to lateral loads. At a later stage in the testing programme, it was necessary to improve the connection between the veneer and timber frame when being excited out of plane to limit the flexural response. The connections included stiff steel brackets screwed to the timber frame and epoxy glued to the brick veneer around the perimeter of the walls that were disengaged when the wall was subjected to in-plane excitation. Figure 8 illustrates the retrofit bracket locations on the veneer. Figure 9 depicts an example of a retrofit bracket installed at eave level, while Figure 10 depicts examples of retrofit brackets installed along the wall height. Lintels were 100-mm × 100-mm × 6-mm equal angle sections embedded 150 mm either side of the openings and supported three courses of units above both the windows and the doorways. Returns were not included between adjacent veneer walls, and stoppers were bolted to the table at the end of each masonry wall and inside each doorway to prevent sliding. A damp proof course was not included, and quality of construction was deemed typical of on-site workmanship. A sketch of the cross-section of the wall is included in Figure 11.

Figure 7.

Key dimensions of masonry veneer (millimetres): (a) east and west walls and (b) north and south walls.

Figure 8.

Retrofit bracket locations to secure veneer to frame to prevent flexural failure: (a) east and west walls and (b) north and south walls.

Figure 9.

Example retrofit bracket installed at eave level.

Figure 10.

Example retrofit bracket installed on side of masonry.

Figure 11.

Cross-section view of wall detail.

Instrumentation

The data acquisition system comprised 60 channels including 22 linearly variable differential transducers (LVDTs), 29 accelerometers, eight strain gauges and one table signal input. The transducers were connected to two National Instruments SCXI chassis, and data were logged at 256 Hz on a PC running custom data acquisition software written in LabVIEW (Figure 12). A total of 10 LVDTs measured global displacement while the remaining 12 measured relative deformation in the corners of penetrations. Certain transducers required repositioning when the direction of shaking was changed due to a limited number of transducers. Eight full-bridge strain gauges were fixed to the bracing straps permitting calculation of tension. Photographs showing in-plane measurement and mid-wall response are provided in Figures 13 and 14, respectively. Figures 15 and 16 illustrate transducer configurations for LVDTs and accelerometers in the N-S and E-W directions, respectively.

Figure 12.

Data acquisition system comprising National Instruments SCXI chassis and PC running LabVIEW software.

Figure 13.

Instrumentation attached to rigid frame measuring in-plane movement of veneer and timber frame using displacement transducers and accelerometers.

Figure 14.

Accelerometers mounted on timber frame to measure flexural response.

Figure 15.

Transducer stations during N-S excitation (A: accelerometer; D: LVDT): (a) frame, (b) veneer and (c) east/west veneer.

Figure 16.

Transducer stations during E-W excitation (A: accelerometer; D: LVDT): (a) frame, (b) veneer and (c) north/south veneer.

Photogrammetry surveys were conducted to record specimen geometry and included 1671 retro-reflective targets, concentrated in corner regions of penetrations but more sparsely positioned elsewhere. A comparison between surveys permitted quantification of permanent deformation sustained as a result of vibrations. Crack profiles were measured based upon displacement generated normal and parallel to the bed joint. In addition to the reference survey, 26 surveys were conducted during N-S testing and 15 surveys were conducted during the E-W testing.

Shaking table input signal

The testing programme required the identification of a signature trace to be adopted and simulated in the shaker table at a gradually increasing intensity throughout testing. The identification of an appropriate vibration required a response spectrum analysis of mine and quarry vibrations recorded in-field. Additionally, it was desirable to record changes in the natural frequency of the structure to better understand degradation as testing progressed. Following preliminary tests on the test house which included a swept sine wave and multiple pulse tests, it was decided that the pulse tests were superior for identifying the natural frequency and subsequently used throughout the testing programme to monitor natural frequencies.

The dominant frequency of blast vibrations recorded in residential areas is typically 5–25 Hz for surface coal mines and 10–35 Hz for surface quarry blasting (Dowding, 1996; Siskind et al., 1980). If the level of vibration is great enough, cracking damage may be sustained particularly in non-structural components such as brick veneer, since the natural frequency of a residential structure’s racking response (4–12 Hz) lies within the frequency band of typical surface blast vibrations. In this regard, global drift is the most appropriate indicator of structural damage. Higher mode effects normally associated with an out-of-plane response are less critical from a damage perspective.

The selection of an appropriate blast vibration signature trace was made based upon a response spectrum analysis of the 110 vibration traces. These signals were recorded at distances of 100–2885 m from blasts at two surface mines, namely, Bengalla and Rix’s Creek and six quarries, namely, Bacchus Marsh, Colac, Geelong, Lilydale, Pakenham and Wollert. While the duration of recordings typically lasted between 2 and 4 s, several unusual records were included with durations as low as 1.5 s and as high as 6 s. A response spectrum analysis was conducted on each normalised trace, and a selection of the vibration generating the greatest response spectral displacement (RSD) from each blast location is summarised in Figure 17. By establishing the RSD for each normalised vibration record (normalised to peak component velocity (PCV) = 1 mm/s) and comparing the response for a 10-Hz single degree of freedom structure, the vibration generating the greatest response from each location was short-listed for consideration as the signature trace.

Figure 17.

Vibration records from the two mines and six quarries which generated the greatest RSD from each source.

The Bengalla vibration, having a dominant frequency equal to 8.0 Hz and 4 s in duration, was adopted as the most appropriate input signature based upon an idealised single degree of freedom response and a target natural frequency of 10 Hz. The signature trace contains two distinct components with different frequency content reflecting separation of the compression, shear and Rayleigh waves and was scaled by magnitude only to preserve the frequency profile (Figure 18). In addition to the Bengalla signature trace, the other vibrations shown in Figure 17 were used for comparison at low-intensity vibration levels, while pulse tests were conducted after each test to monitor the structure’s natural frequency.

Figure 18.

Bengalla trace used as primary input, normalised to PCV = 1 mm/s: (a) displacement–time history and (b) power spectrum density.

Overview of test procedure

The sequence of testing of the test house comprised three main phases: (1) preliminary identification of dynamic characteristics (phase I); (2) low-intensity excitation (phase II); and (3) progressive amplification of signature vibration (phase III). Testing commenced when the masonry was at an age of 188 days and completed at an age of 315 days. Phase I incorporated low-intensity pulses and swept sine waves to ascertain preliminary estimates of the natural frequency of the frame and veneer. Phase II simulated a suite of eight mine and quarry vibration traces selected for their damage potential. Peak component velocities at approximately 5 and 25 mm/s were simulated in the N-S and E-W directions to directly assess damage potential to the masonry veneer within levels stipulated by AS2187.2-2006 prior to onset of cracking. Phase III progressively scaled the signature trace to observe crack initiation and propagation and changes in damage state beyond full crack length development, and testing was concluded when the wall was considered to be in danger of collapse.

History of shaking intensity

After completion of phases I and II, testing resumed for phase III by ramping the PCV up from 30 mm/s with regular rotations between the two orthogonal directions of excitation. Secondary traces at lower PCVs were run through the table when the PCV of primary vibrations exceeded 150 mm/s (Figure 19) to investigate the propagation of existing cracks. Upon completion of N-S shaking, a total of 221 N-S tests had been conducted for a table PCV in the range of 1–239 mm/s. Shaking in the E-W direction then recommenced from approximately 130 mm/s and beyond crack initiation at a PCV equal to 303 mm/s; secondary vibrations up to a PCV of 150 mm/s were run after each primary vibration (Figure 20). At the conclusion of testing, 239 blast vibrations were simulated in the E-W direction with the table PCV in the range of 1–383 mm/s. An increase in the scatter of the PCV is observed beyond approximately 200 mm/s that is attributed to the test setup, namely, the limited ability of the hydraulic actuators to accurately reproduce the vibration trace at increasing amplitudes. Table 1 summarises all vibrations simulated during E-W and N-S shaking.

Figure 19.

PCV of N-S vibrations generated during phase III of testing.

Figure 20.

PCV of E-W vibrations generated during phase III of testing.

Table 1.

Summary of vibrations generated in the shaking table throughout testing.

Vibration trace	Type of trace	Dominant frequency (Hz)	East–West		North–South
Vibration trace	Type of trace	Dominant frequency (Hz)	Number of simulations	PCV (mm/s)	Number of simulations	PCV (mm/s)
Bengalla^a	Mine	3.8/8.3	188	383	150	231
Rix’s Creek	Mine	10.3	17	25	16	25
Bacchus Marsh	Quarry	14.0	16	22	15	24
Colac	Quarry	4.3/11.8	19	25	15	24
Geelong	Quarry	14.8	15	23	17	22
Lilydale	Quarry	7.0/22.3	16	24	16	23
Pakenham	Quarry	11.0	15	28	18	31
Wollert	Quarry	10.5	15	25	16	27

PCV: peak component velocity.

Signature trace used for primary excitation.

Identification of natural frequency

Each pulse test comprised three sets of half sine pulses at 5, 10 and 20 Hz which enabled identification of natural frequencies from visual inspection of the power spectral density (PSD) calculated from accelerometer recordings placed at eave level. At the commencement of testing, the natural frequency of the superstructure was estimated to be 7 Hz in the N-S direction and 9.4 Hz in the E-W direction, reducing to 6.5 and 5.9 Hz in the N-S and E-W directions, respectively. The gradual reduction in stiffness was attributed to a softening of the frame and degradation of plasterboard which was exacerbated in the E-W direction by the roof slab being supported on the north and south walls. The natural frequency of the east and west veneers was calculated to be approximately 21 Hz with the final measurement of 14 Hz in the west veneer and 16.3 Hz in the east veneer shortly after first cracking. However, brackets restraining out-of-plane response of the east wall during E-W excitation were not properly disconnected, preventing the estimation of natural frequency of the east wall in its final damage state. The gradual reduction in natural frequency of the east wall suggested progressive micro-cracking while a sudden reduction in natural frequency of the west wall reflected macro-cracking development although visual identification had not been made. Similar tests were conducted for the north and south veneer walls with the north veneer reducing from an initial 22.6 to 18 Hz at the conclusion of testing while the south veneer was initially 21.3 Hz and reduced to 15 Hz reflecting greater damage.

Structural response of test house

The relationship between drift and PCV remained approximately linear for the frame and veneer up until cracking developed in the veneer. Small variations were present which were attributed to degraded tie integrity and variations in pre-tensioning of the bracing. Additional brackets were installed between the veneer and frame to suppress the out-of-plane response to avoid flexural failure. Torsion of the frame was examined but found to be negligible until excessive veneer drift developed due to severe damage.

The racking amplification is an indicator of how responsive the structure is to a given vibration. Prior to the onset of visible cracking, the maximum racking amplification for the east and west masonry walls was 1.3, while the maximum racking amplification in the north and south walls was 1.5. Beyond visible cracking, maximum racking amplifications were up to 5.3 in the veneer, 2.0 in the frame in the N-S direction and 1.9 in the frame in the E-W direction.

A comparison of racking response for different vibration traces was made in the pre-cracked region at a PCV of 25 mm/s. In the N-S direction, the Pakenham vibration generated the greatest response in the frame while the Colac and Bengalla vibrations generated the greatest racking response in the east and west veneers, respectively. During E-W response, both frames experienced the greatest racking due to the Geelong vibration while greatest excitation in the veneers occurred due to the Bengalla vibration.

Cracking damage sustained by the specimen from testing was confined to the veneer with no visible cracking observed in the plasterboard lining. The peak drift measured in the frame throughout testing was 0.42% for a PCV = 383 mm/s vibration although no cracking was observed in the plasterboard joints or around the cornice. The only sign of damage to the plasterboard was loosening of the screw heads around penetrations.

Initiation of cracking

Prior to the onset of visible cracking, peak deformation measured using LVDTs in the corners of penetrations were in the range of 0.13–0.62 mm with an average of 0.3 mm. During N-S excitation, the onset of visible cracking occurred in the east wall (labelled as EWCB in Figure 21(a)) at γ_p = 1/850 caused by a vibration with PCV = 127 mm/s which resulted in full crack propagation. The greatest drift preceding visual cracking was γ_p = 1/960 caused by a PCV = 148 mm/s vibration. Detection of cracking in the west wall occurred during a vibration with PCV equal to 173 mm/s and γ_p = 1/390 with cracks labelled as WWCA1 and WWCB1 being observed (Figure 21(b)). The greatest drift measured prior to cracking in the west wall was γ_p = 1/620 from the preceding excitation with the same PCV = 173 mm/s. Similar tests were completed in the E-W direction with results for the north and south walls shown in Figure 22 and Table 2.

Figure 21.

Identification of location, vibration number and PCV immediately prior to observation of each crack during N-S shaking: (a) east wall and (b) west wall.

Figure 22.

Identification of location, vibration number and PCV immediately prior to observation of each crack during E-W shaking: (a) north wall and (b) south wall.

Table 2.

Greatest PCV and drift recorded prior to and at crack initiation.

Wall	First cracking in wall	Pre-cracking		Crack initiation
Wall	First cracking in wall	PCV (mm/s)	γ_p	PCV (mm/s)	γ_p
East	EWCB	148	1/960	127	1/850
West	WWCA1, WWCB1	173	1/620	173	1/390
North	NWCB1, NWCB2	291	1/700	303	1/870
South	SWCA1, SWCB1, SWCC2	297	1/550	303	1/370

PCV: peak component velocity.

Permanent change in crack width

Permanent changes in crack width were monitored using LVDTs in the corners of penetrations for every test. During N-S excitation, the earliest permanent crack offset measured 0.05 mm and developed across EWCA due to a γ_p = 1/700 and a PCV = 173 mm/s. The earliest crack offset observed in the west wall was 0.06 mm and developed across WWCA1 due to a 189-mm/s vibration generating γ_p = 1/440. During E-W excitation, the first crack offset registering in the north wall was 0.36 mm across the crack labelled NWCA (Figure 22(a)) due to γ_p = 1/550 generated by a PCV = 383 mm/s vibration while the earliest crack offset observed in the south wall was 0.14 mm in the crack labelled SWCA1 (Figure 22(b)) which was due to a PCV = 147 mm/s vibration causing a γ_p = 1/3100. These results revealed a wide range in γ_p (1/3100–1/440) and PCV (147–383 mm/s) required to generate a crack offset with a summary provided in Table 3. Photographs of the cracks at the conclusion of testing are shown in Figure 23 (east wall), Figure 24 (west wall), Figure 25 (north wall) and Figure 26 (south wall).

Table 3.

Summary of the first recorded crack offset in each veneer wall.

Wall	Crack	Crack offset (mm)	PCV (mm/s)	γ_p
East	EWCA	0.05	173	1/700
West	WWCA1	0.06	189	1/440
North	NWCA	0.36	383	1/550
South	SWCA1	0.14	147	1/3100

PCV: peak component velocity.

Figure 23.

Photographs of cracks in east wall at the conclusion of testing: (a) crack EWCA and (b) crack EWCB.

Figure 24.

Photographs of cracks in west wall at the conclusion of testing: (a) crack NWCA, (b) crack NWCB1 and NWCB2 and (c) crack NWCC.

Figure 25.

Photographs of cracks in north wall at the conclusion of testing: (a) crack WWCA1, WWCA2 and WWCA3; (b) crack WWCB1 and WWCB2; and (c) crack WWCC.

Figure 26.

Photographs of cracks in south wall at the conclusion of testing: (a) crack SWCA1, SWCA2 and SWCA3; (b) crack SWCB1, SWCB2 and SWCB3; (c) crack SWCC1 and SWCC2; and (d) crack SWCD.

Permanent damage categories according to AS2870-1996 damage categories

The damage state of each wall after excitation was examined using photogrammetry measurements and compared to damage categories stipulated in AS2870-1996 and summarised in Table 4. The damage category (DC) is based on maximum crack width and is correlated with the peak drift recorded during the excitation immediately prior to the photogrammetry survey with a summary provided in Figure 27 for each wall:

East Wall. The greatest residual damage recorded in the east wall falling within DC 0 was generated by γ_p = 0.04% with peak residual crack width recorded at 0.06 mm. The peak drift due to the next vibration was 0.07% which generated a peak crack width equal to 0.13 mm and within DC1. By interpolation, the DC0–DC1 boundary occurred at approximately γ_p = 0.06%. The remaining excitations for the east wall did not cause residual widths exceeding DC1 although the penultimate excitation during N-S testing generated a peak residual width equal to 0.99 mm and γ_p = 0.12%, just within the DC1 boundary.

West Wall. Damage classification of the west wall was lower due to widespread cracking compared to the discrete cracks that developed in the east wall. The greatest drift recorded while peak residual crack width remained within DC0 was 0.14% with a sole DC1 measurement registered following a drift equal to 0.12% where peak crack width equalled 0.6 mm. Consequently, the increase in damage was not linear due to multiple cracks developing and an approximation to the DC0–DC1 boundary could not be identified. The final excitation caused a γ_p = 0.23% and peak residual crack width equal to 1.38 mm, well within the DC2 category.

North Wall. Residual crack widths in the north wall remained small throughout testing. Drifts up to 0.17% were realised while residual crack widths remained within DC0. The first peak residual crack width (0.12 mm) registering within DC1 occurred due to a vibration generating a γ_p = 0.17%.

South Wall. The south wall experienced the greatest damage classification of all walls. The wall responded to excitation with drifts up to 0.07% without exceeding DC0. When the drift reached a peak value of 0.1%, the greatest permanent damage recorded was equal to 0.1 mm corresponding to the DC0–DC1 boundary. The following survey recorded the final damage state within DC1 with a γ_p = 0.27% and peak crack width equal to 0.42 mm. The next survey revealed a peak residual crack width of 1.3 mm (DC2) at a γ_p = 0.92%; thus, by interpolation the DC1–DC2 boundary falls at 0.72%. Damage did not exceed DC2 for the rest of testing.

Table 4.

Summary of damage categories’ wall cracks as per AS2870 (1996).

Description	Crack width (mm)	Damage category
Hairline cracks	<0.1	0
Fine cracks not needing repair	<1	1
Cracks easily filled	<5	2
Small areas of wall requiring replacement	5–15	3
Extensive repair of large sections of wall	15–25	4

Figure 27.

Peak wall damage in accordance with AS2870-1996 relative to peak drift: (a) east wall, (b) west wall, (c) north wall and (d) south wall.

The DC0–DC1 boundary occurred at a range of drifts from 0.04% to 0.17%, while the DC1–DC2 boundary occurred at drifts in the range of 0.12–0.72%, showing wide variation.

Performance relative to vibration standards

A comparison of the drifts generated by each vibration type was made at limits including the 5-mm/s limit to minimise disturbance to occupants of houses (Figure 4) and the 25-mm/s residential limit to avoid damage to residential structures during phase II. The lowest drift at the threshold of damage for the veneer was generated at a γ_p = 0.10% (1/960; Table 2) which occurred in the east wall at a PCV = 148 mm/s. In contrast, the greatest drift generated during N-S excitation among all the vibration types simulated at PCV = 5 mm/s was 5% of the threshold of damage (west wall generated by a 5.6-mm/s Pakenham vibration). When the intensity of each vibration was increased to the PCV = 25 mm/s limit, the Bengalla vibration generated the greatest drift in the west wall, which was equal to 14% of the damage threshold. During E-W testing at the PCV = 5 mm/s vibration intensity, the greatest generated drift equated to 3% of the damage threshold (north wall generated by a PCV = 4.4 mm/s Pakenham vibration). The Bengalla vibration generated the greatest response close to the 25-mm/s limit with the south veneer generating the greatest drift which was 8% of the damage threshold and caused by a PCV = 24.5 mm/s vibration.

Peak drift to cause cracking measured in the veneer containing door penetrations was nearly twice the drift generated in the veneer containing window penetrations at each of the limits investigated, demonstrating the inappropriateness of PCV as a suitable indicator of damage. Current maximum vibration standards of around 5 mm/s are considered conservative with peak drift demand from these vibrations at around 5% of the damage threshold level. These findings have contributed to a proposal developed by the authors to introduce a capacity spectrum approach as the basis of a more rational method for determining suitable vibration limits rather than relying on the existing deterministic approach (Heath et al., 2015).

Conclusion

A unique shaking table study has been undertaken of the performance of a single room brick veneer structure in pristine condition having characteristics typical of Australian residential dwellings. The test house was subjected to gradually increasing levels of vibration recorded from mine and quarry blasts to examine the in-plane response of the veneer. Eight different vibrations including a designated signature trace were selected for simulation based upon a response spectrum analysis of 102 vibration records. A total of 564 blast vibrations were simulated with table PCV in the range of 1–383 mm/s.

Maximum amplification at the top of the veneer and frame with respect to racking was 1.5 and 2.0, respectively. A comparison of drift generated in the veneer from eight different traces revealed that the signature trace had the greatest damage potential in three out of the four veneer walls, further demonstrating the importance of the vibration frequency content and the inappropriateness of a single PCV-based vibration limit. The lowest drift recorded prior to the initiation of first visible cracking was 1/960 due to a 148-mm/s vibration which had been preceded by 347 vibration signals. The peak crack width measurement recorded with an LVDT prior to the onset of visible cracking was approximately 0.3 mm while the lowest drift observed to cause a permanent change in crack width during N-S shaking was 1/1880 and 1/3100 during E-W shaking. Photogrammetry measurements revealed that the lowest drift at the DC0–DC1 boundary was 0.04% for an isolated crack but as high as 0.17% when multiple cracks were present while the lowest drift causing damage at the DC1–DC2 boundary was 0.12%.

At the 5-mm/s vibration limit imposed by AS2187.2-2006, the greatest observed drift was 5% of the veneer damage threshold whereas at the 25-mm/s limit the greatest drift recorded was approximately 14% of the veneer damage threshold. This comparison demonstrated the conservatism of existing velocity-based vibration limits (PCV) intended to avoid damage to structures and the need for a more rational approach that considers the response spectrum of the vibration signal to deduce damage potential.

Footnotes

Acknowledgements

The authors wish to acknowledge the generous contributions including supply of vibration records made by the project industry partner, Terrock Consulting Engineers, and Australian Research Council (ARC) Linkage Grant No. LP0211407. This research project was undertaken at The University of Melbourne.

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship and/or publication of this article.

Funding

The author(s) received no financial support for the research, authorship, and/or publication of this article.

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