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回采巷道复合顶板破坏机理与支护技术摘要：随着开采深度的增加，顶板条件渐渐由坚硬顶板向复合型顶板变化，以往在坚硬顶板条件下采用的单纯锚栓支护方式已显得不适应，曾发生过多起冒顶事故，安全生产没有保证。为此，本文分析复合顶板下的回采巷道围岩的应力分布状态，变形特征；研究复合顶板在各种锚杆支护方案下的变形、破坏机理；分析复合顶板回采巷道的锚杆支护机理以及合理的控制技术与措施。关键词：复合顶板；联合支护；高预紧（应）力锚杆1 引言我国国有大中型煤矿煤层巷道应用越来越多，回采巷道布置在煤系岩层中，受力情况较复杂，受采动影响较大，是煤矿巷道围岩控制的难点，当围岩强度及其稳定性较差时，部分回采巷道会发生大范围的支护失效和顶板冒落现象，不仅给井下作业人员人身安全带来威胁，而且导致矿上企业巨大的经济损失。复合顶板巷道在煤矿中有广泛的分布，复合顶板的支护是国内外巷道支护的难题之一。由于煤系地层为层状沉积岩层，因而复合顶板是一种较常见的岩层赋存形态。复合顶板极软煤层巷道围岩为差异性很大的非均质层状赋存，在围岩应力作用下表现为顶板极易离层、冒落，难以形成承载结构，强烈的两帮移近、片帮及整体下沉，导致复合顶板下沉而离层破坏，顶板、两帮变形相互作用，形成恶性循环。该类回采巷道采用传统支护方式如工字钢支架、U型钢可缩支架支护时不仅在掘进影响期间围岩变形剧烈，而且在掘后较长时间内难以趋于稳定、变形量大，在服务期间需多次返修，巷道维护极为困难。随着采深的增加，这类巷道的维护问题将越来越突出，因此，寻求一种有效的支护方法具有重要的现实意义。在此，对回采巷道复合顶板破坏机理与支护技术的研究与探讨，为煤矿企业在复杂地质条件下的高产、高效安全生产起到了积极促进作用。2 复合顶板巷道变形破坏机理2.1 复合顶板的概念及特点所谓复合顶板，其实质就是离层型顶板。即煤层上面有总厚0.33m易与上部岩层离层的顶板。这种顶板岩层的岩性主要是页岩、砂质页岩和薄层细砂岩，同时往往夹有薄层煤层、炭质页岩、泥质岩、植物化石碎屑和镜煤条带等。复合顶板由下软上硬岩层构成。下部软岩层可能是一个整层，也可能是由几个岩层组成的岩层组。这里的软岩层与硬岩层只是一个形象的说法，实际上是指：采动后下部岩层或因岩石强度低，或因分层薄，其挠度比上部岩层大，向下弯曲的多，而上、下部岩层间又没有多大的粘接力，因此下部岩层与上部岩层形成离层；从外表看，似乎下部岩层较软，上部岩层较硬。典型的复合顶板有“软”、“弱”、“薄”三个特点：1）煤层顶板由下软上硬的不同岩性的岩层所组成；2）软、硬岩层间夹有煤线或薄层软弱岩层，易离合，差异较大；3）下部软岩层的厚度一般较薄，通常情况下不小于0.5 m，且不大于3.0 m，冒落后不能充满采空区。除此之外，当伪顶厚度超过0.5 m时（超过0.5 m厚的随采随冒的软弱岩层不能叫伪顶），现场往往采用“托伪顶”开采。在这种情况下，煤层的顶板就是复合顶板。应用留煤皮方法采煤时，如果煤皮厚度在0.53.0 m之间，而煤皮与顶板又易分离或煤层有伪顶，则采煤工作面也处在复合顶板之下。厚煤层应用倾斜分层下行垮落开采时，第二分层以及以下分层可能处在再生顶板之下。如果再生顶板的厚度在0.53.0 m，其上为较硬岩层或咬合住的断裂岩块，再生顶板与它又没有多大的粘接力，则在回采第二分层及以下分层时，该分层也处在再生的复合顶板之下。复合顶板下部岩层与上部岩层在力学特性方面的主要差别是他们之间的厚度（包括互层中的分层厚度）、变形模量、泊松比及内摩擦角、黏聚力等方面相差较大。2.2 复合顶板的岩性及结构复合顶板的一半特性是下软上硬，容易发生离层。在软硬顶板之间存在有薄弱面和光滑面，软硬岩层之间的粘接力很小，极易分离。当采掘工作进行后，松软顶板的下沉与坚硬顶板的下沉不同步，因而两层见发生了离层，尤其在锚杆预应力很小时，离层现象更加突出，其巷道顶板岩层的层间离层如图1所示。如图1的情况是普遍常见的结构。还有一些特殊的形成状态，如冲刷造成的顶板由一种岩性，分成两种不同的岩性。两层的层理分明、中间夹有薄弱煤岩或光滑面，使非复合型顶板，变为复合型顶板。如图2所示。在实际生产工作中，因地质情况调查不清楚，而未能采区相应措施，出现大型的冒顶事故。在一些倾角较小，而水平断距又很大的各类断层的作用下，也会出现不同类型的复合型顶板。如图3所示。此外，在近距离煤层群中，上一层煤的采厚由可采变为不可采，且该层的直接顶比较硬，下一层的顶板比较软，使得下层的顶板与上层煤层共同构成了复合型顶板。如图4所示。图1 巷道顶板岩层的层间离层图2 由非复合型顶板变为复合型顶板图3 由不同类型断层形成的复合型顶板图4 下煤层的顶板与煤层构成为复合型顶板2.3 复合型顶板的矿压显现特点巷道开挖后，岩层抗水平应力的截面减少，在水平应力作用下煤层沿水平层理面向巷道挤入，致使巷道帮顶受水平应力作用而破坏；围岩中节理构造面的存在对围岩的承载能力及其稳定性影响很大，尤其是节理面与最大主应力方向斜交时，岩体最容易沿节理弱面破坏而失稳；巷道开挖后，围岩受力状态由三轴应力变为单轴压应力状态，由于岩石单轴抗压强度低，致使围岩产生塑性破坏或沿节理弱面破坏、随着锚固岩体的变形、离层和弯曲，巷道中部的锚杆始终受力，若锚杆的长度、刚性越大，会使之受力越大，锚杆受力即可达到强度极限而破坏，则岩层发生破坏。复合型顶板矿压显现特点可总结如下：1、顶板变形量大。由于复合顶板由松软岩层组成，岩层的弹性模量低，在长期的地层压力作用下，岩层内积聚了大量的弹性变形能。巷道开掘后，巷道表面围岩应力解除，岩层内积聚的弹性变形能释放，从而产生变形。岩层愈软，变形量愈大；地层压力愈大，变形量愈大。复合顶板的变形量可达200400mm。2、顶板变形时间长。巷道掘出后，巷道表层或浅层的弹性变形能会迅速释放，表现为巷道开掘初期变形速度快。巷道围岩弹性变形能的释放，是由浅而深逐渐发展的，围岩的变形会由快变慢、长期发展，但不会像坚硬岩层巷道那样停止变形，围岩的变形范围会越来越大。岩层愈软，巷道变形延续的时间愈长，围岩变形的范围愈大。复合顶板的变形范围可达58m以上。3、顶板压力大。必须对顶板岩层内大量的弹性变形能的释放予以控制。过度的释放会导致顶板岩层的离层和破坏，而释放不足则会极大地增加顶板的支护压力。在不采取任何措施的情况下，锚索支护属于刚性支护。锚索支护系统的刚性系数大，只允许顶板有很小的变形。当顶板岩层内的弹性变形能得不到适当的释放，顶板压力会越来越大，直至锚索被破坏。这种大变形的顶板矿压显现特性与锚索的支护特性显然不相协调。2.4 复合型顶板的变形破坏特征复合顶板具有双重力学属性，在外载作用下，随其自身状态不同呈现两种变形特征，即相对完整状态与类似散体软岩状态。2.4.1 相对完整状态的复合顶板变形移动规律所谓相对完整状态，是指煤层顶板的变形移动发生在顶板组成分层之间，而顶板岩层分层自身则保持完整的稳定状态。此时，顶板岩层是由层状岩体组成的复合顶板。层状岩体属于典型的横观各向同性体，分层层面是其主要弱面。层状岩体的分层层面属闭合节理，根据剪切试验，在顺层理方向剪切下，闭合节理表现有峰值和残值，即层状岩体的分层层面有一定的剪切强度，且在分层之间发生滑移错动后，同样保持微小的残余强度。部分煤系岩层层面的力学参数见表1。表1 部分煤系岩层层面的力学参数层面两侧岩石切向刚度系数KS/MPa切向刚度系数Kn/MPa黏结系数C/MPa内摩擦角/()石灰岩与煤层24.59810.235煤层与页岩交互层14.75880.125页岩互层与砂岩14.75880.125此时顶板层状岩体的强度可表示为： = tanr+Cr （2-1）式中 Cr层面处的黏结系数； r层面处的内摩擦角。当顶板层状岩体在外载作用下，层面处的法向应力与层面的剪切应力满足式2-1时，顶板将沿分层层面发生剪切破坏，相邻的分层沿层面相对滑移。因此作为回采巷道顶板，在无支护条件下，其移动破坏可分为一下几个阶段。1）沿层面的分层相对滑移阶段巷道开掘后，岩体内原有平衡状态被打破，巷道顶板最初以整体岩体承担上覆岩层的重力载荷。在原岩应力及掘巷集中应力的作用下，巷道顶板岩体发生变形，同时内力增加，当层面处的内应力、满足式2-1时，岩体发生剪切破坏，产生层间滑动，由于分层层间滑动，使层面的力学参数进一步减小，整体强度降低。层面滑动范围与掘巷前的应力场及开掘巷道的断面形状有关。当巷道跨度L与顶板分层厚度h之比较小时，即在分层厚度相对较大的情况下，只会在巷道两侧的顶板支撑区产生层间滑动，在岩体中表现为与层理近乎垂直的张开裂隙，如图5所示。当巷道跨度L与顶板分层厚度h之比较大时，即在分层厚度相对较小的情况下，滑动区可能波及整个跨度内的直接顶板岩层，此时下位岩层在中心处与其上位岩层有脱开而离层的趋势。2）分层层间离层阶段顶板分层沿层面的相对滑动，将改变其自身的应力状态。若此时变形后的围岩体与上覆岩层形成的载荷仍不能取得力学平衡，则直接位于巷道上方岩层的变形将继续发展，纵向变形（下沉）增大。由于下位岩层的下沉速度大于上位岩层的下沉速度，变形发展到一定程度，即形成巷道顶板岩层的层间离层，离层范围及离层岩层数也逐渐增加。3）顶板岩层逐层弯曲折断阶段随着离层范围的增加，与上位岩层离层的岩层必须单独承担自身的全部重量。当岩层悬空跨度达到极限跨距，岩层内最大拉应力达到岩石的抗拉强度时，岩层将会发生弯曲张拉破坏。断裂后的岩块若不能形成横向作用力下的铰接平衡，便从巷道顶板垮落下来。离层岩层的这一活动过程可以按均布载荷作用下的固定梁或简支梁变形破坏过程的分析加以说明。由于离层区域自下而上减小，则巷道顶板岩层各分层可能出现的悬空跨距也自下而上减小。当顶板下位岩层强度较低、分层厚度较小时，顶板岩层从裸露在外的第一层开始，向上逐层垮落，如图6所示。2.4.2 类似散体软岩状态的复合顶板变形移动规律煤层顶板整体性破坏后（如出现局部冒落），顶板破坏急剧发展呈现类似散体的破碎状态，其变形移动过程可分为如下几个阶段。1）水平移动并形成大量超薄子分层顶板冒落空间形成后，空间四周各分层岩层向冒落空间水平移动。各分层岩层在水平移动过程中，其内部含超薄复合分层产生滑移错动，形成了数目众多、厚度极小（小于10 mm）的子分层，使薄层复合顶板下位分层成为具有工程意义的软岩。2）下位软岩持续水平移动冒落空间周围大范围的下位软岩，向冒落空间挤出与弯曲，形成较大的破碎变形压力。3）薄层子分层随着顶板分层水平移动，因其自身强度低而形成类似散体的破坏。图5 沿层面的滑动图6 巷道层状岩层顶板冒落形状2.4.3 无支护条件下巷道顶底板层状岩体垮落高度的估算根据A鲍里索夫的建议，开掘在层状岩层的巷道，在长期无支护情况下，其顶板岩层的垮落高度H为H=(L(0.04khk/nk)1/2)tan/2 （2-2）式中 压缩蠕变系数,取0.50.7； hk直接位于冒落空洞之上的第k层岩层分层厚度； k、k第k层岩层的顺层抗拉强度和岩石重力密度； 垮落梯形的边界腰线与层面间的夹角，参照软弱岩层的破断 角，其值可取6080； n承载能力安全系数，可取n=4。其中，可按下述步骤确定冒落空间上方裸露分层的层位k：1）确定巷道顶板各分层因其下分层垮落形成的悬空跨距Li。对巷道顶板岩层各分层自下而上编号，即巷道直接顶板由第i=1，2，3，k，n层岩层组成，第i分层层厚为hi，巷道顶板各分层因其下分层垮落形成的悬空垮落步距Li为Li= Li12h i1 cot （2-3）显然，L1=L，即第一分层的悬空跨距等于巷道断面上方的横向尺寸。2）计算巷道顶板岩层各分层的极限跨距Lji。设巷道顶板岩层各分层的顺层抗拉强度及重力密度分别为i、i（i=1，2，3，k，n），则每一分层岩层在长期无支护情况下，保持自身平衡而不冒落的极限跨距为Lji=(0.04khk/nk)1/2 （2-4）3）求k值。依次比较Li与Lji（i=1，2，3，k，n），若LjiLi，则该分层将发生垮落，直到某一分层i时LjiLi，则该分层将不冒落，其所在层位序号i即是所求的k值，即该分层是冒落空间上方悬露的顶板分层。2.5 复合顶板变形破坏机理最大水平主应力理论认为，围岩层状特征比较突出的回采巷道开挖后引起应力重新分布时，垂直应力向两帮转移，水平应力向顶底板转移，从而引起水平、垂直应力的相互转换叠加，引起应力集中，引起顶板离层、底板鼓起及两帮外移。图7为开掘巷道后岩体的受力情况，N为垂直应力，F为原水平应力。巷道开挖后，巷道的顶板变为一双支梁，开挖段的压力向两帮转移，造成两帮压力增大，引起两帮的水平应力F与原F叠加造成水平应力集中，导致两帮煤体的向里位移。巷道的水平应力，由于顶、底、煤体的层理作用，煤体向里位移，层面间就产生剪切力(摩擦力)F，剪切力随着顶底板的加厚而减小，剪切力F与原F叠加，从而造成顶、底板水平应力集中，特别是与煤层接触的一层顶、底板水平应力最大，从而引起顶、底板弯曲，即顶板下沉，底鼓。严重者造成巷道破坏。复合顶板巷道的变形主要是顶板的挠曲下沉与底板的鼓起，巷道两帮的位移量相比顶底板的位移量较小。巷道顶底板表面都有较大的位移量，其中巷道顶板中央的位移量最大，方向垂直向下。巷道围岩的位移由巷道的顶底板表面至顶底板的深部逐渐减小。图7 巷道围岩受力情况由于巷道可以简化为为平面应变模型，用FLAC数值模拟软件中的弹性模型模拟了翟镇煤矿3416面轨巷的应力分布情况，如图8所示。其应力在平面内存在水平应力、垂直应力及剪应力。因而可以将巷道内各单元的应力转化为最大主应力(拉应力)与最小主应力(压应力)。由弹性力学的知识可知，最大主应力与最小主应力的方向是相互垂直的，如图中所示。在巷道的帮角处，受较大的剪应力作用。巷道围岩中的最大拉应力为2.1 MPa，最大压应力为25.3 MPa。在巷道顶底板中出现拉主应力的绝对值大于压主应力的情况，并且部分单元的拉主应力的方向是水平的，根据抗拉强度远小于抗压强度，岩体中一旦出现拉应力区，该区域往往会最先破坏的原则，说明巷道顶底板中部分区域的破坏将是拉应力破坏。巷道帮角处出现了较大的应力集中，观察到巷道帮角处的主应力的方向为x轴与y轴成45或135的方向上。根据最大与最小的切应力其值为，（12）/2发生在与x轴及y轴成45的斜面上，可以断定在巷道的帮角处的最大主应力即为剪应力，所以巷道的帮角处容易发生剪切破坏。如图9所示，巷道顶底板中出现了范围较大的卸压区(应力降低区)，巷道顶底板表面的垂直应力降到了上覆岩层应力的1/4左右，由顶板表面向上及由底板向下垂直应力降低程度都逐渐减小，直至恢复到上覆岩层应力水平。巷道两帮为应力增高区，应力集中程度最大达到2.25倍，在巷道两帮大约1/2巷宽，即1.7 m左右范围内，应力集中系数仍达到了1.25。侧压系数人对围岩的应力分布影响较大，16 mm 锚杆长度1.61.8 m 间排距0. 81.2 m 设计锚固力6480 kNII稳 定 顶板较完整：单纯锚杆支护 顶板较破碎：锚杆+网支护 端锚： 杆体直径1618 mm 锚杆长度1.62.0 m 间排距0.81.0 m 设计锚固力6480kNIII中等稳定 顶板较完整：锚杆+钢筋梁或桁架 顶板较破碎： 锚杆+W钢带(或钢筋梁)+网，或增加锚索； 桁架+网，或增加锚索 端锚： 杆体直径1618 mm 锚杆长度1.62.0 m 间排距0.81.0 m 设计锚固力6480kN 全长锚固： 杆体直径1822 mm 锚杆长度1.82.4 m 间排距0. 61.0 mVI不稳定 锚杆+W钢带+网，或增加锚索 桁架+网，或增加锚索 全长锚固： 杆体直径1822 mm 锚杆长度1.82.4 m 间排距0. 61.0 mV极不稳定 顶板较完整： 锚杆+金属可缩支架，或增加锚索 顶板较破碎： 锚杆+网+金属可缩支架，或增加锚索 底鼓严重： 锚杆+环形可缩支架 全长锚固： 杆体直径1824 mm 锚杆长度2.02.6 m 间排距0. 61.0 m5）帮锚杆长度Le的确定帮锚杆长度确定基础为巷道破坏最大深度、锚杆外露长度和锚固段长度之和，即：Le=Lc+Ld （4-5）式中 Lc巷道破坏最大深度；Ld锚固段长度与外露长度之和。6）锚杆直径D的确定根据材料力学计算出锚杆直径:D=(4.4P/Jb)1/2 （4-6）式中 P取0.1 MPa;Jb螺纹钢锚杆屈服点。7）锚索长度Ls的确定根据悬吊理论及承压拱理论计算，即：Ls=L1+L2+L3 （4-7）式中 L1锚的外露长度； L2锚索的外露长度，对顶板而言，L2为冒落带(或离层)的高度h； L3锚索的锚固长度。8）网带的选择目前，锚喷支护中所采用的“网”，按其材料不同，可分为金属网(包括铁网和钢筋网)和塑料网。其中，金属网按其网格形状分为方格网和菱形网，按其制作工艺分为焊接网和纺织网。焊接网整体性好、强度高，但刚性较大、加工成本高，在我国很少使用。目前使用最多的还是方格编织网和菱形编织网。近年来，国内外广泛使用钢带作为锚杆的联系构件。在一般锚杆支护的基础上，若干根锚杆共用一条钢带作为辅助托板，使它们互相联系，以形成整体结构，增强对围岩的控制能力。目前，常使用的钢带有：钢板钢带、W型钢带和圆钢钢带。5 主要结论本文通过对复合顶板巷道的相关文献资料的归纳总结、理论分析，取得了以下主要结论：1）复合顶板顶板变形量大，变形时间长，且顶板压力大；2）复合顶板的变形破坏主要是因为支护不力导致顶板初期离层量和下沉量较大，顶板的稳定性恶化，加之复合顶板各岩层节理裂隙发育，且各岩层间黏结力较弱甚至无黏结力，顶板迅速离层、下沉,并逐步向顶板深部扩展，最终导致复合顶板的破坏；3）复合顶板煤层巷道顶板和两帮变形相互影响、相互作用，其中任意一方失稳，必然导致巷道围岩整体失稳，两者支护控制同等重要；4）为充分发挥锚杆支护的主动支护作用，要及时支护并使用专门的风动扭矩放大器加强锚杆支护的预紧力，提高锚杆预紧力是防止顶板离层的重要手段；5）采用锚带网索联合支护方式，可以有效地提高围岩的整体性，防止锚杆周围岩石的松动对锚杆支护作用的影响，而且通过钢带的作用使支护系统具有很高的安全性，防止突发性冒顶事故。锚带网索联合支护方式，优化了支护参数，降低了巷道的维护费用，取得了较好的技术及经济效益；6）复合顶板煤巷锚带网索联合支护要按照科学设计、现场实施、矿压观测、修改完善的程序来进行，在保证支护系统安全可靠的基础上，实现技术效果与经济效益优化。英文原文Triggering of Seismicity Remote from Active Mining ExcavationsBy S.D.McKinnonDepartment of Mining Engineering, Queens University, Kingston,Ontario, CanadaSummaryObservations of seismicity and ground control problems in the Sudbury mining camp have shown that late-stage (young) sub-vertical strike-slip faults are sensitive to small mining-induced stress changes. The strength-limited nature of stress measurements made in the region indicates that these structures are in a state of marginal stability. Numerical continuum models are developed to analyze the behavior of such structures. In the models, shear strain localizations (faults) evolve such that there is close interaction between the fault system, stresses, and boundary deformation. Fault slip activity in these systems is naturally sporadic and reproduces the commonly observed Gutenberg-Richter magnitude frequency relation. It is shown that a relatively minor disturbance to such a system can trigger significant seismicity remote from the source of the disturbance, a behavior which cannot be explained by conventional numerical stress analysis methodologies. The initially uniform orientation of the stress field in these systems evolves with increasing disorder, which explains much of the scatter commonly observed in data sets of stress measurements. Based on these results, implications for stress measurement programs and numerical stability analysis of faults in mines are discussed.Keywords: Triggering, microseismicity, stress analysis, rockbursts, fault stability.1 IntroductionThe majority of seismic events around deep hard rock mines occur close to excavation boundaries. These events are related to mining-induced stress changes leading to damage involving fracturing of intact rock or slip along pre-existing discontinuities. Extraction layouts leading to highly stressed structures such as pillars and abutments are particularly prone to induced seismicity. With appropriate calibration of rock mass strength, numerical stress analysis can be used to estimate the extent of fracturing and therefore the extent of near-excavation seismicity (Beck et al., 1997; Potvin and Hudyma, 2001; Beck and Brady, 2002). Characteristics of near-excavation seismicity include swarms of events triggered by production blasts (which cause a rapid change in the stress field), followed by a gradual decay in event frequency to background levels over a period of hours or days. The regularity in the frequency and location of near-excavation events makes this type of seismicity a manageable mining problem.A certain amount of seismicity also occurs further away from mining excavations and appears to be uncorrelated in time and space with mining activities. Events have been recorded hundreds of meters away from active mining. On the basis of source locations, it has long been recognized that these events are the result of slip on preexisting structures such as faults, dykes or contacts (Smith et al., 1974; Gay et al., 1984). Although the number of events close to mining excavation boundaries vastly exceeds those further away, the latter are of great concern to mining since they tend to be of larger magnitude, increasing the risk of rockburst damage. Since neither their location nor magnitude can be predicted in advance, mines must consequently make more extensive use of heavier ground support to control potential rockburst damage than would be required if events were only located close to active mining excavations.Due to the complex geometry and geological environment of most mines and the availability of numerous commercially supported codes, numerical stress analysis is the tool of choice for the majority of mining stability analysis. However, only a limited amount of success has been obtained in using numerical modelling to understand seismic events on faults and other geological structures. In particular, numerical stress analysis has not been able to explain the occurrence of seismic events remote from mining. Due to the widespread use of numerical stress analysis in the mining industry, it would be desirable to develop a methodology that would enable modelling to be used to explain both types of seismic events. The objective of this paper is to present an investigation into the cause of seismic events remote from mining and the implications for applying numerical stress analysis to the problem. As would be expected, this type of seismicity is strongly influenced by the geological environment in which the mine is located. Motivation for the approach to modelling mining-remote seismicity is taken from stress measurements and known geological controls of seismicity in the Sudbury mining camp, which is a region of intensive mining activity in Ontario, Canada.2 Structural Geology and State of Stress in the Sudbury StructureBased on a variety of stress indicators, the stress field in mid-continental North America is characterized by a horizontal major principal stress with an orientation approximately ENE (Zoback and Zoback, 1980; Zoback, 1992b). Arjang (1991) found a similar orientation for the major principal stress based on overcoring stress measurements made in mines in the Eastern Canadian Shield. Using a more extensive database of stress measurements, Arjang and Herget (1997) noted that although there are reasonable consistencies in factors such as the orientation of the major principal stress and the ratios of horizontal to vertical stresses, there was a very large scatter in the trend of the sub-horizontal major and intermediate principal stresses.A significant portion of the stress measurements available in Eastern Canada were made in mines located in the Sudbury Structure. Poles of principal stresses measured in these mines are shown in the lower hemisphere projections of Fig. 1(a). Stress measurement data used in Fig. 1(a) were taken from Arjang (1998), plus unpublished data from INCO Ltd. (Galbraith, 2002). The minor principal stress is typically subvertical, which shows the least amount of scatter in orientation. The major and intermediate principal stresses generally dip sub-horizontally and have significantly higher variation in orientation compared to the minor principal stress. This is shown more clearly in Fig. 1(b), in which the data is restricted to those measurements with minor principal stress within 20of vertical.When displayed in strength space, Fig. 2, the stress data is strongly suggestive of a linear envelope (for the regression line, R2= 0.85). There is also a reasonably high degree of correlation (R2 = 0:70) between the magnitude of the two sub-horizontal principal stresses, Fig. 3. The implication is that the state of stress is strength limited, similar to the strength envelope formed by laboratory compression tests. If stresses were not in a state of limiting equilibrium, they would not necessarily define a failure envelope. Since large scale structural discontinuities form the weakest link in the fabric of the rock mass, this correlation reflects a limiting equilibrium relationship between the state of stress and the strength of some of the faults. Simple Andersonian ranking (Anderson, 1951) of the principal stresses (the commonly used geological method of determining style of faulting that would be expected for various orientations of principal stresses) suggests that thrust faults would be the most likely candidates for marginal stability.Fig.1.Lower hemisphere projection of principal stresses, a Sudbury region stress measurements, b subset of measurements with minor principal stress within 20 of verticalThis leads to an apparent discrepancy between the type of faulting most favored in an Andersonian sense, i.e. stress measurements indicating a thrust faulting regime, and the apparent marginal stability of the late-stage strike-slip faults and fractures. Cochrane (1989) addressed this problem by noting that hydrofracture stress measurements made in the Paleozoic cover of the mid-continental stress province, Haimson and Doe (1983), indicated that at depth the vertical stress is greater than the minimum principal horizontal stress, suggesting that strike slip faults originated at depth and that subsequent erosion has reduced the vertical stress to the minor principal stress. The relationship between structures and stress conditions in Eastern Canada has also been complicated by repeated episodes of glaciation post-dating the formation of most structures, which suggests that the neotectonic stress field is not completely consistent with the structural fabric. However, based on the strength-limited nature of the stress measurement data and the sensitivity of the late-stage faults and fractures to mining induced stress changes, it would appear that these structures, as opposed to any other system, are in a state of marginal stability with the regional tectonic stresses. Fig.2.Major and minor principal stress Fig.3.Sub-horizontal principal stressmeasurement data, Sudbury region. measurement data, Sudbury region.This geological setting provided the framework within which numerical stress analysis was used to investigate the occurrence of seismicity remote from mining excavations. The observations of Cochrane (1989) show that geological structures are not equally sensitive to mining-induced stress changes, and that in the Sudbury Structure the most important structures to account for are the late-stage sub-vertical strike-slip faults. While the importance of other faults in the system is unknown at this time, the focus on the behaviour of the sub-vertical faults enables the numerical representation to be simplified to two-dimensions, as will be described in more detail below.3 Review of Numerical Stress Analysis Applied to Mining-induced SeismicityThe standard approach to assessing slip on faults due to mining is well documented (CAMIRO, 1997). While the particular choice of numerical stress analysis code and method (continuum, discontinuum, boundary element etc.) may vary, the principles involved in assessing fault stability remain the same. The analysis sequence can be described by the following steps: the pre-mining state of stress is initialized, the new state of stress due to the mining geometry is computed, normal and shear stresses on the fault surfaces are resolved and compared to the assigned strength, and the potential for slip evaluated. Once the slip area has been computed, an estimate of maximum seismic moment and event magnitude can be made (CAMIRO, 1997).If elastic continuum codes are used, faults are generally not incorporated into the models and therefore have no effect on the stress field. Slip on these virtual faults is generally assessed using the Excess Shear Stress (ESS) criterion (Ryder, 1987). Examples of this approach can be found in the work of Board (1996); Urbancic and Trifu (1998); Hanekom (2001); Beck and Brady (2002). Using discontinuum codes, faults may be incorporated explicitly into models. However, since fault surface displacement is typically equilibrated with the initial stress field, no history of deformation is accounted for and there is generally little influence of faults on the pre-mining state of stress. Examples of this approach can be found in CAMIRO (1997).A more sophisticated approach has recently been proposed by Wiles et al. (2001). Using a boundary element method of analysis, observations of seismicity are used to prescribe a displacement on fault segments represented by displacement discontinuity elements. In this manner, a pre-strain is incorporated into seismically active fault segments, resulting in a local distortion of the stress field. This approach is promising, but is limited by the requirement that fault displacements can only be accumulated in the models based on estimates of slip computed from observations of seismicity. Significant seismic events resulting in rockburst damage can occur on faults with no prior history of activity. Also, this piecewise modification of the model is not well suited to the mine design process in which an assessment of risk is often required prior to extensive mining.A common limitation of these methods, and in particular the issue of explainin seismic events remote from mining, is related to fundamental assumptions about fault strength, the state of stress, and the degree of stability of fault segments. The initiation of a homogeneous state of stress prior to mining and the superposition of the computed shear and normal stresses on fault surfaces (virtual or explicit) does not capture the critical interaction between the structures and the local stress field.4 Numerical Modelling of the Evolution of Fault SystemsIn order to numerically model a system of faults in a state of marginal stability, a number of approaches can be taken. It is tempting to make use of discontinuum codes, in which faults are modelled explicitly. However, there are problems related to the evolution of such systems, in particular the method of establishing compatibility of fault orientation, rock mass strength, and stress field orientation. The approach selected was to use a continuum representation starting from an initially intact material in which faults evolved as strain localizations in response to far-field (boundary) deformation. This ensured that faults would be correctly oriented relative to the evolving stress field, and that the stress field varied locally in response to changes in fault strength as a result of constitutive behavior such as slip weakening. The two dimensional finite difference code FLAC (Itasca Consulting Group Inc., 2002) was selected to carry out the stress analysis as it is well suited to modelling fractures and faulting in rock (Cundall, 1989; Hobbs and Ord, 1989; Cundall, 1990; McKinnon and Garrido de la Barra, 1998).To enable faults to form in an initially elastic material, a strain softening constitutive model was used in conjunction with a large strain finite difference formulation for computing gridpoint displacements. A strain softening constitutive model was not essential, but sharpened the shear bands. The memory of where faults occurred in the material was provided by permanent deformation of the grid due to the large strain formulation. The two-dimensional nature of the code resulted in the formation of linear strain localizations, corresponding to strike-slip faults, which are the focus of the investigation.4.1 Model ConstructionA common problem with grid-based continuum models is the effect of grid characteristics on the formation of shear bands. Attempts to minimize these effects were made by various model construction strategies:l Circular boundary models were used in order to minimize formation of shear bands at model corners due to stress concentrations.l Relatively large numbers of zones were used in the models to facilitate the formation of shear bands, which typically span two or three zones. Models consisted of approximately 7000 zones.l The initially square gridpoint geometry was randomized to eliminate regular channels of constant size zones. This reduced the tendency of shear bands to follow the alignment of grid axes.l The Youngs modulus of each zone was varied randomly about a mean value. This effect represented the heterogeneous nature of material properties. Randomizing locations of displacement-induced stress concentration reduced grid triggering of shear band formation.l An annulus of softer material approximately 5 zones wide surrounded the central model material. This material represented a resilient compromise between rigid displacement controlled and soft stress controlled boundary conditions, and was thought to be a geologically more realistic alternative to conventional boundary conditions.A simplified example of the resulting model grid is shown in Fig. 4.Similar numerical experiments were also carried out with substantially larger models than those described here. In those models, the density of shear bands was higher and fault growth (generally initiated in the interior of the models) could be tracked in more detail. However, the behaviour of these larger models was essentially the same as those described, leading to the same overall conclusions.Fig.4.Example of gridpoint randomization used to reduce grid effects on shear band formation4.2 Material Properties, Initial and Boundary ConditionsElastic properties for the models were chosen to represent a typical hard rock found in Ontario mines. These properties are listed in Table 1. It was found, however, that the main conclusions drawn from the models remained the same for other values of hard rock parameters.Table 1. Material propertiesPropertyValueUniaxial compressive strength c（intact）150MPaRMR69Youngs modulus Erm30 GPaPoissons ratio 0.2Cohesion c4.3MPaFriction angle 55Tensile strength t0.5MPaThe Youngs modulus of each zone was varied randomly about a mean value of 30GPa with specified variation chosen from a triangular distribution with a maximum deviation of 20% of the mean value. The tensile strength was set to eliminate tensile failure, since the objective of the analysis was to promote the formation of shear bands representing strike-slip faults. The initial state of stress was set at the yield point of the material with a high mean stress of 25.0MPa, major and minor principal stresses of 47.9MPa and 2.1MPa, respectively (compressive stresses positive). This avoided unnecessary cycling of the model from an unstressed elastic state, and avoided tensile failure. Note that these yield stresses are based on the rock mass properties shown in Table 1. The rock mass strength implied by the stress measurements shown in Fig. 3 applies to a larger scale and is not related to the engineering scale rock mass strength. However, the general conclusions from the results of these models do not depend on these details.The prescription of appropriate boundary conditions was complex. In the crust, deformation is generally some combination of pure and simple shear. The proportion of each component of deformation is significant, as this determines the orientation of the principal stresses in the elastic regime. Using the definition adopted by McKinnon and Garrido de la Barra (1998), shown in Fig. 6, the orientation of the major principal stress for a particular combination of pure and simple shear is:where is measured clockwise from the plane of simple shear displacement to the direction of the major principal stress (Fig. 5). Using an equal contribution of both pure and simple shear, i.e. = 45, the initial orientation of the major principal stress was computed as= 76.7. Boundary gridpoint velocity vectors reproducing the initial state of stress and deformation are shown in Fig. 6. Gridpoint velocities were continuously adjusted such that they did not exceed a value of 110-5 m/step. This velocity corresponds approximately to the maximum fraction of the gridpoint spacing that can occur in one time step. By experimentation, this velocity was found to result in a small and controllable unbalanced forces at gridpoints.Fig.5.Definition used for boundary displacement when combining pure and simple shear, and the resultingorientation of principal stresses (from McKinnon and Garrido de la Barra (1998)Fig.6.Boundary displacement vectors used to reproduce equal proportions of pure and simple sheardeformation5 Modelling ResultsFour aspects of the modelled systems were of interest: fault evolution, characteristics of seismicity, triggering, and characteristics of the stress field in between faults.5.1 Fault FormationThe fault pattern that developed in the modelled system, shown in Fig. 7(a), was similar in orientation to classical Riedel conjugate shears, i.e. close to the theoretical Mohr-Coulomb fracture angle of 45 /2 from the orientation of the major principal stress. Faults shown in this figure are represented by contours of accumulated shear strain.Following initial fault formation, deformation in the model was accommodated largely on existing faults in preference to formation of new faults. Faults in an otherwise intact material represent weakness planes along which slip occurs more easily than formation of new faults. However, continuous slip caused fault terminations to interact with the softer annulus material surrounding the model. Resistance to fault slip gradually increased to the point where boundary deformation was accommodated more easily by formation of new faults. Deformation of the system therefore followed a complex pattern involving both slip on existing fault surfaces, and formation and slip on new faults. The resulting relatively periodic spacing of faults that evolved over time tended to distribute boundary deformation throughout the system.5.3 Influence of Fault Slip on Stress FieldIn addition to changes in the stress field within fault zones (shear bands), stresses are also modified outside of faults in the form of principal stress axis rotations plus magnitude changes. Additional models showed that the effect on stress field was more pronounced as fault strength-loss was increased. More quantitatively, a history of mean major principal stress orientation (tve clockwise from east) from this model is shown in Fig. 12(a), and the corresponding change in standard deviation of the orientation is shown in Fig. 12(b). With increasing deformation and fault slip, the stress field becomes increasingly disordered.These results were used to interpret the scatter in principal stress directions in the Sudbury Structure measurements, shown in Fig. 1. The larger scatter in the sub-horizontal principal stresses compared to the minor (vertical) principal stress can be partially explained by the effect of faults. If the sub-vertical late-stage strike-slip faults are currently active in their accommodation of regional tectonic deformation, displacement along those features would lead to scatter in the orientation of stress components in the orthogonal plane, which for strike-slip faults is the sub-horizontal plane, (see Fig. 1(b). Although scatter in the orientation of principal stresses is common with overcoring stress measurements, this source of error would be equally distributed amongst principal stress components. The significantly higher scatter in the horizontally as opposed to vertically oriented components of the stress field is consistent with the interpretation of marginally stable strike-slip faults.中文译文采矿活动诱发微震SD麦金农加拿大安大略省皇后大学 采矿工程系摘要萨德伯里矿区微震观测数据和地面控制问题显示，后期垂直走向断层对采矿活动诱发的微小应力变化很敏感。该地区有限的应力测量表明，该地区构造处于边际稳定状态。开发数值连续模型分析这种构造行为。在这个模型中，剪应变局部（断层）变化，这样，断层系，应力，和边界变形之间有密切联系。在这些系统中断层滑移是很少的，并共同遵守Gutenberg-Richter频率关系。结果表明，对该系统做相对较小的干扰即可在远离干扰源的地方触发地震，用常规数值应力分析方法不能解释这个现象。该系统应力场最初统一的方向越来越混乱。这就是为什么许多应力测量数据是离散的。基于这些结果，对影响矿井应力测量程序和断层的数值稳定性分析的因素进行了讨论。关键词：触发，微震，应力分析，冲击矿压，断层稳定性。1 导言在深部硬岩矿井中大部分的矿震发生在开采边界附近。这些事件都与采矿引起的应力变化有关，应力变化导致完整岩体的破断或沿原有裂隙滑移。留设保护煤柱等使应力高度集中，这极易诱发矿震。测试岩体强度，数值应力分析报告可以用来预测裂隙程度和诱发矿震的危险性（Beck等人，1997；Potvin和Hudyma， 2001；Beck和Brady，2002）。采掘工作面附近典型的矿震包括生产爆炸（这导致应力场的急剧变化）引起的一系列事故，紧接着是持续数小时至数天的衰减期。这类矿震发生频率具规律性，且位置在采掘工作面附近，因此可解决此采矿问题。远离采掘工作面的地方也常发生一些矿震，这似乎在时间空间上都与采矿活动无关。在远离采掘工作面数百米远的地点也有矿震活动的记录。在震源的了解基础上，人们早就认识到，这是由于断层，岩脉，接触带等构造上的滑移导致（(Smith等人，1974；Gay等人，1984）。虽然采掘工作面附近的微震活动数量远超过远离工作面的，但后者对采矿活动影响较大，因为它们往往具有较大的规模，增加了冲击矿压的破坏性。但现在不能预测其发生位置和规模，因此煤矿应加大力度控制潜在的冲击矿压，而不仅仅针对采掘工作面附近发生的微震。由于大多数矿井的综合因素和地质环境还有编码的实用性，数值应力分析是大多数采矿稳定性分析的首选工具。然而，存在断层或其他地质构造时，应用数值模拟研究矿震只有一部分获得了成功。特别是，数值应力分析一直无法解释发生在远离采掘工作面的微震活动。由于数值应力分析在采矿业应用广泛，这就要求我们制定一种方法，可以同时解释这两种采矿微震。本文的目的是探讨发生在远离采掘工作面的矿震起因，并探究应用数值应力分析解决此问题的方法。正如预期，这类矿震深受矿井所处地质环境的影响。模拟远离采掘工作面的矿震数据取自萨德伯里矿区，这是加拿大安大略省一个采矿活动的密集区域。2 萨德伯里构造带的构造地质学和应力状态根据各种应力测试，中北美大陆应力场的水平最大主应力方向约为北偏东（Zoback，1980；Zoback，1992）。Arjang（1991）根据加拿大东部矿区的应力测试结果发现了一个类似的结论。应用一个更大的应力测量数据库，Arjang和Herget（1997）指出，虽然存在一定的相似因素，如最大主应力方向和水平应力与垂直应力的比值，横向应力非常分散。加拿大东部提供的相当大一部分应力测量数据，是在萨德伯里构造带的矿区测得的。波兰人在这些矿区的主应力测量结果主要分布在预测图的下半部，如图1（a）所示。图1（a）所用的测量数据出自Arjang（1998）文章，和Inco公司（Galbraith，2002）未公布数据。最小主应力通常相互垂直，这里显示的方向最不分散，最大主应力和中间者大致相互平行，相比于最小主应力方向又较大的变化。如图1（b）所示，图中显示与最小主应力夹角在20范围内的数据。如果在强度空间上显示，如图2所示，应力数据存在线性关系(R2= 085)。在两个相互平行的主应力之间也存在很强的线性相关性(R2= 070)，如图3所示。由图知，应力状态强度有限，类似于实验室压缩试验所得的强度包络线。如果应力不在一个极限平衡状态，他们不需确定这样一条包络线。由于在岩体结构中，大规模不连续构造形成了最薄弱的环节，这种相互作用反映了应力状态和断层强度之间存在极限平衡关系。主应力（那些主应力方向预期不同的断层，常用地质方法测量）的Anderson等级（Anderson，1951）表明，冲断层最可能处在边际稳定的状态。根据Anderson理论，这就导致了不同类型的断层之间存在明显的差异，例如，一个冲断层的应力测量，后期滑移断层和裂隙明显处在边际稳定状态。Cochrane（1989）通过研究古生代中期大陆应力测量范围内的水压致裂应力测量数据，解决这个问题。Haimson和Doe(1983)指出，在深部垂直应力大于最小水平主应力。这表明走向滑移断层产生于深度和随之减少了最小主应力的垂直应力。加拿大东部地质构造和应力状态之间的关系，由于冰河后期的反复作用而变得复杂，这表明新的构造应力场与构造不完全一致。然而，应力测量数据有限性和后期断层和裂隙对采矿活动引起的应力变化的敏感性可以显示，这些地质构造由于区域构造应力而处在边际稳定状态，而不是其它系统。图1 主应力投影图，a萨德伯里区域应力测量数据，b与最小主应力夹角为20的数据图2 萨德伯里地区测得最大主，最小主应力 图3 萨德伯里地区测得相互平行的主应力 所有应力都是负的（压应力） 所有应力都是负的（压应力）这种地址结构背景，为应用数值应力分析方法研究远离采掘工作面发生的矿震提供了基础。Cochrane (1989)的观点显示，地质构造并非都易受采矿活动引起的应力变化的影响，在萨德伯里构造带中，最主要的构造是相互垂直的后期走向滑移断层。虽然此系统中其他断层的重要性现在还不知道，将重点放在相互垂直的断层上，可将数值简化到二维空间，下面将详细说明。3 数值应力分析法在采矿诱发微震的应用用来评价采矿活动引起的断层滑移的标准已很好的证明(CAMIRO，1997)。而数值应力分析编码和方法（连续，不连续，边界元素等）的选择可能会有所不同，评价断层稳定性的原则依然相同。分析步骤以下：采矿前应力状态初始化，计算由采矿产生的新应力状态，分析断层表面的应力和剪应力，与制定强度比较，估计产生滑移的可能性。一旦计算出滑移面积，即可预知矿震时间及其强度(CAMIRO，1997)。如果应用弹性连续